-------
Tfl -Tlniiial
44" Bed Depth
Probe Position 32"
ALove Grid
1.6
t.4
1.2
1.0
Time (sec)
Figur* 16. Co .ipcraon of T*mpcr*turซ-Time Trcon M Position
32" Above Grid For Fmnod ซnd Unlmnwi Tuba in
44" Settled Bed at 1.3. and 5 Atmospheres.
176
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a. For a bed height of 27". the 18" heat exchanger bundle yielded signifi.:antly
higher heat transfer coefficients than the 6'2" unfimed U tube for approxi-
mately the sane wetted surface area.
b. To recover the sane heat transfer coefficient, the long unfirmed tube re-
quired twice the bed material at pressurized conditions.
c. For all conditions, the heat transfer coefficient wss substantially hi;-her
for 7 atmospheres than the value obtained at 1 atmosphere.
d. Figures 10a and lOb indicate sore of the neasurenents made with local
instantaneous heat transfer probes. These important results substantiate
che overall results obtained by showing the local coefficient distribution
as a function of height within the bed. For the 27" bed (Fig. 10a> the
maxima heat transfer coefficient is obtained near the bottom of the bed.
This indicates the reason for the higher heat transfer coefficient obtained
for the 18" bundle in (!]. For the W bed (see Fig. lOb) the heat transfer
coefficient is quite low on the bottoo of the bed (due to poor circulation)
and a maxima is reached at almost 27" above the distributor, follrits of lav. an^ short heat
crs and finned and unfimed tubes. The following mmarizcs our conclusions:
a. The effect of pressurizb*; a fluidi^ed bed frcn 1 to 7 atrmphvrvK can in-
crease the heat tr.insfer coi-fficicnt jy as nuch as 73". both lor s:ปrt anu
long heat exchanger configurations.
b. Results obtained previously fir shallow bed.--, whereby the ovt-r.ill !ic.it trans-
fer coefficient was increased by 25T1. by a nonuniforra gas velocity .it the
distribution plate, do not apply for deep beds. The parabolic velocity
distribution seers to lose its effectiveness in deep beds with respect to that
of a uniform pas distribution due to the large pressure gradients associated
wilhin deep beds which dorp initial profile distributions.
c. Heat exchanger configurations in deep beds experience a highly nomffiiform
local heat '.ransfer coefficient, as a function 01 bed height. A leu- heat
transfer coefficient nt the bottoa section of the bed. resulting from a poor
circulation of particles, is followed by an optiian heat transfer coefficient
in the middle of the bed (which corresponds to vnlues cbtainabli- in a slullcv:
bed) which is then followed by a region of ;ow heat transfer at tic top of Uu-
bed (which corresponds to a region of high jorosity).
d. For the conditions tested here, firmed tubing did not contribute to an increase
in heat extraction frin pressurized fluidizid beds compared to unfimed tuhinj-..
It is believed that this will probably also be true for other finned tubing
configurations, and for different sized oarticles. Kmn Uie present results
it appears that on both technical and ei.ona.iic croundj.. firmrd cubing should
not be utilized in the expanded bed region of a fluldi;xd facility.
177
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Settled
Grid
Settled
Bed Hetql
o>
Settled
eป
Settled
2
~
It
rs
"'_.
.
r*
ซ 17. Schematre R^irtunution of Oacp Bed With Long Vtrtiol
Hot Exchanger Configuration and Suit* of Shallow Bซdi
(Hut Eachanga Surface Ann and Partidi Volume* An Equal).
178
-------
e. For a given long unfinned cube bundle configuration, and with a deep settled
bed (say of 57' in our facility) an optirun heat extraction probably would
be obtained by havir,g three separate tube bundles connected in series, each
having its own grid and each having a shallow bed one-third the height of the
deep bed. Such a configuration is shown in Fig. 17. It is felt that the
second schene depicted will yield heat transfer coefficients significantly
higher than the long tube hurdle with a deep bed. Such a scheme does not re-
quire an increase in the mct>er of coal feed points. All the coal could still
be burnt at the bottom, the only additional requirement is that of additional
grids in the U-d to facilitate particle circulation and to keep the local
porosity everywhere near the value which optimizes the heat transfer coefficient
(an : on the order of .6)
f. Other heat exchanger arrangements such as those discussed in [2]vherehori-
zontal finned tubing is placed in the upper portion of the bed (v*iere the
oorosity is very high) as an econonizer rany also yield additional heat
extraction.
VI. RETERQJCES
1. Zakkay, V.. Miller. C.. and Brentan, A., "Hc.it Transfer Characteristics in a Kluidized
Bed Packed with lie.it Exchangers." paper presented at Combustion Institute Uestem States
Section 1977 Fall Meeting on Catalytic Ccrfcustion/Fluidized Bed Conbustion. 17-18 October
1977. Stanford. California.
2. ZakXay. V.. Miller. G.. and Panunzio. S., "The Use of Fluidi:'.ed Beds for Heating Air
for Wind Tunnels," paper presented at the Wth Semiannual t-t-eting of the Supersonic
Tumel Assocl/'ion. SeptenbcT 14-15, 1977. Toulouse, Franco.
179
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QUESTIONS/RESPONSES/COMMENTS
ROBERT BROOKS, CHAIRMAN: Are there questions?
MR. CATIPOVIC: My name is Nick Catipovic from Oregon State
University, and we are doing some similar heat transfer studies. I
would have a question regarding the transient heat transfer measure-
ments which you showed, with the instantaneous heat transfer coeffi-
cient oscillating between zero and 100. What were the particle
sizes?
DR. ZAKKAY: The particle size was l/16th inch.
MR. CATIPOVIC: 1/16?
DR. ZAKKAY: Between 1/16 and 1/8, actually.
MR. CATIPOVIC: And what was the minimum fluidization velocity?
DR. ZAKKAY: I think it was about 2-1/2 to three feet per second.
MR. CATIPOVIC: Okay. 2-1/2 to 3 ft/sec. And when the bubble
engulfs your probe, you get a coefficient of zero.
DR. ZAKKAY: You don't get exactly zero. You get the heat
transfer from the gas. That's all. Thus heat transfer is much lowe1*
than the heat transfer from the emulsion. It's about an order of
magnitude lower. /
MR. CATIPOVIC: Yes. But the air is going at a velccUy of
maybe six feet per second or more.
DR. ZAKKAY: Yes. But if I expand the scale, you know, then you
would be able to see the heat transfer from the bubble as it passes
the probe. You don't get zero; you never get zero. But, in comparison
to the heat transfer particles, you can consider it practically insigni-
ficant..
MR. CATIPOVIC: Okay. The reason I am asking you is because we
have some very similar results, and for 4-millimeter limestone or
dolomite, our coefficients oscillate from say, 20 to 50 Btu's.
DR. ZAKKAY: 20 to 50. Yes. Is this atniospheric pressure?
MR. CATIPOVIC: Atmospheric.
DR. ZAKKAY: Well, we seem to see values considerably below 20.
180
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SPEAKER (not Identified): Could ycu tell rce what is the differ-
ence between a shallow and a deep bed? You have a 27-inch bed which
you call shallow, and a 5[-i ci bed which you call deep.
DR. ZAKKAY: i.'el 1, I think this is actually a characteristic of
the bed itself, of the diameter of the bed. The proper parameters
bed height to bed di^neier, I believe.
SPEAKER: Do you have any such statement, that this ratio is
smaller, then this is shallow, and
DR. ZAKKAY: I woild say not right now; but I would guess that a
H/D of 1 to 2, I would call a shallow bed; for H/D ป1, this would
probably be a deep bed.
DR. BAR-COHEN: Bar-Cohen from HIT. Could you comment on the
size and frequency of the bubbles observed and the impact on the
heat transfer rate at the ti-be wall?
SPEAKER (not identified): Well, I think you nay well be right;
but I think there is another feature which perhaps ougltt to be taken
into account. There has been some work published recently about the
effect of higher pressure on the convective component, when you have
large particles.
DR. ZAKKAY: The convective component of the gas or the particles?
SPEAKER: Of the heat transfer in the gas between the particles;
as you get larger particles, of course, that becomes more important,
and it becomes mor.ป important when you get hiy.ier pressures.
DR. ZAKKAY: I see.
SPEAKER: So I don't think one could explain your results on
that alone; I think you may well be right about the reduced bubble
size.
DR. ZAKKAY: Well, we are trying to improve our in-bed diagnos-
tics to determine exac*ly wnat is happening. And I know the BCURA
results of the effects of heat transfer due to pressure do not agree
with ours. They did not find this enhancement effect. Of course,
they had a very packed bed with horizontal tubes; and probably the
horizontal tubes suppress the bi.bble formation considerably. There-
fore, when they worked at higher pressure, there were no bubbles to
suppress and thus the heat transfer was unchanged.
181
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SPEAKER: I have another comment, really, by which perhaps you
might be able to explain some of your other results about the bottom
section of the bed not giving high heat transfer. We have noticed
that, particularly with large particles, when you get a fairly deep
bed, you get a phenomenon which has been described as "solid slugging",
where you get horizontal voids across the whole of the cross-section.
Now these break down further up into the bed into wall slugs, which
give you very good particle circulation; and it may be that's what's
happening.
DR. ZAKKAY: But I didn't get any slugging at the bottom of the
bed. In fact, our observations indicate that the bottom of the bed
was only percolating, the air was just going through a porous media.
Finally at a certain height, the bed acts like a fluidized bed. But
at the bottom, there was very little movement.
SPEAKER: Yes. We agree, I think. Okay; thank you.
ROBERT BROOKS, CHAIRMAN: Yes, next?
MR. WEBB: Ralph Webb, Penn State University. "Did you account
for fin efficiency, in your heat transfer coefficient?"
DR. ZAKKAY: The fin efficiency?
MR. WEBB: You are transferring heat in this fin, so there must
be a temperature drop due to conduction in the fin material?
DR. ZAKKAY: I'm measuring the Q-dot, actually. I am measuring
how much heat I am extracting and comparing one condition to the other
with the same temperature distribution in the bed. But I am supplying
more heat, actually, at the bottom in order to account for that, so
the bed distribution of temperature is the same for the unfinned as
the finned.
MR. WEBB: All right. If you had a fin of infinite thermal
conductivity, would you get a higher Q-dot.
DR. ZAKKAY: We should, yes; in a gas media but perhaps not in a
fluidized media.
182
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INTRODUCTION
GEORGE WETH, CHAIRMAN: Our next speaker will be Lawrence Golan.
Lawrence got his Bachelor's and his Master's in Mechanical Engineer-
ing at West Virginia University and his Ph.D. at Lehigh University.
He is the Senior Project Engineer at Exxon Research and Engineering
Company. At Exxon, his present responsibilities include technical
supervision of the flow visualization, coking, and heat flux studies
on the FBC task force; so with that, we'll have Dr. Golan.
183
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Industrial Application of Fluidized Bed
Combustion - Single Tube Heat
Transfer Studies
D. C. Cherrington, L. P. Golan, F. G. Hammitt
Exxon Research and Engineering Company
Florham Park, New Jersey
ABST ACT
The purposes of this paper are (1) to review the objectives and status of
the DOE supported program on FBC process heaters underway at Exxon and (2) to present
some preliminary data from the program on heat transfer coefficients tor large
horizontal tubes immersed in a fluidized bed of large limestone particles.
The goal of the Exxon/DOE FBC program is to make a thorough technical and
economic assessment of the FBC concept for refinery and chemical plant fired process
heaters. The underlying philosophy of the program is to build on the developments
of the boiler oriented programs, concentrating on areas of difference between boilers
and process heaters. Two significant areas are being investigated - effects of tube
size and hydrocarbon coking. The R&D work in progress includes ambient air fluidiza-
tion studies on tube bundles containing tubes from two to six inches in diameter,
coking studies of a hydrocarbon process stream, and peripheral heat flux measurements
on tubes immersed in a fluidized bed.
An experimental study where the heat transfer coefficient for large, single
horizontal tubes immersed in a fluidized bed was measured has been completed. The
heat transfer experiments were conducted on 2 inch diameter tubes with 390 L*, 1000 u
spherical glass beads, and 2, 4. and 6 inch diameter tubes with a blend of limestone
particles ranging in size from 200-4000 ป. The measured heat transfer coefficients
displayed a dependence on particle diameter, tube diameter, and fluidization velocity.
The heat transfer coefficient for the limestone blend exhibited less dependence on
the fluidixatior: velocity than the glass beads. Visual observation of the solids
movement around the tube circumference and the peripherally measured heat transfer
coefficients indicate at least three different heat transfer zones on the tube wall.
PART I: PROGRAM OBJECTIVES AND SUMMARY OF PROGRESS TO DATE
During the past decade. r..any agencies have been evaluating the use of
fluidized bed combustors as a Eeans of achieving more efficient fuel consumption and
for utilizing coal and other lover quality fuels in a more environmentally acceptable
manner. These investigations have had as their end application electric power
generation. The industrial sector where nearly 1/3 of the total United States energy
is consumed was not receiving attention. Primarily for this reason, the Energy
Research and Development Administration (now the Department of Energy) has awarded
contracts for developing Fluidized Bed Combustion (FBC) for Industrial Applications.
The goal of the Exxon FBC program is to make a complete technical and economic assess-
ment of the FBC concept for refinery and chemical plant fired process heaters. This
portion of the paper contains an outline of the overall Exxon program, including a
status report of the ongoing R&D activity.
Objectives and Scope of Work
The purpose of this program is to extend the state-of-the-art of fluidized
bed coal combustion, which at present, addresses the generation of steam to applica-
tions where oil passing through immersed tubes in the bed will receive heat and be
heated to a required condition. This purpose will be achieved by the successful
completion of the following program objectives:
a. To conduct an R&D program necessary to provide the engineering data and know-
how for designing a fluidized bed process heater.
184
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b. To conduct an economic analysis necessary to evaluate the economic attractiveness
of fluidized bed combustion for indirect fired process heater applications.
c. To demonstrate the operation of a coal fired fluidized bed heater in an actual
refinery environment for an extended period of time.
d. To prepare a complete Design Specification and Control Cost Estimate for a
commercial sized fluidized bed coal fired process heater.
The basic approach to be followed in pursuing the objectives of this program
will be to build on the fluidized bed technology that is now available and under
development by others in the related area of fluidized bed boiler applications. Effort
in this program will be concentrated on doing the incremental work necessary to extra-
polate the boiler oriented technology to refinery and petrochemical plant type indirect
fired process heaters. The areas of technology common to both steam generating boilers
and process heaters will not intentionally be advanced by this program. However, the
state-of-the-art and the results of complimentary programs in the boiler area will be
used in the overall technical and economic assessment of potential fluidized bed process
heater applications.
The two principle areas of technology that have been identified as being
peculiar to process heater applications and which are not being addressed in the on-
going boiler oriented programs concern the effects of tube size and hydrocarbon coking.
These two areas will be investigates in this program.
Indirect fired process heater tubes are conventinally two to five times
larger in diameter than boiler tubes. A typical crude oil heater, for example, may
have a multitude of 4" to 8" diameter tubes in the heat pick-up zones, as contrasted
to the 1" to 2" diameter tubes normally used in steam boilers. The effect that these
larger tubes will have on fluidization characteristics and definition of the optimum
or acceptable configuration of a tube bundle immersed within a fluidized bed must be
investigated.
Similarly, the parameters affecting hydrocarbon coking must be investigated.
When heating a hydrocarbon to 600"F+ (as required for separation by distillation or
other typical processes), some degradation ot the oil and coke laydown on the inside
tube wall is unavoidable. The rate of coke laydowr. is affected primarily by the
temperature of the hydrocarbon film on th
-------
Summary of Progress to Date
The program got underway July 1. 1976. The program effort to date has been
devoted to the design and construction of the three major laboratory units that will
be used to generate most of the Phase 1 program data. The first of these units is
the Two-Dimensional Flow Visualization Unit. This is an atmospheric pressure, trans-
parent test chamber where fluidization and mixing characteristics of a fluidized bed
containing immersed tubes can be visually observed and quantitatively measured. A
schematic of the facility is shown in Figure 1. (This unit is described in Part II
of this paper.) The unit construction was completed and testing comr/'.nced in June,
1977. To date, five different bundle configurations have been rested and evaluated.
The first consisted of 2-inch diamter tubes on 3-diameter ceuter-to-cซnter spacing.
This configuration was identical to the bundle array installed in the Rivesville
demonstration boiler and was intended to establish a baseline of performance against
which comparisons could be made with data as it becomes available from that unit.
These tests were followed in cum by bundle configurations which were
designed to evaluate the effects on fluidization performance of varying tube size
and spacing. The additional bundles tested to date were all 6-inch diameter tubes
but arranged on 2-diameter, 3-diameter and 4-diameter cencer-to-center spacing using
a scaleup of the Rivesville vertical and diagonal proportional spacing. One tube
bundle on 2-diameter equilateral triangle pitch was also tested.
In conjunction with the fluidization performance testing, some conductive/
convective heat transfer data were obtained on each of the bundle configurations.
Variations in heat transfer coefficients as a function of peripheral tube surface
orientation and tube location in the bundle have been determined. In addition, heat
transfer data on single isolated tuhes immersed in a fluidized bed have been measured
using a range of tube sizes and bed materials to determine what affects the presence
of adjacent tubes and varying bed materials (particularly bed particle size) have on
heat transfer characteristics. These data are particularly useful in comparing the
program results with data reported by other investigators who predominantly used
single tubes and/or relatively fine bed materials. These single tube data are
reported in Part II of this paper. It is expected that, all aspects of the Flow
Visualization portion of this program will be completed during December, 1977.
The second laboratory unit is the Process Stream Coking Unit (a simplified
flow plan is shown on Figure 2). The goal of the testing conducted in this unit will
be to determine what effect the high fluidized bed heat flux levels will have on the
coking rate of a hydrocarbon process stream.
The major pieces of equipment in the Coking Unit are: pump (1000 B/D,
650 psig), four single tube heat exchangers (0.6 inch I.D., 9 ft. long), four 50 kW
electric heaters, a 2.8 M Btu/hr gas fired crude preheater, and necessary instrumenta-
tion. The coking tests must be carried out in a refinery location because of the
crude handling facilities required and the volume of crude involved (1000 B/D).
The unit is now completely assembled at the test site, Exxon's Bayway Refinery,
Linden, New Jersey. Final pre-startup checkout of the unit is now underway with
testing expected to begin in December, 1977.
The hydrocarbon feed to the coking unit will be a slip stream from a
refinery pipestill process stream. During a test, the hydrocarbon feed to the four
exchangers will come from a sinrle point so that all exchangers receive a feed common
in inlet bulk temperature and composition. The coking tests on these four exchangers
will proceed simultaneously. The need to run multiple coking tests concurrently is
dictated by the fact that the crude feed to a pipestill heater changes with time.
By running multiple exchangers in one test, a direct comparison of the influence of
film temperature on coking rate will be observed and the influence of feed type
eliminated. The ability to cross-correlate data between tests will depend on the
number of crude slate changes during a given test.
The test conditions are designed to simulate conventional pipestill heater
operations and those expected for a FB heater. The start of test tubeside mass
velocities expected to be run in the four exchangers are 150, 300, 450, and
600 Ibm/sec ft^. These mass velocities correspond to nominal pipestill operation
186
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DUST COLLECTOR
VENT
AIR SUPPLY
AIR METERING RUN
TUBE BUNDLE
GRID
Figure 1.2-D Flow Visualization Unit
187
-------
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PROCESS
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Figure 2. Process Stream Coking Unit
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(300 Ibm/sec f t2) . pipestill turndown (150 Ibm/sec f t2) , an? two increased mass
velocities possibly needed in a FB combustor to limit the process film temperatures.
The selected feed temperatures represent nominal pipestill heater outlet conditions
(650ฐF), heater inlet conditions (450ฐF), and nominal crude temperature at the
desalter (250ฐF). The start of run heat flux levels on the hot crude cases (650ฐF,
450ฐF) have been selected as 15.000. 30.000 45.000. and 60.000 Btu/hr ft2. Higher
flux levels of 80,000 and 100.000 Btu/hr ft2 are planned for the cold crude tests.
Hot crude tests at the 15.000 3tu/hr ft2 flux level represents the average heat flux
in a conventional pipestill heater, while a 30,000 heat flux represents a peak flux.
Operation at these heat flux levels will permit tie-in with conventional heater coking
rate experience. The higher heat flux levels for both hot and cold crude represent
the heat flux conditions possible in a fluidized bed
The heat exchanger unit will be placed in operation at the designed test
conditions and run until the pressure drop and energy balance data indicate that
coking has progressed to a significant extent. The individual coked exchangers will
be phut down and removed for visual inspection. A test will continue until all
exchangers are coked or until the test period has expired. The entire unit will
then be phut down, all exchangers replaced, and another run begun.
The third laboratory unit will be the High Temperature Heat Flux Unit (see
Figure 3) . This fluidized bed will be operated :.n the temperature range of 1000-
1600 F. The purpose of the studies to be conducted in this unit are to determine
the local heat transfer coefficients (including radiation) and variations in heat
flux at significant positions in the tube bundle submerged in the bed. The tube
bundle to be tested will be the arrangement determined in tne Two Dimensional
Studies to be of potential commercial interest for the design of process heaters.
The heat flux studies will first be conducted without fuel combustion in
the bed. The hot fluidizatrion gas will be produced by combustion of fuel in a
combustor and ducting the gas to the test facility. In this way, fuel distribution
variables will be eliminated as an element of concern in these studies. At a later
stage in the program, solid fuel will be burned directly in the bed.
The fluidized bed will be capable of firing 12 M B';u/hr. The design of
the unit has been completed and approved by DOE. Vendor proposals have been received
for all major equipment components and orders should be released in the near future.
PART II: SINGLE TUBE HEAT TRANSFER STUDIES
Several heat transfer studies have been conducted where measurements o'
heat transfer coefficients from a horizontal tube or to a horizontal tube immerst^
a fluidized bed have been made. A number of correlations have been proposed as pre-
dictive tools, but the results of different correlations give conflicting trends.
References (l."?_- 3-4)1 review these correlations. Reference (1_) points out that there
is very little agreement between the various correlations or between correlations and
data. An uncertainty band of 100% is found between correlations. It is also
suggested that the wide range of results may have been influenced by the scale of
the experimental equipment. These previous studies have limited applicability to
many industrial applications (2_) in that fine or narrow particle ranee bed material
and/or small tube diameters (1 and 2 inch) have been used. In many industrial
applications, large tubes ranging up to 6 inch diameter can be conveniently used.
In addition, large particle bed material with a wide particle size distribution appear
likely. This is especially true for fluidized bed combustors where attrition of the
limestone will naturally cause a wide particle size distrubution to exist in the bed.
For this reason, this study was initiated to experimentally measure tube to
bed heat transfer coefficients in an atmospheric pressure/temperature bed where large
tubes would be exposed to a bed of large particles in test equipment of relatively
large scale. The experimental variables consisted of fluidization velocity (ranging
up to 16 fps) , bed particle siiie (390 u, 1000 u, spherical glass beads and limestone
with a wide size distribution ranging from 200-4000 u), and tube size (2-4-6 inch
diameter) . The limestone particles tended to be quite angular with the large particles
being rectangular in cross section. The volume average particle size of the limestone
was 1000 u .
1. Underlined numbers in parenthesis refer to the references.
189
-------
<ฃ>
O
, Existirj
I
I From Subusk @ (1) & (5) i
Flow Visualization Studies
Process
Fluid
(Water)
Cooling
System
Blower
And
Driver
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1
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Air Meter
Run
Flue Gas
Discharge
Cyclones
Screw Feeder
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Figure 3. High Temperature Heat Flux Unit Flow Plan
-------
Fluldization Facilities
The heat transfer measurements were made by inserting a specially designed
tube (described in the next section) into one of two fluidization units. The larger
fluidization unit was approximately 1 ft. in depth by 7.5 ft. wide by 12 ft. high
(Fig. 1) . The 1 ft. horizontal dimension was selected to minimize wall effects.
The large walls of the unit were constructed of plexiglass so tnat bed behavior could
be visualized. Fluidization air to this unit was provided by an electric motor
driven compressor which supplied up to 10,000 SCFM of air at 10 psig discharge pres-
sure. Air from the compressor v;as directed to the plenum located at the bottom of
the flow visualization unit. The grid used was that recommended for and installed
in the Rivesville FBC (6) boiler. The grid consisted of 1/10 inch diameter holes
located on a triangular~pitch for an open area of approximately 4%. Downstream of
the flow visualization section the fluidization air was cleaned of entrained parti-
cles before being discharged to the atmosphere. Two 29 inch diameter cyclones
designed for a maximum of 100 fps inlet velocity at 10,000 SCFM total air flow rate
operated in parallel with the solids remc/ved in the cyclones returned to the bed.
The second and smaller of the two facilities was an 8 inch circular, column
constructed of plexiglass. A Riversville grid was also installed in the small unit.
Fluidization air to this unit was supplied by a separate blower and metering system.
Heat transfer experiments with glass beads were conducted in the smaller column
while limestone tests were conducted in both units. The grid to tube centerline
was maintained at 7 inches for all glass bead and limestone data collected in the
circular column. The packed bed height was above tube level. In the large unit,
the tube centerline to grid spacing was maintained at 18 inches. The size range
distribution for the limestone as charged to the unit is shown on Figure 4.
Heat Transfer Probe
The heat transfer data was gathered from a specially designed probe
inserted in the fluidized bed zone. Heat transfer measurements in the small fluid-
ization unit were made on the 2 inch diameter tube while 2, 4(2). and 6(2) inch
diameter tubes were tested in the large unit.
The heat transfer probes consisted of specially instrumented plexiglas
tubes which were installed in the fluidized bed zone of the test unit. Each tube
had a 1/4" x 6" x .005'' thick Nichrome strip imbedded flush with the outside tube
surface. One or two 40 BWG iron-constanran loop junction thermocouples were attached
to the under surface of the strip to monitor strip temperature. (See Figure 5 for
assembly detailed.)
The Nichrome strip which had a resistance of approximately 0.2 ohms was
electric resistance heated to a temperature of 30* to 60s above the ambient bed
temperature. The power input required to maintain this differential temperature
between the strip and the bed was monitored and is a direct function of che mean
conductive/convective heat transfer coefficient over the area of the strip. By com-
paring the relative power required to maintain the temperature of the strip at
various locations around the circumference of the tube and from one tube location
to another. A pattern of maldistribution as a function of tube location and surface
orientation was obtained.
The instrumentation for the heater element consisted on a D.C. power source
with a built-in volt and ammeter and a digital temperature indicator. The Nichrome
strips were resistance heated to a temperature of 30 to 60ฐF above the ambient bed
temperature. The heat transferred was therefore from the strip to the cooler bed.
After the system reached equilibrium, the strip temperature was monitored for one
minute with maximum aud minimum temperatures during this time period noted. The
average of the maximum and minimum temperature level was used in evaluating the
average heat transfer coefficients, so that:
3.14 I2 R
S
ave As "save
~4 inch and 6 inch nearest plexiglas equivalent to nominal pipe size or 4.5 inch
and 6.5 inch O.D. respectively.
191
-------
I-*
u>
ro
5000
6000
0.1
5 10 50
Cumulative Wt." Less Than
99.9
Figure 4. Particle Size Distribution of Limestone Bed For "Rivetville Configuration" Tests
-------
vo
00
30 C.i- 0- F it,
T.t-ซ;.-!( Wll
T'Ti i v '"^:: *
.
. 1/1 \ V6 C.
*-N,;*'S'ซ Sv a
1/4- 6- . .03S Tt..4l
Figure 5. Ambient Temperature Test Probe
-------
P = I2RS - El
to assure that .the voltage, amperage, and resistance of the circuit all balanced.
Each heat transfer tube could be indexed or rotated on its axis so that
local heat transfer data could be collected around the circumference of the tube.
Data for all tubes at all velocities was gathered at 45" degree intervals of tube
rotation.
Probe Calibration
In order to compensate for the thei-rcal losses other than those directly
from the strip surface, each heater assembly was calibrated prior to testing. The
calibration war performed by placing the heated probe in a metered air stream and/or
water where the heat transfer coefficient could be calculated with a reasonable
degree of accuracy. By comparing this coefficient with that back calculated from
the measured power input, a correction factor for each strip was generated. Fox
imposed heat transfer coefficients greater than 30 Btu/hr ft2 ฐF, the calibration
experiments including computer modeling of the probe design indicated a probe error
not greater than 157,. The higher the imposed heat transfer coefficient, the greater
the probe accuracy. At an imposed coefficient of 100 Btu/hr ft2 ฐF. probe error was
much less than 57.. Probe to probe correction factors varied only by 2-37=. In all
calibration runs, the measured coefficients were higher than that predicted. All
heat transfer coefficients presented are as measured with no corrections applied
since the coefficients generally are in the 30 Btu/hr ft2 ฐF and higher ranges where
probe accuracy is within other experimental variables.
Single Tub? Tests Run
The following single tube tests were run in either the 8-inch diameter
column or the Two-Dimensional Flow Visualization Unit.
Tests run in 8-inch column:
Test No. 1 - Two-inch diameter tube: 390 u spherical glass bead bed
material. Kluidization velocities u = .34 to 1.6 ft/see.
No. 2 - Two-inch diameter tube; 1000 u spherical glas?; beads: u = 1.8
to 4.0 ft/sec.
No. 3 - Two-inch diameter tube: 200 u - 4000 u limestone (1000 u wt.
avg. size); u = 2.1 to 8.4 ft/sec.
No. 4 - Two-inch diameter tube: 200 u - 1800 u "fertilizer filler"
graded limestone (650 ป wt. avg. size); u= 2.2 to 3.5 fiz/sec.
No. 5 - Two-inch diameter tube; 1000 u - 4000 u "moon mountain grit"
graded limestone (2500 u wt. avg. size); u" 6.3 to 8.0 ft/sec.
Test runs in Two-Dimensional Flow Visualization Unit:
No. 6 - Two-inch diameter tube; 200 u - 4000 u limestone (1000 u wt.
avg. size); u ป 11 to 15.8 ft/sec.
No. 7 - Four-inch diameter tub>; 2CO u - 4000 u limestone; u = 10.5 to
14.6 ft/sec.
No. 8 - Six-inch diameter tube; 200 u - 4000 u limestcne; u = 11.9 to
15.0 ft/sec.
194
-------
No. 9 - Six-inch diameter tube; 200 u - AOOO u limestone (1000 u wt.
avg. size). Bed depths of 15-inch and 21-inch above grid
level, u = 11.4 ft sec.
By analyzing the results from tests in the above matrix, various conclusions
and observations can be made concerning the effects of fluidization velocity, bed
particle size, tube diameter and bed depth. Each of these parameters will be discussed
separately in the following paragraphs.
System Limitations
All single tube tests were run over a range of fluidization velocities -
the range for each test varying depending on the bed particle size and the physical
limitations of the test facilities being used. For example, in tne B-inch
tests with glass beads, maximum fluidization velocity was limited only to the entrain-
ment velocity of the particles being used. With limestone, on the other hand, a
system pressure head limitation restricted the region of investigation to a maximum
of 8 ft/sec or about 3 times u r.
When testing single tubes in rhe Two-Dimensional Flow Visualization Unit,
it was not possible to operate the unit at low fluidization velocities in the range
from 2.4 ft/sec (umj) up to about 8 ft/sec. In this operating range, the entire bed
tended to alternately lift and collapse as a single mass, putting very high oscil-
latary forces on the plexigias walls of the unit. Operating in this mode was judged
tj be inappropriate for the safety of this equipment.
It is interesting to note that when the unit is operated under the same
conditions but with a tube bundle immersed in the bed, the same phenomena is not
observed.
Visual Observations
During the course of running these test? in both the 8-inch column and the
larger Two-Dimensional Flow Visualization Unit, some visual observations were noted
which may help to explain or interpret the data. These observations will therefore
be described before discussing the test results.
Three distinct zones of Folids flow or solids-to-tube contacting were
observed around the periphery of the horizontal iranersed tubes. Since there is
general agreement among investigators that the relatively high heat transfer coef-
ficients in a fluidized bed are a result of the scrubbing action or contacting of
the solids particles on the tube wall, it would follow th.it these different patterns
or zones of flow could be significant to our understanding of the heat transfer
mechanism.
The first zone occurred on the lower portion of the tube - that is. the side
of the tube facing toward the direction of oncoming fluidization gas (identified as
180* position on all accompanying figures). This area has been described by others
as a bubble shrouded zone; however, in this study, it did not appear as a complete
gas pocket. The zone appeared to be lighter than normal bed density with very rapid
density variation occurring. This zone can be contrasted to the second zone located
on top of the tube (0*). in this second zone, the tube was covered by a dense "cap"
of particles. The cap was not stagnant as oscillations of the cap occurred in a
regular side to side sliding motion. In addition to the sliding, frequent bubbles
crossed the upper surface causing solids replacement. This bubbling behavior
regularly occurred only over the center two thirds of the tube, indicating a possible
wall effect. This wall effect was also noted in the size of the cap. When the bed
was defluidized and solids drained from the bed, the cap of material remaining on the
heat transfer tube was observed to be saddle surfaced, the peak of the cap being two
to three times higher near the wall than tube center. In the fluidized condition,
it VAS not possible to measure the cap at the tube center. However, a qualitative
judgement was made as to cap size by brightly back lighting the cap area and comparing
light transmitted to the height of the cap at the wall. Again, the wall cap appeared
195
-------
taller than at the tube center. The location of the third zone varied with fluidiza-
tion velocity, but was normally locateo at approximately the 45-90ฐ position. This
zone alternated between the dense cap and dilute phase, much like that of the
advancing and retreating of ocean surf. As will be seen later, this "surf" con-
sistenly coincided with the area of highest heat transfer.
Average Heat Transfer vs. Fluidization Velocity and Particle Size
The heat transfer data as a function of fluidizacion velocity and particle
size were analyzed on the basis of both overall average coefficients and the effects
on peripherial maldistribution patterns. The overall average coefficients will be
discussed first.
The average overall coefficients measured for the 390 uand 1000 v spherical
glass beads and the 200-4000 u limestone blend are displayed on Figure 6. The data
are plotted against a velocity parameter of
u
umf
which tends to normalize the data for the differences in fluidization characteristics
imposed by differences in particle size. The umf used for the glass beads were
calculated by the Leva method while a nominal Ujnf of 2.4 ft/sec was used for the
limestone blend.
At the higher end of the velocity ranges tested, the limestone data com-
pared very closely with the 1000 u glass bead data. It is interesting to note that
the weight average particle size for the limestone blend was nearly 1000 u (see
Figure 4 for the particle size history of the limestone bed material used in all
tests).
At lower fluidization velocities, the resultant limestone heat transfer
coefficients deviated substantially fron the 1000 u line. At the lowest fluidization
velocity at which data were collected, the limestone and 390 u glass beads had nearly
identical measured heat transfer coefficients. Based on Leva's correlation for
minimum fluidization velocity, much of the larger sized limestone would have senre-
gated out of these low velocities with only particles of about 600 u or smaller
still in a well fluidized state. Under these conditions, this aveiage size t'luidized
particle would have been 340 ป. It is probable that multi-layer fluidization had
occurred as described by Wen.^Z.' Wen's discussions center on a two-particle system;
however, the same general trend should occur in multi-particle systems. Wen's theory
predicts distinctly separated layers of fluidization when
Dp
K >1.3. In these tests, the particle size range was significantly greater
UP2 than 1.3.
It was observed that both the 390 u and 1000 u glass bead heat transfer
coefficients increased rapidly with fluidization velocity while the limestone blend
displayed much less sensitivity to fluidization velocity.
The coefficients measured for the glass spheres followed the inverse
relationship with particle size of
.. . 0.36
have
fr)
This relationship has also been reported by Zabrosky. -'
A few tests with 106 u glass spheres were also run which further confirmed
this inverse relationship over the entire 10:1 particle size range.
196
-------
vo
ฃ 90
o
CM
^ 80
5 70
8
ra
60
t 50
u
o
o
ฃ 40
u.
(/>
z
<
H 30
20
UJ
o
"
Glass Beads
Limestone Blend
(50/50)
Note: Slumoed Bed Height Above Tube;
Tube Diameter 2" O.D.
I
I
1.0 1.5
2.0 2.5
U/Umf
3.0 3.5 4.0 4.5
Figure 6. Heat Transfer Coefficient vs. Bed Particle Size
-------
Local Heat Transfer vs. Fluidization Velocity and Particle Size
The effect of fluidization velocity and bed partible size on local peri-
pheral hefit transfer can be seen from an examination of Figures 7a, 7b, and 7c.
Figures 7a and 7b are polar plots of peripheral heat transfer patterns for the 390 u
and 1000 u glass beads respectively with velocity parameters of u in the range of
1.0 to 4. umf
An inspection of the data indicate that at relatively low fluidization
velocities the highest rate of heat transfer occurred ar the 90ฐ/270ฐ positions
on the tube where particle aggitation was the most vigorous. The coefficient at the
upper surface (0ฐ) increased rapidly with an increase in fluidization velocity. This
corresponds to the observed increase in mobility of the particles located in this
"cap" region. As the fluidization velocity was increased, the heat transfer profile
became more circular and the cusps or flatness in the "cap" zone disappear. Very
little change in heat transfer coefficient occurred at the bottom (180ฐ) of the tube
since this position was consistently contacted by dilute emulsion phase solids. The
highest heat transfer rates at the intermediate and higher velocities were found
approximately at the 45ฐ/315" position with the exact maximum position depending on
velocity. These high coefficients are consistent with the visual observation of the
rapidly oscillating solids layer at the location of the previously described <:surf
line" and "cap". Rotation of the test probe into and out of the surfe line region
caused substantial and very rapid strip temperature variations at constant power
input.
The local heat transfer data shown on Figure 7c were obtained in the same
8-inch column but using the limestone blend bed material. In this case, runs were
made at u ranging from .9 to 3.0. While the data are not conclusive, it would
umf
appear that at u , the maldistribution patterns are less sensitive to fluidiza-
^w-1
tion velocity for the limestone. This performance would be consistent with data
discussed earlier for overall average coefficients and summarized on Figure 6.
Two additional brief tests were run in the 8-inch column to evaluate the
effects of varying the particle size distribution of the limestone bed material. One
test used a quarry graded "fertilizer filler" limestone with a size distribution of
200 u to 1800 u with a weight average size of 650 u. The second test used a "moon
mountain grit" with a size distribution of 1000 ;: to 4000 u and a weight average size
of 2500 ;i. (It is a 50-50 blend of these two grades of material that is used as the
charge material for all tests run in the Two-Dimensional Flow Visualization Unit for
both single tubes and bundle configuration tests).
The results of these two rests are compared on Figure 8 with the previously
noted data on the limestone blend. i'he data appear to follow the same general pattern
observed in the glass bead tests - namely, that heat transfer coefficients increase
with a decrease in weight average psrticle size. No more definitive conclusions are
drawn from these limited data.
Heat Transfer Coefficients vs. Tube Diameter
The effects of tube diameter on overall heat transfer performance was also
investigated. Since the 8-inch test column obviously could not accocmodate tubes
larger than 2-inch diameter, these tests were carried out in the Two-Dimensional Flow
Visualization Bed. Tests were run on single 2-inch. 4-inch, and 6-inch diameter
tubes. As can be seen on Figure 9. the 2-inch diameter tube data blended very well
with the single overlapping data point obtained on the 8-inch column during the lime-
stone bed tests on that unit.
In these tests, heat transfer coefficients were measured at 15ฐ increments
around the tube circumference and averaged to obtain overall heat transfer data for
each tube at each velocity.
198
-------
LOCAL HEAT TRANSFER COEFFICIENT VS. FLUIDIZATION VELOCITY
390/x GLASS BEADS - 2" TUBE DIAMETER
315
ieoฐ
= 4.4
a
Umf
= 2.5
= 1.2
Figure 7a.
199
-------
LOCAL HEAT TRANSFER COEFFICIENT VS. FLUIDIZATION VELOCITY
lOOOpt GLASS BEADS - 2" TUBE DIAMETER
315
45ฐ
ocal
ave
225
*-ซ.!
Umf
Ua
Umf
Ua
= 3.1
= 1.9
Figure 7b.
200
-------
LOCAL HEAT TRANSFER COEFFICIENT VS. FLUIDIZATION VELOCITY
BLENDED LIMESTONE - 2" TUUE DIAMZTER
a
;
mf
a _
Uซf"
= 1.6
315".
'local
225"
135ฐ
Characterization 01 The Blended Limestone Bed As A Function Of Fluidization Velocity
U
_Umf-
1.3
1.6
2.5
3.8
. Bed Material
Fluldized
70
95
100
100
Figure 7c.
Largest Particle Fluidized'50 ,
Particle
2000 /i /800/i
3000 fj. /lOOO/i
Exceeds Max. Particle In Bed/
1100/i
Exceeds Max. Panicle In Bed/
1100 M
201
-------
vu
IT
f$ 80
u.
DC
* 70
= u
^_
CO
> 60
h-
Z
G 50
CZ
g "
8 ฐ 4o
a:
UJ
ii
S? 30
<
a:
K
< 20
UJ
LJ
< 10
oc
<
Q
i ! 1 1 1 1 1 1
Limestone
50/50 Wt
1 1 1 1
'~
Blend
. Moon Mountain
Grit And Fertilizer Filler
Fertilizer
Filler
,
i
4,^^.-** Moon Mountain
S''"'
-^x^
Grit ;
1
' Limestone Particle Size Distribution
Wt. Ave. Size
~ Fertilizer Filler 650 M
Moon Mountain Grit 2500 n
50/50 Blend 1000 M
Size Range
200 - 2500
1800 - 4000
200 - 4000
Note: Slumped Bed Height Above Tube
u-
1 1 1 II 1 1 1
0123 4567 8
1 1 1 1 1
9 10 11 1? 13 14 If
ACTUAL FLUIDIZATION VELOCITY, U. (fns)
Figure 8. Heat Transfer Coefficient Limestone vi. Particle Size
-------
90
oJ 80
i
LL.
(g
ง70
5
cj
_ป
-
5 50
0
n.
JNป IL
O uj
t*> o 40
0
a:
UJ
| 30
a;
i-
K 20
UJ
X
UJ
o 10
a
UJ
2
I 1 1 1 1 1 1 1 1 1 1 1 !
6" Nominal
Pipe
4" Nominal /
Pipe /
L..~Lr"~~~
-.*"""" """"-
' t
* s \
*
f'" ,2" O.D. Tube
+ '
8" Column Tests
_ 2" Tube: Slumped
Bed Heiyht Above Tube
Data Collected In 20 Bed; 50/50 Wt. Limestone
Blend; Bed Height Equal To Upper Tube Tangent
- In Slumped Condition
-
1 1 1 ! 1 1 1 1 1 1 1 1 1
3 ซJ 5 6 7 8 9 10 11 12 13 14 15 1
17
ACTUAL FLUIDIZATfON VELOCITY, fps
Figure 9. Heat Transfer Coefficient vs. Tube Siza
-------
The data show an increase in heat transfer coefficient with ar. ir-cr^ase it-
tube diameter. Dependence at constant fluidization velocity followed the relationship.
h o (D-.'i '
ave T
These results are in general agreement with results reported by McLaren f.t al and
Kuruchkin^^ but are not consistent with results reported by several other investigators.
Heat Transfer Coefficient vs. Bed Depth
One experiment was conducted to test the sensitivity of the heat transfer
coefficient to bed depth. In this experiment, a 6-inch diameter tube was positioned
18 inches above the grid, while the defluidized bed depth was 15 inches in one run
and 21 inches in the other. In other words, in the first case, the slumped bed was
even with the bottom tangent of the tube and in the second with the top tangent
line. In both cases, the tube was well inundated in the bed wnen fully fluidized.
The results of the heat transfer measurements are shown on Figure 10. The
measured overall average coefficient was about 77. higher for the deeper bed. Also,
there was a rather significant change in the profile or pattern of heat transfer.
The coefficient at the 45ฐ/315" "surf line" positions was measurably higher for the
shallower bed. With the increase in bed depth, the region of highest heat transfer
shifted to the top of the tube, indicating an increase of particle activity in the
"cap" area.
Summary Observation and Conclusions
The following summary observations and conclusions can be made from an
analysis of the single tube heat transfer tests reported here.
(1) Three distinct zones of particle activity or particle-to-tube contacting
can be observed when a tube is immersed in a fluidized bed. These zones
appear to bear a relationship to the heat transfer coefficient measured at
each respective zone.
(2) The overall average heat transfer coefficient to a tube immersed in a
fluidized bed appears to follow a relationship with particle size of
.36
have
(3) For a bed containing a wide particle size distribution, the heat transfer
coefficient is approximately governed by the weight average particle size of
the portion of the bed that is in a well fluidized state.
(4) The hear transfer pattern in a bed composed of a range of particle sizes is
less sensitive to fluidization velocity than a narrow graded size bed
material.
(5) Patterns of peripheral heat transfer to an immersed tube become more symmetrical
with an increase in fluidization velocity. The local coefficient at the
bottom or 180ฐ position on the tube is affected very little by changes in
fluidization velocity.
2
(6) These experiments indicated a h
-------
LOCAL MEAT TRANSFER COEFFICIENT VS. BED DEPTH
BLENDED LIMESTONE - 6" NOMINAL PIPE SIZE ~ = 4.4
- Curve 1
"local
ave
270"
- Curve 2
180ฐ
Explanatory Notes:
_loc
ave
I, Curve 1 - Slumped Bed Depth 21 Indies
r Ft2 ฐF
have = 68.9 Bm/Hr
Curve 2 - Slumped Bed Depth 15 Inches
have =64.4 Bln/Hr Ft2 ฐF
Curve 1 And 2 - Tube Center Line 18" From Grid
Figure 10.
205
-------
REFERENCES
1. Zabrodsky. S. S., Hydrodynamics and Heat Transfer in Fluidized Bed. M.I.T. Press,
Cambridge, MA. (1966).
2. Bottcrall, J. S. M., Fluid Bed Heat Transfer, Academic Press, New York, 1975.
3. Saxena, S. C., etal, "Modeling of the Fluidized Bed Combustor with Ismersed
Tubes," ERDA Quarterly Report No. FE-1787-6. December 1976.
4. Chen, J. C., "Heat Transfer to Tubes in Fluidized Beds," ASME 76-H5-75, National
Heat Transfer Conference, St. Louis (1976).
5. Cherrington, D. C., Golan, L. P., Halow, J. S., Hammitt, F. G., "Fluidized Bed
Combustion for Industrial Applications - Process Heaters," AIChE Paper No. 1,
17th National Heat Transfer Conference, Salt Lake City. Utah, August 1977.
6. "Multi-Cell Fluidized Bed Boiler Design Construction and Test Program," OCR
Contract No. 14-32-001-1237, Quarterly Report No. 1, July/Sept. 1974.
7. Wen, C. Y., Yu, Y. H., "Mechanics of Fluidization," Chemical Engineering Symposium
Series Vo. 62, 1966.
8. McLaren, J., Williams, D. F., J. Institute of Fuel, Vol. 42, p. 303 (1969).
9. Kuruchkin (as reported in "Modeling of a Fluidized Bed Combustor with Immersed
Tubes," ERDA Quarterly Report FE-1787-6, Sept.-Nov. 1976, Department of Energy
Engineering, Illinois University.)
NOMENCLATURE
A - heat transfer area
D - diameter
E - volts
h - heat transfer coefficient
I - amps
P - power
R - nichrome resistance
T - temperature
u - velocity
ft*
ft
volts
Btu/hr ft41
amps
watts
ohms
"F
fps
SUBSCRIPT
a - actual
ave - average value
bed - denotes bed condition
P - particle
S - nichrome strip
T - tube
mf - minimum fluidization
206
-------
QUESTIONS/RESPONSES/COMMENTS
MR. GOLAN: The first question is from Mr. W. L. James, Fennel 1
Corporation. He asks, "Your fluid bed slide showing the equipment
schematic indicated a gaseous fuel using a orecombustor. You also
showed cyclones and ash disposal equipment, and it's an obvious ques-
tion, where does the ash come from?"
During the time we were discussing that slide, I said the heat
flux unit would be used in two applications. Initially we would burn
only a gaseous fuel, and obviously we wouldn't have any ash at that
time. The cyclones are however required to recycle the entrained
limestone bed material. Then, in a later stage of the program, we
would be adding the coal feeding facilities, and at that time we
would need both the cyclones and the ash disposal equipment.
His second question is: "Is the limestone particle size inter-
mixed?" I don't know exactly what he means by that, but our stone
has a rather wide size distribution of 100 to 560ฐ . We buy two
different blends of limestones. One is the Moon Mountain grit
and the other one is the fertilizer filler. The limestones arrive
separately and are loaded into different drums. At t!ie time we
start a test, we determine how many barrels of each limestone we
are going to put into our 2-D unit; normally a barrel of each, which
gives us about 1,000 pounds limestone in the bed. We then fluidize
the bed and the limestone particles intermix.
The next question is frora Dr. Rao of MITRE: "Why does the heat
transfer coefficient increase with the tube diameter?" That is a
very good question, and I don't have the explanation at this time.
As I said, I really don't know if our work's been conclusive, but for
the set of 16 measurements that we've made, the four-inch tubes
always had higher coefficients than the two-inch, and the six-inch
tubes had higher heat transfer coefficients than the four-inch tubes.
We have tested over a limited velocity range. To confuse the issue
a little more, when we did the experimental work with a bundle of
tubes, the two-inch bundle give us higher heat transfer coefficients
than the six-inch bundle. So far we have no explanation for this
phenomena.
MR. Newby from Westinghouse: "How is the heat transfer coeffi-
cient distributed around the tube circumference?" It's difficult to
describe without using a chalk board, but if we were to visuali.:e the
gas stream coming onto a tube which is facing the gas streamand
label that the 180-degree positionwe'd find at this loca on the
207
-------
coefficient is very insensitive to fluidization velocity. A 30
coefficient at three feet per second would still be 30 at 6, 10,
and 12 fps.
The coefficient on the top surface of the tube (that is, at the
0-degree position), however, was relatively sensitive to the fluidi-
zation velocity. At rather low fluidization velocities, those around
one, 1 to 1-1/2 umf, the particles become defluidized at the zero degree
position and the coefficients are rather low. As the velocity is
increased, that coefficient would increase due to particle activity
at the zero position, and in some cases it became equal to the tube
averaae coefficient.
The coefficients along the sides (the 90-desree positions)
are, at the minimum fluidization condition (2-1/2 feet a second),
very high and the "surf line" is located right at that position.
As fluidization velocity was increased, the surf line would move
up along the tube circumference, and the coefficient, then, would
look, as we affectionately call them, like "bunny ears" at that surf
line. The surf line coefficients could be twice that of the 180
degree coefficients.
Mr. Howell asks: "Is there any indication that a different, grid-
plate-to-tube-bundle distance, other than the 18 inches that you have
been using, would be more advantageous?" That's difficult to say,
because we haven't tested anything besides 18 inches. However, our
test plan includes investigations with varying grid to tube spacing.
Right now we've got the four-inch tube bundle on two-diameter equi-
lateral triangular pitch arrangement at the IP-inch level, with the
50-50 blend of limestone. We will be dropping the tube bundle to
nine inches above the grid to investigate just the question that's
been asked. Lowering this distance should reduce the size of the
bubbles that enter into the tube bundle, and that may be helpful. I
really shouldn't anticipate the results at this time.
Mr. Howell*s second question is: "How do the cold model studies
compare to those run by Foster-Wheeler two years ago? Is there as
much channeling and stagnant locations in the bed?"
When we ran the two-inch-diameter bundle, with a duplicate of
at the 3-D Rivesville spacing, we saw that the bed tended to operate
in what I call an oscillatory type mode. The fluidization gases
would swish to this side of the unit, and then swish back to the
other side, so there would be channels of almost pure gas existing
through the bed. As we increased the velocities to approximately 12
208
-------
foot per second, the tendency was to evacuate the tube bundle of
particles and stack the particles up in the upper tube. This appears
to be a very undesirable mode, because all the heat transfer particles
would be sitting on "to of the tube bundle. In all cases, the bed
appeared to be very Mve and there didn't appear to be any stagnant
zones in the bed.
When we went to our six-inch diameter tubes, this oscillatory
motion or side-to-side n;oMcn didn't exist. The packing of par-
ticles on top of the tube appears to depend upon the horizontal
tube spacing, particle size and fluidization velocity.
Are there any other questions?
CHAIRMAN: Thank you very much.
209
-------
INTRODUCTION
ALRFRT A. JOMKF, CHAIRMAN: The next speaker is Joseph Mei, who
received his bachelor of science in mechanical engineering from Chu
Hi College in Hong Kong in 1962, and his MS in engineering science
from West Virginia University in 1967. He received his Ph.n. in
Aerospace Engineering from West Virginia University in 1973. He
joined the Morgantown Fnergy Research Center in 1975, and was involved
in the development of advanced coal gasification processes. Currently
he is engaged in research and development in fluidized bed combustion.
Mr. Mei.
210
-------
Fluidized-Bed Combustion of Lignite and
Lignite Refuse
J. S. Mei, U. Grimm. R. L. Rice.
andJ.S. Halow
Department of Energy
ABSTRACT Morgantown Energy Research Center
Lignites comprise about 29 percent of the Nation's solid fuel reserves and
constitute a major potential source for future' energy needs. Samples of Texas lignite
from the Wilccx formation have boon investigated at the Morgantown Energy Research
Center, V.S. Department of Energy, to assess this fuel as a potential feed stock for
fluidized-bed combuslors. Combustion tests wen- performed over a wide range of
operating conditions to develop fluidized-bed combustion ( FBC) engineering and emissions
data on this low-quality fuel. Desirable combust ion characteristics observed include:
no clinker formation at bod temperature as high us 2OOO"F. high sulfur retention capa-
bility of the mineral from lignite at bed temperature below lt>OOฐF. and freeboard
burning of carryover less than expected. Results of these tests indicate that lignite
may be a particularly attractive solid fuel for industrial KliC applications. One
potential problem with the use of this fuel is formation of very line ash particles
which are difficult to retain in tho bed. Use of an attrition resistant bed material
was required to maintain bed inventory.
INTRODUCTION
The Morgantown Energy Research Center of the U.S. Department of Energy began a
program in Juno 1976 to examine the application of atcospher i-_""l 1 uidi nod-bed combustion
(AFI1C) technology to low-qua 1 i t y fuels. The purpose of this program is to evaluate the
feasibility oi' burning low-quality fuels in AFBCs. Conbust ion character ist ics of
various low-quality fuels such as anthracite culm, sub-hiluminous coal, and lignite coal
have bi.-t.-n examined in an 18-inch diameter f luidi iv'-b'-d combustor. This paper presents
the results of recent tests of Texas lignite and lignite refuse.
I,ignites comprise approximately 29 percent of iln.- Nation's coal reservesi and
constituto a major potential replacement for oil and natural gas in industrial applica-
tions for steam raisins: and process heating. Lignite reserves, however, have been
virtually untouched due to their relatively low-quality. Ti.-xas. which is the largest
lignite producing state-, still generates almost 90 percent of its electricity from
natural gas in spite of the state's rich lignite deposits. Identifii.u coal resources in
Texas have been assessed^ to be about one percent ปf tho Nation's total. It is esti-
mated that more than half of the Texas coal resources are lignite occurring in three
formations in the lower Tertiary of the Gulf Coastal Plain (Figure I). Tho wilcox
formation contains approximately 80 percent of the licnite resources, while smaller
lignite deposits are IOUIHI in the Yegua and Jackson format ions. Two samples of Texas
lignite from the Wilcox formation were examined to assess this fuel as a potential feed
stock for fluidized-bed comhustors. One sample was from a commercially-mined deposit
termed the "Big Brown" lignite. The second sample. "Bon-hole -6". was obtained fron
large boreholes owned by Dow Chemical a few miles fron the "Big Brown" commercial
operation. Both of these samples came from an area near FairfieJd. Texas, located in
Freestone County.
EXPERIMENTAL FACILITY AND PROCEDURE
The experimeTtal facility at the Morgantown Energy Research Center for fluidized-
bed combustion fuel feasibility studies is shown schenatica1ly in Figure 2. It is
basica*"y a ref ractory-1 ined cycl indric.al combustor having an internal diameter of IS
inchos an the 45-inch high bed region and a 27-inch high expanded freeboard section of
24 inch diameter. The combustor contains a set >f hair-pin-type water cooled bed heat
exchanger tubes located in the bed region and a single pass water cooled heat exchanger
In the expanded freeboard. Coal is metered fr<
-------
Ill Hut
II1C01 IIGNIK
;:SJฅ:3 eillMlMOUS COIl
CISON name
Figure 1. Coal Resources of Texas
additional carbon burnup. Flue gas leaving the combustor was continuously sampled and
analyzed. CO and COg aro determined by non-dispersive infra-red analyzers. A chemi-
lurainescence analyzer was used to determine NO ai.d N'OX. SO2 and total hydrocarbon
analyses are nude by a flame photometric analyzer and a flame lonization analyzer,
respectively, and O2 by a paramagnetic analyzer.
Start-up of the unit is accomplished by preheating the lower part of the con-
bustor to 1500ฐF.with premixed natural gas and air. When the desired temperature level
( 1500ฐF) is reached, sfpproximately 50 pounds of bed material is added at a slow rate
until it is also up to a temperature of about 1500ฐF. When this occurs, fuel feed is
started and continued until the desired bed level is attained. At this point, the
natural gas is shut off and the fire is self-sustaining.
Chemical Analyses of Texas Lignite and Lignite Refuse
The chemical analyses of Texas lignite and lignite refuse on an as-received basis
are given in Table I. A portion of the Borehole #6 lignite was subjected to a water
washing, coal cleaning process and separated into two sizes: lumps which .sere 1 x J-
212
-------
LIMESTONE f -\
, HOPPER ; ''"EL HOPPER
!GH HOPPER
SCREA FEEDtR
ASH
LOCK HOPPER
Fip-..-.e2. Fluidized-Bed Combustor
Table f. Analyses of As-Received Ligr.ite for Fiuidizcd-Bcd Conbustion
Constituent ,
pet .
Total carbon
Hydrogen
Oxygon
Nitrogen
Sul fur
Moisture
Ash
Heating value,
Btu/lb (as
recc i ved )
Texas
"Big, Brown"
55.21
5.00
1-1.98
0.94
1.03
10.65
12.19
9.170
Ma t
Borehole *6
19. 13
1.66
13.88
1.09
1.30
16.42
13.52
8.237
rials Used
ป6 Washed
Lumps
51 . 15
5.83
11 .69
0.94
0.88
21.26
8.25
8,920
=6 Washed
-~=-Jฃirn.US5, ^- -*-.
56.09
5.80
11 .53
1 .08
1.00
15.26
9.24
9,375
=C V.'a.shery
Ref use _
4 1 . 33
4.14
9.75
0.72
2.34
11.13
30.59
6,805
1
inch, and a fines fraction -J x 0 inch. The washery refuse had a size range of 1x0
inch. All fuels were crushed to a -i inch before injection into the FBC. The Borehole
#6 washed fines had the highest heating value, 9400 Btu/lb, while the? lignite refuse
had the lowest heating value of 6800 Btu/lb. Ash content ranged from a high of 31
213
-------
percent in the refuse- to a low of 8 percent in the washed lumps. Moisture content
varied from approximately 11 percent in the Big Brown to 21 percent in the washed
lumps. Tin- sulfur content, as seen from this table, ranged from a low of 0.88 percent
in the borehole t*G washed lumps to a high of 2.3'1 percent in the Borehole -6 washery
refuse.
Operating Conditions
Fluid)xed-bed combustion tests of lignite and lignite refuse were carried out
over a fairly wide range of operating conditions. Those; are given in Table II bolow.
TabJ
-------
rates up to 15 percent. Carbon combustion ef fie iencies also increase with bed tempera-
tures up to I600ฐF. Increasing excess air rate beyond 15 percent or raising bed tem-
perature above 1600ฐF has little effect on carbon combustion efficiency. The effects
of excess air rate, and bed temperature on carbon combustion efficiency are reflected
in decreases in carryover of paniculate carbon and in the carbon monoxid" concentration
in the flue gas at higher excess air rates and bed temperatures. Table III illustrates
these effects at bed temperature of 1-1701' to 1750ฐF and excess air rates of from 3.7
to 7.0 percent.
Table III. Effects of Bed Temperature and Oxygen in Fine GAS on Carbon Monoxide
and Part-iculate Carryover of Unturned Carbon
Run ! Bed Temp. ,
N.ป.-- 1, - .ฐ.r _.
7/3 | 1600
8/7 i 1GOO
7/2 | 1-170
8/8 j 175O
Sup. Vol. .
ft/sec
3
1
3
1
Oxvgen in
Flue Gas,
... .J>v-t = ^ . _
7.00
'1.00
4.50
:\ . 70
Carbon in
2nd Cyclone
Ash, pet.
1 .5
8.82
1C. 20
0.32
Carbun in
Bag Filter
Ashj p,-t.
2.7
2.^2
10.81
1 .41
CO in
Flue Gas
pot......
O.OG
0.17
0.40
0. 16
Sii'l fur Dioxide.' Emissions
Sulfur dioxide emissions in the flue- gas are plotted against the Ca/'S mole ratio
in Figure -1. The mineral constituents of these lignites and lignite refuse materials
contain significant amounts of calcium which is effective in sulfur capture. The cal-
cium tc sulfur iu>le ratio was therefore defined as the calcium in the fuel pi us the
calcium content of the fresh stone divided by the sulfur content of the coal. The test
data of lignite and lignite refuse over a bed temperature range from 1450" to IGOOฐF
were correlated fairly well by a single curve despite variation in superficial gas
velocities, bed depths, bed materials, sorbents and sulfur contents in the fuel feeds.
The results indicate that current EPA sulfur dioxide emission limits can be net with
relatively low sorbent addition rates and in some; cases without sorbent addition. Cal-
cium oxide in the lignite ash reduces or eliminates the need for sorbent -addition. At
bed temperature above 1GOO'>F, no consistent effect of Ca/S mole ratio on sulfur dioxide
(.missions is evident. The sullur retention data is plotted as a function of the Ca/S
mole r::tio in Figure 5. As seen, a 90 percent sulfur retention can often he achieved
at Ca/S mole ratio greater than 1.0 and at operating bed temperatures at or below 1600ฐF.
Again, a single curve was used to correlate the tost data.
It has been previously reported'' that sulfur is retained in tho ash material
while burning western lignites. To confirm the hypothesis that ash participated in
sulfur capture;, a portion of the as-received lignite was tested for its sulfur capture
capability by ashing the coal in a temperature range which characterix.es the normal
f 1 uidizcd-bcd combustor. X-ray fluorescence analysis of the ashes showed that SO;j
content of the ashes decrease's from 15.45 percent at 1600ฐF ashing temperature to'0.36
percent at 1850"F ashing temperature confirmin. the FDC results that the ca'cium in the
ash participates in sulfur capture. Tin.- resul.s also indicate that there rr.ay be a
temperature range where tho lignite can be burned without added sorbents for sulfur
emission control.
During the combustion tests with a pure quartz sand bed material, sulfur dioxide
emissions increased in spite of the addition of sorbent. Observation of lignite ash by
x-ray diffraction and SKM testing indicate an increase of calcium-aluminum-silicate
parallel to the disappearance of the sulfur in the ash. These preliminary results
suggest that quart?, sand bed material may affect the sulfur capture capability of a
composite bed of lignite mineral, sorbent, and inert material. These results, however.
are net conclusive.
Nitrogen Oxide Emissions
Nitrogen oxide emission in the flue gas measured during the combustion tests was
low over the entire range of operating conditions. Values of NOX concentration in the
flue gas ranging from 50 ppm to 600 ppm were observed depending on operating conditions
as well as nitrogen content in the fuel. The fluidized-bcd combustor was operated at a
bed temperature between 1450ฐ to 1750ฐF which is well below the temperature level at
215
-------
4.0
g
2 3.0
CD
^
CM
s
1 2-ฐ
8
5
i , ?
1 . C
o
ce I.C
3
0
1
Ca/S
THE F
-
e
9
- 8 *
- g
a
_ *
9
i
1
N
JEL
9
o
v,
X
j
11111
WITHOUT LIMESTONE ADDITION
OWITH LIMESTONE ADDITION
O
~
O
EPA EMISSION STANDARD
-
ฐ ฐ o
^- 0 ฐ ฐฐ
0
1 ฐ 1 I 1 1
2 3 4561
Ca/S MOLE RATIO
Figure 4. Effect of Ca/S Mole Ratio on Sulfur Dioxide Emissions
(Ca derived from ash and limestone- addition)
which oxidation of atmospheric nitrogen occurs-.*' The only significant -source of
nitrogen oxide was fuund to bo nitric oxide. In Figure 6, nitric oxide emissions are
plotted versus oxygen in flue gas. With SiOo and AloO-^ beds, the nitric oxide level
in the flue gas increases with oxygen. Thป* spread of the data suggests that bed
material may affect the nitric oxide emission. The nitric oxide concentration measure-
ments shown in Table IV indicate that in the bed temperature range of 1450ฐ to 1650ฐF,
NO omission is only slightly dependent upon bed temperature. Above 1650ฐF, however,
nitric oxide decreases with increasing bซ.'d temperature. The reduction in NO concentra-
tion observed at higher temperatures is in contrast to the results obtained by Pope,
Evans and Bobbins5 and Sheffield6 which indicated that the nitric oxide emission
increases with bed temperature. The most likely explanation for the NO results in this
study is an interaction of NO and 50% ln tne bed. As mentioned earlier, sulfur reten-
tion capability of the composite fluidizcd bed decreases as bed temperature increases
above 1600ฐF. The nitric oxide concentration, in turn, likely decreases due to the
effect of sulfur dioxide concentration on the NO reduction reaction.**
216
-------
C WITH LIUESTONE ADDITION
O WITHOUT LIMESTONE ADDITION
3 4
Co/S HOLE RATIO
Figure 5. Effect of Ca/S Mole Ratio on Sulfur Retention
(Ca derived from ash and limestone addition)
Table IV. Effect of Bed Temperature on Nitric Oxide
Run
No.
T-Z_ฑi =;ฃ_-=
871
P/2
8/9
8/10
8/3
8/5
8/11
8/13
Bed Temp. .
QF ^ ^.
1450
1600
1500
1650
1700
1550
1700
1650
O-2 in Flue
GaSj_j)ct.
"""5.50'
5.00
8.50
7.10
3.75
3.00
3.50
3.75
Sup. Vel . .
ft /sec
"-4 -
4
6
6
4
4
4
4.5
X2 in Coal
_pct .
1.01
1.01
1.08
1.08
1.01
1 .01
0.72
0.72
NO in Fluo
.Gas^ jJEm _
180
210
125
150
190
210
100
240
217
-------
Q.
Z
I-
00
10
o
0.8
0.7
S 0.6
CD
tO
i
iii
UJ
Q
X
O
O
a:
0.5
0.4
0.3
0.2
0.1 h
3N_0]_CORRESPONDING TO
EPA EMISSION STANDARD
OVD
1 1 1 1 1
I I i
^X. BED
BED\MAT'L
TEMP 'FX^
1450
1500
1550
1600
1650
1700
1750
Si02
D
O
O
A
O
9
Q
AI203
a
A
i
AI203
USED
a
&
-------
CONCLUSIONS
Based on the test data, several conclusions can be drawn from the present
feasibility studies on combustion of Texas lignite and lignite refuse:
1. With the recycling of primary cyclone ash, 95 percent or better carbon com-
bustion efficiency is attainable.
2. Current EPA sulfur dioxide emission standards can be mot with little and in
some cases no sorbent addition provided that the fluidized-bed combustor operates al
bed temperatures at or below 1600ฐF.
3. A 90 percent sulfur retention is attainable with Ca/S mole ratio equal to
or greater than 1.0 at operating bed temperature below or at 1600ฐF.
4. Nitrogen oxide concentration in the flue gas is low and is dependent upon
the excess oxygen.
5. The high volumetric heat release rates comparable with rates obtained in
burning higher quality coal were found.
6. Based on the combustion test data, the atmospheric fluidtzed-bed combustion
technology offers a promising way to utilize Texas lignite and lignite refuse.
ACKNOWLEDGEMENT
The authors wish to thank Dow Chemical, Texas Division, for providing the lignite
samples.
REFERENCES
1. Hammond. A. L. , Hetz, W. D. , and Maugh II, T. 11., Energy and the Future. American
Association for the Advancement of Science, 1973.-
2. Evan, T. J. and Kaiser. W. R. . "Texas," 1976 Keystone Coal Industry Manual, p. 638.
3. Levone, II. D. . and Hand, J. W. , "Sulfur Stays in the Ash When Lignite Burns," Coal
Mining and Processing. February 1975, pp. 46-48.
4. Westinghouse Research Laboratories, "Evaluation of the Fluid)zed-bed Combustion
Process Summary Report-Envi ronmontal Protection Agency," Archer, D. II., et al,
Pittsburgh, PA. November 1971.
5. Ehrlich. S.. "A Coal Fired Fluidized-Bed Boiler," Institute of Fuel Symposium
Series No. 1: Fluidised Combustion, Internaliona'l Conference, London, UK,
September 1975, pp. C4-1-C4-10.
6. Gibbs, B. M., et al, "Coal Combustion and NO Formation in an Experimental Fluidizcd-
Bed," Institute of Fuel Symposium Series No. 1: Fluidised Combustion, International
Conference, London, UK, September 1975, pp. D6-1-D6-13.
7. Burke. D. P., "FBC May be a Better Way to Burn Coal," Chemical Week, September 22,
1976.
219
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QUESTIONS/RESPONSES/COMMENTS
WAYNE A. McCURDY: 1 believe that Mr. Mei has a question or so.
MR. MEI: I have quite a few questions here as a matter of fact.
The first ons is from Dr. Porter from Energy Resources Company.
"What precaution do you take in fuel (lignite) storage and handling
to prevent auto-ignition?"
Some lignite material I have observed will burn spontaneously if
left in a pile for as short a period as two hours. The lignite we
received from Low Chemical is in drums. We have a cold storage
plant, which is maintained at a certain temperature. The lignite was
stored in the storage plant. We did not observe any auto-ignition.
The second question is from Mr. Shelton Ehrlich, EPRI. The
question has three parts; the first one is, "What is lignite refuse
and give analysis." The lignite washe'ry refuse is the residue after
the coal cleaning process, i.e., the residue coming from the washing
process. Actually most of it is ash. I believe I did show the
chemical analysis of lignite washery refuse. It is about 40 percent
of carbon and approximately 31 percent ash.
The second part is, "Your data did not seem to show 90 percent
of sulfur retention at a calcium to sulfur mole ratio equal to one.
Could you review this?" I recall in rny slides that it did show
a calcium to sulfur mole ratio equal to one or greater than one. We
have several runs which show a 90 or better sulfur retention.
The last part of the question is, the NOX data show a very
high variation with a coal ash bed. "What result was found with
limestone?" Hell, as you recall on the NOX emission plot versus the
excess oxygen, in the flue gas, there is one data point which is the
last test for the lignite refuse using ash as bed material. We found
that the NOX emission is rather high. That's why I mentioned in my
talk that the scatter of the data suggests that the bed material may
effect the NOX emission. And the reason for that is not quite
clear at the present. During the lignite and lignite refuse tests we
did not use limestone as bed material.
I have a question from R.G. Seth, Gilbert Associates, "What kind
of particle size is used in the 18-inch combustor, and what is total
pressure drop across the bed?" The size of the coal particle is
quarter inch by zero. The total pressure drop across the bed is
somewhere around 30 to 35 inches of water.
Another question, frcn C. C. Gentry, Phillips Petroleum Company.
"Was the immersed tube plain or finned?" We use just a plain tube.
220
-------
"What were the tube dimensions and material types?" The diameter
for the bed heat exchange tube is 1/2 inch, schedule 160, type 316
stainless steel. And for the freeboard heat exchanger, the tube
diameter is 3/4 inch.
"What was tube layout spacing?" Well, the tubes were arranged
in a triangular pitch. The horizontal distance between tube centers
is 3 inches, and the vertical distance between tube centers is 2
inches.
"What range of heat transfer coefficient was produced in the
immersed tube? Has a dimensionless correlation been developed? If
so, please indicate from your correlation."
Well, during the lignite run, we did not use the bed heat
exchange tube in order to maintain the bed at certain temperature
levels. But we did use the freeboard heat exchange tube. The
average overall heat transfer coefficient for the freeboard heat
exchange tube is somewhere around 10 Btu per hour per square foot per
degree F. So, we did not have any correlation for the heat transfer
coefficients.
"When will an ERDA (DOE now as of October 1, 1977) report be
issued, title arid report number?" Well, we have not finished the
entire report, but v/e expect to finish it by the end of January. At
the present time, we do not know the number of the report.
The last question is from Mr. M. H. Schwartz, Shell Development
Company. "The extremely high moisture and ash content which charac-
terize coals of the Yequa and Jackson formations suggest thdt fluid
bed combustor may be an appropriate utilization technology. Do you
plan to test such coals?" At the present time, we do not have
such plans to test other types of lignites. He also asks, "Since
the fusion temperatures and softening points for these lignites are
extremely low, did you observe any significant fouling, refractory
attack, or slagging when you ran your combustor at extremely elevated
temperatures near 2000ฐF?"
MR. ME I: We ran the combustor at a bed temperature of 2000ฐF
for only several hours. For the rest of the tests, bed temperature
was ranged from 1450ฐF to 1700ฐF. Are there any more questions?
Thank you.
221
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INTRODUCTION
OFORC.F HFTH, CHAIRMAN: Our next presentation will he by
Mr. Heman Mack with the Rattelle Coltnbus Lahoratories. He received
his deqree in chenical engineering from Cooper Union, and is currently
the program nanager of fluidized bed technology at Battelle. So,
Heman?
222
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Battelle's Multisolid Fluidized-Bed
Combustion Process
H. Nack. K.T.Liu, and
G. W. Felton
Battelle Columbus Laboratories
ABSTRACT
A novel fluidlzed-bed combustion process has been developed and called Xultisolid Huidized-Bed
Combustion (MS-1 EC). Studies in a 6-inch-diaceter cocbustor have demonstrated feasibility of operation
at superficial gas velocities exceeding 30 feet per second with good linestone utilization and con-
bust ion efficiency. Scale-up studies on a 400 pounds of coal per hour cc^bustor/boiler are in progress.
Successful results will lead to a demonstration steam-generator plant of 25,000 pound per hour capacity.
A description of the concept and results of studies in the 6-inch-diaceter bench-scale unit are pre-
sented in this paper.
BACKCKOIAU
About 4 years ago, Battelle initiated a program aimed at advancing the state of the art of
fluidlzcd-bed combustion. A technique was conceived and developed that is called the Multisolid
Fluldized-bed Combustion (.VS-FBC) process. The initial experimental development was done in a 6-incli-
diancter conbustor. Presently, we are studying the process in a rectangular 2-square-foot cross section
subscale experimental combustor/boiler under a L.S. DOE Industrial Applications Demonstration Program.
This program is aiL-_-d toward construction and operation of a 25,000 Ib/hr steam plant at Battelle which
will be our lead boiler.
In this paper, we describe the Multisolid Fluldized-Bed Ccr.bustlun process and present some
results of our work in the 6-inch unit. Shakedown of the subscale unit has been completed and vc are
into the test program. However, definitive results are not yet available.
DESCKIH-IO.N OF Hit KS-FBC PROCESS
Multisolid Fluidized-Bed Combustion cay be viewed as a process in vhlch an entrained fluidizcd
bed is superimposed on a conventional dense fluldized-bed combustion process. The dense flu'dizcd-bcd
process (Figure 1), as related to the MS-FBC, has the following characteristics:
Air is used for bed fluidizatlon and coal combustion; the air velocity is high
30 to 40 ft/sec.
The bed consists of high-specific-gravity material.
Crushed coal is the fuel.
Finely ground limestone is fed into the bed to capture the sulfur.
Sulfated limestone and ash are blown out of the dense bed and are collected downstream.
The entrained fluidlzed bed (Figure 2) consists of inert material that:
Is sculler in size than the dense flul^lzed-bed material
Is separated from the exhaust gases and recycled to the dense bed (in the combustor).
Combining these two fluldized beds yields BatteUc's Multisolid Fluidlzed Bed, shown In Fig-
ure 3.
During operation, the MS-FBC has two distinct zones in the cocbustor column. The bottom zone
is the dense bed region, where the heavy bed material is fluldized and retained. Immediately above the
dense bed region is an entrained bed zone, Consisting of relatively fine particles cf the same or
another bed material. The entrained bed material, which is fed into the dense bed region in the cor.-
bustor, is carried upward by the air stream in the burner column an*i into a gas-solid separator; froo
there, the material is recycled into the coabustor.
223
-------
tCUANIU
IlUt CAS
ASH AND
SUIFATED
LIMfSTONE
DISTRIBUTOR
PLATE
DENSE
BED
MATERIAL
CRUSHED
COAl
HNHV
GROUND
HMtSIONE
Figure 1. Dens* Fluidited Bad
224
-------
ACUANIO
'nut CAS
STEAM
WAIt*
BURNtR
DISTRIBUTOR
PIAII
CrUSHID
COAl
r- IINHY
GROUND
IIMISIONE
AIR
Figur* 2. Entrained Fluidized Bed
225
-------
ป SHAM
DISTRIBUTOR
PLATE
Figure 3. Battelle Multisolid Fluidized Bed
226
-------
Crushed coal and finely ground limestone are also fed into the dense bed region where the con-
bust ion of coal and reaction between the linestone and liberated sulfur dioxide occur, both at the
same tine. Like the entrained bed particles, the coal ash and sulfated limestone particles are blown
upward into the separator. Because the ash and sulfated limestone particles are lighter and smaller
th&R the entrained bed material, nose of these particles remain in the flue-gas stream and are col-
lected in one or more high-efficiency cyclone separators. The flue gas emitted to the air is environ-
mentally acceptable. Tne dry ash plus sulfur-bearing limestone mixture is disposed of o- used for
some non-process purpose.
A few operating features of the MS-FBC should be emphasized:
The combustor contains entrained bed material as well as dense fluidlzed-bed material,
coal, and limestone.
The dense fluidlzed-bed material serves to trap or slow down the escape of other mate-
rials that are blown out of the dense bed at the high (30 to 40 ft/sec) air velocities
of operation.
The heat released in the burner is transferred to the entrained fluidized-bed material
and iaises its temperature to 1600 to 1750 V.
The entrained bed material acts as a heat carrier, giving up it? heat to the stean
tubes located outside of the dense bed.
The entrained bed material is separated from the exhaust stream and returned, at a
controlled rate, to the conbustor.
One method of providing for the use of the heat produced in the burner is to immerse the boiler
tubes in the entrained fluidized-bed region (just above the dense fluidized bed) ar illustrated in
Figure 3. This process is called the "Multisolid Fluidized Bed With integral Boiler".
Here, as the hot entrained bed material and exhaust gases move upward iron the dense bed, they
transfer heat to the boiler tubes, thus converting the incoming water to steam. At the same time, the
entrained bed material and exhaust gases cool to about 650 F -- the stack gas temperature for some
industrial butlers.
An important variation of this process involves the use of an external boiler, as Illustrated
in Figure < The process, called the "Multisolid Fluidized Bed With External Boiler", has the follow-
ing characteristics:
The steam tubes are Immersed in a separate fluidized bed within the external boiler,
not in the conbustor column.
The hot entrained fluidized-bed material (at 1600 to 1750 F) is separated from the
exhaust gases and drops into the external boiler.
The external boiler is, a conventional fluidized bed operated at low velocities of 1 to
2 ft/sec (compared to 30 to 40 ft/sec in the conbustor). Little or no conbustion occurs
in the external boiler. Erosion/corrosion of the boiler tubes is minimized under these
working conditions.
The external boiler Is operated to maximize the heat transfer coefficients, thus
minimizing the tube requirements and size of boiler.
The hot gases leaving the separator (and the external boiler) are cleaned via the ash
and sulfated limestone collector, and move to a separate convective heat exchanger
or economizer where the gas temperature is reduced to a reasonable stack temperature.
About one-third of the steam is produced in the economizer, and the remainder in the
external boiler.
OPERATIONAL AM) PERFORMANCE CHARACTERISTICS
Technical feasibility of the MS-FBC has been successfully demonstrated in a bench-scale 6-inch-
diaoeter x 20-foot-high combustion unit at Battelle. Crushed coal of less than 1/4-lnch size and
limestone of less than 100-mlcron size were used. The combustor temperature was maintained at 1600 to
1750 F under atmospheric pressure.
227
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HUE CAS TO
KONOMIZH
SHAM
WAHI
DISTRIBUTOR PLAIE
AIR
30TOซ
RET/SEC
Figure 4. Btttelle Multisolid Fhritizad Bซd
External Boiler
228
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Three types of high-sulfur coal coal fron Ohio, from Illinois, and from Pennsylvania -- were
tested. The process operated similarly with all. The limestones evaluated to date were obtained from
Flqua, Ohio, and Grove, Virginia.
Specullte, a brand of hematite (iron ore}, was used as the high-specific-gravity material in the
dense fluidized bed. Speculite is relatively inexpensive. Is chemically stable in this application,
and demonstrates a low attrition rate.
Fine Speculite, 1icestone, and a rounded sand vere tried as the entrained fluidized-bed mate-
rial. Sand is currently used because it is low in cost and attrition rate, and shows satisfactory
or superior performance in other respects.
As described earlier. In the MS-FBC with either integral boiler or external boiler, the
entrained fluidized-bed material serves as a heat carrier that absorbs combustion heat in the dense
fluldlzed-bed region (the burner) and transfers the heat to heat-exchange surfaces located outside
the dense bed region. In the feasibility assessment of the XS-FBC, the 6-lnch bench-scale unit was
run to simulate operations of the process with an integral boiler. Demonstrated characteristics .ire
discussed below. A comparison of operating characteristics of conventional FBC and MS-FBC is shoun
in Table 1. A dlagrao of the 6-inch unit Is given in Figure 5.
Superficial Air Velocities
The MS-FBC has been operated at superficial air velocities up to 35 ft/sec. This is between
3 and 4 times the velocities used in conventional FBC boilers (8 to 10 ft/sec). There have been no
indications that 35 ft/sec is an upper licit for the process. It should be mentioned that the system
throughput is directly proportional to the air velocity; thus, increased velocity Implies a more
compact boiler design. The MS-FBC boiler would have a cor.bust ion rate of 200 or more Ib of coal
per hr per ft? of bed area.
Sulfur Sorbent Requirement
EPA sulfur oxide enlsslon standards (1.2 Ib S02 per nil lion Btu) have been met consistently
with a limestone requirement equivalent to l.S to 2.2 while burning 4 percent sulfur coal. Figure 6
shows the effect of limestone size on sulfur capture. The effect of bed temperature is shown in
Figures 7 and 8.
It should be emphasized that MS-FBC can eiploy almost any :ypc of limestone or dolomite regard-
less nf its physical characteristics (pore size, attrition resistance, etc.) because the process uses
finely ground sorbents. In conventional FBC, the particle size of sorbent is generally an order of
magnitude larger than that for MS-FBC; therefore, suitable selection of a sorbent for the conventional
process can become restrictive.
Combustion Efficiency
Up to 96 percent coal combustion efficiency has been obtained in the 6-in. experimental
HS-FBC unit.
Erosion/Corrosion of Steat. Tubes
Erosion/corrosion of vertical steaa tubes has been ceasured in a SO-hr test run. In the en-
trained fluidized bed, metal wastage rates of vertical specimens was 0.0054 mil/hr for carbon steel
(A109) at 485 F, 0.00018 mll/hr for 9 percent chromium steel (P9) at 525 F. and 0.00019 mil/hr for
Type 347 stainless steel at 570 F. These cetal wastage rates do not appear to be excessive. The
maximum allowable rate is about 0.0015 mil/hr.
In the dense fluldlzed-bed. erosion/corrosion rates were approximately an order of magnitude
higher. However, heat-exchanger tubes are not subjected to the dense bed in the KS-FBC process.
Part Load Operation
The fluidizatlon characteristics of the dense fluidized bed of high-specific-gravity material
are such that fluidlzation can be maintained over a wide range of superficial air velocities and,
hence, combustion rates. Because the dense-bed temperature is controlled by the reclrculatlon of
entrained bed material rather than by Icsersed heat-exchanger tubes, the range of loads is not
restricted by the decreasing bed temperature, as is the case in conventional FBC.
229
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Table I. Comparison of Operating Characteristics of Conventional FBC and KS-FBC
FBC
MS-FBC
Superficial Gas Velocity
Densu Bed Material
Type
Typical Size
Particle Density
Entrained Bed Material
Type
Typical Size
Particle Density
Sulfur Sorbent
Type
Typical Size
Method of Heat Recovery
Steam Tubes In Dense Bed
Steam Tubes in Free Boar-1
Steam Tubes in Entrained ^ed
Method of Controlling Dense Bed
Temperature
12 ft/sec, max
Limestone or
Dolomite
8 x 30 mesh
2.6 g/cc
Not used
Limestone or
Dolomite
8 x 30
Yes
Yes
By Immersed
nteam tubes
30-40 ft/sec
Iron ore, etc.
6 x 12 mesh
5.2 g/cc
Sand
20-100 oesh
2.6 g/cc
Limestone or
Dolomite
Minus 325 mesh
Ho
Yes
By adjusting
recirculation
rate of en-
trained bed
solid
230
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To 4th
cyclone
ond stock
Deflector plote
Flue gos
sompling
port
Combustor
l ond
limestone
Distributor
(dote
t= A,r
Figure 5. Schematic of 6-lnch-Diar
-------
100
90
80
c 7ฐ
QJ
U
I 60
ง
'ฃ 50
o>
30
20
10
.Federal S02 Emission
Standard
-325 mesh
-200 mesh
Coal: Illinois No. 6
Limestone: Piqua Limestone
O.5 1.0 1.5 2.O
Ca/S Molar Ratio
Figure 6. Effect of Limestone on Sulfur Retention
2.5
232
-------
o>
IOO
90
80
70
tt>
Q. 60
o
' 50
I 4ฐ
30
20
10
O
1550-1700 F 1850 F
1760 F '800 F
Federal
Emission
Standard <<
Coal: Illinois No. 6
Limestone: Pi qua Limestone _
(-325 mesh)
Data Taken From Run 1013
1234
Ca/S Molar Ratio
Figure 7. Effect of Combustion Temperature on Sulfur Retention
233
-------
o
I4
h.
o
o
o>
E
o>
fa
O)
o
o>
Coal: Illinois No. 6
Limestone: Piqua Limestone
(-325 mesh)
I40O I5OO I60O 1700 1300 1900
Combustor Bed Temperature, F
Figure 8. Effect o i Combustion Temperatw* on
Limestone Requirement
2000
234
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The 6-in. experimental unit has been operated at loads ranging from the nominal 100 percent
(corresponding to a superficial velocity of 32 ft/sec) to about 20 percent, l-o difficulty .as en-
countered at any of the loads.
Shut-Down and Standby
The absence of heat-exchanger tubes in the dense fluidized bed of MS-FEC permits the process to
be shut down rapidly by simply stopping the flow of air, coal, and limestone. Ko deliberate cooling
of the bed material is required. Further, the bed material can be retained at high temperatures for
considerable periojs, permitting rapid restarts following an ext< iced shut-down.
IMPLICATIONS AND ADVANTAGES OF THE PROCESS
The MS-FBC process has many implications and advantages. These are presented below as cate-
gorized in four areas of interest:
1. Environmental
Effective SOj emission control, even when more stringent emission standards
are imposed
Low KOX emissions, due to low c-abustion temperature
Small amount of dry-solid wastes to be disposed of, due to low sulfur-sorbent
requl i i-'Cfcnt
2. Operational
Excellent turn-down capability
Fast load-following
Immediate restart and shut-down
Adaptability to a variety of fuels
Ability -.o use any type of limestone sorbent
3. Engineering/Developmental
I ewer ,cale-up problecs, so faster rate of development to a coccerciol
scale plant is likely
Natural circulation of water, because heat-exchanger tubes are vertical
Applicability to retrofitting existing boilers
Greatly reduced or virtually eliainated corrosion/erosion of steam tubes,
because steam is generated in a low-velocity 1luidized-bed external boiler
It, Economic
Low capital cost, because the boiler (a) is compact with a high throughput.
(b) can be shop fabricated and rail shipped even in the case of talrly
high capacity boilers, and (c) requires a smaller Investment in solid-waste
handling and disposal facilities or activities
Low operating cost, due to low capital cost, low requirement for sulfur
sorbent, and small amount of solid waste.
Large-capacity packaged boilers (of the order of 150,000 Ib of steam per hr) are expected to be
feasible for the H',-rEC process. These will provide reduced boiler cost, increased reliability of
construction, and enhanced ease of installation.
TIMETABLE OF CURRENT DOE MS-FBC PROJECT
The Kultlsolid Fluidized-Bed Combustion process was developed under the Battelle Energy Program,
which invested some $600,000 In the development over a three-year period. The feasibility of this
concept has been successfully demonstrated in a 6-in.-diameter (0.22-ft2 bed area) coal-co-bustIon
unit. The L.S. DUE contract with Battslie calls for a two-phasr scaling-up of this process over
six years (Including .'hree years of ope)ating the demonstration plant). The Sub-Scale Experimental
Unit System (SSELS), vl.iro represents a 10-fold scaleup of the 6-in. bench-scale unit, is now in
235
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operation. This pilot-scale unit Is designed to produce about 6,000 Ib of stean per hr from 400 Ib of
coal per hr. The coehustor/boller Is sketched In Figure 9. The full-scale deaonstratios plant, which
will be built adjacent to BattelIt's present steam plant, will represent a further scale-up of about
t> ti-es; it will produce 25.000 Ib o! steao per hr irom 2,500 Ib of coal per hr. Data obtained Iron
this deoonstration unit will be used to design and build coccercial boilers.
STEAM
4000LB/HR
STEAM DRUM
IS STEAM
TUBES, r IMA.
HUE CAS
TO CLEAN UT
CYCLONES AND
STACK
SAND RECYCLE
SOLIDS
DISENCAGEMlNf
ZONE
COAl AMD
UMESIONE fill)
RECYCLED ENTRAINED
MATERIAL
RECYCLED ASM.
SORCENT.ASAND
MAIN
UUIOIZING AIR
DENSE BED ZONE
Figure 9. Battelto Sub-Scale Experimental Unit Fhjidued-Bed
Stsam GMM/ator
236
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QUESTIONS/RESPONSES/COMMENTS
MR. N'ACK: Fron J. S. Cordon, TRH Energy Systems: "Some infor-
ration on pressure drops in your system?" Typically in the 6-inch
unit, we neasure a pressure drop of about 35 inches of water across
the dense bed. He note very little pressure drop, possibly as nuch
as 2 inches of water, due to drop in the entrained zone.
Fugene 5-nyk, Arqonne National Laboratory, asks three questions:
1. "Pcesn't your use of the entrained bed as the heat transfer
neditm result in ni/ch lower heat transfer coefficients, thereby
partially nullifying one of the biggest advantages of FBC?" The
transfer coefficients in our system in the entrained zone have been
measured at about in to 20 Rtu per hour per square foot per degree
Fahrenheit, against about 40 to 50 in conventional systems, and so
it is lower. However, in the integral boiler node, we are using the
freeboard space, essentially, and the spacing of the tubes is not
critical; that is, they can he relatively close together, so that
on a volume basis, we feel that the heat transfer space required is
aboi:t the same.
2. "Poes the limestone recycle7 'f not, how do you keep it in
the bed lorn; enouoh for nood sulfur retention?" Since K^ use a very
finely divided limestone, tiv mechanism of reaction is diffusion
limited, rather than reacti -limited; so reaction time has very
little effect.
3. "Poos unburned carbon recycle? If not, how do you keep the
coal in the bed long enough for good combustion efficiency?" Yes;
we do collect and reinject carbon in t-.oth the 6-inch unit and our
pilot plant unit, and we feel that this increases combustion effi-
ciency by about 6 percent.
A question asked by farl Oliver of SRI. "Have you measured
attrition of the iron ore? How fast is it?" The attrition rate of
the iron ore is about .on? pounds of ore per pound of coal. As I
mentioned, this corresponds to probably less than 2 cents per* million
Ptu's.
Pick Heldhart, University of Bradford, U.K. 1. "How much
attrition of the sand occurs, and does the attrition occur princi-
pally in the cyclone or in the bed?" Sand attrition rate is about
.02 pounds per pound of coal. Vie don't know where it occurs. We
think it occurs in the dense bed, primarily, although certainly
there is some attrition in the cyclone; 2. "Hhat are the bed den-
sities in the iron ore section and in the dilute phase?" The bed
density in the iron ore section is about 50 pounds per cubic foot.
We do not know what the bed density in the dilute phase is. We have
no good way of measuring that at this time.
237
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A. l/omser, Homser Engineering: "Hhat combustion efficiencies
and NOX enissions have you observed?" Our conhustion efficiencies
in the fi-inch unit have been in the range of 87 to 9fi percent. Using
Illinois No. 6 coal, i/e have observed f!0x emission concentrations of
300 to 400. ppn.
Fron Henry Kwon, Porr-Oliver: "Fron the process point of view,
high gas velocity seens to offer a possible alternative, but what
about mechanical aspects? Pue to high gas velocity, the unit nay be
relatively very tall, which would cause a lot of mechanical problems
in construction of a commercial size."
As I mentioned, the tubes can be closely spaced, so that the
height of the unit will not be considered to he excessive. Our pilot
plant unit combustor is 21 foet high, which is not unusual for a
large comhustor. 'cale-up would be primarily by length and width,
rather than height, and so consequently, we don't see this as becom-
ing the najor problem.
From Ken Chipley, Union Carbide-Oak Ridge National Laboratory:
"Ho you have any idea as to the heat transfer coefficient on the
vertical tubes?" I mentioned that it is about 10 to 20 Rtu/hr per
square foot per degree Fahrenheit. It is a function of the recycle
rate, and varies to some extent with that rate, going up as the
recycle rate goes up.
Another question from Ken Chipley: "What is the wall construc-
tion of the pilot plant unit?" The pilot plant unit is constructed
of stainless steel, hot wall construction. It was designed so that
it could be rapidly started up and operated easily and changes made
easily.
From Masayuki Horio, fJagoya University: "Hhat do you expect or
have you experienced with deposit of fine limestone sticking on the
wall duct, which was seen in the dry limestone stack gas desulfuri-
zation?" My answer is, as 1 mentioned, to date we have had over 300
hours of operation on the pilot plant unit, and we don't see any
problem in terms of the ducts clogging up with dry limestone. We
don't anticipate any in view of the high velocities and the presence
of other materials in the flue gases that will scour the limestone.
From A. Saha, Burns and Roe Industrial Services: "!t seems
that you need a high calcium-sulfur mole ratio for meeting EPA
limits, as in simole fluidized bed combustion systems. So, are
you substituting a smaller but more complex system for a bigger and
simpler fluidized bed combustion system?" I think we have demon-
strated, for the most part, that our calcium-sulfur ratios are con-
siderably lower than in the conventional system; aliout 2 to 2.5
238
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seens like a fairly good number that we have routinely experienced.
In addition, we can use almost any type of limestone. The fact that
it night attrite is a benefit in our system so that it gives us a
lot nore flexibility in terms of the source of limestone.
Secondly, Mr. Saha asks: "Hhat percent of carbon loss are you
having?" I mentioned that we have experienced combustion efficien-
cies between 87 and ifi percent. These are not optimized values, by
any neans, as they were obtained in a small unit. He hope that
possibly we can exceed that as we continue in our test program on
our pilot plant.
CHAIRMAN: Thank you very much, Mr. Mack.
239
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INTRODUCTION
ROBERT BROOKS, CHAIRMAN: Our r.ext paper is a report of work at
Battelle on a pneumatic solids injector and startup burner for their
muHisolid fluidized bed combustion process. The paper, by G. W.
Felton and R. D. Giamma., will be presented in parts by both G.W.
Felton and R. D. Giammar. Mr. Giammar received his MS degree in
Mechanical Engineering and has been a research scientist at Battelle
since 1967. His fields of interest include heat transfer, fluid flow,
and thermodynamics, and he has worked on the firing of pulverized and
stoker coal as well as conventional fuels.
240
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Pneumatic Solids Injector and Start-Up
Burner for Battelle's Multisolid
Fluidized-Bed Combustion (MS-FBC) Process
G. W. Felton. R. O. Giammar, H. R. Hazard.
D. R. Taylor
Battelle Columbus Laboratories
Columbus, Ohio
ABSTRACT
Two important and innovative components have contributed to the successful
operation of Battelle's Multisolid Fluidized-Bed Combustion Process: a pneumatic
solids injector ar I a start-up burner. This paper, appropriately in two parts,
describes each equipment's development, dcsiqn, and test.
In Part I, we describe how Battello has developed and successfully combustor-
tested a novel pneumatic injector for feeding solids ir.to the bottom of a fluidized-
bed against back pressures of up to four PSI. The injector was sized to feed 600
pounds por hour of 1/8 x 0 material using 1/2 oound ot air per pound of solids fed.
In the process downstream air from the iniector is bypassed back throuqh a control
valve into the injector solids entrance stream to reduce the solids inlet vacuum to
any desired low level.
Part II describes how Battello selected the qas turbine can-type burner because
it operates stably over a wide range of firing rates (50 to 1). air/fuel ratios (100
to 1) and pressures (up to 15 atm). Due to its high heat-release rate, (8.5 x 106
Btu/hr-ft^ atm), it is extremely compact and provides uniform nas-outlot temperatures.
PART I. A PNEUMATIC SOLIDS INJECTOR FOR BATTELLF'S MULTISOLID FLUIDIZF.D-BF.D
COMBUSTION PROCF.SS
Introduction
The injection of solids into rrorc-traditional fluidizcd beds has already boon
successfully accomplished by a variety of techniques, but the problem of solids injec-
tion becomes more difficult for flattelle's Multisolid Flujdizcd-Bed Combustor (MS-FBC)
when crushed coal d/8" x 0) is fed along with -325 mesh limestone. The necessity to
keep dry the powdered limestone is evident. However, coal with a high surface moisture
will cloq feed mechanisms and injectors for cbvious reasons.
A quick review of somo of the technioues used to feed solids into a fluidizcd
bed include:
Angled needle tubes (used at Rivcsville)
Coal slirjger feeder used for stoker boilers being evaluated for top
feeding of laroe coal fractions (used by Foster-Wheeler)
Pneumatic injection systems of various types.
Potential problem areas for coal-feed mechanisms include:
Coal is allowed to heat up to the plastic state where agglomeration will
plug tubes if coal is not sufficiently cooled
Excessive fluidized-bed velocity that will entrain fine coal if it is top
fed
Mechanical feeders that have a lower reliability factor where more moving
parts and several coal streims exist, especially when feeding coal with
greater than 31 moisture
The type of feed system can be solids size dependent. For example, some
injectors may require finer crushed coal than a slinaer type of system
which will not work well for fine coal
Mechanical solids feed system reliability weighed against initial and
operating costs for a pneumatic injector system
Air requirements for injector systems evaluated with attention to potential
scale-up problems including amount of air used for injection of solids.
241
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A ypical coal injection system will use 1/2 pound of air per r^ound of solids
fed. Horizontal velocities roust be hich enouqh tc prevent saltinq out of solids. To
accomplish this, a velocity of 80 feet per second is generally in a safe region to
prevent solids dropout.
Battclle's MS-FBC requires greater than 600 pounds per hour of solids per
injector. Coal is dryed to 9'' moisture crushed and screened to 1/8" x 0 at Battelle's
West Joffcrson coal preparation plant for use with iniectors. We were aware that nany
mechanical problems were encountered in soli-1-feed systems with rotary valves and
conveyors. To us, a more simplified pneumatic solid-feed system seemed desirable for
tho MS-FEC, especially with the use of -325 mesh limestone along with 1/8' x 0 coal.
We believed that properly designed injectors prevented coal frcr. becoming heated to
the plastic region due to a rapid discharge into the combustor.
To begin with, Battelle wanted a solids injector that met the following
specifications:
Work against a four pound per squara inch back pressure of the fluidized bcrf
Introduce no additional air that was not measured along with solids being
fed
Life time of one year between inspections
Solids throughput of 1/2 pound of air per pound of solids with outlet
velocity of 80 foot per second
Feed 1/8" x 0 coal and -325 mesh limestone.
To meet these specifications, we tested several commercially-available
injectors at Battclle's West Jefferson site, using feeders for a conventional 24-inch
diameter FBC and feeding 260 pounds per hour of crus/-fx} coal with 70 pounds per hour
of limestone (;-cc Figure 1). This testing resulted in nozzle failures after approxi-
mately SO hours of operating time because of erosion of the inside of the nozzle where
feed jir exited and because of erosion of the nozzle bore by discharging solids (see
Figure 2) .
Since most nozzles are designed to evacua;-o a vessel with steam or air flow
through the nozzle, we know very little about material flow rotes versus particle size.
We tested several Penberthy nozzles (designed for steasi operation) for solid flow
rates but were not able to meet Battelle's specifications. We modified these nozzles
by enlarging the body bore and trimming .060 off the nozzle inlet, and this proved to
be a very satisfactory solution as tested in a bench-scale test rig.
Testing
We used a small test set-up to investigate soli;s flow through nozzles into a
tank against varying back pressures. (See Figures 3 and 4). This apparatus consists
of a rotometcr to measure external bleed air to the iniectors. This design proved
unsatisfactory because so much air was required {over 600 SCFH) to control the inlet
vacuum to a desired level. We then developed another arrangement where downstream air
from the nozzle was bled back through a control valve to reduce the vacuum at the
solids inlet to any desired low value. In this process some solids are recycled back
through the bleed valve used to reduce suction pressure, but this did not have any
detrimental effect on the operation of the system and did not require any air that had
not been previously measured. A very low vacuum is maintained easily by controlling
the amount of air recycled.
To increase the operating life tine of the injector nozzle, the body was
fabricated from 4340 steel hardened to Rockwell C 50. This prevented excessive wear
such as was noted on the inside of the test nozzle where exit air had eroded the brass.
RESULTS
Two finished nozzles of the modified Penberthy design, manufactured from 4340
hardened steel, were tested in Battelle's MS-FBC. With over 250 hours of operating
time, they-have performed satisfactcrily. One nozzle injects 1/8" x 0 coal with -325
mesh limestone and the other rcinjects ซ-he flyash to inprove combustion efficiency.
These nozzles should have near 50 pounds per square inch operating pressure and
require an auxiliary air compressor in addition to the main fluidizing air blower.
242
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Figure 1. Coal Injector Nozzte Tot Site
243
-------
Figure 2. Commercial Noitle Failure at 50 Hours Run Tirm at 260 Pounds/Hour Solids Feed
244
-------
I
I
*
3
Figure 3. Injector Test Apparatus
245
-------
~ - ** J-. ~ < - I _;
*<- 2?- 55^
Z i
Figure 4. Injector Test Apparatus
246
-------
The advantages arc obvious: they arc sinple to construct, they work well and they are
reliable. Larger nozzles could be fabricated to feed 1/4" x 0 coal in larger systems
if desired.
A picture of the actual injector nozzle used for both coal/limestone and ash
injection into Battelle's MS-FBC is shown in Figure 5.
This nozzle will feed 600 pounds per hour of solids with 11/16 inch bore at 60
pounds per square inch pressure and 561 pounds per hour at 50 pounds per square inch
with two-inch Hg suction or less at the solids feed point.
PART II. A START-UP BURNER FOR BATTELLE'S MULTISOLID FLUIDIZED-BED COMBUSTION PROCESS
Introduction
An auxiliary start-up burner is used initially to heat the bed to some temper-
ature (generally above 1000 F) so that when coal is injected into the bed, ignition
occurs and combustion is maintained. This start-up burner also generates a stable
flame and a uniform outlet gas temperature over a wide range of air/fuel ratios and
firing rates. Most importantly, it operates against the pressure of the fluidized bed
system. Most "off-the-shelf" commercial oil/gas burners meet these requirements,
except for the last the capabilit/ of firing at elevated pressures. The "off-the-
shelf" burner is generally designed to operate noar atmospheric pressures and specially
designed housings are required for pressures much above 2 psig. The standard burner
accessories such as pilots, controls, and regulators, usually are not designed for
elevated pressures. These burners, with the possible exception of the high excess
burners, also cannot provide a uniform outlet temperature of 1000 F over a wide turn-
down range (although dilution air can be mixed into the gas stream downstream of the
burner).
Accordingly, the start-up burner for Battelle's V.ultisolid Fluidizod-Bed System
requires a specially designed burner or a modification of a commercial burner. Because
of the constraint of a relatively short delivery tine, we, at Battelle's Columbus
Laboratories, decided to design and build the start-up bjrner. We selected a burner
design similar to that of a gas turbine can-type ooaibustor.
This paper provides the background to the desian of this start-up burner.
(For a more detailed discussion, the reader is referred to the Gas Turbine Engineering
Handbook.!)
Start-up Burner Design
The burner design, based on the gas turbine combustor, was selected because it
met all the specifications of stable flame and uniform qas temperatures, over a range
or air/fuel ratios (up to 100/1), turndown ratio (up to 40:1), and pressures (up to
15 atmospheres). In addition, because of its high heat release rates (up to 8,500,000
Btu/hr-ft3 atm), the gas turbine combustor is relatively compact in comparison to an
industrial burner.
We knew that the design of qas turbine ror.bustors had been advanced through
experimental development with qualitative assistance fron theory. We also know that
in any practical combustion system the large number of variables and their complex
interrelations make designs from fundamental principles difficult. Although gas
turbine combustor design and development is still an empirical air, we were prepared
to meet the difficulties.
Gas Turbine Start-up Burner Design
Although gas-turbine start-up burners vary widely in design, all similarly
perform the basic functions. Fuel, either gas or distillate oil, is injected into a
region whore ignition and flame stabilization are provided by a recirculating flow
pattern within the burning flame. Combustion continues in a region in which near
stoichiometric air/fuel ratios are maintained. The hot combustion products are then
mixed with the remaining air in a dilution region to provide a uniform and suitable
temperature profile at the outlet of the combustor.
247
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Figure 5. Modified Penberthy Tyv Injector Nozzle Used to Inject Solids Into BatttNe't MS-FBC
248
-------
Figure 6 shovs the tubular -rorJoustor that Eattello -ised for -ho start--.p
burner for its Mu!tisoliO F'luidized-Eod Systeri. This start-up burner con.bines i..->sr of
the oenc-ral features in coir.bustor de-si in Th.-;- corbustion air passes throu.;h a perfor-
ated plat' to improve flow distribution and is distributed around tho co-irjstor lir.c-r.
The air then enters the annular space b"tween the- liner and casin.i (6-inch pir,e) .3--;
is admitted ar3dually into the space within the liner through nunurous holes :::,2 s! ts.
The cicsiun of these holes and slots provides for distinct rones for Mane stabiliza-
tion, co-bust ion, and dilution, and also provides film coolino of tho liner. i'-;--l :s
admit tc-d through a nozzle that distributes wit.hin an included amile of 60 to 90 decrees
A spar'i plug is used for linht-off only.
In Fiqure 6, the combustion zone is surrounded by .; cyl in-iric-.il shro'jd that
provides a good ci rcun-.f crential flow distribution. This shroud !ui = a soco.idjry eff'--rt
of roc:::cinc: radiation to the outer cj~i:n:, thus ziini^izinq casino ten:.c-raturcs. In
the dilution zone, rows of holes arc arranqea for thorough ict mixi.no and on tra inner.*
to provide nearly uniform outlot tcnperatures.
The coiT.bustor linor is const rur:tc :'. of air-roolod s^r-f-r. ri.nal. The outsid'-
surface is cooled by flow of air in the annular r-r,;ic" around i , and tr.c- insi-io
surface is cooled by a filn of air odnitt": throuuh slurs .-.rr.-inj..-' to blanket the
surface. This filn is trie feature t h^t rak.es 1 iซhtw-i>:h^ netal c-^nst ruct ion :-ossi;jle
as all the heat received by the linor is only that radiated from the :'ia~<.-.
Conbustion air is admitted through tn'- rjius o: r.'idial iets. These iets are
deni'jnc-u so their anole one r-enetr.jt ior: t.ruvid'-s an ir-T.-in-i'-r.-'nt at the cerbustor axis.
This results in an uปst rcuri flow :hicr. ni abi 1 i z<:-r, h" flanr-. i.ach lot entrains air
and b.irnir.j fuel. A reci rculat;ion pattern it. formed th-'it provit.'es intensive turbu-
lence- and nixinn throtrihoct tho conb-js'or. rnrierJ to Jj.iv> ,1 cold flew i rossuro <:r-p of
9" I!2O et desi'in condition of one nilli^n litu/nr j:j['Jt with -in outlet t'>r.pซ.-r.if-.r-.- of
1200!-'. This burr.er , h-ns the following ad.'anta<:fs ปhat iran not b..' n.itrhod by any o'.h- r
conb'jst ion O'tuipnc-nt:
It will opi-rate equally well at any air prt-bsure within th<- rjr.ซ.iu' of
pressures available, f/rr: 1 r> to 100 f.si-i.
It can be operated with .-sr.y desired Jas cut !c-'. t-:::. -raturo. frcr. a!:o:;t
200F ^o about 1ROOP. Gnrr- the air flow is vstabl ishc-d th-.' st3rt--.::J b'jrr...-r
outlet tonpet aturc ran 'if changed within this ran'-i" by .-Jijn.Jir.w fuo! flew.
It will lit:ht off innocli-itely -ind t,or. it ively with .jn .-K-i-trio F-..-ar'-: at any
fuel rate above the niriir.-irti required for ior.ition. which rorros: cr.J^ to a
rather low outlet tr-npcrature.
With ?hซ' burner operatino. it ran he oxp-osc-J to lar*:ซ-- i:xctirsปor.s i:i ->ir
flow, in pressure, and in fuel rate with no danrror of b'owo-.it. Thv cnly
way of blowinc; out the fla.-nc is to redur-e : h'- fuel rlow rate b.-!"w ti:'-
nininun. which corresponds to a burn'--r-fvn lot t c."r^p--r.it in e of -00 to -5CGF.
Tho cas turbine burner is uniu-j^ in that ovr r.ost of tho operating rin.---, th"
overall fuel/air ratio is too loan to i-inite. explode, or b'..rn. This is nado possible
by carrying out cor.bustion with a stoichio.-ietrir nixtures in thi- buincr r-nii of th--
combustcr. followed by suo-ossive dilution by air iets thai- thoroughly nix the com-
bustion pro.-iuct s to a lean. lcw-tcor,erature Dixt-jre. In case of an jntorrur.t icri of
fuel rlow that would put the flame out. or if there w--ro a failure to inniti-, th<-
mixture at the burner outlet, and throughout tho qasifior vessel or rorrbustor, would
be too lean to iqnite, burn, or explode. Thus, the burn'-r is bJ si ally r.-.ich safer
than conventional burners. As a r-onscq-jcnce, only a thermocouple located in exhaust
249
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CO
a
c
o
39
Z
I.I
O >
5
Figun&
250
-------
stream was used to cut off the fuel supply in the event that the aas temperature
deviated above or below the desired set point.
RESULTS
The start-up burner was checked out prior to installation in Battelle's
Multisolid Fluidized-fled Systen. At the desian firing rate, ionition occurred at air/
fuel ratios of about 50 to 1. Once the flaire was established this burner was operated
at air/fuel ratios of over 500 to 1.
After check out. we installed burner in our fluid-bed system. Ignition occurs
readily. Upon establishing the desired air and fuel rates, the burner maintained a
gas outlet temperature to within a degree.
REFERENCES
1. Hazard, Herbert R., 'Combustors'. Chapter 5 of the Gas Turbine Engineering
Handbook, First Edition. Editor, John W. Sawyer. Gas Turbine Publications, Inc.,
Stamford, Connecticut, 1966.
251
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QUESTIONS/RESPONSES/COMMENTS
P.DRFPT RPnnr.S, CHAIRMAN': Are there questions? Could you
identify yourself, please?
MR. W1RKSFR: Al Womser of Womser Engineering. What oil
nozzle did you use, and was that running on oil or gas, then?
MP. filAHMAR: For oil , and we have used a standard air-atonized
nozzle that you would use in a conriercial burner. We have also fired
the burner with a sonic-core nozzle with very fine atorvization, which
perfomed equally as well as a Delavan Swell-Air Nozzles operating at
An pounds pressure with air atonizationnothing really specific or
special with this nozzle.
FIR. FFLTflfl: That rccirculating air stream just serves to kill
the vacuun that you would nomally yet with an injector, an eductor.
And all we are doinq there is recirculatinq air hack to eliminate
that vacuun, because we don't want a larye volume of unncasured air
coning hack in with the solid naterial. This enables us to get a
fjood neasure of conhustion air in our unit.
PORFRT RP.nOKS, CHAIRMAN: Are there other questions? Are there
any questions on any of the four papers we have heard this afternoon?
If not, I want to thank the authors for bringing to us information on
this very important subject this afternoon.
252
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INTRODUCTION
ROBERT CHRUNOWSKI, CHAIRMAN: The first paper on the agenda
will be "Fluidized Combustion of Beds of Large, Dense Particles in
Reprocessing HTGP Fuel" by Derrell Young of the General Atomic
Company. Darrell graduated fron Lamarr University in Beaumont,
Texas, in 1970 with an MS in chemical engineering. He's worked for
a couple of years with Texaco. Since 1973, he's been a process
development engineer for General Atomic and is now a senior engineer.
I'd like to get this started right away, and get right into the
session. Oarrell?
MR. YOUNG: First of all, ny thanks to the Conference coordinators
for inviting our paper. It is on cjraphite burning. We'll depart a
little bit from coal burning and ! d also like to thank you for
your attention at the outset, especially those who might have heard
this paper at the AICHE meeting just last month. I'll try to throw
in some different stuff especially for you, and hopefully make
it interesting for everyone. May I have the slides on please?
253
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Fluidized Combustion of Beds of
Large, Dense Particles in
Reprocessing HTGR Fuel
DerrellT. Young
General Atomic Company
ABSTRACT
Fijldized bed combustion of graphite fuel elements and carbon external to fuel particles Is
required in reprocessing high-temperature gas-cooled reactor (HTGR) cores for recovery of uranium.
This burning process requires combustion of beds containing both large particles and vjry dense car-
tides as well as combustion of fine graphite particles which elutriate from the bed. Equipment must
be designed for optimum simplicity and reliability as ultimate operation will occur in a limited
access 'hot cell1 environment.
Results reported in this paper indicate that successful long-term operation of fuel element
burning with complete combustion of all graphite fines leading to a fuel particle product containing
< 1: external carbon can be performed on equipment developed in this program.
INTRODUCTION
Objective
The High-Temperature Gas-Cooled Reactor (HTGR) Is a thermal reactor concept which uses helium
as the coolant and graphite as the moderator and core structural material. The HTGR fuel cycle employs
an additional energy source by the use of thorium as the fertile material. The U-233 bred material is
recovered for recycle to the reactor with U-235 as the makeup fissile material.
The fuel In an HTGR consists of fissile mlcrosphere particles containing e1.her U-235 for fresh
fuel or U-233 for recycle fuel, and thorium fertile particles. The fissile partu ie? are coated with
a multi-layer pyrolytic carbon-silicon carbide ('TRISO') coating and the fertile particles are coated
with a two-layer pyrolytic carbon ('BISO') coating. These particles are bonded into fuel rods and
are contained in a hexagonal graphite fuel element. 'The fuel element and the fuel particles are
shown in Figs. 1 and 2.
196 18 89 KG GRAPMITI
7ปI8 116 KG IOIปl
711 IB 96KG I01AI CปOBO%
CROiSSiClON
Figure 1. HTGR Fuel Element
254
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fyHOLปIlC CARBON
TRISOHSSILl
Figure 2. HTGR Coated Fuel Particles
The recovery of bred U-233 'in HTGRs for recycle requires reprocessing of the spent graphite-
base fuel. This reprocessing involves removal of the graphite, separation of fissile and fertile
particles, dissolution and solvent extraction to separate uranium, thorium and fission products. The
reference method of removing the graphite is by combustion, which is most readily achieved in flui-
dized-bed burners.
Two types of fluidized-bed burners are bein-) used: the crushed fuel element, or 'primary
burner,1 and the crushed particle, or 'secondary burner.' The secondary burner i. discusbฐd briefly
by Young (l) and in rare detail by Rlctaan (;). This paper presents details of the primary burner
development as its relatively large and widely rancing size and density distribution bed material
presents interesting fluidization and fines recycle problems which are unique in the varied jses of
large particle fluidized-bed combustion.
The Problem
The primary crushed fuel element burner receives a feed stream of intact fuel particles and
crushed graphite produced by a three-stage crushing system. Average properties of the burner feed,
nominally minus 5000 -_m, are shown in Table I. The prinary burner must ignite and burn the fuel
element graphite, the fuel rod pitch, and the fuel particle pyrolytic carbon at 900ฐC producing a
product of discretely separable silicon carbide coated fissile oarticles and fertile fuel kernel
particles containing less than 1 wt. * exposed burnable carbon (Table II). Adequate fluidization of
the largest burning graphite and the large, dense fuel particles requires superficial velocities of
1.01 to 1.37 m/sec and a carefully designed gas distribution device. Graphite fines (-425 ..m)
generated both in the crushing process and in the fluidized-bed burning, elutriate before they are
burned. This carryover of fines is significant, amounting to over 301 of the carbon fed. This large
mass of fines presents an expensive nuclear waste disposal problem: hence, combustion of the fines
within the burner system is necessary.
Table I. Average Properties of Primary Burner Feed & Startup Bed
Tap Density:
Bulk Density:
Angle of Repose:
Avg. Burnable
Avq. Particle
1.25
1.08
35*
g/cm3
g/cn3
Carbon: 817
Size:
1378
urn
1 Burnable
Ut 5
Size Total
-6350
-4760
.
.
-
.
"
869
550
420
375
250
m
m
m
Fl
m
n
m
* 4760 m
* 869 m
+ 550 m
+ 420 m
+ 375 m
+ 250 n
5
60
20
5
4
3
3
of
Sample
Carbon
in
Fraction
100
100
21
78
97
98
98
255
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Table II. Average Properties of Prirary Burner Fuel Particle Product
Bulk Density: 2-6 gr/cc
Size: 200 - 870 -_
Exposed Burnable Garten: < 17
Tie primary burner operates in semi-continuous batch cycles. Operating guidelines ere sumnar-
i;ed below in Table III. This operating mode processes an incoming lot of fu-.-l elements in a series
of burner runs which are separated by brief fresh feed interruptions for unl-.ddii.g a partial batch of
the fuel particle product. The burner temperature is raintained by t>e theial inertia of the par-
tial bed retained and by external induction heating. The burner is completely emptied only after the
specified lot of fuel elements is processed. The burner is then restartec by using the induction
system to heat the initial crushed fuel element loading from the next customer lot to ignition tem-
perature. Thus, the bed composition ranges from one ric'i in burnable ca-bon to one rich in burned-
back fuel particles. Bed bulk density and. hence, bed Height, varies by a factor > 5.
Table III. Operating Guides - Prirary Fluidize-l-Sf-d Baner
Fuel
p.urner
a. operating mode semi-continuous batch
b. startup induction heater
c. gas distributor perforated cone
d. feed system gravity/rotaซ-y valire
e. product removal gravity/pne'jnatic
f. feed/product control .... s*tered feed/batched product
g. heat removal air cooling *ia awtular shroud
h. fines recycle gravity external recycle to
burner via external cyclone/
filter sysfan
i. temperature 900 + 50=C
j. pressure less"than 15 psiq at all
vessel locations
k. materials
- vessel Hastelloy x*
-auxiliaries 304 L SST
* Product of Cabot Corporation
The major design problems have included: the requirenent for a specialized t'on of the exces-
sive irass of fine particles elutriated from the bed; and development of a vife. reliable neans of
operation using automatic controllers.
Scope
The development work described has been done prircipally in a turner of 0.20 n internal diame-
ter (Fig. 3) and proceeds from development documented by Stula, et alt-) arrt Young (0. A 0.4 m
burner (Fig. 4) has also been recently activated and is being operated in conjunction with the O.Z m
burner test work.
These bi-rr.ers are engineering-scale systems for a planned larr-e-scale demonstration facility.
The work is funded by DOE'S Division of Waste Management. Production and Recycle.
Significance
The direct importance of this work is the verification of a difficult process which is vital
to the recycle/recovery of uranium in the HTGR fuel cycle.
256
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Burner showing induction coiK
viewed through safety glass window
Full scale burner
drawing
Local instrumentation
panel
, .
..X
Figure 3. 020 m Bufrm Cabinet
257
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OH GJS f>Hf Cyclone
Fttrl S/sttm
B? r-urSiSHK-jr"
R ' , -H^PJife
t ' /""""nsf --',?>
p/ *- -i.^
Figure 4. 0.40 m Burner Upper Section
258
-------
An indirect significance 07" this work involves the nuch wider applications of fluidized-bec:
combustion and/or gasificv.ion of coal in po;/er plants. The techniques of process control, fines
recycle, removal of ash, avoiding particle fusion and the results of heat transfer, scale-up, and
autcration studies apply readily to such uses of atrospheric and pressurized fluidized-bed combustion
of large particles.
SUMMARY Of RESULTS
Cas Pis tributor Design
The gas distributor design {shown at the botton of the burner in F~'q. 1} was a product of de-
velopment experience with flat perforated plates and single gas inlet cones. It consists of a per-
forated cone and vertex line. The cone orifice size and an'ile were experinenta'ly optimized to
yield even gas distribution through all perforations, limit plenum drainage, maximize solids nixing
at the plato surface, and rrini^ize plate surface temperature, 'lie vertex line was sized for levita-
tion of bed cdterial during normal operation and for free flowing product discharge. Design calcu-
lations for perforated cones have developed fro<-' scale-up information.
leed hopixr
heating susceptoi &
cooling i*cket -
induction "*
healing coil
cold ait in ป. -
perforated cone ซ
93$ d'S'nhutor J
oซvg*n COp S
rotary valve
... . - _ . . __ insulation
finci
secondary
pneumatic product tiansport
Figure 5. Primary Burner Configuration
The general gas distributor functions considered in design were: to support the bed and pre-
vent excessive solid naterial fron flowing down in to the pleriun charbcr; to distribute the nas uni-
fomly over the bed cross section by designing sufficient pressure drop across the plate while main-
taining a naxinuri pressure drop to lir;it the bacr-Dressure on the gas Supply system; to propote rapid
particle rovenent throughout th<" bed and at the plate and minimize solid; stannation on plate. Des'cn
problens specific to the prirarv burner included the cyclic burner operation expected (rultiplc start-
ups and shutdowns). TMs required a distributor v.hich functioned well over extreme ranges of gas
flow. Avoiding stagnation areas of the burning solids and resultant hot spots in Mich varied opera-
tion was further conplicated by the wide rar.-:e of particle sizes and densitir'. in the nixed heavy
metal-carbon beds. The distributor design also had to be within the constraints of burner confiqura-
tion dictated by the mechanics cf rerote hardware, r.inimization of high raterial stress and wear areas,
and safety.
259
-------
Conventional flat perforated plates were tested on the primary burner. Although relatively
simple in their design and fabrication, a disadvantage of flat distribution devices was that solids
stagnation zones were liable to form between the distributor holes and at the wall-plate juncture.
The bed stagnation and resultant fusion of bed solids into defluidized clinkers, or agglomerates,
caused flat distributor plate inal-distribution and hot spots in pri:r.ary burning(-;. Theie problems
with flat distributor plates were a result of the horizontal geometry of the device and the resul-
tant potential for stagnation of solids.
The subsequent distributor designs tested departed from the flat configuration to circumvent
the distributor stagnation failures. These designs utilized sloping surfaces to avoid burning solids
stagnation. A 50" included angle cone with a single vertex inlet gas opening was tested initially.
The device worked adequately with feed containing material '; 4760 \m. However, when a nominal crusher
product containing ป. 5 wt. 1. > 4760 \-.m was fed, intermittent spikeo fluctuations of the bed pre* ire
drop (;.P) occurred. These fluctuations were more than twice the magnitude of the recorded slugg. i
bed ;.P and occurred on 1 min. cycles. The bed ;.P cycles were followed by excursions in bed temp, *-
lure and off-qas compositions and resulted in formation of agglomerated fuel particles which adherec
to portions of the sloping v.all of the cone. It was concluded that these operating problems were
due to cyclic internal gas spouting caused by the intermittent accumulation of the largest feed ma-
terial in the cone distritutor('-). The bed material was apparently in continual transition between
spoutable and unspoutable conditions as described by Hathur(^) and Ge1dart(7). This spouting con-
dition could not be allowed in the combustion process due to the local stagnation of burning solids
at the cone surface radial to the internal spout.
The merits of the flat perforated plate and the single gas inlet cone were then combined to
overcome the problems inherent to the separate devices. The resultant device was a perforated cone
with a gas inlet through the cone vertex opening. The vertex gas inlet opening was sized to levitate
the largest bed material with -. 25 vol. 1, of ป.he total inlet gas flow. The remaining 75 vol. 1 of
the total qas flow was allotted to flow to the plenum chamber and through the cone perforations. The
device worked well with the largest nominal feed^-'.
An experimental program was then undertaken to determine the optimum included angle for the
cone. Included angle cones of W, 90", and 120" were used in ccr.bustion runs at similar conditions.
The 90ฐ cone was selected as the steeper 50' cone surface temperature ran 100ฐ to 150ฐ hotter than
the 90" and 120ฐ cones, and the shalV^.-ir 120ฐ cone produced solids agglomerates which fused to the
cone surface around the periphery of several orifices^).
A 90" perforated cone has proven effective in hundreds of hours of combustion operation on the
0.2 m burner, including many start-ups and shutdowns. A scaled-up version has been operated in ini-
tial shakedown of the 0.4 m burner and has also given good performance. Distributor design from (10).
Fines Recycle System Design and Testing
Disposal of the large mass of elutriated fines ('. 30 wt. X of the carbon fed) Kas been accom-
plished by returning the fines to the burner for combustion in nultiple recycle pasces. Three types
of fines transport systems have been used: 1) a dual-parallel pressurized hopper rtense phase system;
2) a tri-series pressurized hopper donse phase system; and 3) a single gravity hooper-rotary air-
lock system. The fines transport systems are shown in Fig. 6. Parametric analysis has indicated
the main effects of each system on the process and has shown that the gravity system is the best al-
ternative.
Equivalent fines burning capacity was noted with the fines recycle injected through the bed
via the cone vertex gas inlet and with the recycling fines injected above the static bed into the
bed solids disengaging height. However, in-bed fines recycle drastically increased fuel particle
carryover into the recycle loop, thereby breaking more particles and increasing non-combustible
heavy netal oxide buildup in the fines. Agglomeration of fuel particles also increased with in-bed
fines recycle. In addition to these negative a^jt-cts, final bed carbon was highest with in-bed
recycle, indicating poorer bed burnout, ffi neat transfer showed an apparent increase with above-
bed recycle as minimum radial and axi>'1. temperature differences were found in the runs using above-
bed recycle.
'oOth the dual-rv.i-dllel and the tri-series pressurized hopper systems were capable of trans-
porting fines o^iinst the back-pressure of in-bed recycle, with the tri-series system having the
potential o; being smoother in mass flow properties. In-bed recycle of the fines aggravated the bed
sl'-'jing, however, causing the deleterious effects mentioned above. Both of the pressurized fines
transport systems were studied in injection of fines above the bed perpindicular to rising slugs.
This injection seemed to destabilize slugs and, hence, was beneficial in its effects on the burning
260
-------
gravity
single-
hopper
system
recycling fines from
cyclone-filter
- gas in
rotary valve
fines
presurized
dual-parallel
hopper system
to above-bed
or to in bed
recycling fines from
cyclone filter
i
pressurized
-gas in
fines
recycling fines from
cyclone-filter
pressurized
tri series
hopper system
to above bed
gas m
pressurized
Figure 6. Three Fines Recycle Systems Tasted
bed. As fines burning efficiencies were equivalent with above-bed or in-bed pressurized recycle, the
possibility of a simpler gravity above-bed recycle loop was investigated.
Gravity above-bed fines recycle resulted in the least final fines inventory; fuel particle
elutriation, breakage, and agglomeration; oxide build-up in fines; and final bed carbon of any of
the recycle systems tested. The gravity system was the simplest to operate and its single moving
part (the rotary valve) increased the reliability of the system. This system was capable of feeding
fines with large quantities of fuel particles or fresh feed material. (The pressurized systems were
known to plug with such large materJai.) The fabrication cost of the gravity system was less than
half the cost of either of the pressurised transport systems. On this basis, the gravity system
was chosen as the optimum fines recycle technique for further long-term testing (see Fig. 5).
These results are discussed in detail elsewhere(:l).
Statistical Analysis
A series of tests was made on the 0.2 m burner without fines recycle and then with fines re-
cycle using the varied recycle systems described above. These tests were devised and analyzed based
on a statistical design of experiments utilizing a 'fractional factorial' analysis technique as
described by Hunterl'2). The series of runs has been used to quantify the weighted effects of basic
independent process variables such as superficial velocity, particle size, etc., upon burner heat-up
times, fines elutriation, particle breakage and agglomeration, burner axial and radial temperature
profiles, and off-gas compositions.
The most useful information obtained were recommended operating levels for: the induction-
heated start-up phase; the 'equilibrium' phase of burning with gravity above-bed recycle; and the
final product bed 'tailburning' phase immediately prior to shutdown. These results are summarized
in the following text and given in detail in (") and (n).
261
-------
For the 0.2 m burner induction heated start-up: velocities us?d were 1.10 and 1.22 m/s;
bed weights were 32 and 36 kg; ignition 0^ levels were 100 and 150 SLFK; 02 ramp rates to full burn
rate levels were 2 to 8 SLPM/min; feed sires used either had an '- 980 or -. 1300 :.m average particle
size. The results showed that: minimum total velocity and bed weight yielded the fastest heating
to ignition temperature; using minimum f<2 f'ow rate fฐr ignition and ninir-un 02 ramp rates reduced
fuel particle breakage and agglomeration; relatively large feed size, although not a directly con-
trolled variable, was beneficial in reducing both heat-up time and particle breakage-agglorreration.
For the 'equilibrium' phase of burning with above-bed gravity recycle: velocity was either
1.00 or 1.10 m/sec; above-bed or 'secondary' 0? used was either > 40 or > 80 SLPH; fines recycle
either started immediately after the full in-bed 02 (burn rate) was reached or 30 minutes after-
wards; in-bed Oj levels were either 300 or 350 SLPH. The combined results indicate that for opti-
nura operation with above-bed gravity fines recycle, the fines recycle should be started irmediately
after the oxygen ramp is convicted and total in-bed oxygen, above-bed oxygen, and in-bed velocity
should be matched (i.e., at high burn rate, use higS above-bed oxygen flow and high in-bed velocity).
The phase of operation just before shutdown in which the final bed carbon is reduced to < 1
wt. ">. was referred to as 'tailburning.' This phase was not. statistically analyzed but its essential
parameters were varied and the results compared. The parameters were: induction heated assist
versus unassisted tailburn; allowing either 5 or 10 \ol. - 02 in the off-cas during tai'buming; and
either continuing or discontinuing (Fines recycle during the tailburn. The results indicated that
using induction, continuing fines recycle, and allowing raximum off-gas 02 until off-gas CO was no
longer detectable, minimized final system carbon at shutdown.
Other interesting results included: 1) Much lower fuel particle agglomeration (0.001 to
0.015 wt. ? of final bed) was found in fines recycle runs compared to runs without fines recycle
(0.02 to 0.2 wt. ?.). The apparent explanation was that fines recycle (whether in-bed or above-bed)
improved bed mixing, and thus avoided any local or temporal stagnation which can allow fuel par-
ticles to be fused by eutectic compounds. 2) Use of additional 0^ flow through the vertex gas
inlet (the major 02 flow was through the cone perforations) was beneficial to all stuaied dependent
/arables except the desired decrease in the axial bed terperature profile (.-T).
Sys tern^ Automa t i on
The details of the primary burner automatic control system are discussed hy Rode^:^. A brief
sumary of that discussion is given below.
The burner automatic control incorporates a digital process controller with an all analog
input/output interface. The advantages of sur.h a control system to a pilot plant operation can be
su-TTvarfzerf as follows: 1) control loop functions and configurations can be changed easily; 2) con-
trol constants, alarm limits, output limitc, and scaling constants can be changed easily; 3) calcu-
lation of data and/or interface with a computerized information retrieval system during operation
are available; 4) diagnosis of process control problems is facilitated; arv) 5) control panel/room
space is saved.
There are five principal functions (and a host of less significant ones) to be performed or
supervised by the primary burner control system. These functions are suncarized in Table IV.
The control system selected to perform these functions is a centralized digital process con-
troller called DTO'lK'.'F.Xt manufactured by Rosemount, Inc. of Minneapolis. The .'.'.'W.v;y system is a
hard-wired real-time dedicated process controller designed to control up to 100 loops by neans of
a continuous one-second scan of the lo>ps. A 20-loop configuration with capacity for expansion to
50 loops was purchased for the reprocessing cold pilot plant, as shown in Fig. 7.
Each loop accepts a 4-20 ma dc analog signal from the process and provides a continuous <-20
na dc analog output signal back to tt-e process. The analog (A) input signals are converted to digi-
tal (D) form for interfacing with the computer-based central controller, and the digital feedback
instructions from the central controller are reconverted to analog ojtput signals. The A/D and
D/A conversion is performed by process interface modules (PIHs); one PIK is required for each
P/''7T.Vฃ7 channel that has an analog input and/or an analog output.
The configuration of the control loops is set up by the operator beforehand with pins and wire
links on a functionally diagrammed diode pinboard called the channel function matrix (CFK). The
various control schemes which may be implemented include cascade, ratio, feed forward, calculating,
and multi-variable types. The actual operating parameters for these functions are alterable values
262
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Profess Interface
Modules (PIMs)
Central
Controller
OIOGE VfS Operator's
Console
Chjnne' Function
MJIIH i
Figure 7. Pilot Plant Control Instrument Panels
263
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Table IV. Primary Burner Hajor Control Functions
Function
Fluidized-Bed
Heaf'ng
Post-Ignition
Bed Temperature
Control
Oxygen Flow
Fluidiztng Gas
(CO?) Flow
Vessel Cooling
ง
Method of Control
Induction heater cascaded to
bed/vessel/susceptor tempera-
ture
Adjust oxygen flow
Fqual -percentage control
valve
Equal -percentage control
valve; adjust CO? flow set
point based on ted tempera-
ture to maintain a constant
total gas flow velocity
Adjust coolino air flow and
maintain proper coaling air
blower discharge pressure
Type of Sensor
I.S.A. Type K (Chrotnel-P/
alunel )-Thermocouples
Type K Thermocouples
Precision gas flow-
meter (SWIRLKTTER)
Precision gas flow-
meter (SUIRIKETER)
Type-K Thermocouples/
Flow averaging (A.'iNU-
BAR) Pilot Tubes
Control Criteria
(1) Heat fresh feed or
particle beds to ig-
nition temperatures
(700ฐC) in < 2 hours;
(2) Maintain clanp to
hub :.T < 112ฐC for
vessel flanges.
Control tw<>eratures
within ป 10CC of set
points (- * U).
Flow within * 3T of
set point.
Flow within ป 3" of
set point
(1) Maintain full bed
at SOO'C ซnile bum-
ing at 800 ga carbon
per minute;
(2) Air flow within ป
3: of set point;
(3) Maintain tempera-
tures within ป 101 of
set point (-- ฃ i:).
entered through an operator's control console and stored in an internal ccrory. Restructuring of the
control schemes is sirply a ratter of altering the pin and link arrangements on the CFM.
All loops nay be controlled digitally by the central controller under the automatic one-second'
scan arrangement. This is the normal node of operation, called the 'central' control rode. A varia-
tion of this modo is the 'central eanual' which allows a loop to be controlled manually thrcugh he
operator's console. In addition, there is a corcletely seoarate ar.d independent analog 'Nicl'w' con-
trol node, which is effective even if the ccntr*! controller breads down.
There are currently 23 out of SO available channels (CFM) and 13 out of 20 available control
loops (PIHs) assigned to the priryjry burner during the period of initial development work. Possible
later changes tn the nethod of control to provide greater versatility and more automation will con-
sune up to 39 channels and 16 loops for the prtca'y burner.
The .''''rT'.rr system has been in operation since March. 1976. and has been satisfactory in
control of the primary burner. Minor malfunctions have been Quickly traced and corrected by the
manufacturer. The controller and the control philosophy have been very beneficial in simplifying
the burner operator's duties and in providing s.ifer, more constant operating conditions.
Empirical Scale-up and Heat Transfer Studies
Initial operation of the 0.4 n burner has shown significant bed nixing irprovements over the
0.2 m burner. These results may be interpreted in light of decreased bubble bypassfno at the vessel
wall as the vessel diameter increases, as shown by yertherP'). Internal hปat transfer coefficients
during 0.4 ra induction heating have ranged fron 351 to 602 watts/erOR fn the 5^ An(j 47.7 t0 M2.2
watts/mJฐK In the slugging *one(!'). The highest in-bed Internal heat transfer coefficient was ob-
tained with the shallower beds tested
-------
from the increased slugging of deeper beds (L/D ' 4.5). The coefficient increased with increasing
temperature as velocities were fceld constant. An optinun velocity was found at each constant tem-
perature which gave a naxirxm in-bed coefficient. The optinun velocity was dependent on bed weight
and particle size, with deeper beds and larger particles requiring higher velocities for good mixing
and heat transfer.
FUTURE WORK
Both the 0.2 n and the 0.4 n prinary burners will be operated in concurrent tests which will
include studies of scale-up differences, effects of varied feed sizes, hybrid cone distributor de-
signs, extended automatic control, and equipment wear.
ACKNOWLEDGEMENTS
The work described in this report was supported by the U.S. Departnent of Energy, under con-
tract EY-76-C-03-OI67. The results precede from investigations acco'-'plished by technical contribu-
tors over a span of several yea'-s. The contributors include: D. H. Stallings. N. B. Jacavkh.
N. C. H.irvaez, R. M. Earle, T. 0. Wri-jht and U. S. Ricknan in the fabrication and operation of the
0.2 and 0.4 m primary burner equiprcfit; 8. T. Stula in the subsystem design, burner operation and
data evaluation; H. H. Yip in heat transfer an>J related analysis and design work on botn burner, sys-
tens; 0. S. Rode in design of eปperi-ซ?nts. burner operation, and data evaluation; antf J. W. Allen
and R. 0. Zitrrernan in the leadership of the burner development program.
NOTE: Thป parallel sressurizซ?d fines recycle hows were scaled-up from hoppers originally
tested by H. B. Paltrier of Allied Cheoical Corporation. Idaho Falls, Idaho. Discussions with W. B.
Palmer also contributed to the development of the series pressurized hopper system.
REFERENCES
1. D. T. Young. "The Use of Fluidized 6*d Combustion in HTGR Fuel Reprocessing." Proc. Fourth Intl.
Fluidized-Bed Combustion Conf.. McLean. 1975, pp. 657-659.
2. W. S. Rickrcan. "Interin Development Report for Secondary Burning." GA-A13540. Dec. 1975.
3. R. T. StuU.et al. "Interim Development Report for Prirary Burning." GA-A13546. Jan. 1976,
pp. 6-1, 6-2.
4. "HTGR Base Program Quarterly Progress Report for t^e Period Ending Feb. ?S. 1973. W-A1251S.
March 30, 1973. pp. J6--9.
5. "Thoriun Utilization Proorar. Quarterly Progress Report for the Period Endi
-------
13. "Thorium Utilization Proqran Quarterly Progress Report for the Period Ending Aug. 31, 1976.
GA-A14083. Sept. 30. 1976. pp. 4-3 to 4-11.
14. J. S. Rode. "Process Control of an HT&R Fuel Reprocessing Cold Pilot Plant." GA-A14101.
Oct. 1976.
15. Joachim Herther. "Influence of the Bed Diameter on the Hydrodynamics of Gas Fluidized Ceds.
A1CHE Symposium Series on Fluidization. 70. 1S74. pp. 53-62.
266
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QUESTIONS/RESPONSES/COMMENTS
SHELTON EHRLICH, CHAIRMAN: Can we have the lights? Okay, are
there any questions? Any written questions? Any oral questions?
Come on over to the microphone. Identify yourself and the company
you're with.
MR. MARGARITIS: Paul Margaritis with Westinghouse Electric.
You said something about...when your bed spouted that you got agglo-
meration on the wall. Did I hear you right? Could you tell me
a little bit about what the mechanism is that causes the agglomeration
on the wall. Do you think there's hot particles in a bubble that are
getting thrown 'jp against the wall or do you think that just raising
the oxygen content on the wall or what.
MR. YOUNG: It's simply stagnation. All the gas is going up in
the central spout. It's not distributing itself among those solids,
and it is stagnant at the wall. The hot solids--of course, a lot of
us might suspect that solids temperature in a fluid bed is a good
deal hotter than the overall bulk temperaturestagnate, become a
packed bed and really get hot. It simply fuses everything where
the gas isn't distributed. So it's gas maldistribution which results
in stagnation outside that spout, all annular to the spout.
MR. MARGARITIS: Thank you.
SHELTON EHRLICH, CHAIRMAN: Any more questions? I feel kind of
a special relationship with the work that was just described. A
little historical note: About ten years ago when I was working in
Alexandria, Virginia on fluidized bed combustion, one of my associates,
John Bishop, got a telephone call from the General Atomic people
asking for all of the information we had on fluidized bed combus-
tion. So very quickly we wrote them a letter proposal, and never
heard from them until today!
267
-------
Rosoarr.h of Gas Combustion in
od Bod Plants
K /<: Math^rm JififJ A F/ Gliji'fi
IriMitut': O' G'j'. ol UK.- Uf r^i
U lo,?i R
H uit.'i ... tin.- \-.ti\\. t h.ir.K ' i r i ic "I the::,
(.u::i!<*,i-. I i'!" .1 tiปj::.ซij i-neo-j1. ,'>:. -:nnl-;i i r trix.lure in jccli-'i liirou; :i .1
i ซ.::.!.u:;l inn '.;.i r irn :il.i i sy.le:::. iiii" :: ::.
l.il.m'.il "iry -Hid i i: *"
Air
.
0 ,0 0. 0 O
r,.,'..
':*-*?&
'::'ฃฃ*'
** :'/>.'. '-?.'
t ill
u
t L
Air
Figure 1. Deiigrn ol Air and Gat (nation into the Fluidind
Bed. a Homogeneout Gn and Air Mmture: b<
Partial Mixture ol Gas and Air ounide (he Fluidi/ed
Bed: d e Separate Jell ol Gas and Air.
Comh'i slion ซ'f Homogeneous G.is-ond-Air Mixtures
Tlic process of con;bust.ion of hoHior.cncJus R.is-nnd-alr mixtures was studied usin;-.
100 and 160 trm dia laboratory furnaces with an air consumption ratio equaling 0.7- ' 2 . 0 .
and pressure of 1 to lt> aim.
Corundum balls, htf,hly-aliminifcrous ch.iraottc and quart^y sand were used as
pscudof luiai.-.ed heat-carrying agents.
The most favorable conditions for pas combustion arc crc-itcd during the injection
of a previously prepared j;as-and-air r-.ixture into the fluidized bed. Keeping up .1
constant source >/f flame over the fluidi.~-.ed bed of monodisperscd natcri.il, rhe natural
gas is observed to i>;nilc over the fluidized bed surface burning out with a bright blue
268
-------
: I "IT t- ". :*. '. r.c e<... .1:..-;:./ : j: :-11 .s .i:*.c ~; r.iy s .
: r t T.c :.*%;. .-.,.u" vr.v t i-r-j-c-r.**.'.ii'i ป:K-
s tr.e te::;-ซ. r :iure -. : the up:-er p.irticles
I.IM-S t r. 1 ;<.i.il'C tht- i::>:: has I ;<.-:< process
n,- r.ipidlv i.vr.ti'.iv.-'L.eU insi>:e trie i!ui>:i.:i
: i.-i-.-.'i-r.i*.
v:.:^!-. is ;. i-.i- j.j j*'C l..--.ซ-r '. h.in ihi.- i.-niii--:-.
vii r.i"!i>uis;>crsni p.irt ic it-s . KIT i-x.i: .;>!ซ. na:ur.i'
. .IK is ii'ni'.i/ii a: : i.i-^i n: o.j'j-'J.j ;ir.>: i - .! rrr. spin-r ic.ii corur.dum
\i;tf. it !i -s . vi-iU- in .. :mi <>:" O.'io-'i.b :'.::. fr.n.-; iop. ^-i:.-, ::.<: iซ- p.irl icli-s . .'..IN ซ.-n::itiuat. ion
iii.'in:> .v_ 7'JO"C. 7::is IN .iccoui.'-t-o f>>r L.V t hi- I.icl til. it iluriii)' jisi-tiilof luiui.:.iLiปr. ol
;'i-:yซii s;n-r>.i-is ::-ili-ri.'i! -.ซ-ri- is .t r.ir>-:ii->: .-...in- of i.c.iic-ii ;>..i:icit-s lio.itin-,' <>v>.-r thL-
:-*-!-. -'i'.ich ser'.'t-'-s -is i i'.c- . *.".ป. *.-'. /.ir. \.'tj:'.i".ij! inr..
r'i/uri- .' ;.h"V
v.-r.-,us 'hi; -Ii-;.: h <.f
Oi ,
.. '.<>. :: i . Cii.-, i -tini-ent r.n. ion and the ten.per.i
':.- the *_ซ>;. iii:st ion of .1 hoiM>;-ei!i-.n:s i-as-and-a
: ':-:'-;.ire i:: .1 'l.-'i-\.'j "::. bet: of >.-or;::i>!.::-.. i-. i : .;-:owii Th.it the >-.'is hi,-ins to lปurn or.
' :.i- ,-r.itt- *..i-:i the- -i i r cor-.stt".p: ior. r.it. iu - ' i . ! . "."ne v.-or.ci-p. tr.'it ion of :-.!>-t iiane .'ind TJX
<:> cri-.i^-e r.ipidiy :'. r. the i.i-i.-ht of tin- :'! \t i t; i .-.<-,: !n-tj. while those of c.'irhon ilioxitit-
ir.i-Te.'i.se. Tin- c'.r.pos: t io;i o: the n>:-.l.'J:.t icr, pr-uiuc: : :-i-i-o:-.i .-. st.iiu 1 i.'.i-ti -it. .1 tlist.-ir.ce
of 1'J r.:: fn-:-. ::.ซ.- ,-r.iTi.-. !' i :, i.i-n- th.t. thi- t_o::.h-..M inn .:<>.;e end:;. The ilepth t>f t!u
.iciivi- .:oiie is '. r.r... h.ere. ">'i of the , i s !>-.irr.s o-.il . iit-.it r.iiti.it iun in this :irt.-.i is J
;ij.-h .1.-. i ''il>. -)';'i. 'iii') kc.il/r.' per I'.our . 'li'.t- p.-.e'nii'f 1 uitli .u-ti p.-iriicles h:ivi.- no li::.-.- to
fon.jui't ; he r.u:i:it id hc.r. , wiiivii re-.-.ilt.-i ::i -i J'jO-'j'J()"<- t er.j.i-r.iiure ;-r.nl icr.l .-ij>pi-.ir in.-
on the deptl. of the co:-.:>u.-.t i on .:i.i:e. '.-.'i'ii the -iir i-or.:.ur. pt. i on r.it it: of - l.b. the ;
:M-,-in:. tu !mr:i o'j'. in '.i-.e h-ihhli-. .iiul in the rarefied .-.one of the fluidi:-.ed l.*
3
DISTAKCi
16 24
X CRATE
Figure 2. Changt of Temperature (a) and Composition of
Combustion Product! (b) venut th* Fhiidi/ed Bed
D*pth with the Air Contumption Ratio Equaling
a . 1.1 .1 H2. 2 CO: 3 CH4; 4 - Oj. 5 CO2-
The temperatures
the particles. The it--
equalize os a result of
honwgencous mixtures in
because of bad exchange
fact that accounts for
fluidizcd bet' when the
dynamically no nitrogen
vn dy ir.c thurr-.ocouple do not always correspond tป> those of
pcratures of the gas and the "packets" are seen to quickly
intense interphasc heat exchange. During the coi7.bustion of
the bubbles the temperature in them reaches very hif.h values
between the bubbles and the particles (Figure 3). It is this
the }-.esence of nitric oxide in the cor.bustion products at the
thermocouple shows 1300-1400ฐC and at such temperatures thertr,o-
-oxygcn reactions are possible (Figure 4).
269
-------
Combustion of a preliminarily prepared gas-and-air mixture can be organized
both in experimental systems and in szali-size industrial furnaces used for fluicized-
bed heating of inecal to tenperaturts of 900-130GฐC.
Combustion of Gas Partially Mixed with Air
%
It is possible to organize fluidizod-bed combustion of gas partially mixed with
air by injecting the gas-and-air mixture into the ted through the tuyeres (Figure Ib) .
or through the grate (Figure lc>-
Studies of fluidized-bed combustion of gas partially mixed with air were conducted
with the use of 300 ind iOO trim dia experimental furnaces where 0.1-0.5 m fraction
river-sand acted as the,heat-carryin th*
Air Coraumption Ratio Equaling a' 1.1.
1 bubbto; 2 bubble trace; 3-4 gas;
0 0 experimental data
Figur.4.
oncentration in Natural Gas
ranafMuiun Product*.
-' - 1 atm; 0-0 - 5 atm; - - 8 atm-
ซ-ซ-tta?m; a . Q . tSairo
270
-------
The burning ouc of '.he combustible components is not always completed in the
zones of circulation, this process depends on the bed temperature, primary air consump-
tion rate and the nature of pseudofluiuized material; that is why there is always a
snail amount of methane and carbon oxide present in the combustion products over the
bed (Table I).
Table i. Data on Gas Combustion in 500 rat
-------
gas conbustion in the bed sets in when the- temperature of the fluidized bed is as hi^.h
as 650-870&C. At this temperature there is no hydrogen, carbon oxide or no thane- in the
combustion products if the air conburr.ption ratio of gas con.bustion is 1.3-1.7 and primary
nir consumption on the tuyeres is 35-50/i of the steichiometric value. With the increa.se
of secondary air temperature from 100 to 50GฐC co-plete fluidizcd.-bc.-d combustion of gas
is obtained with o = 1.1-1.2. With the tuyeres method of t'luidizcd-bed cor.buction of
gas. a l:iG-2000C thermal gradient is observed to develop 'uctwcen the combustion /.one and
the fluidizcd bed core-.
Gas Combustion with Separate Injection of Gas and Air into the Kluidized Bed
The technique of fluidized-bed conbustion of natural gas with separate injection
of gas and air was developed in experimental and operating industrial furnaces.
dia furnace the process proceeded in the following way. -The air was
1.J-1..I. v i :ปuii L uu:>ervui.iu[i:> :iiiuwi u i-iiciL \>m n.>u i c i ugg ir>u ,
stretched tongues of bright yellow flarae vere jumping, through the lluidizcd bed. their
form being quite different from that or the ,:as bubble. The flame moved both in the
vertical and horizontal planes. The natural gas was burning out above Che bed.
In a 2.75 m dia industrial furnace, the natural gas was injected into a fiuidi/.ed
bed of 3-10 nan limestone particles under a pressure of i atiri and at a rate of 1700 m-Vhour
(Figure Id) through 20 no;:z!es of 7 ran diar.ซ.tcr. I'he nozzles were accommodated on the
periphery of the furnace shaft 200 mm above tin- grate. The air for pscudofluidization
of the particles and gas combustion was injected through the ('.as-lining grate.
The results of '.he test showed that the li:ngth of the jet was 0.5 "'. In the
circular space around the wall a surplus of j-.-in wa.% developed which burned ou'. above
the bed, leaving the center of the furnace without any g !.ป. The substitution of the
cylindrical nozzles for Lavalle nozzles did not improve the completeness of gas ,_ mbuslion.
The failure to obtain fluid i.-.-il-bed combustion of the i
through the peripheral tuyeres is n.-!_ropซ.
the air. The air injected into the bed through the gas-tiir.in,
between the particles in the "packets" and comes to the surfac.
as when it is injected
r mixture of the gas with
grate undergoes filtering
in thซ form of bubbles.
the movement both of the ซ:as and the air bubbles coinciding, neither in time nor in space.
Tlie process of gas and air mixing in the fluidi^ed bed can be improved by injecting, the
gas into the grate zone in the form of thin je^s striking each other. This technique
was used In a 3.8 m dia furnace for roasting antlmonic ore tails.
The gas was injected into the bed through 56 gas caps uniformly fit'.i-d on the
sole of the furnace (Figure le) with a rate of 900-1000 m-Vhour. tvory gas cap was
equipped with 9 air caps situated on a 3^0 mm dia concentric circle around It (Figure 5;
Figure le). The rate of air consumption through each cap was 18.5 la-Vhour. Altogether.
504 caps were installed on the gas-timing sole.
The gas and air jets were directed toward each other, the difference in their
heiphc being 80 ran. Such an arrangement prevented gas combustion in the space between
the caps. During the two months of the furnace's continuous operation the process of
fluldized-bed combustion of the gas was stable, no lumping of the particles or burning
of the caps was observed. The temperature ranr.c of the bed 500 rai above the prate and
at a 200 ran distance fron the wall was maintained at a level of 1000-1020 C.
CONCLUSIONS
An analysis of the work of industrial systems shows that gas combustion in large-
size furnaces can be organized by using the technique of injecting tha gas, ,>artially
mixed with the air outside the fluidizcd bed. through the tuyeres or by injecting t'.it
gas and the air into the bed by separate jets. Such furnaces are of simple structure
and warrant reliable operation.
272
-------
1 '
8 OPKHIKO.: 03 ป
9 CAPo
I
1
o
1
.'
/' '
T. *
*
'ol
i
f
i
3
040/25
t
.'-',-,'
>
t
- ^
0| 3 OP5iKlM3
"~
'
' . '.
oi
i
t
4
ป
.",;
** *
*
1
AliV
t All,
Figure 5. Structure of Gat-Timing Grate with Separate
Infection ot Gas and Air into the Fluiili/ed Bed
273
-------
INTRODUCTION
ROBERT CHRONOWSKI, CHAIRMAN: We've got to continue our tribute
to the late Professor Elliott here. Our next speaker is Dr. J. R.
Howard fron Aston University. I v/ant to read a little bit from
his own biographical sketch.
Dr. Howard's interest in fluidized bed combustion and heat
transfer was first stirred by his talented colleague, the late
Professor Douglas Elliott. Dr. Howard continues to teach at Aston
University and is involved in fluidized combustion and heat transfer
esearch. He has a long history of research and work with industry
and combustion activities, power plant design, temperature control,
and production engineering. The paper deals with combustion experi-
ments within a rotating fluidized bed. The coauthor is Mr. C.
Metcalfe, who is sitting for his Ph.D. at the University now under
Dr. Howard. Dr. Howard?
274
-------
Combustion Experiments within a
Rotating Fluidized Bed
C. I. Metcalte and J. R. Howard
Department of Mechanical Engineering.
University of Aston in Birmingham.
Gosta Green. Birmingham. England
ABSTRACT
Rotating fluidized bed combustors offer the exciting prospects of raising tl-o
rate of heat release per unit volume of bed to dramatically high levels and extending
the range of 'turn down' significantly. 1'esults obtained with a gas-fired fluidized
bed coinbustor are discussed briefly and followed by an account of experience obtained
when burning coal, (anthracite). The latter work exposed some of tho problems to t^e
ox'ercome before the lull potential of this combustion system can be realised. Start-
up, control and segregation problems are described and the direction which future
work should take is discussed.
NOMENCLATURE
c Specific heat of sand particles
(CV) Calorific value
H Geometric mean particle diaroeter
g Gravitational acceleration
B' Ng
Ga' Modified Oalilco number
Q Thermal output power
Ko , Reynolds number at minimum fluid!?.-
mf ation
t Time
T Bed temperature
V Fluidty.jp * velocity
., . ,-[.3 t' ป Minimum fluidization velocity
~~~~ 2 c Particle density
u p
mc Mass flow rate of anthracite ฐf Fluldlzing air density
n.p Mass of sand particles in the bed " Viscosity of fluidizing air
N Integer
INTRODUCTION
Because the particles in a rotating fluidized bed have a high centripetal
acceleration imposed on them due to rotation of the entire bed about an axis parallel
to its free surface, as shown in Figure 1, high fluidization velocities are possible
without nlutriation of particles from the bed. The resulting increased mass flow of
air allows a greater mass flow of fuel to be burnt, hence a greater rate of heat
release within the bed. For a given thermal output, a rotating fluldizeii bed cotabust-
or will be much smaller than a stationary type.
In stationary fluidized beds, combustion can be controlled by altering the
fuel feed rate and fluidizing velocity. The operable range of fluidizing velocity is
limited on the one hand by the mininun velocity at which the bed is properly fluid-
ized and on the other by particle elulriation. With the rotiting fluidized bed. the
operable range of fluidizing velocity can be extended by increasing the rotational
speed to prevent particle clutriation and reducing the speed to lower the minimum
fluidizing velocity.
275
-------
Pneumatic Solids
Transporter
Air Feed
Fluidi/ed
Particles
Plenum
Chamber
Air & Propane
flotameler
Propane
Supply
Coal Hopper
Coal Feed
{
\
*ir Supply
Fan
Pressure
Tapcmg
HoikM Drive
Shaft
Rotary Seal
Figure 1. Schematic Layout of Rotating Fluidiicd Bed Conu-
triation. This limit was then raised by increasing the rotational speed so that the
fluidiซ:ing velocity could be increased. This method of increasing the fluidizing
velocity was eventually limited ty the residence tinve of the propane/air cixture
within the bed becoming too small for the necessary pre-heatinK to ignition temp-
crc'uro. so that the combustion reaction could tako place within the bed. This
condition will be referred to as 'flame out*.
276
-------
The r. i r. i - urn thermal output of the corsbuslor '.vas 1 imitr-d . conceptually., by:
(a) The largest air:propane ratio r.t which conbustion could bo sustained.
(b) Th<- i^ininun i luidiz.at ion velocity at the lowest rotational speed aซ -.i
a reasonably ur.ilors radial thickness of bซ"l would occur. (The shape of the I rซ-e
surface of the betl -xas parabolic at low rotational speeds.)
(c) Fla=ซ-- preparation velocity exceodinK the ' luirli/ai ior. .loci tj . res
in burning upst.-eae of the distributor.
(d) Vass. flex of mixture through tin- distributor iioinj; insufficient to kr-.-p
the distributor cool.
In j-ractice, the f luidi ?.at inn velocity v.-as not alloyed to fall below 1 i vr-
tir.es that rซ-q-ired for incipient, f luidi zat ion of the !> -I . Thus < < ) and (d) dictated
the t-inir.uri practical operatinc the ma I out pi.. . Xซ-vซ-ri nป- !<-ss . the rorr.bi n" 'turn down' rat:-i to
5:1. co.-pared v.-jth about 2:1 ntit ainab ! lith sir. pie stationary l.i-ds . ( ^ ' ) A t;.p:val
oporatine envelope is shov:n in i'ip;ure 2.
The experir.enta 1 work en burning propanerair -ixiur'-s in rotating
bซ-d cor.bustors r.as also der.onstrat<-'J tin- capacity of this >-.ystซ-m >o burn mixtures
which are si p.n : : icant ly weaker than the publ isheil' f la.Tjr.ab i 1 i t y ii^.it ' >) . thus 'jr,-.-rii ni:
up the possibility c! burning w.vak inilustrip) fur."-.s wi.. h a IOV.IT ccnsur.pt ion "I .v.:x.l-
'ary jas. <">
80
70
I 60
1 50
u.
< 40
ง 30
20.
10
Loss of Fluidi/od
Bed Combustion
Maximum
Tempe'dJuie
Limit
_
Data
0 1.0 20 30 4.0 SO 60
Hot Air Velocity (nv'sl
Acceleration
Particle Type
Site Range
Mass
Distributor
: - 30xg
: Silicon Sand
: 180 250m
: 600g
: Conidur No 10
Figure 2. Typical Operating Envelope
277
-------
COAI. COMBUSTION
Te^t Equipment
The rotating fluidized bod comhustor used for experiments described above v:as
u.s---d without modification apart from including a coal feeder. A schematic arrangement
of the tost facility is shown in Figure 1. Thr- distributor was raade of perforated
stainless stnel sheet rolled into the shape of a hollow cylinder 200 mm diameter and
67 mm axial length. Tho lower plate carried the hollow drive shaft through which
fluidizing air was passed to the plenum chamber surrounding the distributor as des-
cribed in Reference 1.
A water-cooled probe was located in the exhaust gas duct through which gas
samples were drawn for analysis. A paramagnetic oxygen analyser and an infra-re'i
CO/CO,, analyser were used to determine concentrations and, for corroboration of these
analyses, a gas chromat.ograph calibrated with gas samples of certified composition
was used.
The coal feeder was of the vibrating platform type, fed from a hopper bolted
to the platform, the coal feed rate being varied by altering the amplitude of the
vibration. The coal particles dropped into the throat of a pneumatic particle trans-
port duct arranged to blow the coal particles through a nozzle on to the free surface--
of the rotating fluidizr-:! bed. The position of the point of coal injection on to the
fluidized bed free surface wan adjustable so as to be able to find the most suitable
point at which to inject the coal.
Initially, GOO g of silica sand particles of size range 180-250 urn were used
in the rotating fluidized bed. The type of coal used was anthracite, crushed and
sieved to a size range GOO - 1000 urn and of proximate analysis, moisture IT., ash -1.5%,
volatile matter 10%, fixed carbon 84.57.. The choice of the latter size range arose
because this produced excellent combustion in stationary shallow fluidized beds and
the anthracite could be crushed to this size economically.
Initial Start-Up Experiments
The combustor was started from cold in the normal way, being fuelled by pro-
pane and allowed to reach a steady state with the bed temperature 950 C. The super-
ficial fluidizing velocity was 1.1 m/s and centripetal acceleration of the bed part-
icles 10 x gravity. When anthracite was fed ปo the bed at the ra'.e of 50 g/min.
(sufficient to give a thermal output of 25 kW), excessive elutriation of fine anthra-
cite particles occurred. Clearly, the rotational speed hart to be increased, hut even
when the centripetal acceleration was 15 x gravity, the reduction in elutriation rate
of the fines was not very substantial. However, the major part of the anthracite
particles reached the bed.
The propane flow was t..en gradually reduced in an attempt to maintain the bed
temperature constant. But even when the propane was shut off completely, the bed
temperature continued to rise, reaching 1140 C, at which point the anthracite feed
was stopped. A non-uniform distribution of anthracite particles could be seen on the
free surface of the bed burning with a yellow flame. More uniform anthracite dist-
ribution was achieved by movinc t!:e coal injection nozzle from the lower end of the
rotating fluidized bed to the upper end. It was also noted that if the exit velocity
of the anthracite particles from the feed nozzle wei-.j increased sufficiently, local
breakdown of fluidization of the rotating bed occurred, and as Figure 3 shows, anthra-
cite became centrifuged towards the distributor where local fusion of particulate
matter had taken place.
These difficulties demonstrated that operation of the above start-up procedure
which has hitherto been found adequate for stationary shallow fluidized bed coal
combustoi-s, is decidedly inadequate for rotating fluidized bed combustors. Two other
manual start-up procedures were also investigated but neither gave satisfactory trans-
ition from propane fuelling to self-sustaining anthracite combustion. However, for
the sake of c ,:tipleteness, the results are reported briefly below. Commencing from
steady state combustion of propane, the procedures and results were:-
278
-------
lllljllll llllllllllllll
M/H ป
ir
i
figure 3. Fused Paniculate Matter from Near the Distributor
(a) Feeding of anthracite commenced at the rate of 50 g/min for one minute.
The propan*? supply was then shut off quickly. This resulted in rapid reduction in
bed temperature and cessation of combustion. During the period of simultaneous
supply of propane and anthracite, local fusion of sand and some clinkering occurred,
disturbing the mixing of particles in the bed.
(b) Simultaneous commencement of anthracite feed at SO g/min and cut off of
propane supply resulted in quenching of combustion alone.
In order to make progress and explore the mechanism of transition from propane
to anthracite combustion further, some ancillary tests to examine particle mixing and
bed cooling characteristics were carried out.
Ancillary Tests
With the bed operating on propane at 950ฐC, superficial fluidizing velocity
1.1 m/s and centripetal acceleration 15 x gravity, the propane flow was shut off
suddenly and the mass flow rate of fluidizing air held constant. The exhaust gas
temperature was recorded continuously for about 10 minutes. The average rate of
279
-------
cooling during the first 30 seconds was used to estimate the mass flow rate of anthra-
cite required to maintain the bed temperature at. 950 C fron equation (1).
rapCp 0-30s - *c (CV) (1)
This simple relationship however ignores the heat input required to the anthra-
cite particles from the hot bed to raise their temperature to ignition point. It also
says nothing about the ma.ss of anthracite u-hich, at any desired thermal output from
the combustor, must reside in the beci. Despite the lack of these data, it was decided
to c:irry out brief start-up trials at anthracite flow rates estimated from equation
(1), to obtain further insight into their effect on the transition process. The
results of these trials are discussed under "Further Start-Up Trials" later in the
text.
Mixing Anthracite and Sand Particles
A half-scale perspex model of the cotnbustor which was capable of imposing accelera-
tions of up to 30 x Gravity on the particles was used to investigate particle mixing
patterns visually with the aid of stroboscopic illumination. The independent vari-
ables were:
(a) Sand particle size
(b) Fluidizing velocity
(c) Anthracite particle si*e
Anthracite and silica sand particles v.ere selected so that each would have
similar minimum fluidizing velocities. The correlation formula used was equation (2)
below, which is due to Wen and Yu.<7>
Remf = (33.72 + O.0408 Ga']S - 33.7 (2)
Equation (2) predicted the final particle size ran^e choices as:-
Silica sand 250 - 355 urn
Anthracite 300 - 500 um
mixed in the proportions 20 grams of anthracite to 200 grams of silica sand in the
bed.
Figures 4a, 4b and 4c are stroboscopic photographs obtained with the perspex
model of the rotating fluidlzed bed combustor with the above particle mixtures at
various fluidizing velocities >vhen the centripetal acceleration on the particles was
10 x gravity. It will be seen, Figure 4c. that at a fluidization Index, U/U ... of
about 2.85, the mixing of sand and anthracite particles is very uniform and there
Is no elutriation.
Conclusions from Initial Start-L'p Experiments and Ancillary Tests
1. Anthracite feed rates calculated from equation (3) below:
Q = Ac (CV) (3)
used for the initial start-up experiments are too large to effect transition from
propane fuelling to self-sustaining anthracite combustion.
2. Selection of both sand and anthracite particle sizes can be critical
because uniform mixing of anthracite and sand is essential for the avoidance of local
hot zones and fusion of particulate matter.
280
-------
Fijdre 4a. Slroboscopic Photograph of Free Surface
U/Umf 0.9 Umf 0.4 m/i
More generally, differences In density, size and shape of particles tends
to promote segregation. Radial segregation seems to be accentuated In high gravlta-
lonal fields, while axial segregation is affected by the positioning of the anthracite
feed point. Where possible, the former differences should be kept as small as practic-
able because good mixing is desired.
3. In the above tests, the radial thickness of the bod was almost uniform to
promote uniform fluidization and only about 10 mm thick in order to maintain a low-
pressure drop, (<150 mm water across the bed at 15 x gravity). However, it has been
shown in rotating bed experiments at ambieiit temperature by Demtrcan and Swithenbank(s
that very good particle mixing can also be obtained when operating the bed in a tumbl-
ing mode.
4. When running continuously on solid fuels, ash removal becomes an important
requirement and further experimentation is required to establish a satisfactory compro-
mise between good mixing of fuel and sand, segregation of ash and the possibility of
excessive air by-pass and non-uniform fluidization.
281
-------
Figure 4b. Stroboscopic Photograph of Free Surface
U/Umf 1.97 Umf - 0.4 m/t
-------
U)
Figure 4c. Stroboscopic Hhotograph of Free Surface
U/Umf" 2.85 Umf 0.4 m/$
-------
Further Start.-fp Trials
New experiments enabled start-up to be performed using the sand and anthracite
particle sizes found to give satisiactory mixing by the ancillary tests. The new
start-up procedure from full propane firing at 950 C with an acceleration on the
particles of 15 x gravity was:
(1) Shut off propane completely and simultaneously commence feeding anthra-
cite ;t the rate of 50 g/min for the first 30 seconds.
.Mi) Thereafter, reduce the anthracite feed rate to 30 g/min.
Results
Sclf-suf>tait ing anthracite combustion was achieved, but the combustion quality
was not always satisfactory. However, the best dry exhaust gas analysis obtained was
0.1% CO, 0.14* II,,, 8.7% C0_, 9t Op. Localised overheating caused some sintering of
the silica sand, see photograph in Figure 5, but there was no evidence of general
migration of anthracite particles towards the distributor as occurred with the initial
start-up xperiments on burning anthracite.
3
The nominal intensity of combustion was 35 MK/m of fluldized bed.
FigunS. Fined Silica Sand
284
-------
GENERAL CONCLUSIONS
1. Although it was possible to obtain self-sustaining cor.bustion of anthra-
cite by manual control of transition Iron steady combustion of propane, some form of
automatic control of the transition process is desirable, particularly if no heat
transfer surfaces are located in the bed for control of its temperature and the
thermal capacity of the bod is small.
2. There is at present insufficient data for design of a control system for
effecting satisfactory transition from propane fuelling to self-sustaining combustion
of anthracite. The data required are the times required for pre-heating the anthra-
cite particles to ignition temperature and to burn to extinction together with the
bed heating and cooling time constants under rotating fluidized bed conditions.
3. When compared with stationary fluidized bed combustors, the rotating type
has a very low thermal capacity for its thermal output, air and fuel flows. This
gives it the characteristics that rapid changes of bed temperature occur with changes
in fuel flow when burning gaseous fuels which ignite and burn rapidly. When operating
on solid fuels, however, heating to ignition temperature and burning of the fuel part-
icles takes a longer time, so that the temperature response of the bed to a chango in
fuel feed rate is not immediate. After this dflay, the bed temperature can then
change in a rapid and uncontrolled manner. This points to the need for an antici-
ptttory control system.
4. Increasing the thermal capacity of the bed will help alleviate the above
problems but the extent to v.-hich this can be done and the best way of doing so is a
matter for further investigation.
ACKNOWLEDGEMENTS
The work described hero is part of a programme supported by the National
Research Development Corporation.
This paper could never be complete without, the authors expressing the debt
they owe to their inspiring colleague, the late Professor Douglas Elliott, for the
encouragement and invaluable advice he gave. His sad death in June 197G deprived the
engineering profession of one of its most imaginative, creative an~ truly courageous
spirits.
REFERENCES
1. r. I. Metcalfe and J. R. Howard, "Fluidization and Gas Combustion in a Rotating
Kluidized lied". Applied Energy (3), (1977), pp 65-74.
2. C. I. Metcalfe and J. R. Howard, "Towards Higher Intensity Combustion:- Rotating
Fluidized Beds". To be published at Engineering Foundation Conference on
Fluidization, Cambridge, April 1978.
3. C. R. Westwood, (1975), PhD Thesis, University of Aston in Birmingham.
4. K. K. Pillai, (1976), PhD Thesis, University of Aston in Birmingham.
5. H. M. Spiers, "Technical Data on Fuel" The Hritish National Committee World
Power Conference, Lor.don, 1961, p 260.
6. British Patent 1 471 598 (27 April 1977).
7. C. Y. Wen and Y. H. Yu, "ft Generalised Method of Predicting Minimum Fluidization
Velocity". A.I.Chem.E. Journal 12 (3), May 1966.
8. N. Demircan, B. M. Gibbs, J. Swithenbank and D. S. Taylor, "Rotating Fluidized
Bed Combustion", to be published at Engineering Foundation Conference on Fluidiz-
ation, Cambridge, April 1978.
285
-------
QUESTIONS/RESPONSES/COMMENTS
SHELTON EHRLIfH, CHAIRMAN: Do we have any questions? Last
call. This is from Frank Hill of Brookhaven National Labs and the
question is addressed to Dr. Howard and Dr. Levy, who is going to be
the next speaker and who is going to describe something not unlike
what we have just heard. "Have you been able to make any observations
on bubble behavior in the rotating system? How might bubble dynamics
in the rotating system differ from bubble behavior in the stationary
system?"
DR. HOWARD: Well, I think I won't answer for Professor Levy.
He will also be able to answer for himself, but I will say that our
experience at Aston is that we certainly observed the bubbling in the
bed when it was taking placewhen the bed was rotating. There were
earlier papers that implied that bubbles would not form; but in fact,
there is a bubbling motion that takes place, certainly, and I
can show photographs of this to prove it.
The other aspect of it, namely, how might bubble dynamics in a
rotating system differ from bubble behavior in stationary systemsor
how might the process be modelledI don't really know the answer to
that one. I would say that bubbles certainly form (and with the beds
that I was playing with, whose radial thickness was quite thin, 10
millimeters or so) the bubbles were fairly small and well-distributed,
and what you might reasonably expect out of a shallow fluidized bed
(a stationary fluidized bed, of course). That is work that remains to
be done.
SPEAKER (not identified): We have made pictures of the surface
of our bed, and we can see bubbling occurring. We don't have any
quantitative information on the rate at which they grow, or the
relationship between fluidizing velocity and bed depth and bubble
size.
SHELTON EHRLICH, CHAIRMAN: Thank you. We are going to have a
coffee break now, if there are no further questions.
286
-------
INTRODUCTION
SHELTON EHRLICH, CHAIRMAN: Ed Levy? Come on up. If you are
here to listen to the second portion of the second session on tech-
nology test installations, you're in the right place. All right.
Can we all get ourselves seated now?
The next speaker, who will talk on "Centrifugal Fluidized Bed
Combustion," is Ed Levy of Lehigh University. He has a degree in
Mechanical Engineering and I think it is a delight to see a mechanical
engineer amongst all these chemical engineers once in a while. He's
been at Lehigh for the last 10 years. He works in heat transfer,
fluid mechanics, and energy conversion. Come on up, Ed.
287
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Centrifugal Ftuidized Bed Combustion
Er'-vird K. Levy, Norman W. Martin and
JohnC. Chen
Lehigh University
ABSTRACT
Centrifugal fluidizcd bed combustion is a relatively new concept for coal com-
bustion, where the bed rotates about its vortical axis of symmetry and the fluidizing
air flows radially inward through the porous cylindrical surface of the distributor.
l.chigh University has been involved in analysis and experiments with centrifugal beds
since l'.'~4 and is presently performing room temperature fluidization experiments. Oata
arc presented on the effects of particle size, bed mass, angular velocity, distributor
taper angle and distributor pressure drop on minimum fluidization and bed pressure
drop. Analyses which predict bed shape, niiniraum jfluidi z;it ion and bed pressure drop are
described and compared with the data.
I \TROIitlCTION
Centrifugal Huidiznt ion is a relatively new ftiiiilizeil bed concept, where the
bed rotates abou. it's axis of symmetry and the fluidizing air flows radially inward
through the porous cylindrical surface of the distributor (l-igure 1). The inward drag
force of the fltiidizing air on the bed material is balanced by the large radial accel-
erations caused by the rotational notion, permitting much larger air flow rates per
unit volume than arc possible with a conventional fluidizcd bed cr crating vertically
against gravity.
ROTATIONAL POWER
ID"
HOT GASES
,-PLENUM
/ VESSEL
AIR
DISTRIBUTOR
S
10"
V SOL I OS REMOVAL
BEARINGS AND SEALS
50 100 ISO 200
ANGULAR VELOCITY, ct/0 (f/mcn)
250
Figure 1. Conceptual Drawing of Centrifugal Flurdirad
Bad Comfauttor
Figure 2. Electrical Capacity of a Combined Cycle
Power Plant
288
-------
The rotational motion of the system gives it a number of distinct operating ad-
vantages, which make this device interesting for coal combustion applications. The
high fluidi:ing velocities provide for a compact system with relatively easy startup
and fewer problems with solids feed and bed mixing. In addition, !>;. varying the ซpeed
of rotation of the bed, the bed temperature, and the fluidi:ing velocity, the thermal
power output of the device can be varied over an extremely wide raime. Some of these
advantages arc illustrated in Figure 2, which shows the theoretical electrical output
of a large utility combined cycle power plant utiliii*;! a centrifugal fluidized bed
coal combustor.1 The combustor operates at 10 atrosphcrcs and 8"(lฐC and has a distri-
butor which is 3.04 m in diameter and 2.28 a high. With a nornal operating point on
curve 5 and an angular velocity of ISO rpm, the system would produce approximately
500 Mh'c. By decreasing the speed of rotation to 75 rpn and the velocity of the fluid-
i:ing air to minimum fluidi:ation conditions, the power output would be reduced to
50 MlVc.
Past re earch on centrifugal f luidizat ion includes experiments on minimum fluiJ-
1^.2.' ion and bed pressure drop performed at I'.rookhavcn National Laboratory and the Air
Force Acrosnacc Research Laboratory,-"'" cxpcr'.nicnt s by Ccl'pcrin in Russia on bed be-
havior,''"6 anu experiments by Farkas and his colleagues of the II. S. Pepartncnt of
Agriculture for food drying applications.7 At present, research is underway at
Brookhavc!> National Laboratory on the use of centrifugal fluidiied beds for flue gas
dcsul fur izat ion,h by Metcalfe and Howard in Crcat liritain on fluidination phenomena
and gas combustion,' and by Swithcnbank ct al. in (Irejt Britain on coal combustion.10
I.ehigh University has been involved actively in analysis and experiments witli
centrifugal iluidized beds since 1!>74. This includes parametric analyses of centrifu-
gal fluidizcd bed combustors,''' studies of fluniiicd bed cycles for combustion appli-
cations,12 and theoretical and experimental studies of minimum fluidi:ation. particle
clutriation, bed pressure drop, and bed startup. l 3 ' "*
This paper describes recent experiments performed at room tcmpeiature and at-
mospheric pressure on f1uidi:ation. A theoretical analysis which predicts the onset
of minimum fluidi:ation and bed pressure drop is described and compared witli experi-
mental data.
RF.VIF.IV OF Till; TIIHORY
The experimental results described in this paper were obtained with tancred
distributors, having the larger diameter at the top. From other experiments,^1 it was
found that the presence of a slight distributor taper angle is very useful in achiev-
ing distribution of bed material over the entire distributor surface.
For given bed dimensions and particle and gas densities, the important varia-
bles effecting minir.um fluid i ;at ion and bed pressure drop arc distributor angle T,
total weight of bed material, distributor pressure drop, angular velocity of the bed,
and particle diameter. Because of the influence of gravity and the coniril shape of
the distributor, the bed thickness varies in the vertical direction. An analytical
model, described in more detail in Reference 15, was developed which accounts for
vertical variations in bed thickness and air velocity by treating the system as a
scries of "n" elements, each of height l\z (sec Figure 31. The total pressure drop
across the grid and bed for each clement is given as
tPT " Pr>RII) * APBF.n (1 *
If the bed is packed locally, the bed pressure drop is equal to
*PBED
1.75(1-^1 ,.. , , * |_i_ . _i_| (2)
If the bed is fluidizcd locally, the voidage c is independent of radius, and the bed
material is in rigid body motion, the expression for bed pressure drop is
289
-------
(3)
Tor thin beds, the radial velocity at minimum fluidizat ion, calculated at the outer
raJius of the bed, reduces to
Ca = [150(l-c)/c
1 1 . 7S/ Jc 3 IRe2
(4)
The calculation procedure assumes a value for APr and uses equation (1J to compute
for each of the n elements. The total mass flow rate through the distributor
m(APT)
(5)
EXPLRIMUNTAL APPARATUS
The test section consists of a 0.3 m diameter, 0.15 in high distributor contained
in a square air plenum. The distributor is fabricated froin reinforced fine mesh screen
and the top end wall is transparent for visualisation of flow. A variable speed motor
is used to rotate the distributor by means of a shaft connected to the bottom end wall
of the test section. Room temperature air is supplied to the plenum fiom a compressor.
In all experiments the apparatus was b;itch operated, where a Riven charge of
atcrial was loaded into the test section and the distributor was accelerated tr<>m
bed material
rest to the desired angular velocity. The experiments were performed with glass
having narrow particle size distributions with mean diameters of 121, 351 and 475
beads
urn.
TIIUORF.TICAL AND LXPURIMF.NTAL RI-SULTS
Theoretical results were obtained on the effects of distributor shape and oper-
ating conditions on bed thickness. Typical results of the analysis arc given in l:ig-
ures 4, 5, and 6 as a function of distributor angle anJ angular velocity. The graphs
indicate the large non'jni formi t ies in bed thickness which can occur, particularly with
small bed masses and low angular velocities. An extreme situation is presented in
Figure 5 for a 2ฐ taper angle and 1.5 kg of bed material, where the bed docs not cover
the top of the distributor at low angular velocities. The opposite trend is shown in
Figure 6 with a bฐ taper angle, where the bed thickness increases in the vertical di-
rection. Measurcir.-nts of hcd thickness made at the top transparent end wail of the
test section are in good agreement with theoretical results. l-'igurc 7 shows bed thick-
ness as a function of air flow rate for.a 3.47ฐ grid with 1.5 and 4.5 kg of bed ma-
terial. Once the bed is fluiJi:cd (at m^f) bubbles erupt from vi.^ bed surface, cuus-
inc an apparent increase in bed thickness.
Ihe vertical variations in bed thickness lead to a nonuniformity of fluiJiia-
tion at transition from a packed to a fluidi.'.od slate. Th'-, is illustrated in I-i cure
8, showing the variation of bed pressure drop with lir flow rate, and l:igurc :>, showing
r, (ป)-
Fijure 3. Mo0el of Bed (or Analysis of Fluidiutien
290
-------
O.IS
u 0.10
u
z
in
6* 10' 2' 4ซ 6' 8* 10* 2ซ 4' 6ซ
5*8
4.5kg
0 0.02 0.04 0.06
BED THICKNESS (m)
Figure 4. Predicted Bed Thickness for Sev ;al Bed Masses at 26.2 rad/s
0.15
0.10
in
a
0.05
x
4
-4.5kg
7.5 kg
0 0.01 0.02 0.03 0.04 0.05 0.06
BED THICKNESS (m)
Figure 5. Predicted Bed Thickness for Several Bed Masses with 2ฐ Grid Angle
0.15
7.5kg
0.01
0.02 0.03 0.04
BED THICKNESS (m)
0.05
0.06
Figure 6. Predicted Bed Thickness for Several Bed Masses with & Grid Angle
291
-------
G
en
Ul
z
U
I
o
tu
00
0.04
O.02
O
O
C
(rt
UJ
z
1C
ซJ
X
1-
o
UJ
Ok
2
.
1
r1
'j
I.5.,O SURFACE
dp ป 4 75 > 10 n* " ^-/
<-. 36.7 rod/, 4.5 k,Q SURFACE
* J.^ f
Deป0.l02m 4.5hg\79U8BLE PEAK
GRID * 1
V
V
El
61
i TT IS g-1-
r ^x^.0..
T , ! . "^
3.0
O.I 02 0.3 0.4
AIR FLOW RATE (ms/$)
Figure 7. Comparison of Theoretical and (Experimental
Results on Bed Thickness
3
S
S 1.0
o
O.3
ua> 36.7 rod/*
ซp -SSOllO"*!!!
a 3.5ป
1.9 kg BED MASS
0 O.09 O.IO O.IS O.ZO
VERTICAL DISTANCE. Z (m)
Figure 9. Axial Variation of Radial Velocity
2,500
Ctป0ซ 36.7 rod/ป
dp ซ350ปIO'6m
a ซ3.5*
I.S kg 8EO MASS
0.38
O.4I
O.45
4.000
- 3.000
o.
o
-------
Theoretical and experimental results showing the effects of angular velocity
and bed mass on fluidization are given in Figures 10 and 11. For these conditions, a
4.5 kg bed mass corresponds to an approximate bed thickness of 2.5 cm.
The analysis and experiments both show a strong effect of grid pressure drop on
fluidization. Several different grid resistances, with the measured grid pressure drop
characteristics shcvn in Figure '2, were used in the experiments. Figure 13 shows the
effect of grid resistance on bed pressure drop as predicted by the theoretical analy-
sis. Generally, the influence of grid resistance is felt only during transition from
a packed to a completely fluidized state. The smaller the grid pressure drop, the
longer the re?ulting transition. In the case of a very low grid pressure drop, transi-
tion from minimum fluidization to complete fluidization may require such a large in-
crease in air flow rate that particle clutriation occurs before complete fluidization.
Figure 14 shows the theoretical axial variation of radial velocity at minimum fluidi-
zation for different grid resistances. A small flow resistance leads to extremely
large axial variations in velocity and as the grid flow resistance approaches zero,
the velocity nonunifornity can be large enough to permit clutriation to occur immedi-
ately after the first ring of bed material is fluidized.
These effects were observed in the experiments shown in Figure IS. Using a
1.5 kg bed mass, the bed pressure drop data and theory for grids 1 and 2 both exhibit
a well-defined transition between the packed and the fully fluidized states. Once
fluidized, the bed pressure drop is relatively insensitive to air flow rate. The low
grid pressure drop case, grid 0, gives radically different bed pressure drop behavior.
Transition to a fluidized state begins at a lower air flow rate and continues to very
high values of flow rate. This very long transition region is marked by a bed pres-
sure drop which increases continuously with flow rate after the initiation of fluidi-
zation. bimilar results are shown in Figure 16 for a thicker bed where particle elu-
triation occurs before minimum fluidization when grid 0 is used.
Through analysis of these results and others, it is concluded by the authors
that the ratio of grid to bed pressure drop at minimum fluidization should be greater
than 0.2 if the bed is to fluidizc with a narrow well-defined transition region.
7.0 OO
6.000
~. 5.000
a
o
e
o
4.000
3.000
O
kJ
O
2.000
I.OOC
Oe- 0.1524m
3P ซ475ซIO~6m
wซ36.7rod/s
GRID- 2
o .3.47*
7.000
6.OOO
5.000
oT 4.0OO
o
cc
o
U
oc
2 3.000
1.5kg
CซO.4I
1,000
BED MASS 0
c~>' 36.7rod/ป
GRID 3
ซ 5.19*
GRID 0
aซ 3.47ป
0 O.I 0.2 0.3 0.4
AIR FLOW RATE (mS/t)
Figure 11. Effect of Bed Mats on Bed Pressure Drop
0 O.I 0.2 0.3 0.4
AIR FLOW RATE Im5/f)
Figure 12. Measured Grid Pressure Drop Characteristics
293
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s.ooo
ฃ 4.000
u 3.000
2,000
I.OOO
BED MASS-3.Okg
De O.IS24m
aซ4ซ
C -0.42
dpซ35lซIO"6m
GRID 0
-GRID I
O.I 0.2 O.3 0.4-
AIR FLOW RATE (ms/ป)
4.000
3.000
u 2.000
1.000
- MEASURED APC
THEORETICAL APe
BED MASS' 1.5kg
3p-475ซIO"6m
W36.7:od/ป
ป 3.47*
DATA | CRIP
0
I
2
AP6. GRID 2
GRID 0
AIR FLOW RATE (m3/ป)
Figure 15. Effect of Grid Resistance on Transition;
Thin Bed
Figure 13. Theoretical Effect of Grid Resistance on
Bed Pressure Drop
10
&J-20.9 rod/*
dpซ35lปIO"6m
a .4*
6 -0.42
BED MASS* 3.0 kg
GRID 0
GRID 2
UMF
0 0.05 0.10 0.15
VERTICAL DISTANCE. Z (m)
Figure 14. Theoretical Axial Variation of Radial
Velocity; Low Angular Velocity
7.000
6.000
AIR FLOW RATE
Figure ia Effect of Grid Resistance on Transition
Thick Bed
294
-------
6.000
5.000
- 4.000
a.
o
cr
o
3,000
2.000
1,000
A
0
a
3P |MF
I2IXIO"6
351 ป IO'6
475xlO~6
TMEO
0.027
O.IO8
0.230
BCD MASS > 3.0 kg
36.7rod /ซ
GRID I
aซ 3.47*
4.000
3,000
o
DC
O
ir 2.OOO
I.OOO
O.I 0.2 0.3 O.4
AIR FLOW RATE tm3/j)
Figur. 17. Effect on Particle Size
AIR FLOW RATE ds and
large particles. Low grid pressure drop and particle noisturo arc believed to nc re-
sponsible for the anomalous fluid i :at ion behavior observed with the snail particles.
The flow resistance of the grid is an important factor in determining the qual-
ity of fluidii.it ion. As with conventional fluidizcd beds operating vertically against
gravity, sufficient Rii
-------
RUFP.RP.XCnS
1. E. K. Levy and .J. C. Chen, Proceedings 01' the International Powder and Bulk
Solids Hand 1 ins a"d Processing Conference, Rosemont, Illinois, 1977, p. 452.
2. Brookhaven National Laboratory, "Rotating Pluidi:ed Bed Reactor for Space
Nuclear Propulsion," Upton, New York, Reports BNL 50321, August 1971; and
BNL 50362, September 1972.
3. L. Anderson, ct al., .>. Spacecraft, Vol. 9, Xo. 5, May 1972: p. 311.
4. .J. Howard, Air Force Aerospace Research Laboratories, "An l:xper incntal Investi-
gation of Heat Transfer in a Rotating Fluidiicd Bed," ARL 75-0115, IVright
Patterson Air Porce Base, Ohio, May 1975.
S. N. Get'pcrin, V. C. Ainshtein and 1. U. Ooikhman, Khimichcskoc i Xoftyanoc
Mashinostro-;nic (translated from Russian), No. 5, 1964, p. 18.
6. X. Gel'pcrin, P. D. l.cbedcv, V. C. Ainshtein and C".. N. Napalkov, Kh imichcskoc
i Ncftyanoc Mashinost rocnic (translated froir Russian), Xo. 5, 1966, p. 5.
7. I). l:. Parkas, ct al., Pood~ Technology, Vol. 25, November 1969, p. 125.
8. P. Hill, Brookhavcn Xational Laboratory, personal communication, August 1977.
9. C. I. Metcalfc and .1. R. Howard, Applied Energy (3), 1977, p. 65.
10. X. Demircan, G. Gibbs, .1. Swithcnbank aniT~l). S." Taylor, Fl'ii dilation,
Cambridge University Press, 1978.
11. E. Levy, ct al., "Parametric Analysis of a Centrifugal Fluidized Bed Combustor,'
ASML Paper 76IIT68, 1976.
12. IV. Shakespeare, P.. K. Levy, .J. C. Chen and V. Kadambi, Proceedings of the 1977
Intersociety F.nergy Conversion Engineering Conference, Washington, D.C.,
August 1977, p. 751.
13. I;. Levy, N. Martin and .1. Chen, Fluidizat ion, Cambridge University Press, 1978.
14. K. Levy and J. Chen, Lchigh University, "Centrifugal Pluidizcd Combustion of
Coal," Quarterly Reports PP.-2S16-1, .January 1978; PE-2516-2, April 1978;
PP.-2S16-3, July 1978.
296
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QUESTIONS/RESPONSES/COMMENTS
SHELTON EHRLICH, CHAIRMAN: Do we have any questions?
WILLIAM REID, EPRI: What is 1000 pascals in terms of inches of
water?
MR. LEVY: 1000 pascals is approximately 4 inches of water.
SPEAKER (not identified): A rather obvious question is whether
you think that this rotating system could be built to work in an
industrial or some sort of practical application eventually. But
what I would really like to ask is whether you know of any precedent
of a moving system in an application where it has to work under
severe conditions for a long period of time and so forth. Is there
anything analogous to this that has been successful?
MR. LEVY: IV sorry. Would you repeat that question?
SPEAKER: Well, the elementary question is: Is it really
practical to think of such a system working inyou showed it as an
application for a combined cycle. Can you really imagine such a
system, with a rotating member, operating year after year that you
could maintain, and so on? But supplementing that, I wonder whether
there is any precedent for something like thiswhether you can think
of some other application where, you know, something of this sort did
operate successfully? Maybe that's too involved a question?
MR. LEVY: I don't know of any other precedent. The gas turbine
is a system that has very high temperatures and a relatively, perhaps
not as dirty an environment as this, and that operates at very high
speeds of rotation. I am not sure I can cite any precedents for
this system. We do believe that it has very good potential for the
kinds of applications that I described. Jack Howard has described
others, and there are other people working here in the States and in
Britain who are convinced that the thing has merit. I don't know
what else to tell you at this point. It is in a very early stage of
development, and we certainly can't point to a "Rivesville" yet. You
would have to give us time to ao that.
SHELTON EHRLICH, CHAIRMAN: Any other questions? Okay. Thank
you, Ed. Bill Reid's question, "How many inches is 1,000 pascals?"
reminds me of a thing that happened to us at EPRI two weeks ago.
We were sent a stack of reports from the Environmental Protection
Agency about that (indicating) high for us to review and comment on
297
-------
today, regarding the justification for revised new source performance
standards. I was given one of these reports to review, and it wos
all in the SI units. And if anyone can tell me what a heat rate is
in 2340 kilojoules per kilowatt/second, when you can come to work for
us. (Laughter).
298
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INTRODUCTION
JOHN WILSON, CHAIRMAN: Our third speaker, is going to discuss
the chemically active fluid bed. Dr. Gerald Moss is with us this
evening. He is an Engineering Associate at the Esso Research Center
in Abingdon, Oxfordshire, United Kingdom. He is the inventor of the
chemically active fluidized bed system for regenerative desulfuriza-
tion of fuels during either full or partial combustion in a fluid-
ized bed of line. He is also the author of some 25 issued patents.
Dr. Moss is currently developing a range of new projects in the
field of fuel processing and energy conservation, some of which were
derived from the chemically active fluidized bed principle. This
evening he is going to discuss with us the Progress in the Development
of the Desulfurizing Gasifier. Dr. Moss?
299
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Progress in the Development of the
Desulphurising Gasifier
Gerald Moss
Esso Petroleum Company Limited
ABSTRACT
The development of the E.P.A. sponsored Chemically Active Fluidised Bed de-
sulphurising gasifier has now reached the demonstration phase and a 20 MWe unit is
under construction at the La Palma plant of Central Power and Light at San Benito in
Texas.
Although the gasifier was originally designed to burn residual fuel oil, it has
been established that solid fuels can also be gasified and desulphurised. The paper
deals with the operation of the Abingdon pilot plant on coal and also describes what
appears to be a feasible and attractive technique for the reduction of SO2 to ele-
mental sulphur within the regenerator of the gasifier.
INTRODUCTION
The C.A.F.B. principle for regenerative desulphurisation during combustion was
disclosed at the First International Conference on Fluidised Bed Combustion and is
described in more detail in a paper given to the Institute of Fuel Fluidised Combustion
Conference in 1975 (1). It is somotimes described at the one-stt-p regeneration process
and is the subject of patents held by Exxon Research and Engineering Company.
The pilot plant which was built with the support of E.P.A. at the Esso Research
Centre at Abingdon has also been described previously (2), (3). The gasifier is a two
reactor unit, both reactors containing beds of lime fluidiscd with air. Bed material
is exchanged in a controlled manner between the two reactors which are of the same-
height but differ considerably in tiiameter. The large reactor is the gasifier. It
operates at about 900ฐC and fuel introduced below the surface of the fluidised bed is
partially combusted. Normally about 80*. of the sulphur in the fuel is retained by the
bed material as calcium sulphide. The bed material also retains metals which tho fuel
may contain. The regenerator is operated at about 1050 C being held at a higher temp-
erature than the gasifier by the exothermic oxidation of CaS to lime and SO0. The
regulated flow of ted Material between the reactors controls the regenerator^ temper-
ature and at 1050ฐC sulphur dioxide is formed in the regenerator at a concentration in
the region of 7-8" by volume. Diagrammatic cross sectional . jrvป of the first unit
are shown in Figure 1.
THE C.A.F.B. GV5IFICATION OF COAL
Under the comparatively mild opex-ating conditions which favour sulphur retention
a substantial proportion of the fuel is merely cracked to produce appreciable concen-
trations of condensible hydrocarbons in the fuel gas. When the fuel is oil, the carbon
rejected by cracking is laid down on the surfaces of the lime particles and is subse-
quently gasified. When coal is substituted for oil the following factors modify the
operating conditions:
Less hydrogen is available in the fuel for the formation of tars
The char tends to remain as discrete particles instead of being
spread over the surfaces of the lime particles
Small char particles arc elutriated from the bed and pass through
the cyclones to the burner
Appreciable amounts of water and ash enter the system
300
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Outer metal casing-
Insulating refractory
Castable refractory\
Cyclone for
gasifier
/
u>
o
Connection between cyclones
Expansion bellows to
absorb vertical expansion
on gas outlets
Removable lid
Gas pulse
Regenerator
Cyclone fines
fed into bed
transfer pipes
Cos pulse
Bed return
from
regenerator
ED.
Bed
drain
Regenerator
drain
Air Supply
Air Supply
Figure 1. Layout of Continuous Gasilier Unit
-------
The fine char v.-hich is elutriated docs not have to be gasified but is burned directly,
whilst the water in the coal acts as an oxidant. Consequently, despite the less
favourable hydrogen to carbon ratio of coal, the proportion of stoichiometric air
which is required to operate the unit on coal is not very much greater than when oil
is used. The presence of ash and ปater in the fuel also tend to hold down the oper-
ating temperature.
The initial coal gasification tests were made using a small batch reactor and
it was I'ound ihat four quite di ffert-nt coals could be "gasified" at temperatures in the
region ol 900 c and fluidising velocities in the region of 1.4 metres/sec., in a bed 0.4
wires deep. The decree of desulphurisation appeared to somewhat exc'.-od the degree
to which the coal had he-en truly gasi f i'.-d, the best figure, bast-d on the sulphur
content of the bed material, being ihout 641. A large proportion of the coal ash *as
elutriated from the bed and cyclic regenerations, made by simply turning off the fuel
supply showed that continuous operation uas feasible.
Before continuous operation could be demonstrated it was first necessary to
solve two practical problems. We had to provide the pilot plant with suitable coal
handling faci 1 i tic:;, and ปe also had to modify the cyclone drainage system to c'.able
it to cope with the larger flow of solids resulting from the presence of the coal ash.
The coal handling system ปas specific to the pilot sc.ilc of the operation and of
limited general interest, t>ut the cyclone drainage system could have wider appli-
cations in situations where head roon is limited.
The system was devised to meet the following criteria:
No valves handling hot solids
A capability for removing largo chunks of carbon from the
system
Provision for fines re-inject ion near the bottom of the
gasi fier bed
Provision for discarding very fine material for fines
i-.iventory control
The rejection of systems incorporating hot solids handling valves was based on
previous exfxrienre with a variety of such systems, one of the problems being the
presence of largo chunks of carbon following cyclone burn-outs. It was considered
desirable to re-inject the fines near the bottom of the gasifier bed in order to
ensure a reasonable retention time, ปhi1st discarding very fine material would reduce
the amount of material escaping through the gasifier cyclones into the burner.
A line diagram of the fines handling system which was devised to meet these
criteria is shown in Figure 2. The pilot gasifier is fitted with two cyclones in
parallel and as shown in the top left hand corner of Figure 2 the solids drains from
these two cyclones were joined to a common diplcg which drained into a small hopper
containing a fluidisable bed. The term fluidisabie is used, because the flow of
fluidising gas to this bed is intermittent, being controlled by the level of solids
ithin the dipleg. When the hopper is fluidiscd, bed material can flow out of the
dipleg into the hopper, displacing material which drains over the weir, but when the
hopper bed is slumped no such flow can occur. The hopper dipleg svstem Is, therefore,
a gas flow restrictor which enables the or ssure within the hopper to be maintained at
a higher level than that within the cyclor.es. Any lumps of carbon which emerge from
the dipleg sink to the bottom of the hopper when the bed Is flnidised and can be
removed through the large drain valve. Uaterial flowing out of the hopper and over
the weir is entrained by an eductor and is delivered at a still higher pressure to a
low efficiency cyclone, of very small dimensions. The solids draining from this
cyclone enter a second eductor and are re-injected into the gasifier. The gas
leaving the low efficiency cyclone passes through a higher efficiency cyclone, then
through a cooler and finally through a blower which recycles it through the eductor
nozzle, cyclone circuit.
302
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to
o
to
"Level Control
Hopper with
fluidised bed-)-'
Nitrogen
Eductor
Low
Ef'iciency
Cyclone
High
Effici
Cycle;;.
Hopper
-Air
Gasifier
Figure 2. Flow Diagram for Fines Return System
-------
Thfs system gave good results during our first runs using coal and is still in
use. Though those runs were of shc-rl duration, they showed that as much as 897 of the
lignite could be combusted without taking stops to rccove* the carbon discarded with
the cyclone fines,and that desulphurisation efficiencies exceeding SOT. are obtainable.
Between coal tests the gasifier was operated on fuel oil, since it v.as found to be a
simple natter to switch from one fuel to the other. This provided some insight into
the ash retaining capacity of the gasifier bed, because it wad found that coal ash
particles continued to drain from the cyclones during the periods when oil was gasified,
indicating that the ash held in the gasifier bed was not on the whole permanently
hound to the bed particles. During one such period when no limestone was added to the
bed.it was found that over 007 of the loss in lime inventory was accounted for by
material removed from the- cyclone drainage system. L'nder comparable conditions prior
to the installation of this system, stone losses from the gasifier into the boiler
amounted to about 2.7 Kg/hour, most of which was retained either within the boiler or
by a flue gas cyclone- fitted between the boiler and the stack. L'sing the new system
the loss amounted to 1.K Kg/hour and only a tenth of this entered the boiler. The
lower loss 1.8 Kg/hour versus 2.7 Kg/hour presumably reflected the improved operation
of the gasifier cyclones with a reduced fines burden due to the extraction of the
finer material Irom the fines recycle stream. Unfortunately, it turned out that a
large proportion of the lignite ash was too fine to 'ซ retained not or.ly by the
gasilier cyclones but also by the flue gas cycloro system, and it would appear that
be'.ueon 257 and 557. of the fly ash went up the stack. Further details of these test
results art- given in Reference 4.
A prob!ซ-m common to all C.A.F.H. i-i-geni-rativo systems is the disposal of the
concentrated SO,, stream from tin; regenerator. Flemontal sulphur may be obtained by
reacting SL>2 ~ with carbon to produce sulphur and vas i.oped that the SO0 formed
in the lime bed would be reduced to elemental sulphur in the char layer floating on
top, thus eliminating the need for the Kesox plant and reducing the cost of the process.
So far only one test has been made using a batch gasifier but the results have
been encouraging and the technique appears to be feasible. Obtaining a mass balance
during a batch regeneration is extremely difficult, sulphur in particular is difficult
to measure accurately. An aspirator was used to sample the tail gas at a steady rate
throughout the regeneration and whilst this enabled the presence of sulphur to be
verified, it did not allow instantaneous conversion efficiencies to be measured.
A normal batch regeneration was run at a fluidising velocity in the region of
O.C met res/second and when the SOo content of the tail gas had risen to about 67 the
bed was slumped and the char was introduced. The effect of this char on the subse-
quent composition of tht tail gas following refluidisation, is shown in Figure 3,
though it should be pointed out that the lines Joining the points prior and subsequent
to the introduction of the char are somewhat questionable and it is likely that sharp
discontinuities occurred. The suppression of SO2 and the production of C02 do not of
themselves necessarily indicate the production of elemental sulphur, since carbon is
oxidised much more rapidly than CaS and unduซj mixing of the char with the bed would
simply put the regenerator out of action. However, although the char layer was
initially only about 4 cms. deep, the indications were that between 30 and 60% of the
missing SOo was converted to elemental sulphur during the period when the SOo con-
centration*~was less than 6*. Whilst encouraging, this result is not by any means
304
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GAS COMPOSITION DURING FLOATING CHAR
REGENERATION
2 4 6 8 10 12 14 16 18 20 22
TIME MINS FROM INTRODUCTION OF CHAR
Figure 3. Tail Gas Composition During Floating Char Regeneration
-------
conclusive, but it is perhaps worth pointing out that in completely combusting systems
the degree of segregation required will be less stringent, since, in this case, the
regenerator must in any event operate under overall reducing conditions. The technique
is, therefore, recommended to those proposing to operate desulphurising fluidised bed
boilers in the regenerative mode.
CONCLUSIONS
It has been established that coal can be treated by the C.A.F.B. process to
produce a hot desulphurised fuel gas.
A reliable cyclone drainage system has been demonstrated which allows the
height of the disengagement sections of fluidised bed reactors to be drastically
reduced.
What appears to be a promising technique for the direct production of elemental
sulphur within C.A.F.B. regenerators has been proven to be feasible.
ACKNOWLEDGEMENTS
The author wishes to thank the Knvironmental Protection Agency for the financial
support which enabled this work to be done, and to acknowledge the valuable contri-
butions made by all the members of the C.A.F.B. project at Abingdon. In particular
he would like to thank those who worked under the leadership of Dr. D. Lyon on the
conversion of the continuous gasifier to coal firing.
KEFEREXCES
1. C. Uoss, "The Mechanisms of Sulphur Absorption in Fluidised Beds of Lime"
Institute of Fuel Symposium Series No. 1: Fluidised Combustion, London
September 1975.
2. G. Uoss, "The Fluidised Bed Desulphurising Gasifier" Proceedings of the
Second international Conference on Fluidised Bed Combustion, Houston Woods,
Ohio. 1970.
3. G. Moss and D.E. Tisdall "The Design, Construction and Operation of the
Abingdon Fluidised Bed Gasifier" Proceedings of the Third International
Conference on Fluidised Bed Combustion, E.P.A.- C50/2-73-053, 1973.
4. D. Lyon, "First Trials of C.A.F.B. Pilot Plant on Coal", E.P.A.- GOO/7-77-027,
1977.
5. Kirk-Othmer Encyclopedia of Chemical Technology. 2nd Edition, Vol.19 p.410.
306
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QUESTIONS/RESPONSES/COMMENTS
JOHN WILSON, CHAIRMAN: Dr. Moss, I believe, has a few questions
as wel1.
DR. MOSS: Mr. Wormser, of Wormser Engineering, wants to know:
"What is the spacing of fuel injectors when you operate on oil?" The
point is" that we've only got quite a small unit, and it only has
one injector. The largest version we built had a bed area of 5
square feet, and that is the sort of spacing which is being used by
Foster Wheeler in Texas. "Would it be the same if we operated the
CFB as a burner?" Well, we do operate it as a burner, of course, on
startup and sometimes on hot standby; but we don't normally operate
in the fully combusting mode and try to trap sulfur in that way.
Probably the National Coal Board people, who have done this sort of
thing in conjunction with BP, are in a better position to answer that
particular question.
"How high above the distributor do we inject the oil?" Well,
basically you get it as near to the distributor as you can, without
squirting it into the nozzles; 2 or 3 inches above the air nozzles.
The distributor itself isn't flat. I think I pointed out at the last
conference that in order to get the fuel distributed over a large
area, as in the case of the Texas plant, we have used a number of
pits in the distributor, so that the oil ducts can be covered with
refractory and not exposed to the environment of the fluidized
bed until they reach the point at which you want the oil introduced.
I've got a whole range of questions here from Joe Yerushalmi.
First, "What is the gas velocity in the gasifier?" We normally work
in a range of 4 to 6 feet a second, going back to our original units,
and it is difficult to say what the maximum velocity would be. You
can only work a fluidized bed, as probably most of you realize, over
a fairly narrow range of velocities because otherwise you run into
pressure drop problems with your distributor, which is designed for
a specific velocity. Now if you blow a bit harder, you're probably
all right until you reach the situation where you spout. What we
have found is that the reaction between H2$ and the stone is
extremely rapid. The material which comes out--'he sulfur which
leaves the bedis present in the form of organic compounds, and you
have to break down these compounds in order to desulfurize them. If
the bed spouts, you can get a situation where you squirt the stuff
through. And that, I think, is probably the limitation.
"What is the mean particle size of the lime?" Well, we operate
with about 1200 microns, but just for convenience. Tne external surface
of lime isn't the controlling factor. Size isn't ail that important,
307
-------
we've found. It's the internal surface which seems to do the work.
Increasing the amount of fines in our bed material has not, of itself,
improved its reactivity appreciably.
"What is the quality of the fuel gas?" The answer to that is
"highly combustible." It depends, of course, on the stoichiometric
ratio which we are operating at. We normally "turn down" by varying
our stoichiometric ratio which may be anything between, maybe 30
percent stoichiometric to maybe 15 percent, with oil; so naturally,
you're going to get variations in the quality of the gas, different
amounts of tar in it, and so on. There is, of course, a variation in
temperature when you vary the stoichiometric ratio, which you can
compensate for by using a bit of flue gas recycle, but we're not
really strapped on the question of operating temperature. It doesn't
make all that much difference to the desulfurization efficiency over
a reasonable range. A range of say 870ฐC to about 970ฐC, I suppose,
gives you reasonable results.
"What is a typical carbon utilization?" was another question.
Well, the only carbon you would lose in the case of oil firing would
be that which came out of your regenerator as ฃ03. All of the fuel
is burned. It's just a two-stage combustor, basically. The efficiency,
if you like to put it that way, of fuel utilization in a large unit,
allowing for heat losses through the wall and for calcining stone and
so on, may be in the region of 98 percent on a thermal basis. But, of
course, there is a power requirement, wh--:h would knock the overall
efficiency down maybe to about 95 percent.
And there is another question here, "What are the proportions of
lime and char in the gasifier?" Well, this relates to the case where
we were using coal, and we have worked with carbon contents in the
range of 1 to 10 percent by weight. The problem is basically that if
you get too much carbon in your bed, then you'll put your regenerator
out of business, because the carbon reacts much more rapidly with
oxygen than does calcium sulfide. So, if you have too much carbon in
the material entering your regenerator, you cease to regenerate. Okay?
JOHN WILSON, CHAIRMAN: Okay. Are there any other questions?
308
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INTRODUCTION
DR. FREEDMAN, CHAIRMAN: We v^ll return to the technical proqram
now. The next paper, which is not printed in your progan, will be by
Dr. Steve Wright, who is the Deputy Program Manager of IEA Services,
which is building the pressurized flcidized bed plane * urimethorpe,
England, and with that, I will turn it over to .eve.
MR. BROADBENT: Hello. Actually, Steve is going to talk to you
in a few minutes. I just wanted to say a few words first.
I spoke to you generally on Monday about the NCB, the National
Coal Board, IEA Services Limited; about its constitution and aims, and
we welcome this opportunity of talking to you now and extending the
discussion of the more technical aspects of the plant at Grimethorpe.
Since most of the design work is not complete, it is possible to
technically describe it in some detail, and we are going to concen-
trate on the combustor and the exhaust gas cleanup system. Time
permitting, we will go on to other things, but I think those two are
the most important parts.
Before handing you over to Steve Wright, who is the project
manager of the Grimethorpe job, I would like to emphasize that this
project is designed to cover a very large span of pressures. It's
designed from 6 to 12 atmospheres; therefore, it is a rather large
span, and a very experimental facility. Dr. Steve Wright.
309
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A Technical Description of the Plant
Design and Project Progress Report
David H. Broadbent
and
S.J.Wright
National Coal Board (IEA Services) Ltd.
HISTORY OF THE GRIMETHORPE PROJECT
The history leading up to the agreement to build a fluidised-bed facility at
Griroethorpe. near Barnsley, Yorkshire in England is lengthy and has been given in the
paper tr the Fluidised-Bed Combustion Technology Exchange- Workshop here in Washington
earlier this year. In summary the scheme for the present programme of building a pres-
surised tluidised-bed research facility stems from the setting up in 1974 ot the Inter-
national Energ> Agency, for which the NCB was invited to establish the Coal Working
Group under che chairmanship of Mr. Leslie Grainger. NCB's then Member for Science.
Five priority agreements emerged from the working group, one of which is the building
of the experimental Fluidised-Bed Combustion Plant. In order to formally manage the
five projects. NCB (IEA Services) Ltd.. a wholly owned subsidiary of the NCB was set
up. and from the signing in 1975 of the implementing agreement, a tripartite agreement
was reached between the United Kingdom, United States, and the Federal German Republic
to build the experimental facility, having commenced in November 1975 and due for com-
pletion in early 1979 w'th the subsequent experimentation for an additional four years.
OBJECTIVES AND SCOPE OF THC PROJECT
The agreement provides for four principal stages in the programme over the period
of seven years Curing which the experimental 3tage provides the most important objective.
The stages are as follows:
Stage I Procurement of the design study with accompanying documents.
Stage 2 Tendering for the construction of plant and appraisal of tenders.
Stage 3 Construction and acceptance of plant.
Stage 4 Operation of the plant.
The major part of the work is now in stage 3 of the programme, whilst soms minor
sections are still in stage 3. Details of these will be described later.
Broadly the objectives of the programme are threefold:
Objective 1 To build the experimental facility to study combustion, heat transfer.
gas clean up, corrosion and energy recovery in a pressurised fluidised-
bed combustion system.
Objective 2 To carry out tests over a wide range of operating conditions and to
make measurements in greater detail than would be either economic
or possible in an intergrated unit.
Objective 3 To provide data for the analytical modelling of design data for
larger commercial plant by empirical data extrapolation.
The experimental facility provides the means to conduct research into a number
of aspects of large commercial plant, and in the immediate future, the following studies
are sought:
310
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(i) Combustion performance using high ash coal.
(ii) Effects of dolomite additives on sulphur retention.
(iii) Study of heat transfer co-effecients, heat transfer element
materials, and geometry.
(iv) Gas clean up and dust removal prior to atmospheric discharge.
(v) Flue Gas corrosion and particulate erosion effects upon gas turbine
blade material, leading to the optimisation of a flue gas driven
gas turbine design.
THE GRIMETHORPE EXPERIMENTAL FACILITY
The facility evolved from the design study will be built adjacent to the
National Coal Board's Grimethorpe power station near Barnsley in Yorkshire. England.
The major components of the facility will be workshop fabricated and transported to
site for erection and interconnection ready for start up in late 1978 early 1979.
Whilst it is not a condition of the agreement, it is interesting to note that plant
and equipment is coming from each of the countries party to the agreement. In round
figures, of the 10m construction cost about 2m has gone to Germany, 1m has gone to
the U.S.A. and remainder to the U.K.
GENERAL DESCRIPTION OF TTE EXPERIMENTAL FACILITY
The facility (Figures 1 and 2) consists principally of the combustor which is
coal rue led trom a coal/dolomite crushing and blending plant. The combustor's own
steam generation powers a turbo compressor pressuring the combustion system. The steair
is condensed, the condensate being processed to the required water quality before return
to the steam water circuit.
Off-gas from the combustor is cleaned of entrained particulate matter by passage
through two cyclone stages, before cooling, pressure let down and discharge to at-
mosphere through a silencer and stack.
Ash disposal from the combustor and gas cleaning system is provided by suitable
batch hold and dumping facilities.
The whole facility is to be housed in a building some 30m high, 15m wide by 51m
long, attached and adjacent to the existing Grimethorpe Power Station. Construction of
the building is already underway and on schedule. Lifting of heavy equipment necessary
during construction and maintenance is facilitated by a 40 ton capacity remotely con-
trolled gantry crane.
TECHNICAL DESCRIPTION OF THE PLANT
Since rost of the design work is complete, it is now possible to technically
describe the plant in the following order: i) fuel preparation and injection, ii) the
combustor, iii) ash handling and disposal, iv) exhaust gas clean up system. Finally.
mention is made of auxiliary items such as the turbo-compressor.
FUEL PREPARATION EQUIPMENT
The contract for the fuel preparation equipment (Figure 3) was let to Head
Wrightson Process Engineering Ltd. early in August with the next step being the
finalization of detail plant design with NCB (IEA Services) Ltd. prior to the commence-
ment of fabrication. This contract is currently on schedule.
The Fuel Preparation equipment provides an installation to accept raw coal from
a 10,000 tonnes stockpile into a 30 tonne capacity surge bin supplied by an armoured
faced conveyor (AFC). Screening facilities for -25mm coal size, conveyor fed from
311
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CO
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ro
MofcSregUrt Ftmon
Vntd Cduitic/Adtf.
Figure 1. NCB (IEA Services) Ltd. Grimetiiorpe Experimental Facility Plant Layout-Elevation
-------
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Figure 2. NCB (IEA Services) Ltd. Grimethorpe Experimental Facility Gn/Solidi Flow Sheet
-------
P* CGNVC1OA
Figure 3. NCSIIEAServicn) Ltd.
Gcimethorpe Experimental Facility
Schematic of Fuel Preparation Plant
314
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the raw coal surge bin, are provided to permit +25mm only feed to a primary 30t/h
Hammer Mill Crusher. All -25ram coal is conveyed to a 16t/h -30t/h fluidised bed
drier acheiving an output of coal at -17. moisture c.f. normal air dried value.
Secondary crushing of the dried coal is provided by two lOt/h Hammer Mills
capable of comminution to -6.4mm or 1.6mm as required, the product being conveyed
to a 370 tonnes Coal Silo. The internals of these mills can be adjusted, as can the
speed of rotation allowing a wide variation in the size spectrum of the product coal.
A 640 tonnes capacity Dolomite Storage Silo will be supplied with pro-prepared
material delivered by road ".cnkcrs. Space has been left in the layout for a 20 tonnes
Additive Silo which will be similarly supplied with pre-prcpared material.
A blending system, integral with and supplied by the avove storage system.
functions to permit batch mixing of coal, dolomite, and additives to any proportion
required by the experimental facility. The blender is equipped with weighing hoppers
sensed by strain gauges, delivery control valves, and the necessary control system co-
ordinated by a micro-processor unit. The blend will be system delivered to the fuel
injection assembly.
FUEL/BED MATERIAL INJECTION SYSTEM
The contract for the fuel injection system was let to Petrocarb Inc., in the
early part of this year and is well into the detail design stage. Problem.0 have existed
with the juxtapositions of the Ash Screw Conveyor which is to be supplied by Head
Wrightson Process Engineering as part of the ash removal system. The Petrocarb fuel
feed lines must have a straight run to the corabustor before a single vertical rise into
the pressure shell to ensure proper pneumatic transport. However, these problems now
appear to be resolved.
THE COMBUSTOR
The contract for the combustor (Figure 4) was let to Vereinlgte Kesselwerke AC
(VKW) early this year, and the detail design work is nearly complete. A number of design
changes have come about motivated by realizations of cost saving, and principally affect
the construction of the pressure shell and resulting logistics of handling the combustor
intervals during shutdown. The testing procedures to be used for the pressure shell
construction will satisfy the statutory T.U.V. authority in Geroany. and have been
accepted by the U.K. Insurance Company.
The combustor has a datum capacity of 80MW operating under pressures in the
range 6-12 bar with the off-gas in the temperature range 600-950 C with the maximum
design temperature of 1100 C. The combustor structurally consists of four principle
elements, namely: i) the pressure shell, ii) furnace water wails, iii) distributor
plate, iv) heat transfer tube bundles. The combustor pressure shell is cylindrical
in shape, 4m in diameter by 14m in height to the gas outlet flange, and will weigh sone
100 tons erected.
PRESSURE SHELL
The pressure shell, designed to operate at a maximum pressure of IS bar. and
fabricated from 17Mn/4 shaped steel plates, butt welded together, has been sub-contracted
by VKW to Borsig in Berlin. The cylindrical shell is terminated by domed ends, of which
the upper is welded to the shell, and the lover is flange fitted permitting removal and
ready interior access for inspection, adjustment and removal cf the combustor internals
during shutdown. The upper pressure shell dome is integral and a refractory lined hood
removing flue gas from the combustor. The pressurised fluidising air intake is to
the annulus between the hood and the main pressure shell. At the lower end the hood
is connected to the upper end of the furnace water-wall. The lower pressure-shell done.
carries the pipework penetrations for the fluidised bed services such as fuel feed.
bed material and ash disposal. The combustor is supported in a vertical position on
anchorages designed to accomodate thermal expansion and vibration.
315
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COM. MLEfS
Figur* 4. NCB (IEA Servien) Ltd.
Grimethorpe Experimental Facility
Combustor/Preiiur* Vessel
316
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FURNACE
The furnace contained within the pressure shell is of square cross section 2m
x 2m. the walls of which are constructed as water '-alls consisting of spiral tubes
(38mm od x 4.5mm wall) fabricated in st 35.3 steel filleted together with 6mm steel
plate regularly perforated to accomodate thermocouples as demanded by the experimental
programme. The water walls, in addition to heat absorption duties, contain the flui-
dised bed and the resulting hot gas stream. The upper connection is to the insulated
hood and gas outlet as previously described.
PROBE FACILITIES
75mm diameter perforations are provided in the furnace u^ter walls in a specific
location pattern, permitting the installation of measurement probes within the in-bed
tube bank as required by the experimental programme. The water walls are deflected about
these probe perforations. To facilitate installation. iC.-noval, and servicing of probes,
tr.anholes are provided in the pressure shell service by both internal and external gal-
leries.
DISTRIBUTOR PLATE
The distributor plate provides the floor of the f'uidised bed and accomodates the
fuel injection nozzles, the fluidising air distribution system, ปsh removal, and bed
material supply. To date two distributor plates, designated DPI and DP2. have been
designed and provide for differing numbers of fuel feed points. The distributor plates
are capable of positioning at three specific levels relative LO the tube bank. DPI
1.32m x 1.96m in size, ! constructed as double plate sandwiching a pattern of vertical
baffles so dividing the space into the nine even size compartments. Propane feed to
these compartments provides the means of bed ingition to any required pattern. DPI,
supports nine fuel feed nozzles perforating the plate, designed to render an even across
bed supply without risk of plugging during operation. The nozzles are designed as
obliquely truncated short horizontal pipes. Bubble caps al.;o perforate DPI and are so
designed and arranged to ensure complete fluidisation of ' bed without risk of partial
collapse or non fluidiscd zones, being fed from the primary pressuring air supply to the
corabustor. Distributor plate DP.' is constructed in essentially the same manner as DPI
but provides for four coal feed nozzles. Thus, both DPI and DP2 provide for a range of
coal feed rates and supply patterns requiring study during the experimental prograrone.
The associated pipework between the lower pressure shell dome and the distributor plate
is so designed that all necessary dismantling, adjustment, and assembly require the mini-
mum of worktime.
HEAT TRANSFER TUBE BUNDLES
Heat transfer from both rhe combustion zone and the off-gas is -.indertaken by j
number of horizontally orientated tube bundles so designed as to present the necessary
surface area for required duty with the minimum disturbance of the in-bcd flow patterns
necessary for efficient operation. All of the heat transfer tube bundles installed in
the combustor, heat exchanger, ,?nd auxiliary heat exchanger are connected to complete
the steam water circuit, which provides steam to drive the turbo-compressor and. under
most test conditions, will also export steam to tv-! power station range (Figure 5).
ASH DISCHARGE
Problematical ash discharging from the combustor. whilst ensuring an efficient
seal has envoked an inclined screw conveyor proposal operating at variable speed to ac-
commodate proper ash discharge rates as demanded by combustor operation. Maintaining
full discharge chutes from the combustor would effect the seal. A precise layout of the
proposed screw conveyor arrangement in consideration of the Petrocarb fuel injection
system ir. in preparation, possibly requiring modification should an incline screw con-
veyor bfe inoperable.
317
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Figur* 5. NCB (IEA Strvica) Ltd. Grimtthorp* Exporimental Facility
Stesm/Woter Circuit. Tube Bank A
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GAS CLEANING EQUIPMENT
The Exhaust Gas Cleaning Equipment contract was let to Head Wrightson Process
Engineering Ltd. early this year and is still in the detail design stage with a number
of technical matters at present remaining unresolved.
The overall design provides for gas cleaning equipment serving to remove par-
ticulate matter that will be elutriated from the fluidised bed, and is interposed between
the combustor and the primary heat exchanger. Flue gas leaving the combustor is divided
between the four primary cyclones arranged circumferentially about the top of the pres-
sure shell.
Refractory lined ducting will connect each primary cyclone individually to a
secondary cyclone of which there are four arranged in a line. The secondary cyclone
outlets terminate in a gas manifold connected to the principle heat exchanger. A
second refractory duct from the secondary cyclone manifold is provided for connection
to the test cascade loop.
Ash disposal from each primary, and each secondary cyclone is provided by
quenching at pressure using a controlled water supply. The resultant slurry is passed
to a pressurised holding and pressure let down vessel prior to discharge to the water
recovery and ash disposal system. Ancillary items include low pressure (L.P.) and high
pressure (H.P.) slurry pumps for the ash handling system, and a silencer in the main gas
path before final exhaust gas delivery to the chimney stack.
The water to be used for the ash quenching is to be re-cycled and requires
chemical processing to maintain it at a specified quality. Plant is to be installed
in the facility to settle particulate matter from the water. Cooling and pH adjustment
of the supernatant water takes place before laggon storage, with intcrmittant biocide
dosing to maintain a freedom from algal growth in the cooling tower.
TURBO-COMPRESSOR
The remaining item of m;. 'or pl.-.nt is 'he Turbo-Compressor, the contract for
which was let to Compair Industrial '.-c. in October 1976. Fabrication of the main
castings is now well in hand am': or. schedule al'.hough some minor delciys have occured
with some of the detail design.
The Turbo-Compressor consists of a low pressure (L.P.) multistage centrigugal
compressor coupled to a high pressure (H.P.) centr ifu)-.:l compressor both driven by a
steam turbine powered by the team fctnerated during norcal combustor operation, although
as previously described the turbine wil.i :;u powered fro-n tf.c ;.iปwer station's own steam
supply during start up.
The Turbo-Compressor has a design delivr-ry of 31kg srr~ dry air mass at a dis-
charge pressure of 12 bar absolute, used to r>ressurlse the corob'-.-itor and supply the
fluidising air flow demanded by the combustc-r operating cond: t ;..ms.
CONCLUSIONS
In conclusion there are three points I wish > :-.ike:
1. The co-opeiation of all concerned with l!-.e project has been excellent. The
sponsor's input--that is government departments ot three nations, the technical input.
the design input, and the contractors have all integrated so effectively "-hat we are
still on schedule for both cost and time. And here I feel that particular mention
should be made of the efforts of the three contractors concerned with the heart of the
plant, namely, Petrocarb of New York, U.S.A., V.K.W. of Dusseldorf, Germany and Heat
Wrightson of Stockton on Tees, Great Britain.
2. The planning of an experiment?! programme for the 4 year period following
commissioning of the plant is now the most important aspect of the Grimethorpe exercise.
This planning is about to commence in detail at the first meeting of the Technical Com-
mittee on Thursday here in Washington and the ultimate objectives must be, in my opinion,
to produce a test programme which will enable the design parameters for say, 200 MW
d-Tuonstration plant.
319
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3. The Operating Agent, the !!CB (IEA Services) Limited, has assembled and in-
tegrated a unique and experienced tcara--A Task Force--of technical people, design people,
consultants and contractors who are capable of extending their activities to achieve
such an objective and I would like to see them charged with the function of carrying
out such an objective so that the "nO MW design parameters could then be translated
into a demonstration plant ns soon as the Grimcthorpe Experimental Facility has proved
the points on the extrapolation curve. This is surely the sensible way of keeping up
the momentum and speeding development of P.F.B.C.
320
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QUESTIONS/RESPONSES/COMMENTS
DR. FREEDMAN, CHAIRMAN: Mr. Broadbent is here now, and he has
to leave in a few minutes. Are there any questions on the IEA pro-
ject and the mechanical design of the equipment and systems?
SPEAKER (from the floor): Can you give the size of the unit?
The dimensions?
MR. BROADBENT: The bed is six feet square in dimension, in plan
area; ai>d it is about 24 feet in height. 'hat'ซ the comhustor withi-i
the pressure vessel.
SPEAKER (from the floor): What is the size of the pressure
vessel?
MR. BROADBENT: The pressure vessel is about a little over 12 feet
in diameter, and about 42 feet, fran dome to dor.ie.
DR. FREEDMAN: The pressure vessel weight is about 100 tons, Mr.
Broadbent says. Since it's an IEA project, Isn't it two meters by two
meters, and not six feet.
MR. BROADBENT: Yes. Yes; two meters by two meters.
SPEAKER (from the floor): Could Mr. Broadbent tell us a little
bit about the particle cleanup aspect of his work?
MR. BROAPBEMT: Well, we have two sets of cyclones in series. There
are, in fact, four cyclones, primary cyclones, arranged around the
pressure vessel; and then behind that there is another set of cyc-
lones. These are what in England is called the "Steerman High-
Efficiency Cyclone," which is, you know, we U .nk the best inertial
system there is around.
DR. FREEDMAN: Well, then, thank you David, and thank .you, Steve.
321
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Instrumentation
323
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INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN: Our first paper this morning is
entitled, "A High-Temperature, High-Pressure Isokinetic/Isothermal
Sampling System for Pressurized Fluidized Bed Applications." This is
a paper produced by the General Electric Company, and will be pre-
sented by Dr. James Wang, who received his Bachelor's degree from
National Taiwan University, and subsequently his MS and Sc.D. from
MIT. He did post-doctoral research at the 'Jniversity of Michigan,
and joined the General Electric Corporate RaD Center in 1972. He has
had multidisciplinary projects and activities including such things
as laser vilisymmetry development, jet noise suppression, gas turbine
exhaust flow studies, and nuclear core flow hydraulic studies. Dr.
Wang?
325
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A High-Temperature High-Pressure
IsokineXc/lsothermal Sampling System for
Pressurized Fluidized Bed Applications
w J.C.F. Wang.C.G. Bingwall
*W c.M. Thoenr js
General Electric Company
ABSTRACT Schenectady, New York
A program has been str^-'g^-ed to characterize the efflux from a Pressurized
Fluidized Bed (PFB) Reactor-. i,
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SAMPLING SYSTEM
A schematic of the high-temperature high-pressure (HTHP) isokinetic/isothernal
sampling system is sKjwn in Figure 1. It consists of three major components:
1. Isokinetio sampling probe and controller
2. Isothermal environment controller
3. Particulate and vapor sampler
Each of these categories is described in detail in the following sections. I'.ost
parts of this HTHP sampling systwn are made of Kastelloy-X which has been confirmed**?)
to provide minimal surface corrosion under a PFB effluent environment.
ISOKIHETIC SAMPLING PROBE AND CONTROLLER
This component group consists of (1) a sampling probe, (2) two static pressure
sensors, (3) a S-type probe, CO a high-temperature shut-off valve, (5) an isokinetic
sampling controller, (6) a control valve, and (7) a flow meter.
Two types of sampling probe were designed and fabricated. The first one is sim-
ilar to that use-* in Reference 1 and is shown in Figure 2. The free jtrean static pres-
sure taps at O.Gi-inch -iiameter are on the 0.125-inch O.D. tubing above the sampling
nozzle. The sampled flow sensor is the second 0.125-inch O.D. tubing on the same axis
of the sampling nozzle with the static pressure taps near the entrance of the nozzle.
The sampling nozzle is made of 0.25-inch O.D. and 0.025-inch thick wall Hastelloy-X
tubing. A picture of this sampling probe assembly is shown in Figure 3- The perfor-
mance of this configuration wan tested in a room temperature wind tunnel. Free stream
air velocity was measured by a pitot-static probe. The sampling flow was established
via a vacuum pump and measu"ed by a Brooks 10 CFM flow meter. The differential static
pressure signal from the free stream sensor and that from the sensor at the sampling
nozzle entrance were zeroed through the throttling adju.-*ment on th< vacuum pump. The
mean sampling velocity was compared to the free stream air velocity and is shown in
Figure iJ. In the air velocity range from 20 to 100 ft/sec, 'he mean sar.pling velocity
was about 8$ higher than the free stream air velocity. This constant bias character-
istic can be corrected by the electrical control circuit. We can preset the sampling
velocity to be BJ higher than the free stream velocity for all the velocity conditions
or adjust the biased value according to the calibration curve at each velocity condi-
tion. Eased on the bench test results, we are able to maintain isokinetic sampling
conditions in the velocity range from 20 to 100 ft/sec to within ป 5$ accuracy. The
effect of blockage of the vertical portion of the sampling tube was found to be only
2% of the mean velocity measured and can also be corrected easily.
Because of a concern for potential plugging of the sampling nozzle and sensors
in the high dust loading PFB effluence, a second sampling head was designed (Figure 5)
The static tap holes on the nozzles inside wall can be made to O.OUO-inch diameter in-
stead of 0.02-inch en the first sensor tube. The center-body type sensor tube on the
axis of the nozzle entrance in the first design (Figure 2) does not exist in this sec-
ond sampling head design. This difference is expected to minimize the possibility of
sampling nozzle plugging in the latter arrangement. Room temperature wind tunnel test
on this sampling nozzle arrangement is presently in progress.
The S-type pilot probe (Figure 3) is used to measure the actual velocity of the
sampled gas stream. The performance of this probi was tested in the rcon temperature
windtunnel and compared against a pitot-static probe measurement. The calibration data
is shewn in Figure 6 and agrees closely with our prediction.
The high-temperature shutoff valve is made of stainless steel 3'0 body with_S
te #6 ball. The center body gasket, seals, a "
operating temperature and pressure ratings are 1750
urc 7 shows a picture of the assembled valve.
a 9-inch O.D., 2-kW electrical furnace is built
Stellite #6 ball. The center body gasket, seals, and gland packing are grafoil. The
750 F and 220 psi, respectively. Fig-
ure 7 shows a picture of the assembled valve. Due to the physical size cf the valve,
to maintain the valve at 1750 F.
The block diagram of the isokinetic sampling controller and peripheral equipment
is shown in Figure 8. The differential pressure between the two static ports of the
velocity sensor (Figure 2) is monitored with a sensitive pressure to an elect! ! trans-
ducer located in a pressurized cold gas environment. The differential pressuie is
327
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KV2ปi VWPLIWi
wtcee
^-,i--ซ ^O
CONTROLUJ FLOW L.. EXHMIST
(TCIMTC
.* AK R
(iBK'.VI
Figure 1. Schematic at the HTHP liokinetic/liothcrmat Sampling Syrtem
328
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FREE STREAM
STATIC PRESSURE SENSOR
KOZZLE INLET
STATIC PRESSURE
SENSOR
SAMPLING NOZZLE
SAMPLED GAS
TO PRESSURE
TRANSDUCERS
Figure 2. Original lioki.ietic Sampling Probe Design
329
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Figure 3. Isokinetic Sampling Probe Assembly
330
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1.2
1.1
*s 1.0
0.9
0.8
BASED ON PITOT TUBE
.1.03
CORRECTED FOR BLOCKAGE
VN = SA.rPl.KiG NOZZLE VELC-CITY
Vs = FREE AIR STREAM VELOCITY
20 40 60 80
FREE AIR STREAM VELOCITY - FT/SEC
Figure 4. liokinetic Sampler Teปt Results 0.319" I.U. Nouto
FREE STREAM
STATIC PRESSURE SENSOR
SAMPLING
NOZZLE
NOZZLE INLET
STATIC PRESSURE
SENSOR
TO PRESSURE
TRANSDUCERS
Figure 5. liokinetic Sampling Probe Back-Up Design
100
331
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OUTPUT -
INCHES H20
TEMPERATURE 658F
PRESSURE 14.7 PSIA
20 40 60
AIR VELOCITY - FT/SEC
80
JOO
Figure 6. S Probe Calibration
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-'= or ST. tf.rl :,t:ซ OT f'l. M,-:!1;ปH ,C'f: ป'ป, ซ'ป|rit 11!ซ ui'tl
Tigur* 7. High-Tempซratur* Shutofl Valvo
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to
SAMPLING
NOZZLES
t SENSORS
I
1
I
P/E
7RANSD.
I
N
PUR
I
7RANSD.
?"
SURE
SEL
PURGE
1
ELECTRONIC
CONTROLLER
(
SAMPLE
i HOLD
POKER
AMP
Figurt 8. Block Diagram of Sampler System
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related to the difference in the free stream gas velocity and the gas velocity at the
sampling nozzle inlet. A differential pressure of zero corresponds to isokinetic flow
in the sampling nozzle. The output from the pressure transducer is integrated (to pro-
vide high steady state gain) and differentiated (to provide dynamic stability to the
control system) by an electronic controller. The resultant signal is then amplified
in a power amplifier coupled to the torque motor of a flow throttling valve. The inte-
gration in the controller maintains the output of the pressure transducer at a steady
state value of essentially zero which corresponds to isokinetic sampling.
The velocity sensing ports are periodically backflushed with dry nitrogen to pre-
vent blockage of the ports. The backflush cycle is controlled by an electric mechani-
cal program with a five (5) minute cycle time. The duration of the backflush can be
varied between 10 to 90 percent of the total cycle time. During the backflush, the gas
flow in the sensing line may produce an erroneous signal from the pressure transducer.
Consequently, during this period the automatic control is interrupted by an electronic
"sample and hold" circuit. This circuit continuously monitors the error signal at the
input of a power amplifier driving the flow control valve. Several seconds prior to
the backflush flow, the error signal is stored and held in the "sample and hold" cir-
cuit. The control valve area is held at the value corresponding to the isokinetic flow
just prior to the installation of backflush for the duration of the backflush cycle.
The periodic backflush sequence can be overridden by the operator at any time.
This allows a continuous backflush mode of operation. When operating in this mode, the
flow must be reduced to levels which do not produce erroneous signals in the velocity
sensor line.
The pressure transducers and electrically operated valves associated with back-
flushing ore located in a cold gas pressurized vessel (shown in Figure 9). The static
pressure in the vessel is maintained just slightly higher than the static pressure in
the fluidized bed reactor. This equalizes the pressure across val^e seals, tube con-
nectors, and the pressure transducers, thus minimizing problems due to leakage in seals
and connectors. The pressure in the vessel is maintained by a tube connecting the ves-
sel interior to the vessel from a check valve connected to a nitrogen supply 10 to 20
psi higher than the sampled gas pressure. The vessel pressure will then be higher than
the sampled static pressure by the line pressure drop required to support the input
flow.
The flow through the sampler is monitored with a commercially available mass
flow meter. The meter measures the true mass flow ind requires no calibration for gas
temperature or gas pressure. The output of the flow meter is an analog electrical sig-
nal which can be stored in the data logger. The flow meter and the flow control valve
assembly are shown in Figure 10.
This automatic isokinetic sampling controller is also desipned to be operated
manually. For u Specific sampling nozzle area, the mass flow corresponding to isokin-
etic condition can be computed fron the measured pressu.-e, temperature, and velocity
of the PFB exhaust at the sampling location. The sampled gas flow can be manually ad-
Justed via the control valve driver to match the computed isokinetic mass flow. This
manual operation provides an independent backup system for the isokinetic controller.
ISOTHERMAL ENVIRONMENT CONTROLLER
To keep the sampling gas at the temperature of the PFB exhaust, the entire sam-
pling system from the sampling nozzle to the entrance of the vapor condenser is heated
via electrical furnaces. The sampling nozzle and sensors inside the sampling "tee"
section are insulated with Johns-Hanville Min-K material. The sampling tube and the
high-temperature shutoff valve between the "tee" flange and the flange on the pressure
vessel are enclosed inside a 9-inch O.D. electrical furnace. This electrical furnace
employs 208-volt 10-amp single phase power. This electrical furnace and shutoff valve
configuration was tested in the laboratory and successfully maintained at 1750 F +5ฐF
after 30 minutes from room temperature conditions. . ~
Tne entire cyclone train for particulate sampling is enclosed in a Inconel 601
pressure vessel (Figure II). Pure nitrogen is maintained inside the vessel at the
pressure of the PFB exhaust to minimize the pressure differential across the cyclone
335
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Figura9. Inkmrtc Sapling Contralto
336
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FLOW M
CONTROL VALVE
Figur* 10. Control Vllv* ปnd Flow Meter Aoembly
-------
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? t? xi
u>
U>
. j
--------- ,
.t .
jxj
,5
V.
FigurtH. Schematic of Cyclona ind Pmsura Veuel
-------
train. The pressure vessel is heated by a 36-inch long and 1U-inch O.D. electrical
furnace which is designed to input t^O-volt, 36-arap single phase power. During the
checkout t'.st, it took 145 minutes to heat the pressure vessel and the cyclor.e train
assembly to 1750 F from room temperature. Temperature variat on at 175GฐF was found
less than ฑ5 F. Figure 12 shows a picture of this pressure vessel and furnace assembly.
The cuter case of the furnace is designed to provide weatherproofing and electrically
safe protection.
PARTICIPATE AND VAPOR SAMPLER
The dust loading at the PFB exhaust is expected to be high, e.p 10 grain/SCF.
In addition to their inherent shortcomings such as particle rebounce, ' 3'etc., commer-
cially available impactors would also require dilution in fluid application. A multi-
ple-stage cyclone train appears to be the best choice in this high dust loading high-
tetrperature application. A multi-stage cyclone train designed and built by Southern
Research Institute (SRI) for the Environmental Protection Agency, has been selected as
the particulate sampler and collector. This cyclone train has five stages, each with
a different cut-size ranging from 0.3 i"" to 7 urn. Beyond the f.'.'th stage, a 0.14-inch
thick Astroquartz mat acts as a positive filter element to collect particulates of less
than 0.3um diameter. The entire cyclone train and the positive filter holder assembly
is made of Hastelloy-X and is shown in Figure 13. Results of SRI performance testing
of these five cyclones at room temp'rature is shown in Figure 1U. >^J
At. the inlet of the cyclone train, a 0.02-inch diameter orifice is installed on
the side port of a "tee" connector (Figure 15). This orifice provides a connection be-
tween the pressure vessel and the cyclone train. At the beginning of each sampling
test, dry nitrogen can be passed through this orifice to the cyclone train and the rest
of the sampling line. At the end of each test, the hot dry nitrogen inside the pres-
sure vessel will again be flushed through the cyclone train and the sampling line.
This nitrogen purge is an important operation to ensure the last sampled gas just be-
fore the close of the shutoff valve to be pushed through the sampling line.
After the cyclone train and filter assembly, the sampled gas is fed into a vapor
condenser which is made of a 2t-inch long and 2-inch diameter Hastelloy-X tubing sur-
rounded by ice and water. The gas temperature at the exit of this condenser was about
200 F during the bench tests at 1750 F upstream temperature. Most of the alkali and
water vapors are expected to be condensed inside the condenser. The condensed vapor
is then analyzed by the atomic absorption technique to identify the alkali setal con-
centrations.
PFB EFFLUX TESTING AND ANALYSIS
The objective of this HTHP isokinetic/isothermal sampling system development is
to employ it to characterize the particulate and vapor constituents at the PFB exhaust
and various points in the cleanup train. Thus, a comprehensive model of the efflux fron
a PFB Combustor and the transfer furstion of the cleanup equipment can ce established
to assess the utilization of the Coai-firec? Combined Cycle (CFCC) Systes as described
in a related paper by Bekofske, et al. An this conference measurements using this de-
veloped HTHP sampling system will be accomplished as part of the existing United States
ERDA-funded CFCC development test series at f.'CB/CURL, Leatherhead, England in early 1973.
Two HTHP sampling systems have been fabricated and assembled to be used at HCB/
CURL. One system will be installed between the PFB exhaust and inlet to hot gas cleanup
equipment. The second system will be installed at the exit of the cleanup device. The
measurements from these two sampling systems will be used to characterize the efflux
from the NCB/CURL PFB and the equipment transfer function of the Aerodyne Cyclone.
The collected particulate and vapor in the HTHP constituents sampling system will
be analyzed to obtain information on the size distribution, constituents, dust loading,
stickiness, and other mechanical characteristics. The following five types of analysis
on the sampled particulate and condensed vapor are the basic information to be deter-
mined.
339
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. ;,,-
.
Figure 12. Cyclone Train Ho'ising and Furnace
-------
Figure 13. Five Stag* Cyclone Sampling Turn and Pontivi Filter
-------
e
t>
o
14
&
100,
8(
S
8
60
4C
20
O C/clone I
A Cyclone II
O Cyclone III
V Cyclone IV
O Cyclone V
, I
0.1 1.0
PARTICLE DIAMETER, micrometers
Figure 14. Five-Stage Cydona Train Room Temperature
Calibration Data (Courtesy of SRI)
10
SAHPLED GAS I
tHLET ป
OUTLET
CYCLOHE TRAIN
Figure 15. Cydona Train Nitrogan Purge Arrangement
342
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Quantity of participate matter (or dust loading)
Cyclone train and microbalance
Paniculate size distribution
Cyclone train and microbalance
Coulter counter
Banco -ounter
Chemical Composition (Ca, S, C, Cl, :-;6, K, N'a, ..=h, etc.)
Atomic absorption
X-ray fluorescence
Stickiness of the particulate
Scanning electron microscope
Mechanical characteristics of the particulate
Scanning electron microscope
SUMMARY
Two HTHP isokinetic/isothermal sampling systems are developed and huilt at Gen-
eral Electric to te employed at t.'CB/CURL in conjunction with the CFCC Development test
series in early 1978. Each component of the sampling system was tested at CF.-CRD with
satisfactory results. The isokinetic and isothermal features of the system were con-
firmed based en our bench test results. The backflushing and automatic velocity track-
ing capabilities for high dust loading environments are uniquely useful in the PFB ap-
plication. The checkout test for the assembly is underway at elevated temperature and
pressure conditions.
ACKNOWLEDGMENT
The authors wish to thank Mr. R. Tourin of Hew York State ERDA for his consis-
tent support of this work and Messrs. P. Niclas, E. Wheeler, and P. Viscosl for their
help in oesign, assembly, and checkout.
REFERENCES
1. Rlnpwall, C.G., "Compact Sampling System for Collection of Particulates from Sta-
tionary Sources," EPA-65D/2-7I4-0/9, Contract Ho. 68-02-0516, April, 197^.
2. Cray, B. and KcCarron, R.L., General Electric, personal ccmnunication, April, 1977.
3. McCain, J.D., "Impactors: Theory, Practical Operating Problems, and Interferences,"
Proceedings of the Workshop on Sampling, Analysis, and Monitoring of Stack Enis-
sicns, FB-2527i48, April, 1976.
it. Smith, W.B. and Wilson, R.R., Jr., "The Development and Laboratory Evaluation of
a Five-Stage Cyclone," EPA Contract Ho. 08-02-2131, (Final Report is in preparation).
-------
QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: Are then, any questions for Dr.
Wang? Dr. Wang, have you accumulated some?
DK. WANG: Yes. I have a question here from Sam Wolosin, excuse
my pronunciation, from Curtiss Wright Company. There are two ques-
tions. One: "Does the baffle for gas mixing remain in the system at
all times? It seems it would be subject to erosion of active particu-
lates in the gas stream."
We intend to have the baffle in the system all the time. It can
be removed. Actually, the first test we are goinq to run will be a
75-hour test, so by the end of the test we should know what damage
erosion would do on the baffle. Actually, there is one time that we
are thinking of using the baffle plate as the erosion-test material.
The system will be hot during i:he- test, so I don't think there
io much more corrosion. By the way, the baffle is not a flat baffle.
It's a curbed, plumb-body type of baffle; it's a spherical surface,
so it should have minimum disturbance to the flow and cause less
deposition on the surface.
So at this stage we don't know yet. The baffle plate itself is
made of Hastelloy X material; and we have some previous experience
with it. It gave us minimum erosion in that PFB environment. That's
the reason the whole system we're using, you know, from the baffle
plate to the nozzle, the sampling sensors, the five-stage cycling
trains, the positive filtered supports, even the condensers, are all
made of Hastelloy X.
The second question is, "Have any bench tests been made introduc-
ing a known amount of particulates and alkaline metal contaminations
into the system and collecting for calibration of the system? What
were the results?"
We are in the process of running a bench test in the laboratory
with the clean gas to start with, and using the same pressure and
temperature. And then we were planning, actually, m?ybe this week or
next, putting aluminum oxide powders of different sizes into the
system and do a hot calibration.
The calibration curve I showed you on the slides is the cold gas
under room temperature gas calibration made by SRI, before they
delivered the system to us. So we do need a hot gas calibration,
because when the gas is getting into higher temperatures and pres-
sures, the characteristic changes and the cut size will be changed too.
344
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The resulcs haven't come out yet. We will be going that before we do
tfiซi tests on PFB.
By the way, I wanted to riake one comment. I think this is really
the first of this kind of sampling probe dealing wi*.h truly PFR
exhaust studies, because it keeps all the conditions the same as the
sampled i'FR exhaust: the temperature, the pressure and everything; and
there are no condensations, there are no dilutions and so forth, and
it reasures the aerodynamic size. Besides, it also provides an
opportunity to take a look at the particulates, or alkali vapors
collected, and correlate with any other method you want to use after-
wards. So I think this is really the first of this kind.
345
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INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN: The second paper this morning is
entitled "Participate Analysis Instrumentation for Advanced Combus-
tion Systems." It is authored by Drs. Muly and Van Valkenburg of
the Leeds and Northrup Company. The paper will be presented by Dr.
Huly, who is a graduate of Johns Hop!:ins University, and from there
went to Northwestern University, where he received both his MS and
Ph.D. degrees in Electrical Engineering.
Dr. Muly joined Leeds and Northrup in 1971 as Program Manager
for Environmental and Industrial Instrumentation. He is presently
Program Manager for Particulate Analysis Products in the Advanced
Business Development Department. Dr. Muly.
346
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Paniculate Analysis Instrumentation
for Advanced Combustion Systems
Emit C. Muly and Ernest S. VanVa'kenburg
Leeds* Northrup Company
ABSTRACT
An on-line paniculate- analvsis inv.r'jmt-nt h.is bee" developed bv Ix-e-ds and
J.'orthrup Conpany under EKDA sponsorship for ronitorinf. the concentration and size dis-
tribution cf particles in the cas .-lean ur> siape of advanced eonbu.st ion systems. This
instrunent utilizes Iou-.inr.le forward scatter In,": of laser illuminated oarticli-s bv
across-the-duct cx-asurements. A prototype instrument has been tested or a fluidized-
bed combustion system at the Arp.onne National Laboratorv. This r>ar>er Dresent.fi a brief
description of this instrument and discusses the results obtained frora the Arj'.onne
tests.
I.'.TRODl'CTIO:;
An on-line particulate analvsis instrument h/is been developed bv \j- -i!s i, "orth-
rup Company under EKDA sponsorship for jn-situ measurement of particle !o;.jinp and si::e
in the product pas sire.ims of advanc-.-d combustion svsters. A nrototvpe in".trurซ.-n'. has
been aesir.ned and constructed co evaluate this means of measuring particles on fluidl-
zed bed combustion systems and field tests have been completed on this unit at the
Arp.onne National Laboratory.
This particulate instrumental ion is based on Leeds f< Northrun's orior research
in low-an^le forward scattcrim; of llrht by r.icron si/-:- particles pusnendcd in fluid
streams. When such particles are optically lllunrnated. the scattered li;-ht Intensity
at any civen an?,le is a function r-f the size, siiape and index of refraction of the
particles. In the case vhere the wavelength is snail in comparison to the size of the
particles, the spatial distribution of the scattered iiy.rit in the far field !s domi-
nated by the volume and size characteristics of the r<.-irt ic les. The desipn of the KRDA
instrument is based on utilization of sirole diffraction theory to cor.vert measurements
of the ccr.oosite Fraunhofer diffraction pattern for a l.irre number of narticlert into
meaningful particle data which characterizes the size distribution and concentration.
This type of instrumentation is needed to evaluate the performance of seconJarv
particle clean up and to measure the size distribution and concentration of n.irticles
at the inlet to pas turbines in direct combustion croal-fired svstens. The latter
application requires instrumentation amenable to measurement of n.irticles in hir-h
temperature (liOO-2000F) pressurized (up to iO atrxjsphercs) cas streams and be adjpt-
able to rather larp.c diameter pas cucts. The pr-'totypc instrunent reets these require-
ments and accommodates p,as ducts up to one foot infernal diaxeter.
iNSTRL-MEirr DESCRIPTIO:;
A schematic diapram of the onticat train of the ^RDA instrument Is shovn in
Fip.urc 1. Inis instrunent consists of ottlc.il elements mounted to an optical bench
which extends under a horizontal run of the combustion system duct. (The duct is not
shown in the schematic, only its vertical center linn.) The illumination source is a
helium cadmiun laser counted to the underside- of the oDtlcal bench. Koldinr. optics.
attached to the left erd of the optical bench, i.reel the laser bean across the duct
where particles in the r.as stream, vhich arc illuminated by the bean, scatter the
li^.hr. The scattered flux Ib collected bv a lens and focused on a set of function
masks. The portions of flux, which are transmitted throur.h the f. cticn masks, avc
focused by a second lens on to the photosultiplicr detector.
The output pover of the laser Is ceasurcd bv a silicon photo
-------
Stationary Kar.k
Rotator Hask
flotfltor Lens
,Folding Prism
Folding Mirror
tInterference
\ Filter
Figure 1. ParticulMa Instrument Schematic
-------
A photograph of the prototype instrument is shown in Figure 2. This instrument
consists of the optical subsystem assembly, in electronics package. laser pover supply,
a .d a digital line r"'nter. All units, except the printer, are contained in NEMA-12
class enclosures to ceet industrial environmental requirements. The laser and recei-
ving optics units contain thermostat controlled heaters so that the ootical subsystem
may be installed on gas ducts which are either indoors or exposed to the weather out-
of-doors.
The electronics package includes a microcomputer, visual display and all the
operating controls. This unit can be located up to 200 feet from the optical subsys-
tem. The digital data printer is used to log the output measurements and can be loca-
ted remotely from the electronics package, if desired.
The opening on the optical subassembly from the folding optics on the left to the
collection lens bezel is 30 inches. The vertical clearance above the optical bench to
the Laser beam center line is 9%= inches. These dimensions were chosen to accommodate
a 12-inch I.D. duct with a tee section viewing port extending on each side of the duct.
Mechanical adjustments are provided on the folding optics for alignment of the
optical train to the duct windows after the instrument is installed. The instrument
can be mechanically mounted to the duct through the load carrying base structure.
All particles illuminated by the collinated laser beam as it traverses across
the gas duct scatter flux off the beam axis. For particles in the size range 1-100
microns most of the scattered flux is in the near forward direction. He collect that
flux, and through an optical process, convert that data into information concerning the
particle size distribution and the particle loading. The particle loading output is
calibrated to be direct reading in parts per million by volume, i.e.. ratio of total
volume of particles per unit volume of gas.
A common means of presenting size distribution is to plot the population of par-
ticles as a function of their size as shown in figure 3A. In order to obtain data more
useful for control purposes, it is possible to describe size distributions in ways
other than number density. Three means of presenting the same particle distribution
arc shown in Figure 3. the area disbribution DA
-------
CO
en
o
Figure? Ptototyiw IMrl'
liut'umenl
-------
':i r: :rl<- K.u! i ir.. .1
Figute 3. Three Methods of Defining Particle Sue Oittnbution
351
-------
Fiqufeb ClnwupVKwol
ANL
-------
353
-------
The average cean volume -jiameters. 37. for the iirpactor samples were calculated
from truncated lop, r.omai distributions obtained via the Anderson Ir.pactor. Since the
MICROTRAC has a linear size response for particles one micron in diameter and larger.
and an attenuated response to submicron size- particles, the lowest channel (submicron
region) data points from the irrsactor were not used. This nrovides directly comparable-
data over the size ranp.e I-'.") cicrons. The average ncan area diameters. MA. were
similarly calculated fron the- Anderson Impactor data. The differences between the
direct reading, on-line observations via XICROTRAC and the impactor data arc tabulated.
The meeiian diarieter, V) t h pc rccn t i I e for lop, norm.nl distributions can be ex-
pressed as Median Diameter * .??/ x MA. The results of this computation are shown in the
bottom three rows of Table I.
These results indicate pood agreement between optical scattering and cascade
tmpactor methods for particle sizing. In all but one case, the difference is less than
one micron.
Two interesting characteristics arc observed, however. The MICROTRAC size mea-
surement tends to indicate slirhtly smaller size for the median diameter and the
MICROTRAC shows the average size of particles coninp. out of the final filter to be
larf.er than at the input. In one of the three tests, the inpactor data also shows the
output particle size to be greater than the input. This anomaly will be explained
later in discussion of test chronology.
The loading, data as functions of time are shown in Figures 6-7 for two opera-
tional tests. Th. in-situ volurttric loadings, as outnutted by the MICROTRAC instru-
ment, are converted to standard pressure/tenr.erature conditions by the following equa-
tion:
I (Graras/ra1) - D(l * 0.003fe7T)dV (2)
p
where D - density of paniculate (p.m/cc)
T ป c.as temperature, nominally 160C
P = Ras pressure, nominallv 'i atms.
dV instrument output in ppa.
The density of the material samples in these tests was 1.2 ;^n/cc.
The data from MICROTRAC are shown ns dots for unit intervals of time. The
loadinp. is always hieh at the b<-s>,inninp, of each run due to material loosened in set tine.
the duct valves It takes about 15 to 20 minutes for this Lo be ourp.cd out of the c.as
strcara.
The MICROTRAC loading da'a show sl^w oscillator-/ variations. Initially, wo
thought that this raisht be instrumental error caused by chanpc in laser beam intensllv.
However, observations of particle size as a function of tine show similar variations
and these data arc, independent cf laser intensity. Therefore. --o conclude that these
oscillations are characteristic of the Ar^onne/Coabustor process.
Sncoth lines are drawn through the MICKOTRAC data points to show the loadinn
' trends. In addition, the loading values obtained by the cascade iirpactor and the nen-
brane filter samples are shown. The horizontal location and lonc.th of lines for the
extracted sample loadings indicate approximate tice and duration for collection of
each of these samples. The vertical locations of these lines dcs'.p.nate the sample
loading measurements for those intervals.
A sip.nificant part of the variance between MICROTRAC and the extracted samolos
may be due to non-uniformity of loading across ..he duct or. nrobably more likelv, due
to problems of achieving isokcnetic sampling with the extractive probe. A oarticular
advantage of the MICROTRAC type of instrument is that it measures all oarticlcs passinp
through the laser beam and t'-.e gas flow is not influenced in any way by this method of
measurement.
354
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He-Irian Filti-r S.-imnlc
Outnut of Cvclones
Output nf "tt.nl Kilter
II
Figure 6. Experiment L&N-4
Outr>ut of Cvcloncs
'.i-lrnn F\ Itcr
Amlcrson !~.".ictor S.'iirplc
iclnan Filter Sample
* ป
Oi:tnut of 'lot.il Filter
Andc-rsnn Inn.ictor S.innlc
1 I I I I I ป I I I
10
:o 30
>! I N U T E S
50
60
Figure 7. Experiment L&N-6
355
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Figures 8-9 show the tine variations of the mean area dianeter (MA) superimposed
en the loading functions. It is noted that the size variations tend to follow closely
the loading variations. Thus, it is concluded that the size of particles is a major
contributor to the total loadings in thij process.
Furthermore, the amplitude of variation in MA is consistently largest at. the
beginning of the metal filter output reasureirsents and the mean size tcnos to decrease
with time. This is characteristic of the performance of the AM. metal filter. The
efficiency cf filtration improves as particulate builds up in the filter media.
It is also observed in these data, that the MA variations at the input to the
final filter are less than ac the output. The test sequence at ANT. soecificd first
measurement of the particulatc at the output of the netal filter, then at the input.
For all three combustion tests, the output measurements were made in the morning, a
few hours after start-up, and the input measurements occurred in the afternoon. One
can now only hypothesize that this combustion system requires a long time to stabilize
and that the netal filter output variations would have been smaller had the test se-
quence been reversed. This also explains why the average size over the total measure-
ment time was larger at the output than the input. For example, at the end of the
IA"J-5 metal filter output run, the average MA was 1.73 microns and the input average
particle diameter, measured two to three hours later, was 2.26 microns, these results
Indicate the desirability of having two MICROTRAC instruments to enable simultaneous
measurements at thซ input and output of the final filter in order to obtain complete
characterization of a coiabustor/f iltration process.
After completion of the tests at ANL. the prototype instrument was returned to
L&N where the calibration of the particulate loading output was rechecked. For a cali-
brated sample of 10.0 ppm diamond dust of nominal 3 microns diameter, the instrument
read 10.11 ppm, while the standard deviation of this calibration procedure is 0.18 ppm.
Thus, the calibration constant for dV did not change in a measurable degree during two
months of operation.
CONCLUSIONS AND FUTURE WORK
The results of those tests on the AJIL fluidized-bed combustion system indicate
that the MICROTRAC data output compares reasonably well with Anderson impact"- measured
samples. Furthermore, differences observed for these two types of measuremc..i.j nay be
attributed to the means of extracting the particulate sample for the cascade impactor.
Three problems were observed with the MICROTRAC instrument and all three pro-
blems seem related to the laser:
(1) The laser output power slowly decreased with time,
(2) The loading sensitivity (0. 1" graii.j/scf on a 1 inch i.u. duct) is
marginal for after final filtration measurements, and
(3) A significant part of the particulate loading is contributed by
submicron size particles which sre not measurable by the prototype
instrument.
The latter problem was anticipated since the instrument response falls rapidly
as the particle size (diameter) approaches the wavelength of the laser emission
(0.442 ;:m). Improvements in submicron response can be achieved in future in.-truments
by use of a highly stable ultraviolet laser which is now commercially available.
The loading sensitivity observed at ANL was about three tines less sensitive
than predicted from our laboratory measurements. This is directly relatable to the
instabilities of the laser output oowcr in th.-.t instrument. Loading sensitivity of
0.001 grains/sr.f should be achievable on a 10 inch i.d. duct at 10 atns. with an
Improved laser which now exists.
The Argonne tests demonstrate the applicability of forward scattering for on-
line in situ measurements of particulate size and concentration in fluidized-bed com-
bustion processes. Current research, which Leeds & t.'orthrup is conducting for the U.S.
Environmental Protection Agency, indicates that use of combined forward and 90ฐ scat-
tering will enable measurement of submicron and larger size particles at lower concen-
trations, in the ppb rather than ppm range. Thus, optical scattering type instruments
356
-------
-13.0
MA AndersSTr*
Innactor
Input """MA MICRnTKAC
Inrut IVICROTRAC
HA Aruil'f .*ion Inr'.TC tor
Outnut I MICROTRAC
I I I I
I I I I I I
30
t i t: v T F. s
Figure 8. Experiment L&N-S
1.0
-i3.0
30 40
MINUTES
1.0
Figure 9. Experiment L&N-6
357
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offer pronisc of net-tin;1, the increased requirements of p,as turbine instrumentation as
advanced fluidized-foed systems evolve.
The prototype forward scattering instrument is ready to bo delivered to Curtiss
Wrif.ht Corporation where it will be installed between the final filter and the turbine-
inlet on the Small t'.ns Turbine System. The f,as duct at that location is a 4-inch i.d.
refractory lines spool niece. With this size
-------
QUESTIONS/RESPONSES/COMMENTS
MR. HANZALEK: Thank you. Dr. Muly. We are a couple of minutes
ahead of schedule, so if there are any direct questions, preliminarily
at least, that anyone would like to address to Dr. Muly, we can handle
one or two at this time. Does anyone have a question on the top of
his head? Tnerr is one. Would you go to the microphone, please?
SPEAKER (not identified): I would like to know at what particle
loading rates your instrument ceases to function?
DR. MULY: The detection limit is about .1 grains per standard
cubic foot for the one inch duct reported on here. The detection
limit is reduced as the ID of the duct or gas pressure increases.
This should result in a detection limit of .001 grains per standard
cubic foot for a 10 inch duct operating at 10 atmospheres pressure.
By increasing either the path length, operating pressure or the power
of the laser the detection limit will be proportionally reduced.
SPEAKER (not identified): What is the maximum sensitivity of
your instrument?
DR. MULY: That again depends on the dianeter of the duct. For
an instrument with aim path length, concentrations as high as 1
thousand parts per million can be measured accurately. The upper
limit of loading measurement also depends upon the particle size
distribution. Particle streams with large particles can be measu :d
at higher concentrations.
DR. RAO, MITRE Corporation: Is the difference between your
instrument readings and the conventional methods due to possible
condensation?
DR. MULY: We did not collect the samples for either the Gellman
filter analysis nor the Anderson impactor analysis. Both the collec-
tion and data analysis were carried out by Argonne personnel. Every
effort was made to keep the temperature at the filter above the dew
point. However, when particulate samples are collected on a filter,
some condensation of the gas phase components may take place. This
could result in erroneously high readings for the solid phase con-
stituents. This might explain the observed higher readings for the
Gellman filter. No effort was expended to determine if this was the
cause for the observed higher loadings given by the filter method.
DR. RAO, MITRE Corporation: How significant is the impact of
condensation on the conventional methods?
359
-------
DR. MULY: We have no experience in evaluating the effects of
condensation on the conventional methods used in this study.
SPEAKER (not identified): What will be the influence of radiant
particles on the measurement?
DR. MULY: The instrument has been designed using a narrow band
pass interference filter (50A at 442 rm). The signal levels there-
fore produced by thermal radiation aro insignificant in the tenoera-
ture ranges of interest between 1500-2000ฐF.
DR. WANG, General Electric Company: Can you explain the reason
why your measurements agree with impactor results? Is there a the-
oretical analysis explaining this agreement?
DR. MULY: The optical method is not responsive to particles
less than 1 micron in diameter. The intercomparison between the two
techniques was, there ."ore, performed by first truncating the impactor
data at 1 micron and then renormalizing the data. The resulting
agreement is therefore, to some extent, expected. No analysis exists
concerning the actual response factor of the impactor device; i.e.,
its shape factor dependancy and therefore no theoretically derived
prediction of the inter-comparison was performed.
360
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INTRODUCTION
FREDERICK HA'JZALEK, CHAIRMAN: We will now proceed to the third
paper, entitled "Particle Field Diagnostics Systems for Fluidized Eed
Combustion Facilities." This paper is offered by '-!r. William
Bachalo of Spectron Development Laboratories, Inc.
Dr. Bachalc received his Ph.D. fron the University of California
at Berkeley, and his degree was in fluid dynamics. He previously
worked at the fJASA Ames Research Center, investigating the phenomenon
of transonic turbulent boundary layer separation, using laser veloci-
metry. While at Anes, he developed holographic interferorcetry for
use in the- transonic flow studies. He is currently in charge of the
particle field diagnostics program at Spectron Development Laborator-
ies in California. Dr. Bachalo?
361
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Particle Field Diagnostics System
for Fluidized Bed Combustion Facilities
William D. Bachalo
Spectron Development Laboratories. Inc.
Costa Mesa. California
ABSTRACT
Particle diagnostics instrumentation that have been developed for application
in fluidized bed combustion plants will be described. The devices have been designed
to make simultaneous on-line measurement;; of particle diameter over two decades of
size range and of the particle velocity for the determination of particle flux or con-
centration. These ruggedized laser-based light scatter detection instruments make use
of angular scattering intensity ratios for sizing particles in the range O.b - lOjm in
diameter and particle sizing interfcrometry (visibility) in the range 2 - lOOura in
diameter. A brief description of the physics involved will be given together with a
discussion of other available techniques.
INTRODUCTION
Particle diagnostics instruments that are to be i-tilized in the hostile environ-
ments of the fluidized bed combustion power plants face a serious challenge. The in-
struments must mako in situ measurements in a flow field environment with temperatures
as high as 1000 to 2000'F and pressure of 1 to 10 atm. A broad range of particulatc
loading and size will occur in these fields. The system flowrate will also vary de-
pending on the facility output conditions making measurements of velocity necessary if
paniculate rate measurement are to be meaningful. At present, no single instrument
exists that can make in situ particle- size measurements over the entire sli:c range of
interest (i.e.. 0.05 to 20nm).
Hone'iiclesr. particle field size measurements are essential in all areas of the
design, development, and operation of the fluidizod bed combustion facilities. For
example, to understand anj mathematically model the combustion processes, including
the formation of soot and fly-ash, an accumulation of data both on the species and
size of the particulates generated must be available. The optimization of particulatc
removal equipment requires upstream and downstream particle content analysis. Par-
ticle impact ton studies on turbines to determine corrosion and erosion rates as a func-
tion of the incident r.ns particulatc content are also dependent on accurate particu-
late size and velocity measurements.
In the operating plant, constant monitoring of the facility will be required.
Real time in situ sizing equipment will be needed to act as early warning sysccms of
unusually high particle concentrations entering the gas turbine or for water droplets
in the steam turbine. Effluents from the operating plant will demand careful assess-
ment of the particle cmmissions. Detrimental effects of particulates '.s. to a large
extent, dependent on the size distribution. Particles in the size range of .Olum to
lOuni are not easily removed from the effluents. It is particles in this size range
that arc effectively deposited in the pulmonary region of the human respiratory sys-
tem, d)
The sizing techniques that are available fall Into categories based on sampling
and optical methods. Sampling requires extraction of a sample from the flow and this
sample Is then analyzed externally with Instrumentation including microscopy, scanning
electron microscopy, weighing, or a combination of these methods. The sampling inher-
ently disturbs the flow field and hence the particle distribution even If the sample
is withdrawn isckinetleally. Furthermore, agglomeration, impaction and deposition
occur in the sampling process. Sample preparation and analysis frequently requires
several hours. Optical methods based on light scatter detection have several advan-
tages over sampling techniques. These include the fact that they produce near real
time analysis, they are non-instrusive. and measurements can be made with high spatial
resolution at exceedingly high rates. A brief review of other techniques can be found
In Reference 2.
Spectron Development Laboratories, Inc. has developed particle sizing systems
based on the scattering of light from highly focused laser beams to make single par-
ticle measurements of both size and velocity at sample rates as high as I0*/sec. With
362
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the proper definition of the sample volume, the measured accumulation rate and the
velocity, a measurement of the particle concentration is available. The systems spe-
cifically dcs'gncd for application to particle field diagnostics in fluidizcd bed com-
bustion plants will be described in this paper.
SCATTERING OF LIGHT BY SMALL PARTICLES
Because the instruments to be described are based on the scattering of light
from small particles, a very brief discussion of the pertinent scattering mechanisns
will be given here. The formal description of the scattering of electromagnetic radi-
ation by homogeneous spheres requires a solution of Maxwell's equations with the ap-
propriate boundary conditions. Such solutions were derived for spheres of arbitrary
size by Mie in 1908. Tins theory, described fully in References 3 and A. is exact for
incident light that is composed of linearly polarized plane parallel waves and that is
constant over the particle. The scattering radiant flux is given as
..n) sin" : + i, (;..,n) cos' M sin -dvd; (1)
where ' is the scatter angle measured from the beam propagation direction. : is the
angle between the plane ci observation and the E-electronagnetic vector. I0 is the in-
cident intensity of wavelength '. , ij and i-> are the Mie coefficients for perpendicular
and parallel polarization components. In the arguments of the Mie coefficients, n is
the index of refraction and < is the non-dimensional size parameter defined as ("d)/>,
the particle circumference divided by the wavelength.
Mie scattering coefficients can be readily generated with the appropriate com-
puter code for spherical particles in the size range of interest. Polar plots of
intensity with respect to the .ingle o . measured from the plane of incidence, called
scattering diagrams arc computed. The scattering diagnms for several particle sizes
shown in Figure 1 serve to show the significant characteristics of the scattering
function. A logarithmic scale is u?ed with each ring representing an order of magni-
tude and tht; dots nark each 5 degrees of angle. These polar scattering diagrams ..re
then used in all phases of instrument design. For example, an instrument designed to
collect forward scattered light would require a knowledge of the variation of scatter
intensity with particle size, collection nngle. collection F/number. index of refrac-
tion. incident polarization and an evaluation of the relative effects of these param-
eters .
Several scatter detection techniques have evolved from a knowledge of the scat-
tering diagrams. Any number uj." scattering angle collection arrangements and incident
sources could be devised as optical counti-rs. Of che techniques available, those
making use of highly focused high intensity beams of light are superior in most appli-
cations. Focusing the light source has the advantage of prouacing a very small sample
volume. Good spatial resolution is available with the high incident intensity on the
particle providing a good signal-to-noiso ratio. Su-h devices sample a single particlo
at a time to produce a measure of the optical scattering diameter. Other considera-
tions of the particle's scattering properties need to he made in order to relate :his
quantity to the actual particle diameter.
Careful consideration of what constitutes a probe volume must he made in all
laser-based scatter detection instruments if a meaningful particle flux is to bo deter-
mined . When using a laser light source operating in the common TCMOO mode, the bea:<
intensity distribution has a Gaussian shape. The incident intensity on the particle
is then dependent on where it passes through the beam. Because there is always a back-
ground noise level due to other particles in the field scattering from outside of the
focal volume, scatter from the access ports, photomultiplicr r.oiue. .inJ other sources.
a minimum detectable signal level exists. Large particles passing through the focal
vuluuc will scatter light above the minimum or threshold level over a larger region
than will smaller particles, as shown in the rough sketch in Figure 2. Dependence of
the effective probe volume on particle properties was recognized by Farmer*') and by
Hirlctran ct al. who analyzed the effect in his dissertation^-. Kith the Gaussian
intensity distribution of the laser illumination, the effective pr^-be dia-.ctcr will in-
crease with increasing particle scatter cross section resulting in concentration mea-
surements that are biased.
363
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Diameter -- 10.0 pm
MAX = 10'
\
Figure 1. Polar Ploti Showing the Magnitudes of tf<ซ Miซ Scattering Functions
lot Typical Soot like Particle) (n - 1.57-056 and X - O5145pm)
364
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II If
1/C2'
-c -
Voltage
Partic'g S.22S
n larqc
-. small
Gaussian Beam
Intensity Distribution
! Signal Output
Threshold
Li'V'71
Figure 2. Dependence of the sample volume on
Scattering Cross Section
To correct t'or the variation in measuring cross section, the effective probe
volume. V(J.pj) where il is the particle diameter and p[ are the relevant particle prop-
pert, ics. must be defined. Because Mle scattering theory assumes the particle to he
uniforrily illuminated, the focal volume must first be formed at typically an order of
maftp.i tude larger in dian-.eter -ban the largest particle to be measured. ;this diameter
is defined to be the rec.ion within which the intensity is above the l/e- of the peak
intensity. Achieving the sufficiently snail focal volume required often makes bca.T.
.expansion necessary. Kocusin,: of the Gaussian beams is treated in detail by Korei-
nik ('). This ai .lysis will ensure that the maximum intensity is incident on the par-
ticle in the probe rcr,ion and that the waves there are plane, satisfying the .'lie
theory assumptions.
Once the focal volume intensity distribution is defined a relationship can be
df-rived to evaluate the effective probe volum^. A sensitive probe tarpit area.
A_(d.n) i * established that is a function of the particle scalte-'ir.g cross section
(uepcndf r.t o^. particle diameter and index of refraction for nor.-absorbing spheres).
In the present case where the particles are primarily soot or other carb naceous mate-
rial, the bcattcrinr, has only a very weak dependence on the index of refraction. To
determine the particle flux a histogram is formed of the particle counts for each
small size interval over a r.iven time period. The sensitive area. A.,(d) is then deter-
mined for each size interval and is used to properly weight the histogram to correct
for the statistical bia. in dctcrrslnin?, particle flux. This phenomenon is a hidden
blcssinr, since there is typically a p.rcater nun-.bcr of particles in the small sizes.
The s:r.all Affective probe arej for the smaller sizes reduces the possibility of mul-
tiple particles existing in the probe volu-.e simultaneously.
Because the refractive index of nost particle fields is unknown, the scattering
instruments should be dcsir.sied to minimize the effects of this parameter. It i:> knov-i
from comparisons to Mie scattering theory and through asymptotic approximations -_o the
theory that the scattering in the near forward direction can be approximated very well
by Fraunhofer diffraction theory. Diffractive scattering is Independent of the^index
of refraction and has been shown to have only a weak dependence on the particle's
shape*-3'. The size measurements of irregular shaped isometric particles (i.e.. par-
ticles that do not have a ^reat difference in their respective dimensions) do not
differ much from spheres of equal projected area.
365
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Availability of f hi- .sc-'i' tor in;- inti.-nsi'y a-: .1 we] 1 -'K-f ::.<: func'. :' on ป:" par'.icle
si/.e h;i:. lซ-:n;'. :< j.ar'. i c ! e- coun'er.s based eiri the
s.casuron.fnf ; of absolu'e scattering intensity. .'-:ซ.- vi-ral d: f f it -jl '. :<-s i~.ป.-d:.-i' el v :.-o-
corxo apparent in i h i :. approach. i'erhaps '.If r..n::< trou:. i <-:;o.~e is li'jt- o _:.ซ_ 'j.-rj.s.sian
intensity ili ;, t r i but ion of the !a:.<-r boan. In a :i-i\ yd i sper.-.e par'icli- field : i .".
po:;sibi(- 10 ;<! si:..i!ar sea', t c-r in;; i nt err; i t i <; . dependin;; on 'ho particle : r.'i j< c r. ory
hrou;;h tljo probe. for p.'irticles vi-ry different in size. Several techni-;';es h.-sv.- been
devised in el forts to (.vซ-n-(.:::t- t ho uncc-r ; ;ii ru y in p.'irticlo ra jc-c'.ory . K'ir t-x.irr.plc-.
.it-roijyn.'i::iic fotusiri;; h;is hcon ur.til lo forco '. hi- p.'ir"- i cits ihrn-iy.'n iht con'i-r ซ!" t he
j/rohi- volurao whi-ro the inltn:> i t / i :; .-ipproxirnt -! y co:is-. /ml . This is nor. iK-fir.'iblo nor
ovi-n po::sibli- if in sinj nif.-i.surcTiont r: ;irc rซ-'j-ii rt.-'i . 'j'hi-r aftt-rip's h.-jvซ- l>ot-n r:.;jdt- i'j
ri-jt-ct .ill tT>issin;;s oiilsido of I he ct-ntr-'il port ion of The foc.'i! '/olimt-. 7::e cri ; ic-il
.'il i i;n:n<-nt required render:; these t <.-chrปi'|uos less th-'in ties i r.'ihle outsidt- of the J.-ปbซir.i-
lory cnvi ronmcnl . In .-'.d;hl by other ;>.-irt iclcs in the p;ith .'ind by acctrr.-.!.-it ion of ;.;ir-
ticul.iie m.-it cr i:i I on the :icce.-;s wint
(he panic1'; velocity. The techniques will be describ.-d in the following .sections.
('ARTICLE SlX.n.'c; IN
I'article si/.in;; by the measurenienl of the laser doppler ;if,nal visibility was
orif.inally invest ij-.at ed by rVimii-r C) jjy that t ice laser velocinetry was oecor-.inf.
well-established as a viable fluid flow measurenenl technique. Analysis of f he scat-
tering phenomena associated with the dual beam laser velocimeter were bein;; investi-
gated in an effort to improve the signal- t o-noise ratio of the instrument. This led
to the real i/.at ion that there was a dependence of the sir.nal modulation on the size
of the scattering particle relative lo the interference frinc.e spacing. Relationships
were derived and analy/.ed to correlate the particle si/.e and properties to tie signal
visibility.
A schom.it Ic drawing of a typical optical arrangement for particle sizinR inter-
feronietry is shown in Figure 3-a. An Arp,on-ion laser (blue, ' - O.&SSura) prcvidcs the
necessary coherent lip.ht source for tliis technique. The laser beam is split Into two
equal intensity beams with a beam splitter and focused with a lens to a crossover
point. Careful optical design is required to ensure that the beani crossover ;:nd the
foci of the beams coincide. At the crossover, the beams interfere to forra a station-
ary set of frinr.es (Figure 3-b) . The spacing; of these fringes can be determined very
accurately either ~y direct measure or from the optical geometry and the following
expression:
2 sin 0/2
where A is the wavelength of the incident beam and u is the bean intersection angle.
366
(2)
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Pa-ticie-LaCcn
Gas Flow
Samc*2 Volume
-PM.T.
Figure 3ซ. Sdwmatiol Diagram of the Pvticle Si* ing
lnttft*romct*f Optra
GAUSSIAN RADIAL
INTENSITY DISTRIBUTION
* - >
BRIGHT FRINGES
1/c2 RELATIVE BEAM
INTENSITY
sin (0/2)
Figun3-b. Entar^ad View ol ttw ftote Region
367
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t'.ir . i c i ซ-. ::.ovi -.; vi'h ' :.< fl'jw and ;,ass:ti;' rhr-.-j;':i '.h<- f'jca! :> ;:.-.<.- <: c a'. . e r
1
.,
r
j
'.". '
r i
el
la
: c.'i
.11 1.
ri;- (-
at l
-,t.
1 :.
1
VI
r )
;.:.a! .
V.-i r i a :
o he f
'.eLted !.y ' .'.<- ;>:.,.
'i(- i ' 'I 1 M(_(. '!
'.:i i n he lien' h of
r i :::ซ s;.ac in;- ( i e
A heuristic '>::>! anat ion
:.e r.oij'ii.v :' or. is a :"]:>-..
. a f;i:ii:t itin <;f '''u.-. w.-'.e
<:" vhv t he- vi ;:!<: 1 : v var
::-.jre '.. -.::.-
" . *
',:: r>:" i.t.- ;,.-ir '
o -J is * : t- :i-ir"
; c : t- .li ;H-
* : c 1 1-
1:: as !oi]o-*-:;. i'ai t i c 1 ซ-s rrj-u.-h st-.a 1 ! ซ-r ' fi.'in '.he fri:';;-!- Mia^in;-, tros-.i.-.;1 * he pro:/*- vul-
'.:~ie will sca-'er li/h' 'o 'rare n-i'. the sinusoidal fringe in'ir.-.i'y i- i s r i I,'.;; i :m :n-
<-ident MI.on it. l:icri-.-r.: :;; i ho p.ir'iclc- -,: .-.e with res;u-ct. to : . :'r:r.,c spacitij' vi !'.
result in 'he nar'ii'ie s; .it i er i r.;- iii-.h' fr';ni the fr:n;-e i'_ is en-<-r i r:;- -i-: -_-e 1 ! as i:e
or.e it is li-avin/.. 'i'.'n- vi ::* hi J i ' v . a-. siu:h. i s the- ra*io rif he ii-'n* sea'. * i.-f'. ;i v:!-.-:>
tin- |i.-iri icii- i:: i:t-nt cn-il over a Ijri;1::. fringe in when i'
! riiii-t- .'imi is ; iven l,y
re.i
vliere I is the jilui' ซ>c;it hfilt- current..
ซj! the multiple IM.-.IRI :;c.-it. I rrin;' |irnci.-s
useful in vi si.ial i .: int' the phennrien.i .
The tht.ir1/ !ป*-iiiiul tin- pheniin'.cti-i
.i::u;t r i i- in-h.iviur n! th
A! t Jj'i'Jrr. I tie r i j'/iro'js r.:it hcr:.'it ic.* I *!esc
:> i :. rather i_i>ra;ilex. 'iiis simple Je ;cr i
r i pt. :or.
:n i on is
/df
wlu-ri- Jj is the Ilessel's function of the first kind. In addition, it vas concluded
that in 'lie case of p.iraxial lii-.lu col led i on, the visibility is independent of tin-
Mil' scat, l.erin;' amplitudes, the particle shape, anil index of re! ra.; t iปn.
Subsequent 1 y, Robinson and Chu('") derived the functional re Lit ionsh i p usiniป, a
Pore rigorous approach hase'J en scal.ir diffraction (henry. biffrjctimi theory does
not. include index of refraction effects and Is valid unly for sc.it teri n;'. centers that.
are several diameters (-.renter than the w.ivelenป'.lh of lii-.lit.. The analysis cor.c luded
that the visibility is a function of the lir.ht collection aperture as woll as the par-
ticle diameter. With p.iraxial forvarc! scatter observations the thecrv vas in i-.ood
agreement with Farmer's resu.ts. More important is the fact that their experimental
data and Farrier's agreed very well with the theory usins; the above arsunpt ions. The
theory was also able, to predict experimental results for particle si::cs down to only
a few microns in diameter.
Evaluation of the possible effects of refractive index, were ~ade in a more re-
cent paper by Chi. anJ KobinsonO 1 hisinc, a di-tailcd analysis of scat.terin?, from par-
ticles in two cr-ssed coherent lii-ht beams. In this study the exact solutions of
Hflswell's equations (Hie cheery) were employed. For optics confirurations wherein
the scalar diffraction theory applied, the analysis was in ar.reer.ent with Krju.it ion 4.
The analysis showed reasonable agreement with the diffraction theory for scatterers
that aro absorptive. The scalar theory did, however, i;ive consistently hic.her values
of visibility than did the !lie calculations. For transparent spheres havin;; little or
no absorption significant oscillations of the visibility function iccurs for frinrc
spacing less than 10.; of
light passing through the sphere with the diffractive scattering. The lobes seen on
the scattering diagrams in Figure 1 arc in parr due to this phc-r.or.ป?non. For compari-
son. Figure 5 shows the Mic scattering diagrams for a particle with and without
363
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(b) V - U.-W
F190*ป 4. Otcillotcopt Tracm of S>rปH ''ซ" "ป Pvttcto d/mซ
lnt
-------
absorption. It is clear that the small lobes arc- due to inter ference with the li;;ht
transmitted thro'jr.h the particle. For Lire, or p.irticlc-s the diffractive scatter is
dominant about the near- forward direction and the main forward scatter lobe is sev-
eral orders of magnitude ;-reater than the off-axis secondary lobes. Thus, in the dual
beam scatlerinc, the interference between the forward lobe and the secondary lobes is
diminished. It is believed that the oscillation in the visibility function is tlut- to
the secondary lobe interference.
Robt-rds( ' ^' examined some further aspects of the desipn of the optical c,eometry.
In particular, the size of the collection aperture and bean stop, with respect to the
beam spacing at. the collection len:;. were shown to affect the visioility. This re-
sult is not surprising, in th.it the entire forward scattered lobes should be collected
if the result is to aj-.ree with the analysis. The dependence was shown to be minimized
with an adequately larc.e collection F/nuraber. His data also showed excellent agree-
ment with the theory.
From these observations and the experimental results, it is clear that the mea-
surement of visibility can dole-mine particle size reliably if the instrument is prop-
erly designed. Instruments based on this concent will produce the size and velocity
of the particle sinul laneous 1 y . We have developed the electronics for processing the
sic.nals to reduce the signals to particle sixe and velocity distributions in near-real
tine. Sophisticated noise discrimination loc.ic has been introduced to reject undesir-
able sic.nals and spurious noise. The p.resent systems can make measurements in parti-
cle field concentrations as hif.h as 105/cc(for s:r.a 1 1 particles) at rates as hii-h as
!0"*/sec anil in flows of very hii'h velocities ( 600 n/sec). A: present, we limit the
lower si/.e ranc.e of this instrument to approximately 'ซ_ra to safely satisfy the afore-
mentioned scattering criteria. However, ri-liable measurements r>i spherical water
droplets as larc.e as *> mr have been made. To extend our measurement capabilities down
I <> approximately O.Oum, scattering intensity rat.ioinc, has been integrated into our
system.
SCATTKKKIG INTENSITY KATIQl.'X:
This '.<.-ch:i t'l'Je. based on the collection of scattered lii'ht at two angles in the
near-forward direction (f-ic.ure ft) . also makes use of a hic.hly focused lic.ht source
allowing, for very hic.h spatial resolution. The .-hsoltitv scattering intensity docs
not enter into the size determinat ion and bซ-cause of the near- forward lic.ht collec-
tion. there is sufficient independence of particle shape and index if "-efraclion (for
absorbing part iclos) .
Figure 7 is a schematic diagram of (he optical c.come I r v required for this sys-
tem. An Arc.on-Ion laser is used to provide lic.ht of wavelength U . <145;ji?.. To obtain
the desired probe volume, beam expansion is used before the he. in is focused by the
transmitting lens to a spot of approximately 50um in diameter. The scattered lic.ht
Is detected by the collection lens and separated by a pair of annular masks into two
anr.lcs of scatter; 2.5* and 'j' In this case. Ratios of tin- sii-.nals from the photo-
multiplier tubes arc taken in the electronics processor for conversion to particle
size.
Reduction of the measured ratios requires the use of the polar scattering in-
tensity distributions computed from the Hie theorv. It is important to note that the
ratios of radiant flux. F. scattered by .1 particle of diamo'.i-r d is independent of
the incident intensity. That is the ratio
F(f'j) M,C"1(i.n> + i,('.,. i.r:)| sin rtj
FT77;! " H, <<-,.. ,n) t i2(-2...n)| sin ,
for two finite collection angles "j and "j. has ! cancel from the expression. This
is of course the reason for taking the ratios. The incident intensity on the particle
Is unknown as a result of intensity variation vltli trajcctorv through the focal volume
and benra attenuation. The plots of the normalized scattering intensity ratios versus
particle diameter for three pairs of collection angles are shown in Figure 8-a Com-
plex index of refraction eivcn by n 1.57 - 0.561 typical of soot was used in .he
calculation of these curves.
Dependence on the index of refraction can become troublesome for non-absorbing
spheres (e.g.. pure water droplctn with n - 1.33 - 01). For such tutorials, the
370
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Figure 6. Expanded Forward Scatter Lobe* Showing Collection
Anglei lot Intensity Ration^
Particle-Laden
Gas Flow
Sample Volume
/Annular Mask Pair
Beam E xpanaa" / Coiiimator ,
Zf Scatter
.2,5" /LJ
Scatter
PM.T.
L_J
Argon-ion Laser
Figure 7. Schซmซtic Oitgnm ol thซ Rttiotnq Optia Sytt*m
371
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intensity ratio is not nonotonic with p-jrticle si/.e but shows sone oscillation lead-
ing, to multivaluedness (Kif.urc H-b) . This effect is due to the interference of the
electromagnetic wave scattered by diffraction with the scattered light transmitted
through the sphere and hence, that is shifted in phase. Fortunately, the oscillations
are damped for absorbing particles. In the case of soot, n = 1.57 - 0.56i. or other
carbonaceous material, the relationship is a well-behaved nonotonic function of 'xe.
Another difficulty occurs when the instrument is used to measure in pop/dis-
perse particle fields. r'or example, with the 10"/5" .scattering anj^le combination.
particles i.,;i;ei than approximately 'J..m in diameter will show similar ratios to
smaller particle:; (figure 8-a) . This is due to the larger angle detector bei'inning
to collect light fron the secondary scattering lobe (r.ce figure I). Hirlesan^ J
suggested using a third collection angle to form two ratios as a means of overcoming
this problem. In our integrated system, the visibility processor will have a size
range overlap with the intensity ratioing system. If the ratioing sixer samples a
particle larger than its design range, the visibility processor will also sample that
particle .simultaneously and will cause the ratioing electronics to reject the sar.ple.
Hirleraan performed a thorough analysis of the intensity ratioing technique and
rr.ade estimates of the instrument accuracy. He quotes estimates as hij'.h as 507, uncer-
tainty for combust ion aerosols with very little absorption. The uncertainty is re-
duced significantly to approximately '/.Q7-. for absorbing particles. Our ratioing system
is expected to operate at a sample rate as high as IG'/scc, in particle concentra-
tions as high as lO'/cc.
Ml'LTII'LE I'ROBK {'ARTICLE DIAGNOSTICS KISTKliMKN'T
It is clear that an in situ particle sixing instrument will be required to op-
erate over a sixe range greater than one decade. The achievement of this is no triv-
ial matter since the scatter intensity goes approximately as the diameter to the
fourth power. For two decades of si;:e range, the scattering intensity would range
over roui-.hlv eight orders of magnitude on the forward lobe. Because of this intensity
range and the fact that, there are typically a larger number of particles in the small-
er end of the sine range, two probe volume si/.es are used.
Tlie Argon-Ion laser can produce both a blue (0.488;.m) and a green f.31'ป5..n)
wavelength of light s Imul l.-.neously. Figure <> is a schematic drawing of this systen.
The wavelength's .ire separated with a dispersion prism and directed throu/.h the respec-
tive optical elements previously described. Since the rat ioing technique is sensitive
to the smaller si;-.e range, that beam is expanded to achieve a focal volume that is an
order of magnitude smaller than the visibility probe volume. The incident intensity
(energy flux per unit area) is thus increased by 10' in the smaller probe. Particles
in the focal volume scatter light simil f aneous ly fron both wavelengths. The collect-
ing optics separate tile signals by wavelength and direct them to the appropriate de-
tectors .
Tills concept n. t only extends the si;:e ranf.e without any need for adjustments
by the operator hut also provides a check for the sizing ambiguity of the ratioini;
system. Our desii;n has been directed toward development of a working system that will
function in the harsh environments without requiring; attention by trained specialists
in optics or electronics.
The entire optics package is compactly mounted on a single chassis to ensure
instrument a 1 is;nraent . A sealed rupgedi.:ed enclosure Is used to prevent fouling of the
optics elements and to protect the components from the environment. Flexible connec-
tions between the instrument ,?nd the facility access windows are used to keep these
elements clean while allowing the probe to be traversed across the test field. In
very hlr.h temperature applications, an additional insulated housing can be provided
and outfitted with cooling by either air or water.
CO:;CUJSIONS
The concepts incorporated Into our particle si/ing instrument have been shown
to offer a reliable ceons of maklni- real t irae in situ particle size measurements.
Both particle sizing interfcrometry and scattering intensity ratiolng have been tested
in real measurement situations and have produced reliable data. Because of the rela-
tive complexities involved in these systems, great care is needed in their design and
applications. Our goal has been to develop a field-ready instrument that will produce
reliable results without requiring operators sophisticated in the physics. To realise
this goal, we have integrated our optics systems such that measurements ovซ;r a
372
-------
3 4 *i ft 7
Diamater m Microns
Figure 8 a. Seafaring Intensity Ratio Venus Particle Size
(or Rtpxsentativa Cclfcct on Angle Pairs
CC
1O
Ofe-
C.2-
nr1.57-0i
= 1.57-0.57.
4ฐ/2ฐ Scattering
O123456
Diamztcr tn Microns(M)
Figure 8-b. Comptnon of Absorptive end Non-ebsorptive
Scancring Ratio*
373
-------
\
'i
X
/ .
* .' .fi-'.s. >f r t-* ',* rj?r *;* r*'V i-' i. .*f . f j *J> ป'
64
Figure 9. Integrated Particle Siiinq Syitem Combminq P. rtide Sumq
Inlctftrometry and Scntei.nq Inttmity RซtK>n 9
realistic p.iriiclc' sixo r.iny.e can he nwide without .inv nan ipul.il ions by the- opc-r.:itor.
We h.'ivc found in p.isl cxperivict- that the requirement for such tvmi pul.it ions ii'.id to
instrument failure or unrol i.iblc djt.'i.
Relali*;:i data rates. The instruments al?o
have the capability of beinp interfaced to the power plant to warn the systcn, for
example, of excessive particle loading.
In the future, we plan to extend the nininum size defection ranr.c down to O.lyra
particles. A candidate method th.jt 'las produced accurate results is based on ralioinr,
the forward and backseattcred li;-,ht. Such -i technique could be incorporated into our
present system.
374
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REFERENCES
1. D. orvce-Sr.i'.h. i'hysics Bulletin. Vol. 25. 197'.. p. 173
?. J. D. Trolinf.cr and W. D. Bachalo. Spcctron Develonnent Labor.V-.nr ies . Ir.c..
"Particle field Diagnostic Systems for Hirh Tempera turo/Pre:. sure Kr.vironr-c-nts".
presented at "KPA/EKOA Sv^iposiuni on Hir,h Temperature/Pressure Paniculate Control".
Washington. D. C.. JO ar.J. Jl September 1977
'j. H. van cl<> Hulst. Light Scattering by Snail Part icles . John Wilev a.id Sons. New
York. 1957 "
4. M. Kerker. The Seal tor ing of Light and Other Elect ror^i^nct ic R.-nliat ion. Academic
Press. New York. 1V67
"3. V. M. rarr.er, "Sample Space for Particle Si/.e- and Velocity Measuring Interferom-
eters". Appi ied Opt ics . Vol. 15. No. 8. August 1976
6. '.. D. Hirlenan. "Optical Technique for Paniculate Character ix.at inn in Combustion
>.vironir.ents: The Multiple Ratio Single ['article Counter", 1'h.D. Thesis. Purdue
L'niversi t v. August 1977
7. H. Kn^elnilc. "ImaRinR of Optical Modes - Resona'ors wit!: Internal Lenses". Be 1 1
Systems Technical Journal. Vol. 44. No. 1-6. 1965
8. C. N. Davies. Aerosol Science. Academic Press. London. 1966
9. W. M. Fanr.er. "TI.e Interieronietric Obse:vation of Dynamic Particle Sir.e. Velocity.
and Number Density". !'h . .' Thesis. University of Tennessee. 1973
10. D. M. Robinson ar.d W. P. Chu. "Diffraction Analysis of Poppler Signal Character-
istics :"or a Cross-hearn Laser Uoppler Velocimeter" . Appl icd Opt ics , Vol. 14, No. 9,
September 1975
11. W. P. Chu and D. M. Robinson. "Scattering from a Movim- Spherical Particle by Two
Crossed Coherent PI ne Waves". Apj; ' ' _'_ Optics. Vol. 16. No. 3. March 1977
12. D. W. Roberds . "Particle Siting Usinp, Laser Inter ferometry". Appl ied Opt ics .
Vol. 16. No. 7. July 1977
375
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QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: Thank you, Dr. Bachalo. We have
a minute or two since we are ahead of schedule. Is there anyone who
would like to ask an immediate question? Would you identify yourself,
please?
MR. KWON: I am Henry Kwon from Dorr Oliver, Inc. Suppose two
particles, side by side, not quite put together, happen to enter that
crossover point. How would your instrument respond to that situation?
That is the first question.
The second one is what is the maximum and minimum range of
particle concentrations your instrument could be applied to?
DR. BACHALO: My answer to the first question is if two parti-
cles pass through the probe volume at the same time, one slightly
behind the other, they will undoubtedly produce a signal that is out
of phase. The electronics processor tests the periodicity of each
signal by comparing the period ratio measured over 5 cycles and 8
cycles to the ratio 5/8 to within a preset acceptable difference.
Aperiodic signals would be rejected.
The second question: If you have a long enough time to wait,
there really is no minimum particle rate. Both of the techniques
described are single-particle counters; they are sampling an indivi-
dual particle at a time. The upper limit is determined by the
requirement of having a high probability of having one particle in
the probe volume at any time. Knowing the size of the probe volumes,
you can determine the loading that will give a sufficiently high
probability of having one particle in the probe volume at one time.
If the number- of multiple particle occurrences is too great, the
distribution will be biased since multiple occurrences are rejected
by the electronics.
To work in flows of high concentration every effort must be
taken to reduce the measurement cross-section. The probe volumes can
be made as small as 10"' cubic centimeters for scattering intensity
ratioing and 10~4 cubic centimeters for the particle sizing inter-
ferometry. This would correspond to measurable concentrations of
106 and 103 particles per cubic centimeter, respectively.
FREDERICK HANZALEK, CHAIRMAN: Dr. Bachalo, I think you had at
least one.
376
-------
DR. BACHALO: Yes; I have two questions. One is *rom Dr. Janes
Wang. His question is: "Can you explain why your neasurenent
disagrees with Coulter counter results? Is there a theoretical
analysis?"
For those of you who don't know, we also had an instrument at
Argonne National Laboratories doing tests at a similar time that Dr.
Muly's instrument was there. Some of these results were presented
last niaht by Mr. Swift.
Our instrument measured particles that were somewhat larger
than the Coulter counter. There are two possible reasons for this.
The Coulter counter is obviously not an on-1 ine device. To make a
measurement, a sample of particle laden fluid must he extracted from
the flow. This sample is then placed into an electrolytic solution
and the impedance of the current flow through an orifice is measured
as particle* moving with the electrolytic solution through the
orifice obstruct the current flow. An electric volume (E.V.) is
measured and related to the actual particle size.
If particles measured in situ by our instrument were composed
of agglomerated sub-particles, these agglomerates could have been
redispersed in the electrolytic solution.
The other Tssihility can be related to a difficulty with
our electronic processor. The system was picking up noise and some
of the size distribution indicate that this was the case. Unexplain-
able himodal distributions occurred during the measurements. Trouble
with the processor was discovered later and corrected.
The other questions were from Jerry Whitfield from the United
Kingdom. The first question is: "What are the fringe spacings in
the green and blue probe volumes?" I may not have made myself
clear. There are two techniques that we have introduced into our
system. One makes use of particle sizing interferometry, in which we
spl. c the beam into two and interfere them; that causes the fringes.
That technique was used for the larger size range. We were
sensitive to sizes from about 5 to 40 microns for this particular
instrument, so the fringe spacing was set at about 45 microns.
In the green, we are using scattering intensity ratio. A single
beam is focused down to a very small point (approximately 80//m in
diameter) and we then look at the scattering at two angles. By
ratioing the two intensities we can get an indication of the particle
size, using the scattering diagrams to relate the collected scatter
intensity ratio to the particle size.
377
-------
Ti.e second question is: "Can you describe the influence of
turbulence in the particulate flow stream on the output signal and on
the estimate of particle size?"
You nay remember from the introduction that I have measured
transonic and supersonic flows with a laser velocimeter and thus have
addressed that difficulty. We have made accurate measurements in
sunersonic turbulent boundary layer separation wherein the turbulence
levels are very high.
If a turbulent cell passes through one beam and not the other
(or through the other later), a phase shift in the interference can
occur which will cause the fringes to move and can also change the
fringe spacing by changing the beam intersection angle. However, the
beams are entering the flow very close together and as long as the
turbulent cells are large enough the time in which this error might
occur is an extremely small percentage of the total time. This is
especially true in the low speed flow found in the fluidized bed
combustion facilities.
As to the size measurement, the effect of the slight variation
in fringe spacing will manifest itself as a proportionate error in
size measurement.
378
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A Particulate Sampling System
for Pressurized Fluidized Bed Combustors
William Masters and Robert Larkin
Acurex Corporation
ABSTRACT
A particulate sampler for high-temperature, high-pressure processes has been
developed and successfully demonstrated. The system uses an extractive approach, re--
moving samples from the process stream for complete analysis of particulate size dis-
tribution, morphology, and chemical composition. System capabilities have been demon-
strated by sampling a pressurized fluidized bed combustor. This paper describes the
extractive sampling approach, the HTHP sampler design, and the data obtained from
sampling operations.
INTRODUCTION
Advanced coal conversion processes present new problems in particulate sampling,
including severe environments beyond the capabilities of conventional equipment. This
paper describes a newly developed sampling system, specifically designed for the high
temperatures and pressures found in pressurized fluidized bed combustors. The system
uses an extractive sampling approach, withdrawing sampler, from the process stream for
complete analysis of particulate concentration, shape, size, and chemical composition.
The capabilities of the new system have been demonstrated in two phases of sampling
operations at a pilot-scale fluidized bed combustor owned and operated by Exxon Corpo-
ration. The first phase of testing was performed with the sample probe in its basic
configuration. The second test phase utilized modified probe internals designed to
investigate possible condensation of alkali metals. The system performed successfully
in a variety of operating modes, producing sample data from both test series.
Acurex Corporation has developed the HTHP sampler for the Industrial Environ-
mental Research Laboratory of the Environmental Protection Agency. The work is part of
a broad program investigating new sampling technology for advanced coal conversion
processes (Contrcct 68-02-2153). The EPA Project Officer for the contract is William
Kuykendal.
The following sections of this paper discuss the extractive sampling concept,
the HTII" sampler design, and sampling operations that have been performed with the new
system.
Extractive Sampling
In extractive sampling, a quantity of particle-laden product gas is drawn out of
the process for analysis. Once extracted, the sample can be thoroughly examined by
conventional methods. If proper care is taken to obtain and maintain a representative
sample, the extractive approach will provide complete, accurate information on proc^-?
constituents.
The sample is typically extracted through a probe inserted into the process
duct. The sample withdrawal rate at the probe nozzle must be matched to the duct velo-
city to avoid biasing particle size distribution measurements (isokinetic sampling).
The error in measured particle content as a function of an isokinetic velocity mismatch
can be estimated analytically (see Figure 1). For fine particles at low velocities the
error is negligible, but for larger particles or high velocities, serious errors re-
sult.
Sample temperature is also a consideration in the extractive approach. Ideally,
the temperature would be maintained at process conditions during particulate separation
and analysis. In practice, however, the sample is usually cooled to temperatures com-
patible with analysis equipment.
The major advantage of the extractive method is that the sample can be analyzed
by conventional techniques. For example, particulate removal and size classification
379
-------
2.0
1.8
O
O
ง1.6
c
O
111
5 12
Z
O 1.0
(C
O 0.8
8
O
u<
c 0.6
O
O
u.
O 0.4
I
0.2
k = 1.0
0.1 0.3 1 3
VELOCITY RATIO (R)
10
Figure 1. Probe Inlet Bias (from Reference 1)
380
-------
devices, trace element collectors, and chemical analysis techniques are all highly
developed (References 1 to 5). Extractive sampling is commonly used in emissions
measurement and combustion studies. . .
Access to the pressurized duct is the main difficulty in extending extractive
sampling technology to hign-pressure, high-temperature processes. The hardware re-
quirements for entering a pressurized process are much more complex than for ambient
pressure applications. The selected design Cor the HTHP sampler is described in the
following sc-ction.
HTHP SAMPLER DESIGN
The new sampler design adapts conventional sampling technology to high-
temperature, high-pressure environments. Key system components are:
A traversing sample probe that can be inserted or withdrawn during process
operation
A probe housing that contains process pressure during sampling
A cascade impactor to both collect and size particulate interchangeable
with a bulk filter (Phase I tests)
An in-stack scalping cyclone and backup filter followed by a "cold" final
filter (Phase II tests)
Conventional trace element collectors (organics trap and impingers)
Measurement and control instrumentation to assure isokinctic conditions
A schematic diagram of the Phase I sampler is shown in Figure 2. The samp'e prcje is
mounted within a pressure-containment housing. The probe can be inserted into the
stream through valves that connect the housing assembly to the process duct. Sample
flow in the probe pacse-j through a cooler and particulate collector (cascade impactor).
Flowrate is controlled by a throttle valve at the piobe exit. After leaving the probe,
sample gases are conducted through the trace element collectors, and vented.
The sampling system is shown in Figure 3. In addition to the probe and housing
assembly, controls, and sample collectors, the system includes a portable hydraulic
pump.
One of the basic decisions in designing the sampler was the choice between fixed
and translating probe configurations. A translating probe (one that is insertable and
removable during process operation) is more complex than a stationary probe, but offers
several operating advantages:
Particulate deposition losses in the probe can be recovered
Nozzles can be changed to maintain isokinetic conditions
The probe can traverse the duct to measure flow variations
Probe exposure to erosive/corrosive conditions is minimized
Inspection and maintenance are possible during process operations
Based on these advantages, the translating probe design v.-as selected for the HTHP sam-
pler. The sample probe and particle collector are shown in Figure 4.
Selecting the particulate collection temperature was a second major design deci-
sion. A number of well-characterized device^ are available for use below 500ฐF, but,
high-temperature particulate collectors are in an early stage of development. Based on
this practical limitation, a collection temperature of 450ฐF was selected with the
awareness that possible changes in particulate composition would have to be considered.
Major changes in composition are not likely above the sulfuric acid dewpoint. However,
changes in trace element concentration are a potential concern. The HTHP sampler has
381
-------
ft'
Flow
Process Duct
Access Valves
Enclosure
' (Pressure Boundary)
JKNV^NLy
Cooler' Impacto
Control Valves
Flexible Line
Heat Tracing
t ..n,
Flow Organic
ontrols Collector
1 1 .^.
Vent
Trace
Metals
Implngers
Figure 2. System Schematic
Probe Drive
Hydraulic Cylinder
Microswitches For
Transverse Control
Inner Tubular Housing
Dowtherm Coolant
Systems And Controls
i
Outer Tubular' Control Umbilical-7!
Housing ซ^
Dowtherm Coolant
Supply And Return
Sample Line
'-T-tff-r Control Valve
u'-r-:r And Operator
_ . - _ . Hydraulic
Control Console Supply System
Figure 3. High-Temperature. High-Pressure Sampling System
382
-------
Figure 4. Exploded View of HTHP Probe
383
-------
been used in an experiment investigating the effect of collection temperature on par-
ticulatc composition, as described in a later section of this paper. The impactor used
with the HTHP sampler is a Mark III, University of Washington Source Test Cascade Im-
pactor, Model D.
The cascade impactor has several advantages over other particle collectors. The
device provides many stages of size classification in a small volume. Also, impactors
classify particlo size based on incrtial and aerodynamic properties that relate di-
rectly to the performance of particulate removal devices. Impactor performance is well
characterized for moderate temperature, ambient pressure operation. The effect of high-
pressure on sizing performance can be estimated using theoretical correction factors.
For fine particles at moderate temperatures, even large pressure increases have little
effect on impactor performance. Collection temperature, however, can affect measure-
ments more significantly. Variations in cut size with temperatures have been calcu-
lated by the impactor supplier for temperatures up to 500ฐF.
In the selected sampler design, the sample probe enters the pressurized process
through 4-inch diameter full-opening valves while process pressure is contained by a
surrounding housing assembly. The housing, shown in Figure 5, consists of two tele-
scoping cylinders which move the probe into and out of the process. Hydraulic cylin-
ders ccnnect the two parts of the housing. Their function is to accurately position
the prooe, and withstand the large forces from process pressure. Scaling at the joint
between the housing cylinders is critical, so redundant seals are used. The tele-
scoping housing is the most complex part of the sampling system and consequently re-
quired the most design and development effort.
The HTHP sampling system also includes the instruments and controls necessary
for accurate sampling. Sample flowrate is one of the important parameters that is
monitored and controlled. Flow nust oe both isokinetic at the probe nozzle and within
the operating limits of the particle collector. For proper control, flow conditions
in both the process stream and sample probe must be measured. A pitot tube and thermo-
couple are mounted on the probe to measure process stream conditions, and a calibrated
orifice and thermocouple check the sample flow. The flowrate is adjusted to particle
collector requirements by a valve near the probe exit. Nozzle entrance velocity is
varied by selecting larger or smaller nozzles. The sampling system includes other con-
trols for sample temperature, probe traverse, trace element collector flow, and other
key operating parameters. System controls are housed in two portable enclosures, shown
in Figure 6.
The trace element collection equipment included in the sampling system consists
of an organics module and impinger train (see Figure 7). Both units are identical to
those used in the Source Assessment Sampling System that is commercially available from
Acurex. The organic module cools the sample gas to 70ฐF and traps organic vapors in a
potous polymer granular bed. The polyjier used in this test series is Rohm & Haas XAD-2
gas chromatographic packing material. The impinger train has four high-volume glass
impingers, three filled with oxidizing solutions and one with silica gel moisture
absorbant. The oxidizing reagents in the impingers collect volatile trace elements by
oxidative dissolution. The reagents are:
Impinger Solution
No. 1 6M H/Ox
No. 2 0.2M (NH4)2S2O8 + 0.02M AgNO3
No. 3 0.2M (N!l4>2S2^i + 0.02M AgNOj
No. 4 Silica gel
The peroxide solution in Impinger No. 1 collects reducing gase1" such as sulfur dioxide
which would lessen the oxidative capability of Impingers Nos. ซ nd 3. The ammonium
sulfate and silver nitrate solutions serve as the trace element collectors in the
impinger train (Reference 3).
PFBC Facility
The new HTHP sampler has been demonstrated in operations at the Exxon Miniplant
PFBC. The PFBC facility is described in this*section.
384
-------
Figure 5. Aarotherm HTHP Sampling Probe and Duct Interface Valve
38b
-------
Figure 6. Control Console!
386
-------
Impingers j/. -J'
Flow
Control
Ovsn
-^E^iP
Figure 7. Flow Control Oven and Gas Train
387
-------
The- V.iniplant is .-ป pilot-scale- pre-ssuri ze-d fluidize-d bod coir.bustor operated for
t h<- Kf'A by the Kxxon pes'-arch "m'i Kr.'fi no'.-r i n-( Conpany in I. indc-n, :.'e-w Je-rsey. The- I-FIiC
procr-KSt shown in Ki'jure H, i:'; a ceiir.bine-d-cycle- coal combustion process. Combustion
occur;; under pressure- in a lime-stone- bed that is fluidiz'-d by incoming air. Fluidiza-
t. ion ijivr-s -food mixinq for efficient combustion, and the- lime-stone- bod ronovos ri'jch of
the- :;ul!'ur role-.'ise-d durin-j the- combustion proce-r.s. Added useful e-ner'r/ con be produced
by e-xpandini] hi'|h-pres:;ure- flue- q.ise-:; in a 'jas turbine-, if particulato loadinq can bo
roduce-d to the- leve-ls (0.0002 t.o O.OOi -jr/scf) ro'iu i re-d to protect turbinr- bla-Je-s. The-
Kxxon Minipl.mt facility in be-in') use-'l to invest i'jat'-- fluirlize-d bod combustion, 'jas
cl'.-.inup devices, and part icul ate-
-------
Gas turbine
Boiler
R*generator
Figure 8. Pressurized Fluidized Bed Coal Combustor System
389
-------
Figure9. Minipiant PFBC
390
-------
Figure 10. HTHP Probe Assembly (rotated at Exxon Mimplant
391
-------
:OO
SO
60
PERCENTAGE UNDERSIZE (BY WEIGHT)
10 20 3C 40 50 60 70 80 90
98
J)
N
4C
10
8
UJ
5
UJ
d 2
10
08
06
0.4
02
0.1
90 80 70 60 SO 40 30
PERCENTAGE OVERSIZE
Figure 11. Partids Size Distribution
10
392
-------
i;l?>''v --vVf-V- '. '
-
., , '*' rf-f," ,'i " j>- i ; 'A'-""- .' ' ""' *>^'-*<. '*: .* " " *"-
<*!;'v ''ซ\. -* ''" " - '"*''' j- '
. . \\ ,-. , ->-;. , , . , .. -
.'ป;.-';--
. J.:. ' .
Figura 12. Imptctor Subflratw
393
-------
STAGE 5
FILTER
Figure 12. Concluded
394
-------
several plugged jets. The substrates from Run 2 are lightly loaded. Those from Run 3
show heavier, three-dimensional deposits.
Examples of particulate photomicrographs, showing particle size and shape, are
shown in Figures 13 and 14. These plots were made by a scanning electron microscope.
The irregular appearance is typical of flyash from lower temperature combustion pro-
cesses (Reference 8). The photos show the trend of decreasing physical size from
Stage 1 to Stage 6, although irr&gular shape and possible agglomeration make visual
interpretation of particle size very difficult.
The chemical composition of the collected particulate was analyzed by dispersive
X-ray fluorescence. Spectra of X-ray emissions from impactor Stage 1 and Stage 6 are
shown in Figure 15. The peaks in the spectra correspond to the number of emissions
detected at characteristic wavelengths of various elements. Results show the presence
of aluminum, silicon, sulfur, potassium, calcium, titanium, iron and copper.
Comparing the relative height of the peaks in two spectra can give a rough indi-
cation of the relative quantities of elements present in two samples. The comparison
of Stoge 1 and Stage 6 spectra shows no apparent difference in bulk composition between
the coarse particles collected (DSQ of about 30 microns) and the finer particles (050
of about 0.6 micron).
Data from the system demonstration tests are discussed more extensively in a
test report submitted to i-PA IERL (Reference 9).
Phase II Condensation Tests
Following the system demonstration tests, a second series of sampling operations
was conducted at the Exxon Hiniplant. The purpose of these tests was to investigate
the effect of sample cooling on measured particulate mass and composition. We were
specifically concerned that trace elements might condense between process temperature
and conventional particulate collection temperature (about 450''F) . Of particular
interest were the alkali metals, primarily sodium and potassium. For these tests, the
sampler was set up to collect particulate at process temperature, so trace element con-
densation could be investigated in two ways. First, the trace element content of
particulate collected at 450ฐF (from the demonstration test series) could bo compared
with the content of particulate collected at duct temperature to see if any significant
differences result. Second, after the particulate was removed at process temperature,
the sample gasses could be cooled and filtered to collect condensation products.
The probe configuration for the condensation tests is shown in Figure 16. A
scalping cyclone and a high-temperature filter are mounted on the front of the probe to
remove particulate at process conditions. After a series of chokod orifices, used to
gradually reduce sample gas pressure, a final filter removes condensed material as well
as breakthrough particulate. The cyclone is a Southern Research Institute model,
designed for much loss severe operating temperatures. This cyclone was readily avail-
able, was small enough for insertion into the duct, and had a very efficient C.6-micron
cut-point. During the tests, however, the cyclone's protective gold plating blistered
and fell off, leaving titanium surfaces exposed to heavy oxidation. Chemical analysis
of the particulate samples showed significant gold contamination in the cyclone and
front filters but none in the rear filter. Titanium contamination was not evident in
any of the samples.
The high-temperat.ure filter following the scalping cyclone is made of saffil
alumina, a material that Acurex is currently testing for high-temperature baahouse
filters. This material c-eems to offer excellent temperature resistance and effective
filtration, but its performance hasn't yet been fully characterized. .Its porfor-ar.co
.'.i the condensation tests was quite good, particularly with a two-filter "sandwich."
The estimated filter efficiency was well over 90 percent of the fine particulate passed
by the scalping cyclone.
The final filter at the sample probe exit is a standard Gel-!an "microquartz"
type with high efficiency and low trace element content. It is possible to use con-
ventional filter materials at this location because sample gas temperatur"> are sub-
stantially reduced by the probe cooler section and by sonic throttling in tho orifice
section.
395
-------
Figure 13. Particle Photomicrograph! Impactor Stage 1
396
-------
Figure 14. Particle Photomicrographs Impactor Stag* 6
397
-------
Figure 15. Particle Chemical Composition-Demonstration Tests
398
-------
CO
VO
Normal
Cooler
4 \ Process Flow
Condensation Test
Process Flow
Impactor
Sample Flow
(Or Filter) Throttling
Valve
Scalping
Cyclone
r"br
1350ฐF
Filter
HT
Choked
Orifices
Cooler I 1- I
i
H)0ฐF
Filter
I
3^7.
Sample
Flow
ปป
Figure 16. Probe Configuration
-------
Four sampling runs were completed in the condensation test series. Conclusions
have fcoen drawn on the available data regarding condensation of trace notals, parti-
cularly the more common alkali metals, sodium and potassium.
The original intention of the Phase II tests was to compare- the elemental con-
centrations found in the filtered natcrial of the in-stack scalping cyclone and backup
filter with the material caught on the rear filter. If the oysten worked ideally, one
could assume all the material on the rear filter was in a gaseous state at stream con-
ditions and would thereby pass through the hot cyclone filter combination and thus be
indicative of condensation products produced within the probe. Unfortunately, the rear
(or cold) filter could not be reliably analyzed. Carryover of glass fiber filter
material during sample preparation, the snail amount of the sample available on the
cold filter and additional contaminations, yielded poor detection limits with the spark
source mass spectrometer (SSMS) . .Moreover, the values reported ฃor tnia iiltcr W(Tซ> '' n
mass units rather than concentration units because a r.et filter collection weight was
not determined. This made any direct comparisons of cold and hot catches from Phase II
results alone difficult.
Samples of the probe wash were analyzed by Arthur D. Little, Inc., to determine
the source of the suspected contamination of the rear filter. A sample from each test
was analyzed by thermal gravimetric analysis (TGA), infrared analysis (IR), X-ray
fluorescence (XRF) , and low resolution mass spectra (LH.XS) . Results indicated there
were three sources of contanination. They were: (1) approximately 30 percent particu-
late -- attributed to hot, in-stack filter ijreakthrouch, (2) approximately 25 percent
sulfuric acid and sulfatc condensate attributed to localized cold spots (measured
at 200ฐF) below the HjSC^ condensation temperature, arxi (3) approximately 40 percent
organics attributed to the disintegration of packir.g material from a valve located
just upstream of the roar filter. The evidence of breakthrough particulate contami-
nation was supported by photomicrographs, which showed similar appearance between
material on the front and roar filters, and by dispersive fluorescent X-ray spectrum
which shows a similarity in chemical composition. The sulfur content did, however,
increase on the rear filter, apparently as a result of sulfate condensation.
An alternate approach was taken to resolve the question of trace netal conden-
sablcs. During the Phase I demonstration tests at Exxon, a test run was made at essen-
tially the same operating and scream conditions as the Phase II tests. A total mass
filter was used in place of the cascade inuactor and was maintained near 400ฐF. North-
rop Services, Inc. performed a SSMS analysis of the bulk filter catch, as they did with
the Phase II cyclone and backup filter supples. Table i presents the results for these
analyses. The measured elcr.ental content is similar to common flyash. A partial,
nondimensionalized comparison of these two sets of results. Phase I and Phase II, is
presented in Table II. Reference quantities Fe and Kc, have been chosen to nondimen-
sionalizc the results because they exist in significant concentrations in each sample
and are not likely to be present as a result of contamination. ''ondimensionalizir.q
was done because there appeared to be a diluent in the Phase I filter catch. The Si
concentrations indicate that the filter material itself nay be the diluent in the
samolc. In fact, the sample analyst acknowledged some difficulty in separating the
sample from the filtering media.
Aside from the Si results, comparison of the other quantities shown indicates
that there was no detectable change in the concentration of Ha or K fros the hot to
cold particulate catches. This limited data would indicate that particulate collection
at low-pressure and 400ฐF yields accurate results for particu'.ates.
CONCLUSIONS
The sampling system described in this paper der^..strates that extractive sam-
pling is a feasible approach for sampling high-temperature, high-pressure processes.
Furthermore, the Phase II condensation test data indicates that sample filtration at
reduced pressure and low temperature (400ฐF) yields accurate results for particulates.
Technology for sampling pressurized fluidized bed combustors is now developed and
available. Future development also will be required, however, to make useful appli-
cation of this technology and extend it to other advanced coal conversion processes.
One of the remaining issues for FBC high-temperature, high-pressure sampling,
is system cost/performance trade-offs. Process developers seem to be interested in
both upgraded and downgraded versions of the sampling system. Upgraded versions offer
400
-------
Table I. Concentration of Elements in Flyash SSMS Analysis (Partial)
Element
K
Ila
Rb
Cs
Al
Si
Fe
Ca
Mg
Ti
Sr
Ba
Au
P
Cu
Zr
Ni
Cr
Pb
Cyclone
(ppm)
8,200
1,310
<70
6.7
164,000
94,000
30,000
20,000
11,400
2,430
810
710
650
276
248
160
120
<90
85
Test No.
Front
(ppm)
8,850
2,500
:68
0.23
94,000
82,600
13,400
19,000
17,800
1,950
555
694
118
223
165
140
100
<140
75
3 Phase II
Rear
(ppm)
15.1
<135.0 *
<0.43*
<0.62
64.2
<2510.0
60.0
<6.6 *
38.0
7.1
0.6
<5.0 *
<1.4
<224.0 *
<1.5 *
<2.1 *
5.8
<13.1
<4.0 *
Rear Blank
(ppm)
5.4
91.0
0.13
--
6.4
1210
2.36
7.3
5.9
3.1
1.15
1.5
68.0
0.91
0.64
0.29
6.4
1.2
Phase I
Bulk Filter
(ppm)
16,260
3,560
866
8.8
Major
310, OO"1
36,300
44,700
28,600
10,600
1,320
1,080
13
1,880
142
334
348
366
86
Notes: < - Natural background limited detection limit.
<* - Blank limited detection limit.
longer sampling durations, quicker turnaround and better operating convenience. Down-
graded versions, such as fixed-probe designs, arc cheaper, but give less information.
gasi
The next objective for extractive sc.-npling is to develop technology for coal
ififjrs. Particulatc measurement is also important for developing these processes,
^ ซ-i i 1. * *_ L *? IC1lt.lv.uJU\.L IIH_ M ^ u I *_ lllrjll l_ JO U J ฃ31^ Itlll-'Ul.l.ClIll. i UI tit: VU i Ul.- i il
and environmental difficulties are even more sovere than for FPBC's
REFERENCFK
1. Lundgrcn and Calvert, "Aerosol Sampling with a Side Per: Probe," Amor. Ird. Hvo
Ass. J. , 28:213 (1967) . ' '.
I. Calvert and Parker, "Collection Mechanisms at High Temperature and Pressure,"
Symposium on Particulato Control in Energy Process, EPA-600/7-76-010, Sept. 1976.
3. Homersma, et al., 1ERL-RTP Procedures Manual: Level 1 Environmental Assessment,
EPA-600/2-76-106a, June 1976.
4. Blake, D. E. , Operating ar.d Service Manual -- Source Assessment Sampling Svsten,
Aerothorm Report UM-77-80, March 1977. ~
5. Gooding, C. 11., Wind Tunnel Evaluation of Particle Sizing Instruments, EPA-600/2-
76-073, March 1976T
6. Operation Manual, Mark III University of Washington Source Test Cascade Inpactor
(Model D), Pollution Control Systems Corporation, Renten Washington, March 1974.
7. Hoke, R. C., "FBC Particulatc Control Practice and Future Needs: Exxon V.iniplant,"
Symposium on Particulatc Control in Energy Processes, EPA-600/7-7G-010, Sept. 1976.
8. Hoke, R. C., Exxon Research and Engineering Company, Linden, New ? rsey, Personal
Communication.
9. Masters, W. Z., "Field Testing of a Sampling System for High-Tenperature/High-
Pressure Processes," Annual Report, Measurements of High-Temperaturc/High-i'rossure
Processes, Aerotherm Report TR-77-55, July 1977.
401
-------
II. Partial Comparison of Front and Rear Particulate Catches
from Exxon Test Series I and II
K
Fc
Na
Fe
Si
Fc
Ca
Fe
Mq
Fe
Sr
Fc
Ba
Fe
K
Mg
Ha
Mg
Si
Mg
Fe
Mg
Ca
Mg
Sr
Mg
Ba
M?
Phase I Tests
Bulk Filter
0.45
0.10
8.54
1.23
0.79
0.04
0.03
0.57
0.12
10.84
1.27
1.56
0.05
0.04
Cyclone
0.26
0.04
3.13
0.67
0.38
0.03
0.02
0.72
0.11
8.25
2.63
1.75
0.07
0.06
Phase II Tests
Front Filter
3.66
0.19
6.16
1.42
1.33
0.04
0.05
0.50
0.14
4.64
0.75
1.07
0.03
0.04
Avg.
0.47
0.12
4.65
1.05
0.86
0.04
0.04
0.61
0.13
6.45
1.69
1.41
0.05
0.05
Ratio
0.96
0.83
1.84
1.17
0.92
1.00
0.75
0.93
0.92
1.68
0.75
1.11
1.00
0.80
402
-------
Mathematical Modelling
403
-------
INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN Our next paper was originally
scheduled for presentation at 10:40. As I mentioned before, we are
interchanging those. So, we are going to go now to the paper on FBC
modeling and data base. This paper was authored by Messrs. Tung,
Goldman, and Louis of Massachusetts Institute of Technology.
The paper will he presented by Dr. Tung, who is a graduate of
MIT. After work at Continental Oil Corporation on natural gas
liquefaction, and where he also set up a catalysis and catalytic
process research laboratory, he joined the MIT energy laboratory
approximately three years ago. His work there has been primarily in
the areas of fluidized bed combustion, energy-to-energy conversion,
fuel-to-fuel conversion, and catalysis. Dr. Tung.
405
-------
-------
FBC-Modelling and Data Base
Shao E. Tung. Jay Goldman,
and Jean F. Louis
Massachusetts Institute of Technology
I. INTRODUCTION
One of the most attractive, near-term advanced technologies for the environment-
ally compatible use of coal is fluidized bed combustion. For that reason, significant
commitment has Laen made by DOE to bring this technology to commercial application. As
a part of this overall effort, MIT .has organized a broad-based multi-disciplinary group
working toward the establishment of a systen model for the technology. The overall
objectives of the project are:
(1) To establish a comprehensive system model that is sufficiently
precise to be suitable for process and engineering design
optimization.
(2) To establish a data base system that will reposit all relevant data
on coal based FBC (fluidized bed combustor) in a single site and
can be used to answer various queries from remote sites.
It is proposed to achieve these broad objectives in several phases. The initial
phase, which is currently being conducted, has the following more limited qoals:
Modelling: To develop an initial system model using mainly the
state-of-the-art information
Data Base System: To select and implement a data base management system
suitable- for FBC.
II. FIRST GENERATION SYSTF.M MODEL DEVELOPMENT
In this initial phase of nxxlelling, a portion of effort has been devoted to
develop adequat.; pro.-ess representation. In this regard, a major finding of our study
is that the behtvior of an FRC in which relatively large particles (e.g., 1mm) are being
fluidized differ; significantly from the behavior of a more traditional fluidized bed
reactor such as i-n FB catalytic cracking un> c where relatively small particles
(e.g., 60u) aie being fluidized. Furthermore, the differences seem to permeate through
all phases of bed behavior'^'^).
When uncertainties regarding process representation and our commitment to put
togother a system model utilizing the available information in the best way possible
were considered together, we decided to proceed with the structure of an initial model
comprised of four component models. Four existing models were selected.
Existing Models Selected
Component Models for Initial Application
fluid dynamics fast bubble model
combustion modified Davidson Model<3>
heat transfer (horizontal tubes) modified Vreedenberg Model!'')
d^sulfurization Borgwardt Model(5), (6), (7)
Along with the Ken and Leva model for vertical tube heat transfer, the Merrick
and Highley'8! models for attrition and elutriation are also being used.
Not all existing models can be used directly. Some need considerable modifi-
cations while some others need incremental development19ป10ปHซ12) . it is also
apparent that all the submodels adopted for the system model use must be compatible and
consistent so that thi;y may be linked together for continuous operation. When the
modifications and link-ups are properly executed, the system raodel has demonstrated to
be capable of approximately computing the following grous bed parameters:
406
-------
o Carbon Loading
o Combustion Efficiency
o Carbon Particle Size Distribution
o SC>2 Emissions
o Stone Utiliz-tion
o Heat Transfer Coefficients of
Horizontal/Vertical Immersed Tubes
o Bed Temperature
o NOx Emission
A sketch of this initial model is shown ir. Figure 1 . ""ป facili - c- the use of
this model, a standard set of values concerning bed proportit operating conditions,
acceptor properties, coal properties, cyclone properties are p.ovidea. This will allow
the users to compute any particular output parameters of interest to him j.n relation to
any input parameter. A segment of the computer output is shown in Appendix I. In this
segment, the combustion efficiency is computed to be 99.lt. for a superficial velocity
(U0) of 1.5 meters/sec, operating under the conditions set forth by the standard values
furnished. Kith a number of such calculations, a sensitivity analysis can be carried
out to investigate the influence of Uo on the combustion efficiency. Other sensitivity
analyses can be carried out in a similar manner.
In one respect, our modelling experience has been particularly encouraging.
After having recognized that the fast bubble is perhaps inadequate to represent the
fluid dynamic behavior of an entire bed, we have discovered an important need to modu-
larize our component models so that a capability ey'.sts to replace each unit individually
at a later time. Yet, due to the complex interactions between the various component
models, we were previously uncertain as to whether modularization would DC successful.
It now appears that the four selected component models are linked together only by
relatively weak interactions. Hence, they can be necoupled.and to that extent modular-
ization is feasible.
Despite the fact that the computed output parameters appear to be reasonable in
magnitude and their trends of variation witn relevant input parameters also appear
reasonable, the simplistic nature of this initial model must be fully recognized. As
such, the model is more suitable to serve as a framework for further refinement, rather
than as a functional model. The present model is still useful for parametric studies.
Nevertheless, introduction of better physical description developed as the key feature
of the process w: 11 upgrade the present model to a powerful predictive tool.
III. FIRST ROUND SYSTEM MODEL UPGRADING
Even though our essential commitment during this first phase is to the application
of state-of-the-art models, our effort has not been bound by the formality of this
commitment. Since the beginning, a part of our effort has been devoted to the develop-
ment of twenty key feature models. These key feature models, together with nine
additional key feature models of which the development will be initiated soon, are
listed in Table 1.
All of these key feature models are beyond state-of-the-art models, and their develop-
ment has provided a broad base for the upgrading of our system model.
In many key areas, our analytical inquiries have already improved our under-
standing in the process representation a great deal. These newly acquired insights,
which are being targeted for our first round model improvement, are briefly discussed
below.
(1) Slow Bubble Region (Fluid Dynamics)
It was recognized at the outset that the fast bubble model probably does not
adequately represent the fluid dynamics of the entire bed. The oxygen profile data
obtained by Pereira and Beฃrd3)at Shefield suggest that a large fraction of combustion
occurs in a region near the distributor plate. In this region, bubbles are either small
or not as yet formed. If fast bubble is assumed for this section, the computed oxygen
delivery to the carbon surface will be unduly small because, for fast bubbles, there is
a significant diffusional resistance for the oxygen to transfer from the bubble phase to
the emulsion phase. Such resistance is extremely small for the slow bubbles ard r.cn-
existant prior to the bubble formation. Hence, the fast bubble model tends to under-
estimate combustion efficiency.
407
-------
-
8
User's Input
(e.g. U0>
User's Input
(e.g. Combus-
tion efficiency)
Fluids Dynamics
Fast Bubble Model
Standard Set of Input
Temperature
Combustion
Modified Davidson
Model
Desulfurization
Borgwardt
Model
Material Balances
+
Concentration
Profile
Heat Transfer
Modified
Vreedenberg's
Model
Physical
Thormodynamic
Properties
Output
Figure 1. First Generation System Model
-------
Table I. Key Feature Models
o
VO
lindcr Development
(1) Coal Dcvolatilization (Combustion)
(2) Char Combustion Kinetics (Combustion)
(3) CO Burnout (Combustion)
(4) NOX Emission (Combustion)
(5) Flow Regimes (Fluid Dynamics)
(6) Slow Bubble Model (Fluid Dynamics)
(7) Bubble Size (Fluid Dynamics)
(8) Mass Transfer Coefficient
(Fluid Dynamics)
(9) Mass Heat Transfer to Single
Particle (Fluid Dynamics)
(10) Pore Plugging (Desulfurization)
(11) Expanding Grain (Desulfurization)
(12) Horizontal Tube Correlation
(Heat Transfer)
(13) Influence of Large Particle in
Heat Transfer (Heat Transfer)
(14) Honuniform Heat Flow to Tubes
(Heat Transfer)
(15) Force on Tubes (Heat Transfer
and Fluid Dynamics)
(10) Thermal and Mechanical Stresses (Ma.nrials)
(17) Creep Deforraantion and Creep Rupture
(Materials)
(13) Fatigue Mode of Failure (Materials)
(la) Corrosion Mode of Failure (Materials)
(20) Transient Behavior
To Bo Initiated
(21) Flow behavior with/with<>ut Tubes
(Fluid Dynamics)
(22) Coal Combustion (Combustion)
(23) CO Emission and Reducing Zone (Combustion)
(24) NO Emission Refinement (Combustion)
(25) Desulfurization Under Varying SO,
Concentration Zone Environment (Desulfurization)
(26) Local Temperature Profiles and Hot Spots
(Combustion and Heat Transfer)
(2V) (Bed Bcnavior of Free Board (Fluid Dynamics)
(2d) Acceptor Calcination (Desulfurization)
(29) u.iat Transfer to Tubes and Bounding Wall
(Heat Transfer)
-------
Significant improvement can be achieved if the bed is divided into two sections:
a pre-bubble/sJow bubble region (where U[jr < t!mfA ^f) followed by a fast bubble region
(where Ugr > Umf/emf). The transition occurs when UBr = Umf/--mf. However, if this
computed transition point is less than 5cm from the distributor plate, we will use 5cm
as the transition height because below that height bubble is assumed to be not yet
formed,and in that pre-bubble region, the bed behaves closely to a slow hubble bed with
respect to oxygen delivery to the carbon particle. Oxygen delivery is important to
combustion efficiency calculation. Such a two-region bed model has already been
implemented, and the improvement achieved in the combustion efficiency prediction is
given in Figure 2.
(2) Devolatilization
Devolatilization and combustion of volatile matter from coal are important
because they exert influence on the oxygen and temperature profiles of the bed and on
the formation of hot spots. The stability range of FB operation is also extended to
lower te-.peratures because of the volatile matter's lower ignition temperature.
The traditional approach to FBC modeling has been either to ignore the volatiles
altogether or to assume that the devolati1ization is instantaneous. Our analysis shows
that devolatilization is a time-dependent process<2ป9ป14>. For a 1mm particle with 2%
oxygen in the surrounding gas and 1173ฐK (1650ฐF), Figure 3 indicates that the
devolatilization is complete after 2 seconds, while the final burning time is 308
seconds. Thus, the devolatilization time is commensurate with the mixing times for
particles in an FD which is typically on the order of 1 second.
This analytical finding is important when feeding from above the bed is con-
sidered. It also implies that both NOX and S02 are more probably being released all
over the bed rather than all at once near the feed point.
(3) Combustion Kinetics (Combustion)
Whether the combustion in an FBC is a diffusion controlled process depends both
on particle size of carbon as well as on the temperature of combustion. The combustion
may be mainly diffusion controlled if temperature is high and the particle size is
large; but will be mainly kinetically controlled if the conditions are reversed. For
1mm diameter carbon particle at 1173ฐK (1650ฐF), the diffusion and the chemical
kinetics roucjhly offer equal nsistance to combustion. In Figure 4, the oxygen con-
centration at the surface over the oxygen concentration at infinity at the time of 50%
burnout in taken as the average fractional kinetic resistance. If the chemical kinetics
are not taken into account, the burning rate for small particles such as elutriated
carbon particles could be too high, and the combustion efficiency will be overestimated.
ft simple version of chemical kinetics (not taking pore diffusion into account)
has already been incorporated into the system model.
(4) KOy and CO emissions
A NOX generation model for a two-region bed has been developed assuming that the
evolution and burning of coal volatiles take place uniformly throughout fie bed and the
NOX produced can be reduced to N2, either by char or by any reducing gases. The model
has already been incorporated into the system model.
A CO burnout model is being developed. Our preliminary results show that there
is a kinetic limitation on CO burnout below 1050ฐK (1430ฐF). However, if the CO
concentration leaving the bed is higher than 1%, the heat release due to combustion may
be sufficient to overcome this limitation. Nevertheless, if the cooling rate is in
excess of a few thousand degrees per second, CO may quench regardless of it-j initial
values. The average experimental evidence supports the kinetic limitation, but does
not show t'.e same extent of burnout. This may be due to, among other things, the overly
simplistic free-board fluid dynamics model employee'.
(5) Pore Plugging and Stone Evaluation
Figure 5 shows our model calculations on the variation of SO2 emissions as a
function of the calcium-sulfur ratio. The lowest curve shows that same variation using
an empirical model developed from the available data sets. It appears that the model
410
-------
100
90
G 80
'&
70
60
SO
Superficial Velocity - ?.5 m/s (7.5 (t'jecl
eซceoAir-20%
Bed Temperature - 1200 K 1170O Fl
800 900 1000 1100 1200
Figure 2. Effect of Two Region Bed on Carbon Efficiency
i.o
c ฃ
5 u-
\
Volatile mjtttr
*nd rniduAf
carbon burr*
T. i
Particle diamrtvf 1m.r>
V Amount of Volatile Released
Vo Total Amount of Volatile*
W Amourn of Total Cartxm Burned
WQ * Total Amount of Carbon
* y Volatile matten
bumai
detactied vlame
t.0
10
Time (
-------
1.0
I .
t\>
ISO
1000
3000
100
90
70
60
SO
40
System Model Stone No 6
Temperature- tt!6e
Tempers .re- 11I6'K I1S50ฐF>
Exploded Bed Height 1M (3 ft.)
FluiUi/ing velocity Uf *
233M<ซclMl'wc.l
1.5
2.0
C./S
2.5
3.0
3.5
Particle Diameter (ป ml
ซ 5 stone is a medium grained, granular arid porous, gray reef type of
dolomite. ซ6 Hone contains 81% dolomites some clay and fine ouari;
silica impurities, medium grained, gramular and microporous.
Figure 4. Dependence of Kinetic Resistance to Combustion on
Particle Size With Ignition of Volatile!.
Figure 5. % Sulfur Removal vs. Ca/S Rafio
-------
calculation yields higher SC>2 removal. This might be attributed to:
o The Borgwardt model does not account for the pore plugging
due to solid volume expansion when CaO is converted to CaSC>4.
o The Borgwardt model applies only to the precalcined stone,
while the empirical model applies to uncalcined stone.
This figure illustrates the importance of incorporating both pore plugging and stone
colcii.ation features into the system model.
Two models, namely a "pore plugging model" and an "expanding grain model" are
being simultaneously developed. Analytical solutions have been attained for both
cases. Figure 6 illustrates the variation of percent conversion versus initial porosity
according to the expanding grain model both for calcine and for dolomite.
The maximum conversion values in Figure 6 are obtained with the assumption that
the radius of the grains (which are assumed uniform) are the sanie for all grains in the
pellet. This is equivalent to assuming uniform reacting gas concentration inside the
pellet; i.e., no pore diffusion. In actuality, pore diffusion is operative and pores
are only plugged at the mouth. This will result in a conversion lower than the maximum
level predicted. Hence, it is not surprising that the experimental data in Figure 6
are less than the model predicted values. Improvement of the model can be attained if
a size distribution is assumed for the grains.
The development of a stone calcination model will be initiated soon.
(6) Heat Transfer Correlations (Heat Transfer)
Heat transfer coefficient data from Pope, Evans and Robbins'16', General
Electric(17)( an<} McLaren and Williamsd2) have been compared with the modified
Vreedenberg correlation used in our system model. The data are not in good agreement
(see Example B of DBMS Applications) with model predictions. It should be noted that
most of these newly acquired data are from beds with high superficial velocities
operating in slugging or turbulent regimes, whereas the original correlation was based
on freely bubbling beds. Additionally, each of the present sets of data were insuf-
ficient in some way for complete comparison with the correlation. Some information is
assigned to fill the information gaps, thus leave considerable possibility of error.
Nevertheless, the comparison appears to suggest that a high velocity fluidized bed
operating in the slugging or turbulent regimes may require a different correlation.
A new correlation comprising a gas conduction model and a parallel convective
heat transfer model has been developed by our heat transfer group. As is shown
Figure 7, agreement with the experimental data is good and appears to have the
potential of developing into a physical correlation for large particle FBC heat
transfer.
(7) Thermal Stresses (Materials - Heat Transfer)
We previously reportedH' that we would have yielding with the 304 stainless
steel tubes immersed in a fluidizcd bed, and would consequently face a low cycle
fatigue problem due to startup-shutdown eye-lie operations.
Subsequently, thermal stresses induced by two-phase flow (water and steam)
within a horizontal tube have been analyzed. From the temperature profiles
calculated 111), it appears that signficant thermal stress might set up in the inside
wall of the tube.
IV. FBC DATA BASE MANAGEMENT SYSTEM (DBMS) DEVELOPMENT
Data should be regarded as vital resources which must be organized to maximize
their value. This is especially true for KBC where tho rate of technical development
is so high that the data file will soon grow at a large rate. Furthermore, it is
believed that the value of data is enhanced when data are assembled together in a data
base setting.
413
-------
I
100
80
0.511
0.676
40
20
0.0 0.1
60
40
20
0.2
0.5
O.G
0.3 0.4
Initial Porosity
Figure 6. Effect of Initial Porositv on Conversion
as Predicted by Expanding Grain Model
(Points Are Experimental Data)
0.7
BCE
DGE
^Mclaren 8, Williams
ฃ Pef
A Bartel & Genetti
CNYL1 HSpsia
9 NYU 40 psia
ONYU ISpsia
9 NYU 75 psia
I l i
3175vm
10
20
40
Re
100 200
Oo
-------
In this initial phase, a portion of our effort has been devoted to the selection
of a JEMS. Our early study suggested that the selection of a correct data uase system
would be a critical step in our overall task of establishing a DBMS, due to the fact
that most of the available data base systems were developed for administrative purposes.
A large body of available well-used systems were found to be unsuitable for our appli-
cation because of constraints in data type, limitations in physical storage structure,
and/or lack of broad capability in user-system interactions, etc.
Our selection criteria were:
1. flexible data base structure
2. concurrent use capaoility
3. on-line query capability
4. data accessible from Fortran programs
5. well developed and well supported
Starting with nine commercial systems, Computer Corporation of America's
Model 204& was finlly selected. A major advantage of Model 204-- over the other
s/tems investigated is its flexibility. Any number of files with almost any structure
(hierarchy, network or even relational) can bo used with equal ease and efficiency.
While Model 204's creation procedure is fairly involved, fine tuning of the data base
in relation to its user's pattern is better facilitated, with a resultant net savings
in cost and increase in performance. Although some other data base systems nay have
slightly simpler (i.e., non-procedural) user's language, our experience has been that
these languages are less able to cope with the kind of users we feel the FUC data base
must support. On the other hand. Model 204'K uses a semi-procedural language whici. can
support airaost any kind of processing ^hile still remaining relatively easy to use.
Thus, ou- conclusion is that Model 204'K best fits the needs of the FBC appl icatioi'.
As soon as we realized that the selection of a FU'table data base system would
be prolonged, we decjded to establish a prototype data base system based upon an MIT
developed KDMS (relational data management system). The establishment of a prototype
system has allowed other tasks such as data input, file design, system-user interface
development to proceed
-------
L'xample A. Getting tho Relevant Information
A user wants to know soxc information that rcicht be of interest to hi- in the
area of Ucsm f urization. 'ihe information is not specified at the beginning: he may
proceed to use DBMS in the following nannor:
First, he may ask DBMS to print out what papers are available in the data
base by typimj in three input lines (underscored):
"cc list_area"
"dcsulfurization"
"yes"
The HUMS will produce Table A-l.
After examininq Table A-l, the user decides that he wants to examine or.e paper,
i.e., "Isothermal Reactivity of Selected Calcined Limestone with SO,", more closely.
He can type in the following two statements:
ec list_figures
name of tne paper - "Isothermal Reactivity of Selected Calcined
Limestone with SO2"
The DBMS then produces Table A-2.
If the user now wants to know more about Figure 1, he can type in the following
3 lines of statements:
oc tab_figure
name of the paper - "Isothermal Reactivity of Selected Calcined
Limestone with SO,"
Figure I
The OHMS will produce Figure 1 in the tabular form. . (Table A-3).
If he wants tho figure plotted, he can type in
ec gr.tuh figui^
name of the paper - "Isothernal Reactivity of Selected Calcined
Limestone with SO2"
Figure 1
Other relevant information DBMS r.ocds
The DBMS will tnen present the plot. (Table A-4).
L'xample b. Check Modi tied Vrccdonborg Correlation with GE Data, ftR Data, etc.
Tnis example shows the utility of a general purpose macro facility which can
either plot a graph or print a list of values of the left hand side and the right hand
side of a correlation. To use the macro, type
"cc check_correlation"
As is shown in Table B-l, the macro wil'. then ask cq_vbl = variable: e.<:..
Uf = fluidizing velocity. The macro will report the above query until you type a
pcriout.).
The macro will next ask for the left hand side and the right hand side of the
correlation. After obtaining the answers, the macro will compute the left and the
right sides for any set of data and plot out as the left side vs. the right sice.
The macro also allows the user to obtain a list of data sources and the user can
restrict the set of data points to those of a particular paper or even to a particular
table/figure, if desired.
416
-------
Table A-l
ec list_arca
For which area? desulphurization
include paper titles with author names? yes
Listing of papers for area desulphurization:
TITLE
AUTHOR (S)
Charachterization and Control of Caseous Emission from Coal Fired Flui
\cdized Bad Boilers
Robison E. B.
Fluidized Bed Combustion: A status report
Jonko A. A.
Isothermal Reactivity of Selected Calcined Limestones with SO2
Borgwardt, R. H.
Kinetics of the Reaction of SO2 with Calcined .Limestone
Borgwacdt, R. H.
Properties of Carbonate Rocks to S02 Reactivity
Borgwardt, R. H. Harvey, R. D.
Reduction of Atmospheric Pollution by the Application of Fluidized Bed
\c Combustion
Jonko A.A.
Supplementary Borgwardt Data
Borgwardt. R. H.
417
-------
Table A-2
ec list_fiqures
For which paper? Isothermal Reactivity of Selected Calcined Limestones \
with S02
List of 16 Fiqures (and Tablesi for dosulplrjrizacion paper
Isothermal Reactivity of Selected Calcined Limestones with SO2':
Figure 1 -- Comparison of Reactivity of Calcines with SO2 at 980C
Figure 2 Sorption of SO2 by Different Particle Sizes of Type 4
\cCalcine at 980C
Figure 3 Estimation of Initial Rate (r3) and Effectiveness Fact
\cor (eta) for the Sorption of SO2 by Type 4 Calcine
Figure 4 Comparison of Reactivity with B.E.T. Surface Area
Figure 5 -- Reaction rate at Constant Sulfation as a Function of P
\carticle Diameter
Figure 6 Reaction Rate vs Surface Area of 150/170 mesh Particle
\cs
Figure 7 Effect of Mode of Calcination on Reactivity of Michiga
\cn Marl
Figure 8 Relation between Effectiveness Factor and Pore Size Co
\cr 42/65 mesh Particles at 988C
Table 1 Kinetic Parameters for SO2 Sorption by Calcines
Table 2 Parameters for S02 Sorption by Calcines as a function
\cof Particle Size
Table 3 Summary of Physical Properties of Calcines
Table 3a Effect of Calcination Temperature on Physical Properti
\ces of Type 2 Calcite Spar, 12/16 mesh
Table 4 Comparison of Reactivates of 150/170 mesh stones at 98
\cOC
Table 5 -- Maximum Percent CaO conversion to which equation 4 app
\clies
Table 6 Physical properties of Raw Limestone, 10/28 mesh
Table 7 Chemical Analysis Of Michigan Marl
418
-------
Table A-3
ec tab figure
Which paper? Isothermal Re-ictivity of Selected Calcined Limestones \
with 502
Whicn .figure? F i g j:o 1
Table of 66 data pomes from
Figure 1 of Isother-nal Reactivity of Selected Calcined Limestones with S02'
fixed Temperature * 980
fixed lower roesh = 150
fixed upper mesn = 173
Bed
5.
5.
5.
5.
S.
5.
5.
6.
6.
6.
6.
6.
8.
8.
8.
8.
8.
8.
8.
16.
16.
16.
16.
16.
16.
16.
Material
S03 absorbed
0
.6
1.1
1.6
2.
2.3
2.7
0
6.5
11. 2
15.2
18.3
0
4.
5.3
6.3
7.
7.7
8 2
0
5.
8.1
10.7
12.8
14.6
14.6
Time
0
20
40
60
80
100
120
0
20
40
60
80
0
20
40
60
80
100
12?
0
20
40
60
80
100
120
419
-------
Table A-4
oc graph figure
Which paper? Isothermal Reactivity of Selected Calcined Limestones \
with SO2 '
Which figure? Figure 1
for what x variable? Time
for what y variable SO3 absorbed
constrained variable Bed Material
value for Bed Material: B
how tr.jny coo--.tca ints niust be satisfied: 1
with interpolation? yes
12.+
I '-data point
. ^Interpolated point
:
r
i
10.4
i
t
9.+
I
I
8.*
I
I +
7.+
I +
I
6.+ ซ
I +
I
4.*
I +
3.+
| +
I
2.* *
I +
I
1.* +
I
1 +
g.t ----- + ----- + ----- + ----- +. ----- + ----- 4. ----- +
e. 17.1 34.3 51.5 68.6 85.7 102.8 120.
420
-------
Table B-l
ec :?v;c
rfiicM ar?i? hoat t r ansfor
?} vbl = 7iri39lo: n _; _ oytsiJo .T? itr.tf.an sf or .ooff ic i ?f>t
oj vol ป vJriillo: Dt ป tab?
31 /bl = ^arii^lo: k ป TJ-; t.lornjl con j_ujtj.yity_
?T v'jl = variiolc: Pr ซ gil Pran-ltl nuibJr
OT vbl ป vjriiblc: -Tosilon = h^i norosity
on vbl ป vjrii'jlc: TI -j * 1 3 o / i s :_o s i t y
C-) vbl ป vjriiblc: Do ป__ ? j r t i - 1 j s 1 7?
JT vbl = v.iriiblo: r'na r. = pirticlo ionsity
C-T ^01 * vjti-ablo: Uf ฐ fluiMsim /clocity
o-j /1>1 ซ viciiolo: i ซ ir.T/itJtiTi.-il cansunt
typ? loft hinl si.'Jc of on'Jition:
( ('i*Dt/Kl /( (Pf *. i) * ( I -on- i Ion) ) )
tyn? ri-jht nnl nil? of i
l:>ir>JJ*( ( Uf*Dtซ^g| /( ( (Oo.MUJ)**i)*rno *g) ) *ซ. 325)
'3o yu wis'i i list of -Jjta sources? y-?3
Jot.i fcoti Ui? fol lowing J o.joors .Jn3 figures :
Eloctric Ot)i qjjrt?rly rooort
; i
H?it tro.nsf-or in pilot pljnt fluiJizoJ bc.l conbastor
lo 1
PER rooort FE123 ?-5/b-2
Tiblo 1
V.iic'.i oioor al 1
ch?ck cortclation: J7 djta points have boon correlated.
do yoj *is:i to plot tin correlation? yes
Jo you wish to tabulate tho correlation? no
421
-------
Table B-l (Cont'd.)
of relation CORRECTION on 12/05/77 Pa-jo 1
126) J
X = lojf ilon) ))
Y = l3i(jJUซ((('Jf*Ot*aiu)/(((Do/10il)**3)*ch3_s*q) )*
C X \XtT V\LU?: loft
3 plottsJ 33 points:
i J;ntifier nine
* riiht
J.OOJf
I
I
2. J4J +
I
2.J5U
lซ
I
2.dUJป
I
I
2. 7r>Jf
2.652+
I
I
2.6UU
I
I
)f
I
I
2. 402 + -----* ซ.__.__ป_--.-ป_- ซ. * >
2.425 2.527 2.632 2.714 2.839 2.941 3.042 1.143
422
-------
Two more examples; i.e., Example C and Example D are given in Appendix II.
VI. PLAN FOR THE ATTAINMENT OF A "SECOND GENERATION" SYSTEM MODEL
The approach to attain a secoiid generation model sufficiently precise for FBC
design optimization is summarized in Figure d. Referring to that figure, the first
step is to develop a simplistic first generation system model using essentially
state-of-the-art models. This development is now coirplete, as is described in the
previous section. Simultaneously, a number of key feature models which go beyond the
current state-of-tho-art are being developed (Table I;. The introduction of such key
feature models in the initial system model leads to a refined and more comprehensive
first generation system model, as is described under another previous section. Doth
tiie key feature models and they system model are being validated against data found in
the developing data base.
In the meanwhile, using the data assembled in the data base, empirical models are
being developed for such bed parameters as combustion efficiency and emission charac-
teristics. When the two sets of models (i.e., analytical set and empirical set) merge,
the combined model will be considered as our "second generation" system model.
VII. SUMMARY
To summarize, our accomplishments during the current phase are:
o Completed a "first generation system model", which o.in approximately
compute a number of FBC parameters.
o Incorporated a slow bubble mojol and a combustion kinetic model into
the initial first generation system model to upgrade the model
performance.
o Developing additional key feature mode-Is to gain an improved
understanding in the physical representation of FBC, and to
establish a broad base for further model upgrading and refinement.
o Evaluated a number of available data base management systems (DBMS)
and selected one most appropriate for FBC system use.
o Implemented a prototype data base system.
o Developed a number of macros for several sample applications
of the DBMS.
ACKNOWLEDGMENT
The desulfurization empirical model was developed outside of this project by
Dr. James Gruhl, to whom we are grateful.
REFERENCES
1. Tung, S.E., A.F. Sarofim, L.R. Glicksman and J.F. Louis, "The Fluidized Bod
Combustor Modeling", Conference on Mathematical Modeling of Coal Conversion
Processes, Washsington, D.C., November 16, 1976.
2. Tung, S.E., A.F. Sarofim, L.R. Glicksman and J.F. Louis, "System Model and Data
Base System for FBC, FBC Technology Exchange Workshop, Rcston, Va.,
April 13-15, 1977.
3. Avedisian, M.M. and J.F. Davidson, "Combustion of Carbon Particles in a Fluidized
Bed", Trans. Institute of Chemical Engineering, 51_; 121, 1973.
4. fl.I.T. Energy Lab Report, "Heat Rejection from Horizontal Tubes to Shallow
Fluidized Beds", rt.I.T. E.L. 74-007, Andeen, B.R., L.R. Glicksman and W.M. Rohsenow,
H.I.T., Cambridge, Mass., 1974.
423
-------
Key
Feature
Models
First
Generation
System
Model
_\/aijdatipn__ __
Refined
First Generation
System
Model
Second Generation
System Model
Data
Base
Empirical
Models
Figures. Modelling Approach
424
-------
5. Borgwardc, "Kinetics of Reaction of SC>2 with Calcined Limestone", Environmental
Science and Technology, 4_ & ^: January 1970, pp. 59-63.
6. Borywardt, "Properties of Carbonate Rocks Relatea to S02 Activity", Environmental
Science and Technology 6 & 4_, April 1972, pp. 350-360.
7. Borgwardt, R.H., "Isothermal Reactivity of Selected Calcined Limestone with SO,1',
presented at International Dry Lir.estone Injection Process Symposium, Paducah, Ky.,
June 22-26, 1970.
8. Merrick, D., and J. Highley, "Particle Size Reduction and Elutriation in a
Fluidizcd Bed Process, AIChE Symposium Series, V. 70, N. 137, 1974.
9. M.I.T., "Modeling of Fluidized Bed Combustion of Coal", Quarterly Technical Progress
Report (No. 2), November 1976.
10. M.I.T., "Modeling of Fluidized Bed Combustion of Coal", Quarterly Technical Progress
Report (No. 3), tebruary 1977.
11. M.I.T., "Modeling of Fluidized Bed Combustion of Coal", Quarterly Technical Progress
Report (No. 4), May 1977.
12. M.I.T., "Modeling of Fluidized Bed Combustion of Coal", Quarterly Technical Progress
Report (Nol 5), August 1977.
13. Gibbs, B.M., F.J. Pereira and J.M. Beer, "Coal Combustion and NO Formation in an
Experimental Fluidized Bed", Institute of Fuel Symposium, Series No. 1, Fluidi/.cd
Bed Combustion, Longdon, Sept. 1975.
APPENDIX I
A SEGMF.NT OF MODEL COMPUTATION
A segment of model computation is shown in Table II.
Referring to that table, the user's objective is to compute combustion efficiency
at Up = 1.5. TPS Ithcrmal power syste-) is a subsystem we use to facilitate the
writing modelling program under TSO environment. As presently structured, the model
first lists a series of subroutines that are executed. It then asks whether the user
would like to print out (1) coal particle size distribution and (2) :;OX profile in the
bed. Alternatively, the display of these stops can be suppressed. The final result of
interest is printed out in the last statement. In this manner, the effect of Uo on
combustion efficiency can be attained with a number of such calculations usinq different
Uo input. Effect of other output parameters by any relevant input parameters can bo
similarly computed.
APPENDIX II
DATA EASE SAMPLE APPLICATION EXAMPLES
Example C. Chock Pigford Model with Borgwardt's Data
An algebraic form of the Pigford model nas been developed by our dcsulfurization
group usinu single point collocation method U2>. the equation is given as:
(f = 4/3.55 (60 + 0.5775 0.5 -f^.2-0.17-.-1 5 + 0.3335 ) (1)
where
F = fractional conversion of CaO
' t
t = time (sec)
F,t and y arc dimensionloss physical panv^ters which are derived
from the model, and are unique characteristics of each calcined
stone type.
In this example, we wish to compare the derived values of time dependent con-
version with actual experimental values. The DBMS contains Borywardt's experimental
data. By use of the macro check_conversion_correl-;Lion, the experimental data is
converted to the comparible form of conversion vs. time. The experimental and derived
points for conversion vs. tine are then superimposed in separate plots for each of
several stones. One such plot is given, (See Table C-l)
425
-------
TABLE II.
XODEL COMPUTATION OF COMBUSTION EFFICIENCY
WHEN U =1.5
o
select uo comef
TPS READY
set uo 1.5
TPS READY
run select
CALL INPUT
OUTER LOOP
CALL PRECAL
MIDDLE LOOP
CALL FLUIDM
FLUIDM: TRANSITION POINT BELOW 5 CM - SET TO 5 CM
CALL HASSEX
INNER LOOP
CALL COMB
CALL DESULF
CALL MATBAL
CALL ENERGY
CALL NOX
CALL OUT
IF THE USER WANTS THE PARTICLE SI7.E DISTRIBUTION
OF CARBON IN THE BED AND ASSOCIATED STREAMS
HUNTED AT THE TERMINAL
TYPE 1. AMD PUSH RETURN
OTHERWISE TYPE 0. AND PUSH RETURN
IF THE USER WAOTS THE NO PROFILE IN THE BED
PRINTED AT THE TERMINAL
TYPE 1. AND i-USH RETURN
OTHERWISE TYPE 0. AND PUSH RETURN
UO - +1.50000E+OO
COMEF - +9.91220E-01
TPS READY
426
-------
Table C-l
Jc ck condor sicncorrol ")t ion
wnich "jol aatoriil i \
esti.notod /aluo fot jan.-na? 52. 65
sstiaat?J //iluo for bot-a? 6.167
Jo you visi 3 list of data sources? yos
data fro.ii tho fallowini 7 oaoors nd fi^uccss:
Isotnifnul R-ปa:ti/ity of 3-jloctoJ CilcinDl Liaostonis vitrt "02
Pijuro 1 Figure 2 Figure 1
Proaorties of Cicbonat? Rocks to >02 R-o-ictivitv
Fijur.? IU Fi-jurซ J Fi7uri 4
Saoolo-nsnt-ity 3orirfacrlt Data
Fijuro Jr]
Viiich o-ioer Supol ?n?nt.iry B
.*riic;i fi-juro Fiiuro 3d
ck_con/crsi3n_corrcl.ation: 16 dat-i points have bco.n correlated.
do /ou wia'i to plot th? ;orrelition? y/?s
do /ou wis^i to t-ioul-ate th3 correlation? yog
Jo /ou wisn the Jati brokon down by source? no
Graol of rotation C^RaEL^TI^^J on 12/UO//; Pa:? I
1 is reporteJ Jat.a
F is co.np'jtod /aluo.
GM1A=i2.65, 9ET\=6. 167
For n-itirial 11. author is "Jor^wardt, R. H.. author's id is 4
1JMERIC X \XIS V\LUE: X
s plotted 35 points:
ilentificr n.an3
Y
P
427
-------
Table C-l (Cont'd.)
d4.
42. ป
I
I
15. *
I
I
2d. ป
I
I
21.+
I
I
14. +'
*
Uorqwardt':;
Data
Compared values
from Pigford Model
I
0.+-
96.
2/4.
. f- - ,
J6J.
. f-.--.
452.
541. 630.
Tabulation of 3'53 Con/crsiDn Correlation for
For natorial 11, .luthoc is Bor-jwirdt, R. H. ,
9jnna = 52.65
beta ซ 6. 167
CoTioute.1 Con/or sion
8.26JJ24
9.dd422ป
10.d3331
14. 00304
17. 16352
19. ซ2792
22. 17531
24.29751
26.24907
28.06554
29.77161
31.33525
32.92002
34.33349
51.60946
7d.86511
luthor ' s i'J is 4
TIME
7.
IU
12.
20
3d
40
50
bJ
70
dO
90
10 J
110
120
270
630
Convor s
3. 734672
4. 7)05d4
7.469)43
11. 70197
16. 1335d
Id. 92234
19.42029
24. 39J35
2b.dU964
30. 37533
32. 36715
32. 11313
34.dba94
36.84376
53.77927
62.49351
428
-------
Example D. Check Desul f urization Empirical Model with ANL Data
The following empirical equation has been developed from a group of data to show
sulfur removal in the KBC.
i sulfur removal =
100-
ฐ'25
- 0.30!
0.420 (fe----
where
C- = parameter for effect of coal source = 1.13 for this application
Ca = Ca/S ratio, valid in ramie from 1.0 to 5.0
lif - fluidizinq velocity, ft/sec, valid from 2.0 to .-1.0
1'jj = bed temperature, ฐF, valid from 1380T to 1650ฐF
Bj = static bed depth, in., valid from 10.0 in. to 70.0 in.
(expanded bed depth = 1.3 y. H(;J
Rs = parameter for number of times sbrbcnt used,
1 + number of times recycle = 1 for this application
This sample application is actually another example of tho use of the
check_corrclat ion mac re . In this case, us is shown in Table D-l, the left side is
simply the neasured quantity of S(j-> while the other side is the equation which
represents the empirical nouol .
429
-------
Table D-l
cc cJiocfc correlation
rfhic!i aroi? dcsulpliur izjtion_
?q vbl * variaolc: c^ซ*Ci to S note ratio
si vbl variible: Uf_ .
oq vbl ซ vari-Vjlo: TปTe.nperature_
eq vbl ป variable: .">l'_s?..tSlel. b.^i fcoi
oq vbl ซ yariiblo: so2 ป 'iOZ rqjipyป?j
cq vbl ป viti-iblo:
typo loft rti.Ti silo of oquition:
typo tiiht *ปjr>1 side af oqj.it ion:
'.J--LV- U*< IJ.b/CT>*lUt**.2j-.
'' '
.-
\c(2. I a -= i) l'7i'2 ? /] ) *' 2 . 0 1 * ( 2 . 24 n 1* - . 2i
do /3J wisi i li-Jt of Jati siur;i>3?
ftoi th? follorftni i aa>;rc mi fiiurcss:
ir j:ht?t irition ml Control of Gmoua Cniziion fron Coal Fired Flul
izod ซc-J Boilers
Tiolc I
i^J DoJ Conbust ion: \ statua report
Fable 1
R?.luw-tion of Ataosph;ri= Pollution by tnc ^ppli;atio.t of KluidizoJ Bo.i
\c Cinbustion
Tible 1
ic:i paoer Fluidiz-jl B.?d Conbuation^^ Jitatuj rcpor_t
fiiurc .Tcble.l^.
ch;ck correlation: 20 data points '10 /" been correlated.
do you wis^ to olot thป correlation? yoa
do you wis>t to tabulate th-; correlation?
should dita bo identified uy source??
P\PER
3IOG RIGMT H^SO 51DE
430
-------
Table D-l (Jont'd.)
3raovi af r?l3tion COaRSlATIO* on 12/06/77
X S3 2
f lOU-O. U'(JJ.6/ca)ป('Jfซ*.25-.3) (l + . 429* ((11.
-------
Tdblc n-1 (Contr'd.)
Flu- HZJ 1 I'll C
Table- 1
445
i'M
441
443
4 J4
446
45)
4s2
447
yl2
4iu
450
4i I
44 d
41 J
9) V
iJu
44 4
454
444
o.-oj-.ti
j
d
j
2'ป.
42.
51.
52.
SJ.
60
61.
6V.
jt.
01.
72.
/4.
74.
7i.
Jo.
ai .
41.
>n: \ ititu-, tcojrt
1UJ
luu
luo
)5. 25779
24. 1-44-i)
5i. 7)7)1
64. 22471
67. 7577
C.5. 67244
55. 4J152
7J. 45244
56. 345dl
36. J45SI
65.42575
08.4)446
70. J=iiH
57. 420J5
11. 215S/
64. 24454
/I. 124U5
432
-------
QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: Are there questions on the FBC
modeling and data base for Dr. Tung? I will ask Dr. Tung to come up
to defend his paper.
DR. TUNG: The one question is by Dr. Saxena: "Could you tell
me something about the inappropriateness of the Vreedenberg Correla-
tion and the direction in which you are working to improve it?"
I shall answer you as I see it, and perhaps Professor
Glicksman, if he is present, would like to make comments afterwards.
The Vreedenberg Correlation was developed for small particles
operating at relatively low superficial fluidizing velocities. In
the present application, we are extrapolating the correlation to cover
the large particle fluidization at much higher fluidizing velocities.
In addition, the Vreedenberg Correlation was developed to correlate
data of a single tube and not for that of tube banks. Hence, if it
does not work, there are plenty of reasons for it.
We are at present developing a new correlation to replace the
Vreedenberg Correlation. The new correlation utilizss a characteris-
tic of a large particle system. As you know, for large particles,
the temperature of the particles does not substantially change during
the heat transfer process. In other words, the bed may be approxi-
mated as an isothermal bed. The conduction component of the new
equation was developed with this assumption. In addition, the new
equation also has a gas convection component. As was shown in one of
my Vugraphs, the new equation correlates the experimental data quite
well.
The second question asked by Dr. Saxena is: "How can your data
base be used by outsiders?"
DR. TUNG: The data base can be used in Chicago as easily in
Boston. The next question, by Dr. Rao of the MITRE Corporation:
"From the mathematical model effort, could you now give some comments
on the scale-up factors that could be employed with confidence?"
With confidence, no. You see, the scale factor is essentially a
matter of fluid dynamics and fluid dynamics is one area about which
we do not feel very comfortable. The physics in fluid dynamics is
not as yet well defined. For instance, the flow regimes under which
a FBC is operating are not well defined. They might be different in
different sections of the bed. All in all, it will be some time
before our model can predict scale factor with confidence.
433
-------
Tne second portion of Dr. Rao's question is: "Does the mathema-
tical model handle pressurized fluidized bed combustors at this time?"
No, not at this time. We intend to do one thing at a time.
After the AFBC is wel1-modeled, we will then extend our model appli-
cation to the PKBC. This v.ill require some modification in the
fluid dynamic aspects of the model, but the kinetic portion of the
model should remain essentially the same.
Mr. McNeese of Oak Ridge National Laboratory asks: "Will you
conduct experiments to oheck your correlation? If so, what type of
experiments?"
The scope of our contract with DOE does not include conduction
experiments, but we are gathering some small-scale experimental data
on NOX emission under an EPA contract. We are also building a 21 x
2* experimental FBC. We realize, of course, that the role of experi-
mentation is paramount and are hoping that we shall have the oppor-
tunity to do more experiments in the future when the 2' x 2' combustor
is built.
A question from Dr. Bastress: "Does the MIT model predict
a particulate emission? Does an information base exist to develop
particulate emission?"
We have a cyclone model, and the cyclone model will treat some
of the particulate problems. We are also working on an empirical
model. The data available are very scanty. The empirical model
attempts to utilize whatever data there are to predict emissions.
Eventually, we shall have good particulate models.
The next question, by Dr. Seth of Gilbert Associates, was in
three parts. The first part of the question is: "What is the
rationale of separating the bed between the fast and slow bubble
regimes?"
The rate at which the oxygen is delivered to the carbon surface
depends upon whether you are operating in a fast bubble regime or in
a slow bubble regime, since with fast bubbles there is a resistance
for oxygen to diffuse from the bubble to the emulsion phase; but witli
slow bubbles the resistance is very small or nonexistant. Of course,
before the bubbles are formed, there is no such bubble-to-emulsion
phase diffusional resistance. Now the rate of combustion very much
depends upon the rate of oxygen delivery to the carbon surface.
Hence, it is important to calculate the rate of oxygen delivery
correctly. It, in turn, means that you have to know whether you are
in a slow bubble/pre-bubble regime or in a fast bubble regime.
434
-------
In a FBC, the bubbles form near the Distributor plate. After
they are formed, they are initially small in size and travel slowly.
For a certain distance, they could be travelling in a slow bubble
regime. After a certain point, because the bubbles grow bigger and
travel faster, they may be travelling in a fast bubble regime. This
transition point can be calculated. We have found that by separating
the model into two regimes, the predictive capability of the model
for combustion efficiency has much improved.
The second part of Dr. Seth's question or comment was: "The
applicability of the Vreedenberg equation to a fluidized bed with
external heat source in the bed is doubtful."
Yes. In uct, for a fluiaized bed in which relatively larger
particles are fluidizing and in which heat is transferred to the tube
banks, its applicability is doubtful even without considering
external heat source in the bed.
The third portion of his question or comment was: "The conven-
tional modeling techniques are not applicable due to the presence of
the heat source in the bed, burning coal particles".
It can be modeled, I believe, but you have to be careful.
Professor Sarofim's calculation has shown that when coal burns, the
temperature will shoot up and then cool down. It has been estimated
that the temperature of a burning coal particle is about 200ฐF higher
than the bed material. All this has to be taken into consideration
if you wish to model the combustion precisely.
FREDERICK HANZALEK, CHAIRMAN: Thank you very much.
435
-------
INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN: Our first paper of the second
session is the one which was originally scheduled as the last of the
first session. This is entitled, "A Model of Coal Combustion in
Fluidized Bed Combustors." This paper is authored by Messrs. Beer,
Baron, Borqhi, Hodges, and Sarofim of MIT and the paper will be
presented by Dr. Sarotim, who received his initial degree from Oxford
University in England, but then subsequently took his SM and Sc.D.
degrees at MIT. He is currently there as a Professor of chemical
engineering. Addo?
436
-------
A Model of Coal Combustion in
Fluidized Bed Combustors
R.E. Baron. J.M. Beer. G. Borghi.
J.L. Hodges, and A.F. Sarofim
Department ot Chemical Engineering
Massachusetts institute of Technology
ABSTRACT
This paper describes the development of a model to predict the carbon combustion
efficiency in a fluid bed combustor (FBC). The model is part of a larger program
aimed at producing a comprehensive model of total system performance.
In order to predict carbon combustion efficiency, it is necessary to be able to
predict burning times of individual particles. Therefore, a detailed model for single
coal particle combustion has been developed. This detailed model has been used to
develop kinetic parameters which can be utilized in the system model and provides an
approximate procedure for accounting for the differences in coal reactivity, as
manifested by differences in B.E.T. surface and effective pore diameter.
System codel predictions of carbon combustion efficiency arc in reasonable
agreement with limited experimental data. The nodel currently proviaes a means for
evaluating the effects on carbon combustion efficiency of mean bed temperature, coal
and acceptor particle sizes and physical properties, superficial velocity, excess air,
and bed height. A nimber of approximations are incoroorated into the systen model
and there is a need for a continuous critical evaluation and refinement of the model
as data becomes available from pilot plant and demonstration units.
INTRODUCTION
Fluid bed combustion (FEC) is a promising technology for burning coal in
industrial and utility boilers in an environmentally acceptable manner (1). One of
the problems with fluid bed combustion Is that the carbon entrainment from t!-.e bed is
high, amounting to as much as 20 percent of the carbon feed (2,3). The carbon
combustion efficiency can be increased by recycling carbon fines, and by such means
efficiencies comparable to those attained in stoker furnaces can be achieved. Higher
carbon combustion efficiencies c.in be attained by use of a carbon burnout cell.
The focus of this paper is on a modeling effort undertaken to predict the carbon
combustion efficiency in order ro assist in the scale up. optimization, and design of
fluid bed cosbustors.
The factors that influence carbon combustion efficiency are shown In Figure 1.
The two major sources of combustible loss are the elutriatlon of solid carbon
(governed by carbon loaning, particle size distribution, and pas velocity) and carbon
overflow (which is determined by rhe carbon loading and accentor withdrawal rate).
Clearly, a model for carbon combustion requires a good description of the carbon
loading in the bed. The carbon loading can be calculated by equating the carbon feed
rate to the sun of the carbon burning and loss rates. The ^olloving section Is devoted
to a description of the carbon combustion rate, since this is a critical input to the
overall systen model.
SINGLE PARTICLE COMBUSTION
The important factors that have to be correctly predicted by a particle
combustion model include the effects of particle size, bed tcnncrnture, oxidant
concentration, and coal-rype on the burnlne times. In the present model the rate of
weight loss of a particle is determined by the rate of evolution of volatile naceri.il
and by the rate of consumption of the residue. Separate models have been develope:!
for these two contributions, but in reco.^nition of the fact that the two processes arc
not necessarily independent of each other, but can occur simultaneously, provision has
been made to allow for a dynamic interaction between the two.
437
-------
^
Acceptor
Stone Feed
Rate
_ Carbon
*^ Overflow
^.
Attrition
Rate
t
Feed Particle
Size Distribution
Bed Fluid \^^
Mechanics |
*
V
-
,-ป
^^
\
^-*
Combustion
Efficiency ป-ซ^
_
Carbon
Loading
-^ - ~~
Supi
* Ve
Bed Carbon Particle
Size Distribution
H
Paitide Combustion
Model
Excess
Air
^~~^ Elutriation/
"* Carbon Recycle
/ f
erficial I/ /
tocity f /
/
^^^^
Particle
Fluid
Mechanics
Figure 1. Factors Affecting Combustion Efficiency.
438
-------
The devolatilization model employed is based on the work of Anthony et al (4).
Simultaneous thermal decomposition reactions are assumed to occur, which are first
order in the amount of volatiles present in the coal and wi.'.ch are described by
Arrhenius-type rate expressions having the -same preexponent:il factor ko but different
activation energies. The large number of reactions occurring justifies the intro-
duction of a continuous spectrum of activation energies. Assuming a Gaussian
activation energy distribution with mean EO and standard deviation a ths amount of
volatiles released up to any time is obtained by integration over the range of
activation energies. The result is:
VR- VR
00
-L-Texp
nsrt J
exp(-E/RT) dt -
E -
dE
where Vp is the amount of volatiles evolved up to time t (grams volatiles/grams
original coal), V-> is the asymptotic value of Vjj for large times, T is the particle
temperature, and R the Heal gas constant. Equation (1) describes the devolatiliza-
tion history of particles :..: an inert atmosphere. When interest is in devolatilization
in an oxidizing nediun, as in the case of fluidized bed combustion, the oxidant attsc!;s
the residue and any as yet unreleased volatiles, so that the value of VR must be
modified to allow for the heterogeneous combustion processes.
The model of residue combustion developed considers a spherical particle wit'.i
cylindrical pores of length equal to the particle radius. For the purpose of
performing mass and energy balances on the particle, values of the Sherwood and
Nussclt numbers are obtained from the Ranz and Marshall (5) corrolation for heat and
mass transfer to a sphere. An oxygen balance on the particle is made by equating the
flux to the particle as expressed by use of the Sherwood number and the total oxygon
consumption due to combustion, given by the sum of the oxygen flux into the pores,
the oxygen reaching the external surface of the particle, and the oxygen needed to
combust the volatiles evolved. For an nc" order reaction at the surface (n ^ 1) the
balance takes the form:
2 DGSh(C.
-Cs>
2R
4TtR2(l-a)knCsn
ATiR2a 2
h(n-H)
OoVOL
where:
h
kn
32VOL
a
Sh
radius of particle.
oxygen concentration in bulk phase.
oxygen concentration at the surface.
molecular diffusivity of oxygen.
pore diffusion coefficient for oxypcn.
pore radius.
intrinsic reaction rate for heterogeneous reaction.
oxygen required for combustion of volatiles.
void surface fraction of particle, taken to be
equal to the porosity of the particle.
Sherwood number for the burning particle.
The term M^VOL 1s calculated from equation (1) with the assumption that the elemental
composition of the volatiles is the same as that of the pnrent coal and that thev
undergo complete combustion. After an initial short heat-up period, this term exceeds
in magnitude the term on the left-hand side of the equation. In this case the
volatiles evolved shield the particle from the oxygen diffusing towards it and no
combustion of the residue occurs. With increasing time volatile evolution decreases
439
-------
according to equation (1) and oxygen can then reach the surface, as is the case during
the initial transient heating, when the particle temperature is low. Equation (2) can
then be solved for the surface oxygen concentration and the oxygen fluxes can be
evaluated. A one-h ilf order reaction at the surface is assumed and the reaction rate
expression used is a modification of that obtained by Snith and Tyler (6) in their
studies with brown coal.
Rn 5 = 96.66 exp(-32.600/RT) C^0'5 ^'^s C (3)
2 cm"sec
This is an intrinsic reaction rate expression which is based on the true surface area
of the coal. The particles are assumed to shrink due to the reaction occurring on the
external surface and to increase in porosity due to reactions occurring in the pores.
In recognition of the fact th*c for short penetration depths of t!ie -xidant into the
pores the par-.icles may appear to burn as shrinking suheres of constant density, a
fraction of oxygen flux into the pores given by (1 - 6) . where P is the depth of
penetration and R the radius of the particle, was also assumed to contribute to the
shrinkage of the particle rather than to a change in porosity. The radius of the oores
is allowed ~o grow due to internal combustion but the cylindrical geometry is main-
tained to simplify the mathematical treatment.
The energy balance on the particle considers raoiation between the particle
and the bed, heat generation due to the heterogeneous combustion reaction, and
convection heat transfer from the particle:
3"R3pCp dTp - 4ปR2aB(TB4 - Tp4) + 4*R2h(TB - Tp) + F^C-AH^) (4)
dt
where Tp, R, o. and Cp are the temperature, radius, density, and heat capacity of the
particle, respectively, f/j, is the molar flux of oxygen with its associated heat of
reaction AHp>. Tg is the bod temperature, Og is Boltzmann's constant, and h the ho.it
transfer coefficient to the particle, estimated from the Ranz and Marshall correlation.
When the volatiles shield the particle from the oxygen there is no heat generation d-je
to reaction and the volatiles are assumed to burn in a diffusion Flarae which surrounds
the particle, so that a term for heat conduction from the flane to the particle is
substituted for the convection term. In consideration of the fact that for the case
of high coal feed rates or low excess air values the fast evolution of the volatiles
aay cause a local depletion of the oxygen concentration and thus quench the flame,
an alternative energy balance which assumes that volatiles burn in a premix.ed mode far
away from the particles has also been formulated. The heat of reaction used in t'lis
energy balance corresponds to reaction oฃ rarbon and oxygen to vicld CO. since it is
known from experiment.il evidence available in the literature (6) that at the
temperatures of interest for fluidizcd bed combustion this is the nain product of the
carbon-oxygen reaction. It is postulated that the burnout of CO occurs far enough
from the particle surface to cause the effect of this exothermic reaction on the
particle temperature to be negligible. This assumption will be tested in the future
by implementing a 'rinetic model for the boundary layer reactions of CO. The solution
to equation (4) describes the transient heat-up of the particles upon immersion into
the bed and their tenperature history during devolatilization and combustion.
The model of coal combustion described stresses the simultaneous occurance of
devolatilization and combustion. Figure 2 shows a schematic representation of thr
Interactions which the model considers. The rate of combustion of the particle is
computed from the rate of devolatilization and residue consumption. The rate of
themal decomposition depends according to equation (1) on the instantaneous volatile
content of the coal, the temperature of the particle, arjd on the specific coal tyne
and composition, sine", the parameters i, ko, Eo, and VR can vary between different
coals. In turn, it influences the temperature of the particle, since the form of ths
energy balance (equation (4)) depends on the magnitude of F02VOL- the oxygen flux to
the particle by equation (2). and obviously the volatile content of the residue bv
equation (1). The rate of combustion of the residue depends on the oxygen flux that
reaches the particle and tJ-" coal composition and affects the anoint of volatiles
present in the residue. The oxygen flux to the particle depends in turn, as shown
by equation (2), on the rate of volatile evolution through 02VOL. on tne physical
properties of the coal through a and h, on the Sherwood number, on the oxygen
440
-------
x*
f
Particle Combustion
Rate "^^_
-/ _! - .^
1 J Volatile Evolution
J[L - Rate
1 Volatile
| Content
Residue Combustion
Rate
T J 1
Particle
Temperature
\
\
Coal
Properties
-j i
Oxygen f\ai
To Particle
i *
Intrinsic
Reaction
Rate
Nusselt
Number
*
^-^
i ซ i
Sherw
Num
i
ood
aer
- Coal
Physical C
Prop.
i /
Fluid
Mechanics
^
1
0:
"Mruxn-
tration
r
Figurt 2. Factors Affecting Particle Burning Rat*.
441
-------
concentration in the bulk, and on the particle temperature by virtue of equation (4).
While the initial physical properties of the coal are generally known, the oxygen
concentration in the bulk and the Sherwood and Husselt numbers depend on the fluid
mechanical characteristics of the fluidization, so that the oxygen profile and
fluidization velocities must be determined before the combustioi model can be applied.
In practice the fluid mechanics routine incorporated In the systems model provides
the combustion routine with the necessary inputs. The predictive capabilities of
this model are shown in the following figures. Figure 3 shows the weight loss and
devolatilization histories obtained for particles 3 nrn. in diameter for an ฐ7, oxygen
concentration in the dense phase at two different temperatures. At 102VK the model
predicts that after a heat-up period of the order of 1 second, volatile evolution
takes place over a time-span of about 18 seconds. Combustion of the residue occurs
over a much larger period of time, 810 seconds, the rate of combustion increasing
due to the effect of pore widening. At the higher temperatures the same character-
istics are visible although the time scale is reduced due to the exponential
dependence of the rates of devolatilization and combustion on temperature. Burning
times obtained with the use of this model were found to agree to within a factor of
two with experimental data.
The devolatilization times shown are commensurable with particle mixing times,
which are of the order of seconds for typical fluidization conditions. This implies
that the assumption of 'nstantaneous devolatilization at the distributor plate
commonly made in connection with coal combustion cannot be considered to be valid
for the large particle sizes and low temperatures characteristic of fluidized bed
operations. Although the kinetics of devolatilization must be considered in
calculating local hot spots and other parameters sensitive to the local fuel/air
ratio, the devolatilization times are a small enough fraction of the total burning
times that they can be neglected when the objective is the limited one of this paper,
notably the calculation of carbon combustion efficiency.
Figures A and 5 pertain to the two cases shown in Figure 3 and illustrate the
detailed information that the model can provide. In Figure A, the porosity, density,
and temperature of the particle as well as the ratio of surface concentration to free
strean concentration of oxygen during the combustion process are shown for the higher
bed temperature, 1323ฐK. The ratio of oxygen concentration is a measure of the
extent of pore diffusion and kinetic limitations in the overall combustion process.
Also shown are the radius of the volatile diffusion flame and a ratio of the reaction
rate calculated assuming reaction to occur on the outer surface of the particle down .
to the true reaction rate. The value if this ratio is a measure of the importance
of pore diffusion in the process.
Following the particle from the time of immersion into the bed one notes a
transient he?.t-up period during which little combustion or devolatilization occurs,
so that the density and the porosity of the particle remain constant. As the particle
heats up, volatile evolution -ets in and a volatile flame ignites. Conduction of heat
from the flame causes a further steep increase in temperature which in turn enhances
the devolatilifc-irion rate. As the volatile content of the coal drops, the flame
eventually receeds and attaches to the surface. As the radius of the volatile shell
decreases, the temperature of the particle reaches a sharp peak due to the proximity
of the flame. The density of the particle drops quickly during the devolatilization
and the oxygen ratio, initially having a value close to one due to the low tempera-
tures and slow surface kinetics, drops to zero while the particle is enveloped by the
volatile flame and then rises to a steady value during combustion of the residue. It
can be seen that the residue combustion is mainly controlled by external diffusion,
the oxygen ratio being low, so that the particle burns in a "shrinking core" fashion
at constant density and porosity. Only towards the end of its life-time, when the
particle has shrunk to a small size, is there an upturn of the porosity and surface
oxygen ratio and a slight decrease in density. For small particles these trends are
nuch more evident. It is however interesting to note that even for the 3 mm. particle
reaction in the pores accounts for most of the combustion occurring, as can be seen
from the low value of the reaction rate ratio. The penetration depth, however, is
apparently too shore to cause significant changes in the particle porosity.
It must be pointed out that the values obtained for the radius of the volatile
shell are too large to be considered realistic. This is due to the fact that the
volatile evolution is an intrinsically transient phenomenon which does not lend itself
well to the quasi-steady analysis of the combustion process that has been used in
442
-------
CASE I : T 1 323'K
KLAME FORMATION t-O.S:scc.
^FI.AMF. ATTACHMENT C-6.91.sec.
E 2: T " 1023*K
Ayr KORM.t- J.-9 sec.
LAME ATT. f]fv)0sec.
II II.)
IUO'>.f)
Figure 3. Bumout and DevolatilUation Histories
of Pwttcla* with Ignition VoUtiles. Particle Radius H = 0.15 cm.
i.c-5r
K.
ฐ>
> H
C W
tea
o
.70
(a)
POROSITY
01 L.
0.1 1.0
1O7D
1CO.O 300.0
6.0
fa. 0
0.1 1.0
TIME (SECi
10*J X TCK) (b)
0.] 10 10.0 100.0 300.0
TIME (SEC)
(c)
VOLATILE FLAME
RADIUS (cm)
10.0 100.0 300.0
TIME
-------
developing the model . A more exzict treatment of the volatile flane is currently
being iaplesented. Figure 5 exhibits the same trends evident in Figure 4. Due to the
lower tenperature. however, it can be seen that the burning of the residue is
controlled to a larger extent by the surface kinetics and pore diffusion resistances.
the oxygen ratio averaging about 0.5 during that time. The nodel predicts that.
depending on operating conditions and particle sizes, the ccnbustion rate nay be
limited by external rr.ass transfer or by pore diffusion and chemical kinetics. Figure 6
shows the ratio of oxygen concentration at the surface to oxyp.en concentration in the
bulk, which is a oeasure of the importance of pore diffusion and reaction kinetics in
the overall process. Since the surface oxygen concentration varies during combustion,
the value selected for che purpose of this graph is that at the half-life of the
particles, when the oxygen concentration has in fact reached a steady value. It
appears that decreasing the particle size and lowering the temperature has a narked
effect of shifting the conbustion process from an external dif fusionally contrclle-f
to a kinetically limited repine. For the temperature and particle ranees of interest
for fluidized combustion, the kinetic resistance is by no means negligible, so that
allowance for kinetics at the surface clearly must be made in predicting burning rates.
carbon loading, and combustion efficiencies.
It is unclear at the present time whether the building of an ash layer on the
burning particle is an important factor that should be considered in modeling the co.il
burning process. Allowance for diffusion through the ash layer will be made in f'irurc
refinements of tho single particle model, and the sensitivity of the results to this
aJded resistance will deterninc whether it kill be considered in an improved version
ot the systems nodel.
The physical implications of the results shown on Figure 6 are evidenced in
Figure 7. where the different behavior of particles burning in different regimes is
dexonstratcd. It can be seen that large particles at high temperatures bum at
almost constant density in a "shrinking core" fashion, while snail particles at lov
temperatures burr, in depth. The size of these particles changes little during
combustion while their density
-------
1.0--
0.9 '
0.8 .
0.7. .
j
i
<0.6-
= 0.5.
o
t-
Ui
?0.4.
gO.3
gO.2'
0.1-
h-
100 ISO
JOCO
PARTICLE DIAMETER (pm(
3000
Figure 6. Dependence of Kinetic Resistance to
Combustion on Particte Sim with Ignition of Votatitev
02
0.4 0.6
TiVE/BURNOUT TIME
1.0
Figure 1. Density Chang* of Ptrtictes with Combustion. Co* of Ignition of Vounites.
445
-------
ks - " K;
When reacrion is fast, it can be shown (7) that definition of a Thiele modulus:
k C n'1
kC
De/R2 (8)
where De is an effective diffusion coefficient for oxygen in the particle, leads to
the result:
n , J_ or n * -5 --- * --- (9)
3ซ 3R
so that n is inversely proportional to particle radius. Since AT is porportional to
particle volume and As to the particle R2 , it follows that:
n A^VA,, rป f(R) (10)
and thus ks is independent of particle radius, as is the case when using the
unmodified first order rate expression. When reaction is slow n approaches one and
thus ks is proportional to R:
kgaR
The detailed combustion model for a single particle can provide values of n which can
be used in equation (7) to obtain a value of ks that will be compatible with a first
order rate expression hut will now depend on the particle radius of interest and will
therefore allow more accurate predictions. Since the different reactivities of coals
depend mainly on variations in the value of their specific surface area and not on
a difference in their intrinsic reaction rate, this procedure in essence enables one
to introduce coal type as one of the important model parameters.
Due to the dependence of the effectiveness factor on the particle size .
-------
5.0
4.0
23JO
in
rป
.1.0
o<
o
I
I
I
I
O2 0.4 0.6 0.8 1.0
FRACTIONAL WEIGHT LOSS
Figure 8. Variation of Nomuliied Effectiveness Factor
During Combustion.
447
-------
A second order approximation is to treat the volatile evolution
as uniform throughout the bed since it has been shown that the
devolatilization tines are comparable in magnitude with the
tines of particle reoirculation in the bed.
2) The temperature differential between particle and bed is
small enough (see Figures 4 and 5), other thrn for the short
initial period of strong volatile evolution, -''at they can
be neglected In the system model
3) The density of the particle can be treated as a constant
over most of the burnout period.
4) The chemical resistance is important but can be acco-jnted
for by use of a simple kinetic model accounting for resistances
in series of diffusion and chemical kinetics:
kdif? kocxp(E/RT)
for a first order reaction, vhore the difference in reactivity of
different coals can be inserted if kinetic data are available on
tiieir oxidation rates. In the absence of such date the variability
which is due to differences in pore structure and surface area can
be obtained by substitution in equation (71. For example, for a
coal particle of radius 0.15 cm. and 250 m-/p. speciftc,,surf3ce area
with a porosity of 3U percent and a pore-radius of 2*> A burning in
the presence of a 2 percent oxygen concentration, use of equation (7)
in conjnction with the results of the single particle model predicts
the following rate constant: ,
Ks - ^.96 x 105cxp(-22.7SO/RT) gatoms C Am
cm-* /
This relation can be compared with the expression given by Field et al (5).
Kc - 1 38 x lO7' x T c-xp(-35.700/RT) gatomsC
lee
At 1323ฐK the rate constant derived for the Smith and Tyler intrinsic kinetics is a
factor of 2.2 larger than that of Field et al (8). This difference may well be
attributable to differences in surface area and pore diameter, in addition to
possible variations in intrinsic carbon reactivity. It should be noted, however,
that the difference in effective activation energy is large. This has a profound
significance for purposes of modeling the temperature dependence of combustion
efficiency as will be shown later.
S/STE.M MODEL
There are four main components in the system model: fluid r.ec'-nnics, combustion,
desulfurization, and heat transfer. The model, receiving information on stone and
coal feed particle size and density, bed weight, temperature, superficial velocity.
etc.. will predict carbon combustion efficiency, steam generation rate, heat transfer
coefficient, flue gas emission levels (SO?. NOX), stone utilization. ,ed expansion.
and the particle size distribution of the'carbbn in the bed, in the recycle leg and
in the elutriant.
A brief description of tr.odel components appears below. This paper will be
particularly concerned with the structure of the combustion conponent. The other
components are treated in a more comprehensive fashion elsewhere (9).
448
-------
Fluid Mechanics
The Cranfield and Geldart bubble frequency correlation (10) is used, along wivh
the necessary gas mass balances, to predict bubble growth throughout the bod. The
distributor plate/slow bubble region (i.n the lower section of the bed) is treated
as one phase plug flow. The upper section, if in the fast bubble repine, is treated
classically; two ohases are assumed with the emulsion gas well-stirred and the bubble
gas in plug flow. The Davidson-Harrison correlation (11) is used to calculate bubble/
emulsion phase gas interchange races in the fast bubble regime.
Combustion
As mentioned previously, a valid first approximation for single particle
combustion in fluid bed boilers can be based on instantaneous devolatilization,
constant density of the burning carbon particle, and the Field correlation for first
order surface kinetics. The oxygen mass transfer rate to the particle was estimated
using the single particle correlation for Sherwood number.
A modified form of the population balance suggested by Kunii and Levenspic-1 (12)
has been used to calculate the bed particle size distribution. The modifications
were necessary to account for the attrition and elutriation rates.
COMPARISONS OF SYSTEM MODEL PREDICTIONS FOR COMBUSTION EFFICIENCY WITH SELECTED PILOT
PLANT DATA.
In order to test the validity of the systems model, three experimental programs
were considered: the Pope. Evans, and Robbins (PER) fluid bed colum (?). the six
inch National Coal Board (SCB) atmospheric combustor (14). and the experiments of
Gibbs . Periera, and Beer (15). These experiments were selecred because they were
simple, with no peripherals such as carbon recycle or burnup cell to complicate
interpretation of the data. They were also selected in order to evalu.-ite the model
over as wide a range of operating conditions as possible. The PER studies were
performed with comparatively large particles and high superficial velocities, arid the
NCS data with small particles and low velocities. The Cibbs. Periera. and Beer
experiments were in the intermediate range.
Figure 9 illustrates data of Pope. Evans, and Robbins. along with system model
predictions for combustion efficiency. The data were obtained over seven hour runs
and the range of data are shown. Average values, when reported, have also been
included. For the two highest bed temperatures the agreement between theory and
experiment is reasonable. However, for the cwo lovest temperatures the model appears
to underestimate bed combustion efficiency. The temperature effect (calculated using
the rate constant of Fields) seems to suggest that the activation enerpy used is too
high. This is consistent with the findings of the detailed model of particle burning.
Figure 10 illustrates the data of Gibbs. et al, alone with system model predic-
tions for combustion efficiency as a function of temperature. The predictions
obtained using the Field's rate constant compare reasonably well with the data.
However, the predictions show a stronger temperature dependence than was observ:d
experimentally. The model underestimates combustion efficiency for the lower bed
temperature. This is another indication that the activation enerpy used in the Field's
constant is too high. This activation enerpy (35.700 cal/pmole) represents an average
value obtained by fitting a first order model to a large number of data. These data
were taken with many different types of carbon, ranging from electrode carbon to coal
char. The single particle burning model previously discussed in this paper supeests
that a much lower activation energy might be appropriate for the bitursinous and
lignite coal used in fluid bed combustors. For this reason, a modified first order
correlation was generated with an activation enerpy equal to one-half Field's value.
The modified correlation was developed by setting the two constants equal at a
temperature midway between the two data points. It can be seen that the modified
constant does a much better job of predicting the temperature dependence of the
combustion efficiency than dees the Field's constant. This emphasizes the need for a
single particle burning model which can take account of differences in the reactivitv
of various coal types.
449
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Run Number C-321
EXCESS AIRU)
UQ (M/S)
BED TEMP (ฐK)
nc
-------
Model predictions are compared with NCB data in Figure 11. Efficiency is plotted
as a function of excess air. Predicted trends appear to be correct, and the predicted
values appear to compare reasonably with experimental data. The relative importance
of chemical kinetics are emphasized with small particles as well as at low bed
temperatures. Since the system model underestimated corhustion efficiency for lower
bed temperatures, it would be expected that the model will also underpredict
combustion efficiency for small particles. This does appear to be the case.
Correlation appears to be Rood. However, it should be recognized that large errors
in carbon loss estimates will be less evident for high combustion efficiency than
for low efficiency. For example, an agreement between theory and experiment of one
percentage point in combustion efficiency would amcunt to an error of about 5 percent
in carbon loss for an efficiency of 80 percent. For an efficiency of 99 percent.
however, this would amount to a 100 percent error in the carbon loss estimate. The
apparent agreement between theory and experiment for the smaller particle sizes
should be considered within this context.
In summary, system model predictions appear reasonable for conditions expected
at full load bed operation. However, for part load (low temperature) operation, it
appears that the first order rate constant of Fields, et al. overemphasizes the rolo
of chemical kinetics. A modified version of the Field's constant, using a lower
activation energy, provides a realistic estimate of the temperature dependence of the
bed combustion efficiency.
CONCLUDING REMARKS
Elutriation of solid carbon represents the major combustible loss in fluid bed
furnaces under normal operating conditions. Carbon loss in the fluid bed combustor
is high. It may, in certain cases, exceed 20 percent of the coal feed (this may be
compared with a combustion efficiency in a stoUer furnace of over 95 percent and in
a pulverized coal furnace of over 99.5 percent). The high carbon carryover results
from .a desire to minimize furnace size by increasing the volumetric he.it release rate
in the bed. This is usually accomplished by running with large sized ccal feed and a
high superficial velocity. Unfortunately, this results in a rapid elutriation of t'.ic
carbon fines generated oy attrition and combustion. Carbon combustion efficiency can
be Increased by recycling these elutriated fines, or by use of a high temperature
carbon burn-up cell. Carbon recycle loops and burn-up cells introduce capital
expenditures which must be considered within the conttxt of an overall optimization
program. The expense of a carbon burn-up cell (CBC) is probably justified in a
large utility furnace, where capital costs are amortized over a comparatively lonp.
time period and there is a great need to reduce operating costs. The CBC would be
less useful in a small industrial boiler, where there is a greater emphasis on
minimizing the initial cost of the boiler.
The system model can play a useful role in performing the cited trade-off
analyses. The model also has a potential use in the design and scale-up of fluii'
bed combustors. The model has the capability to predict the dependence of bed
combustion efficiency on temperature, coal particle size, superficial velocity, and
carbon reactivity. The agreement of the theoretical predictions for combustion
efficiency with limited pilot plant data is gratifying, but there is a need to
incorporate a more refined model for char combustion in the system model which can
take account of the varying reactivity of different types of coal.
ACKNOWLEDGEMENTS
The research work presented in this paper has been supported by the Department
of Energy under contract number E(49-18)-2295. The financial support and the
assistance given by DOE in the course of ihis study are gratefully acknowledged.
451
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] . OII
0 . 9 8
u
So
o
ป-t
b.
b.
U U
7.
O
9f,
94
0 . 9 2
O 0.
o
r.XCKSS A1K 1 (t-xp)~| nc
12. f I" 9>.<>
9.8 1" fi.it "| 94.7
6" DIAMKTF.K
ATM FLUID BED
OF NCB-CRE
REF.:I'ii-Ma ป>;>
KUNS ป 1.2 I, 1.5
> 10
EXCKSS AIK
Figure 11. ompariwn of Experimental Data with System
Model Predictions for Combuition Efficiency.
452
-------
1. Comparative Evaluation of Phase 1 Results frora the Energy Conversion Alternative
Study. NASA TMX-71855. February, 1976.
2. Waters. P.L.. "Factors Influencing the Fluidizrd Combustion of Low Grade Solid
and Liquid Fuels", Institute of Fuel Symposium, Series No. 1, Fluidized Combustion,
London. September. 1975.
3. Cordon, J., et al, "Study of the Characterization and Control of Air Pollutants
from a Fluidized Bed Boiler- The S02 Acceptor Process". PB-229 242. July, 1977.
It. Anthony, D.B. , J.B. Howard, H.C. Hottel, and H.P. Keissner. "Rapid De-volatilization
of Pulverised Coal", Fifteenth Symposium (Int.) on Combustion. The Combustion
Institute. Pittsburgh. 1975, p. 1303.
5. Ran^. W.E. and W.R. Marshall, Jr., Chemical Engineering Progres 48. 1952, p. 141-
146 and 173-180. ~
6. Smith, I.W. and R.J. Tyler. "Receptivitv of a Porous Brown Coal Char to Oxygen
Between 630 and 1812ฐK". Combustion Science and Technology. 9. 87. 1974.
7. Carberry. J.J., "Chemical and Catalytic Reaction Engineering", McGraw Hill
Chemical Engineering Series, 1976.
8. Field. M.A.. D.W. Gill. B.B. Morgan and P.G.W. Hawksley, "Combustion of Pulverized
Coal". BCURA, LEATHERHEAD. 1967.
9. Louis, J., et al. "Modelling of Fluidized Bed Combustion of Coal", Quarterly
Technical Progress Reports I through 5.
10. Cranfield, R. and D. Geldart. "Large Particle Fluidization". Chemical Engineering
Science. Vol. 29. 1974.
11. Davidson, J. and D. Harrison. "Fluidization", Academic Press, 1971.
12. Kunii. D. and 0. Levcnspiei. "Fluidization Engineering", John Wiley and Sons. 1969.
13. Merrick, D. and J. Highley, "Particle Size Reduction and Elutriation in a Fluidized
Bed Process". AIChE Symposium Series. No. 137. Vol. 70, 1974.
14. Reduction of Atmospheric Pollution, Appendices 1-3. National Coal Board. PB-21067A.
15. Gibbs. B. and J. Beer. "Concentration and Temperature Distribution in a Fluidized
Bed Coal Combustor", Presented at the Combustion Institute Europeon Symposium.
University of Sheffield. England. 1973.
453
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QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: We are now ready for the question
and answer period. I think that a number of questio.... have *>een
passed now to each, of the authors; perhaps not to Dr. Saxena, *>ut
please furnish him with them while I go more or less in time ore -.
Dr. Sarofim has had the most time to prepare, and also most people
have had more time to prepare questions for him. You have a few,
presumably. Would you please read from whom the question comes
before you answer it?
DR. SAROFIM: In view of the time limitations for the presenta-
tion, we did not present the details of the model, and the questions
that I have received relate to those. We also didn't have time to
acknowledge prior work in this area, and I would like to comment that
two of the questions come from two individuals who have done pioneer-
ing studies in the modeling of fluidized combustion.
Professor Saxena has a three-part question. The first part
says: "Your external diffusion resistance refers to gas film and/or
ash film diffusion resistances. From where do you get your values?"
We do not allow for an ash film diffusion resistance. We used
for the external gas film resistance the correlation of Rentz and
Marshall. We are aware of correlations that have been presented
for fluidized bed operation showing much lower heat transfer correla-
tions, but we are reasonably convinced and have evidence to support
that the correlations such as the ones presented by Sikali and
Reischenden for fluidized beds are not valid and that Rentz and
Marshall is a much more appropriate one.
The second question says: "Do you account for bubbles in the
bed?" And the third one is related: "How are gas and solids flow
approximated in the model?"
We have relied heavily on the inputs from Professor Glicksman
and his work in modeling fluid dynamics. We allow for the fact that
bubbles are born small, and grow; and as they grow, they eventually
reach a size where their velocity is less than the mean velocity of
the gases in the bubble, or they move into the slow bubble regime.
So we account in our model for a fast bubble regime up to the
point of transition to the low bubble, and allow for the Grandfield
and Gc-lgard correlation for bubble growth, and the Davidson and
Harrison coefficient for exchange between the bubbles and the emulsion
phase. For the circulation of solids in the bed, we assume the
solids to be well mixed.
454
-------
Professor Wen, West Virginia University, also has a three-part
question. He comments first that "combustion efficiency is roughly
equal to the coal feed less the eleutriation rate." He then goes on
to ask: "Have we compared carbon loadings in the bed with experi-
mental data?" Of course, he is referring to the fact that the
eleutriation is proportional to the carbon loading in the bed, and we
have worried about carbon loadings in the bed. The numbers we get
are in the range of 0.1 to 1 percent, and we have compared them with
the very limited data that are available in the literature. We find
good agreement whenever data, such as the British, with small par-
ticle sizes, low carbon loadings are on the lower end of the scale.
The data with larger particle sizes have carbon loadings closer to 1
percent, and again, we predict the right magnitudes. The data are
sufficiently imprecise, though, that we really cannot say that this
is a great check. We do need more information in this area.
He asks a second question: "Did you take into account the
swelling of the particle?" And the answer is, "No." The evidence, I
think, is slight; I would like to hear the data which support very
significant swelling. But at this stage, we do not allow for swelling.
And the third question I have already answered. It relates to solids
mixing.
Dr. Rao of MITRE has a four-part question. I only have three
cards, but 10 questions. "How does the ratio of devolatilization
time to the particle burning time vary with such variables as pressure,
temperature, and particle diameter?"
The devolatilization times really are kinetically limited; and
so we say that the time required for devolatilization is entirely a
function of temperature, and it's only a function of particle size
to the extent that particle size, in fact, also determines temperature.
The burning time follows in the diffusion-limiting regime a
square power dependence on particle size; and as you would expect, as
you get to smaller diameters where chemical kinetics becomes more
important, then dependence on particle size decreases.
The effect on pressure; there we have a first-order dependence
on the chemical rate constant which can vary down to a half-order
rate on the chemical rate constant. On the mass transfer coeffi-
cient, of course, you have the canceling effects between the effect
of pressure on concentration and the effect of pressure on diffusivity.
The second part of the question: "Does the model currently
account for the discrete feed points in real systems?" We do not at
this stage have a solid diffusion model, which is what is required
455
-------
in order to account for the discrete feed points. We anticipate that
in future work that we will, as we get into the details of solid
mixing into the bed, then address this very important practical
problem.
I don't think I should comment on the third one. He says,
"Could I comment on the numbe1* of feed points required for a given
AFB combustor." Well, the rule of thumb is one feed point for nine
square feet. The only comment to be made is that it is a function of
circulation rates, too; the interesting questions are, can you people
have proposed parting beds for overfeed, and I think this is something
that I should leave to -he practical sessions to answer.
"Does the model incorporate feed size distribution and determine
the particle size distribution in the bed and in the off-gas?" We
certainly allow for the feed particle size distribution and use the
population balance mechanics to follow the particle size distribution
within the bed, with allowance for the effects of eleutriation and
attrition. In the off-gases, we do not allow for a freeboard combus-
tion model at this stage, but it is important, and we have plans to
do so in future work.
456
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INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN: We will proceed innediately to
our third oaper of the second half of the session, which is entitled
"Fluid Dynamic Modeling of Fluidized Bed Conbustors." This is
authored jointly by Messrs. Bar-Cohen, Glicksman, and Hughes of KIT.
The paper will be presented by Dr. Clicks-nan, who received his
initial training and BS degree at MIT, wandered all the way out to
Stanford to get his MS, but then came back to MIT for his Ph.D.,
where he is currently on the faculty in the Mechanical Engineering
department. Dr. Glicksman?
457
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Fluid Dynamic Modelling of
Fluidized Bed Combustors
A. bar-Cohen, L. Glicksman.
R. Hughes
Department of Mechanical Engineering
CT Massachusetts Institute ot Technology
The dvnaalcs of ^.v bubbles rising and growing vlthin the sollds/gaa enulslon are today known to
define the operating characteristics of rvtny practical fluidized beds and* in snal l-part tele svstens,
l...ve be.-n sliown to follow the tw>.-phase theory. The present effort focuses on the fluid cv.hanics of
larw particle fiulJization. as Is anticipated for fluldlzed bed co;ahustors. nnd extends the Iwo-ph.ise
the.try to this iapctriant flufdization category. The results sugp*~Kt that tvo-phase theory, when com-
bined with an appropriate bubhle-f requer.cv relation, can prov* !e an accurate prediction uf bubble size
and bubble fraction and their variation with height in the bed. Furthermore. examination of these re-
I. it ions in Hunt of the particle sizes .ind air flow rates proposed for prototype fluidi'ed bed eonbus-
lors suggests that an ippreciahle part of bed operation will be in the slow-bubble regime where ouch
<>t the erajlsion gas f K>wป through tlte bubble void. In this regime, gas flow through the bed can be
noJelled as plug fli-c and thus provides substantially less resistance to nass transfer than is encount-
ered In the conventional, fast-bubble reside.
I".,.- operating characteristics of t lulrilzcd bed toabustors. Including: coabustlon efflcU-ncv.
t'vd txpanslon and he.tt - rets-tva ! tate. are largely deterained by the f lu id-dvnamlc merbanlsBM active in
the ! iii ii' irt.-d oedlun. In bubhleii rlsinK throtich the enulslon are viewed as the Second
ph. ire. Kvtfnilve i>tvซ-]it I ป-ai lป-ns bv I)aviJs\>n and co-workers |1| have shovn that the dynamics of a single
nan biiohli- la : lluldlz>d Bed Inn are Ind l-t in,;u Ishable fro^ Iho-:e of a i;as bubble In a hoaogeneous
liquid and that potenllal t li*w the rv as first developed bv tMvles and Tavlor [2| can be used tci estab-
lish the velite.lt v tieldi In tlie sur r.'.iml I n>; eeulslon aป \Jvl'. as In the rising bubble.
In flutdlrod b4 ! i-i>nhustซrs, I lie coal p.irtlclei ore anticipated to constitute onlv a verv sT.all
fra.iloa (Ivplcallv ir> ot t ! p.ir t I. een obtained In InveBt Igat Ions of relatively snail particles
(tvplc.illv HXM) and l.n.- tluldlrlng veliซc 1 1 lvป. II Is thus Inportant to Identify and quantify the diff-
erence betveen the tluld dvn.iaic
-------
CMUktlON
Figui* 1. SMtch Of Gซ Flow For A Slow Bubfalt.
At S*cn In Bubbte't Rctenno* firm*
459
-------
-.-jr. ! - l\. t. I..} -.-.. !j.,v i
..I'.ii.i:.,-. t-. I!,;. I. !!.ซ
.1.
-.1 !.. !.-: ii,.- '_ .1::-! -.^ I :. v.!.i-i..ir I : I ,. t i nn: t.-I..T iv. t-
m-i int.-iri.i! i.-,-i'-n. I- .;. t i .->!-. .V. . ,t ::. .-.-.. ., .. I.r/.-r
-< - >.:.t.-:'..- t .-(ป it I .it : ;ป .-..! I.--.-. .if i i . !-.-.. i i..1.1 : -i ' t.r ป..
!li- .. ! l.-vln-.'. l!ir.-i.-l. t!..- !-.'.i.|.. v..l-! .!> i-.-.lul it.-! ! ;.,...-.; .. v.-1 i t v ' - : i.
r. .;..-. I t.. Hi.- l.u!.'. I. .iii.l il..r. i -. ! it v ! i , -' _, i. I ii Iv.- t.- t!..- : i ! ! i .< i :! u' - i . !.'. .;:.,
II..- f-.ill!. -.1 l>ivl-l-.'ili .in.) IP.Ili...ll |l|. t! '.- i. ..I Iv li-iit.-.l I., i .. I.,!..! l,.|'.'il.- . ! !.!.! I
|n.il !. !l -..!.l ! t.-.l. l.-.-.'-vi-r . I'i it I1" i..i'.(...- ir.-.i iri.ti' i !--i I'M. i . f.:.!.! I -.. :... !- . ..
i.-. ,......... .. .1... I...M.I.. v.l... it.- ..j.i.i..... i.. : I,..- -.I..,/!., .t lr.,n-.ji ,...
'I..V.-..T. Ill .! I..-.!
.Ii .11 il.ni..t. i , .
Hi.- itil.l-.lili.il ; .-. /.). If- ...ii I..- .l.-t.-rnlit.-l l.v .:: nlii.:. .: .ซ-..-. --I I >..' f'i!.ll .:,.! I . . . n fi.
I '. ! , 1 1., it I i.. in!, r .1 it i. . I .-..-. .! ll v, I . r.i^l .ilv.i ;. I . -.u. .', ... ( I ..... t. r inirun t ! i ..-I . / it i"ii ..!
t!..- 1- itl I. I, . .ii-l ll,.i .
I - I' I' / - I' I"
.- -. nl 7^1 -- t
I...HI imillv .-I -;-.|i-!. I I..VT, . ..i.-.i-,1 In-,: ! iiM |.-w Iv 1 1.- !..il.l. I.- w.i . .ici-l .l.-ut,!-.n
II..-H (t- M-..-.I t.. ..M.iin II..- r.. nil-.il.- ....II. I ;>irlnl>- -.!. ilv. I . .,-
< roitlnltii; ll..-s< .-l.-o.-m > int.. .1 sln.-.li- rvl.it lr .-r-pir I. .11 Iv ili-t .-r= ln.->l u.ili- (r.nti.-n. I h.- ;:.is . .mt |:>H|| v n-l.iti.-ii l--r ->>-
lilt- I lซ*v Is I"..IIIK! .is
I' ' I I --'-rl ซ! 'I I- ,!' ซ Ml * -' ' . ป n . ป'. HI - i )' I ('.I
... nf c >. nt nl h n r:f
Snhst.intl.il si.-s-.'l 1 1 i- .u I. MI n .-.in H.- .1. !i l.-v.-.l hv .iv-iialn^ UK- w.ilซ- fr.i-ll.-n. ,. t.-r -;1i-v
IwihMcs to IM- nrr. I i .- ilป I v ;.n.-i 1 1 , -is sui^.-st o*l hv r* : .-r.-n.-*- I * | . <>r -i-; ^.in s.- i^pl l.-.l t r>T. t hป- .-.-i-. ! li-w
.ir.xiivl .1 :!* h.iM.I. . U'.tli thin .i-;su-;pt Ion .in.1 vuh^i 1 1ซ: ;..n fr->3 l!q. I ,mปl .' fur I !>c '.MiMil.- rise
vt-I.H-itv. K.|. 'i .-.in h.' rx-dlflfJ t.' r.-.nl
l"he rnnllnnity of s lnv-hiibHIe c.iซ! fl
-------
.Ict.iture. ..-.eve r ป ; :i
ut.. :u..Mv lr.-iiu-n.-y ..:..:
**J " * * . * ti * r itt .ri
; j ป..lo:4 , ^-- .1 i tul *:i.' I I . . 1:1 t'.** v.' 1 .*r*. tiic t uLM*.- f !'- ft 1 .il inn , I ; . ?
LuLi- !ป t f ..t 1 iun i r. ii'U.'td to ."J't.tl
-1:1 J i no lu- in;- t ; . M ป t :u*
ior sj'vcifivd nr ซ-i >;.jr i r.i 1 1 / J--tป-r:.m.-J v.il-i.-s nl ป. ..:u! t , t;^- buILU- Ji.ir.etfr auj hui.Ulo (r. action
at v-ir ion-. ucii>tt-t c.iit .i i *iu i ;!ป ->i- !elซ-ri-Tnf.l f . ปr 'ป .tcl. input v.i 1 w nt i.t'perf lci.il f't*ป vปr 1-
i-c il / Ly -.it^ul tanvi'u ป 1 y -.oi vir... i.'\. J .itiJ '>. ttu^ovttr , I i^.- rnn;ปlrx ป!ป'pซ'n*Jrnre -i! i on bub U It* J i,ttu-tป*r
t.-ir.t--. ซ.-x;i 1 .ci t '.oiut 111:1 f -if t .tesซ' p.t r ur-'l crs all tjut injtussiblt.* .ปnJ .in t to rat ivc nuncr teal li-clmi-jut* ,
uซ."ปcr i UvJ ii. Ootat 1 in ri-f ซ-rctici* 1 1 1 * ** upy 1 IcO.
W!ปi 1* .'.ever at t .I'tt - '..u!>blซ' cor re l-*t tuns arซ' av *I 1 a!> li* in the t ltซr,tture and a recent "no due to
I'.trtcn. et al. \'t\ .ป;'pc-.ir.-ป to .irhievc *.i mil tc Jnt ^ccur.icy, none of t!ies-.~ furo-jl.tf.lonb at tetr.pt ii to
*ati.,ty K.IS (low continuity via t:ปป- two-piiu**** motion::. TIw ! re'iu*-ncv/two-piirtปe fomul^tion 'lt'.cu:.*od
-itovc ro-*ซ'nt el furl focuses prlewri ly on * low buLblc
J/n.ป^icj, it is oi interest to extend t.'ic nodcl into t!i*' fast-buMilc rcp.loe by appropriate nod I f ic.it ion
.,* i,,. ;.
liubLtc ~** flow ia 'itis rซ'>>,ia4> Invulvcs r.as rcclrcula^ Ion in tl.e bubble cluud and walco as well
.1, tue i;ปป vulunu rl-,lnr. i i tin- but. Me, liut In cuntr.v.t tป tlic ilow lni'ublc (low. there 1ป no throupii-
f!uw In .in l:. c :
ป>sปunint; once again tltat tiic isolated-bubble results can be used to characterize Che behavior o( bubble
bwarns and coaoininj; bubble (law with r.ai (low in the emulsion pnuie. ซ;aป-(lov continuity (or this
rcKinu can bo cxpfcsscJ as,
Solution tut the (a>t-hubblc dianetcr and bubble (ractlon can now follow the procedure discussed
(or tite uluv-bubblc rcr, Ine by use. of Cranficld and Celdart's I?) (requrncy correlation. As w.is noted
earlier, while this correlation wai developed (roe slow. bubble d.ita obtained wltli larRe-part Icle
(iuldlied buds. 11 agreed <(uile kvll with Culdart's ! -1 1 s=al 1-part Iclc/ (ast- oubb-lซ results ind could.
thui, >crv.-> as a starting point (or a cocpietc 'jujble oodol.
Tlio relations presented above have been baaed exclusively on the potent lal-(lov codcl for an
isolated l-ubbla. Ycl, Ja t!>c relative bubble rise velocity approaches the interstitial Ra> velocity,
L. L(, the transition (roa slow to (.i^t bubbles, Ujvldion and Harrison's aodcl. as veil as later
rcilnencnta, predicts a rapid increase towards infinity In tile olzc of tiic rcclrculat Ion zone. ซ~illc
this conditlcn nay be approxiutvd ' y a single bubl le in a l.ir,-c flui.Ii^cJ red, t'.c size of the recir-
culatlon zone around in individual bu'-ulc In -i uuULlo :i.-.i;n -rut c'.'.irly j;-p.'c.ic'.i x finll>- llnli .1.4 live
461
-------
bubble r Ise ve I'M- it v approach*--; t hซ- f nr erst 1 1 Jal >;ป* v*-lซป i f v. The f lปj Jd-nechanl'- 'li-ta I i s ป;f this nซ're
ซ ormnrilv I'll* nuntered ซ out J ;'.urat i-- such .'* solut ion verv unwield 1 v .
Alternately, In tin- pre<..-nt model, the tvn-phase e*;u.it inns b.isซ-d tin an isol.-(t**d buhble vtrซ- .'i1.-
Mj&4t! lo a[.;ปly as Ion;-, a-, t tt.- ป|ซj i va |t-nl sp!n-rf'.il vulun.- '.! I (M- h-*hM*-, Its vn\rv and Its rlnud ftปr
rซ-f i r< i: I at i<ปn /on*-) d Id iml ovt- r lap t hซ- -i'l jaซ-c-nt hu(ปl>lซ /w.^t-/ซ l*.nfl vซปl urac. Consi-^urnt 1 y, < a Irulat iซm
til r l:i- fjfซ'< I.,e t*uttMt--s I /ซ var I at inns nป .ir t tiซ- s lov/f .isi t ran1. It ion rvjst /tw.t i t a nor*.* ret Jr*-d rป>d*.- 1 and
.ili'l.r-.I. lnt*-rpi*Ial I'.a t>*-tw<->n i In* fast aiwl slow !ปuhhlซ* rt.'1.il Inns ausl suffl'o fซ>r the pri-sซ-nl t Irw,-.
VAI.IIiAII'iN OK KI'HHI.K **/)[]ป,!,
IKJC to tin- v-huhMซ- d-ita In t h*- I i tซ-ratur*-. <..r.plt-tซ- validation of the Iwo-phas**/
I ri-qtifitrv hn'ibU* K'"V* li wซlซ-l d.-rlv.-d in MM- |ir.-vlซ-us section Is r*ปซil dilflrult. There ar*.-. n*-verthe-
less, several central fjin-st Ions vhle answered hy cneijiarlnR the prซ-JIซti'ปn of various f!ov
parameters with put>Ush*-d xp.-r f m-nl a I results.
In the O.infield ;irirl fu-ldart study (?) tmth huhhle f rt-^iiency and huhhle d|.-imt-l*T w^-re measured
hy Independent neans. The re Int Ivc (*rc-dic| I ve cap.ibi I i t v of i he present h;iti!ปle growth mode 1 and
Ikirtun's l.isl-liuhMr fnrnnlnc inn | ** | * .m I hus hi- ex.idfned hv mnp.'ir I son with Cr.mf Iซ- 1 d and C*-lUart*s
1 7 t ex per (men I a I d va lues. 'I he exrel 1 1 -ill at'.rceoent h*-t wป-ป-n t he f r*"?' i*-ncv/t wo- phase t heorv and dat .1
.,hi>vii in Figure ^, --rovltlos 'ป initial iont Irnwition of t!w v.ilidjlv <ปf the two-phase thi'ory In t :te slow
huhhle regime. A 1 1 crn.il ป 1 y . the failure of the- Unt'-n taodel to anuralely predict the Cranffeld arttl
Celdart results h 1 i-h I U-.ht s the limited applicability of this rorrel.'it ion to slow-hubb lc- data
In addition to huhhle t-rn-t Ion dfataeler. McCrath and St re.it f I e Id f 1O| reintrtrd visual r;thvUes
of t he nuciher of Imhltle erupt Ions, N, visible .it .1 *;! ven instant. S ince this nur.ber I s appro:'.Inat ely
.qu.il lซป I he number of bubbles present In a control volume orrupvInK tปซ' tซ'd <-ross*serC Ion and oqual
In hc-lKhl to a bubble diameter. Hie enpirL -; bubble fraction was ra leu lated arrordinp l
and Inserted In the two-phase i^uat Ions to vie Id a bubble K Ize. As c.m be seen In Fl >',. 3, this rovlded
euch Inprovrd a^reem-nt with the McCrath and St rc.it field | 101 data. Furthermore, the resul t InR ^'allies
of bubt>le diameter are ;i(;.'ilii seen to suggest that a 132 reduction In the experimental * effective bubble
Nlze a;iy be appropriate. With Much a corf* l.'it Ion, tbe r^^xilfled frequencv/two-ph^se model (Including the
ftpec I f If, empl r Ira I Iv der Ived bubble f r.ict ion) wi*uld yield cxrel lent resul Is.
The validity of the present frequency/two-phase approach Is ronflnaed, as well, bv examination
of Verifier's data obtained with typical ooan particle diameters of 103U fill. Ttie Investigation w.:s
alraonc exclusively In the fast-bubble refine, but Is one of the few in the literature that provide* an
explIcit bubble-frequency corre Iat Ion. Tit Is eopirleal expression, shown in Kq. 13 below
fฃ - (0.039H + 0.57)"3 (13)
la Bubstantlal1v different frnai the Cranfleld and Celdart relation and was Inserted In the frequency/
two-phase model for this particular calculation* The dramatic Improvement In the predictive accur.ic /
of the present txxicl resulting frott the use of f** Is clearly shown In Kip. 4 where the locus of the
bubblc-dtaocter prediction using f** In cnnpared wfth that bated on Cranfleld and f^eldart's frequency
(Eq. 8) as well as with the Dartnnw(9) hubble-Rrซwth oodel.
462
-------
- 10
ป
S a
ฃ
3 2
sirrscis:
2 5 IO
BUBBLE DUKETER < MEASURED). d (ea)
20
Figur* 2. Infhjtnce Of Frequency On Bubble Sin Prediction
Oau ol ConlnkJ 4 CeUart 1 7 1
463
-------
i/st.-'i'
e o
.2
in
o e
i
464
-------
j;.*j k.: L.--- iuuuo.--.mtal difficulties in .i:tdl/2iur uluv-Lubble Urliaviur i i the need to determine
aad account for t.*e t.troug.if low and rocirculatiou of emulsion r.-*s- CK-aily, at very K>k bubble ri^e
velocities it ii tiie fon^r taut dor.inates, Lmt as t.*e slow/ fast transition is appro ached. the cross-
M:*:t.lonal **rca for tiirouK.if low decreases and rocirculation involves pror.rt--jsivcly nore of ti>* bubble
v,.lmปc. IM.-VVT t.Mflcss, it :u> Leon ^uw.t'Stctl tiiat for tu-.t ซ'f tlie ujov-bul'ble re*:*,ino, it oay be possi-
ble t-ป cx^r^aj i;a:ป flow continuity for a sinp.lr Luablc in sic^lificd forn [I'*} a;> below
04)
i-tar.ia.it ion of L'\. 14 in light of the rare rlgoroui r.as-fl
t.ut t:te lainii and Leซens;iiel fornulation c
A norc JotjllcJ L-iunination of tlic effective tiirou-.lifIbu coefficient for Crjiiflrld anj CelUarc's
invvi>tirซ*tion 17) i*ป Uisplayod in Fii% 6 vtitrre a rvlati0)
[1J a;ปuo^r^ to a^ply over nuc'i of tite tปlov-|jubblc ranp.e.
inu vx,>etirซ.ntiii variation of '., olitalnud t>y Cranflcld and (^.-Idart (7) in two different run* of
a larr,ซ.--. 'article fluldijcd ued operating in t.ic blow l>ubolc regime, witli t. /L. ran^inK froo 0.4 to
>>.0? and 'J. itf to iJ.'^A, rcb;H:ct ively, is u!tovn in Fl^(. 7. T>*e Intllc.itcd data pointa were obtained by
M>lvin>; O\. 14 at *}>cclficd bed lป.-Ii;!it -., v!>erc bul.bic tlze vaa c-xpor Inrntally dctvrnined and reported.
'i.te solid line represent :ป a icni-c-npir ic.il e-juation for c obtained froa the bubble frequency and size
coi relations ;iri:ซenteJ b/ Cranficld and c seen in Fir,. 7 loth tl'O calculated
value j a. id tite locu? of tlic t correlation bliov tiic anticipated trend wltli bed Itviglit and vliile tlio cal-
culated value* arc -e:ierally uelow tiic correlation, t'.ie dlfforencc Ij typically less than II".
Subtle Si/.e I're.l ict iun^
-..UKKv&t!i that tiic tvo-pltase theory of fluidlzed bed behavior can be
Ituc perforaancc of fluidlzed beds 1* closely linked to the flow reglnc prevailing in tlie fluld-
izcd nudlua and varies substantially as the WJ passes Iran Incipient fluldlzation through bubbling
and on to slugging. Lxanination of tiic governing relations suggests that in an open bed of specified
geooctry. particle and gas density, and pxrtlele shaoe tlic superficial velocity at vhlch the transitions
between various flov regincs occur is dependent exclusively on particle size. In succeeding paragraphs
tne approximate loci of cinlaua fluldlzation, slvw to fait bubble transition, slugging, turbulent flow
ซnd particle teroinal velocity will be discussed. The resulting flow-rcgloc naps for two typical bad
configurations are displayed In Figs. 8 and 9.
465
-------
80
[; T>
n
w
(r
- 70
ฃ
n 60
a
in ซ
HAi.^-'rr-fri.fcs I^'ATF. AVK^A-KS
| >j ^At.1:11! ATRO sir'
ERROR PARS: 1 1 STS. DEV.
DATA OP CHAV?IELD A?;0
CEU5AST : V "
e '
DOTS: VALtrES CAL^LATED PRO" DATA
OP C8AKPIELO ANT) 1ELOAHT 71
O.?0
'ERPHIAL VELOCITY (^PESlrENTAL J, u (e?./a j *
g
Figure 5. Comparhon Of Superficial Veloeitiei Calculated f;
With And Without Recirculation. thing Eqnป. (8) and (14). J 0.10
Respectnety-Data of Crcnf ield & Geldart 171 ^
c.or.
0.00
srr.sct
o
o
o
V8.ซ
10 10 10
HEIOHT. f! (c*i
ซ0
Figure 7. Variation Of Bubble Fraction With Height
Above Distributor-Data of Cranfield
& Geldart (71
0.0 0.2 04 0.6
ut,r/ur
0.8
1.0
Figura 6. Throughftow Coefficient At A Function
With Height Above Diitributor-Oatl of
Cranfield and Geldart |7|
466
-------
rrr^
en
o>
e
5
i r J
ง5!
o I
is
si
&
K
467
-------
:!lnlpun riutdlzatloo Velocity
^ille several correlations are available for predicting the superficial velocity required to
achieve particle t luidization. aost, if not all, seem to be based on a force balance between the v>.*ielit
of liic particles and the bed pressure drop, calculated from Ergon's correlation [12].
Tiic graphical display in Figs. & and t reveals tlic anticipated trend of Increasing t'cf vlth
particle size. Comparison of the run! nun fluidizatlon velocities for 2UOu and 2000'j particles accentu-
ates the significant difference in superficial gas velocities encountered in flolnp. fron snail to large
particle fluid! zed beds.
It oust be noted, of course, that aggregative f luidization is not generally achieved at alni-;ua
fluldlzatlon, but ratiier requires sone further increase In superficial gas velocity. Soae early --or*
by bavldson and Harrison |1| suggests that the transition from particulate to aggregative fluldization
iccurs wlien the ratio of predicted bubble diameter to particle dlanseter is in the range of 1*10, but a
precise correlation of this fluidlzed bed .transition is not yet available. In Its absence Cranfield
and Ccldait's [7] Bea->ur>.-d bubbling Incipience excess velocity, UQJ, - U , - 5 en/sec, nay serve as a
point r-f reference for beds operating with relatively large particles.
Transition Between Fast and 1.1 ox Bubbles
As discussed in the previous section, the flow of gas In an aggregatlvely fluldized ooJlum gives
rise to two distinct bubble rugiacs, depending on the ratio of the bubble rise velocity to the Ras
velocity between the fluldized particles. I.e. the Interstitial velocity. The <. -nmonly encountered
fast-bubble regime Is associated wit a velocity ratios greater than one. while for .--ios l"ss than one
the bed is a.- id to operate in the slov-luublc regime. Tor a given particle size, the superficial velo-
city at vhlch tlie slow/fast transition. I.e. L'ur - t'Qf/cDf, occurs can be determined by use of the
f requency/two-pliase bubble raoucl previously discussed, with t'^ equal to
la i.q. '* The ainic~a fluidization velocity relation previously discussed can then be used to relate*
tlic transition velocity directly to the particle lilancter.
examination of Figures
-------
nay not Lซc applicable to larr.e-?art icle, sn.il l-ป!iarx.'tt-r beds, where the slur, rise velocitv nay fall
lปvlo- t.ie Interst i t 1 jl c.iu velocity a:ij thereby p.ive rise- to .1 slov-;;lmป vliicli lias 7ns throur.'.f lo'*'
-ปi::ilar to a slow bubble.
J i iicnnt iuuous i.eglne
l.i-cenl c::.'C*r iijcntal o: servat io::;. Ji'ป, lj. '.'] '.'irj'.est that nt appruxir.itel y or.o-half t;te particle
terrinal velocity tiic slui'j'.inr. b.-havior ceases and is replaced Sv a ttir'iilent, c'l.iotic flni: durlnf,
wliicii erratic voids and snail flow channels as veil a-j vigorous ..- ircul.it Inn of Mu- .-ol Ids In fin !><'!
tซoco;ic a:>narunt. .'Jut to the linited nun!ปซ^r of tur'-nlfnt flow cxporinent^ i*crforr^-d thus far, tiie
Ijoundary of this post-sluf.^inj; flow rt-rlnc r.ust at this tine tnป viewed as tentative.
Turninal V<-locity
Tlio uppc'-Lound on t'te fluidlzatlnn velocitv for a solids/etas <.*rtilston is attained when the
superficial (;as velocity is equal to t!ie t-rr.ilnal volocltv of the particles and thry are elutriated
froa t!ie bed. In a fluldlzod radius wltli a sicnlficant particle size Jlst r li.ut Ion, tlie >-lut ri.it'nn
process occurs over a ram**.* ปif superficial velocities, hut t!io terminal velocitv of the nean particle
size can l>c used tf> characterize this final fluidized bed transition. Following Kunii and t.evensnlel
j!>)t the particle terninal velocity can he found !w iteration nsinr a different rซ*l.ition for each of
t'aree r anr.es of particle i'^ynolds nunher and :lio locu* of this final flov-rerlr:e l.oundary Is shown In
UPS. a and 'i.
These flow regime naps reveal that vlth ir.crnaslnp partic-le size, the oncratinr. ranpe of tlic
fluidized l>ed , i.e. the region between ninintun f luid i ?.at ion and terminal velocitv, decreases and is
occupied by a progressively larger slow-t>uht*1c rซ'r.irin.
Tim Influence of I'art iclc Size
Lxanlnat ion of two typical Kluldized Red Conbtistor f low-repime naps rcvo.ila the Innortant role
played by particle size In dctcrninlnp the !>cd im-r.it I IIP. n-i''"'1. Aป particle size Increases, the
operating ran|-e of fast bubbles Is found to docrca-ie and the s).>w-bubMe rer.inc bccoirci progressively
r^>re inportant, Furthernore, in relatively larrc-nart icle, :ir..l 1-dlar.eter bods, the slue rise velocity
nay fall uelow tin- interstitial velocity and give rise to a "slow-slui;" rer.lr.e. Alternately, since
riobl expertnent.il fluidized teds to-datc 'i.we been designed to operate with reldtlvclv snail particles,
the f lov- repine nap can be u*cd to explain the prevalence of fast-buhblc models and the relative
scarcity of :*lov-Lubblc formulations.
The results of the present work yurc.eat the following conclusions:
Uirr.e-partlclc fluidized beds linv flow regimes radlrally different fror snail-particle beds.
The slow-bubble flow rer.lnc and the transition fron sloซ to fast 1-ubMe flov rerlraa are very
inportant for larr.e particle svstctn. The two-Dlia.sc hyvo thesis appears tc be valid for larrc-nart Ic le
beds.
Tlic classical bubble oodels used In the two-phase theory have llnlt.it Ions in large-part Iclc
aystcns due to overlapping clouds near flow transition and the inapplicability of U'Arcy's law.
The bubble size can t-e approximated by the use of exnerlocnt.il frequency correlations.
The effort reported heroin is part of a comprehensive FDC modelling study being conducted at
JUT under UU)A Construct :lo. E(19-l 8)- 2290.
A - bed crots-scctlonal area, co
D - bed dlanctcr, ca
d - particle dlasietcr, en
469
-------
d - vuluoetric iM-an Uootle diaoeicr, cm
f - level frequency, s
f - Cranfleld & Ccldari's point - freiuซicy correlation, s
f" - '.Jurtlier' :* correlation for frequency per unit area^ c=i s
;; gravitational acceleration, 'JซJQ.*>^3 en/s~
ii - level above distributor, cm
II, - Halting bed lielp.i.t used In !iacye*s & Celdart's sluฃzฃar. criterion, CD
u - throughflow coefficient
N - number of bubbles la sooe specified volioae
iu.1 - particle Keynolds fza&et
f - superficial fluidizuig velocity.ca.'s
L - bubble rise velocity ftr a freely uubbliag bed, cm/s
Li ' relative bubble rise velocity, cn/i
t. fy(. - bubble rise velocity at the slow/fast bub-ble transition, CB/ ;
b' _ - interstitial gas velocity, absolute velocity of p.as in tiie eaulsion phase, era/*
L' - absolute velocity of teas in the eon;!Bion s*uue at aininoo lluidlzatIon, L ,/c .. ca/i
I *J KX
L , - superficial (;as velocity at wliicii brjbbliij; begins, ca/t
L, - superficial f.al velocity at mint cum fluidiution, cn/ป
L - velocity of solids ta the en-ulslon phjse (defined positive for upward not 1cm). c*/*
1. - terminal velocity of particles, co/ซ
V - bubble volume, CD
V - cloud volun:, CD
V * uako voluDc, en
ft - bubble wake fractloa V^/V
ii - bubble cloud fractioa V /V,
c b
f - voluoctrlc fractioa of Internal reetrculaiioa reeion
n
i - volunctrlc fraction of external recirculatioo region
t - bubble fraction
c , - bed voldagc at niniasa Huidlzatiaa
ml
REFERLIJCtS
1. J.f. Davidson and D. llarrlton. fluldlzed Particles. Caabrldge University Press, Cc&brldge. 1963.
2. K.H. Uavles and C. Taylor. Pruc- Roy. Soc. A. 20u. 375 (1930).
1. X. Ilughci, S.M. Thesis, Uc^artaent of Mechanical Eagloeerlng. KT, Cttfcrldge, Mass, (in preparation).
4. U. Zuber ซnd J.A. Flndlay, Trans. ASIIE - J. of Heat Tr.nnfer. vol. 37C. fr. 453
S. D. kunil and O. Uvcnsplel, floidtiatlon E;:jtneซrtag. Wiley and Soas. Sev York
470
-------
6. :.. Olciart and '...?.. CranfielJ , C'.ien. r.n.-.rtun, K.U. La.iiiizo. .1.1. N.iviJson, nr.J :>. !!arrl:;on. Tr.i^.s. I. r:ier.. !.. . Veil. 3j, p. 37
(I.-77).
l.i. :.. JtUrat'i and K.I!. '.trc.it field, Tr.ins. .:" I"-.t. c.f Ciiva. Tn.-.. '.". 7" ('. 7:;
II, J. .*.rt:ur, in I luiui^jati- T 1.'c'.r.;>It>.-'-' (.-tJitft! i.v ;>.'.. :\-.ii r*i-. ; , '.'!. I, r. 2K> ^]'17').
11. .1. !Ucyc-.-.:, and 1>. ^-Id.trt. i:'ion. !"n-. :cl., 2~>, 2'jj (l''7i).
14. J. "jtrai-is, :;.:!. Tit*.'Si5. Ik-p.irtr.i-nt of .%-c':ianic.il !'.nr Ineer inr , "IT (in ''r^^'ir.ic ie-i ).
li. c.ir.ad.1. ';.:;., ::.;:. "cLju-.'.ii:. .m..1 !...'. :iau!>, r:inซr :'.>. I'UJ. (.'::. ..;f.: :...-, -tim-. c'i:c.n-o, in.
1^76.
l'"j. Ver-A^':ialni , J. , :i.I'. "urnur, .trv! A..'l. fi^ulros. In. I. fnr. C'IPL . , Pr^^os :*ei. I.'ov. . 1>. A7 (!'>7^i
17. ฃit?t>;lu<-Y, Veil. 1. 1>'.7. n. '.I.
471
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QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: Dr. Glicksrnan.
DR. GLICKSMAN: First of all, I would like to acknowledge the
efforts of my colleague, Professor Bar-Cohen, who has been working
with me for the last year on this particular program and is sitting
in the back of the room. That is called "sharing the guilt," if I
can't answer some of the questions.
The first question is from Ralph Wood of General Electric, and
the question is: "Have you tried other coordinates to describe the
flow regime map?" I presume that what he is referring to is a way we
can use a non-dimensional coordinate systen which will generalize the
flow regime map? We have looked at that, and there is no general
coordinate system one can use for different phenomena which are being
controlled by different physical variables. For minimum fluidi-
zation and for the transition to this turbulent flow or diffuse flow,
we are concerned about the drag on individual particles. When v/e are
talking about transition from slew to fast bubbles, we are concerned
here about bubble frequencies, and the partition between flow between
the bubbling phase and the dense phase.
We chose this particular set of dimensional coordinates just to
give people a physical "feel" for what flow regime they would expect
to encounter for the particular particle size and velocity.
Now there are a couple of other parameters hidden in the flow
regime map. One is the height above the distributor plate, which has
a modest effect on flow regimes, and also, of course, the density of
the particle. This nap is constructed for standard atmospheric
conditions; for a heated bed with calcined limestone, the flow
boundaries are going to shift.
His second question was: "How do you reconcile the poor predic-
tion of combustion efficiency that you presented with the good
predictions shov/n in the preceding MIT report?"
The point I was making is that when we used just the one-region
bed model , the one where we assurae fast bubbles from the bottom of
the bed to the top, you get lousy predictions. That is confirmed by
the earlier work that Professor Sarofim and his colleagues got. In
some cases when we have gone to this two regime nap with a slow
bubble at the bottom of the bed and fast bubbles at the top as
Sarofim nas done, you start getting much better predictions of the
combustion efficiencies. So this paper should have been before
Sarofim's, to set the stage for vdiat he was then going to present,
using that same two-regime flow raodel.
472
-------
Finally, a question by Dr. Rao of MITRE: "Can a small error in
temperature measurement result in the lower predicted efficiency
shown?" Temperature has some effect on the combustion kinetics, as
you saw from the results of Professor Sarofim. The slope is rela-
tively modest, so one would not expect that that is the problem. I
think the problem is, and I would like to reemphasize this, a fast
bubble model for a bed containing large particles doesn't work. And
the reason it doesn't work is it's not modeling the correct physics
within the bed. Certainly we haven't modeled the entire physics. We
must include the area near the distributor, where the jets from the
distributors are penetrating a certain distance, giving higher
particle Sherwood numbers because of the higher agitation around the
distributor. That is an area that we hope to model in the future.
FREDERICK HANZALEK: Dr. Saxena?
DR. SAXENA: Can I make a comment about this?
MR. HANZALEK: Surely. You're the cleanup speaker; you have
every privilege.
DR. SAXENA: Now, there is no hassle here. Probably there is a
difference in the terminology and I want to make sure that I under-
stand things correctly.
You know, in the slow and fast regime models, what we are
talking about is essentially the thickness of the cloud around the
bubble; and that thickness also depends upon the bubble size, so you
don't have to have two compartments. What you have to have, actually,
you build in, and when the bubble is small, the cloud is thin. As the
bubble grows, the thickness of the cloud increases, and therefore the
diffusion resistance builds up, and you've got to account for that and
the same computer program can do it. In fact, we did this type of
report in my own modeling work. I'm sure most of you are familiar;
it was out in the Fuel Journal only a couple of months back.
DR. GLICKSHAN: This is not entirely correct. Both slow and
fast bubbles can have clouds. The key distinction is that fast
bubbles produce a significant mass transfer resistance between the
bubbles gas and dense phase while slow bubbles allow the gas to
freely flow from the dense phase to the bubble, eliminating any
bubble to dense phase mass transfer resistance.
473
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INTRODUCTION
FREDERICK HANZALEK, CHAIRMAN: We proceed immediately to the
last paper of the session, which then will be followed by the
question and answer period. This paper is entitled, "A Mechanistic
Model to Explain Ash Agglomeration in Fluidized Bed Combustors."
It is coauthored by Drs. Rehmat and Saxena of the Institute of Gas
Technology and the University of Illinois. It will be presented by
Dr. Selena, who received his initial degrees, bachelor's and
master's, from Lucknow University in India. He then went on to his
Ph.D. at Calcutta University in India. For several years he was
a research associate, and became research officer of the Bhabha
Atomic Research Center in Bombay, India. Coming to the United
States and serving at Purdue University, he assumed his present
position at the University of Illinois at Chicago as a full
professor in 1968. Dr. Saxena?
474
-------
A Mechanistic Model To Explain Ash
Agglomeration in Fluidized Bed
Combustorsand Gasifiers
S. C. Saxena
Department of Energy Engineering
University of Illinois
Chicago. Illinois
A. Rehmat
Institute of Gas Technology
Chicago. Illinois
ABSTRACT
A mechanistic model for ash agglomeration has been proposed In which the higher Internal temper-
ature "f a reacting part'cle for a given UeJ temperature under certain favorable conditions will celt
the Inorganic ash and which in view of its high surface tension will ooze out of the particle and will
Ber^e with similar ash beads to torn agglomerates. The frame work of matliematical equations describing
..nls model are derived and have been solved for a range of operating and system variables. The experi-
mental data generated at 1CT have been discussed in detail and explained on the basis of proposed
theory.
INTRODUCTION
During coal combustion or gasification an Improved carbon conversion efficiency can be achieved
if the loss of carbon in the MSh is minimized. One way of .ichievlng tilts is to separate ash and carbon
from the coal particle, segregate coal and ash particles within the reactor and remove such segregated
ash particles from the reactor. The combustion or gasific.it inn process itself separates the carbon
and ash from coal. However, segregation and the selective removal of ash from the combustor or gasl-
fier still remains a major industrial problem. Using a fluidized bed reactor and agglomerating ash to
the extent that the agglomerated ash Is made heavier than the rest of the bed material offers as one
of the solutions to the above mentioned problen. The agglomeration of ash generally requires that at
least part of the total ash produced during combustion or gasification acquires the ash melting temper-
ature which is generally in the range of 1150'C to 1400SC. However, operating the reactor at such
high temperatures may result in the formation of a massive clinker. A much desired method of agglo-
merating ash at relatively low bed temperature is proposed here.
Since the carbon-oxygen reaction is exothermic. It is possible to achiev.- higher than average
bed temperature within the coal particles by appropriately selecting the operating parameters. When
the temperature in the range of 1150'C to 1400':C are attained within the particle, the inorganic ash
melts and some ash beads ooze out of the cracks. The molten ash has a high surface tension and does
not wet the coke particles readily. Thus, the ash beads that ooze out of the coal particles separate
from them and merge with similar ash beads from the nearby coal particles. The molten ash beads also
wet and trap the solid ash particles present within the bed. An ash agglomerate grows by collecting
inore and more of such liquidous ash beads. It Is prevented to grow too lurge by the quenching effect
of relatively lower bed temperatures and by limited collisions with other hot ash particles due to
turbulence present in the Huidized bed. After the ash agglomerates grow sufficiently large, they
drop to the bottom of the reactor from where they c.m be withdrawn selectively.
The ish fusion temperature depends sensitively on the composition of the inorganic ash, and
may vary considerably froa one coal to the other. In order to be able to agglomerate asli from a given
coal feed, its fusion temperature must be known to establish the conditions in the reactor to achieve
these temperatures within the particles. The internal temperature of the coal particle is generally
a function of the average bed temperature, the initial size of the particle, the oxidant concentration
In the reactant gases, and the atjount of ash produced during reaction. The latter is Incorporated In
the size factor Z, which Is defined as the ratio of the volume of ash produced to the volunsu of coal
reacted. For most of the coals, this volume ratio varies between 0.07 to 0.2. The ash agglomeration
can proceed at bed temperature much below the ash fusion'temperature, if the factors affecting ปw
Internal temperature of the partible are optimized. With a view to adequately understand '
menon of ash agglomeration, the effect of each of the above factors has been investl"
paper to determine the favorable conditions to produce'ash fusion temperatures w1
to initiate ash agglomeration.
475
-------
a^glozierat ion model which nay be applicable to all the processes in various forns to enable correlation
of the data obtained 510 far and also predict the conditions necessary to obtain ash agglomerates under
a wide variety of operation conditions.
In the following section, a mechanistic model has been presented to enable the determination of
>f the core temperature :>f a coke particle which would determine the plausibility of ash aKf.loraer.it icn
mder the given condit fairs of size, partial pressure of oxygen, the ash content of the Riven feed stock
md the average bed temperature.
und
THEORY
In the derivation of equations which mr.y enable the evaluation of the internal temperature uf
the particle, we r.hall assune that the bed is composed of raany individual feed particles suspended In
the fluldi~cd bed maintained at a constant temperature Tc>. The bed is fluidl/.ed by a gas havinj; an
oxidnnt concentration of Co. Since our interest is confined mainly li\ establishing the maximum tempera-
ture attainable within the particles, we shall focus our attention to a single particle being exposed
ti> the fluidlzing ftas. of oxldant concentration of C0. Thus, tlie problen reduces to that nf a single
particle of arbitrary radius Rn and hence use can be made of the heat and mass balance equations for
the single pellet of changing size derived by Rehmat and Saxena\ These equations have been normalized
with respect to the initial particle size R0, the bed temperature T(), and the partial pressure of I he
oxidant, P0, present In the reacting gases. The quasi-steady state approximation will be employed for
the nass transfer equations, and the unsteady state heat transfer equations will he used to obtain a
more accurate estiiute nf the temperatures that develop within a particle.
The temperature within the particle is assumed to he sufficiently high so that the rea'-tlun
C + J ฐ2 cf) (1>
predominates. For simplicity it has been assumed that rhe feed is In the form of char. The rate
expression, the effectIve dlffuslvity, cas film diffusion coefficient and the activation energy have been
used b-ised on the work of Dutta and Wen''.
The reaction of equation (1) is primarily diffusion controlled at tbe temperatures of present
Interest. The char particles are assumed to be spherical and the p.as-solid reaction takes place
following the shrinking core model. The reaction Is confined to a thin layer at the Interface of coke
and ash. A detailed description of shrinking core model is given elsewhere^. According to Dutta and
Wen*", the rate of reaction Is given by
S - *,v
Here, S Is the specific external surface area of the particle and k is the rate constant based on
Reic B
external surface area of the reacting solid.
Based on the above rate equation and under the assumption of quaslsteady state, the material
balance for oxygen Is given by the following expression:
476
-------
with the following boundary conditions
el-.,
and
^f = S (1 - .- ) at ; = -,
d . Sho s s
cxp
R T
(4)
(5)
Equation (3) in conjunction with tin- boundary conditions of equations (4) ;ind (5) can be solved
to yield the expression for the concentration of oxygen at the reaction surface. Thus
= 1 + -^ (1 - ~) uxp
tj
R T
; exp )'- (1 - -V-)
/R T Lc
(h)
' Sho's 'c
Similarly, the heat balance equation with transient term Include') in the analysis is given by
A .
(7)
witli Lin: following boundary riinUitlous
and
The initial condition is given by
R T
!,' ซ U =1 at = 0
(8)
(9)
(10)
In the above equations, the core is assumeu to maintain a uniform temperature, and the walls of
the reactor are maintained at the. bed temperature T . The pressure is maintained constant throughout
tiie reactcr. The ideal gas law is assumed and the effective, diffusivlty of trie oxidaiu through the
ash layer as well as the effective heat conductivity of the ash layer arc assumed to be proportional
to the temperature.
The rate of conversion, X, of the solid recctant is already >:iven by equation (J). Since for
spherical particle, the conversion is related to the system parameter /.^, as
X - 1 - -.' (11)
The expression for the rate of change of '. follows directly from equation (2). Thus,
, : u
X^ vป i ? /1 1 i
i exp i 11 - j
4 ,-. / R T c
with the following boundary condition:
1 at 0 0
(13)
477
-------
Equations (12) and (13) in conjunction with equations (6) to (10) arc to be solved simultan-
eously to Ret the conversion-tine relationship as well as the interior temperature of the particle.
Obviously, Because of turbulence in the fluidiztd bed, none of the feed particles are going to remain
in contact with the incominK Rases and get completely reacted at one stretch. Each particle may get
only a fraction of the total tine required for complete reaction, to remain in contact with the in-
coming gases. Therefore, it is essential to solve the above equations for the core temperature us a
function of time.
The following relations are also required for the solution of the problem:
Volume; of ash formed /1 A \
' * Volume of coke consumed
f.* - Z + (1 - Z)C* (15)
In the fluidized bed, the solid particles are mostly in the emulsion phase where the gas flow
is generally laminar. Therefore we ran approximate the heat and mass transfer crefiicients for the
single pellet to correspond to the laminar regime and employ the following simple expressions for Nc
and S.. : " 10
NuO
K: K:(N )1/3(sB fl)1/r
H-. - 71 + 2 hc ... ROฐ (16)
Sho !. f 1/2
''a
and
"NuO f, T fl/2 Vi"
vliere N_ , N and N are evaluated at the constant bed temperature T and initial particle size R .
Since the bed temperature in the present case is ma intalned constant, equal icns (l*ป) and (17)
can be further simplified as
N - 3- * 3- (18)
s
and
K' K
1 4
NNuO * Z^ * fl/2 <19)
's
where
and
The following notations are used in t.he above equations for simplification:
NNu - 2hR/k (23)
Nu " OR T3/k (24)
NuR oo
T0ke(To)P8
478
-------
2 p (-'.H) D (T )
ro eo
T E,
o 1
(26)
4 C M T
pc s o
(-.-.H)
(27)
R T R k (T }
o o s o
2 D (T ) M
(28)
and
i. R
s o
4k (T )p
s oo
(29)
t/i
(30)
DISCUSSION
The above heat and mass balance equations have been solved by numerical method for the range, of
values of the various parameters of practical Interest. The rate data of C-0 reaction reported by
Dutta and Wen6 have been utilized In the solution of these equations. The size oฃ coke particles
rarising from 2.5 cm to 0.006 <-m has been used. The ash content of the coke is varied from If) percent
to 20 percent which corresponds to the values of Z from 0.1 to 0.2. The partial pressure of oxyp.cn
Is varied from 21 percent as it occurs in the air upto oxygen enrichment of 42 percent. The average
temperature of the fluidized bed is varied from 900'C to 1000'C. Tho value of K. is taken as unity.
The other parameters viz., K^, K^, ;, C, A, .-.and N%. vary with the change in transport p.-opertles due
to temperature and with the change in the size of If lie particle. K and K vary very little when
temperature is changed from 900-'C to 1000{'C and hence constant values h,we been used for both of these
parameters at these temperatures. In Tables 1 through IV, only those values of conversions are listed
which yield core temperatures close to ash fusion temperature of approximately 1200JC. The calcula-
tions have been found to be Insensitive to the value of dlDenslonlesu parameter A, which depends upon
the values of the various physical properties of the reacting system.
Table I.
Internal Core Temperature for an Initial 2.5 cm Coke Particle.
(K. 1.0, K, - 2.0, K ซ 0.57 and N.. u ป 5.0)
T p Z
0 *0
ฐC atm
1000 0.21 0.1
1000 0.21 0.2
900 0.21 0.1
900 0.42 0.1
900 0.21 0.2
4 i. A G t
- sec
315 0.024 40 0.86 2000
3300
5500
6900
315 0.024 40 0.86 2000
3600
5400
7600
102 0.02 43 0.79 3400
5600
7000
102 0.04 86 0.79 990
1070
2700
3450
102 0.02 43 0.79 5500
7600
9000
X
percent
70
80
90
95
60
70
63
90
80
90
95
70
80
90
95
80
90
95
T
"C
1020
1130
1300
1420
1100
1175
1275
1410
1000
1160
1285
920
1090
1400
1600
1130
1260
1350
479
-------
Table !I. Internal Con- Temperature for an Initial 0.62 en Coke Particle.
(Kj = 1.0, K3 = 1.0, K4 - 0.3 .ind N.,uf. = 1.25)
T p Z : .-.AC l
0 ro
^C atm - - - - - sec
1000 0.21 0.1 79 0.024 40 0.86 225
300
480
600
900 0.21 O.I 25.5 0.02 43 0.79 230
325
490
610
1
Table III. Internal Corn Temperature lor an Initial O.Ofc rm Coke
(K, * 1.0. K, = 0.3, K. =0.1 .ind Nu _ = 0.125)
1 3 ** ."iiln
T p 7. : .' A C t
o o
'JC aim - - - sec
1000 0.21 0.1 7.9 0.024 40 0.86 5.0
7.3
8./
900 0.21 0.1 2.55 0.02 43 0.79 4.0
5.9
8.4
Table IV. Internal Core Temperature for an Initial 0.006 c:ra Coki
(Kj = 1.0, K3 ป 0.1. K/( * O.O3 and NN(|R = 0.013)
T p 7. : .1 A C t
0 0
';C aim - - - - ser
1000 0.21 0.1 0.79 0.024 40 0.86 0.113
0.14
0.15
1000 0.42 0.1 0.79 0.04ป 80 0.8f> 0.038
0.050
O.U67
0.078
1000 0.21 0.2 0.79 0.024 40 0.86 0.107
0.128
0.160
0.177
900 0.21 0.1 0.255 0.02 43 0.79 0.22
0.256
0.266
900 0.42 0.1 0.255 0.04 86 0.79 0.118
0.130
0.141
900 0.21 0.2 0.255 0.02 43 0.79 0.207
0.23t>
0.276
X
percent
70
80
90
95
70
80
90
95
Particle.
X
percent
80
90
95
70
80
90
Particle.
X
percent
80
90
95
70
80
90
95
70
80
90
9r>
80
90
95
80
90
95
70
80
90
Tc
rjc
1050
1100
1250
li/0
920
985
1105
1190
T
C
'C
1090
1215
1310
901
970
1060
T
C
'T.
1070
1160
1230
1005
1115
1275
1430
1085
mo
1240
1290
3iO
1030
1120
980
1100
1150
960
1000
1055
480
-------
Some of the qualitative trends that follow fron the analysis of data presented In Tables I
through IV and which favor the creation of high core temperature and thereby proMotu agglomeration
are: the large size of the coke particle, the high ash content of the coke, the high partial pressure
of oxidantt or the high average bed temperature. It therefore follows that depending upon the
properties of the given coke feed stock, different conditions may be necessary in order to agglomerate
the ash. For example, If the feed contains large particles, the ash a/glo:rerat<-ซ; ran be produced at
relatively lower bed temperatures. When fines are fed to the reactor, it requires higher bed tempera-
tures and or higher oxidant concentrations to promote, the formation of ash agglomerates. Based on the
results contalmJ in Tables I through IV, the following rvcomraendations are offered for the operation
of ash .iggloDfr.it ing oor.bustor or gasifler.
The contact time of the coke particle with the oxldant Is an Important factor in ob:.lining
sufficient conversion and bunco suf f ic lont ly high core temperatures to fuse the ar.lt. Generally
In the fluidized bed combustor, the oxidation zone is confined to a very narrow band above tile vrld
due to very high rat-.- of carbon-oxygen reaction. If the ratio of the superficial velocity In the
reactor to the minimum fluidization velocity of the given feed stock Is low, t!iซ" segreeat ion yf the
particles occurs, Saxena a- i Vofe,.;'. The large particles settle down at the bottป,-n and henre tin-
smaller particles get a very little chance to react with oxygen. Therefore, in order to agglonerate
a feed stock of mixed particles of a wide size distribution, caution oust be exercised Iti maintaining
3 good solids mixing in the bed so that all size particles get an equal chance to re.let with oxvgen.
Good solids circulation also helps In controlling and limiting the overall size of the agglomerates
produced. If the agglomerates grow too large, they are not easily removed from the bed and cause
local defluidlzatIon of the bed resulting in hot spots and eventual clinkering of the entire bed.
Although the time required for complete reaction of 0.06 cm particle is small, it is not
always possible for such a particle to renain in contact with the react.inl gases even for such snail
periods because of either segregation or elutriatifjn. This would imply that the presence of large
particles would promote ash agglomerat ion. Ash forsk.-d on small particles can stick to the fused ash
formed from larger particles and thus sustain ash agglomeration. When the elutriated flues are
recycled, these should be injected where the oxygen concentration is maximum and the rate of injection
should nlhii be sufficiently low In order to allow maximum contact time between the fines and the
oxvgen.
In many gasification processes, the steam is added to Inprovo the quality of gas being produced.
Since the carbon-steam reaction is endothermic, it may have considerable effect on the core 'emperature,
In such cases it may be necessary to attain ash fusion temperatures either bv raising the average bed
temperature or by Increasing the partial pressure of oxidant In the react.inl gases.
COMIY.RISON WITH K;T DATA
Some of the data obtained on the IOT ash agglomerating gasifler are presented in Table V. The
average particle size of the feed, the pattljl pressure of oxygen in the <:as feed, and the fluidized
bed temperature are important paraneters and are recorded in Table V. It has generally been observed
that when the average size of the feed is Increased, It requires lower bed terrer.-.ture to agglomerate
ash If the other parameters such as partial pressure of oxygen and the ash content of the feed stock
are maintained constant. Also when the fines are recycled Into the bed, thereby decreasing the
average feed sl/.e, the agglomerates could only be produced by increasing the bed temperature by about
60'JC and almost doubling the partial pressure of oxygen In the agglomeration zone. With FMC char,
which contained particles smaller than 40 mesh (0.042 cm), no ash agglomerates could be produced
either at the bed temperature of 10?S"C or the oxygen partial pressure of O.4 atmosphere because it
Is not possible to achieve the ash fusion temperature within these particles. Further* because of
turbulence In the bed .ind the sire of the particles, none of these partlclos stayed within the oxida-
tion zone long enough to attain ash fusion. The result* suggest that as the size of the feed particles
approaches zero, the bed temperature should approach the ash softening temperature In order to
agglomerate the resultant ash. All these observations are completely consistent with the predictions
of the- analytical model proposed here and the conclusions drawn on its basis as discussed in the pre-
vious section.
CONCLUSIONS
The mechanistic model proposed here for the agglomcratli n of ash In a fluidized bed combustor
or gasifier suggests that it Is possible to agglomerate ash at bed temperatures much below the ash
fusion temperatures- The factors that control the agglomeration process are the size of the coal,
oxidant concentration in the reactant gas, the average bed temperature and the ash content of the
feed. The presence of endothermic reactions such as carbon-stean reaction nay inhibit the ash
481
-------
Table V. Lxper irae-.ital Data from ICT Ash Agglomerating Gasifler.
(u = 5 cro/sc-c. Tr , =1183 "O
fusion
Feed Avg. Size
cm
Coke 0.170
Breeze
Coke 0.085
freeze
Coke 0.085
Breeze
FNC 0.022
Char
T* p
C aim
1000 0.21
1040 0.21
lOftO 0.40
1070 0.44
Fines Remarks
retboved tgxlomerales were
removed agglomerates were
increased T
o
increased p
torm-d
forced with
removed no agglomerates were formed
Minimum bed le. pcralure for agglomeration.
agglomerate: formation by retarding the temperatures developed within the coal particles during
combustion. The proper design of the reactor Is crucial to segregate agglomerates from the rest
of the bed material so that these can be selectively removed. Enough turbulent . should be provided
not only to control the size of the agglomerates but also to provide opportunity to the jnaller coal
particles to react and participate in the agglomeration process. When the fines are recycled into
the reactor, enough residence time should be provided within the reaction zone to consume them.
ACKSO'.n.EDCEMKIJT
This work is supported by the United States National Science Foundation. Grant So. KSC 77-01780,
under a joint USA-USSR research program.
NOMENCLATURE
A - Dlmensicnless quantity defined by equation (25)
C ** Specific lieat of bulk gas at constant pressure
C Specific heat of coke
DC '
C Specific heat of ash
D Molecular diffusivity of oxvgen in the bulk gas phasi>
D Effective diffusivity of oxygen In Che ash layer
E. Activation energy
C - Dlmcnsionless quantity defined by equation (27)
h Convectlve heat transfer coefficient
111 Heat of reaction per mole of the reactant and is negative for exothermic reactions
k Thermal conductivity .'f the bulk f.as
k ซ Effective thermal conductivity of the ash layer
kn Mass transfer coefficient for oxygen across the gas film
ks - Rate constant b.'iHed on the external surface area of the reacting solid
Kj A nuffierical constant which occurs in the correlation of Shrrwood and Nusselt numbers
K, - K,/2
K. ป A numerical constant which occurs In the correlation of Sherwood and Nusselt numbers
482
-------
K- - K2/2
K.. ซ Dinensionless quantity defined by equation (20)
K, 3 Dimensionless quantity defined by equation (21)
M ป Molecular weight of coke
S., N'usselt nunber 2Rli/k
Nu
N o " DIocnsionKss quantity defined by equation (22)
X ป Dinenslonlcss quantity defined by equation (24)
NuK
N * Prandtl number ซ C (../k
N_ a Reynolds number ซ* 2uR, /;.
NR!O ' N>RO(RO/R)
NS<; ป Schmidt sber = '../. D
N . ป Sherwood nunber - 2Kk /D
Sn ci
NSho -NSh(V2R>(D(V/DefToป
p * Partial pressure of oxygon
p * Value of p at the unre.'irlcd trore S'irface
p a Vnluv of p in the bulk jt.'is
p, ป Value of p at the outer surface of tlie partlclv
r = Kadi.il distance fron tlic center of the spherical particle
r ซ Rndius of the unreacted core
c
R ป Particle r.idius
R( ป Initial particle r-.dius
R ฐ Gas Constant
S ป Spc-clfit: external surtace area of tlit- particle
*cx
t Time
T Ti-mper.riture
T. - Temperature of tlie unre.u-ted i-ore
T * Average bed temperature
T Particle surface tenperature
u ซ Flow velocity of bulk g.-is
U Reduced temperature T/T(>
U ป Reduced core temperature ซ T /T
c c o
U - Reduced particle siirfaru tenperature, T./T
X * Conversion of solids defined by equation (11)
Z " A parameter to characterize Rrowth or shrinkage pf the particle, defined by equation (14)
Creek Letters
r. B Dimcnsionless quantity defined by equation (26)
'! - Reduced time defined by equation (30)
11 Viscosity of bulk gas
t. * Reduced distance ซ r/R
f. ป Reduced core radius of the particle = r ,/R
Reduced size of the particle K/R0
v * Density of the bulk gas
i1 Density of the solid react.int
4C3
-------
C'l.ir.ictorlst ir tin.- m (2'tt
, - Dlni-nsl'inlrss "Mantlf/ dellni-1 by ซ"j.:/it Inn '2>i)
ป ki-durfd valui- of ;> ป (p/p >
Ri-d.in-d valui- r,f p_ . fiV/P0>
ป Ri-duot-d value* nf p a (p _/p >
: - R.idl.ilivr li.-.it tr.insfi-r < o,:t I i< |.-i.t
KKH-IRKNCKS
I. I., .li-qulcr. I., l.onj-t-Ik-irnhnn nnd (J. V;m Ih- Putn-, .1. Insi. Fui-1 TJ. No. 12. 19'>0, 584.
/. W. A. S.-ind'tlrtin. A. Rftira.-it .-m:l W. f;. li.ilr. "Tin- C:isi f Ic.lt Inn of Oi.il Cluirs In .-i Fluid I iri-d-Rc-d
Af.li-AKr.ltimiT.it Inn ^;isifi*rt" I'.'ipor pr<-sซ*ntefl /it tin- *>9th Annual Hcfllnf; of the A.I.Ch.K.,
Cl.li-.is'.. UvstinKhoiiHi.- KliTtrli: Corporation. "Adv.mrcd Coal Casl f Irat Ion Systi-m fnr EK-ctrlr Power
r,.wrat Ion." Monthly l'roi>ri-iis Rซ'port fnr Si-ptcmbi-r 1977. prfp.-ired for U.S. EntTKy Ri-si-artli
and OfVirlopncnt Administration, tindi'r contrart No. KF-77-C-01*! *il4.
>. A. Rt-hn.it and S. C. Saxi-na. Ind. Knป;. (;h.-n.. Pron-ss l>-s. I)ซ-v.. IS. 197'i. Tซl.
h. S. Dulta and r Y. Ui-n, Ind. Kn^. Clii-ra. Pr-n-i-ss l)t-s. l>vv., 16. 1977, 11.
7. S. C. Saxfiia .. d (i. .1. VnKr 1 . Chfra. Knc.. J. I''. 1977. 'j9.
484
-------
QUESTIONS/RESPONSES/COMMENTS
FREDERICK HANZALEK, CHAIRMAN: Dr. Saxena?
DR. SAXENA: Okay; I will now try to answer the question that
has been asked me by Dr. Haldipur from Westinghouse Electric. If
I have pronounced your name wrong, please excuse me. He asks: "Does
your model neglect fluid dynamics effects in the reactor?" In the
Westinghouse agglomerator, we have tested fully agglomerated FMC
chars which were identical feedstock to the ones used by IGT. A
very good question.
No, we have not neglected the fluid dynamics, but we have taken
its simplest form for our present illustrative purpose. It is assumed
that the particles are suspended in the fluid which flows around it in
a laminar regime. As a result, the correlations developed by Ranz and
Marshall for gas film mass transfer and convective heat transfer
coefficients have been employed in the calculations. It is also
implied that the particle stays continuously in contact with the gas
which maintains as constant oxygen concentration equal to its initial
value. This ideal requirement will not be met in a real fluidized bed
and, therefore, the calculated core temperatures may differ from the
actual temperatures developed in the particle while in the fluidized
bed. However, in order to agglomerate ash from coal, all the efforts
should be made to achieve gas-solid contact conditions as close to the
ideal us possible.
The Westinghouse agglomerating gasifier, unlike the Institute
cf Gas Technology gasifier, operates at a relatively higher oxygen
pressure. Consequently, the amount of oxygen available per unit
carbon in the Westinghouse gasifier is considerably larger than the
Institute of Gas Technology gasifier. As a result, the Westinghouse
gasifier is capable of generating higher temperatures within the
particles than the Institute of Gas Technology gasifier. These
temperatures will also depend upon the scheme of introducing steam in
the gasifier, an aspect we have not considered at all in this work.
We propose to undertake such calculations and will report our findings
at some later date.
Then, there is a second question: "Could you comment on the
effect of the agglomeration of the insoluble iron content in the
ash?" Only in general terms. The composition and nature of ash
present in the coal influences the calculations in a sensitive fash-
ion. The intraparticle diffusion coefficient which controls the rate
of oxygen transport to the surface of the reacting coal particle will
depend upon the physical properties of the ash through porosity. The
thermal conductivity of the ash also appears in the calculation and
all of these properties will be controlled by ash decomposition.
485
-------
So to the best, what we have done, actually, is provided material
for thought.
KARL BASTRESS, CHAIRMAN: Thank you. Is there anyone who
submitted a question that did not get answered, or does anyone have a
single burning question of general interest? (Laughter) Please
introduce yourself and make your question brief, please.
MR. YERUSHALMI: The gentleman from Westinghouse made me inter-
ested. Oh, I am Joseph Yerushalnii from the City College. Could the
gentleman from Westinghouse describe the shape and the othet physical
characteristics of the ash agglomerate they made from FBC cfur? And
number two, does he agree with Professor Saxena that the surf.-ce
temperature of the char does not exceed the temperature of the bed,
the surrounding bed?
MR. HANZALEK: 1 am afraid I may have let this get out of control.
Go ahead. Please introduce yourself.
MR. HALDIPUR: Golan Haldipur from Westinghouse Energy Systems
Operations. I will answer the last question first, because I remember
it. We don't have any means of measuring radial temperature distribu-
tions in the particles in the fluidized bed agglomerator. We have
attempted to measure actual temperature gradients in our reactor,
but we haven't as yet got a thermocouple which could penetrate our
agglomerator and determine the hottest zone in the reactor. And
regarding particle temperature distributions, it's virtually impos-
sible in our reactor to measure it. The diameter of the reactor
is 20 inches, and we have a high-pressure reactor, 15 atmospheres,
which is different from the IGT reactor. I don't think there is any
more to add on this, unless you have a question.
MR. YERUSHAMI: What was the shape--
MR. HALDIPUR: Oh, the shape. Okay. We ran three feedstocks
from the FMC coal, the Western Kentucky, and the Utah; and we also
ran coke breeze. When we ran a combination of FMC and coke breeze.
we got sub-angular particles, and when we ran the FMC-Utah char, we
got well rounded particles. The size was approximately about 500
microns, and the angular particles were about 800 to 1,000 microns.
MR. BASTRESS: I think the speakers will be here for a few
minutes if you have specific questions, and we will devote the rest
of the time to extending the lunch hour. Fred and I would like to
thank the speakers in particular, as well as the projectionist, for
helping us close this session on time. Thank you very much for your
interest.
486
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INTRODUCTION
FREDERICK HAN7ALFK, CHAIRMAN: Our next paper will be presented by
Dr. David Berkowitz of The MITRF Corporation. It is entitled, "Dynanic
Modeling, Testing, and Control of Fluidized Bed Systens." Dr. Berkowitz
received his Ph.n. in physics from MIT. He brings impressive creden-
tials in that he organized the Seminar in Boiler Modeling, held in
November of 1074 and v/as subsequently editor of its Proceedings. He
first joined the MITRF Corporation in 1%1, and is now the firoup Leader
o^ Power Plant Dynamics and Control. Dr. Rerkowitz.
487
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Dynamic Modeling, Testing, and
Control ot Fluidized Bed Systems
D. A. Berkowitz, A. Ray,'/. Sumaria,
and M.Wilson
The MITRE Corporation
Bedford, Massachusetts
ABSTRACT
A nonlinear dynamic modeling technique for fluidized bed combustion (FBC) sys-
tems has been adapted to the Alexandria Process Development Unit and validated in a
steady state and transient response test program. Datr. from all signal channels were
sampled every few seconds, and are available on magnetic tape. The modeling technique
can be adapted to other FBC plants to evaluate their design and operation, and to ana-
lytical design of their control systems. Several frequency and tine domain methods are
discussed, including classical single-input single-output control, optimal regulators
incorporating a performance index, and decoupled multivariable control using state
feedback.
INTRODUCTION
The purpose of this paner is to suggest initial control development guidelines
for fluidized bed combustion systems. Control system development requires a detailed
understanding of the process, p.-jrticularlv its transient response characteristics as a
function of operating load level. If one's understanding of the process is embodied in
an analytical process description, or sr.-callcd process model, then control systems can
be developed analytically. The methods considered here involve development of a vali-
dated process model as a critical first step, followed by its subsequent use for con-
troller design. This technique - modeling, validation, controller design - has been
successfully applied to solving performance problems of existing conventional electric
generating units. It has been used in specification of new units and simulation of
their performance. Ard it has also been applied to fluidized bed combustion, although
control design in that case is far from complete.
The paper is divided into three parts. Part I (Modeling) discusses purposes for
modeling and the type of modeling method selected for fluidized bed systems. Part II
(Testing) describes data acquisition requirements for model validation and an experi-
mental program at the Alexandria Process Development Unit, in which measured data is
compared with predictions of the Alexandria process model. Part III (Control) deals
with the manner in which the model is used to explore process characteristics and to
develop control system strategies.
PART I: MODELING
Why Model?
In the case of a large steam-elecrric generating unit, design and integration of
its many subsystems is a complex process involving several organizations: the owner-
utility, architect-engineer, boiler manufacturer, turbine manufacturer, control system
manufacturer, and other major component suppliers. To increase productivity, greater
attention must be focussed on systems design and integration to insure that major com-
ponents function together, within specifications, under normal steadv state and tran-
sient operating conditions. This nned has been recognized for conventional nuclear and
fossil fueled generating units [l,2] . There is no less a need in the case of fluidized
bed power plants - perhaps a greater need, because the technology is less mature; the
process less understood.
The work described in this paper deals with extension to fluidized bed combus-
tion systems of what a utility company calls "Design/Operation Evaluation Models" [2].
These models permit:
(1) Increased understanding of the process, including subsystem interaction and
integration;
(2) Improved design and analysis capability for plant control systems to ensure
that major components arc maintained within design limits during normal
488
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steady state and transient operating conditions-.
(3) Verification of desired performance capabilities for different modes of
plant operation: and
(ft) Pre-start-up simulation capability for development of operational procedures
,->nd resolution of operational problems.
Models are not new. Modeling ond analytical capabilitv does exist in the indus-
try and is customarily used (1) as part of the basic plant de-sign function, (2) for
nuclear plant safety, and (3) for operator training. Various models which have been
developed differ in their objectives, the extent to which they include different plant
r.ystems. and the level of mathematical detail required to neet the objectives. A de-
sign code developed to predict performance details \.-hich influence pipe sizing, mate-
rial selection, etc.. requires a high level of mathematical detail, but it's limited
to the boundaries of only one component. A safety related design code requires signi-
ficant mathematical detail in those systems which influence plant safety; other systems
are not included, or only included superficiai.lv. A training, model disnlays to the
operator hov the plant will perform under specified transient conditions, which re-
quires the least mathematical detail, but covers the greatest number of plant systems.
A Design/Operation Evaluation Model is intended to represent plant operation
(particularly under automatic control) during normal plant operating conditions when
safetv is not an issue. This model includes more sv;tems than a safetv code and less
than a training simulator; but it's: more detailed than a training simulator and less
detailed than a safetv code.
TraditionalIv. nower plant development has occurred bv trvins things out, then
fixing them latr.- [2]. This process has resulted ir. generating units that perform ade-
quately, rather than ontimally. and utility personnel end un accenting numerous operat-
ing peculiarities as inherent performance 1 imi t.-itiors. Kxisr.ir.g design approaches need
to be supplemented with systems oriented analysis techniques, such as Design/Operation
Evaluation Models, to assure higher plant availability and .performance.
.Modeling
The method most suitable for Design/Operation Evaluation Models is deductive
(sometimes called .? first principles method), based on fundamental phvsical laws for
energy, mass, and momentum conscrvat ion. and well kr.own empirical correlations for
material properties, heat transfer and flow coefficients. Models of this tvpe have
these advantages:
(1) All model parameters can be determined from available design and steadv
state data;
(2) Model structure is independent of available data:
(3) The model provides insight into the nature of the process that is so vital
in design of new systems, particularly when scale up from pilot systems is
required; and
(4) It establishes a rational basis for increasing or decreasing model complex-
ity.
Mathematically, a deductive process model is: r.on-1 ine.ir. tine-invariant, de-
terministic, continuous-tirn . and in state-snace forr.. The actual process generally
consists of distributed parameter dvnamic elements mathematically represented bv non-
linear partial differential equations with snace anci time as independent variables.
To obtain numerical solutions of these equations bv digital computer, the partial dif-
ferential equations are apprximated hv a finite set of ordinarv nonlinear differential
equations with time as independent variable. Thus, the fundamental modeling assumption
is that a distributed parameter process can be approximated by a lur.ipcd parameter mod-
el. This approach has been shown to be adequate in the simulation of fossil and nu-
clear power generating units at Philadelphia Electric Companv (Cromhv Station I'nit No.
2) [3]', Boston Edison Company (New Boston Station irr.it No. 2) [4], and Public Service
Conipony of Colorado (Fort St. Vrain Unit Nc. 1) (>) . and has also been shown to be
adequate for fluidized bed combustion at the Alexandria Process Development Unit [6] .
Additional applications of the approach by Philadelphia Electric Company include tddy-
stonc Station furnace implosion studies. Peach Bottom Reactor feedpump start-up analy-
sis, and plant dynamic simulation for the proposed Fulton Station HTCR unit, and those
bv Ontario Hydro include Nanticoke Station pulveriser control and deaerator control at
CANDU nuclear power plants [2] .
489
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It is not sufficient to create a large computer program simulating the deductive
model of a thermofluid process: the model must be validated. For a plant in the plan-
ning and design stage, validation is with respect to plant heat balance and other
steady state design calci. ."ions made by the manufacturer, which represent the most
that is known about the ' ..it prior to actual start-up. For an operating plant, vali-
dation can be with mear ^rec> data from steady state and transient response tests, as
well as heat balance esign cita. The process of reconciling model predictions with
actual measurements .an proviae deep insight into less well understood portions of the
process. Performar-e of the er.tire system becomes an important diagnostic in under-
standing process retails.
PART ii: TEST:NG
Data Requirei. V - for Model Validation
Experimer. . /allelation of a dynamic process model suggests the need for large
amounts of data, rded as a function of time, preferably available in machine acces-
sible form.
An important i'eafire of the data is its completeness. That is, the measured var-
iables must constitu:e : set of process outputs that permit an accurate, complete de-
scription of what went on. For example, there is little value in recording tempera-
tures every five seconds if air, water, steam, and solid flow rates are only recorded
manually every 15 or 30 minutes. There is little value in sampling gas analyzers con-
tinuously, if fuel and air flows are not known continuously, as well.
Frequency of observation is also important during start-up and transient opera-
tions, whether planned or unplanrซ?d. All measured variables should be sampled at a
rate several times faster than the fastest natural response time of the process during
the transient periods.
The type of data acquisition
-------
Figure 1. Block Diagram of the Data Acquisition System (DAS)
each channel.
The scan list for a typical run ar. Alexandria includes 76 process variables (36
temperature*. 10 flows, 13 pressures. 1 level. 3 control signals. 8 gas analysis sig-
nals, and 5 wtight signals), plus 18 channels for calibration and electrical check-out.
Data Reporting Tapes for nineteen test runs at Alexandria are now stored and can be
readily copied for anyone who is interested.
Alexandria PDU Results
Model predictions and test results were compared at different steady state load
levels, and for several transient conditions. Table 1 shows the comparison for two
load levels, corresponding to different values of bed level and coal firing rate.
In transient response tests, the plant is allowed to stabilize while maintaining
constant controlled input values. Then, a single input is shifted to a new value, and
data is recorded as a function of time until a new steady state has been reached.
Tests are repeated at different load levels for each controlled input. In this manner,
the ability of the model to predict the complex natural dynamic response of the process
is evaluated. Figures 2, 3, and 4 show the agreement between predicted model responses
and measured values of bed temperature, drum steam pressure, and steam flow, respec-
tively, for an 18-minute interval following an 8.22% step decrease in coal feed rate.
Bed temperature results (Figure 2) agree closelv with experimental data. For drum
steam pressure (Figure 3). model predictions appear faster than test data. This may be
due to inaccuracy in drum water level (that is, drum water thermal capacitance). Steam
flow data (Figure 4) fluctuates due to instrumentation noise, but the average profile
agrees closely with model predictions.
Figures 5, 6, and 7 compare model responses and measured data for bed tempera-
ture, drum steam pressure, and steam flow following a 6.757. step increase in main air
flow. For bed temperature (Figure 5), model results and test data are initially in
close agreement. About ten minutes after the disturbance, test results show an upward
491
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ro
8.22 Z STEP IN COAL FLOM
ป MCCEL RESULTS
O MEASURED VALUES
Figure 2. Bed Temperature Transient Response
Following 8.22% Step Decrease in Coal Feed
Ran
.22 Z STEP IN COAL FLOM
ซ MODEL RESULTS
O MEASURED VALUES
4 6 e 10 12 ii iซ is
Tine IN MINUTES
Figure 3. Drum Steam Procure Transient Re-
iponse Following Step Decrease in Coal Feed
Rate
-8.22 X STEP IN COAL FLOM
* MODEL RESULTS
-------
TABLE I Comparison of Steady-state Model Results with Test Data
Bed depth 0.508 m (20 in)
Process Variables
Model results
Test data
Bed Temperature
Drum steam pressure
Main steam flow
Coal feed rate*
Limestone feed rate*
FD fan air flow
Carrier air flow*
884.4 C(1624 F)
0.7623 x 106 N/m2 (110.7 psia)
0.3404 kg/cec (0.7504 Ibn/sec)
0.0857 kg/sec (0.189 Ibm/sec)
0.0367 kg/sec (0.0809 lbm/~ec)
0.5770 kg/sec (1.272 Ibm/sec)
0.0506 kg/sec (0.1114 Ibm/sec)
882.2 C(1620 F)
0.7653 x 106 N/ri? (111 psia)
0.3402 kg/sec (0.75 Ibn/sec)
0.0857 kg/sec (0.139 Ibm/sec)
0.0367 kg/sec (0.0809 Ibm/sec)
0.5783 kg/sec (1.275 Ibm/sec)
0.0506 kg/sec (0.1114 Ibm/sec)
Process Variables
Bed depth 0.304 m (12 in)
Model results
Test data
Fed Temperature
Drum steam pressure
Main steam flow
Coal feed rate*
Limestone feed rate*
FD Fan air flow
Carrier air flow*
920.56 C(1689 F)
0.7543 x 106 N/m2 (109.4 osia)
0.3329 kg/sec (0.7339 Ibm/sec)
0.0844 kg/sec (0.186 Ibn/sec)
0.0367 kg/sec (0.0809 Ibm/sec)
0.516 kg/sec (1.238 Ibm/sec)
0.0506 kg/sec (0.1114 Ibm/sec)
918.33 CU685 F)
0.7584 N/m2 (1110 psia)
0.3311 kg/sec (0.73 Ibm/sec)
0.0844 kg/sec (0.186 Ibm/sec)
0.0367 kg/sec (0.0809 Ibm/sec)
0.5171 kg/sec (1.240 Ibm/sec)
0.0506 kg/sec (0.1114 Ibm/sec)
-Given input to the model
drift due to a small, inadvertent increase in coal feed rate, which was confirmed by a
decrease in flue gas oxygen concentration. This drift is obvious in drum steam pres-
sure response (Figure 6), but cannot be identified in steam flow response (Figure 7)
because of noise. Following the main air flow increase, both model results and test
data show an initial increase in steam pressure (Figure 6) which then relaxes to a low-
er value. The increase occurs because of an increase in convective heat transfer co-
efficient in the freeboard region. Later on, as bed temperature decreases, flue gas
temperature drops reducing heat transfer. Test data show a larger increase than the
model which is probably associated with some inaccuracy in formulating model equations
for overbed combustion.
PART III: CONTROL
Analyzing the Model
The nonlinear process model in state-space form can be analyzed by sceci.il pur-
pose codes to find steady state operating conditions, to derive linearized nodels at
any load condition, to deteimine system eigenvalues of the linearized models and fre-
quency response plots for any input-cutout pair. etc. For example. Table II shows
system eigenvalues of the linearized Alexandria PDU model at steady state conditions
before and after the 8.227ป decrease in coal feed rate. The eigenvalue magnitudes de-
crease with reduction in coal feed rate indicating that, at lower firing rate, the
process slows down. The smallest eigenvalue is strongly associated with thermal re-
laxation of the fluidized bed and corresponds to a tine constant of approximately 220
seconds before the change and 235 seconds after the change, which illustrates the ef-
fect of process nonlinearity. The -0.100 eigenvalue corresponds to an assumed 10-sec-
ond time constant, which is'related to average coal residence time in the bed during
combustion. The largest eigenvalues, corresponding to time constants of approximately
1.9 and 1.3 seconds are strongly associated with waterwall metal temperatures in the
freeboard region and in the fluidized bed, respectively. These states are fast be-
493
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cause of the very high rate of heat transfer through the metal walls.
Frequency response plots of bed temperature, drum steam pressure, and steam
flow rate with respect to coal feed rate are shown in Figures 8, 9. and 10, respective-
ly, for the plant conditions existing before the coal feed rate disturbance. Figure 8
shows that the transfer function of bed temperature versus coal feed rate can be ap-
proximated by a simple structure of two finite poles and one finite zero. The plot of
magnitude versus frequency shows that process gain attenuates rapidly with increasing
frequency. Thus, a proportional-integral-derivative (P-I-D) controller would result
in fast response, closed-loop action. The transfer function inferred from Figure 8 is
useful for single-input sinj?' e-output controller design. Though the fluidized bed sys-
tem is interactive and multivariable, single-input single-output controller design
methods may be useful as a first step.
Figure 9 shows that the drum pressure versus coal feed rate transfer function
can be approximated with 3 finite poles and 1 finite zero, and Figure 10 shows that
the steam flow versus coal feed rate transfer function could be represented by two fi-
nite poles and 1 finite zero. In these two cases, also, single-input single-output
controllers (possibly P-I controllers) can be designed with the aid of standard algo-
rithms. Frequency response plots provide some insight into the fluidized bed process
dynamics.
Alternative Controller Design Methods
One major objective of state-space modeling is analytical controller design and
its evaluation by digital simulation. The plant model allows determination of a con-
trol law with respect to certain specified performance criteria, as well as input and
state constraints. Various controller design methods are available. Four of them are
reviewed briefly in this section with some technical depth, with respect to commerical
scale fluidized bed power plant applications. These methods can be classified as fol-
lows :
(1) Design based directly on the nonlinear model;
(2) Time domain design using a family of linearized models;
(3) Frequency domain design using a family of linearized models; and
(4) Design on the basis of reduced order models.
Method 1: Design based on nonlinear model
Exact optimal controller design solutions for nonlinear systems with a quadrat-
ic cost functional is very complex and extremely difficult to implement [7] . It in-
volves 2-point boundary value problems. One powerful solution method that has been
employed for low order nonlinear models is "continuation" or "imbedding" [8] . In this
method, a parameter or a set of parameters is introduced in the original problem such
that a class of similar problems is obtained with the property thar the solution to
one of them is known. By successive variation of the imbedding parameter(s), the de-
sired solution can be obtained from the known solution. Jamshidi applied an imbedding
approach to control system design for a 6th order power system model' where solutions
of 2n(n + 1) = 84 differential equations were necessary [8]. For larger order system?.
this approach is too expensive for practical application.
TABLE II. r.igenvalues of linearized model before and
after -8.227. change in coal feed rate
Before After
-.00451 sec"
-.00642
-.0103
-.0168
-.100
-.535
-.780
-.00426
-.00641
-.0103
-.0162
-.100
-.524
-.764
sec"1
494
-------
Wl
100
so
-so
-too
-ISO
-so
-no
J0-l JO'1
10' 10*
IN nno/scc
JO'' JO'1 10' JO* 10*
u *7
S
-ISO
-Z25
-300
g -ซ
-180
-J60
to-' io-'\ to1 toป to1 2
eo
a jo
-30
-60
10"
rncoucNcr IN RBO/SCC
io-ซ ID-' ID' toป to*
FREQUENCY IN RRO/UC
10'* 10"' 101 JO* 10*
Figure 8. Frequency Retponse of Bed Temperature Figure 9. Frequency Reiponje of Drum Steam Prei- Figure 10. Frequency Retponte of Steam Flow with
with Rotpect to Coal Feed Rite turt with Rnpect to Coal Feed Ron. Roipoct to Coal Feed Rate
-------
Lccper and Mulholland presented a method for optimal control of nonlinear sin-
gle-input systems where no 2-point. boundary value problems need be solved, and closed
form solutions for optimal control can be obtained [y] . However, an extension of this
method has apparently nor. been achieved for multivariable systems.
It is felt thar controller de-sign for a cocmerci.i 1 scale fluidized bed power
plant by direct use of the nonlinear model is not a feasible alternative.
Method 2: Time donain with linearized node Is
The linearized model of a nonlinear process represents its local characteristics
around the operating point of linearization; a family cf 1 ir.car: ::eH -^Hpls at a number
of operating points approximates global characteristics of the nonlinear process o.er
the operating range.
McDonald and Kwatny designed a fossil fueled drum boiler plant controller fro:?
a nonlinear 1'tth order plant model linearized .it several operating points [10] . The
design algorithm was based o.. optimal linear regulator theury using a standard quadrat-
ic cost functional perfomanco index. To compensate for model inaccuracies, -i random
noise bias vector, whose derivative is white noise with ^ero mean, was introouced; the
a priori unknown bias became constant as the derivative noise approached zero. Linear
stochastic theory was applied to svnthesi.-.e a deternin: si ic optimal controller such
that the plant output tracks a constant desired value wi
-------
tton cime requirements, for example). During the last ten years. significant prepress
has been made in identifying low order models of large time-invariant systems [22-27] .
Yu and Siggprs demonstrated application of a reduced order linear deterministic model
to approximate a nonlinear, single generating uni-. infinite-bus power system model
(28) . A control system designed with the low order model was : hen shewn to be ade-
quate for the original high order model.
A deterministic high order nonlinear process can be adequately represented by a
family of low order linear stochastic models [29-32] . Tn a survey paper, Astrom and
Eykhoff have discussed various ways of choosing model structure in simple linear forms
and subsequent identification of model parameters [2?] . Moore and Schwenne formulated
a stochastic model of !:he nuclear reactor in a pressurised water reactor (PWR) power
plant [42] . A detailed high order analog simulation of the plant *:hat included n&n-
linear effects, disturbances, and measurement noise, was developed. Frcm data >ener-
ated by the analog plant simulation, a priori unknown parameters of a li-w order linear
stochastic model were identified by the maximum likelihood criterion. The reduced or-
der model was used for reactor controller design. This method is worthy of additional
exploration.
SIT-C'-ARY
The method of process modeling nd model validation loading to analytical con-
trol system design, which is finding increased application in conventional fossil and
nuclear generating units, can advantageously be applied to fluidized bed combustion
systems. The modeling and testing program at Alexandria has demonstrated that the
modeling method adequately represents process dynamic interactions, and information
useful for analytical control system design, such as input and output matrices of th-'
linearized system at several load levels and frequency dependent input-output transfer
functions, can be derived readily from the validated nonlinear prcccss model. In gen-
eral, fluidized bed systems are nonlinear, multivariahle. and interactive. Although
classical controllers based on individual single-input single-output control loops mav
he specified for such systems as a first step in formulating the controller structure.
alternate analytical control design methods seem more promising. Time domain dc-ign
using a family of linearized models and a quadratic cost functicnal performance index
has already been applied to gene-rat ing unit controller problems. Krequencv <'.onain de-
sign usir.g a family of linearized models and state feedback to achieve decoupling of
output variables seeT.s very appropriate, and offers a way to apply conventional single-
input single-output frequency response irethods more effectively.
ACKNOWLEDGEMENT
This work was supported by the U. S. Energy He-search and Development Adminis-
tration (Department of Energy). Coal Conversion and !_'t i 1 i ;:at ion Division, under Con-
tract No. EX-76-C01-24S3. The authors express their appreciation to R. Reed and the
Pope. Evans and Robbins staff for their pp'ience and help during the experimental pro-
gran at Alexandria.
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26. A. Kuppurajulu, and S. Elangovan, "System Analysis by Simplified Models", IEEE
Trans. Automatic Control, AC-15, 234-237 (1970).
27. T. C. Hsia, "On the Simplification of Linear Systems". IEEE Trans. Automatic
Control. (Correspondence), AC-17. 372-374 (1972).
28. Y. N. Yu. and C. Si&gers, "Stabilization and Optimal Control Signals for a Power
System". IEEE Trans. Power Apparatus and Systems PAS-90. 1469-1481 (1971).
29. K. J. Astrom, and P. Eykhoff, "System Identification - A Survey". Autcmatica ]_.
123-16?- (1971).
30. F. C'. Schweppe. Uncertain Dynamic Systems. Prentice-Hall (1972)
31. R. L. Kashyap, "Maximum Likelihood Identification of Stochastic Linear Systems",
IEEE Trans'. Automatic Control. AC-15. 25-34 (1970).
32. R. L. Moore, and F. C. Schweppe, "Model Identification of Adaptive Control of
Nuclear Power Plants". Automatica 9. 309-318 (1973).
499
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QUESTIONS/RESPONSES/COMMENTS
K HAMZALFK, CHAIRMAN; Pr. Rerkowitz.
np. RFRKWIT7: Thank you. I've qot one card and two questions
fron Ralph Wood of General Flectric. "What numerical methods do you
use for solving your nodel?" is the first question. (ซir model equa-
tions are written in CSMP; we qenerate an ordered Fortran load nodule
to simulate systen response as a function of tine. He give the node!
a set of initial values for the state variables, go down through all
the equations evaluating the process variables as well as rates of
change of the state variables, and then, integrate to the next tine
step. The systems are rather stiff, so we use variable tine step and
Fowler-Marten integration. To find steady-state solutions, we operate
on the Fortran load module with various analytical codes that we de-
veloped or that are available, that use Newton-Raphson or Steepest
descent techniques. Typically, that costs about one dollar for the
Alexandria or CTIll nodels. TO simulate a thousand seconds of CTIH
operation (the Model is 36th-order model), it takes typically about a
minute of CPU time, which costs approximately $?n.
The second question: "Is there any physical significance to the
oscillations that are superimposed on some of the data which you showed?"
It is difficult to control coal feed rate to a fluidized bed boiler,
and there is noise associated with the lack of uniformity of the coal
feed. I assume that these probably are mostly concerned about steam
flow, which was the one that oscillated quite a bit. There were also
oscillations in steam flow. Steam passes through an orifice with RP
cell, control valve, and a long pipe connected to the drum. Oscilla-
tions might very well occur in that line. There might also he some
transducer noise. Typically, on one day, steam flow signals are
noisier than on other days, so I suspect that the signal oscillations
do have some physical significance.
FRFPFRICK HAMZALFK, CHAIRMAN: Thank you.
500
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Participates and Removal
501
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Evaluation of a Granular Bed
Filter for Particulate Control in
Fluidized Bed Combustion
Melvyn S. Nutkis. Ronald C. Hoke,
Michael W. Gregory, and Rene R. Br.-rtrand
Exxon Research and Engineering Company
ABSTRACT
In a pressurized fluldized bed coal combustion system where the due gas Is expanded through a
gas turbine to achieve greater cycle efficiency, it is Important that the particulate emissions be at
a level satisfying both turbine and environmental requirements. A program in progress using the EPA-
Exxon Research pressurized fluidized bed coal combustor is evaluating the effectiveness of a granular
bed filler to remove particulars from the high pressure, high temperature flue gas. A 24-hour shake-
down run was accomplished in August 1977. Main problems are with filter cleaning at system conditions,
filter plugging and loss of filter media.
INTRODUCTION
The successful development of the pressurized fluidized bed coal combustion process is depen-
dent on the ability of particulate control devices to remove particulates from the hot combustor flue
gas to very low levels. This must be done to assure that the expansion of the flue gas through a gas
turbine doos not cause damage to the turbine '^y erosion, corrosion, or deposition of solids on the
turbine blades. Current estimates of the allowable particulate concentration in the flue gas entering
the turbine range from 0.04 to 0.001 g/m3 (0.02 to 0.Of 104 grains/SCF)l. To meet these estimated
requirements, the flue gas leaving a pressurized combustor must first be prccleaned in a two stage
cyclone and then sent to a third stage high efficiency device for final cleaning. The efficiency of
the third stage device must be in the range of 93 to 99.72: to be within the currently estimated parti-
culate loading target range.
In addition to the gas turbine inlet requirements, the U.S. Environmental Protection Agency has
Imposed limits on the emission of particulates from coal fired installations of 0.1 Ib/M BTU coal
fired. This translates to a particulate concentration in the flue gas of about 0.1 g/m3 (0.03 gr/SCF),
somewhat higher than the limit set by turbine requirements, li.^reforc, at the present time, removal
efficiencies are dictated by the turbine requirements. However, the environmental standards are
currently being reviewed and may be tightened in the future.
The present particulate removal program at Exxon Research and Engineering Company, which is
sponsored by the U.S. Environmental Protection Agency (Contract 68-02-1312), will tc.ปt two particulate
removal devices. The first device is a granular bed filter of a design developed by the Uucon
Company, ihe choice of the second device has not yet been made, but will probably be a high tempera-
ture bag filter or electrostatic precipitator.
The Ducon-type granular bed filter consists of a number of small beds packed with suitable
granular filter media such as alumina, quartz, etc. A stack of the filter beds form a single filter
element. A number of filter elements can be used depending on the volume of gas to be filtered. A
photograph of a filter element purchased from the Uucon Company for this program, is shewn In Figure 1.
This element consists of a total of twelve individual filter beds. Dirty gas passes through the
screen sections down into the filter beds immediately below the screen sections. Clean gas from the
beds is collected in a manifold in the interior of the element and then passes to the clean gas outlet
system. As the filtration step proceeds, the pressure crop across the element increases and eventually
the element muot be cleaned by the reverse flow of clean gas. This "blow back" occurs by flowing
clean gas in reverse direction through the outlet gas manifold, up through each filter bed and out
through the screens. The function of the screens is to retain the filter media during the blow back
step, keeping it inside the filter beds, while allowing the fine particulates removed from the filter
media by the blow back gas tc pass through. The fine particulate then settles outside the filter ele-
ments and is collected at the bottom of the vessel containing the filter elements.
This type of granular bed filter was chosen to be tested in the EPA/Exxon Hinlplant since it
was considered to possess certain advantages over other high temperature particulate removal systems,
and development of granular bed filters as a class was believed to be further along than the develop-
ment of other high temperature particulate removal systems. Of the granular bed filters under
504
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01
o
01
Figure 1. Ouoon Granular Bed Filter and Shroud
Figure 2. View of Miniplant Showing Filter Prenura Vessel
-------
development, the Ducon filter had been previously tested at relatively high temperatures and pressures
and shaved some promise in meeting the high particulate removal efficiencies required by the gas
turbines3. The Ducon design also had the desirable feature of retaining che granular filter media
in the filter vessel during tie cleaning step. In all other high pressure granular bed filters, the
filter oedia is removed, cleaned externally and recycled b&ck to the filter vessel.
The objectives of the granular bed filter test program at Exxon Research were to measure the
outlet loading from the granular bed filter, determine if the removal efficiency was maintained with
use, measure operational stability of the filter e.g., can low pressure drop across the filter be
maintained, does the filter plug, is the aaount of blow back gas needed to maintain steady operation
within reason, etc. and linally, to measure the long term life of the filter hardware. The primary
operating parameters were the filtered gas flow rate, usually measured as the gas velocity entering
each filter bed. the blow back gas flow rate, the duration and the frequency of the blow back step,
and the type of filter media used i.e., particle size, shape, density.
DESCRIPTION OF EQUIPMENT
The granular bed filter was Installed in a pressure vessel tied into the flue gas exit system
from the EPA/Exxon Hlnlplant pressurized fluidlzed bed combustion unit. The Hiniplant was described
in detail in a previous report2. Briefly, it consists of a comhustor, 33 en (15 in) inside dlaaef!?
and 10 n (32 ft) high, capable of operating at pressures up to 1000 kPa (10 atm abs), temperatures
up to 980ฐC (1800ฐF), superficial gas velocities up to 3 m/s (10 ft/sec) with coal feed rates up to
(1200 SCFM). At a typical fluidization velocity now anticipated for a PFBC boiler (about 6 ft/sec),
the flue gas rate is 18Sm3/nin (650 SCFM). The partiiulate loading in the flue gas entering the filter
is about 2.3 g/c>3 (1.0 gr/SCF). The mass median particle size Is 5 to 7 nlcrons.
The pressure vesr I housing the filter cle&ents consists of a refractory lined vessel
approximately 2.4 m (8 fw) in diaseter by 3.A (11 ft) high. Figure 2 is a photograph of the Mini-
plant showing the pressure vessel. Access to the interior of the pressure vessel can be made through
a 70 era (27 in) manhole. The vessel can hold up to four filter elements installed through four flanges
at the top of the vessel. Each filter element is contained within a shroud, as shown In Figure 1, in
the inside of the pressure vessel. Inlet gas is piped to each shroud, passing through a measuring
orifice which determines the flow rate to each filter element. The gas then enters through the flanged
openings shown in Figure 1. Clean gas exits from each shroud through openings at the top (Figure 1)
and fills the interior of the pressure shell. Blow back air enters each filter element through the
top flanges of the pressure vessel and flows in reverse direction through each filter element. Parti-
culates removed from the filter element during blow back impinge on the inside surface of the shrojd,
fall to the bottom and are collected in lock hoppers. The blow back gas leaving a filter element flows
in reverse direction through the inlet gas system into the filter elements which are in Che filtration
step.
A natural gas preheat burner was Installed to heat the interior of the pressure vessel to a
temperature above the dew point of the combustor flue gas before starting a filtration test. The
burner fires into the vessel through a side port for an 8 to 12 hour period prior to the start of a
run. The filter vessel is at atmospheric pressure during this period. The burner was Installed
after Initial operation of the filter resulted in condensation of moisture on the filter elements
during heat up with flue gas from the coal combustor. Condensation in the presence of flue gas parti-
culates caused plugglrg of the filter and particulate removal lines.
A number of filter element designs and blow back methods were tested during the filter shake-
down tests. Each design and the results of the tests are described in the following Section.
RESULTS
CBF Model 1
Initially, three filter elements were purchased from the Ducon Company. Each element was 20 cm
(8 in) in diameter ty 1.8 m (6 ft) long and contained twelve beds. The Inlet screen size was 50 mesh.
The nominal flow capacity of each element was 8.5Sa^/min (300 SCFM). One of the Ducon elements was
designed to be blown back by short pulses of high pressure air. The pulse duration was approximately
0.5s. This blow back method was cot tested. The other two Ducon elements were blown back with a
larger volume of air for longer durations. The Intent was to fluidize the filter beds rather than
shocking them with a shore pulse of high pressure air. The volume of blow back air was sufficient to
506
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fluidize the filter media. The duration was designed to be less than 10s. Blow back was accomplished
by Isolating one end of a filter element by a blow back nozzle and seal plate and blowing back with air
at a pressure slightly higher than filtration pressure.
After Installation of the filters, shakedown began with ambient temperature testing of the
system.
The objectives of these preliminary tests were to (1) check combustor pressure control with
the filter on line, (2) pressure test the system, (3) check the alignment of the blow back nozzles,
(4) check out the operation of the blow back flow system, and (5) measure the distribution of flow
to each one of the filter elements. A number of mechanical problems were discovered (leaks, mis-
alignments, etc.) and had to be corrected before further testing could be resumed.
A number of high temperature runs were then attempted but the pressure drops across the filter
were extremely high and all attempts at blow back were unsuccessful. Inspection of the filter ele-
ments after each of these runs showed that a hard filter cake had formed on the inlet retaining
screens. This Is shown in Figure 3. The filter medium was usually participate free indicating very
little penetration through the screens. It was originally thought that the plugging occurred during
startup when moisture was present. The preheat burner was later installed and a run was made to
re-evaluate the Ducon filter. The same screen plugging problems occurred and the original Ducon
filter was deemed to be unacceptable for our application.
CBF Model 2
Discussions with the Ducon Company led to the design and fabrication of a second filter system.
Ducon Indicated that they had encountered the same screen plugging problem with flyash and bypassed
it by removing the screens and designing the individual beds with more freeboard to prevent entraln-
nent of the filter media during blow back. It was also recommended that a fluldizlng grid be used
at the bottom of the beds to assure good distribution of the blow back air. The use of an ejector
to replace the sometimes troublesome plunger type blow back nozzles was also suggested. Filter
elements incorporating these suggestions were fabricated by Exxon Research and shakedown continued
using two of these elements each of which contained five filter beds. A photograph of one of these
elements Is shown In Figure 4.
Dirty gas enters the beds through the openings below the top flanges and passes downward through
the filter beds and out into the clean gas outlet tube in the center of the decent. During blow
back, the blow back air passes up through the fluldlzing grids supporting each bed, fluidizes the
beds and blows the fine partlculates out through the inlet slots. A 18 cm (7 in) freeboard above the
the filter beds acts as a disengaging section for the filter media and prevents its entralnment
through the outlet slot.
Operability of the modified filter system has been demonstrated. The end of the initial shake-
down phase was signified by the successful completion of a 24 hour demonstration run. This run was
preceded by a number of shorter duration runs used to establish suitable operating conditions for the
demonstration run. The successful use of an ejector was also demonstrated during one of these runs.
These runs were successful in that filtration, ability to blow back, ability to calntain low pres-
sure drops and collection of particulates after bior back were demonstrated. Collection efficiencies
of 90 to 9SZ were measured for the first few hours of the runs based on outlet particulate concen-
trations of about 0.1 g/m3 (0.05 gr/SCF). Operation for up to 24 hours was also demonstrated with no
significant increase in baseline pressure drop across the filter. Blow back was usually required
every 10-20 minutes during which time the filter pressure drop had increased 14 kPa (2 psl) above its
baseline value. A range of blov back conditions were used to restore the baseline pressure drop.
Blow back durations ranged between 2 and 30 seconds and superficial velocity between 0.15 and 0.75
m/s (0.5 and 2.5 ft/s). Filtration velocities generally ranged between 20 and 24 m/min (60 and 80
ft/nin). Filter media consisting of 300 to 600 micron quartz particles and 840 to 1400 micron alumina
particles were tested. The quantity of blow back air used ranged from 1 to 5! of the filtered gas
rate.
The particulates passing the filter had a cass median eizr of about 3 microns with about 101
larger than 10 microns.
Problem Areas
A number of problems were defined during the shakedown and operation of the filter. Demonstra-
ted particulate outlet concentrations are still higher than the tentative gas turbine inlet require-
ments. However, firm turbine requirements have not been set as yet and it may be too early to reach
507
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Figure 3. Filter with Plugged Inlet Retaining Screens
Figure 4. Modified Filter Element
-------
any ccr.clusior.5 regardir.g the suitability cf the filler to protect a nas turbine. The lover out! .
particulate loadings .->f 0.1 ฃ/r.i (0.05 ฃr/Si,i ) r.eet the curre-t Ll'A eirission standards, h'.-.wever, ir.
all runs, :t was observed liiat the outlet loading increased -'it:: tir.e. Inis is =.howr. in Figure- r>
for two recent rur.s usir.t: the alnrina filter r.edifi. in these recent rur.i" the increase has been less
than tl:at observed in earlier runs due to -ore frequent blow backs at higher velocities. It has r.ot
as yet been uer.onb'rated that the tPA emission standard can be set for r.cre thai: a few hours of
operation.
'.nsfad of forcing a layer or. the top of the beds as had been expected
f line part ic'jla
process explains tin-
es ฐ. Mri.u/houl the 1 liter b
.'t also ;-cปs:bly
.
the increase :n outlet particui.ite cor.v'er.trat i
Ar.t'tlier factor vhic!. kir-jlii be resj.or sil.i*.- !(-r ti:ซ.- retention ct par t .ci'latet; in the filter bco:.
is the rec.>cl in,: uf part :cu! jit,3 : rc-r. bed tu l.d ซ'urii:,- the bK;-. Lacf. step. lhe ;-hrouJs ui:ich surrour.
e.'icl. ilt.Tviil r.a> :r.^e'ie the ;;et!.iir.;- ol t::t- ;'.irt u uiatos bleuT: fro-, the cle:.ป-r.'s. Uiib co-jj-: result i
hifih ir.len.'il rec ircuiat iun rales Let'-eer. ป it.-r.er.ts, or. :.:::ฃ t!.e ;>ar t ii:ulutes to buiid up it", "lie t%ec.ป
anJ thereby c.'iu.sir.^ a~, ir.creji.e :i. t:ie outlet ;ป;ปri icu'.ate cuncentiat ior. . II. is viji be examined it: t:ie
future.
Anttt.er r--turri::K probler, vas t::e los.-> .if filter r.edia Jurin, bl'-'v rack. This occurred ever.
vhen blov i-ack velocities ve:e relatively In:-. .:old r.odei testa indicated that the iot>& vas prouuoly
caused by instantaneouslv hii.ii velocities at the becinninj; o! the bluv back step. i::e use o: sicwer
cper.in^ b lev back valves ar.d sufiic.er.t sur^-e volur.e between the vilves ard the I liter beds prevented
the loss of the filter r.edia. L'ecruasini; tr.e sire ol tiu.- inie1. ^as slt-t at the top o! ti.e liltei beds
also decreased t i:e exte:.t ol the filter neuia loss.
Another pctentiai proyle". vltii the curre'it uesipn is its vjl t-.er.ib '. i itv to uj'sets. If upsets
occur, such as bed pluฃฃir.g cr loss of lilter -edia, the operatiiuj problems caused by sue:: upsets
usually require shutdevn of t!:e syf.ieE. it is usually not pcssible tc take corrective action vr.ich
restores gocd operation. Another probler. which -ay be unique '.o the >!ini;-:ant was the interaction
of thy ^ror.ui.ir bed filter witn the rest of the r EC syster. during blow rac< cycle. .*>n increase in
sysler, presaure was noted during blow back resultir.i; ir. probier:* with the .:.;al i oi syster. vliic!: Is
cor.trullcd bv tin- differential pressure between tnc coal feed vessel ar.d cc-bustcr. Vhiป required
codifications to the coal feea control systes. to r.ini=.i;e ti;e effects.
Ci'KKtNT l'KO(-KA.V
In addition to the tl'A sponsored progra" to evaluate hot i:as cleanup systers, LXX-JII Research
is also under coi.tract to the Depaitr^nt of tner^-y to carry out extended >-as tur'rir.e and boiler tube
tuterlals tests. Preparations are now In progress for a 100 hcur shakedown test in whic'r. the current
granular bed lilter systec will be used to clean the flue gas before it is expanded through a static:,
ary turbine test section. The test section was provided by the General Liectric Cesp.ir.y. Boiler
tube spcci-cns for this te-^t were provided by Westir.ghouse Research Laboratory. Fallowing the 100
hour test, additional tests will be carried out ur.der KHA sponsorship alr.cd at i-provinn the filter
efficiency and efficiency nair.tenar.ee. These tests will primarily concentrate on i:.provinc the
509
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C 0.12
u
0.08
I
0.04
O H1N1PLANT RUN 63
O MIN1PLANT RUN 65
EPA EMISSION STANDARD
0.3
0.2
0.1
u
cj
I
4 6
TIME INTO RUN (HRS)
10
Figure 5. Time History of GBF Performanea
510
-------
cleaning of the filter during the blow back step. The effect of the shrouds which surround each
filter element on the settling of the particles will also be examined. The couple don of the DOE
sponsored materials test program, consisting of a 1000 hour exposure test, will await the outcome
of the above programs.
In addition to the CBF evaluation program, Exxon Research will also evaluate alternate parti-
cjlute removal devices for the EPA. Current plans are to carry out small scale tests on a flue gas
side stream using two devices being studied under EPA sponsorship. The devices are a high temperature
bag filter studied by Acurex and a high temperature electrostatic precipitator studied by Fesearch-
Cottrell. After the small scale tests are completed, one of the devices will be tested in larger
scale on the Hiniplant.
REFERENCES
1. 0. L. Realms, et al. Westinghouse Research Laboratory, "Fluidlzed Bed Combustion Process
Evaluation," EPA-650/2-75-027-C, September 1975.
2. R. C. Hoke, et al, Exxon Research and Engineering Company, "Studies of the Pressurized
Fluldlzed-Bed Coal Combustion Process," EPA-600/7-76-O11, September 1976.
3. AlChE Symposium Series No. 137 Vol. 70 pp. 388-396 (1974).
511
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QUESTIONS/RESPONSES/COMMENTS
MR. 7ESORIERO: Al Tesoriero of the Ducon Company. As you are
aware, the element that has been shown on the slide is not, in fact,
the current Ducon design. The freeboard height, which could explain
the loss of media in the testing, appears to be approximately six
inches. Secondly, the media that has been used, I think you said
800 or 1400 microns, and a space velocity of 66 to 100 ft/min; and,
of course, you realize that both of those variables will affect
performance. Those are the two comments that I have. The question
that I would like to address to you -is what you are going to use
that 250 to 600 micron filter media which is needed to get good
filtration that the face velocity is a problem and the freeboard
height is a problem.
MR. HOKE: The question was by Al Tesoriero who is associated
with the Ducon Company. The first question was concerned with the filter
bed freeboard. The units that we use have a freeboard of six
inches. Al mentioned that Ducon1s design now specifies 10 inches.
We have looked at the question of filter media loss in some cold
model Plexiglas units. We think that the problem is not the free-
board, but rather the fact that during the initial part of the blowback
step, unless you're very careful, a flow surge occurs which lifts
the bed and Mows some media out. If we can control the flow surge,
there is no problem with loss of filter media.
This is the approach that we are taking now. We are putting
in slower acting valves, and we are trying to locate the valves some
distance from the filter in order to get some surge capacity in the
inlet lines. We think that if this is done, we can run without
losing the filter media.
As far as the filter media particle size is concerned, from a
purely theoretical standpoint, you would like to use finer filter
media. However, we feel that the problem at the present time is not
in getting good filtration, but rather in cleaning the particulate
from the filter media. To balance the situation, we are trying to
get a filter media which isn't lost during the blowback.
We have run with smaller granules in the filter bed. We
haven't really seen any difference in filtration efficiency; and
the reason for that, I think, is that we are seeing penetration
by the dust caused by mixing in the bed which is independent of the
particle size of the filter media.
If we once get to the point where the beds are being adequately
cleaned during the blowback step, we will then use smaller filter
media particle sizes.
512
-------
DR. ZENZ: I might point out that you realize that if you make
the filter media particles smaller, you clean the bed easier.
MR. HOKE: Well, again, we have tried to balance two considera-
tions. We had a considerable loss of filter media during our first
runs with the quartz because it was a smaller particle and it had
lower density; so we've gone in the direction of using larger particles
and higher density in order to minimize that problen.
DR. ZENZ: What is your support screen like? Is it just a plain
screen?
MR. HOKE: The screen at the bottom of the bed?
DR. ZENZ: The bottom of the bed, yes.
MR. HOKE: No, it's a screen backed up with a fluidization grid.
DR. ZENZ: What's the fluidization grid like?
MR. HOKE: It's a perforated plate.
DR. ZENZ: You don't know the percentage of open area, do you?
MR. HOKE: I don't recall exactly, but it's designed to give a
pressure drop which I think is about 30 percent of the pressure drop
through the bed. It's conventional design practice.
SPEAKER (from the floor): Just vo ask Dr. Zenz a question.
When you say "optimizing the particle size to 250 to 600 microns,"
do you mean that you have to go to that size to prevent re-entrainment.
of fines after you have completed the blowback?
DR. ZENZ: If the particle size is too big, we don't fluidize.
We could use so much gas on the fluidization to clean it that it
becomes impractical and yet we are not cleaning the filter media as
we'll. If we go down too fine, then it gets kind of ridiculous; the
pressure drop gets too high. So w& more or less optimize out at 250
to 600 iiiicrons.
513
-------
MR. HOKE: We have looked at the effect of velocity during
blowback. In ambient temperature tests in Plexiglas units.
Again, the problem appears to us to be the adhesive nature of the
particulate. Enough of it sticks to the outside of the filter media
so that it is mixed back in the bed, and eventually, we get a uniform
distribution of the particulate from top to bottom.
We've looked at the beds after an extended run, and the filter
beds are loaded with the particulate. Sometimes as much as 30
percent of the weight of the filter media is the particulate, and it
is a uniform distribution from top to bottom. We think the reason
for this is the fact that the dust is very adhesive, and it doesn't
get completely cleared off of the bed during blowback.
CHAIRMAN: Are there other questions? Yes.
SPEAKER: Just out of curiosity, who supplied the filter and
the media?
MR. HOKE: The initial filter was purchased from Ducon. The
second filter was designed by Exxon after consulting with Ducon, and
we constructed it and tested both systems.
CHAIRMAN: Are there any other comments or questions? Thank
you very much, Ron.
514
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INTRODUCTION
DALE KEAIRNS, CHAIRMAN: The next paper will address Granular
Bed Filters for Participate Removal at High Temperature and Pressure.
This paper is by personnel at Air Pollution Technology, R.D. Parker,
S-C Yung, R.G. Patterson, and S. Calvert; and it is also coauthored by
Dennis Drehmel of EPA. Dr. Parker will present the paper.
Dr. Parker received his Ph.D. in mechanical engineering from
Duke University, and he has been with Air Pollution Technology since
that time. He has been carrying out work in terms of developing and
understanding high-temperature, high-pressure particulate control
equipment. Dr. Parker.
515
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Granular Bed Filters for
Paniculate Removal at High
Temperature and Pressure
S.C. Yung. R.D. Parker, R.G. Patterson,
S.Calvert
A.P.T.Jnc.
D. C. Drehmel
AHSTRACT US EPA
The technological status of granular bed filtration for particulatc control
at high temperature and pressure (H'l'l') is critically examined with regard to appli-
cations for fluidizcd bed combust ion (I;KC) processes. A performance model for
granular hod filtration is presented and compared with experimental data fron lab-
oratory and industrial scale filters. The potential of granular bed filters for
I IT I* particulatc control is evaluated with re>,:rd to the cleanup requirements based
on turbine limitations, emissions regulations inr current information on the size
and mass loading of particulatc emitted from l:!iC processes.
INTRODUCTION:
Cranular bed filters have been proposed for the removal of particulatc matter
from high temperature and pressure pases. One such application would be as the
tertiary collection device in a pressurised fluidiied bed boiler power plant as
illustrated schematically in l-'ic.ure 1. The gas leaves the boiler at a temperature
of approximately '.>!MIฐ(; and a pressure of about II) atm. l;irst it passes through a
primary cyclone which removes larger particles (including unburnt carbon) and recy-
cles these particles to the comhustor. The pas leaves the primary cyclone and
passes through a secondary cyclone or mil tic lone separator. This removes nore
large particles and reduces the mass loading of particulatc to the order of 1 gr/SCF.
A tcrtiaiy cleanup device is necessary to reduce the part iculate loadini;
sufficiently to protect the gas turbine from exc-.-ssivc erosion and corrosion damage.
It is also desirable fur the )'.as at this stage to he sufficiently clean to satisfy
all emissions regulations. However, if necessary, it is possible to satisfy the
emissions regulations by cleaning the gas downstream of the turbine using convc-n-
tional control technology.
An idea of the size of equipment required for a fnll-scr.le power plant can
be obtained from the Phase II I'.lIAS study. In the h'est i nghousc report (Reedier et al.
I'.lTfi)1 a design for a (>K(l "!KO power plant was presented. Four pressurized fluidizcd
hod boilers fed gas to two gas turbines. liach gas turbine handled 7M) Ih/s of gas.
llach boiler had four granular bed filters associated with it. The granular bed
filters each handled HH.75 Ib/s or approximately 30,000 CI-'M at 1,77<)ฐF and 10 atn.
CI.F.AMIP KF.QIIIRBIF.XTS
It is generally supposed that the paniculate removal requirements imposed by
the gas turbine tolerance will be more stringent than the emissions regulations.
This may or may not he the case depending on what turbine tolerances arc specified.
It has been suggested (West inghousc7) that a mass loading of 0.15 gr/SCF for par-
ticles smaller than 2 um would allow a satisfactory turbine life. If there were a
sufficient loading of particulatc smaller than 2 urn it would be possible for the gas
to be cleaned sufficiently to protect the turbine while still exceeding the emissions
regulation of 0.1 Ib/lO'-RTII (approximately 0.06 gr/SCF).
The best available information concerning the size and mass loading of par-
ticulatc entering the tertiary collection stage of a pressurized fliiidizcd bed pro-
cess has been reported by lloke et al.' The mass loading has been found to be
typically about I gr/SCF. The size of this part iculate is represented in Figure 2.
Approximately .iO', of the mass is smaller than 2 um so that about H.5 j.r/SCF can be
expected in this size range. It may he possible to protect the turbine with 50'.
collection of these fine particles, but the current emissions regulations require
516
-------
AIR
I
COMPRESSOR
6AS TURBINE
>STACK
TERTIARY HEAT
COLLECTOR RECOVERY
CYCLONE
en
BOILER
STEAM
FEED
WATER
Figure 1. Pretturized fluiditad bed boiler
-------
50
I 10
i s
1.0
0.5
10 30 50 70
Wt % undersize
Figure 2. Particle tile distribution from Exxon miniplant
90
518
-------
at least 80** collection. If the emissions regulations were reduced to 0.05 lb/106
BTIJ at least 90ป collection would be required. This assumes effectively 100ฐ collec-
tion of particles larger than 2 urn.
f.ccently Svcrdrup and Archer* proposed that a mass loading of 0.002 gr/SCF
(for all particle si:es) is required to protect the turbine. If this criterion were
ado.-tcd, clearly the emissions regulations would be satisfied. However, a tertiary
collection efficiency of at least 99.8ฐ would be required to reduce the loading
observed at the lixxon miniplant to this level.
Increasing the efficiency of n tertiary collection device from about 80 to
90' (emissions regulations) to 99.8', (turbine limitations) is a very significant
increase affecting both the cost and the availability of technology for high tem-
perature and pressure (IITP) ^articulate control. A firmer understanding of the
degree of collection required to protect turbines would greatly facilitate the
development of 1ITI1 particle collection technology.
C.RA.NULAR BI:D I-UTKRS
l-'or our purposes, granular bed filters arc -Icfined as any filtration system
comprised of a stationary or slowly moving bed of sc|.:'ratc relatively close-
packed granules as the filtration ni-dium. There arc three basic approaches to
granular bed filtration currently being developed. The filter being tested at the
lixxon miniplaiit is in the class referred to as "fixed bed" filters. The bed of
granules is kept stationary during filtration. It is cleaned by a reverse flow
of gas which fluidizcs the bed and entrains the collected paniculate. The Uucon
and Rexnord granular bed filters are examples of fixed bed filters. The principal
advantage of the fixed iicd approach is that the granules do not have to he recir-
culatcd and thus can he used for many filtration cycles. I'or this reason, fixed
beds have potentially lower operating costs.
In a second type of granular bed filter, referred to as the "moving bed"
filter, the gas flows through a slowly moving bed of granules. Ttie collected
dust particles arc carried with the bed and later removed from the granules before
the granules are recirculatcd. The Combustion Power Company's "dry scrubber" is
an example of a moving bed filter. Moving beds have the advantage that the granules
and dust arc separated externally and arc free from the problem of particle buildup
and bed plugging. An inexpensive but effective means of rccirculating the granules
could significantly lower operating costs.
The third type of filter is the "intermittently moving" filter. This filter
uses stationary granules for filtration and causes the granules to move through the
system intermittently during the cleaning cycle. The panel-bed filter developed at
the City College of the City University of New York (Squires and Pfeffer"1) is an
example of an intermittently moving bed filter.
The major problems with each approach arc summarised in Table I. Many of
these problems can be resolved through further research and development. Granular
bed filters are used successfully to control emissions from clinker coolers in the
cement industry and on hog-fuel boilers in the forest products industry. They
operate in the range of 100-200ฐC and near atmospheric pressure. However appli-
cations for controlling fine particulatc at elevated temperatures and pressures are
very scarce. Granular bed filters are attractive for these applications because
they can be made to withstand high temperature and pressure environments relatively
easily. However more work is needed to determine whether granular bed filters are
satisfactory from the point of view of primary collection efficiency, cost, and
reliability.
Granular Bed filter Model
In order to help evaluate the potential of granular bed filters for IITP
particulatc control we have developed a performance model which has been used to
predict collection efficiency for lab-scale and full-scale filters. The model is
based on the collection of particles in a clean granular bed. The collection
519
-------
Table I. Granular Bed Filter Problems
FIXED PF.D
Plugging of retaining
grids and possible par-
ticle buildup in bed.
2. Particle seepage through
bed during cleaning
cycles.
l-'luidization redisperses
fine dust during clean-
ing.
4. IITP valving required
for reverse air
clcanirg.
S. Temperature losses
proportional to vol-
ume of cleaning air.
MOVING BHI1
1. Particle re-entrainment
in moving bed.
2. Granule rccirculat ion
may cause high opera-
ting cost.
3. Difficult to form a
cake in moving bed.
4. Rrosion of retaining
grids and transport
systems.
5. Temperature losses
proportional to heat
capacity of recircu-
lated granules and
rccirculation rate.
l.\'Ti:RMITTi;\T BED
1. Low gas capacity
can cause high
capital cost.
2. Granule rccircu-
lation may cause
high operating
cost.
3. Need form sur-
face cake to avoid
plugging problems.
4. Erosion of retain-
ing grids and
transport system.
5. Temperature losses
proportional to
heat capacity of
recirctilatcd
granules and re-
circulation rate.
520
-------
efficiency would be affected if there were a significant filter cake on the bed sur-
face and within the bed. The presence of a surface cake, however, has not been
noticed by many investigators working with large-scale filters, h'e feel the clean
bed nodel predicts a conservative estimate of the efficiency attainable by granular
filtration and is a satisfactory model for filters which operate primarily without
the presence of a filter cake.
The granular bed can be envisioned as a great number of impa:tion stages
connected in series as illustrated in I-'igurc 3. Particle collection is by
inerttal impaction and is similar to collection in a cascade impactor. The jet
openings arc the pores in each layer of granules. It is assumed that the jet diam-
eters in the granular bed arc of uniform si:e and are equal to the hydraulic diam-
eter of the void space. The gas velocity in the jet is the average superstitial
gas velocity.
If ' TI' is the collection efficiency of one inpaction stage, the particle
la i| *^< llll- vvfl*l-l,llYJII Vltlt-lvlt
penetration for the granular bed will be
Pt, = (1-nV
(1)
where Pt
penetration for particles with diameter d , fraction
single stage collection efficiency, fraction
number of impact icn stages
As in some cascade impactors, each layer of granules serves both as the jet
plate and as the collection plate. Therefore, each lay -r is an impact ion stage
and "X" is equal to the number of granular layers in a bed. l:or a randomly packed
bed
X
3 Z
where Z = bed depth, cm
AC = granule diameter, cm
and
Pt
(31
The impaction collection efficiency, n, is .1 function of, K , the Lnertial
impaction parameter. The impaction parameter is defined as '
where C'
since
9 v.
Cunningham slip factor, dimcnsionlcss
particle density, g/cm3
particle diameter, cm
*
u. = jet velocity, cm/s
u_ = gas viscosity, poise
dJ
jet diameter, cm
"r.i
(5)
521
-------
969
TAR6ET
Figure3. Impaction model
522
-------
and dj ป 4rH = | (^J dc (6)
where u^.j ซ average interstitial gas velocity, cm/s
c > bed porosity, fraction
TJJ ฐ hydraulic radius, cm
dc = granule diameter, cm
, ,. . C' p d ' ur
we have K J (1 *-2I} (7)
p 2 \ e' / 9 ur d
The relationship between, r,, and, Kp, can be evaluated once the flow field
is defined. Flow fields reported in the literature; e.g., Ranz and Konp.' and
Marple7 are adequate for Kp>0.15. For Kp<0.15, there is no suitable flow field
reported in the literature. Therefore, the relationship between, n, and, K ,
cannot be calculated analytically from the available literature. p
The single stage collection efficiency has been calculated as a function of
Kp from equation 3 and experimental data. Figure 4 shows the results. There is
scatter in the lower end of the curve. For Kp<10"r, n is very sensitive to experi-
mental data. A few percent scatter in data will cause, n, to fluctuate greatly.
Figure 5 compares the experimentally determined, n, versus, K'p, curve with those
reported by Ran: and h'ong*. Stern et al.', and Mercer and StuVford.' All reported
curves arc for Kp>0.15. As can he seen, the present study is consistent with other
researchers's results and is a continuation of their curves into the raiiฃe most
likely to he important for high temperature and pressure filtration.
P?.ป.tsky ct al.10 and Kncttig and Bceckmans1' studied the collection of
monodispcrscd aerosol particles in granular bed filters. Their data were trans-
formed into, Kp, versus, n, plots as shown in Figure 5.
Knettig and Bceckmans11 used 42S urn glass beads as granular material. Bed
porosity was 0.38. Aerosol particles were 0.8, 1.6, and 2.9 ;im in diameter. As
can be seen from Figure S, their data are in close agreement with the results of
the present study.
Parctsky ct al.10 investigated the filtration of 1.1 urn diameter polystyrene
latex aerosols b/ beds of sand. They used a bed of 10-14 mesh (1,200-1,700 urn)
angular sand and a bed of 20-30 mesh (500-850 urn) sand at superficial gas velocities
between 0.3 and 80 cn/s. Bed porosities were 0.41 and 0.43, respectively, h'e have
calculated single stage collection efficiencies from their data. In the calcula-
tion, the granule diameters were assumed to be the arithmetical mean of the smallest
and the largest granule size in the bed. The results are plotted in Figure 5. For
a given incrtial parameter, Paretsky ct al.'s10 data give a higher collection effi-
ciency than reported in this study. Their data would be close to that of the pre-
sent study if the smallest granule dianeter were used instead of the arithmetic
mean.
The design model is based on particle collection by a clean bed. If there
is no filter cake formed on the surface and the collected particles are uniformly
distributed in the bed, the model should be applicable. The design model has
been used to predict the performance of a Rexnord gravel bed filter and of Combus-
tion Power Company "dry scrubber." The predictions are compared with field samp-
ling data taken from the literature.
Rexnord Gravel Bod Filter
McCain" conducted a performance test on a Rexnord gravel bed filter. The
gravel bed filter was installed to clean emissions from a clinker cooler in a
Portland cement plant.
523
-------
O.lci i i i < ui
0.01
o>
"5
BB
CO
J2 0.001
O 1.09ft
A 0.76ft
O 0.5/t
0.0001
_l I I I I I I I I
0.002
0.01
0.1 0.2
Figure 4. Experimental data
524
-------
1.0
0.5
1.0
ฃ 0.05
tป
o
0.01
.005
0.001
20-
PAHETS8Y
DATA X
10-14 (11ESH
STEBW
IMERCER
STAFFORD
I I I I I
KUETTI6 and
0.002
0.01 0.1
Impaetion parameter
Figure 5. Compcriton with pravious wortc
0.5
525
-------
Saraples were taken simultaneously at the fi!ter inlet and outlet with cascade
inpactors. The operating conditions of the gravel beJ were:
Crave! diameter * 4 nm
Face velocity = 73 cn/s
Gas temperature ป 175ฐt:
Pressure drop ซ 25.4 cm K.C.
It was assumed that there was no surface cake ami that the pressure drop across the
bed was 80* of the overall pressure drop. Mrgun's equation fHrgun'1) was used to
estimate a porosity of 0.25. Kith this be.' porosity, the r.rade penetration curve
was calculated for the operating conditions listed above.
Figure 6 is the predicted Rrade penetration curve along with that neasurcd
by McCain17. The prediction is in good agreement with the data.
Combustion Power Company Moving Crave! Red
Hood1* reported the evaluation of the Combustion Power Company (CI'C) moving
gravel bed filter on the control of participate emissions from a hog-fuel fired
boiler. The gravel bed filter was a prototype unit with suggested capacity of
1,133 Am Vein (40,000 ACFM). The bed was packed with an intermediate-size gravel
which was retained on a 3.2 mm 11/8") wire mesh and passed at b.4 ma (1/4") nesh
screen. The bed was a single down-flowing annulus 2.6 m (8.5 ft) O.I), and 1.8 n
(6 ft) I.D.
During sampling the unit was operated at a How rate of 1,558 nVmin (55,000
ACFM). The gas temperature was 177"C (350ฐl-). Figure 7 shows the penetration
curves for three sampling runs.
The avciage granule diameter was assumed to be 4.6 mm and the bed porosity
was calculated to he 0.25. The predicted grade penetration curve is shown in
Figure 7. The predicted penetration is higher than that measured. Recent data
obtained by A.I'.T. on the CI'C moving bed filter is shown in Figure 8. These data
agree well wit!, the model's prediction.
PERFORMANCE PREDICTION FOR IH T APPLICATION
Exxon Research and Engineering Company installed a Ducon granular bed filter
at their fluidized bed coal combust or miniplant. The bed is packed with Agsco no.
2 quart: (400 um mean diameter) to a depth of 3.8 cm (J.S in.). The temperature
of the flue gas from the combustor is 870ฐC (1,600ฐF) and the pressure is 10 atm.
According to lloke1*, there is no surface cake formed and the fly ash is
uniformly distributed in the bed. lor this condition, the clean bed model can
be used to predict the performance of the Ducon granular bed in the miniplant.
From the pressure drop data reported by Exxon, the bed porosity was esti
Dated to be 0.24. Figure 9 shows the predicted performance of the miniplar.t
granular bed filter at high temperature and high pressure and at ambient condi-
tions. A particle density of 1.5 g/cm' was used in the calculation.
If the particle si:c distribution is known. Figure 0 can be used to estimate
the overall collection efficiency of the Ducon granular bed filter. The equation
relating the efficiency for collecting one particle diameter to the overall
efficiency is:
E
1-Ft ป 1- I Pt
id f
-------
1.0
ฃ 0.1
0.01
MCCAIN'S
DATA
as UD s
Particle dia.
10
Figura& Compvison with d*ta for Raxnord fitter
527
-------
1.0
0.5
0.1
0.05
PREDICTIiH
H000S \ ^^
0.1 0.5 1
Particle dia.pm
Figure?. Comparison with d*t* for CPC moving ted fitter
528
-------
0.5
0.05
0.011 I M . I 1
0.4 1 5
Particle dia./imft
Figure 8. Comparison vrith CPC moving bed fitter
529
-------
1.0
0.5
0.1
0.05
I i i i i
0.01
20UC
Graoale tiiaaeter:400/A
Bed Bepth:3.8caj :
.
p = 1.5 g/eซ3
c = 0.25
lOatn
870ฐC
latra
870ฐC
0.1 .5 1.0 5
PARTICLE DIAMETER, pn
FigunS. PrtaSctwJ C8F Pwfeปmaneป
10
530
-------
Hokc' reported data on the particle sire distribution leaving the secondary
cyclone (Figure -). The mass median diameter is 5.5 urn and the geometric standard
deviation is 2.9. The mass loading is approximately 1 gr/SCF. The overall collec-
tion efficiency for this size distribution was calculated graphically to be 95.11
(4.9* penetration!. The emission, therefore, is predicted to be about O.ns gr/SCF
which is satisfactory to meet current emissions regulations, and is within thi
measured range reported by iloke'1.
As can be seen from Figure > '.he granular bed filter should be very effi-
cient for all particles larger than 1 to 2 urn in diameter. Whether or not this
filter is efficiency enough to protect the gas turbine will depend strongly on the
turbine's tolerance for submicron particles.
CONCLUSIONS
Granular bed filter technology has the potential for resolving the 1ITP
paniculate control problem. Collection efficiencies should be sufficient to meet
the current emissions regulations for particulates. Performance may be satisfac-
tory for protecting gas turbines, however, this will depend strongly on the amount
of subnicron particles a turbine can tolerate.
There arc many operational problems and uncertainties which need to he
resolved before HTP granular bed filters can be considered sufficiently reliable
and economical for commercial application. These problems include:
how to prevent particle seepage through the bed (during cleaning or
fill rat ion).
how to reduce temperature losses (especially during cleaning).
how to improve the efficiency and reduce the cost of granule regeneration
and recirculation.
how to reduce pressure drop across the bed.
how to prevent attrition of granules causing particle re-entrainment.
how to prevent sintering of granules.
how to prevent plugging of retaining grids.
Resolving these problems will not only provide a solution to the HTP par-
ticle collection problem, but will improve granular bed filter technology for many
other applications, especially where hot, corrosive gases arc encountered.
ACKNOKLI-DCMIINT
This work is supported by the U.S. F.nvi ronmental Protection Agency.
LIST OF SYMBOLS
C1 ป Cunningham slip factor, dimensionless
d = granule diameter, cm
d. jet diameter, cm
dp * particle diameter, cm
d ป aerodynamic 'iiaroeter ซ d (p C1)', uraA
pa P P
E ป overall collection efficiency, fraction
f(d ) particle size frequency distribution
K ป incrtial impaction parameter, dimensionless
X ซ number of impaction stages
Ft" ป overall penetration, fraction
Pt , penetration for particles with diameter, d , fraction
531
-------
r.. = hydraulic radius, cm
u_. = average interstitial gas velocity, cm/s
u, = jet velocity, cra/s
Z ป bed depth, cm
r = bed porosity, fraction
n = single stage collection efficiency, fraction
PP ป gas viscosity, poise
P = particle density, g/cm'
LIST OF REFERENCES
1. Bccchcr, O.T. et al. Energy Conversion Alternatives Study fECAS), h'cst in-house
Phase II Final Report, Vol. Ill, "Summary and Advanced Steam Plant with Pres-
surized Fluidizcd Bed Boilers", Report NSF/RA-760S90. NTIS PB 268-558,
November 1976.
2. Westinghouse Electric Corporation, "Clean Power Generation from Coal", O.C.R.
84, NTIS PB 234-188, April 1974.
3. Moke, R.C. et al., "A 'rtegcncrativc Limestone Process for Fluidizcd Bed Coal
Combustion and Desulfurization", Monthly Report 87, 1977.
4. Svcrdrup, E.F. and I).II. Archer, "The Tolerance of Large Gas Turbines to Rocks,
Dusts, and Chemical Corrodants", presented at EPA/ERDA Symposium on High
Temperature and Pressure Paniculate Control, Kashington, DC, September 1977.
5. Squires, A.M. and R. Pfoffor, "Pane! Bed Filters for Sinultaneous Removal of
Fly Ash and Sulfur Dioxide", -I. APCA 2_0: 534-538, 1970.
6. Ranz, N.i;. and .I.B. h'onp, "Inpaction of Dust and Smoke Particles", Ind. Eng.
Chcm. 4ฃ: 1371-1381, 1952.
7. Marple, V. The Fundamental Study of Incrtial Inapctors. Ph.D. Thesis,
University of Minnesota, 1970.
8. Stern, A.C. , U.K. Zcllcr, and A.I. Schckman, "Collection Efficiency of .let
Impactors at Reduced Pressures", Ind. and ling. Fundamentals, _[: 273, 1962.
9. Mercer, T.T. and R. G. Stafford, "Impact ion forn Round Jets", Ann. Occupational
llygicncc, ^2: 41-48, 1969.
10. I'aretsky, I.., L. Theodore, R. Pfeffcr, and A.M. Squires, "Panel Hed Filters
for Simultaneous Removal of Fly Ash and Sulfur Dioxide: II Filtration of
Diluted Aerosol by Sand ซcds",'j. APCA, 2ฑ: 204-209, 1971.
11. Knettig, P. and J.M. Becckmans, "Capture of Monodispcrscd Aerosol Particles
in a F'ixed and in a Fluidizcd Bed", Canadian J. of Chcm. ling., 52: 703-706,1974.
12. McCain, .T.D., "Evaluation of Rexnord Gravel Bed Filter", EPA 600/2-76-;64,
NTIS PB 225-095, June 1976.
13. Ergun, S. "Fluid Flow Through Packed Columns", Chcm. Eng. Prog., 48: 89-94,1952.
14. Hood, K.T. "Evaluation of the Combustion Power Company Moving Gravel Bed Dry
Scrubber on the Control of Paniculate Emissions from a Hog-Fired Boiler",
NCASI Special Report, September 1976.
15. llokc, R.C., "Ducon Gravel Bed Filter Testing", presented at EPA/ERDA Symposium
on High Temperature and Pressure Particulate Control, Washington, DC,
September 1977.
532
-------
QUESTIONS/RESPONSES/COMMENTS
SPEAKER (from the floor): I guess I an developing some objec-
tions to both DOE's and EPA's continued, I'll call it, plowing the
sane ground, relative to granular bed filters. I think the state of
the art in the technology really hasn't been advanced at al1 in the
last three years. As I said, I keep hearing the same conclusions,
the same problems, meeting after meeting.
Now relative to some of the fundamental issues here, when you
are talking about the pressurized fluidizcti bed, I think that one of
the underlying issues relative to the entire cleanup system is,
indeed, what can the gas turbine tolerate? I know I have heard at
least three papers today saying, "Here is the range of estimates." I
think considerably more work should be put into determining just what
a reasonable tolerance limit is for a gas turbine. That is something
that's been lacking right from the beginning of the whole effort. I
think when you are talking about meeting the EPA emission standards,
if you meet them while you are protecting the turbine, fine, but we
certainly don't need high-pressure or high-temperature cleanup systems
to meet EPA standards when you can always do it downstream of the
turbine.
I also think relative to the loading specs that it's very
obvious that the loading is a function of the size of the particles;
and the gross range per standard cubic foot of the tolerances are just
not applicable. It's very dependent upon the size of the particulate.
So again, that gets back to the issue tnat someone better, at some
point, start taking a good h.ird look at establishing some tolerances.
If one does know what the gas cleanup has got to do, then I would like
to see both DOE and EPA take a more mature look at this background and
stop really supporting a lot of surveys that have gone on and on. I
want to see some firm new ideas and new concepts developed for parti-
culate control. Thank you for the soap box.
CHAIRMAN: Other comments? Yes.
SPEAKER (from the floor): We have been trying to build a turbine
which would eject particulate material for 40 years. The permissible
limits for particulates in the gas stream from the standpoint of tur-
bine bucket erosion, I think, are very soundly based on that experi-
ence. I don't think there is any point in going into the details here.
But let's not kid ourselves: we are not going to be able to eject
large particulates from the gas turbines, no matter how you dislike
it. Unfortunately sandblasting is a very effective way of removing
material from the turbine and the effects of particulates on gas
turbine buckets.
533
-------
CHAIRMAN: Thank you. Other questions? Go ahead.
MR. SCHWARTZ: Michael Schwartz from Shell Development. This
is essentially a general question. What consideration has been
given to generalizing this technology to additional applications?
What I am thinking of in particular are high-temperature, highly-
fouling particulate streams which arise from incineration.
CHAIRMAN: Are there people that would like to comment to that?
SPEAKER (from the floor): I guess not very much consideration.
CHAIRMAN: I think there was another comment or question here.
SPEAKER (from the floor): I will ask a simple question. Dr.
Parker mentioned several times in his talk that he does not think
that at high temperature and high pressure filtration is by filter
cake. Could he elaborate on what basis does he expect this. Just to
look at it or what?
DR. PARKER: The question was, what was my basis for saying
that I did not think there was any filter cake formed in these appli-
cations at high temperature and pressure. I did not mean to imply
that there never could be a filter cake formed. But in what has
been reported on the experiments at Exxon, the tests of a Squires
granular bed filter at Morgantown, and the tests of Combustion Power
Company, is that there has been no evidence that there was a cake
formed. And we are using that as the basis for justifying, using a
clean bed model for these predictions.
SPEAKER (from the floor): Then how did they arrive at that,
that no filter cake was formed. By looking at the bed, or what?
How do they see it?
DR. PARKER: Well, all they say is there is no evidence that
there was a cake formed. What does that mean? They did not see any
cake; any evidence of a cake. That is, they don't see large clumps
of ash in the hopper after cleaning. They don't see any cake, I
guess, if they stop a test in the middle, they don't see a cake
formed before cleaning the bed.
SPEAKER (from the floor): Did you see a cake of fine particles
or big particles?
DR. PARKER: Well, a filter cake would be a cake like the cake
that Ron Hoke showed on his screen, on the Dueon filter. It will be
on the surface of the bed, and it will be separate from the bed
material; there will be a cake of granules.
534
-------
SPEAKER (from the floor): Doesn't cleaning cause the pressure
drop to go down?
DR. PARKER: Uell, the pressure drop will go up as the bed loads
up anyway, no matter where the particles are collecting.
SPEAKER (from the floor): Can I comment on that?
CHAIRMAN: Sure.
SPEAKER (from the floor): Well, we have taken clean sand in the
Ducon filter, and we have put in dust. We look at the surface; and
let's take an ordinary kerosene flame, where we sucked the smoke right
through the bed, we get the surface black. Now, I don't know whether
you want to call that "cake7 or not. Let's put it this way. We see
penetration into the bed about a half-inch, and it's a gradient in
color; so the surface is much, much blacker than anything else. So
you look at the bed, and at the top the surface is black, you say,
"Oh, well, it's all there." But if you want to call that a layer,
it's a hell of a thin layer. It's so thin thatwell, it's hardly a
cake. Now, we only horse around with pressure drops on the order of
six inches at most. If we let the pressure drop build up to several
pounds maybe you would see a layer that would be a more significant
layer. But just taking the ordinary fixed bed pressure drop correla-
tions, and assume a particle that has collecteda dust of a few
microns diameterand you will see that the layer has to be almost out
of sight, and you've already got a tremendous pressure drop.
So I don't know whether I am answering the question; but the
layer that you would see, if you had a layer, would be hellishly
thick, put it that way; and I suppose that's really what Dr. Parker
is saying. He got away with this correlation on the basis that it
is more impaction than layer.
DR. PARKER: Right. Well, a lot of people have proposed that
granular bed filtration would be similar to fabric filtration, but
this seems to be a basic difference between the two. It is a lot
more difficult to build up a discrete cake on the surface of a bed
of granules.
CHAIRMAN: Yes.
MR. GOREN: Simon Goren from the National Science Foundation.
With regard to your bench-scale experiments, would you tell me what
the aerosol test particle was, the gradient velocities, the grain
size, and to what extent were your beds loaded, or cleaned very
frequently.
535
-------
DR. PARKER: We used polystyrene particles ranging fron 1/2 to
two microns, and in very low concentrations, so that the total
amount of participate collected was small enough to assume that the
bed was clean. We did not take the granules out and wa'.h them after
each test, but there were very low concentrations of polystyrene
particles. Superficial space velocities were around 40 centimeters
per second, probably ranging from about 25 to 65, but I'm not certain
about that. We tested various types of bed materials. We looked at
glass beads about 1 millimeter in diameter. We looked at iron shot
ranging fron 500 microns to 1 millimeter. We looked at sand, which
was an average of about 400 microns; coarse sand, and there may be
others. I can't think of any now.
CHAIRMAN: Thank you very much. Were there other consents or
questions? I thought I saw one more hand over here. Okay. Let's
see. Maybe just a couple of comments.
I think certainly the need for experimental data, as it has
been made in several comments, is very critical and very important
in terms of understanding and developing an effective particulate
control system that is going to be required for pressurized fluidized
bed combustors. I think the other point that was raised over here is
certainly something that we would agree with, and I think most other
people would concur with, in terms of the need to have more work to
extend our understanding of what the tolerances are for particulates
and for deposition from fluidized bed combustion systems, building
upon the broad base of knowledge which does exist, but has not really
been fully understood for systems where you have limestone beds with
char and ash coming off the beds, an area which certainly requires
more definitive work if we are really to expeditiously move forward
and develop commercial systems.
536
-------
INTRODUCTION
DALฃ KEAIRNS, CHAIRMAN: We will now have three papers on noving
beds and finish up with a paper on participate control for fluid bed
combustion processes. The first paper which we will have on moving
beds is entitled "Particulate Renoval from Pressurized Hot Gas."
It will be presented by Hiroshi Teradu of the Babcock Hitachi Company.
He graduated fron the N'agoya Institute of Technology and entered
Hitachi Research Laboratory in 196n. He is mainly engaged in research
on conbustion oscillation, nitric oxide emission fron combustion pro-
cesses, and fluidized bed combustion. He is now the chief researcher
for the coal gasification program contracted with the Resources and
Energy Agencv.
537
-------
Paniculate Removal from Pressurized
Hot Gas
Reijiro Yamamura
Japan Coal Mining Research Centre
Tokyo.Japan
HiroshiTerada
Babcock Hitachi Company
Hiroshima, Japan
ABSTRACT
This paper describes cold and hoc test programs for developing a new parciculate
removal system which contains moving bed filters and fludiized bed regeneration of
alumina balls. Based on cold modelling ttscs studying si ch problems as movement of
ball, in the bed. control of ball discharge- rate and dust removal efficiency, we have
concluded that this system should be very feasible. Hot gas programs have been carried
out in the use- of a one-sragc test filter set for the raw gas line of a 5 T/D pressurized
coal gasifier in Vubari City. We confirmed that we could attain high removal efficiency
of about 90 to 97%. This indicates that a two or three-stage parffculate removal system
will satisfy the inlet condition of a gas turbine.
As you know, high temperature dust removal systems are required for fluidized
bed combustion or for gasification in combined cycle power generation as well as for
hij^h temperature desulfurization. Dust removal systems will require rapid development
as electrostatic precipitators and other conventional techniques will not be enough.
In the United States many projects are under study as is demonstrated by the attendance
here today. In our project, we are studying and developing a filtration system usinp
a moving bed consisting of ceramic balls. Although, in Japan, no pressurized FBC
projects have been started yet, gasification projects have been undertaken since 197"ป.
We are focussing much attention on the development of high temperature dust removal
equipment as one of the important elements of the gasification project. 1 will explain
the outline of our program here.
By the way, we at Babcock Hitachi and Coal Mining Research Centre in Japan have
been assigned to this research by the Resources and Energy Agency of the Japanese
Government.
Figure 1 shows the concept of our system. Dirty gas will be cleaned after passing
through the moving bed filled with alumina balls. Dust containing alumina balls will
enter into a separator beine controlled at a suitable flow rate. In this separator the
dust is removed from the balls, and the balls are heated up if necessary. They will
then enter the moving bed filter again. The important technical points of this system
are mentioned on the figure.
Figure 2 shows our research and development schedule of a moving bed filter to
realize the concept mentioned in Figure 1. Our schedule has been set up to keep the
pace with the development of a coal gasifier. Starting from a feasibility study and a
conceptual design in 1975, we launched tests using cold model equipment in 1976. In
this period, we tested the dust removal performance using simulated dusts. In 1977, we
ran a recirculation test in the system which included dust separation. In parallel with
this, we are carrying out a hot gas test with a test rig connected to a 5 T/D gasifier.
The plan is underway to construct an 8,000 normal cubic meter pilot plant for a 40 T/D
coal gasifier.
Today's report is focused on the cold test and hot gas test results (see Figure 3).
In the cold model tests, we tried to have a smooth flow of alumina balls with a satis-
factory flow rate control, and also measured dust removal performance. The main parts
of equipment are made of clear plastic. Alumina balls supplied from the feed hopper will
move through the moving bed filter with a moderate speed, and the dust will be separated
in the fluidized bed separator. After that, the balls return to the moving bed filter.
The simulated dust supplied from a constant dust feeder will be suspended in the air
stream.
538
-------
HEAT
EXCHANGER,
ATMOSPHERIC |
SEPARATOR
[SEPARATOR
AND
[ BALL HEATER
T
I
(5) DUST SEPARATION
1(6) HEATING OF ALUMINA BALL
A ALUMINA BALL
-At-, 'N
GAS OUT
(1) STRUCTURE OF
MOVING BED
(2) C'JST REMOVAL
FLOW RATE
CONTROLLER
| ALUMINA BALL
a OUT
Figure 1. Contact of Dim Removal System* end Its Problems
FEASIBILITY STUDY
CONCEPTUAL DESIGN
COLD MODEL TEST
PRESSURIZED HOT
GAS TEST
PILOT PLANT
8,000 Nm 3/h
GASIFICATION
COMBINED CYCLE
POWER SYSTEMS
'75
T
>
i
t
OESI
CON
'76
t
\
>
t
F
'77
rousT
k
Y.
'78
'79
REMOVAL
FRECIRCULA
STSTEMS
1
WTfซ:f RIO
iCONSTRUC
1 T
ION
f
MOT CAS TEST
ON AN
STRUC
0
TIONj
RUN
51/0 GASiriER
i
301
k
NINC 1
T
ซOI/oGASlVlCATIC
c
I* PO
KERF
LANT/
1
rEST/
'80
ION
>
k
>N PliNT 1
ป
REMARKS
5 t/e GASIFIER
(YUBARI CITY)
40 t/o GASIFIER
(YUBARI CITY)
RESOURCES AND
ENERGY AGENCY
COAL MINING
RESEARCH CENTER
Figured R & D Schedule ol Moving Bed Filter for
Presuirind Hot Ga Cloning
539
-------
OUST MONITOR
FEED HOPPER
FABRIC
FILTER
MOVING BED FILTER
DUST MONITOR
CUST FEEDER
CYCLONE
FLUIDIZEO BF.O /
SEPARATOR
COMPRESSOR
Figure 3. Schematic Diagram of the Cold Model Test Rig
Figure U shows an outside view of the cold test rig. In this case, beU depth is
500 mm. The partition plate is one of the important elements of the iroving bed filter
as the movement of alumina balls near it will tiiifer significantly depending on the
type of partition plate. In this case it is a louver type.
In Figure 5 you can see the pneumatic recirculation line for system rccirculation
tests. The fiuiJizcd bed separator is located above this line.
In the cold model tests the selection and supply method for the simulated dust, is
very important, because the simulation of their physical properties affects the success
of this test.
In Figure 6. a sample of char collected from the test gasificr is shown (Column A).
We also used the simulated dust as shown in Columns B and C. We are going to mention
the results of our cold model test with several figures now; but, we would like to say
beforehand that our tests results can't be assumed to be quantitatively a true simulation
of dust removal from actual gases. Our test was only for setting up a qualitative con-
cept for a design. Therefore, the quantitative study should be made from the results
of operations with actual dust-lader gases.
Figure 7 shows the effect of bed depth and alumina ball diameter on outlet dust
concentration. Three different ball diameters were investigated, namely 0.5, 1, and
2 mm, untjer the condition of a gas velocity of 0.5 n/sec, and the inlet dust concentration
of lgm/mj have been used. As you can see, the outlet dust concentration decreases
logarithmically as the bed depth increases. It is very interesting here that the outlet
dust concentration increases in the smallest diameter of 0.5 mm compared with 2 mm.
The crussplotted graph of outlet dust concentration against alumina ball diameter
directly shows this tendency. In short, the outlet dust concentration is the least for
the ball diameter of 2 mm in this case. Consequently, the highest dust removal efficiency
can be obtained in the 2 mm ball diameter. This tendency can be seen as the type of
dust varies; therefore, ic can be considered that an optimum ball diameter must exist.
540
-------
Fijurt 4. Phosoflrtph of tht Cold MocM Tซt Rifl
FigurcS. Ptiotogriph of th* Moving Bซd FilMr of Cold MocM
541
-------
to
05
BWTIOE
Figured Pfoptfty of Sซnutat*d Dun for CoU Model Tut
00.1
100 30O 500
BED DEPTH (ren>)
12345
ALUMINA BALL OIA. (mm)
Figure?. Effect ol Bซd Depth & BaO Du. Outtat Out Cone.
542
-------
OUTLET DUST CONCENTRATION (g/Nmป)
P P - w o
ป v * wo o O
MOVING RATE OF BALL
6 Cm/min
INLET OUST CONG.
1 g/Nm'
_jf -~8~"
^r
BALL DIA. 1mm
BED DEPTH 100mm
A a~"
A^HT
X-5T
BALL OIA. 2
BED DEPTH 30Omm
0.5
6AS VELOCITY (m/ซ)
1.0
^
z
o
til
o
o
o
50
o
ujO.1
GAS VELOCITY: 0.5m/*
INLET DUST CONC.1g/Nm>
^ -g
BALL DIA.
HFn OFPTH
9
1mm
Jf^* *"
BALL DIA.
BED DEPTH
1 1 1
Zfflffl
300mm
0 2 4 6 8 10
MOVIN3 RATE OF BALL(cm/min)
Figurซ 8. Effect of Gas Velocity and Moving Rate of Ball on Outlet Dust Concentiation
Figure 8 shows the effect of gas vclocit/ .tnd moving rate of balls on the outlet
dust concentration. We can sec from this figure that ns the gas velocity becomes
low, the outlet dust concentration decreases; and as the moving rate of the ball in-
creases, the outlet dust concentration will increase to some extent also. This increase
seems to be larger with the increase of the ball diameter. But as the moving rate of
the balls continues to climb, its ultimate effects seen to be minor.
The dust accumulation on alumina balls of a unit volume in the moving bed varies
with the conditions of the moving bed and inlet gas. In other words, when the inlet
dust concentration is high or the moving rate of the ball is low, accumulation will be
high. As can be seen from Figure 9, the dust removal efficiency increases with the
increase of an average dust accumulation in the early stage; but, it reaches a peak at
some point and afterwards it decreases with the increase of average dust accumulation.
Therefore, we can see that there is a critical point in the relationship between
average dust accumulation and dust removal efficiency. For this reason, we consider it
totter to think carefully on these critical points.
When connecting with a gas turbine, small diameter and good quality are required
for dust at the turbine inlet, as well as for the amount of dust. Aside from the
quality of dust, the diameters of dusts collected from the moving bed filter outlet
were compared with those at the inlet (see Figure 10). As you can see, the average
dust diameter decreased at the outlet compared with that at the inlet. And therefore,
this moving bed filter can be considered to be effective for the dusts smaller than
10 microns.
Figure 11 shows the pressure drop through the filter for varying conditions of
ball diameter and distance from gas inlet.
In the following several figures, we will mention hot gas tests connecting our
moving bed filter with a gas line of 5 T/D gasifier under investigation at Yubari City
in Japan.
543
-------
0 5 10 15 20
AVERAGE OUST ACCUMULATCN (Q/L BAa)
ttyan 9. Oust Removal Efficancy vs. Atmagt U-
-------
f
too
ao
60
40
20
0
INLET OUST CONC.
0 ex
GAS
GAS VEL.
MOVING RATE
20 40 60 60 100 120
DISTANCE FROM GAS INLET (--)
Figure 11. Prcuun Drop in Moving Bed Filter
Figure 12 shows a flowsheet of the hot gas test equipment. The design condition
of gas flow rate is 200 KnjJ/hr. the pressure is 10 Kg/craZ, and the temperature is 700ฐC.
The main parts of equipment will be pressurized. The flowing rate of alumina balls was
controlled by a rotary valve, and dust concentration was measured with a cylindrical
filter paper at botn the inlet and outlet of the moving bed filter. As the Oow of the
alumina balls was run through, the running time was only 1 to 3 hours.
Figure 13 shows an outside view of the hot gas test equipment. Here we can see
the moving bed filter and feed hopper. The gas-supply sectional area in the moving
bed filter is 1 foot by 1.5 feet, and the bed depth is 8 inches.
Figure 14 shows an outside view of the rotary valve used to control the flow rate
of alumina balls. We can see a spent ball hopper under the rotary valve.
Figure 15 shows the dust removal performance in a hot pas test. We can see. just
as in the cold test, that the outlet dust concentration increases with the gas velocity
in the range of 0.2 to 0.6 m/sec in the moving bed filter. The curve of the dust
removal efficiency shows a high rate of 90 to 97 percent. The figure on the right
side shows the effect of the moving rate of alumina balls. The efficiency decreased
with the increase of moving rate of the balls, but it eventually becomes asymptotic.
As you can see here, these tests were carried out under the condition of pretty
high dust accumulation, so it will be expected that the removal efficiency must be
raised further. Anyway, these values cannot meet the demand of the gas turbine. In a
commercial operation, it will be necessary to provide for 2 to 3 stages of novinz KcJ
filtration.
Figure 16 shows micrographs of dusts sampled at both inlet and outlet of the
moving bed filter in the hot gas tests. These are char consisting of about 50X fixed
carbon, as already mentioned.
After completion of the test, the alumina balls were taken out through the
fluidized bed separator. In this case, the char is very flaky, and it easily separates
from the balls. During the present tests, the gas temperature is about 500ฐC. There
is not ouch adhesion of tar to the balls.
545
-------
RAW GAS LINE
PRESSURE
CONTROL VALVE
FLUIOIZEO BED
SEPARATOR
GAS FLOW RATE : 200 Ntn 3/h
PRESSURE : 10 Kg/cm*
TEMPERATURE : 700 ฐC
Figure 12. Simplified Equipment Flomhett of Hot C i Tett
Figure 13. Photograph of Moving Bed Filter of Hot Gซ Ten Rig
546
-------
F.jure 14. Photograph of Rotwy Valw of Hot CM Tซ Rig
1
BED DEPTH 300
BALL DIA. 2-
INLET OUST CONC
X
BED DEPTH 300"
BALL DIA. 2"
INLET DUST CQNC.
0.2 OA OS OJB
GAS VELOaTY (V.)
02468
MOVING RATE OF BALLfX.)
Figun 15. Effect of CM Velocity and Moving Ratป of Ball on
Outlet Dint Concentration
547
-------
M.B.FILTER INLET
ME AN PARTICLE DIA.9.MI
M.B.FILTER OUTLET
MEAN PARTICLE DHL
-------
AiUM.'NA BALL
MCVltC BEU
FILTER VESST.I
RA.V
^v
AI.UMIMA BAIL
FUJIL>2LD BED
BALI RtGOCRATOR
LIFT GAS
: 710 *C
COLD MOOEL TEST
IT V
-------
INTRODUCTION
MR. KEAIRNS: This is the session on Emission Control and
Participate Removal. I am Dale Keairr.s with Westinghouse R&D Center,
and I'm co-chairman of this session with Dennis Drehmel, who is with
the Particulate Technology branch of EPA. It is a pleasure to have
the opportunity to chair this session, which I think covers a very
important aspect of the problem of fluidized bed combustion, and a
very challenging problem in the sense that one has a variety of
moving targets. Depending on how you design the bed and what beds
you are talking about, the particulate control people don't know what
they have to deal with coming in, and the turbine tolerance require-
ments are somewhat of a moving target, depending on the design, and
as we increase our understanding of what the tolerances are. Then
there are the environmental aspects, in terms of what tolerances will
be required from an environmental point of view, both from dust
loading and particle size distribution. So I think the boundary
conditions on this problem present a very interesting and a dynamic
problem.
I would also then like to mention, and hopefully it is being
announced in the other parallel sessions this afternoon, that the
first paper will not be presented. Professor Gut finger will not be
here this afternoon, and so that paper will not be held, although it
will be included in the proceedings of the conference.
550
-------
Particulate Removal from Hot Gases
Using the Fluidized Bed Cross-Flow
Filter
Chain Gutfinger, G.I. Tardos and
David Oegani
Department of Mechanical Engineering
Technion - Israel Institute of Technology
Haifa. Israel
ABSTRACT
This work describes a granular filter for separation of particulars from a hot gas.
The **rtvice operates in a steady state mode by means of continuous introduction of clean filter
granu.es from the top and removal of dirty granules fron the bottom. The qas flows trans-
versely across the moving active filter material. A bench scale sand filter of 100 cm; cross-
section was constructed and tested during this work and its filtration and hvdrodynamic charac-
teristics are reported. A theoretical model for the computation of filtration efficiencies is
also presented.
IirTRODUCTION
This paper deals with filtration of hot gases containing particulatcs in the size
range of 0.01 to 10 urn by cleans of a moving, continuously regenerating cross-flow filter. The
existing methods for rcraov.il of particles in this size ranqo are electrostatic precipitation
(0.01 - ID urn); packed beds (0.01 - 100); high efficiency-cyclones (0.1 - 3.0); cloth collec-
tors (0.1 - 60); and liquid scrubbers (0.3 - 100). Unfortunately, these methods do not lend
themselves to continuous operation at elevated temperatures either due to their very high cost
(electrostatics), filter clogging (packed beds), high erosion rates or material limitations
(cyclones, bag filters, etc.).
During the past several years the fluidizcd bed filter has been extensively investigated
by different authors [l, 2, 3, 4, 5, 6]. The reported results ir.di.atc clearly that fiuidizcd
t*d filters collect efficiently particles in the size range O.01 - 1O urn and have the advantage
of regeneration by continuous withdrawal and replenishment of bed material. The disadvantage of the
fluidized bed as a dust collecting filter is in its limitation to low qas velocities. These in
turn require high filter areas for flowrates common to industrial systems. Thus, even thnuqh it
was clearly demonstrated that a particle-bod is a useful means for cleaning dirty gases, it was
necessary to expand further effort toward the investigation of a regenerative granular bed filter
J.ble to operate at high gas velocities and elevated temperatures encountered in power plants,
cement factories, fluidizcd bed combustion systems and/or gas turbine cycles.
The present work evaluates the cross-flow filter as a continuous separating device of
small particulatcs from a gas stream. Experiments performed with du~ty gases at 30 C and TOO C
indicate that the cross-flow filter features good separation characteristics, low pressure drop
and high gas flowrate capacity. The experiments and theory reported in this paper can serve as
a basis for the design of -. dustrial granular filters.
FILTRATION EXPERIMENTS
The laboratory scale cross-flow filter device used in this work is presented schematically
in Figs. 1 and 2. It is essentially a moving sand bed. The filter operates in a steady state
node by cieans of continuous introduction of clean granules from the top and removal of dirty gra-
nules from tho bottoa. The gas stream flews transversely across the moving active filter material.
Fluidization of the granules is used in order to allow the removal of dirty filter elements froa
the bed. The sand bed is held in position by two 45 mesh brass screens on the gas inlet and out-
let sides. Thซ active filter thickness is about 10 cm and the surface area normal to the gas
flow is about 100 cm*. The device is operated with quartz sand in the size range 2O - 40 mesh
(40O - 810 u). This sand is capable of continuously withstanding operating tcnperatures up to
1200ฐC.
551
-------
FILTER ELEMENTS
INLET
FLUIDIZATION
AIR
GAS
OUTLET
...._. /30MESH
SANฐ 140 MESH
FILTER ELEMENTS
OUTLET
Figure 1. The fluidizad bed croo flow filter
552
-------
Figure 2. View of the f luidued bed cross-flow filtration system
553
-------
The device is used to clean the exhaust gases coning from a 2.5 HP "Peter' Diesel engine.
These gases were chosen in order to test the filter with industrial pollutants and to reveal
specific difficulties which nay arise during industrial use of the system. The test lovp is
given schematically in r'ig. 3. The exhaust gases pass through a cooler and a cyclone before they
enter the filter. This precaution is necussary in order to renove coarse oil and water drops
from the gases before they reach the inlet screen of the filter. For high temperature .ests the
cooler is disconnected. A gas dilution system with compressed air is also connected to the gas
inlet in order to increase the gas velocity in the bed. The dirty gas nay be passed either
through the filter or through a by-pasr. The gas is sampled isokinetically from a point downstream
the filter. A pressure tap upstrean the filter is connected to a differential manometer.
The pressure drop through the filter was measured as a function of the superficial gas
velocity. These experiments were carried out with compressed air. The results are presented in
Fig. 4, for differei.*ii sand sizes. The total pressure loss through the device varied between
8 - 21 en HO for all gas vel'.'.-ities between IS - 60 cra/.~?c. The "smoke* produced by the Diesel
engine was characterized usir a TSI and a Royco particle size and concentration tr/mitcrs [?].
A typical smoke characteristic is reproduced in Fig. 5 in terns of the number of dust (carbon)
particles per cm- ot ga:; vs. cho particle size* for the specific case of 1 IIP load o. tl.i engine
and JOOO rpm* Au seen, the maximum numh'.r of carbon particles in the smoke is in the diameter
rai.gซ of 0.1 - O.iy, while particles coarser than lu are almost absent, from this figure the
weight concentration of particulatcs was calculated as 1.4 g/n . These data arc subjf.- *. to appre-
ciable changes depending on the engine Ic**!, amount of dilution-air and particle con*:nt.
Kig. 6 presents dust separation characteristics of the fluidized bed cross-flow filter
obtained at roon temperature (JOฐC). The filter was tested with dust naturally present ir. a com-
pressed air stream and with smoko produced by the engine. The difference between particles from
these twn sources is in their sticking properties. The compressed air dust is dryer ar.d therefore
less sticky. This is the reason why carbon particles are separated with somewhat higher efficiency
than atmospheric ttust. The filter was o|>cratcd in the packed and fluidized bod mode with gas velo-
cities between 18 - SO cm/sec, for about 50 hours without a sensible pressure drop increi-e in
time. Approximately 30 kg of sand was continuously circulated through the system during this
period. No efficiency decrease was observed in time throughout the experiment. As seen in the
figure, tho efficiency of the packed bed decreases from about 70\ to OO\ as the cas velocity
increases while the efficiency of the fluidized bed decreases from about un\ to 71* for similar
conditions. The ouch higher efficiency of the fluidized bed is due to electrostatic charges
produced on tho bed granules during fluidization [&].
Thcie experiments show that the cross-ilOK filter has the good properties of the
fluidizeti bed filter such as high efficiency* low pressure drop and continuous operation and
in addition allow-, high gas velocities through the bed. Another good feature of the filter is
its vertical configuration which requires less floor si>ace as compared to tho fluidized bod filter.
Tho active filter area may be increased by increasing the height of the bed. The filler material
(sand) consumption between two consecutive cleanings proved to be very small.
Fig. 7 nresents total efficiencies, n, of the cross-flow filter as a function of dust parti-
cle diameter, with -.he gas temperature as a parameter. The experiment was carried out with snoV.e
from the Oi> jel engine, diluted with hot compressed air. The gas velocity was U =87cm/sec. The
results were obtained at mean gas temperatures of 30, 195 and 205ฐC. The influence of elevated gas
temperature on the total efficiency may be observed: efficiency increased for fine dust particles
(2r <0.5u), whereas the efficiency of coarser particles, (0.5<2r
-------
rn.TOป
n
MTUKTUt
-04-
CTOOK
tXBTT riLTOI
O.CMMT *CMOWU.
THUS
MWTKUS
VC MDMU. VATCM
fllCT
1'
,
HTia co
-------
Sot*
SOmttk-
40lMih-42O-StOu
Got moeity to H* bed.
Fig/an 4. Relationship between pretsure drop *nd vปtoerty in the Ihjidind
bed cross flow filler
..
aoi aoz o-os o-i 0-2
Owl porlicto )< 2rp [/*]
Figura 5. Oust parlid* conesntrttion n pvtid* til*. Sonka produced by
a 2.5 hp diซel engtn*.
POUMT: 1 hp Rotations: 3000 rpjit.
556
-------
IUU
90
80
T 70
u
1 60
*:
S I 50
il
40
0
i iii iii
a
o
0 A A o o -
A oA o oo
x ' X
f .
X X
A A -
A A A
-
Eiporiitwital
ro-u0- IB cm/tec \cซ>mPซซ8ซd (L-l5cm)
Packed K.v.nsm/MC ฐ"
. . 'ปmoke (LซIOcm)
Fluidizedf A-U0ซ 60 cm/we
69(1 \a-Uo'87cm/*ec J
I III III
01 002 005 01 02 05 1-0 2 5
Dust particle diameter 2rp [p]
Figure 6. EKiciencv of the crou llo* filter
2040 ireth sand d -=04)
-------
100
80
S
I "
20
--- 30C
Sar.d 20-40 oe*h (t-0.4)
CAS velocity U -87 cn/iec
Filter thickness L-10 cm
0.01
0.02
0.05
0.1
0.2
DUST PARTICLE DIAMETER 2r (u)
P
X
0.5
1.0
Flgur*?. EHiciซncy of aoo-flow filtar at aWvstad ga tempoituto
-------
where, c, is the porosity of the filter bed and a, is the filter element (granule) radius.
Typical values of granular-filter porosity, c, are in the range O.4 - O.S, hence, the total
efficiency n, depends mainly on three paraaeters: filter height, filter element dianeter and
single filter element efficiency, E. Of these only the latter is know a priori and has to be
either computed or measured experimentally. We suamarize now the basic steps of the filtration
theory leading to the computation of, E. A more detailed exposition on this topic is given in
the literature [?, 8].
The analysis considers a single dust particle approaching a spherical filter element. It
is assumed that any dust particle hitting a filter element is retained by the filter. The mecha-
nism of separation nay be inertial, interceptional, diffusional, electrical, etc., depending on
the combination of forces which act on the dust particles. Trajectories of these particles in the
vicinity of the filter element nay be calculated. Out of all trajectories one is interested in
the "limit trajectory," i.e., the trajectory which just misses the collector, (filter elenent).
Due to the axisynaetrical nature of the problem, any trajectory contained within a streaatub*
generated by the limit trajectories will meet the collector resulting in particle separation.
Hence, for a filter element of diameter 2a and limit trajectory streamtube of diameter 2b. the
single element efficiency is defined as:
E - (|,2 ,2)
The equation of the trajectory of a dust particle approaching a filter element, was presen-
ted by different authors, [l. 3. 8, 12]:
where x is the position vector, i ,
and u is the gas velocity. Eq. (3) may be put in non-dimensional fora by using
U u>Uo. x - x/(a * r ), T U0t/(a +r )
which yield
..2? 1 * *.. - ^ ?__a(l * "J
(4)
Here, U is the dioer. .ionloss gas velocity, St * 2Cp UQr 2/(9ua) is the so-called Stokes number and
R r /a is the Interception parameter. The equation of motion as written above is general enough
P P
to incorporate any effect that influences the notion of the dust particle. This is accomplished by
introducing the proper effect through the external force F . If several effects coexist, they
may be Added together at this point.
Equation (4) Bay be written for the particular case where only the effects of inertia,
interception and diffusion is considered as [7. s]:
(5)
where the diffusion velocity Vn can be computed as follows [l3]i
D
U Pe N dR
o
As seen, V depends on the Peclet number, Pe. the dust particle concentration, N, and the concen-
ts
tration gradient, dN/dR. The latter two quantities are calculated fron the diffusion equation and
substituted bock through Eq. (6) into Eq. (i>. CT. (S) nay be integrated numerically. This integ-
ration leada to the limit trajectory froti which the singlo filter element efficiency, E, follows
directly froa Eq. (2).
559
-------
rig. 11, presents confuted values of E as a function of the S'_c>:es nunbcr. St, with Peclot
, Pe, as art better than the predicted efficiencies. This
ifference it: due to electrical charging of the filter granules by friction effect, which improves
erration ef ficiei.cy. Fig. 10, reproduced fron a previous work on fluidized bed filters L^J*
how:ป the effect of electrical chare on filtration efficienc.
show:, the effect of electrical charge on filtration efficiency
It can bซ clearly seen that the theory predicts reasonably well efficiencirs in uncharged
filtration systems. It is known that charging due to friction is reduced at elevated tempera-
tures and/or high gas humidities, hence, the theoretical results are to ix: preferred for a con-
servative design of such systems.
Use of Present IXita to Predict Filter Pertoi.-n.vico
Fig. 11 is a design nomograph based on experimental data given tr. Fig. f>, together with the
measured pressure drop data. The designer first selects a qas velocity and a filter thickness,
A horizontal line at constant value ol L will yield n as a function of dust pjrticl'i diameter as
will as the pressure drop through the filter. Alternatively, givon the minimum ace. ptablo
efficiency for a certain dust diameter, the designer car look >ip directly tho fil'.ci depth, L,
arid the corresponding pressure drop. A similar noirogr-iph nay bo obtained for the caปe of high qas
ti'ioiH-ratu.';3 from Fig. 7 or by using theoretically predicted data.
CONCLUSIONS
The cross-flow granular filter rpovcd to separate snail p->rticulates from n eoi.tomirated gas
stream with reasonable efficiency. Hy ซi corresponding increase of tho filter thickness this effi-
ciency may be dramatically improved. Tho device can bo used at low as well as at higii gas tempera-
tures in places where dirty gases arn produced.
A full sire filter may be designed using the experimental or theoretical data rc.*x>rt:ed in
this |Hii>er. Tho design should be carried out using Fig. 11 for the case of low gas tei-jซraturcs
and using theoretical data from rig. h for tiio case of high qas temperatures.
ACKNOWLEDGEMENT
Tho sponsors of this research. KFA JOli-h and tiCRD. Israel arc gratefully acknow;edged.
NOMENCLATURE
a - filter clement (granule) radius
L - distance of limit trajectory froc* OX. axis
C - Tunninqhan correction factor
il - particle diffusion coefficient
f>(b/a) - sinqle sphcr*> efficiency
E ป4 (~) PC - pure diffu3ic:ial deposition efficiency.
F - external force acting on particle
L - filter thickness
m - rasa of dust particle
- dust particle concentration
- dust particle concentration, dinensionlcss
560
-------
0-002
0001
OOOJ
0-01
0-1
Stokm Numbtr St
10
Figum 8. Theoretical fJngta fitter element oHicfency varan ttckn nuntbtr coraktering
intfliti and diffinkxvB) dปpoซtion n coraunt twd porotity. c 0.4
561
-------
2
Ul
u
ฃ
u
V
Ul ?
rs>
a.
w
oป
to
50
20
10
05
02
0-1
005
002
001
Theory (C'O 4)
Uo "18 cm/Me
- Uo 33 cm/we
.- Uo*60cm/stc
I i i I
x -U0* 33 cm/sec
A-u0ซ60cm/ปซc
1 I - U0ซ 87 cmAec
I _ i _
> smoke
002 005 0-1 02 0-5 1-0 20 50
Oust particle diameter 2rp
Figure 9. Co(np*iion of xpcrimental dau with theory
10
-------
I
05
g 02
bl
0-1
U
^005
>ป
U
'5
0-5
0-1
005
002
Theory Eซ0-4, Uo" 18 cm/sec
Experimental
Uoฐ 18 cm/sec
Aerosol
Lotex oorosols
Carbon powder
Atmospheric dust
Zinc powder
Neutralized
bed
ซ
a
e
Charged
bed
*
Chorqcd Bed
001 005 0-1 0-20-51 235
2rp[/i] Dust particle diameter
Figun 10. ExpVBiwntsI ซnd tti*cw*ticซl vn<$* tphaซ eHicitncici in ntutnlind ซnd
ctaryd and bซd of 660n diamtttr grtnulcs (porosity c - 0.4)
563
-------
Prusure drop. Ap(em
O 10 ?O 3O 4O 50 60 TO 8O 9O tOO I!Q 120 130
80
70
I 50
o
ฃ 40
ฃ 30
20
10
^999%
FhjidiMd_bjd
Cos vtlocify
Uoซ60 cmAtc
Uoซ PT em/ปc X
/ ^J^.%
/
001 OO2 005 01 02 05 10
Oml pwlid* dnmtlcr 2fp [M]
10
Prtuurc drop. Ap[cm H,0)
-y y ?
0
-------
- dust. jซirticle cor.cer.tratio:. at ir.fir.ity
e-.'au /. - Svdot r.'jnbcr
o .
- ra'iius
- dinur.sior.loas radius
- at.-r'iiol Ciiust particle) radius
- lr.turci.-1-tio.'i
Strike.-; r.'jnfccr
t ine
- dirw;nsior.lซ-:;s tictu
- 1*0-; it ion w:tor
- i*>sitiQi. vector, ciiner.sionless
- fluid velociv/ vector
- fluid vป_-l\ por
- t'jt.al flltt.-r ( f i
Mui'l vi>.c*j';it.y
1. 1.. l-arutaky. L. T(iซ-xl-r.-. K. l-f<:{(ur .ind A.M. :;<|uir<:s. .t. Air Pol Int.. r.mtrol A-.a.. ฃ1^, I't71.
f. 204.
i. II. !. M^i^sivur ar.'l U.S. Micklcy, "Kv.il of Mists and Dusts f roo Air hy hods of Kluidiipd
Solids," Ir.-). Ji.ii l.i.-j:.. 12 IX.
I. '".. Tardon, C. <;uป f iri'(i-r ar.--|K)aition of Du-it Particlts in a I'luidircd Bod Ciltor,"
Ir.rjซ;l .1. Toc-h.. !_.'. 1(74, |.. 1H4.
4. ".. Tarilos, !.'. Al.-i.il and C. ^ut f ir.-icr. "tiif f u::iorul ril'.ratlon of Dust in a Fluidiiod Bed Filter,"
S. t-. Kr.ซ-ttinT and J.M. b'-ci'jwri-j, "Capture of v.modinpor'-.pd A<-ro3ol I'articlos In a Fixed and in a
fluidUcd tuyt.~ Oinadian .1 . _ of Ch<:n. Kn'j. ^2, 1174, p. 701.
fc. D. XcCarthy, A. .7." Vo"i.kr-l, K.ซ. l-at'turr.on and H.L. J.icknon, "Multistaifn Fluidizcd Bed Collection
of Aerosols." Ind. fn-j. rtn.-n. , rr-x-'-.-r-.s ;>?s. rv , IS. i', ri70, |>. 2t>n.
7. ^,1. Tardos, ":".ranulor Hซ?.
11. I.b. Stcchkina and II. A. Fucho. Ar.n. Occup. Hy and
12. H.J. l>ilat and A. Prea, "Calculated Particle Collection r.f f icienclcs of Sinqlo Droplets Including
Inortial lapaction, Brownian Diffusion, Dif fusiophoresls and Thernophoresis." Ata. r.r.v., 10,
l'>7fc, p. 13.
13. B.B. Bird. W.E. Stewart and E.H. Leightfoot, Transport rhcnoaena, J. Wilny I Son*. Now York,
London, 1960.
565
-------
INTRODUCTION
DALE KEAIRNS, CHAIRMAN: Thank you. Our next presentation will
be "Filtration Performance of a Moving Bed Granular Filter: Experimental
Cold-Flow Data." It will be presented by John Guillory of Combustion
Power Company, Inc.
Pr. r.inllory received his RS frori the University of Southwestern
Louisiana in 1ฐK? and t'S in nochanical onqinoorinq fron Louisiana
State, and a Ph.D. in nochanical onqineerinq fron DHahona State. Ho
is a roqistored profossional enqineor and holonqs to tho Anerican
Society of npchanicol Fnqinoors. His specialty is combustion and air
pollution control. Or. fiuillory is a part-tine instructor at the?
nivision of Fnqinoorinq, San Francisco State University. Hn is
currently tho supervisor of thomal and filter sciences at the
Conbustion Power f.onpany, Menlo Part, Talifi
566
-------
Filtration Performance of a
Moving Bed Granular Filter:
Experimental Cold Flow Data
J.L.Guillory
Combustion Power Company
ABSTRACT
Data fron jn KKDA- sponsored c-xporimt-nt.il st.jdy rel.itinq certain mechanical .in j 1 I untrjted in Fi-jure 1.
iv r t i c 1 ! collection i :; uccon[>l i shed l.y c.'iusinซ| the part icu l,,t <- laden <|.is to
.-ove r.t'ii.illy '.-jtvjr'l in cror.Mflow through an .intiulur, of ijranular collecting material
(nedial. Th" dovr.ward flow rate of the <;ranular bed is so lee tod such that the
deposition rate of part iculat e onto the irntiia and its removal rate from the active
filtr.itir.fi zone r'-sult in r.teaily-.Ttate operation which is --| excessive deiosition of [xirticulate ITI the gas
p,-i;;.".,i'l<:S. The outer screen is r.ized for complete retention of media.
In older to
-------
Figural.
568
-------
corrosion associated with certain na^eous pollutants can t>c nil. in: zed through
nainter.ance of olo/atcd temperatures.
Relatively Flat Collection Efficiency/Particle Size Curve.
The geor.etry of tho GBF is such that the various capture nui-cham s:ns (e.g.,
iripaction requiring high gas velocity arid di: fusion r<->:uirir:g r.orc1 modest
velocity) can be made effective wi;.nin the Sjrr.e unit. This characteristic
allovs capture of a given size particle L/v '. :ie riost r-ffici'iit Mechanism
thereby r.inir.i zitui utility eonsuniit ion.
Favorable Secondary Pollution characteristics.
Soluble [^articulate is not reintroduced to the i-nviron.Tcnt through use of
liquid scrubbing r.edia. Furthermore, since cuf.tured paniculate is concen-
trated into a small gas stream, it is t/rซซ'?l ie.il to use very sophist ic.iteU
(and ix-jssibly high specific utility consumption) final filtration to limit
atnospheric pollution as: 'Ciated with mofli.i cl'-aiiin-i.
TKST P
The objective of the tost proijran was to quantify the- Affect ซ,f n-rtiin design
and process variables on the particulate rol lection pซ.-r t'Tr.anc-'- of a ORF. The testing
effoit w,is divided into "subexper imento" represent ing specific curb mot ions of Jesign
i!iid '.perat lonal para-m.-ters. The following subexper iments will be considered in
detai 1 : 'J
ThicK Hซ'd Suhexper incnt
t 1 ', . J in
II - 51 in
;rr> = t.'i inai (122 lป>/tt3 bulk density)
NoTij tiaj _ Bed Sub- (peritnent
t = ' .'.- in
If = ^) in
"P. - 1. 'I rjn (132 Ib/ft3 bulk density)
:'.ri a 1 1 1cซlij Sulicjxyer imcnt
t = 7.f, in
II = 51 in
'n - 0.8 iwii (100 lb/ftj bulk density)
i:ach oubexpor ir.ent consinted of a series >{ tests in which the ir.let part icul jtc
concentration (Li), approach velocity (V) , and m -dia rate (ft) wore varied independently.
Selection of independent variable combinations to be tested within a given subexperi-
mcnt were hosed on a Latin Square cxpf.-r incntal design.
FACILITY DCSCPIPTIOK
The system constructed for this test proaram ia trhcrvit ical ly illustrated in
Figure 2. A positive-displacement blower capable of continuous operation at about
Tt". totin tost proarari consisted >'. nine subexperincnts representing Iu4 test points.
Conplctc res ilts of _ach test arc av.-ilable in reference 121.
569
-------
r-
-------
5 psig and 4000 scftn supplied air to the GBF and auxiliary systems including dust
injection air, media circulation air and fluid bed air but excluding instrusent air.
System pressure was controlled by an automatically-positioned bypass valve. The-
fraction of air which flows through the main duct to the GBF was controlled by a
pneumatically-actuated damper. An integrating pitot-static flow sensing element
located in the main flow branch provided both flow indication and input to the main air
flow control system.
One of the principal functions of the auxiliary air system was to provide air to
the dust injection station. The purpose of the dual-venturi dust injection system
was to provide a means of feeding dust into the main air line (which was u^der positive
pressure during operation), deagglomerate the dust from its packed condition and remove
large foreign or agglomerated particles from the feed.
This was accomplished by setting the flow through the dust injection system to
about 350 scfm by means of a manual valve and flow sensing element (which also pro-
vided input to the main air valve control loop). This flow passed through a venturi
which caused a subatmospheric throat pressure (about - 70IK at 5 psig blower pressure).
A hopper threaded to the throat of this venturi allowed dust to be introduced into the
air stream with a variable-speed vibratory feeder.4 The dust so introduced was
subjected to very high velocity (about Mach 0.6) which deagglomerated most of the
compacted caterial.
The dust-air mixture flowed from the feed venturi into a cyclone containing
three baffles in the annulus. This cyclone served both to complete the deagqlom-
erating process and separate any remaining large particles. The underflow from the
cyclone was captured in a sealed container at the bottom of the conical section which
was emptied periodically. The venturi in the main air line produce j a depression to
help overcome losses in the feed venturi and the cyclone as well as provided a high-
velocity region in which to disperse the dust-air mixture prior to deceleration into
the main duct.
The remainder of the auxiliary air was used in the media transport and cleaning
system. Both injector and transport air flow to the "L-Valve" below the G3F were
automatically and independently controlled. Calibration of this system for a given
media size and density allowed continuous monitoring of media flow. A manually-
controlled air stream to the bottom of the fluid bed provided final media cleaning
and uniform distribution of the clean media into the GBF. These air flows and the
entrained particulate removed from the media were then routed to a conventional bag
filter for final collection and disposal.
Dual sample stations were positioned at the inlet and outlet cf the GBF ir
accordance with EPA recommendations. An opacity indicator on the outlet duct aidud in
identifying steady-state operation.
An automatic data acquisition system scanned 40 instrument signals (seven of
which are shown in Figure 2) every minute. These signals were processed by a
Tektronix 4051 computer, converted to engineering units and recorded on macnetic tape.
In addition, a cathode ray tube display of certain parameters (e.g., GBF pressure drop,
opacity, superficial velocity) was. updated every minute to assist in the operation of
the unit. Plots of the parameters with respect to elapsed time (similar to those
shown in Figure 3) could be displayed upon demand by the operator.
4
A mixture of two grades of hydrated alumina (Al203'3H20) was used to achieve the
desired particle size distribution.
571
-------
TEST: CI6S5 D*TE: 8 3i ?r
OUTLET OPACITY ';>
TIME: 1526
13.0
12.0
II. 0
10.0
!>.0
6.0
7.0
ฃ.0
4.0
3.0
0
A
Vv
f
V\JJ
*|
1 ซ
\A
\
20 40 eo
1,
ft.
^Wl
K
Mr
1
^ 4
if1 \f
/v
ซ
I .,
|V
ri
1
60 100 lid 140 KO
ET -nir,.
1
i 1 ,
"VulN
It
i
t
*~
1 0 200 2.
FILTER PFESiUPE C'RijP -1H
a. a
?.o
e.o
4.0
3.0
2.0
1.6
0.6
\n
"V
"^^
,
s~Hi
|
V
0 20 48 eo ฃ0 100 120 140 ItO ISO 200 2<
ET -nit,.
MEM* CIRCULATION P~TE' Ib-'MIt,,.
200.0
180.0
tcO.O
140.0
120.0
100.0
80.0
60.0
40.0
20.0
0.0
h~
f^KfJ
**ปA^v-
^ป-^~
o 20 40 eo
^
%0 10'J
ET
EPA
IIILET
~liilET
EPA
1
HI
ป
*^*' 'ป rf
20 14
r. '
" OUItEI
^yvvk
*^^TV*H
r>.>v>-
0 IfO 160 <-
A\0!RSEII
OUILEI
"VI
o
1
ฃ
Figure 3. Typical data acquisition system record
572
-------
TEST MEASUREMENTS
The process variables recorded by the data acquisition system were all measured
by conventional techniques requiring minimum operator attention. Particulate loading
and size distribution were measured by the discrete sanple techniques listed below:
Particulate Concentation. - A conventional EPA Method 5 isokinecic filtration
sampling system was used to measure loading at the inlet and outlet of the
GBF. In the absence of condensible material in the sample, the impingers
were replaced with a desiccant cartridge. Three MIETO Model ~,2QO control
stations were available to withdraw samples. In accordance with EPA recom-
mendations, each sample consisted of two 90ฐ traverses.
Particulate Size Distribution. - Size distribution was measured with
Anuersen 2000 Hark III impaction classifiers. All samples were taken iso-
kinetically at the duct centerline.
DATA REDUCTION AND PRESENTATION TECHNIQUES
Two important aspects of the performance of a particulate control system are:
The relation of overall particulate capture to utility consumption, usually
expressed in terms of gas-side pressure loss ar.c
The efficiency attainable in controlling specific particle sizes (fractional
efficiency).
Since a relatively large number of variables was involved in these tests,
linear regression analysis was used to correlate the data. It was shown by analysis
of variance that a good and relatively simple correlation between collection efficiency
and pressure drop IP
A.
A0 [if "l ' 1
where Ao and AI are the "best" coefficients determined from linear regression. The
numerical value of these coefficients will be listed for each subexperiment.
Fractional efficiency performance for each subex?eriment configuration will be
illustrated by plotting the capture efficiency for five size ranges corresponding (in
terms of fractional efficiency) to the most favorable ara. least favorable combinations
observed during each subexperiment.5 since each subexperiment addressed roughly the
same variable combinations, the position of the envelope formed by these extreme
observations is a measure of the effectiveness of a particular configuration. The
variable combinations corresponding to these extreme observations are listed.
TEST RESULTS
Table I lists the results of the Oata correlation for the three subexperiments
previously described. Table II lists the best observed collection performance of
particulate associated with erosion/deposition.
5
This is a somewhat subjective choice since different combinations sonetines resulted
in the same efficiency. The points selected were, however, always representative of
the category of combinations associated with "most favorable" and "least favorable*
performance.
573
-------
TABLE I. DATA CORRELATION
SU3EXPERIMENT
FIGURE NUMBER
AVERAGE MEDIAN
DUST DIAMETER (pm)
(0
M EH
6. g A
O U
U
O CJ v
2 W 2 V
M J u
ป-3 rf O
03 O IK Li
U > b
(u
2 J m
O f* < "
M CO 2
MEM
ง20 AP
U.
gซt; v
M _3 2
M m cj
|J < M
o 2 o
CO O M L;
O H >4 H
M CO <
H ซ 3
M U O
a a M
2 H AP
O 2 0
Thick Bed
4
2.6
0.00303
-0.7
151
0.92
0.74
23
80
0.18
1.20
2
Nomiaaj Bed
^
3.2
0.0170
-0.5
120
1.23
0.62
15
59
0.25
1.08
4
Small Media
6
7.0*
0.00589
-0.6
157
0.40
0.90
34
41
0.37
5.01
9
* Correlation was not materially influenced by deletion of four data points
resulting in an average median diameter of 2.Sum.
574
-------
o
K
<
cr
UJ
?
^^
>-
ซt
z
o
_
-
o
t-
u
-I
o
o
0.995 g!
o
575
-------
10
-1
UJ
z
ee.
UJ
g
10
-2
10-3
o
10-
o
O
1C
-1
O
0.9
ฃ>
>-
0.95 ^
z
o
0.99 S
_i
_i
0.995 2j
>
o
I
lO1'
Figure 5(a). Overall particb collection performance, nominal bed tubexperiment
>
u
z
u.
u.
UJ
_i
<
o
ซx
ce
l.O
B.9
0.8
0.7
5 10 15
AERODYNAMIC PARTICLE DIAMETER (>ปM)
Figure 5(b). Fractional particle collection performance, nominal bed subexperiment
576
-------
r
z
o
10-
>
o
10
-2
O
I
I
0.90
-10.95
0.99
0.995
10-3 ID'2 ฃp Li 10'1
TW
Figura 6(a). Overall particle collection performance, small media lubexperiment
LU
z
o
o
o
u.
UJ
tu
z
LJ
Q.
1.0
0.9
0.8
i i i
I
5 10 15
AERODYNAMIC PARTICLE DIAMETER (PM)
Figure 6(b). Fractional particle collection performance, small media lubexperiment
577
-------
TABLE II. CONDITIONS RESULTING IN MAXIMUM
REMOVAL EFFICIENCY OF 2-5lim PARTIO'LATE
Subexper iir.cn t
Small
Media
Thick
Bed
Nominal
Bed
Velocity
(ft/min)
41
95
41
157
121
82
61
74
61
Inlet
Loading
(gr/sdcf)
1.2
0.7
0.3
o.y
0.4
0.4
1.6
2.3
1.6
Media
Rate
()b/lb)
0.6
2.4
0.9
9.7
0.7
0.3
0.6
0.9
0.7
Median Dia.
of Distri-
bution (urn)
1.8
2.0
1.5
3.3
2.7
1.8
1.6
2.9
4.0
Reriova 1
Efficiency,
2-5vm
0.999
0.991
0.990
0.995
0.994
0.993
0.986
0.985
0.984
Outlet
Loading
2jjm +
(PPMW)
6
7
5
4
3
1
39
40
16
Pressure
Drop
(IKd)
15.9
23.5
9.1
23.1
15.5
14.8
17.4
4.3
11.3
DISCUSSION OF RESULTS
Figure 7 shows the penetration/pressure drop correlation for both the thick bed
and nominal bed subexperimcnts plotted on the same axes. The relative position of the
two curves is explained by noting that the product (VM) is, for a given superficial
flow area and gas density, only a function of the media circulation (in, for example,
Ib/hr) and not gas velocity. Therefore, for a fixed inlet dust load and media circu-
lation, a given pressure drop w:'.ll result in higher removal efficiency for the t.iicker
bed. This behavior was anticipated in view of the fact that the t/dm ratio for the
thicker bed was greater (i.e., a greater number of collector sites present in the
principal flow path). The thinner bed had a more favorable t/11 (smaller values pre-
sumably resulting in less bypass flow around the top of the bed) but this effect was
overfiiidowed by the influence of the increased collector sites. It is also observed
that the negative slope of the thick bed data is greater than the thinner bed data so,
at larger values of |"Ap Li "1 (i.e., higher allowable pressure drops), the relative
LV TTJ
advantage of the thicker bad increases.
The importance of t/dm compared with t/H can be further illustrated by con-
sidering the small media subexperiment. Figure 8 repeats the thick bed and small
media data points. However, the media rate measured in the small media subexperiment
was multiplied by 1.3 to produce equivaler1 volumetric- media rates.6 It is noted that
the data thus transformed is very nearly coincident. Although the t/H ratio differs
by a factor of 2, the t/dm ratio differs only by about 13*.. As in the comparison of
the thick and nominal beds, the influence of t/H appears marginal with respect tot/dn,.
The ratio of the bulk density of the alumina media used in the thick bed subexperiment
to the bulk density of the silica media used in the small media subexperiment is
about 1.3.
578
-------
z
U.I
Q_
>
O
10
-1
10
-2
*^
0.90 o
tu
o
0.95 ฃ
u.
UJ
z
o
I
o
0.99
0.995
o
-J
o
10
-3
10-2
ft Li
T7T
10-1
Figure 7. Influence of configuration of GBF performance
o
I
<
z
UJ
Q.
>
o
10
,-1
-2
10
10-3
O 15.3" BED, 1.9 MM MBDIA
Q 7.6" BED, 0.8 MM MEDIA
I
I
ID
"2
.
LP M
TT
Figure 8. Influence of media size on GBF performance
0.9
0.95 t
0.99 3
0.995
o
579
-------
The fractional, efficiency curves show that very high (90*,+) efficiency is
attainable in the submicron size range for all configurations. These conditions tend
to be identified with high velocities, high inlet loadings and low media rates.7 It
is though- that tho annular configuration improves small particulate capture by pro-
viding lower velocities (and hence favorable diffusion collection) near the outer
radius of the bed, even when high velocities (associated with good impaction collect-
ion) are present near the inner radius of the bed.
The less favorable conditions (which are strongly identified with high media
rates and low inlet dust loadings) also have a detrimental effect on large particulate
capture. This apparently results from reentrainment of previously collected material.
It is observed that the "best" performance of the thinner bed still resulted in a small
amount of large particulate escaping the bed whereas the beds with greater t/dm were
capable of capturing essentially all particulate Oibo~ ? 3um.
Table II addresses collection performance on particulate above 2pm which is
normally identified with erosion and deposition problems in energy recovery equipment.
It is again observed that performance is superior (and, on the average, nearly
identical) in tho larger t/dm subexperiments.
CONCLUSIONS
The moving bed granular filter was found to be consistently capable of particu-
late removal efficiencies in excess of 981 for dust loadings (0.2 to 2.0 grains/sdcf)
and size distributions (1 to lOiim median) associated with many combustion operations.
Submicron coJlcctiori above 90?. was associated with high inlet velocities, high inlet
loadings and low media rates. The beds with larger t/dm ratios were most effective in
capture and retention of large particulate.
ACKNOWLEDGEMENT
This work was funded by the U.S. Energy Research and Development Administration
(Department of Energy) under Contract No. EF-77-C-01-2579.
NOMENCLATURE
"m
II
Li
M
t
V
AP
n
REFERENCES
Representative- diameter of media (mm)
Height of perforated inlet screen normal to gas flow (inches)
Inlet dust concentration (Grains/std dry cu ft)
Media rate (Ib media/lb gas)
Radial thickness of granular bed measured parallel to gas flow (inches)
Superficial gas velocity at inlet screen (ft/min)
Gas pressure drop across granular bed (inches of water)
Particulate collection efficiency (dimensionless)
1. J.H. P^rry, Chemical Engineers Handbook, 4th Edition; McGraw-Hill (New York), 19bj.
2. Cold Flow Test Program, Data Analysis and Observations, Special Task Report
FE-2579-15 prepared for the U.S. Department of Energy by Combustion. Power Company
(Menlo Park, California) under Contract No. EF-77-C-01-2579.
7
Loading and media rate are of greater importance in the smaller t/dm case.
580
-------
QUESTIONS/RESPONSES/COMMENTS
DALE KFAIRNS, CHAIRMAN: Do we have any questions? Would you
uce the microphone, please?
SPFAKFR (from the floor): Could you give us sor^e nurbers for
face velocities, pressure drops, grain loadings, those kinds of
numbers, please?
PR. r.lJILLORY. All right. Before we started the test, we picked
a common group of parameters to keep throughout all the configuration
changes. The face velocities we were looking at ranged from 40 feet
per minute to 160 feet per minute. This represented a maximum flew
rate of about 3,000 CFM.
The gain loadings: the lowest was around 1/10 grain per stan-
dard dry cubic foot. The maximum, if I recall correctly, was about
2-1/4 grains per standard dry cubic foot. Realize we are running at
relatively low temperatures, so using grains per standard dry cubic
foot really doesn't mean that much in this case. Presentation in
terms of "actual" cubic feet would have been just as meaningful.
The particle size distributions, the ones that we are showing
here, were around 2 to 3 micron median diameter. The pressure drops
that we observed, of course, being dependently variable, were all
over the map. Of course our better performance tended to be identi-
fied with the higher pressure drops, which I would say, depending on
the configuration, were anyv/here from about 15 or 20 inches of v/ater
in the case of the thick bed subexperiment up to as high as 35, in
the case of the smaller media; at the low end, two or three inches.
SPrAKFR (from the floor): Are these steady-state? :
DR. GUILLORY: Yes, this v/as. In fact, that was quite an opera-
tion in some of the cases, v/here we were dealing with, say, the thick
bed case. We had 16,000 pounds of media, and it took quite a while
to come to steady state. This is one of the reasons that we did use
an opacity meter. We used a Lear-Sigler opacity meter on the outlet,
and this was, of course, one of our important parameters that we were
continually watching to see when v/e finally did pull into steady
state, not only in terms of such things as pressure drop, but also in
terms of apparent outlet loading. We didn't use the opacity as an
absolute number, but it was an indication that we were no longer
changing. Did that take care of all the questions?
581
-------
INTRODUCTION
DALE KEAIRNS, CHAIRMAN: Our next paper is "Mathematical Model
of a Cross-Flow Moving Bed C-ranular Filter." It will be given by
Henry Wigton of Combustion Power Company. Dr. Wigton received his
Ph.n. from the University of Colorado in chemical engineering. He
also has a master's in chemical engineering from Oklahoma State
University, and a RS in chemical engineering from Texas Tech. Dr.
Wigton is chief scientist at the Combustion Power Company.
532
-------
Mathematical Model of a Cross-Flow
Moving Bed Granular Filter
H.F.Wigton
Combustion Power Company
ABSTRACT
As part of an ERDA-sponsored program, the theoretical performance of a granular
bed filter was modeled. Tho magnitudes of individual filter media grain collection
coefficients for the mechanisms of impaction, interception, diffusion, and sedimenta-
tion are estimated from published theory and data. Recent work by Dr. S. Goren is
used to revise these coefficients for clean media grains. The effects of captured
particulate on thec.a filter coefficients and frictional pressure drop are estimated.
Governing differential equations for gas and solids flow patterns, and of media-borne
and gas-borne particulate matter throughou' the filter are derived. Computer solutions
to these equations are used to correlate actual experimental data.
INTRODUCTION
The objectives of this task were to:
Understand and model the mechanisms by which particulatn matter (aerosols)
are removed from a qas by collector grains (spheres).
Model the gas and media flow patterns within the filter as a function of
operational parameters.
Relate the operational factors and the capture equations through coupling
equations which incorporate the effects of captured particulates on flow
patterns and collection rate equations.
This task ontailed:
Formulation of ijovernir:j equations.
Estimation of numerical coefficients.
Updating of numerical coefficients to incorporate the best values from
experimental data.
Discussion
The capture of aerosols by spheres is a subject which is treated by many
authors, such as llerne (3) and Paretsky (4) . Basically, a particle is captured by a
sphere if it contacts the- sphere. Surface forces will then hold the particle on the
sphere, since these forces arc relatively large for small (micron range) particles.
Neglecting electrostatic and thermophoric effects, individual spheres collect
aerosol particles by at least four mechanisms:
Inertial impaction.
Interception.
Diffusion.
Sedimentation.
583
-------
The capture efficiency of an individual sphere is conventionally assessed by
determining that fraction of gas approaching the sphere which is cleaned by particu-
lates. For example, referring to Figure 1, the volume flow rate of gas sweeping by
each sphere is equal to TI . 2 whereas the volume flow rate of gas which is cleaned is
, 4 dc "'
equal to uyc U, with yc being dependent on particle gas properties. For each particle
size (dp), an efficiency equal to 4y 2 may be calculated for either interception or
impaction by determining the limiting trajectories of each particle size.
Impaction
Collection by impaction occurs when the inertia of a particle causes it to de-
part the gas streamline and collide with or graze the sphere (Figure 2). The theore-
tical models indicate that a partijle flow parameter, called the Stokes number, is a
satisfactory correlating dimensionless group for these analyses. Numerical analysis of
the streamlines and particle trajectories which are calculated by Fuks (2), and
Herne (3) and confirmed by Paretsky (4) indicated that no impaction of particulates
would occur on isolated spheres below a critical Stokes number of 1.212. These calcu-
lations were based on the viscous flow pattern shown in Figure 1. Using the potential
flow model (Figure la), Herne (3) predicts better collection efficiency and a lower
critical Stokes number, because the gas streamlines are c- vded closer to the spheres.
This impaction parameter is defined by Paretsky to be:
Stk = -ฃง--* (1)
c
where: C = Cunningham slip factor expressed as:
21 t ~A3d- \
c = : + a^ (VA2 exp ฃ I
P
and A. = 1.257
A, = 0.400
A^ = 0.55
ฃ = mean free path of gas molecules ~0.065 micro meters at 25 C and ambient
pressure
This coefficient is important when the particle size approaches the same magni-
tude as the mean free path of gas molecules. Paretsky, using the free-flow model
proposed by Happol (5), showed mathematically that the confinement of gas flow near the
spherical surface to a volume which is equivalent to the void fraction of the packed
spheres** would decrease the critical impaction parameter to less than 0.1. The theo-
retical curves arc shown in Figure 3 for several values of packing void fraction. How-
ever, his experimental data taken on packed sand grains indicated:
There was no evidence of a critical value of this parameter even three orders
of magnitude below the theoretical critical value for ah isolated sphere.
(Collection was observed for values as low as St=0.001).
Collection efficiencies attributable to individual media grains were essen-
tially proportional to the Stokes number and inversely proportional to the
void fraction of the packed array at the values which he investigated.
The Stokes number as used by Paretsky, Herne, Goren (6), and Friedlander (7), is twice
the similar impaction number of inertial separation number used by Jackson (8) and
Perry's Handbook, respectively.
**In Happel's model, the volume of voids associated with each sphere is represented by
the volume between the two spheres having radii r^ and r , respectively (see Figure 2) .
584
-------
UC 0
Figure 1. Flow Line* Around a Sphere
(Taken from Fuks)
Figure 2. Free Surface Model Applied
to Inenial Impaction
(Taken from Paretsky)
585
-------
Jackson (5) found that the extrapolated values of Parctsky were still con-
servative (i.e., higher experimental values were obtained than could be pre-
dicted). Among the reasons for higher than theoretical behaviors are:
1) The existence of a wake downstream of a sphere at finite values of the
Reynolds number. Baird (9) has shown this to be an important factor
when falling raindrops were -ised as collectors.
2) The channeling effect of the preceding "row" of spheres which tends to
increase the number of streamlines which are initially directed toward
the center of the target sphere. This results in a greater net direct-
tional change of gas streamlines and increases the probability of a p:^' -
icle's impacting the sphere rather than following the streamline. The
effects of this important factor are best determined experimentally,
since the actual computations are too complex for rigorous determinations.
Impact ion Efficiency. - If the Stokes number is modified by dividing it by
the void fraction, the single particle collection efficiencies reported by Paretsky
are obtained directly over the range' of his data (see Figure 3) .
This expression is used as the value for clean media in the initial calculations.
Interception
Interception of a particle occurs when the particle, oven though following the
gas streamline, comes to within one half a particle diameter and grazes the sphere.
Particle inertia is not required for this mechanism of capture (Figure 2).
Interception Efficiency. - The theoretical interception efficiency has also
been treated by llerne and Paretsky. The estimated efficiency is a function of whether
the potential or viscous gas flow models (Figure 1) are used. For potential flow, a
vnlue of n. equal to approximately
3d
(3)
was suggested by Ranz and Wong (10) and also Jackson, whereas the value of
""--HdfJ ปป
is recommended by Paretsky (4) and Goren (6) when using Happcl's (5) viscous flow
model. The variation in collection efficiency is again due to how closely the stream-
lines approach the collector surface. The gas velocity near the surface of a sphere
is proportional ซ-.o the square of the distance from the surface in potential flow and
varies linearly with the distance in viscous flow. This capture mechanism plays a
more significaiit "ole when captured particles themselves can act as collectors.
Jackson (8) foL-Vd .-quation (3) to be a better indicator of actual data. This equation
will be used in "iht initial estimation of capture coefficients for individual spheres.
Diffusion
Brownian diffusion caused by random motions of small particles being bombarded
by gas molecules enhances the possibility of a particle's being collected. Even though
the particle may be generally following a gas streamline, random motion may occasion-
ally allow the particle to approach a collector surface and could cause it to contact
that surface and be captured. Kriedlander (11) recommends that collection efficiency
for a single isolated sphere by diffusion can be expressed as
Dlf .
586
-------
10
-3
1 1
THEORETICAL
"''
> = 0.43
_EXPERIMEtJTAL . = 0.49
//* ' 0.43
/' 20.30 MESH
* t - r\ A \
= 0.41
10-14 MESH
INERTiAL PARAMETER
Figure 3. Companion of the Theroetical
and Experimental Efficiency due to
Inertial Impaction
(Taken from Pareuky)
10
587
-------
where: Pe = Peclet number defined as =
d_U
d as
D = gas diffusivity = ~
16
Ic = Boltzraann's constant = 1.38 x 10
T = absolute temperature, Kelvin degrees
3nd^ g
C = Cunningham slip factor
Sedimentation
Particles, under the influence of gravity, will tend to settle from the gas
stream onto solid surfaces. This mechanism contributes primarily at low gas veloci-
ties.
Collection efficiencies attributable to individual spheres, as recommended by
Friedlander (11) can be defined as:
/U \ ฐ'77
nsed = ^lM(0i) C6)
where: U = terminal settling velocity of the particle calculated according to:
Ut
C 18M
U = superficial gas velocity
g = gravitational constant
The fraction of particles which bypass the sphere (1- n o) will be reduced by
each mechanism acting on the remainder according to:
(1- noTj - (1- nimp) (1- n inc) (1-ndif; (1-nsedi
Since these individual mechanisms arc small, the equation reduces to:
"o =n imp + n inc + n dif + r' sed
and the equation will be expressed as:
no = C1(St)n + C2 (^E) + C3(Pe)"2/3 + C4(Grv)3/4 (7)
As an initial approximation, values for clean media will be:
no = totil collection efficiency (fraction) of a single, clean, media sphere
C. = 1/r. n=l
The rate at which particles are captured will be calculated from equation (8) which is
readily derived from previously defined terms.
The efficiency of an individual media grain (sphere) is defin-d as that fraction
of gas approaching a sphere which is completely cleaned of particulate natter. The
area of the sphere normal to gas flow is nd 2 . If the area swept clean is denoted
4
as "c", the efficiency for a single sphere could be expressed as:
588
-------
4c _ g approaching - g leaving
ltd g approaching ""g "c
The number cf spheres in a differential volume of packed spheres will be
dV
and the total reduction in concentration in a differential volume would equal the
capture by a single sphere times the number of spheres per unit volume
substituting for "c"
results in the capture equation
dL
g t g d
which is the governing rate equation to be integrated along gas streamlines.
where: dV = volume of differential filter element (cm )
ซ = 1-e fraction of spatial volume occupied by solid spheres
Sc = cross-sectional area normal to gas flow (cm2)
dL = differential distance measured along a gas streamline (cm)
d^ = diameter of media collector (cm)
C = the concentration of particulate in the gas in gms per cubic centimeter
dCg = differential change in particulate concentration in the gas stream in
" gms/ci.i3
An improved approximation of the values of C^, ^ , Co, and C^ for clean static
packed aluminous spheres may be made by fitting the experimental data obtained by
Goren with curves as shown in Figure 4.
Reentrainment of particles can occur when media particles move. This movement
has been observed to be by individual media grains falling or rolling into a void and
decelerating within a finite distance, shedding a fraction of the particles which are
on the surface of the sphere. The rate at which particles are reentrained will be
proportional to the gas velocity, the solids concentration and a power function of the
solids velocity. The net removal of particles from the gas stream will be those
captured minus those reentrained, according to equations (3),(9), and (10) listed
below. Equation (11) couples/"t and Cs. Equation (11) may be integrated along any
gas streamline.
Capture
I ป -c
dC_
r=reentrainment coefficient (9)
.&)0-n*oi;
Net removal from gas (along a gas streamline)
U u nC ) 22_ (ID
589
-------
I
1 Eff.
.01
.005
.002
10 20
U. It/min
Figure 4. Gas Velocity in ft. per minute
100 200
590
-------
The three-dimensional global model, Figure 5, as simplified in Figure 6, may be
represented as a rectangular cross-section (ABCD) rotated around the vertical axis of
symmetry. Individual volumetric elements are represented by areas (such as 1,2,3,4).
Similarly if an isoparametric element such as 1234 in Figure 7 is used to
make material balances, equation (12)
- rU.U.
us
may be integrated along a solids streamline (between gas streamlines) to determine the
net change of particulate matter associated with the media. The use of isoparametric
elements simplifies the material balances since only one inlet and outlet component
for each phase need be considered with the particulate concentration in the gas,
specified at the gas inlet face and the concentration of particulate in the media
specified at the solids inlet boundry. Integration along the upper gas streamline can
proceed, yielding simultaneous values of Cs and Cg along the gas streamline (and at
each solid streamline where it is intersected by the gas streamline) . The complete
solution can be found in this inarching mode.
New values of gas velocity (streamline location) can then be calculated and the
process repeated until convergence is obtained. At each iteration, the problem can
be solved as a linear system of equations.
An alternate method of solving the gas flow equations as nonlinear sets is
discussed in detail under computer implementation.
OPERATIONAL PARAMETERS
Gas Flow. - Gas flow depends on geometry, boundary conditions, and governing
equations. While the average gas velocity through the filter is the primary variable
which is fixed by design requirements, the actual velocity at any point in the filter
is a complex function of geometry and other operating variables.
The end effects in the cold flow model apparatus are accentuated because the
ratio of outer to inner diameter is greater than one and the ratio of filter thick-
ness to filter height is significant. The actual gas velocity at any point in the
filter is required to calculate the prevailing capture coefficients and to establish
the gas streamlines which are needed to make integrated (path dependent) material
balances.
For normal operation at ambient or higher pressure and ambient and higher temp-
eratures, the gas may he considered as an incompressible, viscous, Netwonian fluid.
Derivation of equations representative of key variables in the model are as follows.
Flow of a viscous incompressible, Newtonian fluid through packed beds or
through porous materials will obey:
Darcy's Law VP = RU
Continuity Equation 7-U = 0
so that in Cartesian coordinates, the simplified equation:
"Pi) = ป
is the overall governing equation
where: 7P = pressure gradient
R = specific resistivity to flow
V-U = divergence of velocity
591
-------
MEDIA INLETS (4)
SOLIDS FLOW
Figure 5. Moving Bed GBF Geometry
592
-------
GLOBAL HODFL (ABCD)
VOLUMETRIC ELEMENT (1234)
Figures. Control Vc'ymes
593
-------
ซปz
X
/ I I I I I
till ' '
I / / ' '
fill
"S3
I t I
f I I
/ / / ,'~M
x
/ f
I I
I I
f I
I I
-M-
A'
1
*
1
1
1
1
1
1
f
1
f
1
1
f
1
1
f
1
"s8
-ปV
v
V
-~ v
"96
v
~~ "98
"slO
Figure 7. Solids Streซnlines vs. Gas Streamlines
594
-------
To quantify this relationship, this equation may be compared to Ergun's (1)
correlation of one-dimensional flow through parked media.
Ergun Correlation jp
^TT = (a+bU) U (14)
dL
2
where: _ 150 (1-c) y ....
a ~~ liJj
b = j^75 (1-r.) g (16)
3
ฃd c g
jj- = pressure drop per unit length
U = superificial velocity of the gas (based on an empty filter) in the
direction of the pressure drop in cm/sec.
By using both a viscous term "a" and a kinetic term "b", Ergun resolved many apparent
discrepancies in the literature as well as correlating his own extensive d,-ปca. The
viscous term dominates at lower Reynolds numbers and the kinetic term is mDro import-
ant at higher Reynolds numbers. (The granular filter operates at Reynolds numbers in
which "a" and "bU" are the same order of magnitude so tnat both terms ~^st be con-
sidered when determining flow distribution and pressure drop). When Ergun's correla-
tion is resolved in two-dimensional Cartesian coordinates, the following equations
result:
' ) U
ay ' y
where: U = component of velocity in the y direction
Uy = component cf v.-locity in the x direction
/2 2
|u| = total vector velocity \UX +Uy (cm/sec)
dc = diameter of media grains (cm)
i. = void fraction (fraction of the total volume not occupied by media)
I = sphericity of media grain, defined as equal to the ratio of surface area
of hypothetical perfect sphere of equal volume to the actual surface area
of the media grain ,
pg = density of the gas (gms/cm )
n - viscosity of gas in consistent units (i.e., gm/cm sec!
g - gravitational constant
Forchcimer's Law for steady-state flow as discussed by Irmay i'ij) rnciy be COM-
binc-d with the continuity equation and written as:
? / VP__\ = Q (17)
A triple identity will result if
R = a+bU = a'+b'q
so that for Cartesian coordinates
(18)
'"(^) - ฐ
may be evaluated by using the coefficients of the Ergun correlation
R = a+bU
When written for cylindrical coordinates, with no angular dependence (axisymetrical),
the governing equation for pressure distribution throughout the bed is:
_ ^ I + ^1 - '^-\ ,- o <19)
or / ' :>z \ R
595
-------
The stream function (ijr) which is the value of a streamline is defined according
to the equations
The equation for the stream functions may bo solved directly as
j,r.-^J+^r(-^l= 0 (20)
Inasmuch as R is a function of if/, this equation is nonlinear and must be solved
numerically, subject to the following boundry conditions:
ifi = 1 along top barrier (D'-D-C-C'J; See Figure 7
if" = 0 along bottom barrier (A'-A-B-B'j
UJ: = 0 at gas inlet (D'-A'J
|i = 0 at gas outlet (C'-B1)
a similar set of boundry conditions
P = 1 along inlet (D'-A')j See Figure 9
P = 0 along outlet (C'-B')
|| = 0 along barriers CA'-A), (b'-B), (C'-C) & (D'-D)
3P
yr = 0 along media inlet and outlet (A-B)& (D-C)
allow a solution for the normalized pressure distribution if desired. Either of these
equations may be solved numerically by relaxation or direct methods. Since the
equations arc nonlinear (R is dependent on P and 0) iterative methods are required
for precise solution. With no front face spillage, and smoothe walls movement of
media within the filter annulus has been observed to follow a "mass flow" pattern in
which the velocity of the solids is everywhere constant. Front face spillage for
clean media has been found to be proportional to the vertical velocity of media in
the filter and indirectly proportional to the inlet gas velocity.
This flow pattern can Be described by an equation similar to the governing gas
flow equation, with the solid stream function designated as "S"
3SX J_
Tt, Jz
with boundary conditions
S = 1 along outer cylinder radius C-C'-B'-B ; See Figure 7
S = 0 along inner radius upper boundary (D-D*)
S = Sfc constant - a function of geometry solids velocity and gas velocity
-long (A1-A)
S = a function of Z and Sb along gas inlet (D'-A1)
37 = 0 along both top (D-C) and bottom (A-B)
596
-------
A solution of this equation for Rs = 1 and Sb = 0.4 is shown as the vertically oriented
dotted lines in Figure 7. These solid streamlines, when superimposed on the gas
streamlines, form isoparametric volume elements such as represented by area 1234 in
Figure 7 and simplify the material balance equations which are discussed next.
MATERIAL BALANCES
The material balance r>qnปซ- -. ..is (Table I) state simply that at steady state (no
accumulation) the input to any volume within the filter must equal the output from
that volume. Equation (21) is applied to each individual particulate size classifi-
cation as well as the overall particrlate balance and states that any change in parti-
cle concentration in the gas stream ('~g) must appear as a proportional change in the
solids concentration (Cs). The particular form of this equation is advantageous in
implementing numerical computations.
Table I. Material Balance
I 'p^c^u^ * ^C.-0,.) dn = 0 dn = vector norm.
/surface 9 g g s s s
1
'
Vซ(P C U + pcco^o^ dv = 0 Divergence theorem true for any
volume
ggg
vo lume .
and 0=7ซu =7ซC p and p are constants.
L1 = gas velocity in cm/sec
Ug = solids velocity in,cm/sec
P^ = gas density, gm/cm ,
p" = solids density, gm/cm
C = concentration of particles in the gas, gm/gm
Cg = concentration of particles in the gas, gm/gm
(21)
COUPLING GAS FLOW AND CAPTURED PARTICULATE
The flow equation and capture efficiency equation can be solved individually
for the clean media case, but coupling relationships are required to account for the
effects of captured particulate material.
First, a correction is made on the resistivity of media to gas flow. The
presence of particulate matter alters both the effective void fraction of the packed
media spheres and the surface characteristics of the spheres. As can be observed from
the relationships in the Ergun equation coefficients, for the flow resistivity
R = (a+bU)
=, - 150 (1-c)2 ป
a ~~ O 1
(yd ) G Q
b = 1.75 (l-c)Pg
an agglomerate of micron-size particles will offer sevoral orders of magnitude more
resistance to flow than an array of media spheres so that the effective void fraction
between large collector spheres will be reduced by the bulk volume of small particles
deposited within the interstices.
597
-------
The net effective void fraction can be estimated by considering a unit volume
of filter bed
V. . , = 1 = V +r. when clean
total c o
V_ , = 1 = V +c+V
Total c a
The mass of collector particles in a unit volume is Mc = Vt x pbc = pbc- Tne rcass of
ash particles in the same volume is Ma = Va x ()Da. Tne mass or weight ratio of ash
to media
Ma Vba
s = ^ = ~^T
substituting JT x C or V (the volume of ash per unit volume) and 1-ฃ for V (the
volume of collector grains per unit volume. Equation (22) becomes
E ' ^ ฃ Cs
where: p. = bulk density of the media in grams per cubic centimeters
pP0 = bulk density of the ash in grams per cubic centimeters
Del
C = ash concentration in weight of ash per unit weight of media
o
cs = void fraction of clean media
The sphericity (
The mass ratio of particles to collectors (C ) equals the total mass of particles (N)
'
(!i!-ป)
V-6 P'
times the mass of each particle [ p p I divided by the mass of the single media
sphere under consideration ( c p 1. Substituting C c c for N results in:
V c/ s ^T-
p P
where: N = number or particles of diameter d on the collector surface
d = diameter of the ash particle p
pp= densitv of the ash ^article
PP= density of the media grain
Since the effective volume of a large spherical particle is essentially un-
changed by coating it with a monolayer of micron-size particles, only the surface area
changes will be considered in adjusting the sphericity according to:
$' _ effective sphericity _ original area
$ ~ original sphericity ~ effective area
598
-------
. w
Equation (23) could possibly yield a sphericity less than would be obtained with each
media sphere completely coated with particulate spheres. This complete coverage would
effectively double the surface area by replacing circles of area p with hemi-
TI 2 4
spheres of area = d . The mathematical equivalent to this physical limit is obtained
by using the value of $ or ^o whichever is greater as the actual sphericity of dirty
media calculations.
Figure 8 shows how the specific resistivity to flow varies throughout the filter.
The field above and to the right of the dotted line R = 1 remain" as essentially
clean media while the region below and to the left of the line R = 3.0 represents
the highest concentration of collected materials.
Usin-i values of C (ash concentration) specified at each grid point and impres-
sing boundary conditions, a grid of velocity (stream functions) and pressure vectors
can then be calculated using either relaxation methods or direct methods. Once the
velocity, pressures, and the stream functions are known, the flow net of streamlines
and isobars can be plotted as in Figure 9. Such a flew net is required to establish
the path of the gas so that a material balance of the gas can be made as it flows
through the filter. The material balance and rate equations can then be used to calcu-
late flow resistivity profiles as shown in Figure 8 by a method such as is outlined
in Figure 10.
COUPLING MATERIAL BALANCES WITH RATE EQUATIONS
As particulate matter is collected, the particle themselves now function as
small collectors! operating with the same basic mechanisms of collection as the larger
media collectors.
These particles are effective insofar as they protrude into the gas flow
streams. It is important to note that the theoretical models such as proposed by
Gorcn and Paretsky show that the increase in efficiency should be proportional to the
same factors as the increase in flow resistivity. Those relationships are also com-
patible with the frequently observed relationships between the friction factor (f) and
the mass transfer factor (j) in diffusional processes. The equation used to estimate
the coupling of capture efficiency with solids concentration will be
(24)
where: " = collection coefficient for an individual dirty media sphere
n = collection coefficient for an individual clean media sphere
a = viscous Ergun coefficient for clean media
bฐ= kinetic Ergun coefficient for clean media
U = loc-il superficial gas velocity
a = viscous Ergun coefficient corrected for changes in void fraction and
sphericity
b = kinetic Ergun coefficient corrected for changes in void fraction and
sphericity
CONCLUSIONS
The effective particle capture efficiencies of an array of packed spheres is
greater than can be predicted fi.m theories applicable to single spheres. This
synergism is futher enhanced at the Reynolds numbers prevailing in the normal opera-
tional ranges of commercial filters.
The effects of captured particulate material on the pressure drop and overall
filter efficiency are greater than car. be explained by the reduction of interstitial
voids. An additional correction of media grain sphericity as it is applied in the
Ergun flow correlation appears to predict these effects.
599
-------
FILTER r*\
INLET LV
GAS
STREAMLINES
!\ \
\
\ \
A\\\\\
-*\
ป\vป
\ \
\ \ \ \ \ \
\ v \\\"ฐ-
\ \ \ \ \
V \ \ ^
\\>
\x^
\ \ \
\ \ \ \
\ \
\ % \ \
\ t \ \
\ \ \ i
\ \
ปป-\\\\
N
\\
B'
B
Inlet
'3.0
routlet
Figure 8. Solids Loading Factor ซ. Gas Streamltnas
600
-------
MEDIA
P=3
P=10
INLET
SCREEN
GAS FLOW
OUTLET
SCREEN
Gas Flow Model
'out
Figures. Gas Flow Model
601
-------
MEDIA SOLIDS IfiUT
GAS INLET
PARTICIPATE
SOLIDS ~
CONCENTRATION
GAS
CONCENTRATION
PROFILE
MEDIA
CONCENTRATION
PROFILE
A GAS 1
| "1 STREAMLINES 1
I PRESSURE 1
| DROP ง
1 _ .._._
1 fc| SOLIDS 1
1 STREAMLINES |
1 "
1
i
MATERIAL ^ COLLECTION
BALAKCE MECHANISM
1
t
<
'ARTICULATE NEU SOLIDS CONCENTRATION PROFILE
MCENTRATION
GAS OUTLET
CONCENTRATION
ป
/ FILTER \
VEFFICIENCY/
Figure 10. Filter Efficiency Calculation Method
602
-------
Summary of Equations
Material Balance (Isoparametric Volume) :
3C
Vg 3L* dLg
3C
3IT dLs
s
Rate Equation:
Gas Flow:
dC
_i /I 1Z\ + _i /I IP
3r |R 3rJ 3z IR 3z
_i I ^ ill + -JL /^. III "
3r Ir 3r/ 3Z |r 3z/
Solids Flow:
.ป fe- ปd + i. /R- is
3r \r 3ry 3z IF" Sz
Capture:
Coupling:
C2(St)n + C3(Pe)~2/3 + C4(Grv)3/4
R = (a+bU)
fa+bV
Mba '1
ACKNOWLEDGEMENT
This work was done as part of ERDA Contract No. EF-77-C-01-2579. The consult-
ing services of Dr. S. Goren (University of California, Berkeley), Dr. S.K. Friedlander
(UCLA) and advice of Dr. M.L. Jackson are gratefully acknowledged. Special thanks are
extended to L.B. Wiaton (University of California, Berkeley) for his help in the field
of applied mathematics.
603
-------
NOMENCLATURE
A,, A-, A, = coefficients to evaluate Cunningham slip factor
A = total effective area of a sphere when partially covered with small
spheres (cm2)
a, a = coefficients to evaluate the viscous term of the Ergun flow resistivity
0 (gms SP.c/cm4) . (a0 applies specifically to clean media case).
Li, b = coefficients to evaluate the kinetic term of the Ergun flow resistivity
(gm sec^/cm5). (b applies specifically to clean media case).
c = fractional volumetric flow approaching a sphere which is cleaned by a
single sphere.
C = Cunningham slip factor (dimensionless).
C,, C_, C,, C. = coefficients to evaluate ... . ,., . (dimensionless)
1234 int, imp, dir, sea
C = concentration of particles in the gas, gm/gm
C' = concentration of particles in the solids, gm/gm
d = diameter of media grains (cm)
d = diameter of aerosol particle (cm)
D" = gas diffusivity (cm/sec2)
f = 3nd u
C
g = gravitational constant - 980(gms mass) (cm)
(gms force)(sec2)
Grv = gravitational number /U.\ (dimensionless)
k = Boltzmann's constant = 1.38 x 10 ergs/degree
I = mean free path of gas molecules (cm)
L = distance measured along a gas streamline (cm)
Lg = distance measured along a solid streamline (cm)
N = number of captured particles associated with collector sphere
N = number of collector spheres in a unit volume
P = pressure in gms force/cm
Pe = Peclet number dcU (dimensionless)
D
R = resistivity to gas flow = (a+bU)
RS = normalized resistivity to solids flow
S = solids stream function
Sc = cross-sectional area of a sphere normal to gas flow
St = Stokes number
T = temperature in absolute Celsius (Kelvin)
0 = superficial velocity of the gas (based on an empty filter) in the direction
of the pressure drop (cm/sec)
U = component of velocity in the y direction
Ujj = component of velocity in the x direction
|U|= total vector velocity ^U 2+U 2 (cm/sec)
U = gas velocity in cm/sec x ^
V9= volume (cm3)
V = volume of media grains (cm )
V = total spatial volume (cm^)
V = total volume occupied by collected particulates (cm )
" = fraction of volume occupied by media grain
c = fraction of volume not occupied by either media grains or collected
M particulate. c applies to clean media only.
/ = total capture efficiency of an individual sphere
/*? = particle capture efficiency of an individual sphere by interception
/jLp = particle capture efficiency of an individual sphere by impaction
/^ij = particle capture efficiency of an individual sphere by diffusion
/*sed = Part*cle capture efficiency of an individual sphere by sedimentation
/* = total capture efficiency of an individual sphere of a clean sphere
604
-------
Ut = terminal settling velocity of particle
P = bulk density of collected particulate gms/cm
ba
p = bulk density of collected media grains gms/cm
DC
p = absolute density of media grains gms/cm
p = absolute density of particles
p = gas density, c./cm
p = bulk solids density, gin/cm = ph
V = gas viscosity in gms mass/cm sec
Us = velocity of media solids
REFERENCES
1. S. Ergun, Chem. Eng. Prog., 48, No. 2, 1952.
2. N.A. Fuks, The Mechanics of Aerosols, CWL Special Publication 4-12, USDC
59-21069, 1955.
3. H. Herne, Aerodynamic Capture of Particles, edited by E.G. Richardson,
Pergamon Press, 1960.
4. L. Paretsky, L. Theodore, R. Pfcffer, A.M. Squires, J. Air Poll. Control Assoc.,
2_1, 204 (1971).
5. J. Happel, AIChE J., 4, 197, j'>58.
6. S.L. Goren, Consulting Report to CPC to be published as an addendum to Final
Report for Contract EF-77-C-Oi -2579.
7. S.K. Friedlander, Smoke, Dust and Hazo; Fundamentals of Aerosol Behavior,
Wiley-tnterscience, N.Y., 1977.
8. M.L. Jackson and R.G. Patterson, Shallow Multistage Fluidizcd Beds for Particle
Collection, paper presented at AIChE, 68th Annual Meeting, Los Angeles, CA,
November 1975.
9. K.V. Baird, Journal of the Atmospheric Sciences, 31, September 1974.
10. W.E. Ranz, j'.B. Wong, ln~d. Eng. Chom.. 4j4, 1371 (1952).
11. S.K. Friedlander, S.K. , personal commuTTtcation.
12. A.C. Pajfltrtkes. AIChE J., 2^, November 1977.
13. S. Irmay, T., Ancr. Gco. UnT, 39, No. 4, 1958.
605
-------
INTRODUCTION
DALE KEAIRNS, CHAIRMAN: The three papers that we will hear prior
to the break address the problems of high-temperature, high-pressure
particulate control as it would apply to pressurized fluidized bed
combustion concepts. The first paper this afternoon will analyze a
new particulate control technique; in the second paper we will hear
some experimental results on a high-temperature, high-pressure
granular bed filter operation; and then the third paper this morning
will provide some input into an overview in terms of the technical
status of granular bed filtration and the potential for that technology.
So I think we have an opportunity to increase our understanding of
these areas here this afternoon.
The first speaker is Dr. Ken Tsao. Dr. Tsao is a professor
in the Energetics Department at the University of Wisconsin in
Milwaukee. He obtained a Ph.D. in Mechanical Engineering at the
University of Wisconsin in Madison. He has an interesting background
and one that is applicable to the fluid bed combustion development in
that as part of his background he served as a power plant supervisor
with a petroleum corporation, among other activities. The title of
his paper is "Multiple Jet Particle Collection in a Cyclone by
Reheating Fluidized Bed Combustion Products." Dr. Tsao.
606
-------
Multiple Jet Particle Collection in a
Cyclone by Reheating Fluidized Bed
Combustion Products
Ken C. Tsao, Kuang T. Yung
and Jeffrey F. Bradley
The University of WisconsinMilwaukee
ABSTRACT Milwaukee, Wisconsin
A new particle collection technique is analyzed and presented for its potential
application in high temperature and high pressure gas cleaning systens. The new tech-
nique is based on the probability of particle collision and agglomeration phenomena by
reheating fluidized bed combustion products near coal-osh fusion temperature. The en-
trapped solids after impactions will agglomerate and adhere together to form into
larger sizes for effective separation linger centrifugal action. A mathematical model
is constructed leading to the design of an experimental high temperature cyclone. Ef-
fect of particle jet geometry and velocity distribution is discussed for the highest
rate of generation of new particles.
INTRODUCTION
Effective use of fluidized bed combustion products in a combined gas-steam tur-
bine cycle depends upon the ability of cleaning particulatc concentration in the high-
temperature, hiqh-pressure flue gas to an acceptable degree for the safe operation of
gas turbines. Presently, there are numerous research and development projects involving
cyclones, granular bed filters, molten salt scrubbers, and other hybrid processes such
as charged filters in modified electrostatic precipitators dซ2). However, some spec-
ific problems such as the effect of sticking on adherent particles in the efficiency of
clean-up apparatus result in making hot gas clean-up a major technical challenge. It
was proposed that a new approach^3' utilizing the self-agglomerating phenomena of car-
bon ash particle near its fusion temperature to a modified multi-inlet, multi-pass cy-
clone with insitu combuo'-.ion be investigated. Collection efficiency of submicron par-
ticles could be increased further in such an apparatus with additional collection mech-
anism of impaction of solid particles.
The combustion products from the fluidized bed boiler when passing a region of
high temperature zone in the cyclone such that coal/ash particles would enter momentar-
ily a partial moltun state. The particles will coagulate, agglomerate and stick to-
gether after impaction to form into a greater size but subsequently will be separated
out under centrifugal action.
The goal of this paper attempts: 1) to establish a mathematical model to simu-
iute the particle collision and agglomeration phenomena occurred in the proposed
multi-jet cyclone, and 2) to estimate the concentration of agglomerated new particles
by taking the effect of particle size and the jet velocity of incoming fluidized bed
combustion products.
FORMULATION OF COLLISION EO.UATIONS
The formulation of mathematical model is based on, firstly, the collision of an
elastic collision and then modified with a term called "probability factor" for inelas-
tic impaction. The particles are considered to be removed from the main dust stream
after the first collision process. Consider a particle with a diameter DO moving at a
velocity v into a group of particles of the same size, DO and concentration ng- The
number of elastic collision between a single incident particle and the group of par-
ticles, as shown in Figure 1, is
A = n0 r,OQ v (1)
where n0 is the particle number concentration per unit volume; v, the velocity of
moving particles; and o.^ = ^~ (Da+Db) , the total cross-section of hard spheres of
diameters, Da and DD- The subscripts, 00 or ab refer to two groups of particles of
a chosen reference size or particles with diameters D. and D-, respectively. To extend
607
-------
JL
nflCJ *
T
-4ฎ O
Tn ฐ o
ifi . a"
o
O *
O^o
o
x PARTICLE
DIAMETER. Dfa
PARTICLE
CONS., nb
COLLISION RATE
Figure 1. Single Particte Collision
608
-------
the single particle collision into a system of collisions between two groups of par-
ticles where each group contains no particles of uniform size DO, the number of colli-
sions per unit time becomes
oo
ฐoo no 3oo
where Qgo is the "probability factor" to account for the effect of inelastic colli-
sion. Since the particles after striking each other are considered to be removed from
the main dust stream, no secondary collision would occur. Thus, the number of colli-
sions is greatly reduced depending on the probably striking chance, QOQป among the
particles. Extension of Equation (2) to two groups of dissimilar particles of dia-
and D. , and their respective number density
meters
equation of collision is,
n and n^,
a general rate
(k + k.
a b
where kg and ku are the ratios of diameters of particles a and b with respect
to that of the reference size On. Equation (Z) reduces to Equation (2) for ka = kv>
= 1. And the new particle formed will have a diameter of (k-} + kฃ) */3 DQ-
RATE OF GENERATION OF NEW PARTICLES
Intuitively, the rate of generation OL new particles of greater size would de-
pend on many factors such as the collision rate, the thickness of molten layer enclo-
sing the coal-ash particle, the activation energy of molecules, etc. The rate of gen-
eration of new particles is further governed by perhaps, the operating parameters such
as temperature, velocity of dust ladden gas stream, and the incident angle, etc. in the
proposed multi-jet cyclone. It is postulated that the rate of generation of new par-
ticles after collision between particles a and b is,
G
ab
fab Aab
where fab is referrod to as an "adhesive ability factor," 0 <_ fat, <_ 1, and fajj = f^a-
In the case of a system composed of two groups of particles with number concentration
particles is
where Pa, Pb (
in each group; Figure 2, the rate of generation of new
,+k.
^ab
If
aa
2ฃab papb
.,
fbb pb
oo
* a a *b *b
are the percentages of the number of particles
i-nd
b with respect to the total number of particles in the gas stream. Generalization of
Equation (5) will lead to Equation (6) that for a system of groups containing particle
concentrations of nj, n2,..., nN with particle sizes of U\, D2,..-> DM and number
percentages of PI, P2, --- , PNป Figure 3, 2
EFFECT OF JET VELOCITY PROFILE ON A NEW PARTICLE GENERATION
One of the most controllable operating parameters in a multi-jet cyclone with
reheating is the dust ladden gas stream velocity. The offeet of jet exit geometry on
the new particle generation is of particular interest to the experimental design, set-
up and testing of the proposed new gas clean-up technique. Two ca..cs of jet velocity
profile (4) were incorporated for a case study. The profiles are:
(1) for a plane jet.
"n x
-S =2.48 (ฃ +
50
0.6)
"1/2 "0 1/2
{ )
^- = exp 1-75
(7)
609
-------
DO, "a
Db.Hb
V
D0fn0
*.S
WHERE Dj kjDซ
Gob
Pi I
Figure 2. Rate of Generation of New Particle*
Figure 3. Rate ol Generation of New Particles with Sin Variation
-------
(2) and for a round jet
J!= 6.3 (ฃซ) &)
5o
= exp [-96 (ฃ) 1 (3)
ปm
where UQ, u_, u are the mean gas exit velocity, the mean axial velocity, and the
local jet velocity; r-Q, aป ^nc densities of dust lodden gas stream at nozzle exit,
of the entrained air and of the cor,l-ash particles: h, do. tlr nozzle height of a
plane jet and the diameter of a round jet, respectively.
In cooperation of the jet velocity profile and the incoming coal-ash particle
mass concentration, the number of particles at a given section along the jet axis is
calculated. Hence, the rate of collision at any location of (x, y) in a plane jet is,
2
357.12 G-C . 3/2 2
A00(x,y) = ( - - !-2) (-ฐ) < + 0.6) exp 1-148 (^) |Q (9)
fur a round jot,
:D0
A00(x,r) = < ~> (-) (-) exp i-211 <) 1 Q (10)
Substituting Equations (9) and (10) into Equation (6), the rate of generation of new
particles is obtained.
DISCUSSTO:; OF RESULTS
For a given mass density of coal ash particles in the gas stream. Figure 4 shows
the effect of particle size upon the collision rate. The smaller the size of the in-
cident particles, the greater will be the collision rate. This is of interest to the-
partical application of the proposed multi-jet cyclone, since its intended purpose is
to clean up the submicron particles.
Figure 5 shows the effect of particle size with constant adhesive ability fac-
tor on the rate of generation of new particles. There appears that the values of f
is not sensitive to the rate of generation when the particle size exceeds the reference
particle of 2.. . On the other hand, the rate of generation would differ by approxi-
mately one magnitude of order when the particle sizes are varied fvom 1;; to 2...
It can be viewed that the proposed hot gas cleaning process is favorable toward
the smaller particles. The effect of non-uniform adhesive ability factor is shown
on the same figure by the dashed line.
With a given sot of f values (fn = f^2 = ^22 ~,0.5), the effect of percent
of particles containing two different sizes of particles in each group is shown in
Figure 6. The trend is f>"i:?cnt that the greater the percentage of the smaller size
particle.-; contained in a gas stream, the greater wTIT~'be th'o generation rate of new
particles.
Figure 7 presents the plot of AQO with respect to a plane jet stream profile.
AOO decreases as the jet travels Lurthcr down its axis. The graph can be utilized to
calculate the number of new particles produced at any given section of jet stream.
This will be beneficial for selecting an optimum geometry regarding the effectiveness
cf the proposed cyclone design.
8
As a practical example, let us take the probability factor Ogo = 0.5 x 10
and a plane jet velocity profile to estimate the actual rate of generation of new
particles. The QOO vai_e was calculated and based on the probability of that among
7500 possible choices, there are 75 red balls for which any random draw of 75 balls
from the 7500 total containing 10 red balls in one draw is 0.5 x 10"8. Assume also
^11 = fl2 = f22 = ฐ-5ป PI = ?2 = 0.5 and kj = 2, kj = 1: we find from Figures 6 or 7,
that II t* 0.27 at y_ _ x = 5.0. Further using Figure 7, AQQ is found to be 5.2 x 10~2.
x ~ h
611
-------
0.1 0.ซ 1.0 L4 1.0
RATIO OF DIAMETERS. k,=
D.
Figure 4. Ratio of Collision Rate for Single Component
612
-------
4H
30
10
50
3.0
I.O
05
0.25
WHERE Ps - PERCENTAOE OF MASS
, CONCENTRATION
fs-ADHESIVE ABILITY
k's-PARTICLE SIZE RATIO
1.0 2.0 3.0
PARTICLE SIZE RATIO, k=-Sป
Figure S. H Value* Versus Particle Sin
613
-------
BO
as i.o 1.0
PARTICLE SIZE RATIO , k,
Figure 6. Effect of Particle Mass Percent on '1 Values
614
-------
.6'
.52
io3
D.-DIA. OF REFERENCE
PART1CLE.2/4
s-V/x)x,62lฐ
Figure 7. Dimemionlesi Rata of Collision. A0
15
615
-------
Hence the rate of generation of new oarticl.es is
_, 357.2 G0 Cjj p 3/2
G = (5.2 x 10 Z) [( 4 ฐ ฐ ) (-2) J 0QO
n Dj p2 ('a
For un = 33 m/s, Cn = 0.001 ltom/ft3, o = 454 lbm/ft3, Dn = 2 u , G becomes (i.78 x
14 ft 6
10 ) Qnn. Further take Qn. = 0.5 x 10 , then G = 0.89 x 10 particles/sec. In the
UU UU j i
above calculation, tho incoming total number of particles of 2 u is 1.86 x 10 and
that of 4 p is 2.32 x 1010.
ACKNOWLEDGEMENT
This research was sponsored by the U.S. Energy Research and Development Admin-
istration under Starter GrantsUniversity Projects in Coal Research. Use of the
computing facility through a grant from the Graduate School, The University of Wis-
consin at Milwaukee is sincerely acknowledged.
REFERENCES
1. EPA/ERDA Symposium on High Temperature/Pressure Particulate Control, September,
1977, Washington, D.C.
2. Wade, G.L., "Particulate Removal from Hot Combustion Gases,* Proc. of 4th Inter-
national Conference on FBC, Washington, D.C., 1975.
3. Tsao, K.C., "Particulate Removal at High Temperature/High Pros., re by Self-
Agylomeration Process in a Cyclone," Coal Research Starter Grant with ERDA,
September, 1977.
4. Field, M.A., et.nl., "Combustion of Pulverized Coal," The British Coul Utilization
Rescrach Association, 1967.
616
-------
QUESTIONS/RESPONSES/COMMENTS
MR. SMITH: Ken Smith, Exxon Research. I've got two questions.
The first one is: have you considered the relative velocity difference
between the different size particles?
DR. TSAO: At this moment, we are talking of one velocity para-
meter only, "V". Certainly, the model can be incorporated into two
different velocities, in essence the relative values between the
particles. Yes, it can be done.
\
SPEAKER (from the floor): But if the relative velocity is small,
that means that the particles won't collide.
DR. TSAO: Yes. I follow your question. If the two particles
are one in front of the other, and suppose they are moving at the
same velocity, certainly they won't have any chance of colliding at
all. In order to incorporate this condition, if you are trying to
see that, the particles, the greater the size, the greater the
surface drag; therefore, we do not expecc that the various sizes
will move at the same velocity at all. In e.>sence, I am saying that
the model, to assume one uniform velocity, perhaps may not be realis-
tic. However, this has to be verified in performance under the
experimental conditions.
SPEAKER (from tha floor): The smaller particle will have more
drag.
DR. TSAO: Yes, under the condition of equal momentum achieved
for different size of particles in the cyclone.
SPEAKER (from the floor): One other question is: are you
worried about refragmentation of the particles after you have agglom-
erated them, at the high velocities in the cooling section of the
cyclone?
DR. TSAO: There is such a possibility. I think I agree with
you. I wish to enter this factor, our "F" factor, if you recall.
The "F" factor could be varied between zero and one. In essence, the
worst case would be F equals zero: no colliding, no collision at
all, which would be the worst case. I hope I have answered you.
MR. KEAIRNS: Let's see. I think we have another question over
here, and could you repeat the question into the microphone so that
all could hear? I don't think there are microphones around. Yes,
Professor Beer.
617
-------
PROFESSOR BEER: I would like to ask a question about the
super-imposition of rotating flow in the cyclone and the temperature
gradient introduced as you have mentioned it. Now, it is known that
if one had a positive density gradient, radial density gradient, in
the rotating flow field, that this can cause a density-stratified
emission, because the low-density gas in the middle cannot get out
and therefore the flow is laminarized, thence the turbulence is less.
Conversely, if you are introducing the high-temperature region
outside, you get highly unstable situations. Now, do you believe
that either of these, that is, a laminarized core in the cyclone or a
highly turbulent situation with the high-temperature zone outside,
might help you in reaching your objectives?
DR. TSAO: If I may repeat the question, if I can repeat it
correctly
MR. KEAIRNS: I gave you a tough assignment to begin with.
DR. TSAO: I am here and happy to learn, in essence; and Professor
Beer raised the question about density stratification as well as the
distribution of the temperature across the cyclone; and also because
you do have a swirling effect, swirling velocity, in the cyclone,
therefore what would it be? Should we have a laminar flow in the
central core, or will we have turbulent flov: in the outer core, or how
the penetration of the particle is from one region to another, if this
is the question. If I may answer your question, based upon my specu-
lation: In the experimental cyclone which we propose, we do have
auxiliary jets on the sides We will hope introducing the auxiliary
jets around the peripheral section will help to stabilize the tempera-
ture distribution in the center core, which we intend to maintain the
uniform temperature. By introducing the additional peripheral jets,
hopefully the jets will have enough pressure gradient which would
penetrate into the center core of the cyclone. This is something wa
hope. We still have no experimental data. We have no verification;
but it's based upon my intelligent conjecture.
CHAIRMAN: Let's see. There was a question submitted for Dr.
Tsao. Perhaps he could comment on that.
DR. TSAO: Thank you. I have two questions that were submitted
by Mr. Henry Kwon from Dorr Oliver, Inc. The first question is: "At
the temperature range you have shown, some elements of ashes can be
softened, which may cause scale buildup on the wall. If so, your
scheme may not work as you hope. Did you look at the problem from
this direction?"
I was anticipating this question to be raised. I believe the
boiler manufacturer may be able to answer this question better than I
613
-------
can. However, based upon ny experience with the boilers, we do have
the wet-bottom, also called cyclone burners, and with that type of a
boiler you night be plugging the "ashtray," or should I say the
discharge duct of the wet-bottom with ash. But it didn't happen.
However, I cannot assure you in the case of this cyclone, whether it
will happen or will not happen. There is no experimental evidence
yet; so my answer to that question would be "yes or no." I hope that
it v/il 1 not happen.
The second question is: "At the turn-down rate of flue gas
flow, do you foresee a significant reduction in cracking efficiency?"
The question here depends upon how you control your temperature zone.
The heating and the residence time of the particulate, the size of
particles traveling level in the high-temperature zone will have
tremendous effect upon the cracking efficiency. Therefore, I would
inject, if the design of the cyclone itself can be met at the high
temperature zone, some type of path is available. It depends upon
your loading conditions. Therefore, under that condition, I would
hope the cracking efficiency will remain at its design point.
MR. KEAIRNS: Okay. Thank you very much, Dr. Tsao.
619
-------
Feasibility of Barrier Filtration
Using Ceramic Fibers
Michael A. Shackieton
Acurex Corporation
Mountain View. Calif.
ABSTRACT
Barrier filtration using ceramic fiber filters offers a promising solution to
the problem of controlling particles in the high-temperature, high-pressure environment.
Industrial experience has proven this technique is capable of high efficiency particle
control, including fine particles, in near ambient temperatures and pressures. Exam-
ining those particle removal mechanisms which apply to barrier filtration indicates
that adverse effects caused by increased gas viscosity at high temperatures can be
compensated for in the design of the filter medium and in the design of the filter
system. Ceramic fibers are available which have smaller diameters (3..m) than conven-
tional fibers used for filters (10 to 20 ..m) . Analysis indicates that using these fine
diameter fibers should make it possible to produce filter media having weights less
than or, at most, equal to conventional media.
This paper reports on work being performed under EPA Contract 68-02-2169 to dem-
onstrate the feasibility of high-temperature, high-pressure particle control by filtra-
tion. Tests at room ambient have shown the filtration capability of ceraraic fiber
beds. Tests at high temperature and pressure have demonstrated several ceramic media
configurations capable of withstanding in excess of SO,000 cleaning pulse cyclos. To
solve the high temperature gas cleaning problem for TDC application rapid development
of this technology is clearly needed.
FEASIBILITY OF BARRIER FILTRATION USING CERAMIC FIBERS
INTRODUCTION
Many advanced technology processes currently being developed require removing
particles from high-temperature and pressure gas streams. An objective of developing
these processes is to increase coal use by making it economically efficient and environ-
mentally safe. These processes, such as pressurized fluidizcd-bcd coal combustion,
involve expanding the high-temperature and pressure gases across a turbine to generate
power to produce electricity. Such applications require removing particulate flyash
from the gas streams before expansion across the turbine. Techniques to accomplish
the required particle control have not yet been demonstrated.
Under normal environmental conditions, barrier filtration is an effective method
of achieving the required '^vel of particle control. However, at high temperature
(815ฐC) and pressure (10 atm), barrier filtration and other conventional particle control
methods arc limited by materials capable of surviving in the environment and by effects
of changes in gas properties.
Under EPA Contract 68-02-216y, Acurex Corporation is investigating the suitabil-
ity of commercially-available ceramic fiber filters for high-temperature filtration.
This work is sponsored by the Particulate Technology Brunei: of the Industrial Environ-
mental Research Laboratory at Research Triangle Park, North Carolina.
Major goals of this program are to:
Design and build a filter media test facility capable of operating at 815'C
and 10 atm pressure
Test available ceramic fiber forms (woven cloths, felted mats) to determine
if any can survive mechanical displacements and accelerations likely to be
encountered in online cleaning of high-temperature filter applications
620
-------
Develop preliminary performance data for those configurations which
appear most promising for high-temperature filter applications
Make recommendations based on the experience and data collected
Barrier filtration with available ceramic fibers is likely to be a good tech-
nique for particle control at high temperature and pressure. To illustrate why this
is true, a short review and discussion of barrier filtration theory is helpful.
Figure 1 is taken from a report titled "rffects of Temperature and Pressure on
Particle Collection Mechanisms: Theoretical Review" by Seymour Calvert and Richard
Parker (EPA-600/7-77-002), January 1977.* This figure shows a calculated fractional
efficiency curve for a fiber bed. Minimum efficiency is indicated for a particle size
of about 0.5 vm. The dip in the curve occurs because of the interaction of the three
collection mechanisms which apply to barrier filtration. These mechanisms are direct
interception, diffusion, and inerti.il impaction. For particle size less than about 0.5
um, collection by diffusion is increased, improving the efficiency of the filter bed.
For particle size larger than about 0.5 -fm, collection by inertial impaction is improved,
increasing the collection efficiency of the filter bed. It should be remembered that
this curve applies only to initial performance of a clean fiber bed. That is, it does
not include the increased collection efficiency that results from the filtration of the
accumulating dust cake. Note also that the Ho. 3 curve indicates that the inertial
impaction parameter for high-temperature and pressure conditions .should show a small
decrease in performance. To understand the magnitude of this effect we can compare the
performance of standard filter media when tested with Dioctylphthalate smoke (D.O.P.) to
its performance when tested after a stabilized dust cake has been developed. A D.O.P.
smoke penetration test is a standard test to measure the efficiency of high performance
filters such as those used to filter "Clean Room" air or to collect biological contami-
nants. This test measures how efficiently a filter ro>nov3s a 0.3 um diamcte' D.O.P.
smoke particle. Woven or felt filter media of the type corr.r-.only used for industrial
filters will collect only 10 or 20 percent of 0.3 urn D.O.P. smoke. Yet, after develop-
ing a dust cake, these same filter media will collect submicrometer particulate at an
efficiency of greater than 90 percent. Thus, compared to the c-'.ianges in performance
which take place in a filter media during the conditioning process, the changes pre-
dicted as a result of high-temperature operation are small.
Available ceramic fibers offer unique advantages for filtration, since many of
these fibers have finer diameters than conventional filter fibers. Conventional fibers
are usually 10 or 20 yin in diameter, while ceramic fibers are available with average
diameters of only 3.0 ym.
Collection efficiency can be improved simply by making a filter bed thicker, thus
increasing the basis weight of the filter (its weight per unit area). However, to
achieve high collection efficiency in this way can lead to high operating pressure drops.
Collection efficiency can also be increased by reducing the fiber diameter, which can
result in decreased basis weight and filter bed thickness. The importance of fiber
diameter is illustrated in the following equations which describe the three primary par-
ticle collection mechanisms applicable to barrier filtration.
dp
Interception parameter K- = -r (1)
c.r- d 2b
Impaction parameter K = 9P ?* <2'
f * kT
Diffusion parameter K. = = . , (3)
a 3 ~ j cl u ci,-
g p g f
These equations describe the collection mechanisms, but are not collection efficiency
equations. However, when expressed as above, an increase in any of the mechanism pa-
rameters (K , K , Kd) will result in an increase in efficiency.
621
-------
100
90
80
U 60
iu
O
t 50
UJ
I 40
20
10
0
CONSTANT FACE VELOCITY
NO. CONDITIONS
1 20ฐC. 1 atm
2 1.100ฐC. 1 aim
3 I.IOO-C. 15 atm
0.1
0.5 1.0
PARTICLE DIAMETER urn
50
Figure 1. The effects of high temperature and pressure on the
collection efficiency of a fiber bed
622
-------
The interception parameter is not a function of temperature and pressure, but
it is a function of fiber diameter. Changing from a 20-vm fiber to a 3.0-^m fiber will
increase the interception parameter by a factor of 6.67 times.
The impaction parameter is a function of temperature and pressure, essentially
through changes in the gas viscosity (ug). For air, increasing temperature froTi 20 to
815ฐC increases viscosity by about 2.5 times. This reduces the impaction parameter by
a factor of 1/2.5 or 0.4. But, che change in fiber diameter from 20 urn to 3.0 ;.m in-
creases the impaction parameter by 6.67 times. The net effoct of the two changes is
to increase the impaction parameter by 2.7 times.
The diffusion parameter is a function of temperature and pressure through changes
in the ratio of (C'T/ug) . When operating at 815ฐC and 10 a>Ti pressure, this ratio tends
to remain unchanged or to increase slightly. But, the diffusion parameter is also a
function of fiber diameter and a change in fiber diameter from 20 um to 3.0 urn will
increase the diffusion parameter by 6.67 times.
From the above discussion it is evident that if we make a filter using 3.0-ym
diameter ceramic fiber (which is commercially ivailable), it is reasonable to expect
that even at high temperature and pressure this filter will have high collection effi-
ciency without excessive filter bed thicknesses or basis weights. Using the method
developed by Torgeson, .-t- is possible 1-3 calculate collection efficiency for a given
particle size and fiber t^ed parameters. This calculation was performed for a 0.5-..ni
diameter particle with a density of 1.5 g/cm3 (as measured at the Exxon Miniplant),
for gas temperature of 815ฐC, and pressure of 10 atm. A fiber bed composed of alumina
fibers with 3.0-um diameter and fiber density of 2.8 g/cn3 was assumed. Results of
this analysj : for two filtration velocities and two solidities (a - the volume frac-
tion of the Tiber bed which is solid) are plotted in Figure 2. This analysis indicates
that a 3.0-um diameter ceramic fiber filter bed with a basis weight of 500 to 600 g/m2
will collect submicrometer particulate with an initial (clean) efficiency of about
90 percent at high temperature and pressure. Recall that a typical industrial filter
media (20-um fibers) would collect such particles at only 20 percent efficiency for
the same basis weight. Another way to look at this is to note that to achieve col-
lection efficiency comparable to commercial industrial media will require a ceramic
fiber filter media with weights only one-tenth that of the commercial media. Another
interesting feature of the analysis is that efficiency decreases for increasing veloc-
ity. However, by adding fibers, the given efficiency can be maintained as velocity is
increased. The quantity of additional fiber required is relatively small, especially
if only 20-percent initial efficiency is adequate for a 0.5-um particle.
Most commercially-available ceramic fiber structures are produced for insulation
applications. Consequently, these materials are generally characterized by in open
fibrous structure. That is, they have low solidity, with perhaps only 2 percent of
the volume occupied by fibers (a = 0.02). A solidity of a - 0.10 is more typical of a
structure designed for filtration. Figure 3 shows the effect on fiber bed thickness
for changes in solidity and air-to-cloth ratio. For solidity typical of insulation
materials (i = 0.02), a fiber bed about 1 cm thick should achieve high initial collec-
tion efficiency of submic^oiieter particulate, while a more compressed media with
o 0.10 would achieve this efficiency with a bed thickness of only 2 ran. If effi-
ciency typical of industrial filters is adequate, very thin layers of the 3.0-^m
diameter fibers will suffice. Note also that filter media thickness is not a strong
function of air-to-cloth ratio, indicating that high filtration velocity should be
possible. Of course, higher filtration velocity will result in increased pressure drop,
but this may be acceptable in a PFBC application.
Room Ambient Filter Media Tests
Available ceramic fiber configurations can be classified into the following three
groups of materials:
Woven structures - cloth woven from long-filament yarns of ceramic fibers
Papers - Ceramic structures produced from short lengths of fibers, generally
held together with binders.
Felts - Structures produced to form mats of relatively long fibers. These
materials are known as blan.':ets in the insulation industry. They tend to be
less tightly packed than conventional felt materials.
623
-------
3.0* m DIA FIBERS
2.8 g/cm> FIBER DENSITY
05* m OIA PARTICLE
1.5 9/em'
815'C
10 ATM
90
oป
j
I 60
5
y
D 70
i
uj 60
50
40
20
0
(i FT/MIN) *\
2.54 cm/tec)J
(25 FT/MIN)
1 12.7 cm/Me
c
I I I I I I I
IB ox/yd'
100 20>) 300 400 500 GOO 700 SCO 900 1000 1100 1200
BASIS WEIGHT ~g/M>
Fijure 2. CtlcuUted ptrformanet 3.0 pm alumina fibtf tad
-------
1.4
ra
1.2
1.1
ro
O 09
I
8 08
w
2
O 0.7
z
g 06
a
ฃ O.S
Q
as
02
01
o
0.5 Km PARTICLE
81S*C
10 ATM
4 6 8 10
AIR-TO-CLOTH RATIO - CM/SEC
14
Fi8urป3. Fiber bad thidcrwa
625
-------
A larqo number of ceramic fiber filter r.edia candidates have been subjected to
a scries of filtration tests at room ambient conditions. These tests included some
examples of conventional filter r.edia for co~.parison. Included among the tests were:
Diootylphtalatc ssoke (D.O.P) penetration as a function of air flow velocity
LX?termination of maxir.um poro size !in micrometers)
o Measurement of j>ermeabil i ty
Flat-sheet dust loadini tests using A.C. Fine test dust. Over-all collection
efficiency and dust loading roquirca to develop 3.7 KFa {15 in l^O) pressure
drop are detcrnined froci this test which is operated at 10 cm/sec (20 ft/min)
air-to-cloth ratio.
Data collected fron these tests are summarised on Table I.
Penetration tests using D.O.P. smoke measure the ability of the clean fiber bod
to stop fine particles. The- D.O.P. smoke generator is adjusted to provide a noninal
particle size of 0.3 .,m diameter which is a "roost penetrati.-.g" particle Fizc because
of the minimal effect of diffusion and inertial impaction at this j.aiticle size. The
D.O.I-. tost results should correlate well to the results prc-dicted by analysis since
particle collection is provided only by the fibers and not hy the dust cake. Figure 4
provides a plot of the D.O.P. efficiency as a function of air flew velocity for all the
media tested. Ceraraic media el^ta arc plotted in solid lines and conventional rac-iia in
dotted lines. Numbers on the curves refer to *bose on Table I. Several interesting
observations cass be ir-adc eoncernir.q this data:
Several of the ceramic natorials, especially the ceramic papers and felts,
are capable of higher efficiency collection of fir.e particles than are media
normally used successively in commercial filter units.
Many of the woven ccraaic materials, had zero D.O.P. efficiency at low velocity
and hiqher D.O.P. efficiency at highc-r velocity.
This is contrary to what theory suggests and to tho Lc-hjvior normally seen
in tests of conventional filter mati-r ials. A likely explanation for this
performance is that it is caused by t.>c presence o: man-/ large pores in the
r.edia. {examination of the pore size 'Jata in Table I shows that the woven
ceramic materials as a group are charicterized by larger pore size than are
conventional filter materials. Thus, at low airflc/w velocity, most of the
flow passes through the larqo |>ores ar.u little filtration takes place. As
velocity is increased, flow through the large pores, becomes restricted ^nd
some of the flow is caused to pass through smaller pores where more filtra-
tion can take place.
The D.O.P. data also supports the theoretical analysis. Efficiency as J
function of basis weight for selected ceramic materials is plotted in Figure
5. The materials selected arc ceramic uapcrs and felts. These materials
provide a fiber bed sirailar to that for which the analysis summarized in
Figure 2 was based. Figure 5 shows that the nominally J -n fibers do indeed
provide higher collection efficiency on a woight-per-unit area basis than
conventional media produced with larger diameter fibers.
Maximum pore size data shows that many of the woven ceramic materials had pores
larger than those characteristic of commercial filter materials. Also, many of the
felt and paper materials had pore sizes similar to those of conventional filter
materials.
Permeability is measured as the flow per unit area at a co-rtant pressure drop.
Thar, a material with low permeability offers a high restriction to gas flow and one
with high permeability allows more gas to pcnetrste for a given pressure drop. Table
I shews that some ceramic materials arc available which have low permeability, while
othe;s have high perceability. Socc of the woven materials have low permeability and
large port size, while others have high permeability and larce pore size. Most of the
paper and felt materials have permeability similar to that of commonly used filter
materials.
626
-------
TAHLE I. SUMMARY ROOM AMBTEN'T TEST DATA
627
-------
TABLE I (ConcJudcd)
...,(,'.-.I.,* * Mt-ltt .ป '-(.4 "I . rป
I ,|*-i 'viin i,iMtfi> #/>ป!
r>-i <-it (, i ! t ' I'.t.iii
.I. <-~it.,ป-.t. I.M KiU-i I r.i> I'-/ -I't.ft ' I/.41*.
li| i-ij-t fr. . l.l.^Utl l/'i Alt ' '".'. 1/4. ป4.
tปil '>:'*w.*v-;i lfjM.ni* I/'.
628
-------
100
90
80
70
60
| 60
n
d
g
.B
S 40
30
20
10
5 10
Airflow Velocity cm sec
Figure 4. D.Ot. effideney in air-flow velocity
629
-------
1200 -ป00
Figure 5. D.O.P. e ficiencv'n basis weight.
630
-------
Flat sheet dust loading tests were performed as follows: A 7.62 cm (3 inch)
diameter disc of media is suspended across an air stream which is maintained at 10.16
cm/sec (20 ft/min) velocity through the filter media. In this test the media supports
itself against the pressure drop (no screen is used) . Standard A.C. Fine test dust
(0-80 yn silica) was fed to the media at a nominal rate- of 0.883 g/m3 (0.025 g/ft?)
until a pressure drop of 3.73S KPa (15 in H20) is reached. Pressure drop as a function
of time is monitored during the test. This data is presented in Figures 6, 7, and 8
for selected materials. From the aata collected, dust loading (g/m2) necessary to
cause a given pressure drop 3.735 KPa (15 in f^O) is determined. Examination of this
data in Table I shows that so.v.e of the woven materials reached high pressure drops
while collecting only a small weight per unit area of dust. This is true also of the
commercial woven materials (items 31 and 32). Other woven ceramics were penetrated so
severely that they would not develop a oressure drop of 3.735 KPa (15 in HjO).
Two of the non-woven samples (which were unsupported) fractured as a result of
the pressure drop across them. Several of the ceramic paper and felt materials
exhibited dust loading, similar to that which is expected from conventional filter
papers and felts.
The flat sheet leading tests also provided overall collection efficiency (mass
basis) data for the tested materials. Dust penetrating the mcuia was collected in an
absolute filter downstream of the test media. Table I reveals that most of the woven
ceramic materials did not achieve high collection efficiency in this test. On the
other hand, woven commercial materials were only moderately efficient. Several of the
ceramic paper and felt materials, however, did provide collection efficiency of 99
percent or better. The two materials which fractured would have provided higher effi-
ciency performance had they not fractured. The test was stopped as soon as the fracture
was detected.
General Conclusions from Room Ambient Tests
Several of the ceramic paper and felt materials are capable of removing fine
particles at high efficiency without excessive filter basis weights.
The ceramic paper and felt materials have filtration characteristics and
performed similar to paper and felt commercial filter media in a scries of
filter media tests.
The ceramic woven materials in general were characterized by large pores and
poor collection efficiency in the dust loading tests. The range of parameters
exhibited by the various materials, however, indicate that an acceptable
woven ceramic filter media can proiably be fabricated, b.it such a filter
media would have the sarcc limitations as currently available woven filters.
That is, acceptable performance is only probable at low air-to-cloth ratios.
"Blanket" ceramic fiber materials (felts) consisting of small diameter fibers
(3.0 urn) appear to be the most promising materials for high temperature and
pressure tests because of their combination >*. good filtration performance
and relatively high strength.
High Temperature/Pressure Tests
Two major questions concerning the suitability ot ceramic fibers for filtration
need to be answered. These are:
(1) How durable are ceramic fiber structures when subjected to environmental
conditions associated with filtration applications.
(2) How well do ceramic fibers perform as filters in the HTHP environment.
Some preliminary answers are available concerning the first of these questions.
Three ceramic filter media configurations have survived a test during which the
filter elements were subjected to 50,000 cleaning pulses. The objective of the tests
was to simulate approximately one year of operation of mechanical loads on the media
at high temperature and pressure. Test conditions were as follows:
631
-------
oo
ro
10
20
n *
I
30 40
Time ~ Minutes
50
A.C. FineTซtOuปt
0883gM3
A C 10 16cm sec
ป33 Conventional Fe'l Filter
60
70
Figure 0. Dun loading of wramlo ftto.
-------
A.C. Fine Test Oust
0883g/M3
A/C 10.16 cm/see
Figure?. Dint loading of ceramic paper.
633
-------
U)
AC F mt Tm Dint
ORlUfi.'
A.CIOI6.TOป<-
31. 1? Ciw>vnliซn
-------
Temper a t'J re - 815 *C
Pressure - 930 KPa
Air-to-cloth-ratio - five to one (2.54 cm/sec)
Cleaning pulse pressure - 1100 KPa
Cleaning pulse interval - ~ 10 seconds
Cleaning pulse duration - 100 a second
Dust - recirculated fly ash
The three filter media configurations tested were:
Saffil Alumina Mat contained between an inside and an outside layer of 304
stainless steel knit wire screen. Figure 9 shows how easily the residual
dust cake was removed from this media after the test.
Woven Fiberfrax cloth with nichrome wire scrim insert. Figure 10 shows the
dust cake following the 50,000 pulse test.
Fiberfrax blanket contained between an inside and an outside cylinder of 304
stainless steel square oesh screen similar to common window screen. The
ceramic fiber blanket was held in position txitweon the screen with 302 SS
wire sewn between the screens. This resulted in <-
-------
Figure9. SaHil aluminป-pon ttft dust ata
(Cliarnd strip using vacuum cleaner)
636
-------
Figure 10. Woven (iberfrax-post test dun cakt
637
-------
v'^-ff^iL^fjr^Wj
^V'^^7ฃ55l
//^t-^iซgg5jig!tg
Fibปflf
633
-------
639
-------
High Temperature, High Pressure
Electrostatic Precipitation
Paul Feldman
John Bush
Myron Robinson
Research-Coitrell. Inc.
Bound Brook. New Jersey
ABSTRACT
This paper presents results of worr. conducted by "osearch-Cottrel 1 under FPA
Contract 68-02-2104. Tho purpose of the work completed to date was to demonstrate
tho ability to qenorate stable corona a*. temi>oratures to 2000ฐF and oressures to 500
psiq, thus establishing the feasibility of electrostatic precipitation as a means of
particulate removal from the effluent of fluidized hod combustors or coal nasifiers
at hiqh temperature and pressure. Tho work was quite successful in demor.strat inq
stable corona qeneration and in dcfininn ranges of temperature and pressure over
which the stable discharge can be maintained.
Oases investigated were air, flue nas, and simulated (noncombustible) fuol
las in coaxial wire-pipo electrodes. Pino diameter was ^ixed at 3 inches; wire
diameter varied from 0.062 to 0.125 inches. Results are reported for both tx>larities
in terms of curront-vol ta'te characteristics, corona onset and sparkovcr voltanes,
and critical siq. By
technical feasibility is meant the ability to qenerate stable corona over the ranao
of temperature and urossuro indicated. This was accomplished in a laboratory scale
tubular precipitator in no-flow, particle-free oocration. ^hiT method of oocration
was chosen because it allowed a we 11-control led, economical evaluation o' the
electrical characteristics of the system over the total ranoe of the variables. A
second uhase uroqram is needed to carry the work further into evaluation of narticulate
collection characteristics.
The primary variables studied were tonoeraturc, oressure,
-------
There is a fundamental problem encountered in desionino a precioitator for a
given hiah temperature/pressure service, "his is our incomplete knowle'lqe of i) the
ranqe of variables (pressure, temperature, electrode aoometry, qas composition.
polarity) over which a stable corona discharge can be maintained, and ii) the current-
voltace characteristics in that ranoe. In particular, there exists for the oosik.ive
discharae a critical pressure above which soarkover alone, without antecedent corona,
prevails. When the discharqe polarity is neoative, the critical ohenomenon is not
so precisely defined, and a postcritical discharae (often unstable) may be found at
pressures extending beyond the critical value.
Two opposinq effects are responsible for the phenomenon of the critical
pressure. First, shorter mean-free oaths at elevated oressures imoede ionization *w
collision and so tend to raise the sparkover level. Second, the doriser oackinq o?
aas molecules renders photoionization more likely and reduces ion diffusion. Thus,
pressure facilitates strcaxer proportion from the Anode across the qap and, at the
critical pressure, sparkover results.
The likely explanation of the relatively low value attained by the positive
sparkover voltaqe and its concomitant lower critical density is as follows:' Intense
ionization of the qas is produced in the hiah-field renion in the vicinity of the
discharge wire which attracts and removes the hiahly nobile electrons. The heavy
positive ions are repelled from the wire and move rlowly toward the collectinq
electrodes. However, on the far side o.r the i
-------
rXPF.RIMKNTAl, APPARATUS
Fi'iure 1 shows t.hc- configuration of the tost orccipitator used in this nrooran.
It is a wire-pipe d.?siqn (inclosed in a pressure vessel. ^ho nressure vessel was
desiqned for pressures to VJO psi'i and was assembled in three sections, each servina
a specific purpose. The top section contains the feedthrouah hushini for aonlyinq
hiqh voltaqo to the discharge electrode and a pressure relief line to protect from
over pressurization. The bottom section has a sitie access openinn for adiustments
and observ.it ions, a bottom support insulator to center the discharoe electrode, and
the Mas inlet. The center section of the vessel holds the urecioitator tube surrounded
by a three-zone heater used to reach the desired ouoratinq temperature. A layer of
Kaowool insulation separates the heater and the pressure vessel wall.
The collection tube electrode is a 7.26 cm internal diameter, Inconol 00ฐF to 2000ฐF for all oas
compositions, and, with air, ambient temperature data were also taken.
Pressure was varied in 50 psi intervals from ami ,ent to 500 osiq. In novin
-------
E?:t>
m
LU
Tor :::si;LA7jF BCSHIN
DISC1SASGE ELECTRODE
TUDE ELECTF.O^E
HEA7EP
ACCESS POPT
3O7TO.". SuPFORT
GAS INLET
Figure 1. Laboratory Precipitttor and Pressun Vessel for Test Program
643
-------
Table II. fias Composition
(Volume %)
Component
co2
He
ฐ2
N2
"2ฐ
Substitute Fuel fias
23.0
18.5
53.5
5.0
'Cfombustion f,as
9.2
2.8
83.0
5.0
Data were t^ken for three discharge electrode sizes as shown in Table I with
air. For the combustion gas and substitute fuel gas mixtures, only the 2.34 mm wire
was used.
The primary data taken were current-voltaqe curves for each of the experimental
conditions. The current-voltaqe curves were obtained on an X-Y recorder by recording
the curves for both increasing and decreasing voltaqe levels and repeating each
trace. Sparking voltage was determined as the final voltaae attained after being
held at sparking for two minutes. Corona starting voltages were determined by (1)
observation of voltaqe at which corona pips disapp, ircd on the oscilliscope with
decreasing applied voltage and (2) extrapolation or the current-voltage curves to
zero.
RESULTS
The raw experimental data, consisting of curves of linear current density
(mA/m) vs impressed voltage (kV) ire reproduced in Figures 2 throuah 5 for air,
Fioure 6 for simulated combustion gas and Fiqure 7 for substitute (i.e., noncom-
bustiblc) fuel gas. Corona-starting and sparkover voltaaes, derived from these
curves and independent measurements, are shown as functions of relative gas density,
<*, in Figures 8 to 10 for air. Figure 11 for combustion oas, and Fiaure 12 for
substitute fuel gas.
The first and most important objective is to examine th data for the puroosc
of establishing temperature or pressure limits to a stable corona discharoe. Such
limits may be caused by: i) excessive currents at low voltages resultina from
thermal ionization (where "excessive" and "low" are taken from the point of view of
practical prccipitator operation) and ii) the disappearance of (stable) corona due
to the manifestation of the critical pressure.
Examination of the data shows that catastrophic high-temperature currents are
not observed in this study under any conditions. This significant point is evident
over the full range of experimental pressures and temperatures and both polarities.
The relative gas density 6 is taken with respect to atmosoheric oressure and room
tenperature (294ฐK)
644
-------
Figure 2. Current-Voltage Curva Taken in Dry Air at Temperatures ol 294 K. 533 K.
end 950 K for 2344 mm Wire Electrode.
645
-------
Figure 3. Current-Voltage Curves Taken in Air at a Temperature of 811 K for
Wire Electrode! of 1.575mm. 2.344mm. and 3.17Sm.n.
646
-------
Figure 4. Current-Voltage Curves Taken in Air at Temperature of 1089 K for
Wire Electrodes of 1575mm. 2344mm. and 3.175mm.
647
-------
Figure 5. Currant-Voltage Curvm Taken in Air it a Temperature of 1386 K for Win
Electrodes of 1.575mm, 2344mm. and 3.175mm.
648
-------
o\
f*
VO
Figure 6. Current-Voltage Curves Taken in Simulated Combustion Gas Mixture for
Temperatures of 533 K, 811 Kt 1089 K, and 1366 K Using a 2.344mm Wire Electrode.
-------
at
vn
o
*
i.
c
r
' a
ปuttt<*l
1. JLU^>
11 ซU M M MM 'f
ซt !ซ.*
*
S%
Ij
a
-f r - .
uป'ปi
'b
i.
I
c
di
<
1.
j;
- , , ,
"f
\
Figur* 7. Current-VolUge Curvซi Taken With a Substitute Fuel Get Mixture at
TemperaturM of 533 K. 811 K. and 1366 K for ?.344mm Wira Eloctrodt.
-------
I
a i
Figure 8. Sparking and Corona Starting Voltage in Air at a Function oป
Relative Air Density at 294 K and 533 K.
B v'
i 3 JTrl.
I S ..
j ~r" g **
1
-, -.*
'! * 5 "*'
> o
j . .-; -
>/ I ';
rr_ f
['-. :::'....
Figure 9. Sparkwig and Corona Starting Voltages in Air as a Function of the Relative
Air Density at 811 K and 950 K for Wire Electrodes ot 3.175mm, 2344mm. and 1.575mm.
651
-------
H
1
13 :"
Hi*
'.ir
iv a*
--.-I *
J i
I f
*""
W. -
13
-- . .,_
, . . i ~ -
. ..
'
.
Figure 10. Sper*ing and Corona Starting Voltage* in Air a Function of Rctetitt
Air C*nuTf it 1089 K and 1366 K end for Win Ekjctrodnof 3.17Smm.2.3C4mm. and 1.575mm.
652
-------
'!,-::::'.':: I l,-:
7f:._'.-_.
^T '~"'-;-
- h :
q .
T ซl
5.:.:::...i *.'
Ftgufป11. Spsrkcng and Coron* Suning Vohagn fora Simulated Combustion Gaj Miปturซ ttป
Function of Rclnivt Gn Ocmtty ซ 533 K. 811 K. 1083 K. tod 1366 K fof 2.344mm Wtiป Elwtrode.
:,; ITI::':"
i
..j
i
]
L:.;.-:.-.:.-.:.::
d fTCl'" b
fZ,
iป. .
'I '.:" . "
3u.u:-:"ej -
3~" ..! 5?
'i?:::' f:-:
/T-il"' d !
:
1
,.."::...!
.
5-1
"^r:
Figure 12. Spwfcng ซ) Coran* Stirling Voluget in a Stdatitut* Fud Oซ Mixtur* M Ftmctiofl
of Rซl*tnป On Ownity M 633 K. 8J1 K. 1089 K. end 1366 K lor 2.344mm Wnป EtoetrwJfc
653
-------
; -:::;> :..- x:..!-:ซ >. .':.ป r ;.--..iw.iy T'jrr'T^s ^t. low volt i-if? ..re ~',s'. lik-lv to
oco-;r at tj.i- I'^v-'-'vT ':.!- '!e:isi* iซ-" II''V.-'-"ซ-r for t * < r'-'iu* --*- '"iror.a ซj* ' l*-ss ?ha*:
.il.'jj1 -inity, t :; r-Y-T-.e i:: :ซ.-rซ.-r.i 1 iy '..;. i.e.. .'.' :>r'-r.'.,irk'ivซ-r c'jr r'-nป:: .ir'.- ruch
lซ.-sr. h,i:. at. iii':!.--i :..:; i ป.!ซ;. !'..' i.'jsitiv t.ol.ir i .-/. r.ซ- low-' :.r.-r-.:,.ir i-.ov.-r
sifJ'-ri.'i'i the posit i ve ti i ::r7n.ir':--, the data show '.ha*, c'jr.b'i::'-ior
|e:;sซ-r extent , l',vซ ;*. d.-rv.ir ie:;. (-.'it , in .ir:y eve.-.t, -i :-r'.ซk,lep
as*:->c > .it ed with low 'ier.-. i t. jes d'>e;; r:f>t -irir.e.
l.-ant .it lower tซ-ri<-r.i'.'jr<-:., ซ>..it Mi-- or i t ir7,i 1-:,rer.-:'jr<- :ik!"p.o.-?er.oii woul'! set an
iif.i>'-r-:.ri-:;:;-ir<- lir>it t-j th<- i-'initiv- cur rent -/> 1 t.i-ป" t:urv-s an-i t h'- .T.;so':iซiป.o'l
vol t jซit tlej.;; i ty ป-.irv<-!: -- uiv,n '-or-v.-rt i r;'i density to :-:: !->r e.ich (xilarity, i! i r. clear li.it 'he r.e'T.it i vซ- -7rttic.il
iปri-s-:urt- .ilw.iyr. exrt-edr: t !iซ'- i.oritive. It is furt'i'.-r .iptj.irent t h.i'- he ne-iative
critic.il iire;;sui.-, like '!: |>o-.it.ive, i nere,-ir:<.-s with tennซ-r.'iป.!jre. ! r\ other virds,
'he tiiuln'-r the r--r.:--i at i;r<.-, the >ire.it.er the r.iti'ie ri! :,rซ.-:;r.urer. for r.t.ahle r:ซ.-M.it i v>*
corona.
Oinn.ir i son ol t t.e sn.irkover-vol t ,i'ie vs ;' ricr.jres H- 1 0 r-'Vals
a t"t,.|i-n':y t<:r tre s,-)-;itive r.i-.irkover volt.i'ir to exceed the :i."iative .it t emiier.jf ure-,
of ri ! I r and li inner and tor low air densities (. lc;;:; t.han 1 or 2). "}.ซ data .ire
not une'iui vi<:a I on tin:; ..< if in each c.isr , hut 'he trend sirens clear, u.irl icular 1 v
in-native than jiositivo ป:urrป-nt.n pn-vailinn at a nivi.-n voltane at ปhe hi'iher tt-mner.it'irci
are, howevei1, unrii :.t akahle.
r.omewhat hinder in-nalive than Kosit.ive current:; that m.'iv he ol>served at ปho
lower ten|ป-rat uro-; are, in part, to |,e at.t r I b.Jted to the s inn i f i cant rree-e Icot ro.i
comf*onซ-nt of t he cjrrent presertt ?-ir ri-lat i velv lonn mean ''r'-i'- uat hs (low : ) and
narruw in'erelect r<ปle sn.icinn.
AM.11 n, it rs-iy hi: nen- :allv (thounh not inv.iriahly) :;eer. rror\ Kiitur'"ป H-10
that, above an air d'-n:;ity of 1 or 2, the nena' iv soar ko-.-er volt.Ki*.- i : hinhiT * han
_he ixDGi'ivi.-. f-inre i noro.ir.ed density reduces tho mean f roe oath", and nobilities of
the charne carriers, enhanced oloct ron attachment and increase.! neuat.i ve-ion snace-
charne itensity ninht ho ex|x-cted to li.-ad to hi'iher nenativr r.narkover voltam-a.
That in, hinh pressure in combination with hinh temperature restores, in a r.enso,
tho low-tempi-raturo situation.
In tho c.tso of substitute fuel nas (Kinurt; 12) the iปositi\o suarkovor voltace
excoof1:! tho noaatjvo, over tho full temperature ran'io shown, 1111 to .1 density of 6 or
7. For combustion aas (Finurc 11) th-: transition occurs at about a density of 4 for
temperatures of, or nroator than, loa9 K.
As temperature and pressure arc increased toq<;thor, 'or all of tho oxoor ipen'.il
situations, it is clear from the data that precipitation is i:or.siblo at sinni f Scant ly
hinhcr voltanes than at normal conditions. This is a most important fact whon
assessing tho viability of electrostatic urecipitat ion 'or hiah tenuorature, hioh
pressure particulato removal applications, esoecially in comnarison to other collection
devices. Tho reason for this is that the rate of particle collection in a orocioitator
is rounhly proportional to the square of the electric field stronath in tho orecinita-
tor. The field strennth in turn increases with npolii.-d volta-:o. Tho not effect of
an increase in particle collection efficiency, or a decrease in :>r.--cioi t.'tv.: size.
Thus precipitation becomes noro efficient as temoorature and urossuro increase
together.
654
-------
Other particle- c >1 lect ion i!o>vicfs such -is filters of various tyuos, cyolor.es,
otc. -Jo not bonofi". fro->. increasing temperature and orossuro. In fact, -oerforr.anco
deteriorates in th^so -ovicos becauso of i r.crea.s ir.q iias viscosity ar.J dorreasi r.!' thซ.- na -or conclusions dt-rivcd fron this work:
1. Thoro arซ- r.o torir.cr.it uri.- or pressure linitations to cl>-ct ror.tat ic ure
ipitation cvor tho ran'ic stu-iiod.
2. Precipitation boccr^s noro ot'ficicnt with increasing tor.uซ.-rat ure ami
.ross'jre. This is in diri-cf contrast to the trend of other ocrticl-- collection
ir;V ic<:^.
1. Critical pressure increases with tonpcrature.
4. !.''"'iat ivซ- critical pressure is hi'ihor than nositive.
rj. ;.'cซi.ปli .< currents arc hit>i r.so.-., M., "I'.lectrontat ic I recipitat ion" in Air ''ol lulipji_Contฃol , K, Strauss,
-I., Voi. 1, V.'i ley-Interr>c-ii'nce, New York, lrป71, "in/." 227- US" ~
2. Cooperrvjn, I'., "Stxintaneous lonization of f,ar,es at Hioh Tonijoraturo, Tonforonco
tMper C1-17J. .--n-r. IMS'. i:lec. Knurs., 1*165
1. H row: >. f. r . .inii w.ilK^r, A. h., Teasifoi 1 i ty IV'nonat rat ion of Klectros-.atic
I'recipitat ion jt 1700ฐK" ) . Air Pollution Control Ansoc. ?1 , 617-620 (1971)
4. roopeman, I1., "Stjont.'inซrous lonJ7..ปtion of Oasos at Hiqh Temperature," Paper l"S-
MO:J-6, Ir.st.. I'.'octron. Kn.iru., l'ป71
r>. hnliinson. M. , 'Critical Pressure.-, of "he Positive Corona Between Concentric
Cylinders in Air,' .1. Aniil. !'hys. 40. 5107-5112 (1969)
6. H'jwoll. A. H. , "Breakdown Studies in Compressed r.asos," Trans. Am. Inst. Klec.
Knars.. S?, 153-204 (I9J9)
655
-------
Corrosion and Erosion
657
-------
INTRODUCTION
STAMT.Y DAPK'mAS, CHAIRMAN: Our next speaker will bo Anne Rowe
of NASA-Lewis. The title of her paper will he "Corrosion/Frosion of
Turhine Blade Materials in the Hiqh-Velocity Effluent of a Pres-
surized Fluidized foal Conhustor." Coauthors on this are Zellars and
Lowell, also of f.'A.SA.
659
-------
Erosion/Corrosion of Turbine Airfoil
Materials in the High-Velocity Effluent of
A Pressurized Fluidized Coal Combustor
Glenn R. Zellars. Anne P. Rowe. and
Carl E. Lowell
Lewis Research Center
National Aeronautics and Space Administration
Cleveland. Ohio
, ,; /_..-.-> ,.-,:' . . vi ...- f . . ,. /'
' ' - '.ป I.. I..* /.!.*! .. ... > . - .., , / .
in K'l'. ', tr." I'.r.t rr'. r^'>n !:: ! >!.. i !n :. '. :> !*<:-
ro it.cl.
^orpriri'^r^t.r;, rtr. "o^n In r'it*. P. K',*jrl*_- -j I:; :i .*.?}. '-.".Tt I ' -ir'-.w! :;ป :;h"iv;Inr vri" :'ซ%1';" \'_, LT,
o 03 en (9 In. to ^1 !n.) '.r. ^0? rn (6-1/2 ft). ":;? ti".i .lonJrinfrs rrojoclca *.h.-jt t.:.c
660
-------
Figurtl. Lewi* PFB Tot Facility.
661
-------
FUTURE
CARROUSEL CYCLONE
WEDGE AND
TESTUNIT-v FILTERH
TV ViCTj
CAMERA
^
SOLIDS
REMOVAL
SCREW
Oi
ro
FILTER
"r
Figure 2. High Bay Ten Area.
Figure 3. Schematic Drawing o< Corrlnntoi and Aumli.iry Coi
-------
.. .' ' -.- :..,- .:' :. >. : .:. : . -. -,.: I':.'-. ; /.;...;) :,- !;..- .-: of :
*:.:: ':.:.: :.-:-.. -:.-sr: ':-.::.:.: ป.-; :.::. r.;--".--:.: -i i .::; x-:--iy :! :':': , ':':. .:.<";.-.-1
:'. v -i i : .'.. ., i-:.-::-- ':-. K--;'--., ':.'! -::.or '.r^';:.'2 :,:' "
-.; -.:'.' \i.:." :' :;.; '.'.;: . '. r :.:'. i"i". :..:''.,'... -.\\',;-.:; -.-.'.-r'.>. L-.---,-:n--
_
663
-------
Figurt4. CVYT AftซTซt.
664
-------
COMBUSTOR EFFLUENT
o
OS
Qฃ
JLJ
Q_
1100
10
20 30 40
TIME, hr
TOTAL
AIR FLOW
Ib/hr kg/hr
650-1300
600 J 275
50 60
Figure 5. Representative Temperature and Flow Variation!
During the Second Week of the 150 m/iee 745 C Test.
TABLE!
COAL AND LIMESTONE ANALYSES
COAL
CONSTITUENT wt %
FIXED CARBON
VOLATILES (INCLUDING 1.86%
ASH
SILICA
ALUMINA
FERRIC OXIDE
LIME
PHOS PENTOXIDE
TITANIA
POTASSIUM OXIDE
SODIUM OXIDE
SULFUR TRIOXIDE
MAGNESIA
UNDETERMINED
SULFUR
JOULES
Btu
53.92
SULFUR) 38.07
3.74
2.04
1.37
.32
.03
.06
.10
.03
.14
.06
.14
1.86
3.157xl07/kg
13586/lb
LIMESTONE
CONSTITUENT wt %
SULFUR 0.02
SILICA 1.28
ALUMINA <. 10
CALCIUM CARBONATE 96.32
MAGNESIA .51
UNDETERMINED 1.77
665
-------
TABLE II
SUMMARY OF TEST CONDITIONS
TEST
NO.
I
II
III
IV
GAS
VELOCITY,
m/sec
150
270
150
270
SPECIMN
TEMP.
c
745
720
800
790
AVC
SOLIDS
LOADING,
g/scm
3.9
4.4
2.3
3.7
AVG GAS COMPOSITION
ฐ2-
%
7.7
7.6
11.7
6.6
co2.
%
9.5
10.6
7.5
11.1
CO.
ppm
20
14
9
4
NOX,
ppm
122
110
231
239
so2.
ppm
450
420
258
365
THC.
ppm
2
2
2
1
CS-77-2681
Fiซura6. Effluent Pvticta from PFB.
666
-------
TABLE III
MASTER HEAT ALLOY COMPOSITIONS
wt%
Ni
Co
Cr
At
Ti
Ta
Id
A/In
nfJ
jf
L\
ft
et
>l
c
B
S
Uf
MI
IN-100
HAL
14.90
9.30
5 ?S
4.90
? sn
C. ou
08
. Uo
a 21
on
.07
ff,
. U5
a i?
.014
.C92
U-700
BAL
15.50
14.20
4 on
3.25
A A(\
a 10
a 06
.016
.004
IN-792
BAL
9.20
1Z70
a cjt
j. Jt
3.90
4nc
. (fj
1 OA
1. VO
4 in
. 1U
nA
. UD
a is
a 10
.015
.005
on
. CU
MM-509
9.90
BAL
S. 4
a 28
i in
3. /U
A QC
O.'O
JIH
.*u>
a 32
a 59
.006
.007
CS-77-2475
667
-------
and MX-509 as an e/.a.r.ple of a Co-base vane alloy. These par-t Icular alloys were
selected because they have boon used ox tons I vely !n corrosion studies lr. our labora-
tories and thus offer possi bl 1 it ier of data correlation.
Figure 7 is a sketch of the wedge :;p.ec I men. The fc.-jse Is held in the carrousel
by a set screw, while the top end Is free. The narrow e*i ;ป; or. the left !c designated
the leading ed;;e; the curved surface opposite In the trail'm: od^e. L-x-.'/iInr. alori,", the
leading '--dee from the free end, the side In view is sailed the left face and the one
behind Is the rl;;ht face.
RECULT:; AJ;D Di"cu::r.io:!
Specimens exposed In the CWT to the high-velocity coal combustion products of
the PFii suffered damar.e fron both erosion and corros:lon, the extent depending prinar-
ily on the velocity of the n/'H stream, the solids loading In the ,~as, the specimen
ter.perature, and to a lesser extent the alloy properties.
As a result of the r.eometry of the test section and entry nor.zle, Inpait of t.he
f;as strenn on the rotating specimens produced unevenly distributed d.-:r.na>;e. I'r.!r.ary
r.aterlal removal occurred at the center section of ee.ch specimen, which was directly
above the noz/.le and thus received the maximum impact of the gas stream and the
erc-sive particles. As can be seen In Fl^. 8, a photograph of specimens fron the
second test run, the leading cdt;c shows the greatest material loss. :'.oth face:; are
eroded, and the trailing edge also loses some material although It is exposed to '.ho
direct path of the stream only a third as much time a:> the leading edr.e. Kvldence
tliat the direct partt'ilo paths aro diverted by the rotating specimens is seen in the
fact that the two faces experience slightly different damage patterns: the rljfht- face
suffers somewhat more damage than the left.
Stronf; dependence of erosive damage on geometrical factors err,phasl;:f.-. the
necessity of testing turbine materials for PKB applications in a turbine confli'uratioa.
At the end of each test the wedge specimens were removed fron the carrousel,
rinsed in cold water, wiped dry, wcir.hed, measured, ptioto;;raphed, and then aee tinned
for notallographlc examination. Loose du.st war, removed by this procedure but there
remained particles of KepOj, r;iOp, and CaoO/i firmly embedded in the specimen surfaces
after exposure at both velocities. I-Mgure 9 shew:: scanning; electron mlcroivypiis; (.':!::!)
and energy dispersive spectra (KD.S) of the elements present on the surfaces of two
typical eroded leading; edr.es. ."EM portrays raised areas, in this case the embedded
particles, as lighter than the background alloy. The KD.'; scans Indicate Ca and ;; from
CaSOlj, Fo from FejO^, and oi and Al from aluminum silicates and SlOp, aloni; with Ti,
Cr, Co and i.'l fron the alloy3. Oxyi;en docs not appeal- on the EDS scans because it is
out of ranr;e of this analysis technique.
Figure 10 Is a cross-sec' tonal view of an lil-100 specimen from the high-velocity
h'Eh-temperature run In which the particles, SiOj In this case, have deformed the usual
cubic Y-Y' morphology of this alloy, by the force of the Impact. The Au peak Identi-
fied In this and subsequent EDS seine arises from the coating applied to make the
specimens conductive for SEil analysis.
The leading edge to trailing edge distance, t, at the center of each specimen
was measured by micrometer before and after exposure. Since Initial examination of
the specimens had shown that erosion wis the primary mechanism of danage, the center
section At data were divided by the average solids loading for the run In order to
assist In evaluating temperature and velocity effects.
The resulting data are listed In Table IV, grouped by velocity to show the
relatively small influence of the temperature differences a;id the much larger Influ-
ence of the gas velocity differences. Duplicate alloy specimens gave results that
agreed within i 0.3 cm/yr : g/scm. Thus differences between alloys under the same
test conditions are In some cases below the level of significance. Minor differences
are really insignificant anyway since the smallest loss rate in the table represents
a loss of over 2 cm In 10,000 hrs, totally unacceptable for power turbine application.
This result was of course not unexpected: these tests were run in order to establish
a base for comparison with results after various levels of gas clean-up.
668
-------
1.27cm
ESTIMATED EFFECTIVE
ATTACK AREA-13.5 cm? \
ZONE OF HOT
GAS STREAM <
IMPINGEMENT
1.27cmDlAM
CS-77-2682
Figure?. Carrotoel Wedge Test (CWT) Specimen.
7.62cm
Figure 8. Test Specimens After Exposure.
669
C-77-2880
-------
-------
VfVfl
SPFCIViFf,
-------
Although differences between loss rates for different alloys and u.: f ferer.t
temperatures are snail, '.he v-;ocl:.y "ffe::t Is -jioar, as seen In .-'ฐ evidence
of reaction products. If i:;-]i)9 reacted, as would be expected from the previous test
results, both reaction product:: and depleted son-- have been eroded away.
In the third t-st, at. 1^0 ryr..,-; ;lnd 80" C for 91 hours, all four alloy:-, devel-
oped continuous layers of react Inri products and depl"t!on -/.ones, a:; seen In Fig. 1 '',
on both faces and trail Ing ..-dgos. These lay.-:; had been larger, oi-oded away f:-.-:m thv
leading edge:;, a portion of wnlch can be seen In the il-700 nir-rograph. Th^ 'i -p:ot ion
zone on :-!M-C09 '-'annot be seen !r; this 1'1,-ur--.' t.-ecause It Iri'/olves fine :;tr:i':ture wh !'-::
Is not resolved by light microscopy; It appears In .".KM mlcroj-raphs.
.'IKK examination of the reaction product arid depletion layers, exar-.p ::; of which
are seen In Fig. l.:i, revealed small particles dm-p In the depletion ::'.n-.-:! '>'.' all f;,:..
In some cases the scans suggest that the su! fides are I.'!", It: snmo Crp."o, in S''"'- botn.
The uutrr layer1 of reaction products appears to be a mixture of corrosion produc:'...; .::;!
effluent particles.
The presence of sulfides on thes-: alloys rr.ay have arl:;ซ-n fr"".m "ondensat Jr.t: of
"aj.SOjj or >o.T)|| on the r.[)eclmens, although ^00 C Is near '.he lower thr'-.-.hcl'J reported
for that su7fldatlon nechanlr.tn. ' Neither of those compound.' war. ':<:' ' ! !:i the
solids analyses, including a sample collected on the test sect. Ion cold f'.s:."er and
analyzed by atomic absorption, but only a trace amount is needed : :: ' nit late th? reac-
tion arid the possibility cannot be ruled out. The coal analyses do Include .", '^i an-i
K, and the !!a did not show up downstream; but the 1 lr:.estone contains a iur.Ir.u:-. sili-
cates, which it has been suggested might be effective in tying up the ulk.-il! :.vtals
an.,1 thus preventing sulfate formation.
An altersiatlve mechanism for sul f i'iatlor; Is reaction with gaseous :;0-. A trial
calculation on the basis of equilibrium gas compositions yielded values for- '.h1"1 ."?.
0;j, and SO, pressures such that sulftdation should not occur, according to the therrr.c;-
chemical diagrams. However, the molecular gas transport mechanism, for example, which
carries SO^ into the oxygen-poor Interior of the alloy, has been reported- to cause
sulfidatlon with an SOj pressure as low as the average calculated for these ru:?s.
Also, that average value ma; well have been exceeded substantially for short periods.
In the final test, at 270 m/sec and 790 C for 36 hours, the s-ime types of corro-
sion products were presumably formed as In the previous run, but the reaction r.ro-ijcts
were removed more rapidly at the higher velocity. Figure 17 shows an example of .'!!-
rich oxides underlaid by a Cr-rich region and then a depleted son? cont.a Snir.~ 'I?-
particles. How-'/er, as seen in Fig. 18, most of the reaction products and depleted
zones have been eroded away. Examination of crocr sections near the free ends of the
specimens Instead of at the centers, where the temperature was perhaps lower than at
the center but the erosive force was certainly less, revealed both reaction products
and depleted zones as shown in i-'lg. 19.
672
-------
HI IN-100
O U-700
^ IN-792
Q MM509
20
^S 18
>J "*
Elk 16
u? 14
1 12
i 10
8
S 6
? 4
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150 270 150
GAS VELOCITY, m/sec
Figure 11. Specimen Ccntar Section MซUI Lou.
270
:^f.,
: >n '><ป? .
C'VS
DIRECTION v->V/<
IN-100
20 urn
U-700
IN-792
MM-509
Figure 12. Face$ of Alloys Exposed it 150 m/ปc; 745 C. 100 houra.
673
-------
RFACTION PRODUCTS
OUTSIDE EDGE
Al\\ %>.:Cr
Si Au ti CO'
AI : , '. \ Ni
Au Ti Cr Co
Figure 13. IN-100 Exposed at 150 m/sec; 745 C. 100 hours.
GAS
DIRECTION
20 Mm
Figure 14. Left Facet of Alloys Exposed at 150 m/sec; 800 C. 91 hours.
674
-------
GAS
JRFCTIOfv ?
TRAILING EDGE
DEPLETED
ZONE WITH
PARTICLES
CAS
DIRECTION
AREA WITH
PARTICLES
Au S ' '. ' Ni
fi C'r Co
Au S fi \ \ \i
C'r Co
Figure 15. Sulfide Part:clot in Depleted Zone on U 700 Exposed it
1 SO in/MC; 800 C, 91 hours.
Figure 16. Microprobe Confirmation of Sulfidaticn on IN-100 Expond at
ISO m/sec; 800 C. 91 hourj
675
-------
.-.' . ;'' *;' "ivi1*'
' '
Figure 17. Resrfon Production Left Face of IN-100 Expowd at
270 m'sec. 790 C. 36hours.
i
11)
Figure 18. Right F
-------
;!ป"(, 1111',
ป
I'.KM'
'.'"': '-
l.i 700
-v '
IN-792
Figure 19 Right Facn Near Emh ot Alloys Exposed at 270 m/sec: 790 C. 36 hours.
677
-------
!'o;;r- alloy:; r,t !.:<: :.:;
r,r.iv :-c:;;:t. Jo:. oi:::<.-./'.-'
(3) I.'..-|.o:::l:: o!' boi' '.!iซ- Tour alloy:: ..:.']
wvr'.1 :;lif,li'- undo:- ;.:!e::o :-.--vo:-'- -onfjlt Ions.
1 . .!. Ct:-!:iC/:r .'iiri ::. Krl !;:, -ti::i-h "'.r>r,'-r':i<.'j:-'- Cor-r-oslor! ! r. Kl-j: ri I::_{ ':':'. S-.r.:i ::::.--i-:;
A".:-::-: wintor ;-:t.,;., :;-.: Y.,:*, i^.r., !?y-:.
678
-------
QUESTIONS/RESPONSES/COMMENTS
STANLEY DAPKUNAS, CHAIRMAN: We have time for a couple of
questions, if whoever has a question will please go to the microphone
and identify themselves and submit their question ir, writinq.
MR. OILS: Ray Oils, National Bureau of Standards. What was the
particle velocity?
MS. ROWE: We did not make that calculation. We calculated the
gas velocity, a d we realized that the particles a>-e slightly silver,
but we have not done that analysis.
MR. OILS: Well, yes. It could be markedly slower depending on
the configuration of the duct in which you place the specimen. How
long did you accelerate the particles, before you impinged then on
the specimens?
MS. ROWE: The entry nozzle is perhaps two inches high. It's
decreasing from 80 psig In the bed.
MR. OILS: Yes. The problem, of course, is the drag on the
particles. You may not have accelerated them long enough. In two
inches, in that type of gas stream, they will be entrained only to a
small friction of the entraining velocity; a half or a third or
something of this nature.
MS. ROWE: I don't know what those numbers are. I thoroughly
agree with you, but I don't know what you mean by "enough." It was
enough to chew up the specimens; and what we're trying to do--
MR. OILS: Yes. But not /50 ft/sec, or--
MS. ROWE: That's the gas velocity; yes.
MR. OILS: What was the actual temperature variation during the
test?
MS. ROWE: The lower temperatures were 720 and 745 Centigrade,
the higher 790 and 800.
MR. OILS: No; I mean during an individual test.
MS. ROWE: Oh, the plot that I showed? About 20 degrees up and
down.
MR. OILS: This is metal temperature?
679
-------
MS. ROWE: The metal tenperatures were within five degrees of it
and, yes, followed it up and down.
MR. DILS: All right; thank you.
STANLEY DAPKUNAS, CHAIRMAN: We have one here from Sheldon Lee
of Argonne Lab directed to Anne Rowe. The question is, "Did you
detect any alkaline metals on your testing specimens?"
MS. ROWE: No, we did not. But let me remind you that the first
thing that I did when I got them out of the bed was to rinse them off
in cold water. We frankly didn't expect to find sulfidation at
that low a temperature; 80D was our top temperature, which is pretty
well at the bottom of the temperature range for the usual sulfidation
mechanisms. The next tine, I'll use the hot water rinse, and analyze
for tra^j metals; but we didn't do it in this case.
STANLEY DAPKUNAS, CHAIRMAN: Thank you.
680
-------
INTRODUCTION
STANLEY PAPKIINAS, CHAIRMAN: Our next paper is entitled "High-
Tenperature Corrosion of Metals and Alloys in Fluidized Red Combustion
Systems," by John Stringer of EPRI. John graduated from the University
of Liverpool in 1955 with his bachelor's. His Ph.n. was received in
'58; and in '75, he was awarded the degree of Doctor of Engineering
by the University.
He has been at Liverpool except for the period of '63 to '66,
when he was at Rattelle Columbus, and since sometime in 1977, he has
been at EPRI.
Dr. Stringer asks that Robert LaNouze and Eddie Rogers of the
Coal Research Establishment at the National Coal Board be credited
for coauthorship of his paper.
631
-------
High-Temperature Corrosion of Metals
and Alloys in Fluidized Bed
Combustion Systems
John Stringer
Electric Power Research Institute
R. D. La Nauze and E. A. Rogers
Coal Research Establishment
National Coal Board
ABSTRACT
A series of tests has boon conducted to examine the corrosion of metals and
alloys in various locations in atmospheric pressure fluidized bed combustors. These
tests l.ave revealed that under certain circumstances severe sulfidation/oxidation cor-
rosion of in-bed comisonents can occur. A compact deposit formed on tho surface of the
in-bed components which was rich in calcium sulphate when a limestone acceptor was used,
but the presence of the deposit alone did not seem sufficient for corrosion to occur.
An additional factor is probably the presence of local low oxyqen a< vity renions
associated with relatively static parts of the bed. The various factors are discussed,
and possible remedies suqqested.
INTRODUCTION
Fireside corrosion of superheaters in conventional pulverized coal-fired boilers
is due to the formation on the metal of a deposit of ash containinq alkali sulfates.
The detailed mechanism of attack is still a matter of some controversy, but tho forma-
tion of the ash deposit is caused by the partial meltinq of the ash particles in tho
hot combustion oases. The alkalis are released from the minerals in the coal, and
react with sulfur .ind oxyqen to form sul fates: it appears possible th.it complex sul-
fatcsf perhaps rontaininq iron, are formed. These require the presence of hiqh local
partial pressures of KO-, which may develop beneath the ash deposits. The sulfatcs are
molten at least in some parts of the deposits.
The relatively low combustion temperature in a t'luidizoi! bed combustor (fb'.'l
should:-
(a) prevent ash fusion
(b) limit tho release of alkalis
In addition, if there is a sulfur sorbent present ?.n the .bed, the formation of
hiqh local SO activities would appear to be unlikely.
For these reasons fireside corrosion is less of a problem in an fbc than in
conventional coal-fired boilers. A literature survey (1), however, showed that very
occasionally severe corrosion ot in-bed components at hiqh metal temperatures was
encountered. As a result a uroqram was initiated at the Coal Research F.stabl ishmcnt
(CRE) of the National Coal Board (VCB) under the sponsorship of the Electric Power
"osearch Institute (EPRI) to study the corrosion of a range of alloys in different
nations in a fluidized bod combustor. This proqram referred to as the KPRI tests was
managed by Combustion Systems Ltd., l!K. Tests wore carried out at CRK. Metallooraphic
examination of duplicate sets of specimens wore undertaken by CRE and in America by
General Electric Co. and Foster Wheeler Development Corp.
Subsequently a joint Ncn/EPRI scries of tests have been undertaken as a follow
up to the EPRI tests.
This paper discusses both series of tests. It attempts to outline the important
conclusions that can be drawn from the data, to elucidate what materials problem areas
exist and what steps can or should be undertaken to eliminate them. More detailed
metallographic results on the EPRI proqram appear elsewhere (2).
682
-------
THE EPRI TESTS
This program used the 0.3n (12 in) square atmosoheric pressure combustor at CRH.
Turbine materials were tested for possib'c pressurized fbc applications, and for this
reason the nomine1 bed temperature of 900 C (1650 F) was 50 c (90 F) higher than used
previously (3) (4). The other bed operating conditions were:-
Fluidizing velocity 0.9 m/s (3 ft/s)
Excess air 10-20?
Coal Illinois No. 6
Acceptor Penrith (UK) limestone
Ca:S i tio 3:1
SO cc :ent of exhaust gas 300-400 ppm
Two j '";" h tests wcro conducted under nominally identical conditions.
Specimens of different alloys (Table I) were tested as:-
(i) In-bed air-cooled ' -.ibes
(ii) Above bed (freobo-Ti ) air-cooled tubes
(iii) In-bed uncooleo Trjpons
(iv) Freeboard uncoo" d coupons
(v) Pins in the ex'iaust gases after the secondary cyclone.
The last group of specimens was included to test possible turbine materials.
Three test sections were installed .nftcr the secondary cyclone, the first and second
sections wore designed to operate at the scimc nominal gas temperature, but at nominal
gas velocities of 30.5 m/s (100 ft/s) and 61 m/s (200 ft/s), respectively. The third
section was Designed to operate at -i gas velocity of 30.5 m/s (100 ft/s) and at a lower
temperature. in practice, because of heat losses through the walls of the system, it
proved extremely difficult 'o attain the required temperatures in the turbine test sec-
tion. Some natural gas was burned in the freeboard to raise the exit gas temperature,
but the amount of this was limited by the need to i:void sintering of the dust and to
minimize the change in exit gas composition. For the tests, the nominal temperature
.';i the first and second test sections was 800 C (1470 F) with 720 C (1330 F) in the
third section. It seems likely that tho temperatures were tco low and the times too
short to aive a realistic estimate of the likelihood of hot corrosion in a gas turbine.
Similarly, tho velocities were almost certainly too low to give an estimate of possible
erosion in the turbine. Because of this, the primary node of attack in the turbine
naterials was difficult to ascertain. Oxidation was felt to be the r.ajor form of
attiick although sulfidation was detected *n several specimens. C,~ 25-J1 demonstrated
the best resistance to attach, while r,TD-lll, i:;-713 LC ar.d i;;-738 also performed well.
Inconel C71 exhibited very severe local corrosion.
RESULTS
Tn-Bod Air Cooled Tubes
Specimen temperatures
There were four in-bed specimen tubes, each compiled of rino scaments 19.n (0.75
in) long: 50rni (2 in) OD and approximately 42mm (1.65 in) ID. The nominal temperatures
of these fcjr tubes were 540, 650, 760 and 840 C (1000, 1200, 1400 and 1550 F)'. The
two lower temperature tubes were chosen to study conditions appropriate to steam supor-
heater tubvs, th<_- upper two to simulate conditions Appropriate to air cycle applica-
tions.
The tube temperatures were measured with thermocouples in vel's drilled in the
rinos, and thus wore a measure of the mid-wall temperature. The maximum surface
temperatures for cacn tube were higher', as indicated below.
During the experimental runs the temperatures were measured with four thornxi-
couples along the length of .ich tube at the 3 or 9 o'clock positions. For an initial
proving run, two rings on a i 'be carried four thermocouples; one at the top, one at the
bottom, and two at opposite si.'es in the 3 and 9 o'cป k positions. These revealed a
circumferential temperature varidtion which increased as the temperature of the tube
decreased. The two rings showed different raaanitudss of the effect, suggesting that it
683
-------
Table I. Typical vJomposition of Alloys Tested
Ferri t ic
Steels
Cor- Ten B
2V Cr-1 Mo
9 Cr-1 Mo
Type 405 SS
E-Brito 26-1
GE 2541
Austenit ic
Steels
(Cr-Ni-Fe)
Nitronic 50
Type 321 SS
Type 310 SS
Type 347 II SS
Type 329 SS
21-6-9
Incoloy 800
.''anauritc 36X
liK-40
Nickel-Base
Alloys
Inconel 690
Inconel 601
Inconel 617
Inconel 671
Hastclloy X
RA 333
U-500
U-700
IN-738
GTD-111
IN 713 LC
Cobalt Base
Alloys
HA 188
X-40
XSX-414
C
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
0.
2
1
1
05
001
C2
05
06
04
06
05
03
04
4
4
05
04
07
05
1
04
08
07
1
1
1
0.08
0.
0.
5
2
Cr
0.
2.
9
12
26
25
22
18
25
18
28
20
19
25
25
27
22
22
48
21.
25
19
14
15
14
13
22
25
30
Ni
5
2
-
0.3
0.1
-
12
11.6
17.2
9.5
4.2
7.2
31.5
34
20
64.6
61
56
51
5 bal
45
bal
.6 bal
.7 bal
bal
.6 bal
.2 22.3
.8 10.6
10.8
Fe Co Mo
bal
bal - 1
bal - 1
bal
bal 0.01 1
69.5
bal - 2
bal - 0.3
bal - 0.4
bal 0.24 0.4
bal 0.2 1.5
ba 1 - 0.1
46.5
bal
bal
8 - -
15.1
0.2 12.2 8.7
0.4 -
18.5 2.1 9
18 33
0.3 19 4
0.1 15 4.2
0.2 8 1.7
0.1 9.7 1.5
0.2 0.5 4.6
1.8 bal
0.2 bal
1.2 bal
K M Ti
-
- - - 0
0
-0.1-0
- - - 0
4.76 - 0
_
- 0.4 1
- - - 1
- - - 1
- - - 0
-
0.4 0.4 0
- - - 1
1
0
1.5 0.4 0
-1-0
0.2 0
0.7 0
- - - 1
3.0 3.0 <0
4.4 3.5 '0
2.7 3.5 3.5 0
3.0 5 '0
6.0 '0.9
-------
varied along the tube, but both showed the same general form of the temperature distri-
bution.
From this information the outside skin temperatures have been calculated to
exceed the nominal tube temperatures by a maximum of 86, 73, fiO and 50 C (155, 132, 108,
90 F) on the 540, 650, 760 and 840 C (1000, 1200. 1400 and 1550 F) tubes respectively.
Generally, the hottest part of the tube was towards the top, and this may at
first sight appear surprising. As only four point temperature measurements were made,
it is only clear that the tubes were hotter towards the top. This may be consistent
with measured circumferential variations of heat transfer coefficients, see for example
Noack (5) which shows high heat transfer rates from the 10 o'clock to the 2 o'clock
positions. Clearly, a circumferential temperature variation of this magnitude is
worrying, and from the point of view of materials selection it is important to estab-
lish whether this is a general characteristic of tubes in fluidized beds.
Tube Deposits
At the end of each test period the tubes were found to be covered with a fairly thick
polished deposit (Figure 1). This was reddish brown on the upper two tubes and darker
on the lower two tubes. This color difference was a function of bed position not of
tube temperature, because the position of the 650 C (1200 F) and 760 C (3400 F) tubes
were exchanged for the second run. The deposit consisted largely of calcium sulfate;
there appeared to be some calcium oxide present together with other components derived
from the coal ash. There was a very small amount of cnrbon detectable. No systematic
variation in composition could be associated with the color difference.
The deformation temperature of the deposit WPS found to be 1220 C (2230 F) which
is well above the bed temperature suggesting that no significant part was molten at
temperature. However, this measurement was done under oxidizing and neutral conditions,
and as will appear loter, reducing conditions might have been more appropriate. It is
common for the deformation temperature of coal ashes to be lower in reducing atmos-
pheres than in oxidizing, although the difference is not usually greater than 100 C.
CRE have been attempting to measure the sintering temperature of fine ash particles as
distinct from the standard deformation temperature. Although this work is still in the
development stage, indications are that fine Illinois No. 6 ash particles sinter in a
mildly reducing atmosphere at a temperature close to that of the bed.
Under the scanning electron microscope, the deposits appeared remarkably compact
and pore-free (Figure 2). composed of particles whose size appeared to be in the range
l-5:-m. This was very similar indeed to the particle size in the limestone used (Figure
3). High pressure mercury porosimetry on a deposit sample showed that there was less
than 1% porosity for pore entrances with diameters in the range 58 - 0.014um, support-
ing the scanning electron microscope observati ns.
The severity of attack of the metal di'l not seem to be influenced by variations
in the deposit thickness.
Corrosion
The most obvious feature of the corrosion of the in-bed tubes was the presence
of sulfides beneath the surface of the metal in all the alloys. The stainless steels
exhibited a fairly uniform band containing discrete, more-or-less spherical sulfides.
A typical example is shown in Figure 4, a specimen exposed during the NCB/EPRI tests.
lor the three hotter tubes, the more sophisticated alloys such as Inconel 617, incoloy
800 and Haynes 188 showed considerable grain-boundary penetration of sulfides (Figure
5). These alloys also showed local breakdown of protective behavior, with thick oxide
scales forming beneath the deposit and very considerable sulfi.de formation boneath the
oxide. Often -voids appeared to be present in the metal, which could have been due to
loss of particles, e.g. sulfides or oxides, during specimen'preparation. Between the
oxide and the deposit there was frequently a bright white phase which electron probe
raicroanalysis showed to be a sulfide of the base metals iron, nickel and cobalt, (Fig-
ure 6). Since these sulfides are not very stable, it demonstrates that the sulfur
activity at this point must have been high.
Figure 7 shows typical data for the scale thickness and internal penetration.
The temperature dependence of the corrosion did not appear to be very great. Incoloy
685
\
-------
Figure 1. The General Appearance of tin- Tubes
Alter Withdrawal from the Combustor After the
First 1000 lu 1 f,t. Showmq the Polished Deposit
tffl -
*.:
Figure 2. The Appearance of a Fracture
Cross Section of the Deposit
(scanning electron micrograph: x 2000)
Figure3. The Surf ace of a Fragment of
Pennth Limestone (SEM: x 2000).
686
-------
Fiqi*rปป4. A Cross Section of a Sp^umen of Type
347M s \ ( ซ|iriซM lot 1".() hr ,il a Met ll T.", J>.M.,
!..! til 760C (1400r-i ป BOO
figured. A S|>**cimen of Hayrw& 188 ฃ
lot 1000 hr at a MeTal Teni|>etdtuie ot 180C
(1550FI The D.irk Upp>-r Section r. the [>|>nsit
Pnnciprilly Sulphate. The Liqht Colored Ph iw
InitneilMtely Below the Deposit is a Cobalt
Nickle Sulphide, the Uiqlit Grey Beneath Thii
is the Oxide The Metal. With Internal Penetration
ol Guide and Sulphide !ป 300)
Figures. Specimen UR4| Inconel 617 Enposed for 250 h in NCB/EPRI Test 2. Metal Temperature
843 C (1650 F). fcr.oi.jY Dispersive Analysis ot Liqhi Globular Phase on Outside ot Suie.
687
-------
HASTEUOV X
INCONEl 617
INCONEL 601
INCONEl 671
MAXIMUM MAXIMUM
CORROSIVE SCALE
PENETRATION
Figure 7. Typical Data for the Maximum Corrosion of
the Alloys (second 1000 h test).
688
-------
800 was more severely corroded on the tube with a nominal temperature of 760 C (1400 F)
than on the tubes at 6SO C (1200 F) and 840 C (1550 F). but the difference lay within
what might be expected to be normal scatter.
The extent of attack en the 540 C (1000 F) tube was less than that observed on
the higher temperature tubes. This may be partially due to the use on that tube of
alloys somewhat less sensitive to sulfidation attack.
There was a marked variation in the attack around the circumference of the tubes,
some regions appearing quite free of accelerated corrosion, while others exhibited
severe sulfidation. There was a tendency for the attack to be greater at the top of
tubes, particularly for the sensitive alloys such as Incoloy 800 and Inconel 617. It
seems probable that this could not be attributed to the temperature variations alonซ,
since:-
(a) the temperature dependence- of the corrosion appeared to be snail,
(b) the circumferential variation of attack was also present on the nominal
840 C (1550 F) tube, on which the circumferential variations of temperature
were small.
It was particularly interesting that Inconel 671, an alloy well-known for its
resistance to aggressive oxidizing environments, suffered catastrophic corrosion. This
was the worst of the alloys investigated (Figure 8).
The low-alloy feiritic steels, Corten B and 2V Cr-lMo oxidized rapidly. However,
in this range the oxidation of these stools is markedly dependent on temperature, and
the considerable temperature variations on tho nominal 540 C (1000 F) tube nakes it
difficult to pass judgement on the behaviour of these materials within che fluidized
bc-1.
Effect of Exposure Time
It. had originally been the intention to expose most of the rings for 2000 h,
removing a small number at 1000 h and replacing them with duplicates. In the event,
the rings were sufficiently disturbed during cooling at the end of the first test to
fail a leak tost, so only 4 rings on tho 650 C (1200 F) tubes wore resubmitted foi the
second tost. Those showed much less than twice the attack of those exposed for 1000 h
implying that the rate of attack was diminishing with time.
Corrosion Criterion
The Central Electricity Generating Board in the UK use a corrosion criterion
which is equivalent to the loss of 7 mm of metal in 200,000 h, corresponding to the
total loss of tho tube wall in tho lifetime of a boiler. Since generally corrosion
rates fall with time, this gives a reasonably conservative criterion for the extrapola-
tion of tests of around 10.000 h length. For a 1000 h test, 'his would he equivalent
to a maximum loss of structural notal (metal lost by scaling plus the metal affected
by internal sulfidation or oxidation) of 35:.m assuming a linear rate of loss. This is
almost certainly an excessively severe criterion. For example, if it is assumed that
the rate of metal loss varies with the square root of time (parabolic oxidation) and
that an acceptable loss after 200,000 h is 10% of the wall thickness, 700_m, then the
acceptable loss at 1000 h would be 50;.m. Both criteria would give 70um at 2000 h.
Very few of the alloys tested match these criteria. At 650 C (1200 F) it scons
likely that several alloys, includina the austcni!-.ic stainless stools, arc reasonably
close, but at the higher metal temperatures no alloy meets those criteria. Some alloys
may be acceptable for use at the higher temperatures if a less stringent corrosion cri-
terion can be tolerated.
Above-Bod Air-Coolcd Tubes
The two above-bed tubes had nominal metal temperatures of 650 C (1200 F) and
760 C (1400 F) . The deposit which forr.ed on these tubes was relatively loose and
friable. The tubes did not show appreciable corrosion, and several alloys appear cap-
able of meeting performance criteria (2).
689
-------
Figur* 8. A Corrosion Pit on an Incond 671 Specimen
Exposed for 250 h it Metal TwnfMriture of
840C(1550F)
-------
The Uncooled Coupons
As might be expected from the results of the cooled <;Decimens, several of the
coupon materials exhibited severe sulfidation/oxidation attack in the bed, notably
Inconel 671, '2). This alloy also underwent severe corrosion in the freeboard.
Several other alloys had sulfidation attack in the freeboard. This may be due to the
accumulation of deposit on the coupons, or to the proximity of the natural gas flame.
However, the attack in the freeboard was generally oxidation. Within the bed, several
alloys appeared to be corroding slowly enough to be acceptable as uncooled structural
component materials, notably a G.E. alloy, GE 2541 which resembles Kanthal, and Haynes
Alloy 188. Two cast austenitics, UK 40 and Manaurito 36X were also marginally accept-
able.
DISCUSSION OF THE EPRI TESTS
Clearly the most serious aspect of the EPRI tests was the sulfidation/oxidation
of the in-bed tubes. This contrasts with an earlier series of tests under the sponsor-
ship cf EPA using the same combustor and similar nominal conditions which revealed
little sulfidjtion (3) (4). The earlier test used a lower bod temperature 850 C
(1560 F) and the significance of this is discussed later.
Calcium sulfate can decompose to release sulfur, by a process such as
CaSO. ป CaO + ปs S, + 3/2 O,
4 22
and it is obvious that the lower the oxygen partial pressure (activity), the higher the
sulfur partial pressure (activity) and vice versa. The normal excess oxygen levels
would correspond to an oxygen partial pressure of 10 atm and at the bed temperature
this would be equivalent to a sulfur partial pressure of 10~ atm or less. This is
insufficient to sulfidizc any of the components of the alloys.
However, it is well known that fluidizod beds consist of two phases, a bubble
phase and a dense or emulsion phase which is like the bed at incipient tluidization.
The greater proportion of the fluidizing gas passes through the bed as bubbles. Since
the burning carbon particles reside in the dense phase, the oxygen concentration in
this phase depends amongst other things on the rate of interchange of gas between the
bubble and ciense phases. Fluctuating oxygen concentrations in a fluidizcd bed com-
bustor have been demonstrated by Gibbs et al (6) using a fast response mass spectro-
meter. The oxygen concentration in the dense phase was found to be significantly
lower than that of the bubble phase.
Experiments with a stabilized zirconia probe (7) have suggested that oxygen par-
tial pressures of 10 atm may be present locally in the bed which corresponds to a
sulfur partial pressure of around 10~ atm. This is sufficient to sulfidize almost any
alloy and certainly high enough for nickel sulf Ide to be stable. The fact that iron
and nickel-containing sulfides arc found between the oxid? scale and the deposit proves
that in this location the sulfur partial pressure does indeed reach values close to
those calculated. Figure 9, the Ca-O-S thermodynamic stability diagram, and Figure 10,
the stability diagrams for the major alloy elc icnts (8) , show that this condition
produces an environment which is close to the oxide/sulfide/sulfate boundary.
Low local oxygen partial pressures may be attained in several ways.
(i) High concentration of coal, due for example to the proximity of a coal feed
port.
(ii) A relatively stagnant region of bed material which becomes oxygen starved.
(iii) Non-uniform flow of air through the bed, with air channelling through
certain paths.
(iv) Uneven coal feeding: if the coal feed increases markedly for a short time
the oxygen partial pressure can drop abruptly-
(v) Cn the injection of coal, the vclatiles are released quite rapidly and can
pass upward through the bed largely uncombusted; they burn in the freeboard
691
-------
-60
Figure 9. Phasa Stability Diagram for the Ca-O-S System at 1000
andZOOOK (1310 and 1670F).
IV
s
0
-5
ซx 10
&
S .,,
-20
-25
-30
-35
-4O
Cf
Suil.de / *'// '
/ /^Sฎฐ
~ ** /
/ s
Fe. N. / /
Cr
Mn
-
-
Meul
-
1 1 1 1 1
/
//
/
\ \
J
|
'//
S
M-O-S
114< K()600;F)
Onxit
1 1 | 1 1
-55 -50 -45 -40 -35 -30 -25 -20 -15 -10-5
5 10 20
Figure 10. Flute Stability in Fe. Ni. Co, Cr. Mn-S-O System et 1144ฐK (1600ฐ F).
692
-------
immediately above the bed. Clearly this is equivalent to a low oxygen
activity zone passing through the bed.
(vi) Defluidization of parts of the bed can produce local regions starved of
oxygen.
(vii) The dense phase, as discussed above, is itself relatively oxygen deficient.
(viii) Beneath an adherent deposit the oxygen activity may drop if oxygen removal
by the oxidation reaction is relatively rapid in comparison to the influx
of oxygen through the deposit layer. This will clrarly be favored by
thick, compact, deposits with low levels of porosity.
In the present case, the deposit alono seems insufficient to produce the low
oxygen partial pressures because some areas of the specimens wore covered witn thick
deposits but showed no signs of corrosion. It. is of course possible that the deposits
had different porosity in different regions, but this was not apparent under the
microscope. The deposit, consisting largely of calcium sulfate, is however obviously
of great importance as a source of sulfur.
The 'stagnant region" hypothesis would imply that only the tops of the tubes
would be attacked. As mentioned before, thc-re was indeed a tendency for the top of
the tube to be rather more corroded, but s-ilfidation was not confined to trioso regions.
This therefore appears to be a factor in p. Jducing a corrosive environment, but is not
in itself sufficient to explain the results.
Thore was no obvious tendency for the attack to bo more severe near the coal
port; that is the 650 C (1200 F) and 760 C (1400 F) tubes were not obviously more
corroded when they were in the lower row as opposed to the upper row. Neither were
the center rings on the tubes more corroded than those towards the ends. These con-
clusions may be altered when the experimental data are subjected to more detailed
analysis.
There was no evidence of uneven coal feeding or dcfluidization in the tests.
Non-uniform air flow should be reflected in a non-uniform pattern of attack. While the
tests were not dosijnod to reveal such non uniformity, it would probably have been
detected.
On balance, therefore, it seemed likely thit the major factors contributing tc
low oxygon activity, in order of importance, werc:-
1. Relatively stagnant regions of bed material on top of the tubes.
2. The presence of thick, compact deposits.
It follows that possible solutions to the corrosion problem include:-
(a) Prevent the build up of a calcium sulfato-higli layer on the tubes.
(b) Prevent the establishment of stagnant regisns Above the tubes by adjustir.j:-
(i) Fluidizing velocity
(ii) Tube dimensions
(iii) Geometry of the tube array
(c) Try to ensure good gas and solids mixing in the bed to reduce reqions of
low oxygen activity. At the moment there appears to be little information
on the level of oxygen activity in the bed, nor arc there combustion models
of sufficient complexity from which the effects of operatini variables on
the oxygen level can be estimated with confidence. Factors which might be
of importance include:-
(i) Excess air
(ii) Fluidizing velocity
(iii) Distributor dcsinn
(iv) Method of coal feeding
(v) Form and location of coal feed port
693
-------
f-.
(vi) Macro solids flow patterns
THE NCB/EPRI TESTS
A joint NCB/EPRT program has commenced to determine the important factors con-
tributing to the corrosion observed in the EPRI tests. A run time of 250 h was
selected, because it was believed that now the pattern of corrosion had been recognized,
a shorter time could be tolerated. Attention was focussed en the in-bed tubes, with a
rather more restricted range of alloys but including more stainless steels of the 18/10
type. The first three tests had the following conditions:-
Tost 1 was essentially a duplicate of the EPRI conditions, to be sure that sul-
fidation/oxidation could indeed be detected after 250 h.
Test 2 used a bed temperature of 850 C (1560 F), so that the conditions were
nominally the same as those in the earlier EPA supported tests, (3) which had
produced only slight sulfidation/oxidation corrosion. This was to test the
view that the reason for the severe corrosion in the EPRI tests was the higher
bed temperature.
Test 3 used a bed temperature of 900 C (1650 F) without limestone addition.
The bed consisted entirely of coal ash.
PRELIMINARY RESULTS OF THE NCB/EPRI TESTS
The data from those three tests have not yet been fully analyzed, and the
results presented here must be regarded as preliminary. However, the major points are
unlikely to bn altered by more detailed study.
The first test produced corrosion essentially similar in character to that
reported for the EPRI tests, with sensitive alloys such as Inconel 671, Inconel 617
and Incoloy 800 showing significant sulf idation/ox'idation corrosion.
The second tost showed a very similar degree of attack, suggesting that bed
temperature in this instance is not a major variable in determining the extent of
corrosion. In detail, it appears that whereas the relatively resistant materials, such
as Type 347 s.s., and the very sensitive material Inconel 671 seemed to behave in the
same way in both tests, both Inconel 617 and Incoloy 800 showed reduced corrosion in
the second test, though this reduction was not great. This suggests that the latter
two alloys may be sensitive indicators of corrosive potential in fluidizcd beds.
The third test resulted in very little corrosion indeed. For most specimens,
the outer surface of the tube appeared less oxidized than the bore. Even Inconel 671
showed no attack. No specimens showed any evidence of sulfidation. This demonstrates
the importance of the CaSO./CaO equilibrium in producing the sulfidation attack. A
deposit did form on the tuBcs, and it was of a similar thickness to that formed in the
other tests; the grain size nlso seemed much the same under the nicroscope. However,
electron probe microanalysis showed it to contain largely aluminum and silicon, with a
smaller amount of iron and no calcium or sulfur. This is consistent with its being
coal ash.
It is clearly important to determine the difference between the second of these
tests, which produced significant corrosion, and the EPA tests, which did rot. The
test conditions were the same, but the coals and limestones used were different. The
iimestone used in the EPA tests was U.S. limestone No. 18, which contained 15% quartz
(SiO^; as relatively large crystals (around 200 urn) intimately mixed with the 1-2 urn
calcfte crystals. The limestone used in the tests described in this paper was o U.K.
Penrith limestone containing 99.8% calcite. The coal used in the EPA tests was Pitts-
burgh Humphrey No. 7, whereas that used in the present tests was Illinois No. 6; the
differences in the compositions of these two coals do not seem great. The effect of
the limestone and the coal on the corrosion will be examined in future tests. The
lower sintering temperature of Illinois No. 6 as reported earlier also requires further
consideration.
694
-------
CONCLUSIONS
It is clear from the experiments described above and from other experience
reported in the literature that while fluidized bed cornbustors can operate with no
corrosion at all, there is a risk of sulf idation/oxidation corrosion of high tempera
ture metal components. Once initiated, this form of attack can produce very rapid
local degradation of alloys sensitive to this type of corrosion. The conditions1
are likely to lead to this tvpe of attack are:-
1. The presence of calcium sulfatc as a deposit on the metal surface.
2. The existence of a local region of low oxygen activity near the metal.
3. The presence of a sensitive alloy in the low oxygen activity region.
4. Metal temperatures above 650ฐC.
At th<_- moment, the factors that determine whether cr not a calcium sulfate layer
forms on the tubes are not understood. Some of the factors likely to affect the exist-
once of low oxygen activity regions have been listed earlier, and it may be possible to
optimize bed operation to eliminate the corrosion risk. The most important contribu-
tion of the present investigation has been to establish that alloys such as Inconel 671,
Incoloy 800, Inconel 617 and perhaps Haynes Alloy 188 arc sensitive to corrosion in the
bed; but the stainless stools such as Types 304, 329 and ?47 are relatively insensitive;
they do sulfidizc but the morphology of the sulfidation does not appear to load to
catastrophic attack.
Further work is clearly required on these aspects. Experiments should be con-
ducted to determine the local oxyqen activities within the bed, and the effect of
operating variables on the oxygen distribution. The testing of the nore resistant
materials should be extended to longer times to ensure that the sulf idation/oxidation
attack does not eventually become catastrophic. The possibility of using coatings to
resist attack should also bo investigated.
It docs not seem that in-bed corrosion is an insuperable problem limiting the
technology; solutions are almost certainly available. But a note of caution has been
expressed which can be overcome if the development of a proper understanding of the
corrosion problem is undertaken.
ACKNOWLEDGEMENT
The experimental work was supported by the Electric Power Research Institute
under contract numbers RP 388 and RP 979-1; and by the National Coal Board. The con-
tributions of Mr. A. J. Minchcner, Mr. J. C. Holder and Mr. A. J. Page of the Coal
Research Establishment and Dr. D. P. Whittle of the University of Liverpool were of
great value. The views expressed are those of the authors and not necessarily those of
the Electric Power Research Institute or the National Coal Board.
REFERENCES
1. J. Stringer and S. Ehrlich. ASME paper 76-WA/CD-4 (1976) pp 11.
2. J. C. Holder, R. D. La Nauze, E. A. Rogers and G. G. Thurlow. Paper to Eng. Found.
Conf. on Ash Deposits and Corrosion. Henniker, New Hampshire, 26 June - 1 July
1977.
3. Fluidized Combustion Control Group, NCB 'Reduction in Atmospheric Pollution' Final
Report of the National Caol Board to EPA, Reference No. DUB 060971 (Sept. 1971) .
4. M. J. Cooke and E. A. Rogers. Paper No. B6 Inst. of Fuel Syrap. Set. No. 1. Sept.
1975.
5. R. Noack. Chemio Ing Technk 1970, 42, 371.
6. B. M. Gibbs, F. J. Pcrcira, J. H. Beer. Paper D6, Inst of Fuel. Symp. Ser. No. 1.
Sept. 1975.
7. M. J. Cooke, A. J. B. Cutler and E. Raask. J. Inst. Fuel, 45 (1972) 153.
8. P. L. Hemmings and R. A. Perkins. Report No. LMSC-D558238 on EPRI Project No.
RP716-1 (March, 1977).
695
-------
QUESTIONS/RESPONSES/COMMENTS
STANLFY PAPKUNAS, CHAIRMAN: Thank you very much, John. Are
there any questions? Pon't run away. Please go to the nicrophone.
PROF. PEER : .1. Reer, MIT. A couple of years ago, at Sheffield
University, they carried out and reported sone experimental data on a
one square foot fluidized conhustor, in which detailed species
concentration and distribution, tine-resolved and spatially-resolved
distribution, were measured in a fluidized bed.
We found that at about 18 inches above the distributor plate,
there were strongly fluctuating conditions when the time-average
oxygen concentration was around 3 percent; the fluctuation ran
between 4 to 4-1/2 percent and reducing conditions. It seemed to us
that these are extremely unfavorable conditions for any metal, and
they are very favorable for sulfate formation. I wonder if Dr.
Stringer agrees.
Ry the way, I would like to mention that this work was spon-
sored by the National Coal Roard, but somehow they didn't very much
look at the results which they have received.
PR. STRINGER: Thanks for the piece of information. It's very
useful. I was aware of one measurement, the measurement I mentioned
using a zirconia probe, which was done by the National Coal Board
with some help from CEGR; and that was published about 1972, which
again showed exactly the thing you say, short-term fluctuations. And
yes; I agree entirely. That would he very bad.
PROF. REER: It's with one cycle per sec. So fluctuations
between 4 percent and zero, or reducing at one cycle per second, at
18 inches above the distributor plate.
PR. STRINGER: I think that would be consistent with the single
zirconia probe.
PROF. REER: And with 20 percent excess air overall in the bed.
PR. STRINGER: Yes. Thank you very much. That emphasizes what
difficult conditions one can have under circumstances where one
wouldn't dream they would occur.
STANLEY PAPKUNAS, CHAIRMAN: We have a written question here
for you to answer while the gentleman is going to the microphone.
696
-------
PR. STRINGER: Okay. This is fron J. Mogul of Curtiss Wright,
who wants to know what results we had on the iron-chroniun-aluminum-
yttrium alloy in the test. That was the GE 2541 alloy, which is
basically a kanthal containing yttrium to hold the alunina scale on.
This was just used as an in-hed component of the test for support
pieces, and it looked good. It was far and av/ay the best of the
in-bed materials that we had, and would certainly be acceptable as an
in-bed supportive material.
MR. YF.3USHALMI: Joe Yerushalni, The City College. We have
heard in the past day and a half that the partial slumping of the bed
night prove a means of achieving turndown. Now, John, are you
suggesting that that will cause disaster at the same time?
OR. STRINGER: I have been trying to find out from a number of
my more expert colleages what precisely happens to the metal when you
slump the bed on it. My initial reaction would be that the temperature
would go up fairly sharply, but I don't know how long it lasts for,
Joe. And yes, indeed, as the last speaker said, once you start
sulfidation, it will propagate by itself. So a little spike can be a
bad thing for this particular type of attack. So you know, we have
to look, I think.
STANLEY DAPKIINAS, CHAIRMAN: One more question.
MR. LEON: Okay. A.l Leon, Dorr Oliver. I have two questions.
Based on your corrosion mechanism, would you comment on the effect of
corrosion rates in an AFB versus a PFB, where a PFB would have higher
oxygen partial pressures?
DR. STRINGER: Yes. We very much hope to get some data on a
PFB, because there are two things about that. The higher oxygen
would look to be good for you. That would be a first guess. However,
don't forget that in that case we are just raising the oxygen pressure
by a factor of 4 or 5 overall, whereas the oxygen potential differ-
ence between the oxidizing and reducing regions is enormous. It is
several orders of magnitude, from say 10~12 atm going up to one
atmosphere. So it may not be as much of a help as one would hope,
but it's certainly going to be in the right direction.
However, in a pressurized fluidized bed, there are two other
things different. You don't use a limestone acceptor. You use a
dolomite acceptor, of necessity; and that is going to change the
chemistry a bit, in a way that I would not like to anticipate. And
secondly, the general distribution of the bubble and the emulsion
phases, I am told, are different; and that would introduce a further
factor which might change things.
697
-------
MR. LEON: Okay. The last question was on the effect nf verti-
cal tubes as against horizontal tubes. Horizontal tubes, you mention
a cap on top of the tubes. You wouldn't have this with vertical
tubes.
PR. STRINGER: Absolutely. You would hot have it with vertical
tubes, so long as the tube was continuing straight up to infinity.
If you have to bend the tube over at any point, then the position of
the bend might well be an excellent location for forming something
nasty.
STANLEY F1APKUNAS, CHAIRMAN: Thank you very much, John.
698
-------
1
INTRODUCTION
STANELY DAPKUNAS, CHAIRMAN: Our next speaker will be Leon
Glicksman of MIT, who will present a paper on Thermal Stresses and
Fatigue of Heat Transfer Tubes Immersed in a Fluidizert Pซvj Combt jr.
Dr. Glicksman received his Bachelor's from MIT in '59, r. Master's
from Stanford in '60 and his Ph.D. from MIT in '64. He is on the MIT
faculty in the Mechanical Engineering Department.
699
-------
Thermal Stresses and Fatigue of Heat
Transfer Tubes immersed in a Fluidized
Bed Combustor
N. Decker. L. Glicksman, R. Pe !oux. T. Shen
Department of Mechanical Engineering
Massachusetts Institute of Technology
ABSTRACT
Thermal stress conditions are investigated for horizontal tubes In a fluldized bed cocbustor.
Circumfcrcntially non-uniform temperatures and high frequency cyclic temperature distributions are pro-
jected from observed licat transfer variations on tubes immersed In fluldized beds. The thernal stress
produced by the external heat transfer variations arc shown to be less important than that caused by
Internal variations due to the boiling two-phase flow within the tube. High cycle fatigue due to ex-
ternal heat transfer fluctuations is sliovn to be unlikely.
Low cycle fatigue due to cyclic micro-yielding during each start-stop cycle of the bed was also
Investigated. Lains fatigue data for 304 stainless steel, it can be concluded that low cycle fatigue
failure will not occur.
INTRODUCTION
Fluldized bed combustors cormonly use a horizontal tubular Ucat exchanger imersed within the
Led for generating stcau. With this arrangement the gas side heat transfer coefficient la strongly
influenced by the motion of solid particles and, compared ultl. tubes in conventional boilers, there
will be differences in tiierm.il behavior which will be important to the selection of tube materials and
dimensions. Tubes within the *>cd nay be subject to corrosion and dynamic forces set up by tlte inter-
action with the bed THUS, i.i some designs of fluidizcd bed combustors the tube vails will be thicker
than a corresponding tube in a .mvcntior.al boiler operating at the same pressure. The thicker vails
will experience more severe thernal stress during operation.
In the particular case of horizontal tubes, the particle motion is not uniform around the cir-
cumference of the tu'jc. The tube experience* a spatially non-uniform heat transfer coefficient whicu
produces a tioc-averagu tncrmal .stress field which varies Circumfcrcntially .truuad the tube. The
particle notion is unsteady causing the local film coefficient cf heat transfer to fluctuate with : 'TC.
The alternating thcrrul transients induce a time varying thermal stress field in the tube wall.
In addition to these external effects, the two phase flow within the horizontal tube can becoae
stratified with the vapor occupying the upper portion o'. the lube or a vavc-like liquid flow can exist
in the tube alternately wetting and drying out the upper portion of the tube. The internal flow
conditions can adversely affect both Instantaneous and time-averaged thermal stress in the tube. A
schematic figure ot all of ch?se effects along witli a graph showing a typical external heat transfer
coefficient distribution [1), are shown in Figure 1.
Two conceivable tube failure modes arc investigated here. The therail stress caused by a large
uniform gas-side film coefficient is evaluated along with its influence upon possible low cycle fatigue
of thick walled tubes due to startup and shutdown cycles. The additional effect of spatial variations
in the time-averaged thermal stress field is considered with regard to the low cycle fat'gue problca.
Higher frequency fluctuations in the temperature and stress distribution for both internal and external
variation-, in the heat transfer coefficient are estimated to determine the likelihood of high cycle
fatigue.
Therm-.t Stress With Steady Uniform Heat Transfer
To form a basis of comparison for the non-steady and non-uniform cases, consider first a heat
exchanger tube with steady heat transfer coefficients which are uniform over both the inner and outer
tube surfaces. In addition, the bed temperature and the bulk temperature of the fluid within the tube
will be considered constant. For these conditions analytical expressions exist fcr both the temperature
distribution in the tube metal and the corresponding thermal stress. Both :re fractions of radius
alone. The temperature distribution, assuming a constant value of thermal conductivity Is given as:
700
-------
Defluidized particle cap
Possible separated
two phase flow
Intermittent contact
h fluctuates ~ 1 Hz
Defluidized gas pocket
Typical circumferential
variation of h
Figure 1. Conditions at Tube Walls (or a
Horizontal Tube in a Fluidized Bed Combustor
701
-------
TCr)
In r/r.
jj-j-yi-
where
T 1 TBED-*STn If""^
s "n ro/ri k
11/r ho + 1/^h, + p-\ I |
and
Ti ' TSTM
TBCD " TSTM
tn ro/rt
(1)
(2)
(3)
Manson (2) provides expressions for the stress within a cylinder of linear elastic material In plane
strain.
60 tt-vlr'
2 ^ 2
(T(r) -
2 _ 2 .r
(T(r) - T;rdr - (T(r) -
rl
1_^L_ f
^ - 'i2 7 r
- Trc{)rdr - / (T(r) - Tref)rdr
ฐrr> - QE(T(r) - Tref>
(4)
(5)
(6)
Integrating the stress equations us inc. the steady state temperature distribution, the stress Is found
to be:
i ฐE "TCBE ['.'<'' * rJ2) l * ln r/ri]
"08 ' l-v ^2^^ . r2} - 2 ln ro/r1 J
T ฐc "TTJBL fr.2<'2 " rt2> tn r"i ]
ฐ" ' l~v [2r2(ro2 - rt2) " 2 tn ro/rlj
(7)
tn r/r
ฐrr> - ฐE "TUBE
Ti ' Tref
1/2
(10)
The thermal stress for two tube sizes and two materials is given in Table I for the tube ends uncon-
strained. T.ie stainless steel can uc seen to experience a greater tticrrul stress due to its poor tlicr-
702
-------
T.iMv 1. Calculated Siresica - CtrciUufcrcntialiy luifora iซ--at Flux
.laterlal Sire (ut/ 10) .'.F(*F) c "(?sl) o '"*T(;>st)
1 1/4 cr-Ulo
304 SS
2 1/4 Cr-Ulo
304 ib
2.
I.
2.
2.
50/1
5J/1
O'J/1
oa/i
75
75
64
64
42
8S
19
41
i:;
6770
18510
2950
8230
on
-5340
-14600
-2580
-725d
IS
9680
21420
ao40
13300
OUT
-3430
-U700
1500
-3160
Condlciorts: t^m - 50')*F l>n - 3350 DTl7hr-f t"F
Tltli> " lSuo*r ''EXT ' 30 CTV/hr-ft2*F
P - 1000 psia
703
-------
conductivity. Vac cori'jiiied thermal anil mechanical stress exceeds its yield stress of 21,000 psi in
larger, tuicker walled tube.
i ij'ure 2 illustrates iiou these stress quantities vary with nouition In c'tc tube wall. It is
iAnt to note titat the :>tress at tiie inside surface i:; tensile and chat at the external surface is
co^vruasivc. it should also be noted that for a given inside ur cutsld*: diancter and fixed bod and
stean Lcuperaturcs a thinner tui-e vail will have smaller thermal stress, but larger raeclianic.il stress
due to t.te internal pressure.
T.iernal Stress with Circunfercntlally Varying Heat Transfer
Die idealized axisyunetric situation analyzed in ttie last section serves as a standard to which
the results for non-uniform itcat fluxes may be compared. Analytical solutions are net available fur
cite general non-uniforo case, theicforc. the temperature distributions were determined using finite
difference computer programs for specific cases.
The external heat transfer coefficient varies around the circumference of a horizontal tube due
Co Che varying degree of local particle mocion.Thc he;it tr.insfer-ttt Inroerscd surfaces in fl-:idized beds
can be related to gas and solid properties, and particle replacement frequency. I [any investigators
iiave reported frequent exchange of particles at tltc sides of horizontal tubes, but long residence times
for particles in a stagnant wake region or cap on Che top of tubes and a long residence time of voids
at Cue lower surface of Che tuh'fs. Thus, as expected, the upper and lover portions of the tube surface
have been observed Co liavc lower local film coefflcicncs 11].
A sketch of r.othcrmal lines within the Cube wall is shown in Figure J for an extnoc case in
wulch Cue cxteiual film coefficient was taken as zero (adlabaclc) over Che upper region of the cube.
Lvcn with this abrupt change in the film coefficient no incense thermal gradient is established. The
naximun thermal gradient is seen Co be well away from Che adiabatic zone and has as its eoper bound
Cite thermal gradient of the ail.pier axisyonptrlc case with a uniform heat transfer coefficient equal to
the local coefficient ac Che side of Che Cube.
Tiie effccc of a non-uniform excernal film coefficient, Chen, is Co produce stresses no greaccr
titan those of a uniform film coefficient, Che stresses for which can be obtained from Che analytical
onu-dicvnsional relations.
Tiic horizontal orientation of the tubes may produce anochcr effccc IndependenC of Che fluldized
bed -jeiuvior. Alchough horizontal Cubes are not usually used in conventional boilers, lltilte.! dac.i
indicates tiiat severe problems oay occur if a sufficiently hl^li quality, low flov rate two-phase mix-
Cure of water and steam flews within the Cube (1). Under these flow conditions gravitational scpara-
Ciun of cite phases can keep liquid from vcccinf* Che upper portions of the tube wall, leaving vapor with
its low thermal conductivity In contact wlch Che wall. If this vcrc to occur in a steady fashion, a
tuiaperacure distribution similar to that shown in Kip.urc 4 would result. Here substantial circumfer-
ential tu.-rnal gradiencs are established at the three pltase InCcrfacc. In realicy the position of this
interface would fluctuate wi .h tine causing large lcu-.il temperature excursions which will be discus-
sed in another section.
7i>c corresponding thermal stress in cite Cube cross section was found by a finite element numeri-
cal solution technique [4] assuming Che material was Isotroplc and llnc.irly clastic. Figure 5 shows
Che clicrmal stress pattern associaced with the temperature distribution of 1'lgurc 4. Gi.ovn hero It the
equivalent seres:; as calculated by e-jualiou 10. The largest equivalent stress occurs in the region of
greatest Ceuperacure. A large conprcsslve scress in the axial direction is the primary contributor to
Cite cquivalenc scress in Chis region. The equivalent stress throughout Che rcsc of the tul/e is consid-
erably sculler and suggests that Che upper region of the Cube experiences the most crucial conditions.
The use of equivalent stress, however, disguises Che local stress components. As .in example, tin:
stresses in the viclnicy of Che liquld-vapor-wall Interface are o.Q - 12.10U psi. o - 730 psl,
oz< - -6OIM psi while cite equivalent stress is only -10.QUO psl.
1C must be recognized that this situation has been greatly simplified, but that magnitudes of
these stresses clearly indicate che severity of the problem. The problems creatjd by flow stratifica-
tion are not likely Co be solved by material selection alone. Rather, measure* will have Co be taken
Co avoid operation in this particular two-phase flow regime. 1C should be no;ed in addicion tliac tliu
presence of 1HO* "ILairpln" bends, especially in the vertical plane, has been observed 15] co intensify
Che problem by causing Che Cube wall co dry ouC near Che exit of the bend.
704
-------
3000
Tangential
Radial
Axial
- Equivalent
-3000
0.82 0.85 0.90 0.9S 1.00
Radius (IN)
Figure 2. Thermal Stress Distribution Within
a Tube Wall for Axi-symmetric Conditions
ID = 1.64
OD=2.00
AdiabJtic
Figure 3. Temperature Distribution due to
Non-uniform External Heat Transfer Coefficient
ID-1.75 IN OD-2.50 IN
TBED = 1500ฐF
hป80
hIN
HRFT2ฐF
Figure 4. Steady Temperature Distribution
within Tube Wall due to Stratified Two-Phase
Flow Inside Tub*
Figures. Equivalent Stress for Tube due to
Thermal Conditions of Figure 4
(Stresses in KS1)
705
-------
High Cycle Fatigue
The possibility of high cycle fatigue due to the fluctuating nature of the heat transfer coef-
ficient oust be considered. The local coefficient at the side of a tube in a fluidized bed has been
observed to alternate bctwccr. high and lov values at a frequency of about 1 Hi. This is caused by the
frequent passage of a bubble over the tube surface; .Beeping away solid particles, leaving the region
temporarily bare and then replenishing the surface with hot particles. Large particle-: vhich will be
used in fluidized bed conhustors tend to produce fairly constant local heat fluxes during the period
of contact. The flln coefficient is reduced to nearly zero when gas alcne contacts the tube and thus
if gas and packet residence tines are roughly equal, the instantaneous film coefficients during eaul-
slon contact nay be roug!il\ tvice the time average flln coefficients at these locations. When radi-
ation if included, the rininun and avcrar.o coefficients will both increase.
The upper and lower Units of instantaneous tcnpcrature distributions are shown in Figure 6
where a regular cyclic variation of t'.c file coefficient occurs with equal gas and emulsion residence
periods. It can be seen that temperature only varies over a sm.i!l r.ingc. The amplitude of the fluctu-
ations Is greatest at tlic external tube surface and tlic amplitude decreases rapidly <itii penetration
into the tul>c Jail. Coupling this with the earlier observation that the inner surface is (n ten-ion
while tlic outer surface is in compression, it becooes clc.ir that the temperature fluctuations are
greatest where they arc lease likely to assist in Che propagation of a crack. The actual Instantaneous
temperature distributions within the tube wall fall within the envelopes shown in Figure 6 and have a
sinusoidal appearance with one quarter to three quarters of a wave within the oaterial. Instantaneous
thermal stress distributions (the circumferential component only) are shown in Figure 7 for four in-
stants at equal intervals within a 1 !lz cycle. The theroal stress fluctuations are not confined to the
externil surface, but they are small. Thus, high cycle fatigue caused by fluctuations of heat flux
on the outside of the tube does not appear to be a likely node of failure.
High frequency thermal fluctuations are also possible at the inner surface of the tube due to
alternate contact with liquid and vapor at a given location as was suggested earlier. A rough esti-
mate was made of the temperature fluctuations at inner and outer surfaces If the wetting and drying
occur within the range of periods observed by Lls and Strickland (5). Conditions and results are given
in Table II, and the limits of temperature excursions are shown In Figure 8. It can be seen that in
this case the temperature changes arc relatively large and located at a surface under tensile stress
where a crack is likely to r.row. Figure 9 shows the circumferential stress distribution across the tube
wall (or the thermal conditions of Figure 8. The tensile stress is seen to cycle over an appreciable-
range. Figure 'J is based on an internal coefficient which flu,-tunics between 230 and 2000 BTV/hr-ft *F.
For these conditions the Inside tube wall teenerature fluctuates a maxima, of 22*T; whereas, for the
conditions of Figure 8 the maximum temperature fluctuation ii }2ฐF. The results should only be vleved
as approxlnatc since the correct values of the cyclic heat transfer coefficient and frequency are a
function of the particular steam quality and flow rate and the tube size.
Low Cycle Fatigue
If the steady state operating conditions generate combined thermal and nechanlcal stresses which
locally exceed the elastic Mrolt of the material, micruvicldlng will occur on each start-etop cycle.
The cyclic mlcroyieldlng in the plastic strain range will ultimately lead to fatigue failure. The low
cycle fatigue behavior of type 304 stainless steel under fully reversed, axial strain-controlled con-
dition was Investigated by Cheng ct al 16). and the data arc reproduced in Figure 10. Though the tube
temperature normally does not exceed 700*F, fatigue data at 1000'F which are available can be used to
predict a lower bound of the tube lite. For the thick walled stainless Cube given In Table I the ther-
mal and mechanical stresses will produce a total strain amplitude of the order of 0.1X. As a conserva-
tive approximation, a total strain range equal to twice the total strain amplitude, i.e. 0.2". will be
assumed because there may be fully reversed yielding In some part of the fluidixed bed tubes. In Fig-
ure 10, It can be seen that at a total strain amplitude of 0.2Z, Nf approaclies iafinlty. Tims we can
safely conclude that Type 304 stainless steel will not crack by low cycle fatigue due to a start-stop
operation for the conditions given above. If a steep local teaperaturo gradient exists, such as that
shown in Figure ft, where substantial clrcucferentt.il thcr:jl gradients exist near the two-phase Inter-
face, the thenna.' stresses could possibly produce a large total strain range which could cause crack
Initiation during the life of tho tube.
CONCLUSION
Horizontal steam generating tubes In fluidized beds oay be subject to fluctuating and non-unl-
fortaly varying iioat transfer coefficients around their circumference. JJon-unlfcreitles on the exterior
of the tube do not contribute to increased steady state t her ml itrcss nor do they cause fatigue damage
since the outside tube surface Is in compression. The external fluctuations nay contribute to cpalling
706
-------
560
550
u7 540
530
520
510
500
Effect of fluctuating
external h
hMAX=8ฐ
ID = 1.64
OD=2.00
1 CPS
2 CPS
0.82 0.85
0.90 0.95
Radius (IN)
1.00
Figure 6. Temperature Variation within Tube
Wall due to Fluctuating Exterior Heat Transfer
Coefficient
6
5
4
3
2
1
_ 0
5
.82 .85
.90 .95
Radius (IN)
1.00
Figure 7. Circumferential Stress Dijiribution
for Conditiom Similar to Those of Figure 6.
707
-------
Table II. :Uxlnum TenperatBTe Chinees i.'ue to Fluctuating Internal Flln Coefficient
lntem.il,h
(BTWhr-ft'T)
MAX HIM
External. h
3300
3500
3500
3500
200
200
200
200
(STWhr-ft
60
60
80
80
.
'
Frequency
.2
.5
.5
Wall Thickness AT
(IN) <*F)
IH OUT
.375
.375
.18
.18
54.6 13.3
33.0 2.9
31.0 6.6
23.0 1.8
708
-------
620
GOO
ฃ 560
560
540
620
Effect of fluctuating
internal h
NlAX'3500 BTVJ
SlIN 20ฐ '
10 - 1.64
00-2.00
0.5 CPS .
1 CPS
2 CPS
0.82 O.K.
0.90 0.95
Radius (IN)
1.00
FignreS. Temperature, Variation within Tube
Wan dua to Fluctuating Heat Transfer
Caaff ieient on Inner Tuba Surface.
.82 .85 .90 .95 1.00
Radius (IN)
Figure 9. Circumferential Stress Distribution
for Conditions Similar to Those of Figure 8.
10'
10ฐ
~ i TIMIHI i i T Finn i iiiiiiiT r i niiin i Tiiin:
O (ANU * (BMI) 1000ฐF -
A & (ANL 1050ฐF I
DIANL) (BMI) 1202ฐF_
SR
^c
a10"1
*i**^i C
ป **>
i i i Mini i i i mill i I i Mini i STTfTtil i i
102
10' ?
10ฐ 's
102 103 104 10s
Cycles to failure (Nj)
106
10"
107
Figure 10. Low Cycle Fatigue Behavior of Type
304 Stainless Steel, from Reference (6).
709
-------
ot brittle films 0:1 the surface.
At certain flow conditions the two phase flcv within horizontal tubes will stratify with vapor
at the top of the tube. This will cause severe thermal stresses along the Inside of the tube. If the
wall Is alternately vapor blanketed anu then revet, fatigue failure may ensue. Uxact quantitative
determination of the tube stress and fatigue life is dependent on a detailed knowledge of the flow
behavior within this regime which has not been tlioroup.iily investigated to date. To avoid such problens,
the designer should be sure the internal flew Is outside the troublesome flow regime.
Low cycle fatigue failure Is unlikely to occur due to start-stop operations even for thick wall
stainless tubes which, of the tubes investi.-.ated, h.is ;i mxinmn stress r)o.:..-st to the yielding point.
JJUiiUiCLATUiii:
L - Young's modulus psl
h - film coefficient of heat transfer BTU/hr-fc *F
k - thermal conductivity inu/iir-ft*F
:i, - numLer of cycles to failure
r - radius in
T - temperature *F
TULD - temperature if fluldlzed bed *F
T - saturation temperature of steam *l
AT - temperature difference across tube wall ฐ1ฐ
T .... - reference temperature at which thermal strcj: js are zero ฐF
M.r
a - linear coefficient of tiirrn->l expansion 1/*F
c - strain in/in
t.t plastic strain range in/in
v - Pois,son's ratio
0 - I lt*>n:M 1 strc-SK psl
~^r
o - equivalent ttn-rni.-il sir*-ss psl
Subscripts
00 ' circumferential direction
rr * radial direction
zz - axial direction
i - inside surface
o - outsi.-ie surface
KEFEREJICLS
1. Cclperin, N. 1., and Einstein, V. C. in FluldlzJtIon, ed. Davidson and Harrison, Academic Press,
Hew York, 1971, p. 510.
2. Hanson, S. S., Thermal Stress and Low Cycle Fatigue. McGraw-Hill Co., New York, 1966, pp. 27-33.
3. Styrikovich, M. A. and Iliropol'skii, Z. L.. llydrodynamlc and Heat Transfer During Boiling in High
Pressure Boilers, AEC-tr-44'JO, June 1961, pp. 244-272.
710
-------
4. Uathc, K..AUINA. Import 82448-1, Acoustics and Vibration Laboratory, rieciianlcal Lnginccrlng Depr.rt-
uent, :ilT, Caooridge, ::A, :uy, 1976.
j. Lis, J. and Strickland. J. S., 1Q70 International llcat Transfer Conference, Paris. August 1970.
6. Cheng, C. F., et al. , Low-Cycle Fatigue Eeh.-ivior of Type 3u4 and 316 Stainless Steel at LMFBK
Operating Tcc-peraturc, AST.I STP 520. 1973, pp. 35S-J64.
711
-------
QUESTIONS/RESPONSES/COMMENTS
STANLEY OAFKIINAS, CHAIRMAN: We have one more question that has
been submitted by Al Leon of norr Oliver for Dr. Glicksman, and his
question is: "What tube life do you predict?" Do you want to step
to a nicrophone in the rear of the roon?
OR. OLICK.SMAN: f'n going to have to pass on that one, and ask
that you go hack and talk to our metallurgist friends about what type
of tube life one would predict. In the thermal stress problems with
which we have concerned ourselves, the najor problem is, as I said,
what's going on inside the tube. And this depends on a little bit
better clarification on the flow regimes in there. We just don't
know, at this stage of the game, what kind of problems one will
have.
712
-------
INTRODUCTION
STANFLY DAPKUNAS, CHAIRMAN: Russ McCarron is going to present a
paper on Turbine Materials Corrosion in the Coal-Fired Combined
Cycle. Russ received his BS from Penn State and his MS and Ph.D.
fron the University of Pennsylvania, and he is presently manager of
fossil energy material for the Energy System Program Department of
the General Flectric Company.
713
-------
Turbine Materials Corrosion in the
Coal-Fired Combined Cycle
R.L. McCarron, A.M. Beltran,
H.S. Spacil and K.L. Luthra
General Electric Company
S:henectady. New York
ABSTRACT
The alkali (tla + K) in the vapcr from a Pressurized Fluidized Bed Combusto- (PFBC)
may be up to two orders of magnitude greater- than the acceptable limit based upon present
gas turbine liquid fuel experience. The corrosion conditions developed in the PFBC and
gas turbine are analyzed and compared to classic hot corrosion produced by petroleum
fuel combustion. Available literature regarding materials performance in PFBC environ-
ments is also reviowtd, and the prolable magnitude of the corrosion problem assessed.
Cladding of superal loys with new hot corrosion-resistant materials is a key elc-ment in
the solution to the corrosion problec which could result if conventional turbine alloys
and/or coatings were exposed to this environment. The development of new cladding alloy
compositions to resist corrosion fro-s the direct combustion of coal and preliminary re-
sults of corrosion evaluations of these alloys will be reviewed.
INTRODUCTION
The General Electric concept of the Coal-Fired Combined Cycle (CFCC) includes
a Pressurized Fluidizod Bed Combustor (PFBC) which is cooled through the use of steam
tubes in the bed which supply a stean turbine generator. Limestone or dolomite is ;d
to reduce sulfur emissions. The partially cooled combustion pases exiting from the com-
bustor drive a gr>s turbine after pas; ir.g through a hot-gas cleanup train. The low steam
tube temperatures in the bed 'about 1000 F) will significantly reduce the potential cor-
rosion problem for ht-at transfer tubes located in the bed. The major materials problem
to be overcome is the potential for corrosion/orosion of the hot section parts in the
gas turbine duo to carryover of contoninants in the combustion products from the F'FBC.
Thir. ;.:-oblem will be common -.o any concept which ii.eludes a gas turbine in line with
a i Kb".
A key part of the General Electric (GE) CFCC Development Program funded by the
Department of Energy (DOE) is alloy and process development for tr.e cladding of turbine
blades. In the clad approach, a thin sheet (10 mils) of a corrosion-resistant alloy
is bonded to the airfoil surface of a strong superalloy blade. This is tl-.o only work
of its kind devoted exclusively to th< development of turbine materials for application
in the PFBC gas environment. Blade ceding alone cannot be counted upon to solve the
potential corrosion problem because ti'e contaminants in goal can produce condensed specie;
which can cause corrosion at temporaljres as low as 1100 F, and even after cleanup the
gas may contain enough particulatc matter to plug film cooling holes.
An important part of the clad alloy development is corrosion t?st;r.g and evalua-
tion of candidate alloys. General Electric is conducting corrosion tests of advanced
clad alloys in small burner corrosion rig.> using a synthesized environment and is pre-
paring to test the same materials as airfoil specimens in cascades in the Exxon Kiniplant
PFBC and in the Coal Utilization Research Laboratory (CURL) ?. foot x 3 foot PFBC at
Leatherhead, England. The emphasis at the present stage of the program has been en corro-
sion evaluation of the illoys. Some erosion data may be forthcoming from the cascade
tests mentioned above, but future tests of rotating hardware in the effluent from a PFBC
are required to assess erosion resistance properly.
ALKALI CARRYOVER AND IT3 EFFECTS OH GAS TURBINE HOT SECTION PARTS
Cas Turbine Experience with Al kal i-Cor.tami nated Petroleum Fuels
In a gas turbine the heart of the machine is the turbine section, especially the
first stage buckets. Its integrity has more influence on machine output, efficiency,
This work is sponsored by The Department of Energy under Contract No. EX-76-C-01-2357.
714
-------
and the capability to burn a range of fuels than any other component. The integrity
of the first-stage bucket is largely dictated by mechanical and corrosion limitations.
Corrosion limitations are determined by sulfur and the trace metal contaminants in tlie
combustion products. These come from the fuel or the environment in which the machine
operates.
In petroleum-fired gas turbines the trace metals of most concern are sodium (Ma),
potassium (K), vanadium (V), lead (Pb), and calcium (Ca). If they are present in the
combustion products in significant amounts, the first four can cause turbine tlading
corrosion while all five can cause fouling due to deposits.
Although all five eler.ents are critical, sodium and vanadium generally are the
two most frequently found in petroleum fuels. In coal fuels the two critical elements
will be the alkalis, sodium and potassium. A necessary condition for the catastrophic
attack known as "hot corrosion" to occur is that the combustion products are super-sat-
urated with alkali chloride, hydroxide, and/or sulfate at the temperature which prevails
in the vicinity of the first-stage buckets. If this condition is fulfilled then conden-
sation of the alkali salt and reaction with SO-, and 0_ will cause liquid alkali sulfate
to form on the metal surfaces. Following condlr.sation and formation of the alkali sulfate,
the subsequent steps in the corrosion process are as follows (for the case of 113,30^).
1. The con-lensed KapSO-j reacts with the protective oxide film on hot gas path
parts to form a fiouble oxide such aj Ka,CrO..
2. Some of the NapSO.. is reduced to a sulffde, so that the chemical potential
of S is dramatically increased.
3. Sulfur in the forr. of sulfide then penetrates into the metal forming metal
sulfides, particularly Cr S .
<*. The depletion of the allojf Xf Cr by Three above interferes with formation
of a new protective oxide film.
5. Continuing dissolution of the protective oxide allows the entire hot corrosion
process to accelerate.
6. Lead and vanadium in hot gas streams cause an analogous type of attack. They
dissolve (or flux) the protective oxides to a greater extent than :;a,SOy alone.
Cennr.il Electric has conducted a significant amount of research into the relation-
ship between trace netal contaminants and bucket life. The result of this research has
been the formulation of a proprietary system capable of predicting the effect of trace
metal contaminants on hot-gas-path parts lives.
The basis of this corrosion lives system is a correlation between measurements
on installed gas turbines and data from long-time, laboratory corrosion tests. The
correlation involved measurements on over 100 commercial machines, some of which had
service times of about 100,000 hours. The laboratory tests were conducted ir. a srr.all
burner rig facility which has logged millions of specimen hours since the late 19^0s.
Although the correlation itself is proprietary, an exair.ple of its use is shown
for the GE HS-5001 Heavy Duty Gas Turbine in Figure 1. Here, the effect of cr.e cont.im-
inant (sodium) on first-stage bucket corrosion life is shown. The contaminant is ex-
pressed in terms of equivalent sodium in the fuel, even though it cculd come from fuel,
inlet air, or water/steam injection.
The results in Figure 1 show that for 25,000 hours life, the specification limit
for total equivalent (la in the fuel is 1.2 ppm. The 1.2 ppm sodium in the f-jel translates
to 0.02*4 ppm Ka in the va^or phase at an air/fuel ratio of 50.
Two major points to be observed from Figure 1 are:
1. Tne strong effect that I.'a in the range 0.5 to 1.0 ppm ir. the fuel has upon
bucket corrosion life for the petroleum-fired gas turbine.
2. The significant effect that materials selection can have upon bucxet corrc-
sion life. At a value of 1 ppm sodium in the fuel, for example, it is expected
that 111-738 will provide five times the corrosion life of U-700.
715
-------
100
80
LIFE
40
20
0
738
SPEC LIMIT
0.5 1,0 1,5
EQUIVALENT SODIUM IN FUEL, PPM
Figure 1. Sodium Effect on Bucket Life
2,0
716
-------
Alkali Efflux from the PFBC
Trace elements in the coal which will cause corrosion of hot section parts include
sodium, potassium,, sulfur, and chlorine. Sulfur in a PFBC cannot be reduced to levels
low enough to inhibit alkali sulfate formation even though EPA standards on sulfur emis-
sions can be met. The sulfate is the means by which sulfur is transferred to the turbine
alloys. Chloride,which will be present in the corbustior. products of coal and resultant
deposits, accelerates the corrosive attack of the alkali sulfates.
Alkali metals which are introduced into the tied with the coal and dolomite will
be transported through the system in fine particulate ash and as vapor species which
are in equilibrium with solid alkali metal-containing compounds in the ced or in the
hot-gas particulate cleanup equipment. Because of the gas/solid equilibria which will
be established in the system, it may not be possible to prevent unacceptably high levels
of alkali metals from entering the gas turbine even if all the particulate matter is
removed from the gas stream.
A model has been proposed by two of us (H.S.S. and K.L.L.) which gives a first ap-
proximation of the level of alkali metal present in the efflux (vapor phase) from a fluid
bed combustor. Available thermodynamic data on the relevant compounds were used for
the necessary calculations. The coal ash is arsuned to be siiica with about 2% sodium
plus potassium in the ash. Typical hydrogen, sulfur, and chlorine contents are assumed
for the coal, with appropriate sulfur removal by a sorbent. At the bed temperature,
sodium plus potassium (Ha + K = M) is present in the system either as a liquid K-O-nSiO,,-
Si02 solution which coexists with solid SiO,, or as a liquid M-SO^-CaSO^ nolutiofi. Small
amounts of alkali metal can volatilize from either type of solution, however, according
to the reactions
M20-n?i02(t.) * HCKg) < MCl(g) + MOH(g) * nSi02(s) (1)
or
r^SOjjU) ป HCKg) < MCl(g) ป MOH(g) * SO^g) (2)
followed by the reaction
MOH(g) * HCKg) * KCKg) * H20(g)
which converts almost all of the XOH to MCI. Equilibria of this sort are represented
in Figure 2. Here a correction has been made to account for the increased stability
of alkali metal aluminosilicates relative to silicates, even though the presence of
alunina .is such was not considered specifically.
The two curves in Figure 2 represent equilibrium saturation of the vapor phase
with sodium and potassium for combustion of coals with 0.015 Cl and 0.1* Cl. Many United
States coals have chlorine contents of O.if or greater, ar.d the chlorine tends to drive
the alkali metals into the vapor phase. Temperature and the chlorine content of the
coal are the most important factors in determining the level of alkali retals in the
combustion products. For example, at a combustor operating temperature of 1750 F and
with chlorine in the feedstock at 0.IS, the sodium and potassium in the combustor products
will approach 10 ppm. Data from the CURL Pressurized Fluidized Bed at Leatherhead, England
confirm this magnitude of vapor phase alkali content as shown in Table I. These data
were obtained by passing an isokinetically drawn sample of the flue gas through a high-
efficiency cyclone and extracting the positively ionized alkali metal vapors species
on a collecting electrode in an electric field. The dashed vertical line in Figure 2
specifies the regions of stability for sulfate alone and sulfate plus silicate together.
Below about 1715 F the aluminosi1icate solutions are not stable for ccnditicns given
In Figure 2, and one would not expect effective gettering of the alkali metals in such
soluti ~ins.
As the hot combustion products from the PFBC pass into the turbine, the gas temper-
ature decreases, and the gas comes in contact with cooled octal parts (nozzles and buckets).
Cooling of the gas which is saturated with alkali chloride causes the alkali chlorides
to condense and react with S02 and 02 to form f^SO^.
717
-------
10
700 750 300 350 900 950 CC
.1
ALKALI VAPOR
CONCENTRATION
(PPM)
.01
.001 -
0.1 ZCI
0001
/
/
f i 1
1200 WOO
i 1
1600
1
1
1
1
ii l
180
GAS TEMPERATURE
Figuc* 2. Alkali Vapor Concentration
-------
2MC1 (g) * S0? (g) * 1/2 0? (g)* H?0(g);H2SOl4 (1) * 2 HC1 (g) (3)
The condensed M2^l4 '3 underlined since in this case it may exist in a liquid solution
with calcium sulfate.
Calculations have been carried out to estimate the flux of alkali netal sulfate
condensate to the surface of a first-stage turbine bucket for a turbine burning petroleum
fuel and for a turbine in line with a PFEC. It was assumed for the petrcleum-fired tur-
bine that the sodiua level in the fuel was 1.5 ppm (slightly greater than the maximum
1.2 ppm allowed according to the present liquid fuel specification), and the metal teซ-
perature was 1600 F. Fcp the Coal-Fired Gas Turbine it was assumed that the bed operat-
ing temperature was 1750'F, and the chlorine content of the coal was between 0.01 and
0.1%. A comparison of the results r.t these calculations shows that the f'.ux of alkali
metal sulfate condensate to the bucket of the Coal-Fired Turbine is 30 to 300 tines
greater (corresponding to 0.01 and 0.1J CD than for the Petroleum-Fired Turbine. If
the PFBC operating temperature is reduced from 1750 to 1650 F, then the alkali metal
sulfate flux is approximately 10 to 100 times greater than the Petroleum-Fired Turbine
with 1.5 ppm IIa in the fuel.
From these results it is proposed that for the PFBC/Gas Turbine the net flux of
alkali sulfate (proportlens! to Ha + K in the combustion products) will saturate the
nozzle and bucket surfaces such that the overall hot corrosion attack is reaction-rate
limited, while for a conventional Petroleum-Fired Gas Turbine the flux of alxali metal
sulfate is limited by present fuel specifications such that overall hot-corrosion attack
is flux limited. Qualitative plots of part-life versus alkali metal sulfate flux are
shown in Figure 3 for two temperatures. Conventional gss turbine experience includes
the oxidation and flux-limited regimes as illustrated in Figure 3. Note there is a
reversal in the temperature effect on corrosion at low alkali metal contaminant levels.
This is caused by the increased condensate flux at the lower temperature and has been
documented by field observations. As the alkali metal level in the combustion products
increases, the corrosion reaction is limited by the inherent reaction kinetics between
condensate and the alloy (i.e., further Increases in the alkali netal sulfate flux will
not cause a corresponding increase in the corrosion rate), and the predicted part-life
for an allowable metal loss appi caches a minimum. As shown In Figure 3. the PFEC/Cas
Turbine combination is expected to operate in the reaction rate-limited regime. The
bucket surface will be completely covered with a iayer of condensate and can be described
a^ "saturated." As will te described in a later section, this situation will result
in unacceptably short lives for conventional gas turbine nozzle and bucket alloys.
The presence of high levels of potassium and chlorine In the combustion products
from a PFBC creates the potential for further acceleration of corrosion on gas turbine
parts over and above that due to the high-alkali flux. Recent small burner rig tests
conducted by the General Electric Gas Turbine Division have shown that for up to at
least^^O mole % K-SCK in Ha^SO^ there Is a synergistic effect of Na and K on corrosion
rate. That is, K does not Simply substitute for tla in the corrosion phenomena, but sig-
nificantly increases the rate of corrosion as It is substituted for Na in the deposit.
Potassium has not been a real problem for gas turbines burning liquid fuels because the
alkali comes in with the fuel or air mainly as sea salt which is primarily NaCl with
only a small fraction of KC1. The potassium level in coal, however, will for the most
part be as great ->r greater than sodium so that increased corrosion rates may be antici-
pated from the combined effect of sodium and potassium when compared to sodium alone.
One of the important corrosive contaminants in coal is chlorine. Unlike a con-
ventional gas turbine where any chloride salts in a condensate are completely converted
to sulfates the high chlorine/alkali metal ratio of the PFBC will cause the alkali netal
chlorides to be incorporated partially into the deposit. Chlorides have been identified
in deposits from a PFBC. Alkali chlorides are important in hot corrosion in that not
only do 'hey increase the corrosion rate in combination with sulfates as demonstrated
by crucible ti:^ts, but they hawe oeen reported to attack cobalt base alloys more vigor-
ously than nickel base alleys. In conventional gas turbines where condensed chlorides
are generally not present in deposits, the cobalt base alloys are the most resistant
materials to corrosion attack.
719
-------
OXIDATIO'l T
1
' ELIJX-LIHUED
! (FIELD DATA)
OXIDATION
LIFE
REVERSAL
^CaWEHTKXIAL GT GAS
AND OIL EXPERIENCE
REACTION RATE LIHITED
(LABORATORY DATA)
PFB/GAS TURBINE
MA, K CONCENTRATION III COMBUSTION PRODUCTS
Figure 3. Life Vป Afluli Conctntietioa for Hovy Duty
Gซs Tufttin* (Schematic)
720
-------
PRIOR EXPERIENCE
During the last several years tests have been run in which gas turtir.o hot-section
alloys have teen exposed to the products of co-bustion of a PFBC or Atr-ospr.-rr ic .- lvii
-------
,*OOA)
722
-------
of three rones and was fabricated fron AI3I 310SS tube which was n3r.ir.3lly 3 inches
I.D. Turtir.e rr.nterials specimens were 3-inch by 1/1-ir.cc diameter pins which
were inserted through the wall of the test section at 90 to the gas flow.
Flow through the section war atout 100 feet per second. The list cf materials
is givf:r. :n Table I!. Two tests totaling 2000 hours of operation were completed.
Test ecr;Jit:?.ns for- both tests were nominally the same. The bed temperature
was 1fir.OwC, coal was Illinois S6, and scrbent was limestone. The temperature
in Zone j of the turbine lest section was K80WF, and the temperature in Zone 3
was 1320 F. All cf tl.o alloys were exposed continuously for 2000 hours. Most
cf the alleys wer- exposed in all three zones of the test section. Duplicate
samples of three of the nlloys, IN-73S, X-^O, and IN-713C, were rer.oved after
the first 10CO h^urs and replaced with new samples for the second 1000-hour
test.
After both 1000-hour tests, there was a soft deposit on the tube wal" and lead-
ing, and trailing edge deposits on the specimen pins. Dust buildup w^s a maximum of 15 mm
in some locations. The deposit was light and fluffy and easily removed. The alkali
content cf the c;as was measured downstream of the second ?yclone and before the turbine
test section. The ccasurement technique was similar to that used in the PFBC tests at
Leatherhead. The technique is not well-developed, and there could be some error. The
codiura content of '_ht f.as was 1.W ppm by weight, and the K level was 0.5 pptr.. These
levels are comparable to the alkali content measured in the earlier tests at Leatherhead
(Table I).
The turbine alloy specimens were returned to the Gas Turbine Division of General
Electric for post-f :.. tallurgical evaluation. The significant results are ar, fol-
lows:5
1. Maximum ccrrcsior, attack on most materials in all three zones was 1 to 2 mils
in 2010 hcurs.
2. The test conditions in the turbine lest section, primarily temperature, were
not severe enough to permit a clear-cut distinction between the oxid.it ion/cor-
rosion be'avior of r.o:jt alloys.
3. GE 25ซ1 (FeCrAlY) showed the least attack. The moat extensive attack was
on I'i 671 (Hi-50Cr) which showed oeep localized sulfidation pitting.
i). The oxidation/corrosion attack of about 9 mils in 2000 hours on the above-
bud coupons (^ alloy.-i) was core extensive thgn on similar pin specir.er.s.
This was due to the higher temperature (1650 F) and different environment
above the ted.
CPC Investigations
For the past three years, Combustion Power Company (CPC) has been carrying cut
a progran of testing in a Four-Atmosphere Fluidized Bed Combustor coupled to a Direct-
Fired Turbine, and a smaller 2.2 square foot atmospheric model Fluid Bed Corbustor; since
March 1975 this testing has been directed tcwa d the accumulation of data with regard
to hot corrosion in the ccal-combustion process and the development of methods for dealing
with it.
Materials evaluations have been carried out In the One-Atcosphere Model Combustor.
The Four-Atmosphere Combustor has accumulated about 600 hours burning coal; prior to
this, garbage and wood waste were burnt. Materials evaluations in the PFBC were about
to begin when the granular bed filter in the system failed in December 1975. The inlet
temperature of the turbine, however, is limited to 1^50 F, so that materials evaluations
in the turbine would be of marginal value.
Material specimens are hung in the freeboard of the Atmospheric Combustor. Exposure
temperatures have varied between 1600 and 1700 F, but included some excursions to tenpera-
tures as high as 1760 F. Testing to date has teen conducted with Illinois ป6 coal and
Kaiser dolomite. Alloys tested are listed in Table III. Crude attempts at measuring
the alkali content of the gas have yielded an estimate of 2 oprn Na in the gas in the
freeboard.
723
-------
Table II.
Turbine Alloys Tested in the BCURA AFBC in
Stoke Orchard, England
Nickel Base
U-500
U-700
IN-738
CTD-111
IHCO 713C
Hast X
Inconnel 617
Cobalt Base
X-UO
FSX-41H
HS-188
Protective Coatings
IN-738/Pt-Al
IN-738/BC-29E
FSX-im/Pt-Al
IN-738/BC-21P
GE 2541 Clad
Inconel 671 Clad
724
-------
Table III. Alloys Tests in AFBC at CPC, Kcnlo Park, California
Alloy
SS 301)
SS 309
SS 310
SS 316
SS 310
SS 333
SS UU6
Inconel 600
Inconel 601
Inconel 706
NI55
I.ncoloy 800
Incoloy 825
Nimonlc 80A
Rene 77
tnco 713C
U-700
IN-738
HS 31
FS 14 11
M 509
PWA 68
Probable
Use
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Nonturbine
Stator
Nonturbine
Nonturbine
Stator
Rotor
Rotor
Stator
Turbine
Turbine
Turbine
Turbine
Coating
Ni
10
114
20
ID
35
148
0.7
76
61
12
20
32.5
1414
76
59
7U
55
6 'I
10
10
10
--
Co
--
3-1
20
--
15.0
17.0
8.5
56
55
58
Bal
Cr
20.0
23.0
25.0
18.0
19.0
25.0
25.0
15.5
23.0
16.0
21.0
21.0
21.5
19.5
11). 0
12.5
15.0
16.0
25.0
29.0
23.5
19-25
Fe
Bal
Bal
Bal
Bal
Bal
18.0
714.0
8.0
114. 1
uo.o
33.0
145.0
30.0
--
.-
i.O
--
Composition
Al Ti
--
1.35
0.20
--
0.38
0. 10
1.30
14.30
6. 10
14.00
3. no
--
12-15
1.75
--
0.38
0.90
2.50
3-35
0.80
3.50
3.nc
0.2
--
Ko
2.5
3.0
3.0
3.0
14.?
n. 1
5.0
1.8
--
--
W Cb
3.0
2.5
2.0
2.~6 II
7.5 1.0
7.5
7.0
..
Zr C
0.05
0.05
0.05
0.05
0.05
0.05
0.10
0.08
0.05
0.03
0.15
o.or.
0.03
0.06
O.OH 0.07
0.10 0.12
0.06
0.17
0.50
0.25
0.5
..
ro
VI
-------
After 3bout. JrOOO r.r.urs of materials testir.g in the One-Atmosphere Ccr.tiusf.or very
little corrosion has Lion ob-erved.6 Th-j results dc net distinguish between JO'iSj and
111-736 or I1I-713C 'lorro.- ion resistance of IK-7-S is vastly superior to 30U3S or It.'-713C
in conventional ho'.-ccrres i or. tests). A series of preliminary snort-duration tests showed
that significant, hot-corrosion attack aid occur if the temperature in the freeboard ex-
ceeded &Lout :70r;"r. The CPC tests show S'..r.e ovidr-rce that the addition of an aiunino-
silioate additive- to the bod suppresses the relea.-.o of alkali sulfate with a redaction
in the potential for hot corrosion. Krosicn of material specimens was not observed i r.
the freeboard of ปr.e AMnoah^ric i:cmbustor .low gas velocity).
The main conclusion from the CPC work to dato in that, tonperaturo i.-. the ?ingiซ
no"t important factc" in -Jetern; ininf; whethoi- or not hotr corrosion will occur in an Atr.o-
r.pher ic Coa 1-Cor.Lur.tor Cystc-K. At temperature- o-f 'fjOO'F hot corrosion did not occur,
while temperature excursions above 1700JF produced significant corrosion attack.
Summary of Prior Experience
The most rr.eaninpful data obtained so fa>- aro from the PFEC t^sts at Leatherhead,
England. The rea.-on ic that conditions of tr-p'-rature, pressure, and Ras velocity were
the closest to what :s expected in a real r^stem although there were several c^-rrperature
excursions due to coal feed problems, and these excursions could have contribu'.ed to
the observed attack. As previously indicated, hot corrosion was observed in these tests
after only ?00 hours. Also the level of alkali analyzed in the gas was in the ranp.e
of what is expected based upon thermodynamic estimates.
MATERIALS DEVEI.OrKF.IJT FOR THE COAL-FIRED COUBTI.'hD CYCLE
The current state-of-the-art for particulate removal and alkali vapor inhibition
or removal is such that major breakthroughs will be required to achieve acceptably low
levels of alkali netal contamination. Therefore, the achievement of acceptable turbin'.
blade parts liver, in the expected alkali environment will require at a minimum, substantial
improvem-jiit in corros i on-res i -.tant alloys and hot gas cleanup technology. Because of
the advanced devel opr.eiit cha'ienpe in .ill possible solution."! to the ccrros-ior. problem,
it is likely that a combination of advancements in each technology will be required to
achieve an acceptable system. In this section present work on a materials development
approach will bo reviewrrl as n solution to the corrosion problem.
Materials development in the present DOi-I sponsored General Electric CFCC Develop-
ment Program i:-. directed towards development of sheet claddings for use on gas turbine
buckets anu nozzles that will operate in the environment generated by a Pressurized Fluid-
I zed Bed Combustor. A schematic of the cladding process is shotin in Fipure 5. The overall
philosophy of the technical approach involves a systematic combined study of cladding
chemistry, procers variables, and resultant properties which will lead to a significant
improvement ovซ>r existing surface protection schemes for use in PFBC-type environments.
This program, which is an extension of previous cladding studies at General Electric
is divided into tasks as shown below:
Task 3.1 - Cladding Alloy Development
Task 3.2 - Cladding Process Developaent
Cladding Alloy Development
Th? primary objectives of this task are to:
1. Select and characterize clad alloy coirpositions with expected superior corro-
sion resistance in the fluidized bed combustion environment;
2. Assess the corrosion capability of the best candidate claddings;
3. Characterize the clad substrate interactions as a function of exposure con-
ditions.
726
-------
CORROSION RESISTANT
SHEET ENVELOPE
CREEP-RESISTANT
AIRFOIL BODY
CREEP AND CORROSION RESISTANT
CLAD TURBINE BUCKET
Figure 5. Sketch Illustrating Cladding Concept for Corrosion-Resistant
Gas Turbine Bucket
727
-------
Three reference cladding alloys have been -jred as controls to form the basis for
the development of new alloys. The composition of these alloys are listed fcelow:
Weight J
Alloy Fe m_ Co Cr A_l_ Y Other
OE-2511 Bal - - 25.0 U.O 1.0
Inconel 671 - Bal - 50.0 -
S-57 - 10.0 Bal 25.0 3.0 0.15 5.0 Ta
Modifications of these alloys were selected with Cr an-i Al contents as m^jor
variables since these elements control the type of protective oxide which will form.
Rare earth element additions were made to selected compositions to improve scale adhesion.
Some of the alloys were formulated for sheet fabrication by conventional cast
and wrought metalwcrking practices and, due to anticipated processing probless some of
t!ie alloys, were fabricated from more readily processed pre-alloyed power. Initially
about 15 clad alloy compositions were identified. In an attempt to increase the Al con-
centration of certain of the clad alloys, without reducing fabricability, they were
aluminided by the standard pack cementation process. Finally 11 of the 15 original
alloys were processed to 10-mil sheet which was suitable for the cladding of corrosion
specimens. Compositions of nine of these alloys are shown in Table IV. The two alloys
which are not shown are presently under review Tor possible patent action.
IN-738 (Hi-base) and FSX-114 (Co-base) have been utilized as substrate materials
since these alloys are current gas turbine bucket and nozzle alloys. Prisary emphasis
has been placed on IN-738. Small burner rig corrosion disks of these alloys have been
machined and fully clad with the candidate sheet clad alloys. The specimens were pre-
pared by diffusion-bonding the cladding to the corroson disk using the reference glass
Hot Isostatic Pressure (HIP) process. This process utilizes eolten glass as the pres-
sure transfer media while in an evacuated, sealed container. Hot Isostatic Pressure
Diffusion Bonding conditions were 2100 F, 15,000 psi gas pressure for a one r.our hold
time. Figure 6 is a Schematic of the Hot Isostatic Pressure Autoclave. . Selected alloys
were pack aluminided by established techniques. This does not show in Table IV. The
small burner rig test is being used to screen the elaa opecicens. A schernau-.- of the
small burner rig is shown in Figure 7. Test conditions are givon in Table V. To date.
Test Number One shown in TaMe V has been used. This test environment includes the
alkali contaminants expected from coal, sodium, and potassium at levels which will cause
a saturated layer of alkali sulfate condensate at one atmosphere on the specisen sur-
face. As discussed earlier, this is the expected condition for the PFBC/Gas Turbine
combination. Test Number Two shown in Table V is still under development. This test
will include levels of chlorine which closely simulate those expected from th<ป combus-
tion of coal. An .nitial 1500-hour test has been completed in the small burner rigs,
and a long-time tes\. to develop data as a function of exposure to approximately 5000
hours is now under woy. Following exposure in the small burner rig, the disc-shaped
specimens are metallographically prepared and measured for maximum depth of penetration
as illustrated in Figure 8.
Based upon the 1500-hour Small Burner Rig Corrosion Teat, five claddings have
been selected for testing in the EXXON Hiniplant PFBC facility. A test section which
is a series of four stationary cascades designed to simulate gas turbine conditions
has been fabricated by General Electric and installed in the Exxon Miniplant as part
of the DOE Fireside Corrosion Task II Program. A schematic diagram of the test section
is shown in Figure 9. A summary of the Miniplant test conditions is given in Table VI.
Airfoil-shaped specimens of impulse and reaction design have been procured for the test-
Ing. The impulse-type simulate the gas turbine bucket, and the reaction-type simulate
the gas turbine nozzle (Figure 10). The airfoil specimens are now being clad with the
selected claddings. A similar test section is being built for installation in the CURL,
?. foot by 3 foot PFBC in Leatherhead, England. Clod airfoil specimens will aiso be ex-
posed in this facility. It is anticipated that exposure of the airfoil specimens in
the Exxon Klniplant and the CURL PFBC at Leatherhead, in addition to corrosion data,
will provide an assessment of the potential erosion problem.
728
-------
YOKE FRAME)
MOVEMENT
FURNACE WNNN6S
MOSMCERS
HYPR&UUC
CYL1MK8S
Figure 6. Schematic of Hot bostatie Pressure Autocta*
729
-------
Table IV. CFCC - Clad Alloy Developaent
Alloy
CE-2541
GE-25
-------
CROSS
0 f- SECTION
LINE
EXPOSED
TEST SPECIMEN
AFTER TEST
MET ALLOGRAPH ICALLY
MOUNTED SPECIMEN
AFTER TEST
IMACE OF MOUNTED SPECIMEN
IN OPTICAL COMPARATOR
READ
WtTAL PENETRATION
*CR SIDE (MILS)
METALLOGRAPHIC
MOUNTING
MATERIAL
SET CURSOR IN
THICKNESS BEFORE TEST
BO MILS
FiguraS. Schematic of Small Burnt* Tซt Specimen! and
Mซanxl of Meultographially Measuring Depth of Common
731
-------
Table V. Small Burner Rig Corrosion Screening Test
Temperature - 1600ฐF
Pressure - Atmosphere
Velocity - 70 fps
Fuel - ป2 Distillate
Cor.taninants - 1ปS, Na, K Added to Fuel HC1 Added to Mr
Deration - 1500 Hours for Selection of Materials To Be
Tested in the Exxon Kiniplant
7000 Hours Long Time Test of Selected Materials
Test ป1 -79 ppm Na, 112 ppm K in Fuel
Test 12 - 79 ppm Na, 112 ppm K in Fuel Plus 15 ppm
HC1 in Air
18 IMPULSE AIRFOILS
5 ATM.
I550ฐF
955 fps
0.83 Lbs/SEC.
REACTION AIRFOILS
EXIT
2.9 ATM.
J550ฐF
1910 f pi
Figured. Turbine Ten Section
732
-------
Table VI. Test Conditions Exxon Miniplant
Pressure: 8.7 - 9.1 Atmospheres Absolute
Temperature - In Combus'-or: 1700ฐF - 17bOฐF
Temperature - At Turbine Test Section Inlet:
155CTF - 1570ฐF
Gas Mass Flow Rate: 0.7? I/sec - 0.92 */sec
/ be Hec
Section
Combust ion of Methane may be Hecos.sary to Assure
1S50JF at Turbine Teat ฃ
Up to 8 Clad Alloys
Time - 1000 Hours
Figur* 10. Airfoil Corrosion Specitnmi (or Turbine Tot Section
733
-------
Cladding Pro .ess Devi-1 oprr,er;t
The cladding process ceveiopr.ent effort is ar. if>tซ-gral part of the overall effort
to develop advanced ciaddir.? r.ateria;s for protection of "*-.:; 7-jrbi:'.- Hot-^ect i en par'.s.
Th-.-r.e studies recognize the ir.r.-.-rer.t physi'.--jl and -^-hani-:%j pr>.perty o-'.aricter i st ics
of the individual cladding alloys -jr.-o- tne ir.portanc'; of opt Jr.; z ing tr.o total clad/sus-
strate system for turbine bucket appj icatior.r-.
Presentiy, thtrr. are five key activities under cisl'ir.-g process development.
These are:
1. Clad sheet forming
?. ourface preparation technj^L :s
"J. Optimization of diffusion -ondir.p parameters
'J. Total bucket cladding (tip. platforn, airfoil)
Since an adequate treatment of the cladding prooens Jevelopr.ent effort is legiti-
mately the subject of another complete paper, it will not be discussed further here.
Corrosion Results
Corrosion test results to -.'ate for the cladding alloys consist of those from the
1500-hour test at IfcOO'F in the ssalj burner rig. Trteso are ::r;own in 7aL!e VII for r.o.T.e
but. not all of tho cladding alloys. At the bottorr. of Table VII i.-. a result for hare
111-73*} in the same test. It is otviojs fron ti-.eso- preljsiisry results that several of
the cladding alloys offer the prczi.o-^ of a significant decree of protection for the
present first-stage bucket alloy, IM 738. "ho corros ior. rate of GiC?^-1 ir the cast and
wrought form was 1.1 mi Is/1000 hours and ir. trie powd'-r r.eta, i urgy forr: was 0.ซ milr./loOU
hour:;. Those w<-re clearly the mo::t resistant nater i a Is in this first test. Since trie
small burner rip. is being operated uncer cone!", t i on:; whicr. saturate thซ- specimen surface
with condensate (corrcsponoinf; to trie reaction rate-1 ir.: ten regime in rigur*- 3 ', small
burner rig data i:an be use?: directly to estir.ate the life of f < rst-stage tickets for
the I'FEC/Oas Turbine. Presently the cladding process is b<-;;:g developed to arply 10
mil thick claddings to first r,t;jge buckets. 'Jsing the re.:u.ts in Table VII, a 10 mil
cladding or GKZOII PM would provide 25.000 hours of protection at ItOO^F while GE 2v'li
would provide 9000 hours of protection. On the other hand, th>- corrosion of t:are III-
738 would exceed 10 mils in j>;.-s than 1000 hour:;. f!JJ-7'* is one of tne r.ost corrosion-
resistant alloys in a conventional Oil-Fired -"as Turbine). "hซ. above estimates for cla-i-
dings suggest that there is a pood chance that thick corrosion-resistant claddings will
offer the corrosion protection required in the PFI'-C envi ronK.ent. hesults from or." test,
however, can be misleading. 8e-ults of the add] t i or: jl corrosion tests planned for the
program described here, together vith additional long-time tests of at least 10.000 hours
duration in a real PFRC cnvi ronsent, arc re-quired before the life of the corrosion-re-
sistant cladding materials can be predicted with confidence.
SUMMARY AND CONCLUSIONS
There is the potential for vapor phase alkali tr.etal in the cor.bustion products
from the PFBC which is several orders of n-'.gnituae greater th^n present limits
for petroleum-fired gas turoines.
Preliminary test results suggest that several of the cladding compositions may
have excellent corrosion rr -stance to the PFEC environment.
Long-time corrosion tests of at least 10,000 hours in the coal environment are
required to confirm the corrosion resistance of candidate materials.
Erosion testing of cladding alloys i: also critical.
The overall solution to the problems identified here will require a combination
of improved hot section materials to resist c-rrosicn/erosion and significant
advances in hot gas clean-up technology.
ACKNOWLEDGMENT
This study i.s supported by the US Department of Energy under Contract Ho. EX-76-
C-01-2357 issued by the Fossil Energy Progras. George C. We'.h of DOE/FE is gratefully
acknowledged as the Program Manager. The authors are particularly ind'bted to Mr. H. von
E. Door ing. Gas Turbine Products Division (GE), Cor many useful discussions.
734
-------
Table VII. Cladding Corrosion Data
from 1600 F Sea 11 Burner
Rig Test with 79 Pt-E !.'a
and 112 pps K in the Fuel
Cladding on IN-738 Hi Is per 1000 hr
CE-25
-------
REFERENCES
A.I-. Foster. H. f.oering, J.W. Hickey, "Fuel Flexibility in Heavy Duty Gas Turbines,"
Gas Turbine Products Division State of the Art Paper Ho. GER-2222L, 1977.
General Electric Cocpany, Gas Turbine Products Division, "High Tenperature Gas Tur-
tine Engine Corponent Materials Test Program - Task I, Quarterly Technical Progress
Report No. 5, EHDA Contract Ho. E( 148- 18}- 1765.
J.W. Schultz and W.R. Hulsizer, "Corrosion Resistant Nickel-Base Alloy for Gas Tur-
r-ines," "etals Engineering Quarterly, p. 15, August 1976.
National Research Development Corporation, "Pressurized Fluidized Bed Combustion,"
OCH Contract lu-32-001-1511, Finil Report, November 1973.
L.3. '.bsak and H. von E. Doerir.g, "Host-Test Evaluation of Gas Turbine Alloys in
Fluidized Bed Combustion Gases at CRE," Gas Turbine Division Report 77 GTD-3, Jan-
uary} 1977.
Combustion Power Corporation, "Hot Corros.'on in the Direct-Coal-Fired Gas T.jrbir.e"
(A Supplemental Heport} ERDA Contract E(^9-18)-1536, September 1976.
736
-------
Sorbent Regeneration
737
-------
INTRODUCTION
f'P. rปAMAN, rHAIPMAfi: Our next paper is entitled Themodynanics
of Peqeneratinq Sulfated Lino to he qiven hy Or. Ton '-'heelock.
-------
-------
Thermodynamics of Regenerating Sulfated Lime
Firoz M. Rassiwalla and Thomas O. Wheelock
Chemical Engineering and Nuclear
Engineering Department
Engineering Research Institute
Iowa State University
ABSTRACT
The thercodynamics of a high temperature reductive decomposition process for
regenerating the lime sorbent which has become sulfated in a fluidized bed combustor
arc analyzed to reveal the process characteristics and to show the effects of various
operating conditions on process performance. For this analysis it is assumed that
the reaction system is in thersodynamic equilibrium. The effects of temperature,
pressure, and reducing state on the extent of desulfurization, sulfur dioxide con-
centration, and the fuel and enerpy requirements are shown. Also the effects of using
different types of fuel including coal and cethane for regeneration are indicated.
INTRODUCTION
The future application of fluidized bed combustion systems for coal may depend
on the successful development of a process for regenerating the lime used to absorb
sulfur oxides in these systems. Otherwise the problem of supplying these systems
with lime and disposing the sulfated sorbent may be horrendous.
One of the nost pronising regeneration processes under development involves
decomposing the sulfated lime at high tentperature in a reducing atmosphere produced
by the partial combustion of coal or other carbonaceous or hydrocarbon fuel. This
process is based to a considerable degree on earlier work at Iowa State University
which was directed toward the development of a process for decomposi .g gypsum and
anhydrite, the naturally occurring minerals of calcium sulfate^-*'. .'he reductive
decomposition process was demonstrated with a scall pilot plant supplied with anhyd-
rite and natural ?,as at Kent Feeds Inc.8. The anhydrite was treated in a fluidizcd
bed reactor in which natural pas was also partially combusted to supply both heat
energy and carbon monoxide and hydrogen for reaction with calcium sulfate. A major
improvement in the process resulted with the discovery at Iowa State University that
calcium sulfate can be decomposed advantageously in a two-zone fluidized bed reactor
in which the material is alternately exposed to reducing conditions and oxidir.ing
conditions'. By this procedure the reaction driving force can be kept large without
resulting in the production of an excessive amount of undesirable calcium sulfide
by-product.
The application of the reductive decomposition process co the sulfated lime
produced in fluidized bed combustion systems has been underway for sometime. Signifi-
cant process development prograns have been carried out at both Exxon Research and
Engineering Co.10"1-' and Argonne National Laboratory 14-13. At Exxon this effort has
reached the stage of trial runs with a continuous flow "miniplant" regenerator fueled
with natural gas and coupled to a pressurized fluidized bed coal combustor!3. At
Argonne the prograa has advanced to operation of a smaller continuous flow regenerator
fueled alternatively with isethane or powdered coal!6.18. The Exxon unit has operated
under pressures up to about 10 atn. while the Argonne unit has operated at lower pres-
sures (1.1-1.5 ata.). Continuous regeneration of sorbent lime has also been demon-
strated in an atmospheric pressure unit operated in tandem with one of the first
experimental fluidized bed combustion systems for coal which was built and operated
by the firm of Pope. Evans and Bobbins Inc.IS. This unit was fueled with coal and
operated continuously for several days during a trial run.
Although these developments have been most encouraging, more information is
needed for process analysis and design. More specifically, the performance char-
acteristics of the process are needed. While these may be determined experimentally,
the amount of experimental wcrk entailed is so great that use of theoretical models
of the reaction system to predict these characteristics should also be considered.
For the work reported here, a nodel of the system based on thormodynamic equilibrium
was analyzed to predict the effects of temperature, pressure, and reducing state of
740
-------
the system on the extent of desulfurization, sulfur dioxide concentration, and the
fuel and energy requirements. Also the effects of using different types of fuel
including coal and methane were investigated. The present study relied heavily on
previous studies of the process thermodynamics made by Wheelock and Boylan3, Skopp
eฃ aj.,10. Vogel tฃ ซa.l<>. and Engel20.
REACTION SYSTEM
The most successful demonstrations of the reductive decomposition process
have been made with fluidized bed reactors supplied with calcium sulfate. fuel and
air. with the amount of air being less than that required for complete combustion
of the fuel and the whole system operated under steady-state conditions. Under these
conditions the fuel has been converted to a mixture of carbon monoxide, hydrogen.
carbon dioxide, and water vapor with a corresponding release of heat. Thus gaseous
reducing agents have been available to react with the calciun sulfate at high tempera-
ture and the following endothermic reactions have probably occurred:
CaSO^ + CO = CaO + C02 + S02 (1)
CaSOA + H2 CaO + H-20 + S02 (2)
In addition to these reactions some of the calcium sulfate has been reduced to calciun
sulfide by exothermic reactions such as these
CaSOA + 4 CO = CaS + 4 C02 (3)
CaSO^ + A H2 = CaS + 4 H20 (4)
Under these reaction conditions the solids have been well mixed and in intimate
contact with the gas phase.
If the components of the preceding reactions are in thermodynamic equilibrium,
it can be shown that the intensive state of the reaction system at 1 atrr. can be
represented by the phase diagram of Figure 1. The diagram was constructed of values
which were calculated from basic thermodynamic properties under the assumptions of
ideal gas behavior, unit activity of each solid component, and both isothermal and
isobaric conditions. It shows that the intensive state of the system depends on fix-
ing three parameters such as temperature, pressure, and the ratio of carbon monoxide
to carbon dioxide. This ratio is a measure of the reducing potential of the system
since the gas phase becomes more highly reducing as the ratio increases. Alterna-
tively the ratio of hydrogen to water vapor could have been used since the two ratios
are related by the expression
(5)
where Kg is the equilibrium constant for the water gas shift reaction shown below.
CO + H20 - H2 + C02 (6)
Figure 1 also shows that the number of solid components which are present
depends on the temperature, pressure and reducing potential. Each sclid component
occupies a separate phase. Thus in the region represented by area 1 of the diagram
only calcium oxide is present, in the region represented by area 2 both calcium sul-
fate and calcium oxide are present, and in the region represented by area 3 both cal-
cium sulfide and calcium oxide are present. All three components can coexist only
along the boundary separating area 2 and area 3 so this can be regarded as a co-
existence line. Along this boundary the partial pressure of sulfur dioxide is deter-
mined completely by the following reaction:
5 CaS04 + 5 CaS ฐ CaO + S02 (7)
741
-------
2400
2200
UJ
a:
UJ
a.
2000
1800.-
1600r
SO,
.7
.3
.1
.05
.02
.01
.005
.003
.001
.0005
.0001
.00001
.000003
.000001
0.04 0.05
Figur* 1. Equilibrium phn* diagram
of the system rt 1 Mm.
742
-------
Since the equilibrium constant for this reaction is
V = P fO\
K7 PS02 (8)
it follows that the partial pressure of sulfur dioxide is dependent only on tempera-
ture when all three solid components are present. On the other hand, in area 2, the
partial pressure of sulfur dioxide is determined by reaction 1 and the partial pres-
sure of sulfur dioxide depends on both temperature and the reducing potential as
shown by the expression below defining the equilibrium constant K^.
(9)
Moreover in area 3 the partial pressure of sulfur dioxide is determined by reaction
10.
CaS -f 3 C02 ป CaO + 3 CO + S02 (10)
and therefore
where K\Q is the equilibrium constant for this reaction. From equations 9 and 11 it
can be seen that for a given temperature the sulfur dioxide partial pressure varies
directly with the reducing potential in area 2 and inversely with the reducing
potential in area 3. Moreover from Figure 1 it can be seen that for a given tempera-
ture the maximum sulfur dioxide partial pressure is obtained along the three solid
component coexistence line.
A phase diagram similar to Figure 1 was presented by Vogel et al. '^ for a
total system pressure of 10 attn. At the higher pressure the region represented by
area 1 is not present while two other regions corresponding to mixtures of calcium
carbonate with calcium sulfatc or calcium sulfidc respectively are present. The
calcium carbonate containing regions are present at temperatures below 1950ฐF. If an
inert gas is present, the calcium carbonate containing regions are limited to lower
temperatures than this.
THEORETICAL MODEL
A specific type of reaction system was assumed to provide a basis for analysis
and prediction of the process characteristics. Thus it was assumed that calcium
sulfate, fuel, and air would be fed continuously into a steady state, isotherr-al and
isobaric fluidized bed reactor where the solids would be well-mixed and all of the
products leaving the reactor would be in equilibrium. It was assumed that the react-
ants would enter the system at 77ฐF and the products would leave the system at the
reaction temperature. Also it was assumed that the reactancs would be fed in such
proportions that essentially all of the calcium sulfate except for a trace would be
converted to either calcium oxide or calcium sulfide. In addition it was assur.ed
that the gas phase would exhibit ideal gas behavior and that the activity of each
solid component would be unity.
The effect of various types of hydrocarbon fuels was investigated. It was
assumed that each fuel would react with oxygen in the fluidized bed to provide a
mixture of carbon monoxide, carbon dioxide, hydrogen and water vapor. It was further
assumed in the case of bituminous coal that this material could be represented by the
formula CHn.8 and that its heat of formation would be negligible.
Although the system was not limited initially to the components involved in
reactions 1-4. it became apparent subsequently that those components plus nitrogen
were the only ones likely to be present in significant amounts. Analysis of the
system by the Cibbs free energy minimization techniquc21 showed that under highly
743
-------
reducing conditions produced by feeding methane, none of the following components
were present in significant amounts: methane, carbonyl sulfide, hydrogen sulfide,
elemental sulfur, or sulfur trioxide.
As long as the preceding assumptions apply, the same model can be used to
analyze eitr.?r the one-zone or two-zone fluidized bed reactors. Thus it should not
pake any difference whether the bed is divided into various oxidizing and reducing
zones as long as the solids are well mixed and all of the products leaving the reac-
tion system are in equilibrium.
COMPUTATION METHODS
The performance characteristics of the model system defined above were deter-
mined by analyzing a series of equations representing the equilibrium state of the
system and the material and energy balances. In doing this the independent para-
meters which were usually specified to fix the state of the system included tempera-
ture, pressure, type cf fuel, and percent air.
The percent air is the percentage of the stoichiomeiric amount of air required
for complete combustion of the specified fuel to carbon dioxide and water vapor.
Hence, it is another measure of the reducing potential of thซป system and it is a more
convenient independent parameter than the Pco/Pco-> ratio because the percent air can
be controlled directly by a plant operator.
The performance characteristics which were calculated included the composition
of the gas leaving the system and particularly the sulfur dioxide concentration of
the gas, the percent desulfurization of the solids, the fuel requirement of the pro-
cess, and the thermal energy requirement. The thermal energy requirement is that heat
added to the system beyond that supplied by the fuel which is fed to the reaction
system.
The performance characteristics were determined for two operating regions A
and B. Region A corresponds to area 2 of Figure 1 while region B corresponds to the
boundary between area 2 and area 3 of this figure. Thus operation in region A would
lead to essentially complete desulfurization of the solids while operation in region
B would lead to incomplete desulfurization because of significant conversion to cal-
cium sulfide. However, in both regions the solidi* were assumed to contain a trace of
unreactcd calcium sulfatc. Operation in a region corresponding to area 3 of Figure 1
was net considered because it would lead to large conversions of calcium sulfate to
calcium sulfide as well as waste fuel.
For the equilibrium analysis in region A, the series reactor technique22 was
used. In applying this technique it was assumed that reactions 1 and 6 take place
sequentially with the system first coming to equilibrium with respect to one reartion
and then the other. The process is repeated until no significant change in the state
of the system is observed. Since this is an iterative technique, the calculations
were performed by a digital computer.
For the equilibrium analysis in region B, an explicit solution of the algebraic
equations was obtained. This involved solving the equilibrium defining equations for
reactions 1, 3 and A together with the appropriate material balances.
After the equilibrium state of the system was determined and the material
balances were solved in either region, an energy balance was used next to find the
additional thermal energy requirement of the process. The expression for the energy
balance is shown below.
Q - r n^K. - r m.H. (12)
p i i R J J
This expression indicates that the heat requirement Q is equal to the difference
between the enthalpy cf the produces and the enthalpy of the reactants.
744
-------
PERFOR>JA:;CE CHARACTERISTICS
The calculated performance characteristics of the model reaction system are
presented in Figures 2 to 10. The first of these diagrams shows the two selected
operating regions A and B for the case where methane is the fuel and the total
operating pressure is 1 atn. From this diagram it can be seen that in region A the
partial pressure of sulfur dioxide in the product gas is virtually independent of
temperature and therefore almost entirely dependent on the percent air supplied to the
system while in region B the partial pressure of this component is entirely dependent
on temperature. In region B the sulfur dioxide partial pressure is determined by
equation 8 and in region A by equation 9. Equation 8 shows wh the partial pressure
depends only on temperature. On the other hand, equation 9 indicates that the partial
pressure of sulfur dioxide should also depend on temperature in region A because of
the effect of temperature on K.\. However, the partial pressure of sulfur dioxide is
also dependent on the PCO/PCUT ratio in region A and from Figure 2 it can be seen
that this ratio decreases as the temperature increases. Hence. the effects of temper-
ature through KI and PCO/PCOT largely cancel out.
Figure 2 also shows that in region A the Pro/Pco- ratio depends on both temper-
ature and the percent air whereas in region B this ratio depends only on temperature.
In region A for a given temperature the Hcn/1'cfi ratio increases as the percent air
decreases wh<"h is not surprising since these quantities are both measures of the re-
ducing potential. However, in region B for a specified temperature the Pco/Pcoj ratio
does not change with the percent air because in this region the PCO/PCO- ratio is
determined by equation 13 below which is based on reaction 3.
PCO/PC02 l/K31/4 <13>
Hence, as the percent air is reduced the Pco/Pco.. ratio does noc increase but instead
more calcium sulfate is converted to calcium sulfiJe in place of calcium oxide.
Although Figure 2 applies specifically to the case where methane is the fuel. ..
the relationships portrayed by this diagram are not much different for other fuels.
On the other hand, changes in total system pressure have a marked effect on the loca-
tion of the boundary separating regions A and B as can be seen from Figure 3. Even
though this figure was developed for the case where coal is the fuel, the location of
the line separating regions A and B at 1 atm. is not far from the location of the
corresponding line in Figure 2 for methane. However, increases in system pressure
nove the boundary separating regions A and B to higher and higher temperatures thus
Increasing region B at the expense of region A. Therefore at higher pressures there
is less opportunity to operate in region A than at lower pressures.
The concentration of sulfur dioxide in the product gas where coal is the fuel
is shown by Figures 4 and i for different operating conditions. For a total pressure
of 1 atn.. Figure U again indicates that the concentration of sulfur dioxide depends
almost entirely on the percent -ilr in region A and on temperature in repion B. Al-
though very high sulfur dioxide concentrations arc theoretically attainable in reelon
A at low percent air. the thermal energy requirement for these conditions is imprac-
ticably high. The constant heat input lines or isocalorlc lines plotted in Figure ซ
represent a reasonable range of heat input. For an adiabatlc reactor at 1 atm. the
maximum attainable sulfur dioxide concentration is about 9.5% and this would be
obtained with a temperature of 1850ฐF and 777, air. By supplying 40 kcal./m. CaSOi
additional heat, the maximum sulfur dioxide concentration could be increased to about
lAx while the temperature would have to be raised to 1S70ฐF and the percent air re-
duced to bo'!.. Or. the other hand, for a heat loss of 20 kcal./m. CaSOi from the
system, the maximum attainable sulfur dioxide concentration would be about 3',. For
any specified heat input, the maximum sulfur dioxide concentration is obtained along
the boundary separating regions A and B.
The effect of total system pressure on the theoretically attainable sulfur
dioxide concentration Is shown in Figure 5 for two different temperatures. 1900 and
2100ฐF. In this diagram the Isobars are drawn as solid lines in region A and dashed
lines In region B. It can be seen that in region A the total pressure has relatively
little effect on the sulfur dioxide concentration while in region B It has a major
effect with the concentration falling off as the pressure Is raised.
745
-------
2400
2200-
o
jg
i
2000
1800-
1600-
i 20 40 60
AIR. '-
Figure 2. Sulfur dioxide partial pressures resulting
from the reaction of CH4-air-CaSO4 at 1 atm.
80
2300
2200
2100
2000
1900
1800
1700
COAL
-PRESSURE = 10 atm
_ CaO-CaC03
COEXISTANCE LINE
[__ _ ^
30 50 70
AIR, 2
Figure 3. Effect of pressure on
the boundary ceparcting reojora A and B.
90
746
-------
COAl
PWSSURl - 1 atm
Figure 4. Sulfur dioxide eonetnUMions retutting
from the reaction of coal-air-CaSO4 ซ 1 atm.
'.PRESSURE ซJL atm
. kcal/ro CaSO^
40
50 70
AIR. I
40
32
24
16
8
TEMPERATURE ซ 2100 QF
ESSURE 1 atm
kcal/m CaSO.
o
O O
30
SO 70 90
AIR. :
FigurtS. Sulfur dioxid* eonccntratiom for vartout
prcoum baitd on coal.
747
-------
COAi.
PttSSWE 1
-------
11
cow.
PRtSSL'RE ป 1 au>
FigurtB. Fuel requmnwtt* ol lhซ proom brad on out
SO 70
AIR. '.
Figun9. FoH requirvntnti it ปซnouป pnourtt bnx) on coat
90
749
-------
65
ปซ
J- 50
35
,ซ 20
o~
W 10
.. 20
ป
o
x*-
10
0.1 r
0.15
0.05
30 50 70
AIR. I
Figure 10. Product gป competition bซsea
on cod or rnttham el 2000'F and 1 (tin.
90
750
-------
The percentage desulfurization of the solids where coal is the fuel is shown
in Figures 6 and 7 for different process conditions. In region A the solids are shown
to be completely desulfurized because this was one of the underlying assumptions on
which the analysis is based. In region B the solids are incompletely desulfurired
because they are partially converted to calcium sulfide. From Figure 6 it can be seen
that lower temperatures and lower values of percent air lower the percentage desul-
furization and favor the production of calcium sulfide. Thus at 1500ฐF and 307, air
almost all of the calcium sulfate would be converted to calcium sulfide. Figure 7
shows that higher pressures also inhibit desulfurization and favor the production of
calcium sulfide. In this diagram the isobars are again represented by solid lines
in region A and dashed lines in region B. Figure 5 and 7 show that for a system at
10 atm. total pressure it is not very practical to operate at a temperature as low as
1900ฐF because it would only lead to very incomplete desulfurization and very low
concentrations of sulfur dioxide. However, at 2100ฐF it would be theoretically
possible to operate at 10 atm. total pressure in region A with coraplece desulfuriza-
tion of the solids and product: a product gas with about 1\ sulfur dioxide.
The fuel requirements of the process in the case of coal are shown in Figures
8 and 9 for various process conditions. The fuel requirements are lower in region A
than in region B because the desulfurization reactions (1 and 2) consume less fuel
than the calcium julfidc forming reactions (3 and 4). Also the fuel requirements are
lower for lower values of percent air in region A than for higher values because less
of the fuel is converted to heat energy and more heat is supplied to the system from
an external source. Figure 9 shows that pressure has only a slight effect on the fuel
requirements in region A h.it a major effect in region B because of the increased pro-
duction of calcium sulfide which pressure favors.
It has been noted above that the use of various hydrocarbon fuels does not
greatly affect the relationships indicated by Figure 2. In other words, for a given
temperature, pressure, and percent air, the equilibrium concentration of sulfur
dioxide in the product gas changes only slightly as the hydrogen-to-carbon ratio of
the fuel changes. Thus as this ratio increases, there is a tendency for the sulfur
dioxide concentration to decrease with the change being more pronounced at low percent
air than at high percent air. For example, at 2000ฐF. 1 atm.. and 80?. air the con-
centration of sulfur dioxide in the product gas would be 8.57. if coal were used and
8.0% if methane were used. Similarly with 607. air the concentration of sulfur dioxide
would be 18.8% with coal and 17.27. witl, methane.
The concentrations of other components in the product gas are affected more
than the concentration of sulfur dioxide is affected by the hydrogen-to-carbon ratio
of the fuel (Figure 10). As night be expected the gas would contain higher concentra-
tions of hydrogen and water vapor and lower concentrations of carbon dioxide and car-
bon monoxide if methane were used than if coal were used under similar conditions.
Although the nature of the fuel would not affect the percentage desulfurization
of the solids in region A. it would have some effect on this parameter in region B
with the percentage desulfurization increasing as the hvdrogen-to-carbon ratio of the
fuel Increases. For example, at 1900ฐF. 1 atm.. and 55/i air the solids would be 96'.
desulfurized if coal were used and 997. desulfurized if methane were used.
The fuel requirements of the process would also be affected to some extent by
the nature of the luel. In general, the moles of fuel per mole of calcium sulfate
treated would decline as the hydrogen-to-carbon ratio of the fuel rises. This effect
would be due to the Increasing amount of hydrogen present per r.ole of fuel.
SELECTED CASES
The anticipated performance of the equilibrium desulfurl.-.atlon system for se-
lected process conditions is presented in Table I. A comparison is provided between
operations with different types of fuel, temperatures, pressures, and heat inputs
(or losses). Also a comparison is provided between systems with no heat recovery and
systems with maximum heat recovery. The latter would Involve recovering sensible
heat from the products to preheat the feed. Although both calcium sulfate and air
would be preheated, the fuel would not be preheated because It eight undergo decom-
position and It would be small In amount compared to '.-he other materials. The maxi-
mum heat recovery would take place when one of the following conditions is achieved:
751
-------
Table I. Performance of Dt-siil fijri ?..-it ion Sys!ซ:n L'ndtr Different Process f.'onJi i ions
Fue 1
Type
Coa 1
cn/t
CH,]
CH,.
CH^
Cil,
Coa 1
cn4
Cil;
CH,]
C"/,
CH,;
Ter.p. .
"K
2000
2000
2000
2000
2300
230')
2000
2000
2000
2000
2300
2300
Tress. .
a i ^ .
1
1
1
10
1
10
1
1
1
10
1
10
g"
kcal ./m.
.-;.) lie.
0
0
-10
0
0
0
Ma xi muni
0
0
-10
0
0
0
Ai
it
1
r .
S0^
IKS ul
f . .
rr.
CaSOT
::. A i r
Recovery
.0
.5
.5
.0
.0
. 5
at
.5
.
. 70
.(If,
.>,')
.66
'i. 3V
'.'J.46
11 .63
:o. c.6
1 3 . 36
16.60
3.21
3.7-S
4.27
2 . OS
3. 76
3.67
heal input, kcal./m. CaMC^
(l^ the calcium sulfate and air .ire preheat ft! to the reactor temperature or (2) the
products arc cooleii ti> ambient temperature.
A comparison of the: anticipated results with and withoijt heat recovery shows
the importance of the lat'.er. Hy recovering the maximum possible amount of heat, the
sulfur dioxide concent "ation would he more than doubled and the fuel requirenents more
than halved. In addition the air requirements would be reduced by about two-thirds.
The smaller air rate would permit usini' a smal ler diameter fluidized bed reactor.
Most of the cases listed in Table I are based on adiabatic operation. However.
it can be seen that a heat loss of 10 kcal./m. CaSOA treated would increase the fuel
and air requirements and reduce the sulfur dioxide concentration sij'.nificantly.
Therefore heat losses should be minimised through adequate insulation.
Operation with coal would be somewhat more attractive than operation with
methane because it would require less air and provide ป Mr.hcr concentration of sul-
fur dioxide. It would be impractical to operate at 10 atm. total pressure usinR a
temperature oป 2000ฐF because of Incomplete desul furi;-.aLlon. Satisfactory operation
at this pressure would require a higher temperature such as 2300T. On the other
hand, at 1 atm. votal pressure it would be more efficient to operate at 2000ฐF than
at 2300ฐF.
In regenerating material which is only partially sulfated. the unsulfated lime
would behave like an inert component and it would affect the energy balance. For
example, if adolomitic lime were sulfated to the extent that half of the calcium
oxide portion were sulfated but none of the magnesium oxide portion and the material
was at 77ฐF when supplied to the regenerator, ihc overall effect would be similar to
supplying the system with completely sulfared lime at 77ฐF but with a heat loss of
37 kcal./m. CaSO^. On the other hand, if '..iis partially sulfated material was at
1700ฐF when supplied to the regenerator, the overall effect wo-ild be similar to
supplying completely sulfatc-i lime at 77ฐF but with a heat ^air. of about 25 kcal./m.
CaSOtf. In either case the fuel requirement would be based on the actual quantity of
calcium sulfate supplied.
752
-------
DISCUSSION AND CONCLUSIONS
The theoretical performance characteristics of a reaction svster. for re;
in;- sulfated lime have been described. These characteristics are based on an i
brium model which m;?;.- or may not truly represent an actual syste
evidence presented by various j-.roups is cotif 1 ict in;-, so it is not
whether or not the assumption of equilibrium is valid. In some
tory experiments conducted by the Esso >'.roupl' it was reported t
cent rat ions of sulfur dioxide were obtained while re.'enerat inr, s
ever, more recent operation of the Exxon "miniplant" has produce
cc-r.t rat ions which are only about half of the calculated equilibrium
therm.ore the Exxon j;roup-J has questioned the accuracy of the >-er.cr
published free energy data for the solid components of the react
the work of Curran e^ al.-l. The latter t:roup experimentally me
partial pressure of sulTur dioxide provided by reaction 7 and fo
ably lower than the value predicted by the generally accepted tr
resolve this dilemma, more basic research on the t hermodynaniic r>
ci.um sulfale svsten needs to be carried out.
'enerat-
quili-
m. ".he experimental
possible to judt;e
o: the e.'irly labora-
hat equilibrium con-
ulf.lted lime. '!iow-
d sulfur dioxide con-
ium values'1. Fur-
llv accepted
ion system because of
asurt'd the equilibrium
ur.d it to be consider-
ee ener-'v data. To
roper!ies of the cal-
Althou,'h the accuracy of the results may he open to question, the results can
still provide some useful insi.-ht for process development and system design. Thus
the results surest that the sulfur dioxide concentration may be limited as much by
he availability of thermal energy as bv equilibrium because the equilibrium con-
centration of this component can be increased by employing less air and increasing.
the reducing, potential of the system. However, increasing the reducing potential
also requires supplying more heat to the system. Recovering heat from the reaction
products and usinr, it to preheat the reactants not only conserves er.erry but also
serves the same purpose as supplying heat froni another source. The results also S'-IK-
t;est tii.it for an adiabatic system or for some specified level of hv.it input the maxi-
mum sulfur dioxide concentration ar.d most efficient oper.it ion will result from oper-
ating ai the boundary between the A and B regions. Since hii'her pressures force this
boundary to higher temperature levels, operation at higher pressures requires operat-
inj; at higher temperature's. Operating at temperature lewis and reducing potentials
which place the system in rvy,i< n H will only lead to wasteful conversion of calcium
sulfatc to calcium sulfiUe. Finally the results suc.r.est that the hydror.en-lo-carbon
ratio of the fuel supplied to the process is not very critical, but a fuel such as
coal has an edy.e over a fuel such as methane.
ACKNOWLEDGE: IE:.T
This work was supported by the Knj;inecring Research Institute, leva State
University. Ames. Iowa.
REFERENCES
I.
2.
10.
11.
p. 87.
Wheclock and D. R
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Wheclock and D. R
Hanson. G. F
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T. U. Wheelock and D.
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T. D. Wheelock and i). R. Boylan.
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T. D. Wheelock and D.
1968.
T. D.
T. D.
T. D.
A. M.
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753
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hi-rt rand. M S. N'I'IMS. I). '}. Kin/.ier. I.. A. H'i''i. '.'.. '<. 'Iris'ory. .'inli /.ซ.-tl-hซ.-rrr>'js* I'm .'ind H<-;'<--ii'-r.'i' io.i of S'.il f :r-C'):T. :t inini- A'idi! i vc-s.
ซi.-;ii.rt :.'o. A:;i./i:s-f:ป;::-ioo> .:!7:-.f-:m- 'ซ;,-.
C. .1 Vo;-cl. K. I.. ซ:.-irl.s. .) . Ai-f-i-r:-!.-i:i. .". li.i.'is. J. Ki-:.'. . C. B. Schnf f s'.oi 1 .
J . lii-;i:n-r 1 v. .-inซl A A Joi^t.-. Ar,-o:iru-. 111.
I'j . Arj-"nn<- N.it inn.'i I !..ilior.'ii orv. "A Dcv<-lป:>r:ซTi' I'rซป;-r.-im on Pr<.-ss'.ir i xt-il K! uidi x.t-d-P.i'il
Conixis? ion." Hi-por' .'In. A::i./KS-CK::-1011 . A:u:':.-,l St-purt. Julv 1. 1 '* 74-J-jni.- 10. 1'> T
(;. .1. "iii'i-l . 1'. C>!tinin;-l,.i;:i. .1. Kishi.-r. .1. H-i!i!ปlซ-. S. I.i-e. J. I.t-nc . J. Mont ;u-n.-|.
A. I'.-itu-k. T. Sclmf t :.r ol I . S. Sii-i-i-1 . <;. Sniith. S. Srci'.h, .'.. Snydi-r. S. S.ixt-n.i,
.1. St (ii.-kii.-ir. W. Swi i t . (;. Ti-.its. I. Wilson, .irui A. A. JonV-.i-. Arj-onnt-. 111.,
July \'U',.
!'ป. .1. C. r!.ปi>t .ii-.n.'i. .1. I'. I.CMI.-, C. .1. Vo-.'t-l. fj. Thoijos. .-miJ A. A. JonV-i-. Fri-prinf .
K'i'irih Int i-rn.-il ton-il Oinfซ-n-iii-ซ- on I'l uiili xi-d-Hc-d Oimbustinn. Xcl.f.an. V;i. . IK-C .
1-11. \->r>.
17. Ar;;oiitK- r.'.ition.il I..il>or.'iiorv. "IK-con;iosi t ion of C.'ilcisim Sujf.-itt': A Ri-v'li-w of t hi-
!.: ti-r.ituri-." Ki-'mrt ::.ป. AM.-1><- 1 J'}. W. M. Swift. A. K. I'.'ini-k. G. W Srith.
C. .1. Voci-l . ,-ind A. A. .lonK-. Arj-onni-. 111.. Uc-c. l'<76.
Irt. .1. C. :-liinl.-if.n:i. W. M. Swift, f,'. W. Smith. C,. .1. Voi-c-1 . .'ind A. A. .Joni'.c.'. Prt-or in: .
ti'tih Annu.-il Mi-i-t in;-. AlCliK, Chic.ij'o, 111.. Nov. 2 it-Dec. .'-. I'i7d.
\'l. I'ojii-, Kv/ins .-ind Koliliins, Inc., "Stud;- of the Ch/ir.ic K-r Ix.-it ion of Conti'il of Air
I'ol 1 ut.-nil s from .1 l-'lnidi::fil-Kfซl Hoili-r -- The SO/ Acceptor I'rocess." I'.S.
Knvi ro(v-.ซ-rt.'i 1 l'r< -. KI'A-K.'-/.'-O^l . J. S. (Jonlon.'R. D.
CU-nn. S. Khrlich. K. Kden-r, J. ซ. Itishnp. -IIK! A. K. Scott. Atex;iiidr i.i. V;i. .
r>7.'.
.'0. R. K. Kin-el. M.S. Thesis. lowi St.ite fni v.-rs i t y, Ar.cs. lov.i. t'*7f>.
21. B. Ce.irc.e. L. I1. Krown. C. II. F.-irmer, I', ilutliod. and F. S. M.-innin;>, Ind. Knr..
Chem. I'rocess Dt-c. Uev.. I'>. A:n. Chcm. Soc. , V.';ishinj-,ton. D.C., July 1976. p. M2.
ii . M Hodell ,-ind R. C. Reid. Thi-rmiulvn.'tmics and 11 s Appl icat iuns . Trent ice-Ha 11 .
Inc.. Kn^.U-wood Cliffs. N.J.: \W.' ~[~RW. ""'
23. C. !'. Curran. C. E. Fink, and K. Corln. in KUP]_ O;isJ fi c^tj on. F. C. Schora. Jr.
(editor). Advances in Chemistry Series 69. An.' Chora .~~5oc.. Washington, D.C..
1967. p. 141.
754
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INTRODUCTION
MR. DAMAN, CHAIRMAN: Our next paper this norning is concerned
with the pressurized fluidized bed coal combustion and sorbent
regeneration. This is some of the work that is coming out of the
Exxon activity. The paper will be given by Dr. Ruth, who is a
graduate of the City University of fiew York where, among other
things, he studied under Arthur Squires, who many of you know and
who's been very active in this fluidized bed development program.
Dr. Ruth has been with Exxon since 1972 and he's been very, very
active in the work that's going on there in fluidized bod combustion.
So, Larry.
755
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collected in t:.o second cyclone are rcr/,ved thrcuj-h a lock i.c^er. The flu* y.as Is analyzed -..or.tir.-
uously lor :;oj , LO, COj , *'OXi UI-^ |JJ usln^ or.-lir.e ir.sirurxr.t^. i't-rloijic samples art- ta^er* icr
;.ar t Kulate cor.t* ntrat lor. u-ji:.x **ri iso*.ir.et it pr-.^t a;...4 4 total t i Iter .
Coi^ijst ion studies '-ere r-ade witi. tv<> Loal:.. ot. Eastern bilur.ir.oui Ml tsi*%:r>;i> Sean coal ccr.-
taininft J.U wt . * sulfur and an Illinois r.o. tt bitur.Ir.ous coal cor.tair.iry. 4.ป' vt . '1 sullur. i:.e
Luslt ?n to;*! was screened to a size- ra:.?,e of 200 to ปCOO ..r.. The I * i ir.cii toai vas sireeix d ;o a si *
rar.Ku ol 700 to J4MO ซ.c. iuu :*orLents were used, a Virginia linestor.e ('.rove. ป'.> .''>. IV/jj anlh were scn.tned to a siz*.- rar.ye of feiO to J^GO _^.
ihe rc>;e iterator vas r.ol operated at the tl=.ซ the coz^ustion Mudies vert: carried out.
iJbJecl ives
'Uie objซcl*ves of tl.c conbustloo studies vere to as^esu the e^ls&ซonfป fro^ a pressurized
f luidized Led coal cotXust ton system And provide add i I ional er.i* Ineer ln^ data tor U.e de-.i*T. ol larป er
units*, huns vc*re ruซde in vtticii the sorbent to coal rซiio 'expressed as i.a/S =jol-ir ratio;, ctrLu>t'-r
tt'C.peralure , ^tes Jure, super! icial velocity^ expanded ted height , so r bent part it: !e si/e, ar.d ซ.-vi.e'.s
air level vere v.irit-d u-iir.t; llic two cc.als and sorbenla described in ll:e previous section. A seties ui
runs vas alsu r^idซr vith tt.e i_alcir.ed tore of the licestcr.e sorbeiit. llut- ^as composition u.is re.isureซJ
and the emissions uf So t> , buj, ::0X and CO vere detvrcined. Carbon cor_tuttion e! i ic ier.i: y anu -.t-rall
heat tratisler coel t ic ier.ls Lelveeti tlie ted and the cooling coils vere also measured.
Keaujjb an^ Uiscussiyn
^ t M te. i s^s i o i i a . So^ cr;is!iiotis results cbtafr.t-d wilt* the Lasteri; and Mlir.uis cuals arid t'ti/ci
tlolucite !>orbvM are .-.Uovn in * i,;urc 3. In thi:. IJ/.ur*.-. the perci-rt reduction in ^'>j c^Uoicr.t. ia
plot led against it it ur i bent to coal t ecd rat lo expressed as the ฃjolei> oi calcium t ed in thv sur Lent to
the c-oleb *ปt sulfur led in ihv coal. lite btoichlowet r ic Ca/S ruolar ratio for the iJcsul f ur i zat ion
react ion Is 1.0. AJ seen in I iป;ure 3, the* resu i ป tป tor all the runs are correlated 1 a i r 1 y vi-1 ! * a
:. iM^U- lliu* dct-^ite a variation in tec[>eralure Iron ซ4O to (>^tJ*C (IVO to l?>Ofcl>. .1 varialior. ir; ;-a'.
l>ha-.e resldenrf tint* 1 roE O.M lo J.O se.:, a variation in pressure Iron COO to V JO kt'.t ',*j to > ^tc. a!;:*),
a two'lold v.irialloa in Morbvnt part ic le si /c and a v-ir lat ion in coal source and sulfur lO'teut. 1 for.
t'i;;ure J, it -;an be cone ludcd that the Ca/S cellar ratio is I lie pr ls-iปry variable at tec t in** the SOj cr i^-
sionu and the other variable* play a secondary role. It can also be concluded lti.*t t hi- ratt* o! tl;v-
rcac t ion between :>> ป ซmd thv so i bent 1 ป approx ic^ite ly f ir^t order in !*O-> conccr.t rat loi; ^inn; t hซ: t'J^
retention is independent ot the *>ul f ur content ol the coal . 'ih i s i ir.d IIIK lids been reported pr evioi;- 1 y
by others. t U',urc> j aluo indicate:* that very M^it SUj reduction levels can be reached with -JuJo-ili-
sorbcntu at fairly low i.a/S ratiub. for c>xae.ple, a VOT reduction in VJ-v wo*ild reiufru a ta/S r^tio
of about ป .0.
1l:c effect ot tc&pcralurv on SOj vrUftttionn wai* dctvrnlned usiny; the above data arid additional
data obtained at tt-'C|ปeralurvป as low AH 6^0*C retention dropped Iron &j+ at a Ca/S ratio of 1.^
to under 301 aa the te&pcrature decreased froo 9OO to */yU*(, . It.c effect of r.-ปป pnase residence tir.c
on Suป cDti;3lons vaป also cva^urcd with Jolo<e sorbunt. It Is known, b-v..ซfd on earlier vor>. that
the detful f urizat ion react ion rate decrease* an the uu If at ion levr 1 of the so r bent increases.
therefore, any reaction rate expression oust include not only the usual react ml concentration tvrฃ-s,
but Dust alปo account (or the vtfvct of the sorbent aulfalion level on the rate. The eiicct of the
doloaitc sulfalion level on the reaction rate can be determined by calculating first order rat*- con-
vtantu und plotting thvo a^Jlnst the caiciua sulfatlon (or utilization) level, '.his was done not only
for data obtained in the Dinlplant but for data published by Ar^onne National Laboratory and the
National Corl Board Coal Utilisation Research Laboratory. The results arc shown In I lt;ure ซ. Aซ ivcn,
a Kood correlation reปultet desalte wide varl.itions in the geottetry ot the lluidized bed cocbusloro,
coal and liorbcnt source, and operating conditions. With this information, the effect of reiปidence
tine on SO ^ retention can be calculated. This was Cone and compared to ceasured effects in tixuru 3.
Aป shown in figure I/f reasonably good a^reecer.t was obtained between the measured and predicted effects
of gas phase residence tlcซ on iปUi enibsions. Die ca^nltjde of the residence tlcc cifeit la such that
a six (old decrease tn the residence tioซ fron 3 to 0.5 s will cause a decrease in the SO; retention
fron 90 to about bbl at a Ca/S ratio of 1.3. Or, at a 90Z SO2 retention level, decreasing the
residence tlce froa 3 to 0.3 tป would require doubling the Ca/S ratio Iroc. 1.5 to About 3.0.
SO; caisfiions were also measured using Grove No. 1359 1 item I one as the sorbent. The results are
shown in Figure 6. Contrary to the result* seen with do loci to sorbent, a narked effect ot temperature
occurs with increasing tc&pcrature giving higher SO, rccoval levels. Also, the degree of data s;attcr
-------
Figure 3. SO, Retention Uting Dolomite
Sorbent
\
V ' :
'.
TM i/MK>% I
Figun 4. Vootiofi of Rปtป Comtant with
dtctum Utiluamn (or Sulldton of Dolomit*
Ftgurt 5. Etfteซ of Rcudcnoi Titn* on SOj
Retmlion-Ootomiu Sortxnt
SOj RctmtiOfl Using
759
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Is mere pronounced and S02 retention levels are lover compared to results measured with dolomite sor-
bent. ihese effects are due to the inability of the limestone to calcine completely, i.e. for the
carbonate to decompose to the oxide, under pressurized combustion conditions. Calcination greatly
increases the )>urcslty of the limvstone mar.ini; the Interior surface of the stone more accessible to the
SOv rractanl. At the higher temperature conditions, the limestone undergoes extensive calcination,
ai.d although tiie sorbenl is not as active as doloslte, it Is considerably core active than at tre
lc>er t *.'~per;iturcs where the stone is largely In the carbonate fore.
i'cl loving the tests with Grove lines tone, a series of runs was cade in vhlch the limestone vas
calcined outside the combustor and fed to the cocbustor in the calcined fore along ulth the coal.
l!:e activity of the precalclned limestone vas found to be significantly higher than that of limestone,
even w:.cn the ccmbustor vas operated at th.; lover temperatures vhlch *'. not protect In-sltu calclr.a-
lion. li.cse results are shovn in figure 7, vhlch cocpares the 50^ retention meas'-rew vlth precalcined
limestone will, the results shown previously in figure 6 for limestone and Figure 3 for dolomite. As
seen in figure 7, the precalclned limestone is as active, at an equivalent Ca/S cx>lar ratio, as
dolomite.
Although a Uuldlled bed utility boiler vould noraally be expected to operate In the tempera-
ture rani'f of about 8iO to 9iU*C (1550 to 1750*F), operation at cuch iDver t*cperature, i.e. dovn to
ibout 7iO*C (KGO'r), vould be required to turn dovn the boiler output to Batch a decrease in the
electrical power dcnand. A series of runs vas made uelng doloelte, limestone and prccalclr.eC lice-
btone at titperalure* near 7iO*t to determine the behavior of the HiC system at thc&e lover tenpซra*
lured, in particular to deturnine the effect on SO> removal. Soae runs vere also made at tenpcratures
as lov as b90*C to deteralnv the lowest licit of operabillty. Ihe effect of operating at these low
temperatures on SOj retention iH shown in figure 8. In this figure, the curves for higher temperature
operation vlth doloeite and limestone are Known fur comparison. As seen, the activity of dolomite and
calcined lieestone arc cooparablv but loซrer than the activity ceasured in the normal cocbustlon tts-
peraturv ranf.e. llu; dashed line la figure t) represents doloelte and precalcloed Ilex-stone results in
the bVO*7bO*C tecperature rany.e. However, the results with limestone indicate that it I* coepletely
inactive at the "turndown" letperalures.
The effecllver.es* -if calcined lieestone Is believed due to the formation of very larKe pores
during ttie calcinalluo procedure used in this study. Calcination took place at high temperature (870*C)
and pressure C*(X> H'J) and at a hl^h LOj partial pressure (~110 U*<). According to result* reported
ty WestIngliousv Research Laboratory, these conditions, especially high CO> partial pressure, produces
a favorable porv structure vhicti alni&izes diffusional effects. Itte porem are apparently large
enough that they are not constricted by the lubucquunt furcation ol CaMJ^ and CaLUj to the extent that
the sorbent becoaeM less activu.
Ihe olnlBua teepcrature at vhlch cosbustlun wus stable va* 6VO*C (l.'/O'l). At tcepcraturca
bซlow 690*C, CO concentration In the flu. '.*ป increased and te&pcralurei in the fluldlied bad bccaae
mutable.
Other variables exanlned In thla studv had no oiRnlflcant effect on SO^ ซnls>lon>. Thli
Included total preciure Jtilch ranged froซ 60O to 940 kfa (6 to V ato aba), excel* air vhlch ranged
froa & to 1101, and Borbent particle slle vhlch vaa varied by a factor of 2 froo on* batch screened
to a size range of UOO to 240O i-e to a second batch screened to size range of 79O to 1400 -,D.
Aa a result of the above studies, the sorbsnt requlreoents needed to satisfy the current CPA
new source pcrforoance standards for SO; eelsslons froa coal fired boiler (1.2 Ib SOWM BTL' coat
fired) can be esticated. The estliute is shown in Table I. The estimate was based on rซ phrse
residence lice of 2 s and boiler temperature of 9)0*C (1700'F). As seen in Table 1, doloslte and
calcined lloeslone are oore effective than lloestone on a molar basis. However, on a weight basis
llcesione Is slightly acre effective than doloalte vlth a coal containing 2! sulfur. Llccstone and
dolomite are equivalent tor a 3! sulfur coal. For coals containing core than -K sulfur, doloelte is
a>re affective than liaestane even on weight basis. Uovever, calcined llccstone is core effective
than dolonlte for all sulfur levels. Doloalte sorbent requirement can be estimated for other g'ป
phase residence tiers using data given In Figure 5. Vlth this information, a process design basis
can be set for a pressurized fluidized bed boiler in which the relationship between the dolomite
requirements, fluidization velocity and expanded bed depth can be determined.
KOi Emissions. SO, calisIons were also ceasured and were found to vary from 10 to 200 ppa or
0.0* to ".17 g (as :O:)/XJ (0.1 to 0.4 Ib/H BTV). The data *i- shown in Figure 9 where NO, eclssioat
are plotted against percent excess a
-------
Table I. Sorhent Kปquire:*r.ts for Once-through
Pressurized Fluidlzet! Bซd Combustion
Coal SUi
S (':) Kซrttr.t .<-) Li^-stoi-.i:
2 b'i 1.3
3 73 2.1
4 7* 2.8
j 84 3.2
Kv%1diT.ee lie*- 'i !
iur-ptrature '930*C
iolor.ltt? Liru-stonc Lircstir.c Uolo^itc
0.8 0.8 8.2 i.O 1C
1.0 J.O 20 V.4 20
1.2 1.2 34 15 2V
1.3 1.3 SI 20 40
i
i
X-
C
Figun 7. SO2 Retention thing
761
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Figure 8. Effect of Low Temperature
Turndown" Conditions on SO2 Retention
K> IWSVDW
O.I
O.J
0.4
5 ฐ-5
i"
B* ฐ"'
O.I
0.
60 CO
IJTCISS Ai*. X
Figui e 9. Cor^btion of NO, Emiaiom with Exoen Air
762
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of excess air. The temperature effect in t:.e 670 to 940'C (1250 to 1750ฐF) rar.ce was secondary ard
caused only a 251 increase in the emission level. The emissions are well below the Li'A r.ew source
performance standard of 0.3 g (a* NOi)/XJ {0.7 Ib/M 31',.) and have an average value of only O.C9 g/.H.1
(0.2 ib/X Bit) at 151 excess air. the level cost liki.y to be used in a commercial size boiler.
Other Emissions. SOj emissions in the flue gas were fou-.d to vary wiceiy, usually over a r is reduced to Cau and Sui in a fluidized
bed by reaction with a reducing gas at about 1100'C (2000'*F) and rCO-1000 ซPa (7-lO~atn abs) pressure.
Our objective is to determine if regeneration Is technically viable by studying a continuous com-
bustion-regeneration system-. Preliminary work involved batchwise recent, rat ion of sulfated limestone
in a fluidized bed vessel of eight cm diaccter^. This paper gives the results of batchwise rซ^ซ--.etj-
tlvn in the 22 cm diameter Diniplant regenerator and of continuous regeneration in a systen cor.nist ir.g
of the ciniplanl regenerator coupled to the 33 cc .ilnlplinl combustcr. Operabillty of this 4ystec was
demonstrated in a run lasting over 100 hours during which sorbcnt was continuously rcclrculated
between the cocbustor *nd regenerator.
Regeneration Theory
The principal reaction; involved in the one-step regeneration of '.'.iSO^ are:
CaS04 * CO CaO * CO, + SO, (1)
CaSO 4- 4CO CaS * 4CO, (2)
*t .
JCaSO4 * CaS 4CnO + 4SO. (3)
Reactions (1) and (2) are written with CO as the reductant but other reducing a^or.ts (e.g., H_.. C) can
be used with r.lnilar effect. Reaction (1), the desired reaction. '.ป endothercic and is favored by
high temperature. Reaction (2) is undesirable and it best avoided by high temperature and low CO
concentration. Reaction between CaSO^ *nd CaS can alco occur, although reaction (3) is not independent
since it can be written by combining reactions (1) and (2).
The caxlcun partial pressure of SO) is produced in an equilibrium syttrs containing all thre
-------
In our fluidlzed bed regenerator, fuel Is Introduced at the bottom of the bed, creating a
reducing zone. To minimize the amount of CaS present, a stream of secondary air is added about half-
way up the bed, producing an oxidizing zone in the upper portion of the bed. In the oxidizing zone,
the following reactions can teke place:
CaS + 202 - CaS04 (4)
CaS + 3/202 < CaO -t- S(>2 (5)
Particles containing CaS are alternately exposed to oxidizing and reducing environments. In this way,
the amount of CaS is kept small.
Experimental Equipment and Procedures
The regenerator vessel, shown in Figure 10, is constructed from 46 cm (18 in) Schedule 40 steel
pipe, refractory lined to an Inside diameter of 22 cm (8.5 in). The fluidizing grid is a water-cooled
stainless steel plate containing 89 holes of 3.6 mm (9/64 in) diameter. The height from fluidizing
grid to gas exit is 5.8 m (19 ft). Air and fuel (natural gas) are supplied to a burner located beneath
the fluidizing grid. Supplementary air is added higher in the column through a 1.3 cm (1/2 in) tube.
Flow rates of supplementary fuel and sir are typically about twenty percent of the respective fuel and
air flows supplied to the burner, but this can vary considerably depending on the air/fuel ratios
desired in the oxidizing and reducing zones. Gas leaving the regenerator is passed through a cyclone
to remove particulates and a single pass double pipe heat exchanger to cool the gas. Pressure is then
reduced across a control valve and the gas is then piped to a scrubber for cleanup before venting to
the atmosphere.
During a typical run with the regenerator alone, a batch of sulfated sorbent is charged and
heated, under oxidizing conditions, to the temperature desired. Rc.Jucing conditions are then esta-
blished by increasing the flow rate of supplementary fuel. Air and fuel flow rates are adjusted
during the run to maintain nearly constant fluidized bed temperature. The concentration of SO, in
the regenerator off-gas Is recorded, with the shape of the SC>2 vs. time curve appearing similar to
what is shown In Figure 11. After the run, the solids are removed fron the regenerator and analyzed
for Ca, 504-2, s~2 and total sulfur.
Combined operation of the combustor and regenerator required development of a transfer syster
to circulate sorbent between the two vessels. The system, which can be seen in Figure 2, was designed
to accomplish this by utilizing high bulk density (stick-slip) flow of sorbent in transfer lines.
Pressure in the regenerator is maintained slightly higher than that in the combustor. Solids in the
regencrator-to-combustor transfer line move into the combustor when a pulse of nitrogen is applied
to the lower end of the transfer line. The solids' flow rate is controlled by adjusting the frequency,
duration, and intensity of the pulse. Two slide valves are used in the combustor-to-regenerator
transfer line in order to prevent backflow of regenerator gas up the line. These automatic valves
trap solids in the piping between them. Solids are discharged into the regenerator when the bottom
valve is opened. The manual slide valve In the regenerator-to-combustor line is used in the event of
upsets in order to isolate the combustor fiom the regenerator.
Results and Discussion
Batch Operation. Eleven runs were made in which batch charges of either sulfated limestone or
dolomite, prepared in the fluidized bed coal combustor, were regenerated. Pressure was about 910 kPa
(9 atm) and average bed temperature normally ranged from 1027-1120ฐC (1880-2050ฐF). Fluidized bed
height was 1-1.5 m (3.3-4.9 ft) and gas contact time 1-2 seconds. The objective of these rms was
to determine the extent of bed agglomeration, if any, the concentration of SC>2 produced In the off-gas
from the regenerator, and the degree of reduction of CaSO^ to CaO.
Bed agglomeration was invariably associated with excessively high temperatures (above 1150ฐC);
when temperature was well controlled agglomeration did not occur. One should appreciate that control-
ling bed temperature is more difficult in batch than continuous operation because in batch operation
the rate of regeneration, and hence the heat requirements for the endo thermic regeneration reaction
(1300 kJ/kg or 560 BTU/lb of CaSO^ converted) varies during the run. Hence, fuel and air flow rates
had to be continuously adjusted to compensate for the variation in heat load as the charge of CaSO/,
was regenerated. In runs made at average bed temperatures of up to 1100ฐC, the extent of agglomera-
tion was insignificant.
764
-------
WY FUEL
FLJU1D1ZING GRID
^f- BtKfCR
1 4|0 Jl BXHSK FUCL
Figure 10. Miniplant Regenerator Fuel
and Air Inputs
Figure 11. Typical SO2 Emission for Batch
Regeneration Run
765
-------
SC>2 levels measured during batch operation reached 3.0 cole percent (3.7 percent on a dry
basis), or about half of the calculated equili'-riuii. value. However, some tice ago, Curran^ determined
equilibrium SO^, levels for reaction (3) experimentally. I'slng Curran's levels, which are lower than
the calculated levels, the SO2 concentrations which we treasured averaged about 80 percent of equilib-
riun. In fact, equilibrium was reached in several runs, using Curran's data.
After each run the fractional conversion of CaSO$ to CaO and CaS was calculated from analyticil
res'-'ts for calcium, total sulfur, sulfate, and sulfide. An average of 94 percent of the sulfate was
reduced to oxide. The acount of CaS present was negligible.
Deaonstration of a Fluldized Bed Combustion System with Continuous Sorbent Regeneration. After
operating the regenerator batchwlse we linked the cocbustor and regenerator vessels and made a run to
deixM.etrate continuous operation. Operating conditions are given in Table II. Conditions (especially
regenerator temper., ;ure) were conservative so as 13 provide the best chance of reaching our goal of
100 houis continuous operation. All variables were kept constant except that makeup limestone sorbent
was added to the combustor at the rate necessary to make up for losses caused by attrition and entrain-
ment from the fluidized beds.
SC>2 emissions from the combustor are given in Figure 12. To establish baseline operation, the
regenerator was operated under oxidizing conditions for the first 24 hours with scrbent recirculating
between combust or and regenerator. Emissions gradually increased to about 5.SO ppm (1.1 Ibs S02/106
Bit). Since the S02 ecissions would have been about 1330 ppm at zero retention, SSO ppm corresponds
to about 60 percent retention. This is a much lower level of S02 emissions than would have been
expected had the combustor been operated at the same conditions, but without sorbent recirculating to
the regenerator. The low combustor emissions can be explained because, even though no regeneration
was occurring during this period, the regenerator was acting as a calc'ner and supplying freshly
calcined sorbent to the combustor.
The concentration of S02 in the regenerator off-gas was nearly steady throughout the run and
averaged 0.53 mole percent (dry basis). This is very close to the concentration predicted by a sulfur
mass balance based on the feed rate and sulfur content of the coal entering the coobustor. The cal-
culated equilibrium concentration at the operating conditions of the regenerator was 2.9 percent;
hence, higher SOj ,'evels would probably have been achieved by burning in the combustor more coal of a
higher sulfur content.
Samples of bed were taken from the combustor and regenerator after the run and analyzed. The
combustor bed contained 35.8 mole percent CaO, 18.5 percent CaCO-,, and 45.7 percent CaS04. The
regenerator bed was 80.5 percent CaO, 2.3 percent CaC03- and 17.3 percent CaS04. Because air was
blown through the hot regenerator bed during shutdown, the composition of these solids may have
changed. Any CaS, if present would have been converted to CaSOA, and possibly CaO.
A sulfur mass balance for the demonstration run Is given in Table III. Recovery of sulfur was
103.5 percent. The sulfur balance is very sensitive to the sulfur content of the coal, which would
have needed to be only 0.07 percent higher to obtain a sulfur recovery of exactly 100 percent.
SYNTHETIC RECENERABLE SORBENTS
The conventional sorbents which have been used In fluidized bed combustion are limestone and
dolomite. There are many problems with these natural materials. Only a small fraction of the calcium
contained in limestone or dolomite is utilized (converted tc sulfate), attrition rates are high,
regeneration requires high temperatures and deactivation begins after only a few cycles of sulfur
sorption and regeneration, and different stones vary greatly In their reactivity with S02 and in
attrition resistance. Improved sorbents are needed which arc superior to Huestone and dolomite
in all of these respects.
For nearly two years we have been conducting an experimental program to Identify and develop
regenerable sorbents that are superior to limestone and dolomite. The Impetus for this work was
several paper studies in which the thermodynamics of a large number of compounds were screened In
order to identify those compounds which could absorb sulfur at the conditions of temperature, pres-
sure, and gas composition that prevail in a fluidized bed coal combustor and be easily regenerated
(5,6,7). About thirty such compounds were identified but experimental results to confirm the thermo-
dynanlc predictions were lacking. Also, there was 10 Information available on sulfation and regen-
eration rates, or on activity maintenance. Thermogravimetric analysis was chosen as the basic
screening tool used to determine which of these materials warranted further study.
766
-------
500
400
0 1.1 ?0 30 40 ^0 60 70 10 'JO 110 130
HOURS INTO RUN
Figure 12. Combustor SO2 Emissions During
Combustion-Regeneration Demo Run
Table 11. Operating Conditions During Demonstration Run
Pressure, kPa
Bed Temperature, Average, ฐC
Bed Height, Expanded, Avg., m
Superficial Gas Velocity, m/s
Combustor
760
900
3.4
1.5
Regenerator
770
1010
2.3
0.6
Solids Recirculation Kate, kg/hr
Residence Tlae of Solids, Avg., hr
Makeup Acceptor Addition Rate,
Equiv. Ca/S, Average
Range
Cor.bustor Coal Feed Rate, kg/hr
Coal Type
Stone Type
79
0.55
0-1.3
1-1/2
Champion (Pittsburgh Seam), 2.07. S
Grove Limestone, BCR No. 1359
767
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Table III. Sulfur Balance for
Combustion-Regeneration Demonstration Run
of Sulfur Entering
Sulfur Entering System
Coal 100
100
Sulfur Leaving System
Regenerator off gas 47.1
Combustor flue gas 20.3
Combustor bed reject 9.0
Combustor overhead solids (flyash) 11.6
Regenerator overhead solids 1.0
89.0
Sulfur Accumulated (ฃ Inventory)
Regenerator bed
Combustor bed
Z S .lecovery 103.5
Table IV. SOj Pickup Per Unit Mass of BaTiOj, CAC, and Limestone
No. Mass SO-j Sorbed Per Unit
Sorbent Cycles Mass of Sorbent After t.'o. Cycles Indicated
BaTi03 50 0.22
CAC 25 0.09
Lioestone 5 0.04
768
-------
Over thirty materials were screened using TGA equipment. The compounds and materials tested
included:
oxiues
aluminates
carbonates
cements
titanatea
composites containing CaO, Al-0,, and SiO-
Screening of potential sorbents was accomplished by subjecting each sample to reaction condi-
tions representative of those in a fluidized bed coal combustor: 900ฐC, 0.1-0.252 S02> 52 02, balance
N2- Regeneration of sulfated sorbents was accomplished in an environment of 1100ฐC, 5* CO, balance
N2. Positive identification of sorberts, sulfated sorbents and regenerated sorbents was accomplished
using x-ray diffraction. Initially, sorbents were tested in the form of fine powders. Subsequently,
the powdered sorbents which performed best were fabricated into pellets and evaluated.
The best sorbents were the titanates of barium and calcium and calcium aluminate cement.
These sorbents have constituted the dual development centers of our program.
Barium titanate, EaUOj, was the most reactive sorbent tested. It reacted more rapidly with
SC>2 than all other sorbents and along with CaTiOj was found to be fully regenerable, maintaining its
activity for any number of cycles. Calclun aluminate cement (CAr:) is a structural material thut is
widely used in high temperature applications. We found that pellsts with high attrition resistance
and good activity could be made from CAC.
The clear superiority of BaTi03 and CAC to the conventional sorbent limestone Is shown in
Figure 13. Pellets of all three materials were cycled between Identical sulfation and regeneration
conditions. The time period for sulfation was arbitrarily selected as 75 minutes. Each material
regenerated completely after several minutes. The utilization during sulfation for limestone
declines to less than 5 percent aiter only five cycles, however, the utilization of CAC !. still
about 16 percent after 25 cycles. Moreover, CAC declines only slightly in activity whereas the
activity of limestone declines sharply. On the other hand, 63X103 actually increases la activity
and, after fifty cycles, the utilization is still over 60 percent.
Rather than comparing these sorbents on the basis of percent utilization, a comparison can
also be made based on the mass pickup of SOj per unit mass of sorbent. Table IV gives this
comparison at the number of cycles indicated for each sorrฐnt. On this basis also, BaTIOj and CAC
are both clearly superior to limestone.
Several techniques, including pressing, extrusion, and granulation, have been utilized to pre-
pare pellets from the titanates and CAC. In a s=all attrition rig designed to simulate fluidized bed
operation at room temperature, scne CAC pellets have proven to be core attrition resistant than natural
sorbents. However, the methods used to prepare pellets were only first attempts. We believe that
substantial improvement in the strength properties of pellets are ye- to be realized.
Evaluation of CAC in the TGA has shown that CAC pellets prepared without any additives
to increase activity are already acre activa sorbents than limestone. Indeed, the CAC pellet of
Figure 13 contained no additives. However, the pore volume of CAC can be increased by using core water
to mix the cement or by using burcables such as carbon black when formulating the cement. During the
heating process the carbon burns out leaving pores. In one test of a CAC pellet, we found that adding
one or two percent carbon black increased the utilization about 40 percent compared to a pellet pro-
pared without carbon black.
Sorbent pellets have also been prepared from mixtures of CAC and fly ash, with the expectation
that the fly ash might act both as a burnable powder to increase porosity and activity and as an
aggregate which would make the pellet stronger. However, the magnitude of the effect of fly ash on
the activity of CAC was unexpected. After four cycles, the utilization of a pressed pellet prepared
from equal volumes of CAC and fly ash was over three t jes the utilization of CAC prepared without
fly ash.
Because of the surprisingly large ioprovement in utilization when fly ash was added to CAC,
we investigated the possibility that fly ash might be chemically promoting or catalyzing the sulfation
reaction. An experiment was performed in which a pressed CAC pellet was rolled in fly ash so that
the surface of the pellet was covered. The pellet was then cycled In the TGA. The result was that
769
-------
BARIUM T I TANAU
CALCIUM ALUHINAU CEMENT
CONVENTIONAL SO.''.BENT (GROVE LIMESTONE)
5 6 7 8
CYCLE NUMBER
SO
AFTER
SULFAIION
AFTER
REGENERATION
Figure 13. Comparison of Utilization of Limestone and Pellets of
BaTiO3 and Calcium Aluminate Cement
770
-------
the presence of the snail quantity of fly ash increased the utilization of the CAC by 60-80 percent.
This suggests that the action of fly ash on the sulfation of CAC is not due to pore formation but
may be a chemical effect, and is perhaps cataiyt.1-. A further implication is that the performance
of CAC as a sortent may be considerably better in a real coal combustor, where fly ash is present,
than in the "sterile" environment of a TCA.
CONCLUSIONS
The primary variable affecting $03 emissions from a pressurized fluidized Jad coal combustor
is the Ca/S molar ratio in the incoming feed streams. Gas phase residence time axd! temperature also
affect SOj emissions but to lesser degrees. Residence tine effects can be correlated using a first
order reaction rate expression which satisfactorily described data from a number of laboratories,
using a number of coals with varying sulfur contents. The effectiveness of dolocite, limestone and
calcined limestone' for SC>2 retention was determined. Dolomite and calcined limestone are equally
effective at the same Ca/S molar feed ratio basis. However, at the same Ca/S wel^t feed ratio,
calcined limestone is much more effective. Limestone is less effective than either dolomite or
calcined limestone due to the inability of the limestone to calcine extensively uufer pressurized
FBC conditions. Dolcmite and calcined limestone are also effective at very low cecbustor tempera-
tures whereas lir.estone is completely inactive. The effectiveness of the calcined limestone is
believed due to the formation of very large pores which occurs when the stone is calcined under
high CC>2 partial pressure conditions.
The data presented in this paper permits the estimation of the sorbent requirements and gas
phase residence time requirements needed to control 502 emissions to any desired level.
Other emissions from the combustor are fairly low. NO* emissions vary froa 50 to 200 ppm,
increasing somewhat with excess air and temperature. The emissions are well wlthฃn the EPA emis-
sion standard.
Pressurized regeneration was studied Initially by reacting batches of sulTzted limestone and
dolomite with a reducing gas at 9 atm pressure and about 1100ฐC in the 22 ex dia-^-ter miniplant
regenerator. The gaseous effluent from the regenerator contained up to 3.7 mole jercent S02 (dry
basis) and conversion of CaSO^ to CaO was nearly complete. Formation of CaS was avoided using the
technique of adjacent oxidizing and reducing zones.
Subsequent to the batch studies, the regenerator vessel was coupled to the combustor and the
system was run with sorbent recirculatlng between the two vessels. An important question is what
reduction in sorbent requirements can be realized by adding regeneration to a once-through system.
This question cannot yet be answered precisely but results of the continuous run provides a clue
to the answer. The average SC2 emission from the combustor was 310 ppm, corresponding to a retention
of 11 percent, at an average Ca/S makeup ratio of 0.55. Figure 6, which gives S
-------
REFERENCES
1. M. S. :.'utkis, et al, "Evaluation of a Granular Bed Filter for Particulate Control In Fluidlzed Bed
Combustion," presentation to Fifth international Conference on Fluidized Eed Combustion, December
1977, Washington, D.C.
2. M. S. Nutkis, Proc. Fourth Intl. Conf. on Fluidized Bed Comb., publ. The Mitre Corp., McLean,
Virginia, 1975.
3. L. A. Ruth, Proc. Fourth Intl. Conf. on Fluidized bed Comb., publ. The Mitre Corp., McLean,
Virginia, 1975, 425-38.
ft. C. P. Curran, et al, Fuv>.l Gasification, Advances in Chemistry Series, 69, Aeerican Chemical
Society, 1967, 141-65.
5. E. P. O'Neill, et al, Westlnghouse Research Laboratory, "Experimental and Engineering Support of
the Fluidized Bed Combustion Program, Task 2, Environmental Control L'slnp Alternate Sorbents,"
monthly reports prepared under EPA contract 68-02-2132, Feb.-April 1976.
6. P. S. Lowell and T. B. Parsons, Radian Corp., "Identification of Regenerable Metal Oxide Sorbents
for Fluidized Bed Coal Combustion," EPA-650/2-75-065, 1975.
7. J. A. Cusunano and R. B. Levy, Catalytlca Associates, "Evaluation of Reactive Solids for S02
Removal During Fluidized Bed Coal Conbustion," EPRI project TPS75-603, 1975.
772
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QUESTIONS/RESPONSES/COMMENTS
MR. DAMAN: Thank you, Larry.
MR. DAMAN: Dr. Ruth, I think you have some questions, right?
DR. RUTH: I have a question from Dr. Macek of DOE. "According
to your results, the gas residence time is very important in S02
sorption. Yet I understand that recent CURL results in pressurized
fluidized bedcombustion shov; no Increased absorption with increased
bed depth."
I have not yet seen this data but I v-nll try to give some possible
explanations for the apparent difference in the results. The most
significant effect of gas phase residence time that we observed was
at lower residence times, about one half to one second. Beyond one
second, the effect seems to decline. I don't know what range of resi-
dence times CURL's data represents. Secondly, we varied the residence
time by a factor of six. I'm not sure what range of variation there
was in the CURL data.
DR. MACEK: About two and a half.
DR. RUTH: And was the CURL data with dolomite or limestoi.e?
There is a lot more data scatter with limestone because of calcination
effects, so that it might be difficult to pick up an effect of resi-
dence time with limestone.
DR. MACEK: I would assume dolomite but I don't know.
DR. RUTH: I have three questions from Dave Henzel of Dravo. The
first one, "Do you have a cost comparison for the barium titanate
versus calcium aluminate cements? What would be the source of calcium
aluminate cement?"
All I can give you at the present time are the costs of preparing
the synthetic sorbents. The overall cost picture would also have to
take into account differences in regeneration costs and sulfur re-
covery costs. We would expect that there would be large differences
in those costs as well. For calcium aluminate cement we would esti-
mate, and cement makers confirm this, that the cement would cost about
$150 per ton in the form of pellets suitable for the combustor. The
barium titanate is more expensive, of course. There we've estimated
costs between 600 and 1500 dollars per ton. Some barium titanate
makers believe they can make it for less than 600 dollars per ton.
As for the source of calcium aluminate cement, it's made by fusing a
mixture of limestone and bauxite.
773
-------
The second question from Mr. Henzel. "Is the precalcination of
limestone done external to the FBC unit and under what conditions?"
We actually carried out the precalcination in our combustor
vessel after we had removed most of the cooling tubes. The fuel used
v/as natural gas, and conditions were about 8 or 9 atmospheres pressure
and somewhere around 1700 or 1750 degrees F.
His third question is, "Would you repeat your contract numbers?"
Well, here goes again. The EPA contract was 68-02-1312 and the
National Science Foundation Grant number is AER75-16194.
I have another from Ray Costello, Burns and Roe Industrial
Services Corporation. "Concerning the effect of gas residence
times on S02 capture your slide only went to three seconds. Would
you expect a significant increase in capture at longer residence
times."
As you can see from Figure 5 of the paper, as the residence
time increased, the incremental increase in sulfur retention became
smaller and smaller; our model which relates gas phase residence time
to sulfur retention (see Figure 4 of paper) also predicts this effect.
MR. DAMAN: Thanks.
774
-------
INTRODUCTION
MR. DAMAN, CHAIRMAN: All right, o :r next paper, also on regen-
eration of limestone by John Vogel. John is a prominent member of
the Argonne team working on fluidized bed combustion problems. He
started off in life messing arcurH with nuclear energy and then
apparently saw the light and switched to coal. He's been at Argonne
since 1956. John, why don't you go right ahead?
775
-------
Development of a Process for Regenerating
Partially Sulfated Limestone from FBC Boilers
John C. Montanga, Franklin F. Nunes. Gregory W. Smith
Eugene B. Smyk, F. Gale Teats, G. John Vegel,
and Albert A. Jonke
Argonne National Laboratory
ABSTRACT
In fluidized-bed combustion of high-sulfur coal, a natural calcium-containing
stone such as a limestone or dolomite is used as the bed material and acts as the
sulfur-accepting agent, forming CaSO,, . A means of significantly reducing waste volume
froir. the combustor is regeneration of the CaSO., to CaO and reuse of the regenerated
stone in the combustor. A fluid-bed, reductive decomposition regeneration process has
been developed in which the sensible heat, the heat of reaction, and the reducing gases
are supplied by partial combustion of coal in the bed at 1IOOฐC. The developmental
stages are reported for-. (1) selection of the process, (2) evaluating the regenerator
performance with limestone, (j) performing cyclic (sulfation-regeneration) liir.estone
and dolomite life studies, and (<4) incorporating the experimental results in a process
flowsheet for a 200-MWe KBC boiler-regenerator system.
'INTRODUCTION
Kluidized-bed combustion of coal is currently being developed for electric power
and/or steam generation since the current national goal, to become less dependent on
foreign energy resources, is heavily dependent on increased utilization of domestic
high-sulfur coal. In the fluidized-bed coal combustion process, the coal is combusted
in a fluidized bi.d of a sulfur-accepting sorbent such as limestone or dolomite. The
sulfur released during Combustion reacts with calcium in Che stone to form CaSO,. . The
primary reasons for using natural calcium-based stones are thoir acceptable reactivity,
low costs, anc1 bountiful supply throughout the United States.
In the fluidized-bed process, approximately 500 pounds of limestone is required
per ton of combusted coal (37., sulfur) to maintain the SO? concentration in the flue gas
below the EPA emission standard. The sulfated limestone product (which is mixed with
coal ash) rnay have commercial uses tor cement block manufacture, landfill, or agri-
cultural lime. Even so, methods for reducing the amount of limestone used in the
process would have the advantages that less stone would have to be quarried and less
waste discarded from the process. Less stone would be used if (1) laboratory-scale
tests could identify the more reactive stones so that these may be used, (2) the reac-
tivity of nonrcactivc stones could be increased by physical or cheni-al modification of
the stone, or (3) the stone could be regenerated for reuse in the combustor.
A fluid-bed reductive-decomposition limestone-regeneration process has been
developed in rDU-scale equipment. In th:.s process, both the heat and reductants are
supplied by partial combustion of coal in the bed at 1100"C. The developmental stages
are reported for: (1) selecting the process, (2) optimizing the process, O) performing
cyclic (i.e., sulfation/regeneration) life studies on the sorbent, and (4) incorporating
the experimental results in a process flowsheet for a 200-MWe FBC boiler-regenerator
system.
SELECTION OF THE REGENERATION PROCESS
Thermodynamic and kinetic information were obtained from the literature on
processes for converting calcium sulfa;e to calcium oxide.'-2 Of the more feasible
processes examined, thermal decomposit-'.on or calcium sulfate was not considered to be
a viable process because of the high temperature, >1200eC. needed to obtain a high con-
centration of SO, in the off-gas. At thif, temperature, ash and sulfated stone would
fuse to form unusable clinkers.
Another process (studied in laboratory-scale equipment) consists of two steps.
CaS that has been produced in the first step by reacting CaSOi, vith a reducing gas at
v-870'C is reacted with steam/CO, at v560ฐC to produce CaC03 and H2S. Cyclic processing
776
-------
by alternate sulfation and regeneration showed that the extent of regeneration of the
CaSO_ decreased significantly in succeeding cycles.' Attempts to understand the
mechanism and to develop the process fu'.'ther were abandoned when it became necessary
to select a process for larger-scale development.
Selected for further development was the reductive decomposition process in
which CaSO., is heated in a fluidizcd bed to --IIOO'C in the presence of reductant gases.
Two solid-gas reactions by which regeneration occurs are:
CaSO/, + CO CaO + C02 + SOj
H2 CaO + HoO + S02
(1)
(2)
At lower temperatures and under more highly reducing conditions, the formation of CaS
is favored:
CaoO.; + 4CO - CaS + 4CC2
CaSO/, + U\\2 CaS + 4H20
(3)
The products are CaO. which is reused in the combustici step, and a SO.--containing
off-gas, which can be processed for its sulfur content. The calcium sulfide concentra-
tion in the product is maintained below 0.17a by circulating particles into an oxidizing
zone where CaS is converted to C".iO. .
EQl'IPMLNT
Figures 1 and 2 illustrate the Process Development Unit (PDU) combustion and
regeneration systems used in this investigation. The combustor has a 15-cra ID and is
approximately 3.'* m high. The regenerator consists of a nominal 20-cm-dia pipe,
refractory-lined to an ID of 10.o cm. Bubble-type gas distributor plates are flanged
to the bottom of each reactor and accommodate fluidining-air inlets, solids feed and
removal lines, and thcrnucouples for monitoring the bed temperatures. In each system,
the coal and sorbent are metcred separately to a single pneumatic transport line which
discharges these solids into the fluidizcd bed above the gas distributor plate.
Expanded-bed heights of -.90 cm in the combustor and -.46 cm in the regenerator are main-
tained with overflow pipes.
To Gas Analysis
System
Test Filter
Stainless Steel
Screw
/Compressor
"-^fl
Pressure
Control
Vjlve
.?_ Ventilation
TJ
Steel
Filter
Exhd'ist
Secondary
Cyclone
Primary
Cyclone
Figure 1. Simplified Equipment Flowsheet of Fluidiied-Bed Combustion Process
Development Unit
777
-------
To Gas
Analy/ers
,
I"
Filter
Sample >/ |^_ rL^JPres
Gas Conditioner crrzpCyclones l-T pon
Sulfated-Sorbent
Cou< Hopper 57 Hopper
Rotary Valve
Feed
Air -,
-
Filter
I Pressure
Control
Valve
xhaust
I Regenerator
! (10.8cm 10)
Surge
i v.LinjJ
"Product -r-
Collector
Figure 2. Experimental Sorbent Regeneration System
Other components of the experimental systems are an electrically heated heat
exchanger for preheating the fluidizing gas and a solids-cleanup system for the off-gas
consisting of cyclones and porous metal filters. Most constituents (SO? . Q2 , CO, H>,
CH,, , and NO) in the off-gas are continuously analyzed.
In the regeneration system, the solids transport air constitutes "-407. of the
total fluidizing gas in the reactor. The remaining fluidizing gas is a mixture of pure
nitrogen and oxygen. Oxygen and nitrogen are metered separately and are mixed to pro-
duce the required oxygen environment in the reactor. Thus the oxygen requirement at
different experimental conditions can be satisfied without changing the fluidizing-gas
velocity.
MATERIALS
The two sorbents tested in these studies were Tynochtee dolomite and Greer lime-
stone. The dolomite contained DO wt 7, CaCO,, 39 wt % MgCOj, and 2.1 wt 7. Si as
received. The limestone contained 41.2 wt '/ CaO, 32 wt 7. COj , and 4.27 wt 7, Si. The
nominal size distribution of each limestone was -14 +30 U.S. mesh.
The coal used in the sulfation of Tymochtee dolomite was a Pittsburgh seam coal
which (as received) contained ^-2.3 wt 7, S. -v-7.7 wt % ash, and -ป-2.9 wt 7, moisture and
had a heating value of 7,600 kcal/kg and an average particle size of 320 urn. In the
regeneration steps with Tymochtee dolomite. Triangle coal was combusted under reducing
conditions. It is a high-volatile bituminous coal with a high ash-fusion temperature
(1390ฐC under reducing conditions).
Sewickley coal was used in both the combustion and the regeneration stepc of the
Greer limestone study. As received, it contained 1.4.3 wt % S, 12.7 wt "L ash, and 1.1
wt 'I. moisture and had a heating value of 7,200 kcal/kg. The Sewickley coal has a
relatively low ash-fusion temperature (v.H20ฐC under reducing conditions) .
REGENERATION EXPERIMENTS USING GREER LIMESTONE
The dependence of (1) CaO regeneration and (2) SO? concentration in the regen-
erator off-gas on key variables such as regeneration temperature and solids residence
time in the fluid-bed reactor was studied for once-sulfated Greer limestone to aid in
optimizing the regeneration process conditions for this limestone. Regeneration of
sulfated Tymochtee dolomite had been studied earlier.3 The experimental conditions and
results for these experiments are given in Table I. Regeneration of CaO, calculated
from chemical analyses of the steady-state materials, was based on the calcium to sul-
fur ratios in (1) the sulfated limestone feed and (2) the regenerated product. These
calculated regeneration values are compared in Table I with the values based on the
778
-------
ratio of the sulfur contained in the off-gas no that in the sulfated limestone feed.
Table I. Experimental Conditions and Results for the Regeneration of Greer Lime-
stone by the Incomplete Combustion of Sewickley coal in a Fluidized Bed
Nominal fluidized-bed height: "-46 cm
Reactor ID: 10.3 en;
Pressure: 129 kPa
Coal: Sewickley (4.3 wt 7. S) ; ash fusion temperature (initial
deformation) under reducing conditions: 1119 C
Scrbent: -14 +30 mesh sulfated limestone (7.7 wt 7ปS)
Exp.
No.
RGL-1A
RGL-1B
RGL-1C
RGL-1D
RGL-1K
RGL-1F
?Based
"Based
Bed
Temp,
"C
1050
1050
1050
1100
1100
1100
Fluidizing-
Gas Feed
Velocity. Rate,
m/s kg/hr
1.23 8.2
1.21 15.4
1.21 26.3
1.23 9.1
1.23 15.9
1.29 25.3
Reducing
Solids Gas Cone.
Residence in CaO
Time, Effluent, Regener. ,
rain 7. 7.* / *,ฐ
22.54
11.93
6.99
20.23
11.59
7.12
on flue-gas analyses.
on chemical anlayses of limestone
3
2
3
3
3
2
sai.iples .
.2
.9
.4
2
!5
.9
72
43
28
79
72
66
2/83.0
1/53.3
3/27.3
5/92.0
3/82.8
9/70.9
Major Sulfur
Compounds in Dry
Off -Gas. 7.,
S02
3.3
3.7
4.3
3.9
6.0
8.4
I'-2S
0.09
0.09
0.05
0.2
0.1
0.06
COS
0.07
0.08
0.1
0.1
0.1
0.09
CS2
0.06
0.05
0.05
0.05
0.03
0.02
Effects of Solids Residence Time and 3ed Temperature on Exter-.t of CaO Regeneration
Lxtent of regeneration values are plotted in Fig. 3 as a function of solids
residence time for two temperatures, 1050ฐC and 1100ฐC. When the solids residence time
was decreased, the extent of regeneration decreased. At 1100ฐC, the regeneration rate
is higher and therefore the conversion ratio of CaSO,. to CaO is less affected by a de-
crease in reactor particle residence time. With a residence time as low as 7 min, the
extent of regeneration is considerable. ->-707o.
Figure 3. Regeneration of CaO in Greer
Limestone as a Function of
Solids Residence Time and
Regeneration Temperature
A "best fit" equation has been obtained by gression analysis for the experi-
mental extent of CaO regeneration as a function of .egeneracion temperature and solids
residence time:
779
-------
In (1 - R) = A- -t- B- : (5)
where
R = extent of CaO regeneration (R = 1 for complete regeneration)
i = solids residence time (reactor particle contact time)
A x 102 = -12.4T - 3.98 (6)
B x 103 = 3.25T - 1.24
T = (t - 1050)/50
t = regeneration temperature, "C
The values calculated tv the model equation (Eq. 5) compare favorably with the experi-
mental results. A correlation coefficient of '-0.97 -..-as obtained for the experimental
data and the results predicted from the model equation. These best fit results for
Greer limestone are compared with results for similar experiments with Tymochtee
dolomite in Fig. 3. The regeneration rates for these two sorbents compare very favor-
ably. This relationship for the rate of CaO regeneration is used in the model for the
regeneration process to optimize the design process conditions and to scale up sorbent
regeneration systems (described in a later section).
Effects of Solids Residence Time and Temporaturc on SO. Concentration in the Off-Gas
In the fluidized-bed regeneration process, the SO-, concentration in the off-gas
is determined by (1) the feed rate of CaSO., to the regenerator reactor (solids residence
time and sulfur content of sulfated sorbent), (2) the extent of regeneration of CaO,
and (3) the gas flow rate through the reactor.
In this series of regeneration experiments, sulfated limestone (containing
7.7 wt 7, uulfur) was used. The fiuidizing-gas velocity varied from 1.21 to 1.29 m/s.
a small variation. The gas flow rate through the reactor was not affected greatly by
the fluidizing gas velocity in these experiments. Therefore, the variation of SO, con-
centration in the dry flue gas was due to the mass rate of CaO regeneration.
The experimentally obtained SO, concentrations in the dry off-gas are plotted in
Fig. 4. At 1050"C. the SO-, concentration increased from 3.3% to A. 37. as the solids
residence time decreased from 22.5 min to 7.0 min (i.e., as the sulfatcd-sorbcnt feed
rate increased). At 1100"C. the SO. concentration in the dry off-gas increased from
3.9 to 8.47. as the solids residence time decreased from 20.3 min to 7.1 rin. At the
longest solids residence time (-20 min), the SO concentration was found to be only
slightly higher at the higher termerature. With this long reaction time, most of the
CaO was regenerated at both temperature levels and hence the SO- concentracion in the
off-gas was dependent on the CaSO,, feed rate. For the shorter reaction timo (7 min),
the SO. concentration was much higher at the hipher temperature (IIOO'C) due to the
higher rate of CaO regeneration.
Effect of Regenerating Limestone at a Temperature of 1150">J
One experiment was performed at a temperature of 1150ฐC, the highest temperature
at which a regeneration experiment has been performed. The extent of CaO regeneration
based on solids analysis at the higher temperature did not increase significantly in
comparison with data for 1100ฐC. In contrast, there was a significant increase in CaO
regeneration observed with an increase in temperature from 1050ฐC to 1100ฐC. This
suggests that regeneration at temperatures greater than 1100ฐC would not significantly
improve the performance of the regeneration step. Possibly, higher temperatures accel-
erate the sintering process and render the sorbent less reactive in subsequent sul-
fation steps.
Effect of Bed Temperature. Particle Size, and Reducing Gas Concentration on Bed
Oetluidization Velocity
Bed defluidization velocity, i.e., minimum velocity required to prevent particle
agglomeration, was studied in a statistical experiment performed with Greer limestone
and Sewickley coal. Details of the study were reported earlier.'' The effects on de-
fluidization"velocity of bed temperature (1050 and 1100ฐC). total reducing gas concen-
tration (2.5 and 5.07=). and feed sorbent particle size range (-10 +30 mesh and -14 +30
mesh) were determined. Defluidization velocities in this study ranged from 0.9 to 1.6
m/s. Defluidization velocity as a function of temperature and reducing gas concentra-
tion is shown in Fig. 5, along with minimum fluidization velocities. Bed temperature
780
-------
S.-IU* I4ป* Tlซ*. *:n
Figure 4. Experimental S02 Concentration for the
Regeneration of Greer Limestone as a
Function of Solids Residence Time and
Temperature. Pressure: 129 kPa; Fluid-
izing-Gas velocity: 1.21-1.29 m/s;
Limestone sulfur content: 7.7 wt %
O CRCUX CORRELATION
(BASED OH ROOM TEMP EXP)
O 0% REDUCING GAS
-O 1% REDUCING GAS
0 2 S>. BEOuCISO GAS
050% REIMCIKG GAS
MINIUUV rlUIOliATlON VELOCITY
I .I-..-!. ...... I , J
BOO 900 1000 1100
TEMPERATURE.*C
Figure 5. Oefluidization Velocity and Minimum Fluidization
Velocity vs. Temperature
and reducing gas concentration arc important in fixing the minimum fluidization veloc-
ity. The effect of particle size is minimal. A least-squares fit of the results is
the basis for the following equation:
Vrl = A.05 - (3.61 \ 10-3T) - 2.6R + (2.54
x 10-3TR) + (5.26 x 10-^F) (8)
where V,] = defluidization velocity, m/s
T = operating ternperat ure, ฐC
R = reducing gas concentration in off-gas, "I,
F = mean particle size of sorbent. ;im
CYCLIC SORBENT LIFE STUDY WITH TYMOCHTEE DOLOMITE
The general procedure in the first combustion half-cycle of each series of
cyclic experiments was to 3ulfatc a batch of fresh unsulfated sorbent. The sorbent was
then alternately processed in the regenerator and the combustor without makeup sorbent
for a total of ten combustion and ten regeneration half-cycles. Steady-state gas and
solid samples from each half-cycle were analyzed to assess changes in reactivity (sul-
fur acceptance during combustion), changes in rcgenerability (sulfur release during
regeneration), the extent of decrepitation, and the extent of coal ash buildup as a
function of utilization cycle.
The nominal conditions for the combustion experiments in each cycle were a 900ฐC
bed temperature, a aiO-UPa system pressure, a 1. '. CaO/S mole ratio (ratio of unsulfated
calcium in sorbent to sulfur in coal), --I?/', excess combustion air, a 0.9 m/s fluidizing
gas velocity, and a 0.9 n bed height. Changes in reactivity of the sorbent from cycle
to cycle were reflected in changes in the SO; levels in the flue gas from the corabustor.
The regeneration step of each cycle was performed at a system, pressure of 15&
kPa, a bed temperature of 1100ฐC, and a fluidized bed height of "-46 cm.
Sulfur Acceptance during Combustion
The level of sulfur dioxide in the flue gas and the corresponding sulfur reten-
tion based on the flue-gas analyses for the ten combustion cycles are presented in
Fig. 6.
781
-------
I'M
r/K-
two-
s'
400-
0!
o
Figure 6. Sulfur Retention in Bed and S02 Concentration in
Flue-Gas as a Function of Cycle Number
Sulfur dioxide levels in the gas incrca-.-.ed from --300 ppm in cycle 1 to --950 ppm
in cycle 10. This represents a decrease in sulfur retention from --88Z in cycle 1 to
o57. in cycle 10. Although there is some scatter in the data, it appears that the
reactivity of the sorbent for sulfur retention decreased linearly with increasing com-
bustion cycle over the 10-cycle experiment. The loss in reactivity for sulfur retention
is in good agreement with results reported by Zielke et al.'' in a similar cyclic com-
l>ust ion-regeneration study in which Tymochtee dolomite was used at 155 kPa.
Samples of the sorbent from each regeneration half-cycle were also tested for
reactivity in a TCA apparatus at 900ฐC. using a reactant gas of 0.3% S0> , 57= 0; , and
the balance N;>. The rate of conversion (sulfation) decreased with increasing sulfation
cycle. After the eighth sulfation cycle, however, the loss in reactivity with succeed-
ing sulfation cycles was quite small. This potential leveling-off oi reactivity was
not detected in the cyclic combustion-regeneration experiments performed in the PDU.
Sulfur Release during Regeneration
The experimental conditions and results for a representative segment of each
regeneration step are given in Table II. In the ten regeneration half-cycles, solids
residence times ranged from 6.8 to 8 1 rain. The extent of CaO regeneration based on
solids analysis varied from 67 to 30/1. There was r.o apparent loss in regenerability
during the ten utilization cycles.
The S02 concentration in the dry off-gas from the regenerator ranged fron tt.87.
to 6.1%. In the first cyclic regeneration experiment (CCS-l) , the SO_-> concentration
was diluted by gas usi;-l t.r obtain the f luidizing-gas velocity of 1.43 m/s (other
experiments were done at -.1.26 m/s). In the three final cyclic regeneration experi-
ments, the lower S(>2 cf "entrations in the regenerator reactor off-gas were a result
of the lower sulfur cr v^entrations in the sulfated sorbent (7.1-7.9 wt % S instead of
1.10% SI. Although t^. combustion steps of these cyclic experiments were performed with
a constant CaO/S moli- ratio of ->-1.5 with no virgin-sorbent makeup, the total sulfur
content of the 3uivat-ed sorbent decreased with each cycle due to lowered sulfation
reactivity of the sorbent.
Estimate of Tymochtee Sorbent Makeup Requirements to Meet EPA Sulfur Emission Limit
To estimate the relationship between sorbent makeup (ratio of makeup CaO to total
CaO) and the total CaO/S mole ratio which would be required in a continuous recycle
operation to meet the EPA sulfur emission limit, an analysis was made using the results
of the cyclic combustion/regeneration experiments along with some previously obtained
data. The greater the makeup, the higher the reactivity of the sorbent for sulfur
retention and the lower the CaO/S ratio required. Sorbent utilization data (fraction
of CaO converted to CaSO..) from selected combustion experiments were used with an
adaptation of a procedure developed by Nagiev6 for describing cyclic processes to obtain
the relationship presented in Fig. 7. The curves in Fig. 7 were developed for a
782
-------
Table II. Experimental Conditions and Results for the Regeneration
Step of the Ten Utilization Cycles with Tymochtee Dolomite
Nominal fluidized-bed height: 46 cm
Reactor ID: 10.3 cm
Pressure: 153 kPa
Temperature: 11001>C
Coal: Triangle (0.98 wt 7. S) , ash fusion temperature
(initial deformation under reducing conditions): 1390ฐC
Sorbent: -14 +30 mesh sulfaced Tvmochtee dolomite
Fluid.- Solids 02 Cone.
Gas Res. in Feed
Cycle Expt. Veloc., Time, Gas,
No. No. m/s min %
1 CCS-1 1.43 7
2 CCS-2 1.26 7.5
3 CCS-3 1.22 7.2
4 CCS-4 1.17 7.3
.5 CCS-5 1.17 7.4
6 CCS-6 1.18 7.8
7 CCS-7 1.16 7.3
8 CCS-8 1.18 8.1
9 CCS-9 1.09 7.3
10 CCS-10 1.24 6.8
?Based on off-gas analysis.
Based on chemical analysis
Analysis not performed.
26
27
36
36
36
41
33
25
36
7
7
5
1
a
i
9
4
0
of dolomite
Red.
Gas
Cone. Sulfur
in Cone . in
Off- Sulfated
Gas, Sorbent.
% 7.
2.8
3.0
3.4
2.9
3.0
2.6
2.9
3.0
3.0
3.0
samples
9.05
10.7
10.3
9.9
9.5
9.3
8.5
7.8
7.9
7.1
CaO
Major Sulfur
Compounds in Dry
Off -Gas. %
>o / /tt oO<) H-jU
L ฃ.
73/71
67/67
63/76
67/69
69/75
66/75
69/77
64/80
53/67
63/69
6
d
8
8
3
8
8
6
6
6
.5
.6
.4
.1
.8
.7
.2
.3
.1
.7
0.04
0.02
0.07
0.04
--C
0.03
0.07
0.06
0.1
0.05
COS
0
0
0
0
0
0
0
0
0
.06
.1
.1
.1
--C
.2
.01
.07
.1
.08
CS2
0.04
0.1
0.1
0.1
--C
0.1
0.1
0.07
0.1
0.09
0 02 0ซ 06 03 10"
MAKEUP CoO/TOTAl CoO. RATIO
Figure 7. Calculated Makeup and Total CaO/S
Ratios Required to Achieve 75%
Sulfur Retention as a Function of
the Makeup CaO to Total CaO Ratio.
Sulfation Conditions: Temp, 871ฐC;
Pressure, 810 kPa; Sorbent, Tymochtee
Dolomite.
783
-------
constant sulfur retention of 757<> as compared with ''70% required to meet the EPA emission
limit for the Arkwright coal used in the combustion experiments.
As an example of using Fig. 7, operating at a total CaO/S ratio of 1.5 indicates
that a sorbcnt makeup, -i, of 0.18 and, hence, a makeup CaO/S mole ratio of '-0.27 are
required for a sulfur retention of 75%. Decreasing , to 0.1 (107. makeup) reduces the
makeup CaO/S mole ratio to -0.20 (a reduction of 25%) but increases the total CaO/S
ratio required to 2.0 (an increase of 33%). In comparison to the once-through CaO/S
ratio of "-1 for 75% sulfur retention, the makeup CaO/S mole ratio of 0.2 for a cyclic
process corresponds to an estimated savings of 80% of the fresh limestone requirements.
It should be emphasized, however, that -i cannot be chosen arbitrarily. The
value of .1 will affect and will be determined ultimately by both the process flow sheet
and che system economics.
Porosity of Dolomite as a Function of Utilization Cycle
The porosity of -25 +30 mesh particles was measured by the mercury penetration
method. The cumulative pore volumes for pores ^0.4 wm and also for pores ^0.04 um in
sulfated and regenerated Tymochtee dolomite are given in Table III. It has been re-
ported by Hartman and Coughlin7 that most sulfation takes place in larger pores (-0.4
;.m) and that pores smaller than 0.4 i,m are relatively easy to plug. During sulfation
of CaO, the pores shrink as a result of molecular volume changes.
Table III. Porosity (cm;/g) of Tymochtee Dolomite as
a Function of Utilization Cycle
Cycle
No.
1
2
3
4
5
6
aPores
Pores
Change in
Porosity on
Sulfated
0
0
0
0
0
0
0
>0
120
120
120
156
144
132
4 i,
04
a/0
/O
/O
/O
/O
/O
m
um.
164b
212
140
164
Ijo
140
Regenerated
0
0
0
0
0
0
269a
2&U
252
244
238
204
/O
/O
/O
/O
/O
/O
340b
308
276
258
260
220
Re-generation (,-.)
0
0
0
0
0
0
14J'
168
n?.
OSI8
094
072
Vo.l76b
70.096
/0.136
/ 0.094
/0.104
/0.080
The porosity of sulfated dolomite was relatively unaffected by utilization cycle,
although the sulfur content decreased from -.10 wt 7. to -"7 wt "L. However, the porosity
of the regenerated dolomite consistently decreased with utilization as did the sulfur
content in the regenerated stones. The difference in porosity of the sulfated and
regenerated samples decreased from 1.0.15 cmj/g (pores ^0.4 ;im) after the first cycle
to ->-0.07 cm:/g after the tenth cycle. The porosity of regenerated dolomite decreased
with cyclic use, and thus its effectiveness as an SOj acceptor decreased.
Coal Ash Buildup during Dolomite Utilization Cycles
The extent of coal ash buildup during repeated utilization cycles was evaluated
for its effect on the SO.;-accepting capability of the sorbent in the combustion step.
Based o.. silicon enrichment determined by bulk analysis, coal ash buildups were calcu-
lated for all ten sulfation and regeneration steps. The results are illustrated in
Fig. 8. After ten utilization cycles, it was found that for every 100 g of starting
dolomite, %13 g of coal ash had accumulated in the sorbent.
Several analyses were performed confirming "he buildup of an ash layer on the
surface of the dolomite particles during repeated u/ilization cycles. Sulfated and
regenerated particles from the first, fifth, and ten;h utilization cycles were examined
for macrofeatures under a low-magnification microscope. The photomicrographs clearly
reveal that coating of even the once-sulfated stones with a crust-like layer was
beginning.
784
-------
16
a*
0
o 6
i
z
8 o
- 9 ojh/IOOg VIRGIN DOLOMITE
COMBUSTION STEP OF CYCLE
0-gosh/IOOg UTILIZED DOLOMITE
COMBUSTION STEP OF CYCLE
-gosh/IOOq VIRGIN DOLOLOMITE
REGENERATION STEP OF CYCLE
0-gcsh/IOOg UTILI/ED DOLOMITE
REGENERATION STEP OF CYCLE
8
i ป
r ป
o
n
M
11
i i
R
A-COMBUSTION STEP OF CYCLE
,0-REGENERATION STEP OF CYCLE
02468
TYMOCHTEE DOLOMITE UTILIZATION CYCLE
10
Figure 8. Coal Ash Buildup as a Function oi Utilization Cycle
A petrographic examination of unreacted dolomite and of samples from the first
and tenth cycles revealed the buildup of a vitreous crust containing magnetite (Fe;O.)
and hematite (Fez03) surrounding irost of the particles. X-rav diffraction analyses or
the crust revealed the presence and accumulation of Ca(Al-.7Fe;. ;);0: with increasing
utilization cycles.
Electron microprobe analyses performed on crosi sections of sulfated and regen-
erated dolomite samples from the tenth utilization cycle confirmed the existence of
the coal ash shell and measured its approximate composition. The racial component
concentration profiles for a typical regenerated particle after ten cycles is riven
in Fig 9 Peak concentrations of -.12 wt % Si. --25 wt 7, Fe. and -/ -ปc , Al were four.a
in the crust. The concentrations of these components in Arkwright coal ash.^which was
used during the combustion step of the experiments, are: 12 wt ป Si. 14 wt /. te. ar.c
12 wt 7. Al The above-measured concentrations of relatively major components in tne
particle crust were not in Che same proportion as in the coal ash; re/Si was --2 in t.-.e
crust and -.0.7 in the coal ash.
The sulfur concentration profile shows that this particle haJ not completely
rcEenera'ed and that very little sulfur was present in the particle crust. Scar.s oi
other reeenerated particles had sulfur profiles which indicated nearly complete regen-
eration. Therefore, the presence of a coal ash shell does not prevent sulfur tron
escaping during regeneration.
Sulfated particles from the tenth cycle were also analyzed with the electror.
microprobe The analysis for a typical partially sulfated particle is given in .- ig
10 In all particles analyzed, the formation of an ash crust was again verifiea. As
in'regenerated particles, the ash crust was enriched in calcium, suggesting the poss.-
bility that calcium diffused from the particle interiors to the ash crust.
785
-------
* 3ฐ~ A
.-20 -A
io -y
Ash Particle
Figure 9. Electron Microprobe Analysis of a Typical
Regenerated Dolomite Particle from the
Tenth Cycle
For the particle whose analysis is given in Fiฃ. JO, the sulfur concentration
below the ash crust is highest near the crust and decreases with penetration towards
the center of the particle. If diffusion through the ash crust or sulfated shell con-
trolled the sulfation reaction, the calcium adjacent to the crust would be expected to
be more fully sulfated (^.10 wt %) in a partially reacted particle. Also, a sharper
radial sulfur concentration gradient would be expected at. the reaction front. The
electron microprobe analysis of sulfur concentration in tenth cycle partially sulfated
dolomite suggests a diffusion resistance at the reaction front. Thus, the loss of
reactivity could be due to a loss of local or microporosity caused by sintering within
the dolomite particle.
Attrition and Elutriation Losses during Combustion and Regeneration
The extent of sorbent losses during the combustion and regeneration steps of
the ten utilization cycles was determined. Sorbent losses for each cycle and from
each stage were calculated on the basis of the amount of calcium in particulates re-
moved in the off-gas sys* *ms and the amount of calcium in the sorbent fed to the
reactor. The results are given in Table IV.
Although the first combustion half-cycle loss was quite large, losses during
the remaining sulfation half-cycles were reasonably small, averaging about 8% per
cycle. The lower losses during regeneration can probably be attributed to the very
brief solids residence time (-v.7.5 rain) in the reactor, as compared with the much longer
786
-------
10-
85
o>
12- !
^ Q .;.......
w 4- 7\: :.:':.:.:'i-:'
i
i .......
-..:: : ! ; : :ys
10
8-
4-
Figure 10. Electron Microprobe Analysis of a Typical
Partially Sulfated Dolomite Panicle
(PS-3) from the Tenth Cycle
solids residence time (-.5 h) for sulfacion in the combustor reactor.
CYCLIC SORBENT LIFE STUDY WITH GREER LIMESTONE
The nominal conditions of the combustion experiments were a 308-kPa system
pressure, 855ฐC bed temperature, -.177. excess combustion air, a 1.0 m/s fluidizing-gas
velocity, a 0.9 n bed height, and a constant sulfur retention of T-84% by the sorbent.
The lov.'er pressure was used to simulate the atmospheric-pressure Cluidized-bed com-
bustion concept. In this study, in which sulfur retention was held constant (in the
Tymochtee cyclic study, the CaO/S ratio was held constant), changes in sorbent reac-
tivity were reflected in changes in the CaO/S ratio required to achieve constant
retention (retention was the dependent variable of reactivity in the first cycle
study).
787
-------
Table IV. Attrition and Klutriation Losses for Tymochtee
Dolomite during Combustion and Regeneration in
the Cyclic Utilization Study
Cycle
No.
1
2
3
A
5
6
7
8
9
10
Loss During
Combustion,
wt 7.
16
A
5
3
3
6
4
7
6
A
Lobs During
Regeneration.
wt %
1.9
1.7
3.0
1.2
3.5
3.7
2.0
2.6
0.9
1.3
The nominal operating conditions during the regeneration steps for each half-
cycle were a system pressure of 129 kPa, a bed temperature of 1100'C, a fluidizing-gas
velocity of -,1.2 m/s, a total reducing gas concentration of -.3.07. in the dry off-gas,
and a fluiuized-bed height of i.A6 cm. The residence time of the sorbent in the regen-
eration reactor was nominally -*-7 min.
Sulfur Acceptance during Combustion
The cyclic calcium utilization (i.e., the percent of unsulfated CaO that vis
sulfated during each combustion half cycle) and the percent of the calcium present as
CaSO,, in the sulfated sorbent product for each combustion half-cycle are shown in Fig.
11. The cyclic calcium utilizaticn decreased from --30% during the first cycle to --9Z
in the tenth cycle. Thus, in order to maintain a constant sulfur retention of 8A*.. it
was necessary co increase the CaO/S ratio, which was --2.97. during the first combustion
cycle, by a factor of r>-3 over the ten utilization cycles. This loss of reactivity
agrees closely with the loss of reactivity observed for the experiments with Tymochtee
dolomite.
TGA sulfation experiments were also performed on regenerated samples frcr. this
cyclic experiment. The results, also shown in Fig. 11, were obtained at a reaction
temperature of 855ฐC and atmospheric pressure, using a simulated flue gas of 0.37. S0;.,
3% 02, and the balance N2. Agreement of the TGA data with the PDJ combustion results
was very good.
Sulfur Release during Regeneration
Results for the regeneration step of the ten cycles were similar to those for
the Tymochtee experiments. The S02 concentration in the dry off-gas varied from 6.1
to 8.6%, and the extent of CaO regeneration varied from A9 to 737. during the ten cycles.
The regenerability of the limestone remained acceptable for all ten cycles.
Porosity of Limestone as a Function of Utilization Cycles
Porosity effects, as would be expected, were essentially the same as those
observed during the Tymochtee series of experiments.
The porosity of sulfated lime'-tcne was relatively unaffected by utilization
cycle, although the sulfur content decreased from 1,8.9 vt % to tA.l wt 7.. The porosity
and sulfur content of the regenerated stones decreased with utilization cycle. Most
of the porosity loss was experienced in the first six cycles. This loss can be attri-
buted to the high-temperature (1100ฐC) exposure of limestone in the reducing environ-
ment of the regenerator. As a result of the loss of beneficial porosity, internal
particle diffusion and reaction with S02 were limited.
788
-------
PERCENT Of CALCIUM AS CeSO4
IN SULFATEO MATERIAL
D 6 in COMBUSTOR DATA
TGA DATA
CYCLIC CALCIUM UTILIZATION
O 6 in COMBUSTOR DATA
TGA DATA
466
COMBUSTION CYCLE
Figure 11. Cyclic Calcium Utilization for Greer Limestone.
Sulfur retention maintained at 84%.
Coal Ash Buildup during Limestone Utilization Cycles
The extent of coal ash buildup was again calculated on the basis of silicon
enrichment in the sorbent particles. It was found that every 100 g of starting virgin
limestone accumulated -<25 g of coal ash in ten cycles. In the Tymochtce dolomite
cyclic experiment, in comparison, -.13 g of coal ash was accumulated for every 100 g
of starting virgin dolomite. Arkwright coal, which was used in the cctr.justibn (sui-
fation) steps of that cyclic experiment, contains considerably less ash. 7.7 wt ?.
than does Sewickley coal, 12.7 wt 7,. In both cyclic experiments, most of the ash was
probably accumulated during the combustion steps (where the sorbent is exposed to much
more coal), rather than in the regeneration steps.
Sulfated and regenerated limestone particles from the first and tenth utilisation
cycles were examined with a low-magnification microscope for riacrofeatures. Particles
from the tenth-uti.lization-cycle sample appeared to contain m.ore ash than did the
first-cycle particles. However, not all particles were encapsulated with coal ash, as
was the case with particles from the cyclic dolomite experiments. Many or the tenth-
cycle Greer limestone particles were visually identical to first-cycle particles, which
would indicate that the ash layer thickness was not increasing and that much of the
coal ash was present as individual particles in the bulk utilized limestone. The
results suggest that the maximum ash buildup to be expected when using Gteer limestone
and Sewickley coal is ^20 wt 7ป (-v.25 g ash per 100 g of virgin limestone) in the utilized
stone.
Attrition and Elutriation cf Limestone Particles during Regeneration and b.'.fatior.
The sorbenr losses from attrition and elutriation of the limestone particles
have been determined for the sulfation and regeneration steps, and the data are given
in Table V. The limestone losses caused by attrition averaged "-2.07. in each regener-
ation step. During sulfation, the loss was '.20% in the first cycle and steadily
decreased to *ฅ/, in the final cycles. The greater attrition loss in the first sulfation
789
-------
Table V. Losses of Grcer Limestone Caused by Attrition
and Elutriation during Sulfacion and Regener-
ation Steps in the Cyclic Utilization Study
Loss = 100 ฃ
A = Ca in feed limestone (sulfated or regen-
erated) . kg/h
B = Ca in particles collected froa off-gas.
kg/h
Cycle No.
1
2
3
4
5
6
7
8
9
10
Limestone
Sultation
20.0
12.0
9.2
7.4
8 6
3.6
4.3
3.8
2.6
4.9
Avg. 8.2
Loss. %
.-^generation
2.9
0.6
_
1.3
.
2.9
2.0
1.6
2.4
1.5
1.9
step can be attributed to calcination. In subsequent cycles, the resistance of the
particles to attrition increased because of (1) sulfite hardening and (2) the partial
sintering which occurs at the regeneration teaperatuie.
The losses during sulfation were slightly hight-r in the Greer licjestone
cyclic experiment than in the Tymochtee dolcaite exj;eriaent. However, combustion
conditions differed in these experiments. The Greer li_-.estcae was fully calcined at a
system pressure of 308 kPa and a bed temperature of 853*". whereas the Tymochtee
dolomite was not fully calcined at 810 kPa and 900ฐC.
The combined losses for Greer limestone caused by attrition and elutriation per
cycle averaged ->-107.. Therefore, a fresh Greer limestone makeup rate of at least '-107.
will be needed to replenish losses. A higher nakeup rate cay be required to maintain
the S02-sorption reactivity in the fluidized bed of the boiler.
Estimate of Crccr Sorbent Makeup Requirements to Meet EPA Sulfur Emission Linit
Figure 12 shows the effect of varying the fresh-to-tocal CaO combustor feed ratio
on the amount of fresh and total Greer sorbent feed nee-ded to capture 757. of the S02
formed. The derivation for this figure is tne same for the Tymochtee dolomite cyclic
study reported here.6 For comparison, data from a cyclic experiment with Tysochtee
dolomite, Arkwright coal, and a combustion system pressure of 8 atm are also presented.
Although the curves for Tymochtee dolomite and Greer limestone have sicilar shapes,
the Creer limestone curves are higher by a factor of three ซ-n a molar basis, because
of the lower reactivity of Greer. On a mass basis, the difference is not as great, but
more Greer limestone than Tymochtee dolomite is still required. The reactivity data
for Tymochtee dolomite was obtained during sulfacion experiments at 8 ac=i instead of
1 atm. ..Hence, the sulfation reactivity data predicted for Tycochtee dolomite for an
atmospheric boiler may be high. Also, the reactivity data for Greer licฃstone was
obtained using Sewickley coal (4.3 wt 7. S) instead of Arkwright (2.7 wt % S) coal, and
thus the required sulfur retention was 837. instead of 75%. Thus, the difference
between theoe two sorbcnts is not as great suggested by the curves in Fig. 12.
REGENERATION PROCESS FLOWSHEET DEVELOPMENT
Process Flowsheet for a 200-MW FBC Process
A pr"cess flowsheet for a 203-MWe FBC process with sorbent regeneration has been
developed based on the performance of Greer licestone in the ten-cycle experiment. The
790
-------
0.2 0.4 O.6 O.B
MAKE-UP CflO/TOTAL CflO
1.0
Figure 12. Effect of Makeup-to-Total CaO/S Mole Ratio
Required for 75% Sulfur Retention. Greer
Limestone and Tymochtee Dolomite.
following base conditions are assumed for the boiler: 242 m2 (2600 ft2) distributor
plate area, 3.05 m/s (10 ft/sec) fluidizing-gas velocity, 3% oxygen in the flue gas,
and combustion of 1620 !1g/day (loOO tons/day) of Sewickley coal, which contains 4.3 we
7, sulfur and has a heating value of 28,500 kJ/kg (122aO Btu/lb).
A process flowsheet for the above boiler conditions and a fresh -orbent feed
CaO/S ratio of 1.14 is given in Fig. 13 and Table VI. The combined (virgin plus regen-
erated) limestone CaO/S feed ratio is --5.69. In the absence of regeneration, a CaO/S
feed ratio of '-3.78 would be required for Greer limestone based or. the previously
described reactivity data.
Sulfated limestone (1152 Mg/D or 1311 I/D) is assumed to be introduced into the
regenerator at lla.6 K (1550T) . the temperature in the fluid bed of the boiler. The
fluidizing gas velocity in the regenerator of 1.4 m/s is T.12% greater than the predicted
velocity required to prevenf agglomeration of sorbent with a mean size of 1500 -_m
(-1/8 in.) when it is regenerated at 1100ฐC with 27. total reducing gas in the regener-
ator off-gas. The fluidizing gas to the regenerator is assumed to be heated to 400ฐC
by waste heat recovered from the regenerator off-gas.
The coal consumption by the regenerator reactor with 843ฐC solid and 400ฐC gas
feed streams was estimated to be 60 Mg/D (66 T/D). The fuel consumption by the entire
regeneration process is obtained by adding the coal fed to the regenerator, 77 Mg/D
(85 T/D), to the sulfur recovery step, 23 Mg/D (25 T/D). and subtracting the fuel
credits for the regenerator flue gas cyclone product, 15 Mg/D (17 T/D), and the sensi-
ble heats of the regenerated sorbent, 13 Mg/D (14 T/D), and tail gas stream from the
sulfur recovery step, 12 Mg/D (13 T/D) that will be routed to the boiler. The SO;
concentration in the regenerator off-gas is predicted to be 9.7% (dry), and the gas
Distributor area for the regenerator is predicted to be 14.0 m2 (150 ft2).
Effect of Makeup CaO/S Feed Rates
Flow diagrams containing mass and energy flow streams have been obtained for
different process conditions. These calculations are intended to evaluate the effect
791
-------
MRTICULATE
REMOVAL
TO ATMOSPHERE
AIR 295ฐK .
95 Mg/D SULFATED"
SORBENT
18 Mg/0 UN8URNED CARBON
136 Mg/D ASH
OFF-GAS
653 Mg/D
AIR
HEATER
-MRTICUi_ATE
REMOVAL
23 Mg/D COAL
>3Mg,
a
0
R
REGENERATED
SORBENT 1025 Mg/D
THERMAL CREDIT-
13 Mg/D COAL
SULFATED SORBENT
1324 Mq/D 1152 Mq/p
I36OO Mg/D AIR
_
ffi HT
675 ฐK V
435 Mo/D
AIR
SULFUR
RECOVERY
(90%EFF)
50 Mg
920ฐK
ELEMENTAL
SULFUR
29 M'/0
REGEN. SORBENT
18 Mg/D
UNBWED CARBON
THERMAL CREDIT
(RECYCLED TO
COMBUSTOR)
DRAW OFF
I72 Mg/0 145 Mg/D SULFATED STONE
^ 27 Mg/0 ASH
COAL 1630 Mg/D
FRESH SORBENT 3IO Mg/D Co/S I.I
Figure 13. Process Flowsheet for a 200-MW FBC Process
Table VI. Base Conditions for Flowsheet (Fig. 13)
Greer limestone
307. C=CO,
20% Inert
Coal
Sewickley coal
2B500 kJ/kg (12250 Btu/lb)
4.3% S
' 10.0% Ash
AFBC Boiler
200 MW @ 37% conversion efficiency (9200 Btu/kWh)
Bed temperature. 1120 K (1550ฐF)
Pressure. 100 kPa (1 atm)
Bed area. 242 m2 (2600 ft2)
Combustion efficiency, 99%
Sulfur removal, 837.
Regenerator
BeJ temperature. 1375 K (2000ฐF)
Pressure, 100 kPa (1 atm)
Bed area, 14 m2 (130 ft')
bed height. 0.55 m (1.8 ft)
Gas velocity, 1.4 nt/s (4.5 ft/sec)
Solids residence time, 7 min; extent of regeneration. 657.
Total regeneration system fuel burden = 60 Mg/D (66 T/D)
(i3.6 of coal fed to the combustor)
Composition of Regenerator Flue Gas
8.9% S02
2.0% CO
19.6% CO2
8.7% H20
60.7% N2
792
-------
of makeup CaO/S feed rates (feed race of virgin limestone into the system) to the
boiler on the size of the regeneration system, on the S02 concentration in the regen-
erator off-gas, and on the fuel burden of sorbent regeneration on the boiler or power
plant.
Figure 14 shows the effect of the fresh-to-total CaO combustor feed ratio on
SOj concentration in the regenerator off-gas and on the total regeneration system fuel
burden. The S02 concentration increases quite quickly but levels out; the fuel burden
decreases quite steadily. At high fresh limestone maKeup rates, the size of the regen-
eration system decreases and the amounts of fresh limestone feed and waste sulfated
limestone increases. Consequently, the fuel required for the regeneration and sulfur
recovery steps decreases. Therefore, a decision on optimum operating conditions must
be made on the basis of economics.
The effect of makeup (fresh sorbent) CaO/S mole feed rate to the boiler on the
operating conditions and size of the regeneration system was evaluated ana is shown
in Table VII. As the makeup CaO/S feed ratio was varied from 0.81 (107. of the total
CaO/S feed) to 1.47 (30% of totoal CaO/S feed), the mass rate of sulfated stone that
has to be regenerated decreased from 1881 Mg/D (2074 T/D) to 1076 Mg/D (1186 T/D),
the sorbent waste stream (combined elutriated and draw-off soibenn) increased from
204 Mg/D (225 T/D) to 371 Mg/D (409 T/D), and the size of the regeneration system
decreased by a factor of about two. The coal required for the regeneration step de-
creased from 85 Mg/D (92 T/D) to 53 Mg/D (58 T/D). (The boiler coai. consumption is
1630 Mg/D (1800 T/D)]. The S02 concentration in the regenerator off-gas increased
fi.om 9.0% to 10.087. over the same range of CaO/S makeup ratios (0.81 to 1.47). Reducing
the power plant's fresh sorbent requirements (and its spent sorbent waste stream) in-
creases the size of the regeneration and sulfur recovery system, decreases the SO;
concentration of the regenerator off-gas (which increases "the cost of sulfur recovery) ,
and increases the fuel burden of the regeneration step on the boiler, A sorbent make-
up rate can only be chosen on the basis of an economic evaluation.
Figure 14. Effect of Fresh CaO to Total CaO
Feed Ratio to Combustor on S02
Concentration in Off-Gas and Total
Fuel Burden of Regeneration System
793
-------
Table VII. Predicted Effect of Makeup CaO/S Mole Feed Ratio
Using Greer Limestone, on Regeneration System
(200-MW FBC Boiler)
Regeneration Conditions
T = 137i K (2000ฐF)Kxtent of regeneration = 65%
P = 100 kPa (2 atm) Solids Residence Time = 7 min
Cou'.bustor
Mole
CaO/S Feed
Ratio
Mass"
Makeup Total
0.81 8
1.14 5
1.47 4
*kg feed
"includes
Includes
.06
.69
.89
Makeup
3.13
4.42
5.70
Total
26.89
18.99
16.42
Sorbent
Feed to
Regen. .
Mg/D
1881
1190
1076
Draw-
Off to
Waste.
Mg/D
204
290
371
limestone/kg S in coal.
elutriated and draw-off sorbent.
thermal credit for hot regenerated
Coal
Used
in
b Regen. ,c
Mg/D
83
65
53
sorbent.
Regen.
Bed
Area,
m2
19
14
11
Regen .
Bed ht,
m
0
0
0
.64
.55
.52
S02
Regen.
Off-Gas
(dry) ,
7.
9
9
10
.00
.74
.08
CONCLUSIONS
Reductive decomposition of CaSOt, at 1100ฐC in a fluidized bed is a technically
viable process for regenerating CaO for reuse in the combustion process. Sutticient
regeneration is obtained in a short time (a few minutes) that the regeneration reactor
can be relatively small. The S02 concentration in the off-gas is sufficiently high
that the sulfur can be recovered using commercially available processes. Data from
the cyclic studies and flowsheet studies have demonstrated that the quantity of stone
required per ton of coal processed is significantly less (-v.1/5) when the stone is
regenerated than when the stone is used only once and then discarded. Costs of the
regeneration process and the once-through processes must be compared to determine
economic viability.
ACKNOWLEDGMENTS
We gratefully acknowledge support cf this program by the Department of Energy
and the Environmental Protection Agency.
REFERENCES
1. G. J. Vogel et al.. "Reduction of Atmospheric Pollution by the Application of
Fluidized-Bee Combustion and Regeneration of Sulfur Containing Additives," Argonne
National Laboratory. Annual Report. July 1972-June 1972. ANL/ES-CEN-1005 Q973).
2. W. Swift. A. Panek, A. Smith. G. Vogel, and A. Jonke, "Decomposition of Calcium
Sulfate: A Review of the Literature," Argonne National Laboratory. ANL-76-122
(19V6).
3. J. C. Montagna et al.. "Fluidized-Bed Regeneration of Sulfated Dolomite by
Reductive Decomposition with Coal," 69th Annual AIChE Meeting, Chicago, Nov. 28-
Dec. 2. 1976.
4. J. Montagna, F. Nunes. G. Smith, G. Vogel, and A. Jonke. "High Temperature
Fluidization and Agglomeration Characteristics of Limestones and Coal Ash Particle
Systems." American Institute of Chemical Engineers Meeting, New York City, November.
1977.
5. W. Zielke et al. , "Sulfur Removal .during Combustion of Solid Fuels in a Fluidized
Bed of Dolomite," J. Air Pollution Control Assoc. 20(3), 164-169 (Ir70).
6. N. F. Nagiev, "The Theory of Recycle Processing in Chemical Engineering," Vol. 3,
International Series of Monographs on Chemical Engineering, McMillan, New York
(1964).
7. M. Hartman and R. W. Coughlin, "Reaction of Sulfur Dioxide with Limestone and the
Influence of Pore Structure," Ind. Eng. Chem. 13(3), 248 (1974).
794
-------
QUESTIONS/RESPONSES/COMMENTS
VR. DAMAN: Don't kill yourself. Thank you, John. I think
you have some questions here.
DR. VOGEL: Our contract number is W-31-lOS-ENu-. . I h*ve one
question here from Dr. Hill, Brookhaven National Lab. "In your
future work, what equipment other than a fluidized bed ao you plan to
investigate? And please discuss reasons for the choices."
We plan to investigate the use of an externally fired, rotary
kiln. And the reason for that is, v/el 1, we actually have a couple of
reasons. In our fluidized bed unit, we lose heat through the walls
which means that we have to add more coal to make up this heat loss.
This effectively reduces our S02 concentration in the off gas. If
we use an externally-fired, rotary kiln, we think we can increase the
S02 concentration in the off gas to 16 or 17 percent. The other
reason for working with the kiln is that we want to study solid-solid
reactions. The calcium sulfate and the calcium sulfide reaction and
the calcium sulfate and carbon reaction lend themselves to a kiln
operation.
The other question is from Dr. Steinberg, also from Brookhaven.
"What is the effect of limestone sulfation and regeneration on the
coal utilization efficiency for power production i:i the process you
describe?"
Well, I can't give you a simple answer. It's true you can pick
how far you want to sulfate the particle and how far you want to
regenerate the particle later and these do affect the amount
of coal that you use. I showed you a schematic of a flow sheet, and
that was just one case. There are any number of cases that are being
looked at and I can just give you a broad range of coal utilization
efficiency. In the case that you saw there, if you have a coal pile
sitting by the utility plant, approximately 3 percent of that coal
is going to be used in the regeneration process while 97 percent is
going to be used for power production. Now, you can change conditions
in the regenerator, larger regenerator or sraller, and you can affect
this coal utilization. You can use as little as one or one and
a half percent coal in the regeneration step and have approximately
99 percent for power production. And there are cases where 5 or 6
percent coal is used in the regeneration step and only 95 percent
in production. It comes down finally to an economic choice. You've
got to go through all the cases and find out which one is the most
economic.
MR. DAMAN: Thank you, John.
795
-------
INTRODUCTION
MR. HARVEY: There are three streams of waste material that pro-
ceed from the fluidized bed combustor. One is feed material, one is
intermediate fly-ash and the other of course, is final fly-ash. Now
we really haven't considered that intermediate fly-ash, have we?
We said we've got to put that back in the bed because it's unburned
carbon. But unburned carbon to us is activated carbon to others and
that has a market value. So I don't think we're ready at this stage
to say what the final outcome of all these waste materials is or will
be.
Generally, we say there are three things we can do with this
waste material. We say we can dump it discreetly. We can utilize it.
Or we can regenerate it. The only reason in tfee world we regenerate
it is because we say that this means we won't use as much of it. If
we regenerate eventually we have to throw something, away, and that
takes us back to the other two. I think it's interesting this after-
noon, of the eight papers, the score is four for regeneration, two for
dumping and two for utilization.
I'm extremely pleased to be able to introduce to you the cochair-
man of the afternoon. When it comes to v.aste utilization, the man who
will now introduce the speakers is as qualified to chair or speak
during this session as anyone I know. The man is John Faber and he is
Executive Secretary of the Notional Ash Association here in Washington,
D.C. John?
MR. FABER: I think this ii like a Las Vegas show. Bill and I
are the warmup for the dog and pony show that comes a little later.
You've already screwed up. It's "utilization and disposal" not "util-
ization and regeneration." I've got to get that in there to get my joke
in. Of course, when you talk about disposal, you want the highway
engineer to take care of it and you want the Department of Natural
Resources to take care of it and three or four other people you want
to take care of it as you produce this gold plated material. And it's
kind of reminiscent of a football game we had a couple of years ago up
in West Virginia with Pitt. Pitt came down to Morgantown and we had
things going pretty good. We had a little boy by the name of Tommy
Jones. He was just tearing them up. Every three or four plays, why,
he'd make five or six yards and then they'd pass or something and the
crowd got the bit on this thing and they'd say give the ball to Tommy
Jones, like we want to give this material to the highway man.
So this went on in the first half and we have done pretty good
but Tommy Jones was getting a little tired. Cone back the second half
796
-------
and the same thing started. A couple of pass plays and the crowd
started up. Give the ball to Tommy Jones. Well, they came back to
the huddle and all at once Tommy Jones came out and called time with
the Ref and ran over to the side and got hold of the public address
system and said "Tommy Jones don't want the damn ball."
So that's where we're at now with some of these materials we're
trying to give the highway engineer. Perhaps this afternoon we'll
find some information that will make him more receptive to carry the
ball for us.
Our first speaker this afternoon is Dr. Ralph Yang. He is
presently with Brookhaven National Laboratories. And he is a moder-
ator's dream because his biological sketch is half a line long. Ralph
T. Yang, Ph.D, 1971, from Yale University. Worked at NYU, Argonne
National Laboratories and Alcoa before joining Brookhaven. That tells
you all about him. He's a young man and when you get my age you hate
all of them so there's not much else you can say for them. There is a
correction ' would like to make. If you'll turn to page 16 in the sum-
mation of Dr. Yang's presentation on your program. The last sentence
in the program reads "five percent S0ฃ was obtained from the regen-
erator at 1100ฐC kiln temperature." That should be 1000ฐC kiln temper-
ature. Dr. Yang's presentation this afternoon will be Regeneration
of Lime-Based Sorbents in a Kiln with Solid Reductants. Dr. Yang.
797
-------
Regeneration of Lime-Based Sorbents
in a Kiln with Solid Reductants
Ralph T. Yang, James M. Chen. Gerald Farber,
Ming-Shing Shen, and Meyer Steinberg
Department of Energy and Environment
Brookhaven National Laboratory
Upton, N.Y.
ABSTRACT
1'rocesses based on apparent solid-solid reactions In a kiln-type reactor for regenerating the
lime-based sorbents are being developed at our laboratory. The specific process investigated is to
react the sulfatcd lime with fly ash, both from the f luldlzed-bed combustor (FBC). The unburr.t carbon
In the fly ash Is used aซ the reductant.
Eight-cycle sulfation-reqencratlon based on this scheme has been experimented using Greer lime
and fly ash from Argorne's 6-lnch FBC as the starting materials. The apparatus Included a rotary-kiln
regenerator and a fluidlzed-bed sulfator, both with a i:a. 10-gram capacity and made of quartz. The
kiln temperature was 1000rO in the cyclic experiments. The SC^ concentration reached the thermodynamlr
equilibrium values at slow gas flow rates. The reactivity of the regenerated sorbent did not cecay
appreciably after eight cycles; it actually tended to increase due to the impurities absorbed in the
kiln. Completion of the regeneration of the 30^-sulfated stone from FBC CGuld be reached in an hour
with a time-averaged SO, concentration of 52 from the kiln. Attrition in the kiln is much less than
in a FB regenerator. More results on the kiln regeneration are presented in this paper.
INTRODUCTION
A major advantage of the f lu idlzed-hcd combust lor with lime additives is Its ability ti> burn 1:0.1!
cleanly and to produce economically a desulfurized hot gas. Recognition of the potential of this tech-
nology has accelerated extensive efforts In research and development In this area. For environmental
and economical reasons, however, the regeneration of the lime additive.*; from the spent stone must bu
considered. The state of the art. Including the major problems involved, has been recently reviewed.
The only major process which Is currently under serious consideration Is the reductive decomposi-
tion scheme based on the Wheelock-Kent Feed?; Process.* This process is being modified and developed
further for application in fluldlzed-bed combustion by Argonne National Laboratory' and Exxon Research
and Engineering Company.^ Briefly, the process consists of fluidizlng the partially sulfated particles
with reducing gases at a relatively high temperature to produce lime and sulfur dioxide.
In the regeneration processes, two Important factors are: (1) rates of regeneration and (2)
SC>2 concentration In the pas ph.ise. Because of the high gas velocity required for fluidlzatlon In the
Argonne-Exxon process, the SO^ concentration is kinetlcally controlled, and It is substantially lower
than the thermodynamlc equilibrium values.$ For example, to produce an economically sulfur recoverable
gas, e.g., greater than 52, fInidlzatloii regeneration requires an operating temperature of 1100ฐC, and
the temperature has to be even higher for the pressurized systems. At such high temperatures, the
problems of sorbent deactlvation, bed agglomeration, attrition, etc. all become serious.
The specific process that is being Investigated at Brookhaven is to react the partially sulfated
lime with the fly ash from the f luldlzed-bed combustor in a kiln-type reactor. The fly ash contains
significant amounts of unburnt carbon which is used as the reducr.ant for regeneration. The basic chem-
istry of this process is:
CaS04 + 1/2C * CaO + 1/2C02 -t- S02 (1)
This reaction has been the basis for manufacturing sulfuric acid snd cement from anhydrite in the
European countries. Recent studies by Turkdogan and Vlncers7 and by Yang et al.8 have shown that the
amount of carbon is the controlling factor for determining the reaction product, i.e., CaS vs. CaO, and
that the reaction proceeds in two consequent steps:
CaS04 + 2C -ป CaS + 2C02 (2)
798
-------
CaS + 3CaS04 - 4CaO + 4S02 (3)
with reaction (3) as the rate-controlling step.
The equilibrium partial pressure of SO-, PSQ,< based on equation (3), as a function of tempera-
ture Is given In Figure 1. This equilibrium partial pressure limits the maximum SO, concentration in
the regeneration process as well as In some other regeneration schemes.^ Cyclic sulfation-regeneration
using coconut charcoal as the carbon reductant In Reaction (1) for the regeneration step has been stud-
lei with a TCA system. After ten sulfatIon/regeneration cycles, the reactivity of the material re-
mained the same as that of the raw lime. Also, the kinetic and mechanistic studies showed that the
rates of reaction (3) is strongly dependent upon temperature, and both steam and sodium chloride catalyze
the lime regeneration process. The catalytic effects on lime sulfation by sndlura chloride anc] by
steam'" are well known.
This paper will present our results on the 8-cycle sulfation-rei;en
-------
T.ฐC
1150 1030 950
1.0 c~
Iff1
Iff-*
1ff3
0.7
0.8
1/T(oK)x 103
09
Figure 1. Equilibrium Partial Pressure of SO2 for Lime Regeneration;
1/4CaS+3/4CaSO4 CaO + SO2
800
-------
oo-
Preheater
S02
FB Sulfator
Flow Meter
Figure 2a. Schematic Diagram of Fluidized-Bed Sulfator
Vacuum
Purifier
Sampling Bulbs
H To Scrubber and Flue
Rotameter
Furnace
Figure 2b. Schematic Diagram of Rotary Kiln Regenerator
801
-------
wet chenical and thermal decomposition methods. It was found that the weight loss of the solid samples
In N- flow In the temperature range of 1100ฐC ro 1200ฐC was correspondent (within 3/0 to the SO^ con-
tent (of CaSO.) determined by the ASTM method. In this study, the thermal decomposition method was
therefore adopted for the purpose of time saving. A high temperature TfiA system was used accordingly
to measure both the contents of CaS and CaSO, In the solid samples. About 60 mgs sample was used for
each analysis.
RESULTS AMD DISCUSSION
Regeneration experiments at temperatures ranging from 950ฐC to 1050ฐC and with flow rates (Ar)
ranging fron 10 SCCM (9.5 cm/mln) to 150 SCCM (142 cra/mln) have been made. In all these experiments
four grams of the sulfatcd stone were used. The off gases were found to be predominantly SO,,, CO., and
Ar. The temperature effect on the SO- concentration was measured using low carrier gas velocity TlO
SCCM). Results of the SO. fraction and C02 fraction at various times are Riven In Figures 3 and 4,
respectively. These figures clearly indicate that reaction (1) is a two-step reaction, and reaction
(3) Is the slower rate step, and this phenomenon Is more obvious at lower temperatures; e.g. 1 = 950ฐC.
As mentioned, the "kiln temperature" Indicated in this report was measured at the center line
of the kiln near the head of the regeneration zone. After elaborate temperature measurements, with
both thermocouples and an optical pyrometer, temperature gradients were detected In both the axial and
the radial directions of the kiln. The temperature Increased by 15 to 20ฐC alone the c-.encer line in
the gas passage direction and the temperature of the quartz wall was about 15ฐC higher than that at
the center line. Results on the SO7 concentrations should be, therefore, studied with the understand-
ing of the non-uniformity of the temperature distribution.
Figure 3 gives the history of the SO- concentration of the off gas at three kiln temperatures.
The results clearly indicate that the thermodynamic equilibrium partial pressures of SO- (cf. Figure 1)
were reached at the gys velocity of about 9.5 cm/mln.
The dependence 01 the SO concentration on the carrier gas velocity was measured at 1000ฐC.
Figure 5 shows that Increasing carrier ฃas flow rate decreases the SO., concentration. Froa the data
shown in Figure 5, the approximate curves of the overall extent of regeneration as a function of time
;it different flow rates were calculated. These results are compared with the limiting regeneration
rates measured in the TCA system (wt.: 100 mg, flow rate: 500 SCCM) in Figure 6. It Is clear that the
regeneration rate is suppressed by the SO, concentration in the gas phai.e. By increasing the carrier
gas flew rate, the regeneration rate is increased. From these figures, one can see that high regener-
ablllty with short solid residence time and with high S02 concentration In the off gas c^n be obtained
from this regeneration process. However, the operating conditions would have to be opt Ini-:..-.! with
consideration of the trade-off between the higher SOn concentration and a shorter solid residence time.
To tent the reactivity of the regenerated ma'crlal, eight cyclic sulfatIon/regeneration reactions
have been made. Sulfatlon rctes were measured In tne 30 mm fluldlzed-bed sulfator. The reactivity was
determined based on the fractional uptake of SO., after 5 hours of sulfation time. Also, in order to
compare with the reactivity of the raw sorbcnt, Greer lime was used as the starting material. The
experimental conditions for both sulfation and regeneration arc shown In the following.
Sulfation
Temperature: 850ฐC
Starting material: Creer lime (prccalcincd at 900ฐC)
weight: 10 grams
size: 16/20 mesh
Superficial velocity: 5 ft/sec
Gas composition: SO?: 0.252, 0-: 5%, N- balance
Time: 5 hrs
Regeneration
Temperature: 1000ฐC
Material: Sulfated Creer -I- coal ash (size 200/270 mesh)
Time: 2 hours Flow rate (\r): 100 SCCM
After each sulfation/regeneration run, a small portion (about 60 mg) of the sample with size
16/20 mesh was taken out for analysis of the contents of CaSO, and CaS. Also, to assure having high
extent of regeneration, the regeneration period was kept for two hours. The results of the SOj content
In the solid sample for the eight cycles are given in Figure 7.
802
-------
0.3
0.2
s
0.1
10
15 20
Time, min.
25
30
35
40
Figure 3. Temperature Effect on the Partis) Pressure of SO2 in the Kiln
Regenerator; Gas Flow Hate (Ar), 10 SCCM. Tc = 1050-1070" C (D).
Tc = 1000 1020 C (D). Tc = 950 970 C (0). Tc was measured along
the center of the reactor tube. 4 grams of suHated Greer lime
(16/20 meih) from Argonne'c FBC were used.
5
<
CM
o
J.I
0.6
0.5
0.4
0.3
0.2
0.1
10
15 20
Time, Min.
25
30
35
Figure 4. Partial Pressure of CO2 from the Kiln Regenerator. Gas Flow
Rate (Ar): 10 SCCM. Tc 1050-1070^C (G). Tc -1000 1020'C
(A). Tc = 950-970'C (0). 4 grams of sulfated Greer lime
(16/20 mesh) from Argonne's FBC were used.
803
-------
i r
I r
>0 15 ?O ?b
Figure 5. Gas Flow en the SO2 Concentration at Tc * 1000-1020ฐC from
the Kiln Regenerator, Flow Rates (Ar): 10 SCCM (A). SO SCCM
-------
30%
20%
5
8
10%
I
IS 'R 2s 2R ป;> 3R 4S 4j} 5s 5R 65 6R 7S 7R 85
Number cf Cycle*
Figurr 7. Extents of Sutfation and Regeneration of Greer Lime from the
Cyclic Sulfator (S) and Regenerator (hi ConditioftS are Shown
in Text.
805
-------
This figure shows that for all the regeneration cycles, very high extent of regeneration can be
reached (88-1007,). It was also found that the regenerated stone did not contain any CaS. This indi-
cated that the regeneration reaction was completed in all the cycles. This high extent regeneration
is very encouraging for this regeneration process.
It should be emphasized here that the 2-hr regeneration time used was only to ensure complete
regeneration of lime so we could study the sulfation char.icteristics in a more meaningful way. In
practice, however, the regeneration time can be substantially shortened. Kor example, based on the
Integrated .mounts of SO evolved, over 802 regeneration was accomplished in 1/2 hr at n fiow of 50 SCCM.
o
The regeneration results at 950"C perform^u In TCA Indicated the existence of CaS in the solid
absorbent after 4 hours of regeneration. The difference in the solid content, compared with the re-
sults obtained in this study, is very likely due to the difference in reaction temperatures. Since
reaction (3) was found to be strongly temperature dependent with about 62 Kcal/mole activation energy,
by Increasing temperature from 9501 C to 1000ฐC, the reaction rates become 2.7 times higher. Thus, the
ซccontl step reaction could be completed in a shorter time. Other factors such as the extent of sulfa-
tion of the sorbent and the sources of carbon material may also nffect the reaction rates.' Snyder ct
a!.11 have found that by using H, is the rcductant at .:bove 1100ฐC, the only product was CaO whereas
below 900ฐC. the product was alPCaS. These results are in line with the above discussion and the
fact that reaction (2) has a lower temperature dependence, as observed by Turkdogan and Vlntcrs.
Figure 7 also shows a slight decrease In SO, absorption ability of the regenerated lime in eight
cycles. To further test the reactivity of the regenerated sorbent, portions of the regenerated samples
were sulfated la a TCA system. These results are shown in Figure B indicating approximately 102 de-
crease in reactivity. The detrimental eifect of this temperature on the reactivity of the lime sorbent
Is believed to be lower than the temperature at IIOO"C. For example, the dolomite regenerated at 1100"C
was found to he 'MX. lower than the material at 900ฐC.
Attrition experiments were performed to compare the strength of the regenerative lime with that
of the fresh lime. The samples compared were the regenerated lime after eight cycles, which contained
3 wt Z of SO.j, and the fresh lime precalclned at 900"C, both 16/20 Tvler mesh in particle size. They
were fluldizcd with nitrogen gas In a fluidIzed-bed for 5 hours at 850"C with 5 ft/sec superficial gas
velocity. Results of the size distribution of the attrituted samples are shk>wn in Table I. It Is seen
that the regenerated lime had a much higher strength than the calcined lime. Tils is because the re-
generated lime contained SO, and some silicates as will he shown later; both would Increase the strength
of the material. Also, the regenerated material had been treated at higher temperatures (iOOOฐC) than
the fresh lime (at 900ฐC). The strength of the regenerated material would tnus he higher. Mere direct
comparisons on the attrition characteristics arc being made.
Besides the regeneration reactions, many other reactions did indeed also take place In the kiln
regenerator. The reactions between CaO and the various minerals in the coal ash are well known to the
material scientists. For example, silicates, ferrites, ferrates, aluminates, ferro-si1icatcs, etc. can
all be formed at below 1000ฐC. We have found earlier that Ke-iO- catalyzes the sulfation reaction of
CaO, because it catalyzes the oxidation of SO., to SO . These reactions arc under investigation in
the authors' laboratory. Only some preliminary results will he presented hero.
The sulfaclon rates were compared of a raw Creer line and a kiln-regenerated lime (1000ฐC kiln
temperature) from the partially sulfatcd Crccr lime from Argonne's KBC. The kiln regenerated lime
showed higher reactivity and which absorbed about 30i more SO, at 150 minutes than the raw lime.
X-ray diffraction analysis and atomic absorption methods were used to determine the concentra-
tions of fhe major Impurities (SI, Fe, and Al) In the limestone particles. For these analyses, samples
of Greer limestone particles 06/20 mesh) were screened after each sulfatIon/regeneration experiment and
the free coal ash and finer particles were sieved out. X-ray diffraction showed that calcium silicates
In Creer limestone are present as B-dicalcium silicate form. It has been reported In our laboratory
that B-dlcalclum silicate Is a reactive form towards SO-.^ The B-dicalclua silicate was not present
in Argonne's once sulfatcd Greer limestone while It was present In an appreciable amount as detected by
x-ray after regeneration of this .sample In our kiln. Also found in all the Creer limestone samples was
alpha quartz. From the x-ray diffraction Intensities, the amount of alpha quartz also Increased In the
kiln regenerated sample. Results of the atomic absorption analyses showed that SI, Fe, and Al all
increased in the kiln regem-ratIon. The concentrations were Increased from: 14.502 Si; 0.8*7, Fo, and
1.90% Al In the Argonnc sulfatcd stone to: 14.92 SI; 1.162 Fe, and 2.142 Al In the kiln regenerated
Argonne stone. Although these analyses were preliminary In nature and may not be representative because
only 100 mg sample was used for an analysis, possible catalytic effects due to the minerals may be the
cause of the observed Increase In the sulfation reactivity of the kiln regenerated stone.
806
-------
0.4
0.3
0.2
0.1
20
40
60
80
100 120
Time. min.
140
160
Figure 8. Comparison of the Sulfation Rates between Fresh Lime and
Regenerated Lime at T = 900 C. SO2- 0.5%. 02: 5%. N2- bal.
Sample: Precalcined Lime (Q). 2nJ Cycle Regenerated Lime
(A). 4th Cy-ie Regenerated Lime (0). 8th Cycle Regenerated
Lime (VI.
180
200
Table I. Size Distributions of the 8th Cycle Regenerated Lino
nnJ Fresh Line after 5 Hours, of" Fluldlzat Ion at 850ฐC
Size
(Tvlcr Mesh)
16/20
20/24
24/f>0
60-
Wclsht Pcrcent.-iBe
Regcnera'.'d lime
86.1
13.1
0.3
O.S
fresh lime
30
20.8
14.7
34.5
807
-------
CONCLUSION
Results oT regeneration of the sulfdted lime in a kiln using fly ash as the reductants have been
presented In this paper. Because of the low gas flow required In this process, high SO. concentration
ran be obtained at relatively low temperatures In comparison with the fluid tz.it Ion regeneration process.
The high rrgcnerabilIty In a reasonably short solid residence time, the less detrimental effect of
temperature on the reactivity of the regenerate*! sorbent and potentially less solid attrition make the
kiln regeneration process look promts In,;. The kiln regenerator may also serve the function of the
c.irhon burn-up cell and thus replaces It.
ACK.NQซLEIX;Mfc:.VTS
Discussions and guidance provided by Or. Andfej Macek of the US Department of Ener<;y are ap-
preciated. Argonne National Laboratory kindly supplied the materials for our experiments. The able
and skillful assistance from Messrs. Frank B. Kalnz and Jacob Pruzansky Is gratefully acknowledged.
REFERENCES
1. A. A. Jonke and U. A. McCurdy, paper No. 62a. AIChE 70th Ann. Mtg., N. Y.. November 16, 1977 (1977).
2. T. D. Wheelock and I). R. Hoy I an, "Sulfurlc Acid from Calcium Sulfate", In "Sulfur and SO- Develop-
ments". Chen. Er.:;. Prog. Tech. Manual, AIChE, N. Y. (1971).
1. G. J. Vogcl et al., "A Development Program on Pressured Fluldlzed-Bed Combustion", Annual Report,
Argonne National laboratory, Argonne, 111.. ANr./ES-CF.N-1016 (1976).
4. R. C. Hokc. ct al.. "Studies of the Fl'iidlzcd-Bed Coal Combustion Process", Annual Report, Kxxon
RMป. Linden. N. J., EPA-hOO/7-76-01I (1976).
5. L. A. Ruth, Proc. 4th Intern. Conf. KBC, Mitre, McLean, Va., Der. 1975 (1975).
6. W. Q. Hull, F. Sr.hon, and H. /irngibl. Ind. Eng. Chem.. 4ฃ. 1204 (1951).
7. E. T. TiTkdogen and J. V. Vlnters, Trans. In-.t. Mining & Metallurgy, In press (1977).
8. R. T. Yang, M. S. Shcn, and M. Steinberg, "A Regenerative Process for Fluldizcd-Bed Combustion of
Coal with Lime Additives". BNL Report 22782, Brookhavcn National Laboratory. Upton, Now York (1977).
9. S. Ehrllch. patent disclosure to Office of Coal Research. Pope, Evans and Robblns, Inc. (1968).
10. R. T. Yang, P. T. Cunningham. W. I. Wilson, and S. A. Johnson, Advances In Chemistry. U7, 149 (1974).
11. R. B. Snyder. W. I. Wilson, C. J. Vogcl, and H. \. Jonke, Proc. 4th Intern. Conf. FBC, Mitre. McLean,
Va., Dec. 1975.
12. R. T. Yang, et al., "Regenerative DesulfurtzatIon of Hot Combustion and Fuel Cases", Quarterly
Reports Nos. 5 and 6, Brookhaven National Laboratory, I'pton, New York, April 1-September 30, 1977.
13. R. T. Yang, M. S. Shen, and M. Steinberg, Env. Sc. Tech.. in press.
808
-------
QUESTIONS/RESPONSES/COMMENTS
'R. FABER: Thank you, Dr. Yang. We do have time for one or two
quick questions if somebody has one. Yes, right there.
MR. HUBBLE: Bill Hubble, Argonne. Ralph, did you measure the
concentration of carbon in your ash?
DR. YANG: Oh, yes. I'm sorry. I didn't mention that. The fly-
ash was supplied by Argonne. It was from Argonne's six-inch fluidized
bed combustor. And the carbon was 13 percent in the fly-ash.
MR. HUBBLE: Did you put enough fly-ash in there that ;'ou had
enough carbon to reduce the sulfur?
DR. YANG: We used a molar ratio of two calcium sulfate ":o one
carbon in the feed mixture.
MR. FABER: Thank you, Dr. Yang.
809
-------
INTRODUCTION
MR. FABER, CHAIRMAN: Our second speaker this afternoon is
Dr. R. A. Newby. Dr. Newby spoke this morning and you heard his
credentials then. Dr. Newby is with the Westinghouse Research and
Development Center in Pittsburgh. The title of his afternoon presen-
tation is The Evaluation of Sorbent Regeneration Processes for AFBC
and PFBC.
810
-------
Evaluation of Sorbent Regeneration Processes
forAFBCandPFBC
R. A. Newby, S. Katta, D. L. Keairns
Westinghouse R&O Center
ABSTRACT
Projections of the economics of regenerative fluidized-bed combustion power
plants (atmospheric-pressure and pressurized boilers) have been developed on the basis
of current estimates of regeneration system performance. Economic comparisons with
fluidized-bed combustion power plants operated with once-through sorbent systems and
with conventional coal-fired pover plants using limestone vet-scrubbing are presented.
Regenerative FBC performance requirements for economic feasibility are projected and
critical development needs are discussed.
INTRODUCTION
Developmental facilities for fluidized-bcd combustion power generation are
presently based on once-through sorbent operation. Although research facilities are
.iddressing the area of sorbent regeneration, the technical and economic feasibility of
regeneration is not yet known. Regeneration of sorbent for the purpose of reducing
the rate of spent sorbent production faces trade-offs in the areas of economics,
environmental impact, plant complexity and reliability, and general technical
performance.
An assessment of the economic potential, the technical feasibility, the problem
areas, and the development requirements of the one-step regeneration ..rocess (reductive
decomposition) as applied to AFBC and PFBC is presented. Capital and energy costs of
once-through and regenerative AFEC as a function of the process sulfur load are
projected.
The economics and performance of two regeneration processes that function to
regenerate sorbent produced in PFBC have been previously reported.1 These are a one-
seep process (reductive decomposition) operated at 1000 kPa pressure and the same one-
step process operated at 1000 kPa pressure. An update of that work is presented in
this report. The projections reflect current performance expectations and revised
component cost data.
Details of these studies are presented in an EPA contract report, March 1975
(EPA 600/7-78-039).
CONCLUSIONS
The following conclusions have been drawn from a technical and economic evalua-
tion of sorbent regeneration processes for AFBC and HFBC:
An integrated regeneration system for AFBC or PFBC has yet to be demon-
strated. Information on critical performance factors for commercial oper-
ation is not yet available.
The sulfur recovery system is the dominant subsystem in the regenerative
process. The pressurized regeneration (reductive decomposition) results
in low SO? concentrations (1-2 v/o) requiring significant amounts of coal
for reductant. complex energy reco" and hence has a substan-
tial advantage over the ptessurizcd regeneration^
In the case of atmospheric-pressure regeneration applied to PFBC, the major
uncertainty lies in the solids transport system.
Presented at the 5th International Conference on Fluidized-Bed Combustion, December 12-
14. 1977.
811
-------
Assuming 2 and 12 v/o of SO? for pressurized (PR) and atmospheric regenera-
tion (AR), respectively. Ca/S makeup ratio of 1.0, a process sulfur load
of 0.003. and sulfur recovery in the forra of elemental sulfur, the following
capital .?osts for regeneration (635 MK'e plant) in terms of $/kW have been
projected :
AR for
AFBC
32.2
PR for
PFBC
66.8
AR for
PFBC
57.2
Using the same bases as above, the following energy costs in terms of
mills/kWh have been projected:
AR for
AKBC:
2.9
PR for
PKBC
4.9
AR for
PFBC
3.55
For atmospheric regeneration applied to AFBC, the following energy costs in
terms of mills/kWh for a process sulfur load of 0.025 have been projected
for three different C;i/S ratios, in the case of regeneration, and for a
Ca/S ratio of 2.2, in the case of once-through option:
Regenerative Option
Ca/S Ratio
0.2 0.6 1.0
Once-through
Option
1.9
2.2
2.5
1.92
If sulfur is recovered as sulfuric acid rather than as elemental sulfur, the
capital cost of the regeneration process can be reduced to the following
extent:
AR for
AFBC
97,
PR for
PFBC
247,
AR for
PFKC
117,
If a sulfur recovery process is developed specifically for regeneration,
the regeneration potential may be considerably improved. The scope and the
need for process innovations in sulfur recovery are evident.
The only regenerative PFBC power plant that is economically attractive, as
compared with a conventional power plant with limestone wet-scrubbing, is
based on the low-pressure reductive decomposition.
The overall environmental performance of the low-pressure reductive decom-
position is superior to the other regeneration processes.
The once-through sorbent operation is superior to the regenerative opera-
tions in all environmental aspects except for the quantity of spent sorbent
produced.
812
-------
RECOMMENDATIONS
The following recommendations are made after reviewing the present technology
applicable to regeneration and the development effort that has been carried out so far:
Studies should be continued on particle attrition, sorbent deactivation due
to the presence of fly ash or due to sintering, particle aggloir.eratic.i due
to eutectic formation and gas-particle contacting i;i fluidized beds.
Studies should be continued on the change in activity and the regenerability
of the sorbent with repeated cycling, and the separation of sorbent and ash
in the regenerator.
The raxiir.u'. percent of S02 in the regenerator effluents that c.-n be achieved
in a continuous operation of the combustor-vegenerator system at commercial
operating conditions needs to be demonstrated.
Development of sulfur recovery processes suitable for different regeneration
schemes under consideration should be initiated.
Exploratory work should be conducted on new schemes, such as the production
of sulfur vapor rather than sulfur dioxide in the regenerator.
The low-pressure reductive decomposition for PFBC appears to have greater
potential than pressurized regeneration. The sorbent circulation system
for the low-pressure regeneration for PFBC, the area of greatest uncertainty,
should be evaluated in greater detail.
The present developmental effort on regeneration should be directed to
correspond to the operating conditions envisaged for commercial operation.
Much of the past effort ,on regeneration appears to have no relevance to
industrial practice.
A regeneration system modeling study is needed to assess the regeneration
technology and process economics in greater detail and to permit the assess-
ment of the experimental data that is being accumulated.
Development of optimum methods for disposal/utilization of the s.icnt sorbent
that meet environmental constraints is necessary.
REGENERATIVE AFBC
The one-step reductive decomposition of calcium sulfate is the most attractive
regeneration process proposed for AFBC. An evaluation was completed to develop per-
formance projections, cost estimates, and critical development rcquircnents.
Regeneration Concept
The following reaction takes place in the one-step reductive decomposition of
CaS04:
H,l fH-,01
oh CaO+\clj+S02 (I)
The undesirable competing reaction involving the formation of CaS al.io occurs:
CaS04 + ^Sl^^CaS + 6tcOฐ} <2)
An oxidizing zone may Uu provided in the regeneration vessel to convert CaS to CaSO^:
CaS + 20, v ^ CaS04 (3)
Process De-script ion
In ;:he regeneration process coal is introduced into a fluidi.-ed-beti rcpcr.erator
for in_ situ partial combust icn to provide the reducing gas and the heat necessary for
the reducTTon of CaSO^ to CaO (Figure 1). The regenerated sorbent is returned to the
fluid-bed boiler, where fresh sorbent will be introduced to make up for reduced activity
and losses of the sorbent by attrition and elutriation. Part of the sulfated sorbent
is discarded for disposal or utilization. The regenerator off-gas (containing about
813
-------
Pig. I68JB!>2
00
Utilized
Sorbent
From
Boiler
Air
Steam
Sulfur Recovery Plant
Resox and Beavon Processes
C.W.
AMAM
Spent Stone
Cooler/Conveyor
S. S. D.
C.F. -Coal Feeding System
C.W. -Cooling Water
S. S. D. - Spent Stone Disposal
U.S. -Utilized Sorbent
R. S. - Regenerated Sorbent
Figure 1. Atmospheric One-Step Regeneration Schematic Flow Diagram of One Module
-------
12 percent S02 at a temperature of 1100CC) passes through primary and secondary cyclones
and then exchanges heae with the incoming air to the regenerator before being processed
in a sulfur recovery plant for the production of elemental sulfur.
Performance Projections
Design specifications are listed in Table I. Material and energy balances are
given in Table II and Figure 2. The single most important variable of the process is
the concentration of S02 in the regenerator effluent, which depends on the type of fuel
used in the regenerator, temperature. pressure, heat losses, and the change in the
utilization of calcium across the regenerator. The concentration of S02 in the regen-
erator effluent was estimated to be about 12 percent. The effect of various factors
on the maximum concentration of SO? that can be achieved has been studied with mate-
rial and energy balance considerations in mind.
Table I. Design Specifications and Assumptions for AFBC
Design Conditions:
Boiler coal rate 240.408 kg/hr (635 MW)
Basis for boiler design Previous Wcstinghouse Study^ '
Sorbcnt type Dolomite
Process sulfur load 0.026
Sorbent disposal Before regeneration
Plant capacity factor 707,
Sulfur recovery Elemental sulfur by the RESOX
Process (Foster Wheeler)
Number of regenerator modules 4
Operating pressure and temperature of 101 kPa and 870ฐC
AFBB
In situ partial combustion of coal in
tHe regenerator
Design Assumptions:
Regenerator temperature 1100ฐC
Dolomite makeup rate 1 mole Ca/1 mole S
Dolomite utilization after 10%
regenerator
Percent S02 in the regenerator 1P%
effluent
No CaS is formed
815
-------
Table II. Heat and Material Balances for AFBC
Stream
Ho.
1 Coal to regenerator
2 Air to regenerator
3 Utilized sorbenc
4 Regenerated sorbent
5 Regenerator off-gas
6 Air to heat exchanger
7 Regenerated off-gas
to SRP
8 Coal to SRP
9 Sulfur (807. recovery)
10 Tail -gas from SRP
11 Waste stone to cooler
12 Waste stone for
disposal
13 Sulfur M00%
recovery)
Temp. (ฐC) /Pressure
(k?a)
93.3/137.8
704.4/158.5
871/103.4
1093.3/103.4
1093.3/137.8
121/165.4
537.8/130.9
93.3/137.8
121/103.4
148.9/110.2
871/103.4
93.3/103.4
148.9/110.2
Flowratc ; Enthalpy !
(kg tnoles/hr) i (kJ/kg molo) ! Comments
5602 kg/hr 86.5 kJ/kg .
1
1207 21.074
-,-,, ke moles of Ca , , 00/ i MgO - 50%, CaSO/ - 17.5%
//J * Kr uu.oa* Ca0 . 32 5% "
,,, kg moles of Ca ,, ,, | MgO - 50%. CaSO, - 5%
//J hr ^ILI,ซI i | Ca0 . 457<
155C i 39.890 i CO, - 18.6%. H?0 - 7.87,
I SO; - 12.47., Nj - 61.2%
l.?07 2,888 '
1553 18,319 j
3549 kg/hr ; 86.5 kJ/kg i
155 j j
1636 : i
3,, kg moles of Ca . nQ &s, ''
,,, kg noles of Ca ;
hr i '
33.7
'
00
(ป
en
-------
00
To Stack
Own. 6392A28
S7R.
Beavon
Process
To Stack * Sulfur j
S.R.
Resox
Process
Sulfur
Coal
Air
Make-Upti
Sorbent
Fluid-Bed
Boiler
870ฐC
101 kPa
Coal
S.R.
Spent Stone
Disposal/Utilization
- Sulfur Recovery
Figure 2. Atmospheric One-Step Regenerative Process Flow Diagram
-------
The required calcium/sulfur nol.ir feed rario depends en the activity of the
sorbcnt in the boiler a-.ci the regenerator and ihe rate of circulation of solids
between the two processing steps. For atnospheric-pressurc operation in on::e-through
systems. Ca/S makeup ratios of 2.8/1 and 2.2/1 have beer, projected for calcined lime-
stone and dolomite, respectively, for a temperature of 616*C.2 Pope, Evans and
Robbins^ estimated a ratio of 1/1. while from recent ber.ch-scale experimental data
Argonne^ projected a ratio of about 0.35.
The combustion efficiencv in AFBC can be expecteii to be about 90 percent wi'h-
oul a carbon burn-up celt, the inefficiency result ine mainly from the carry-over or
carbon fines. 6 In an atmospheric regenerator higher carbon losses can be expected
because of (a) the reducing atmosphere, (b) the absence of any internals ant! (c) Che
reaction of carbon with Ci>> . resulting in the f oroat ion of CO near the top portion of
the bcJs and its subsequent loss through the regenerator off-pas. The cor.pensat ins
factors in favor of regeneration are the highest operating temperature and lower
fluidixation velocity. The heat losses of the regenerator have been estimated 10 he
in the ranye of 0.5 to 1 percent . Tlie energy requirerx-iit cf the reeenerator for a
635 MW plant has been estimated to be approximately i percent of the energy requirement
of the holler. Assuming the carbon loss in the regenerator is about 15 percent, tnis
represents about 1/2 percent of the fuel input to the boiler.
Cost F.st imate
The process investment cost .-,nd the- energy cost as a function of the process
sulfur load are shown in Figures 3 .-ind 4. respectively, for the following basis:
Cost corresponds to the er.J of 1976-
Capital charges plus operation and maintenance at 20 percent of the total
cost.
Contingency at 20 percent and contractor fees at 3 percent of the base
cost.
No interest during the construction period is incluaed.
Co.il at S20.00 per Mg and dolomite at S5.00 pc-r Mp.
Waste stone disposal cost at. S3.00 per ^g
e Klect.ricit y at 23 mills/kWh
Process water at SO. 10 for 3.8 m1 ป
No credit for recovered sulfur
70 percent capacity factor (6132 hours of operation in a v-ar).
The sulfur recovery element is by far the nost expensive system For a process
sulfur load cf 0.025. its cost forms more than 60 percc-n: of the total investment cost.
The sorhent circulation element: is the least expensive o:~ the three elements. The
effect of I'SI. on the cost of the sulfur recovery clciscnt is subst ant ia 1 ly higher thar;
on the cost of the regeneration v It-mem or the sorbent circulation ..lenient. Hence, it
is desirable to have us 'ow a load as possible on the sis! fur recovery element.
Assessment
The tost: of the rcRcnt-r.it ive process is ba^cii on the assumption that a con-
centration oi' SO? of about 12 percent can be obtained. This has yet to be
demonstrated experimentally.
The sulfur recovery system is the dominant subsystem in the regenentive
process.
Comparison of the oncc-thrrxigh option with the rop.enerativc option shows
that the process investment cost is about 20 percent higher and the energy
cost, about 4 percent higher for the latter option. This comparison is
based on the assumption that the spent stone does not need further
process ing.
If sulfur is recovered as sulfuric acid rather than as elemental sulfur,
the capital cost of Lhc regenerative option can he reduced by aboui
10 percent.
The regenerative option might become competitive with the once-t.hrouah
option if the makeup ratio of Ca/S can be reduced to about 0.2 for a sorbent
cost (fresh stone plus disposal) of about S8 per Me.
For a makeup rate (Ca/S) of 1.0. regeneration is liV.ely to break even at a
sorbent cost (fresh stone plus disposal) of S12-16 per ton.
818
-------
Cu'.f 6871li-t
4?
36
^ 30
. 635 MW Plant
cone. oป S0-=ai2
FSL - Ib of s handled by Regenerator
per ID of coa! to comnuslor
0.01 a 02
0.03 0.04
Process Sulfur Load
0.05
0.06
Figure 3. Capital Cost of Regenerative Process as a Function of PSL
REGENERATIVE PFI'.C
The economics and performance of two reductive decomposition schemes, one oper-
ated at about :' : ,-:. (10 .ii-osphcrcs) and one at about 100 to 200 kl'a (1 to 2 atmos-
phc-res) were t-st irvit oil. The tlcsir.ns are conceptual in r.al;:re anJ were not based on
sensitivity analysis or optimisation.
Bjปs_l_s_ of _Evalu."t ion
The power plant basis listed in Table III has boon applied in the assessment.
The process sulfur load, reflecting in part the sulfur content of the coal, is varied
from 0.01 to 0.06. Important process characteristics are yxivcn in Table IV as a func-
tion of the process sulfur load.
The specific process options were selected on the basis of result* of previous
engineering assessments and are presented in Table V. Selected regeneration process
opcrat ini; conditions and projected perfornance levels arc sunsnnri::cd in Table VI.
Sulfur dioxide concentrations of 1 and 2 v/o frora the regenerator are ex.-.-ined for the
819
-------
Curve 687042-A
E
I 3
o
I
635 MW Plant
Cone,
of SCL = 0. 12
PSL - Ib of S Handled by Regenerator per
Ib of Coal to Combust or
0.01 0.02
0.03 0.04
Process Sulfur Load
0.05
0.06
Figure 4. Energy Cost of Regenerative Process as Function of PSL
1000 kPa reductive decomposition process because the achievable level for this critic.il
performance factor has nut been demonstrated. A level of 10 v/o is assumed for the
low-pressure reductive decomposition process. The combustor operating conditions are
assumed to result in calcination of the dolomite. A dolomite makeup rate (Ca/S ratio)
of 0.5 to 1.0 moles of calcium per mole of sulfur fed to the combustor is assumed.
Process Performa-.ce Projections
Some key performance characteristics of the PFBC regeneration systems evaluated
are summarized in Table VII as a function of the process sulfur load. Auxiliary power
requirenents (for the sulfur recovery pror-ess, for the compression of air and stack
gas and for sorbent circulation), the rate of coal consumption for regenerator reduc-
cant. the rate of methane consumption for sulfur recovery, and the rate of steam consump-
tion are estimated. The regeneration processes are large power and fuel consumers, -ir.ci
the process designs must be concerned with maximum energy recovery. The energy content
of the regenerator product gas is used to provide the regeneration process auxiliary
power requirements. No energy is exported from the regeneration process to the plant
power cycle in this evaluation, although this may be called for in an optimized power
plant.
820
-------
Table III. Power Plant Basis
Plant Capacity - 635 ^e (based on once-through sorbent power plant performance)
Plant Heat Rate - 9040 kJ/kWh (8570 Btu/kWh) (based on once-through sorbent
performance).
Combustor Excess Air - 17.5%
Conbustor Pressure - 1000 kPa (10 atmospheres)
Process Sulfur Load - 0.01 to 0.06
S02 Control - Meets EPA standard of 0.5 kg S02/GJ (1.2 Ib S02/106 Btu).
Sorbent Type - Dolomite
Layout - Four pressurized boiler modules, four parallel regeneration trains, single
sulfur recovery plant.
Spent Sorbent Processing - None; sorbent is disposed of following regeneration
Sorbent Circulation System - Dilute pneumatic transport
Sulfur recovery tail-gas - Incinerated and exhausted
Table IV. Process Sulfur Load
Coal Sulfur.
w/ob
Combustor Sulfur Removal
Efficiency, ^c
Sulfur Production Rate. Kg/hr
Sorbent Circulation Rate, Mg/hre
Process Sul fur Load3
0.06 | 0.03 j
7.2 4.0
93 85
282 141
150 85
0.
1
65
42
30
01
.8
Defined as Ws (:i-mXpX where Ws is the sulfur content of the coal (weight fraction).
n is the boiler sulfur removal efficiency (fraction), m is the boiler Ca/S makeup
ratio,and Xs is the fractional utilization of the sorbent material following
.regeneration.
"Based on values of m ป 1.0 and Xs = 0.1.
cfiased on satisfying SC^, emission standard of 0.5 kg/GJ and a recovery efficiency
for the sulfur recovery process of 90"",.
"Based on coal heating value of 3,000 kJ/kg (13.000 Btu/lb).
eBased on a dolomite. 30 percent utilization before regeneration and 10 percent after
regeneration.
821
-------
Table V. Selected Process Options
Reductive Decomposition Processes
Kcducl.'ini K-''s f.enerat. ion - in situ wit.h
Fuel for rcduc'nuL - coal
Sulfur recovery form - L-lcnvcni.il sulfur and sulfuric acid evaluated.
Sulfur recovery process - Allied Cher.icnl process with methane reductant (sec
Section 4)
Table VI. Operating Conditions and Performance Projections
Reductive Decomposition |
, i i
Regenerator Pressure. kPa
Regenerator Temperature. ฐC
SO., Mole Percent a^.o Produced
Sulfur Recovery Kfficiency, ~
Dolomite Utilisation in Boiler. 7,
Dolomile Utilisation ai'ter
Rej-enerat ion . 7.
Dolomite Makeup Rate. Ca/S
KIuidix.it ion Velocity, m/sec
Boiler Conditions
Calcium Sulfide in Sorhcnt . 7,
Pressur i/ed
1000
1100
1-2
90
30
10
0.5-1 .0
1.5
Calcining.
0
i Atmospheric Pressure
150
1050
10
90
30
10
0.5-1 .0
1.5
Calclninc,
0
Table VII. Perforisance Projections
Process Sulfur Load
Auxiliary Power, MWc
Coal Consumption. Percent
of Boiler Coal Input
Methane Consumption, GJ/hr
Technical Uncertainties
Reductive Decomposition
17, S02
0.06 0.03 0.01
70 37 15
69 35 12
272 136 il
Energy recovery,
sulfur recovery
27. S02
0.06 0.03 0.01
41 22 10
34 17 6
272 136 41
Energy recovery.
sulfur recovery
107, S02
0.06 0.03 0.01
21 12 6
6 3 1
260 130 40
Temperature con-
trol . sol ids
circulation sys-
tem opcrability
822
-------
Several Technical unceft aim les exist for each of the recent-ration schemes.
The high-pressure reductive- decomposition processes (I and 2 percent SOj) require very
larp,e coal inputs xml a-ixiliar;' power consumption. The efficiency and operabi 1 ity of
t-nev^y recovery is a technical uncertainty .lion); wi:h r he operability and controllabil-
ity of sulfur recovery with such low SO-> concentrations. The low-pressure reductive
decomposition process consumes power and coal at a lower rate, but it:; oper.ibil ity and
reliability is in question because of the complexity of the solids circulation system.
Capital Investment
Estimates of capita! investment for the regeneration processes have been
developed on the following basis:
' Mid-1977 costs
6") 5 MWe power plant
Interest Jui inj; construction, general items, and engineer in); are .iot
included.
All other direct and indirect cost items are included.
The est inated investments are presented as a function of :he process sulfur load in
Tables VIII t hrous-h X.
The most expens ve process sec'ion for the pressurised reductive decomposition
process is the sulfur recovery or sulfuric acid recovery section. The soiljt-nt circu-
lation section is the -ic>st expensive section for the low-pressure reductive decomposi-
tion process, requirit. complex lockhoppers with water-cooled valves.
Table VIII. Investnent for Pressurized Reductive- Decomposition
Process - 1 Percent S02, $/kW
Process Section r
Ki-k;enerat ion
Sorbent Circulation
Sulfur Recovery
(Sulfuric Acid Recovery)
Total
0.06
15.6
16.9
92.3
(55.3)
124.8 (87 8)
Process Sulfur Load
! ฐ ฐ3 T
10.6
16.1
60.8
(36.5)
87.5 (63.2)
0.01
4.8
15.1
29.5
(17.6)
49.4 (37.5)
Table IX. Investment for Pressurized Reductive Decomposition
Process - 2 Percent S02 . $/kW
i
Process Section |
Rcsenerat ion
Sorbent Circulation
Sulfur Recovery
(Sulfuric Acid Recovery)
Total
Process Sulfur Load
0.06 | 0.03
10.5 -
16.9
68.2
(44.2)
95.6 (71.6)
5.7
16.1
45.0
(29.2)
66.8 (51.0)
0.01
3.2
15.1
21.8
(14.2)
40.1 (32.5)
823
-------
Table X. Investment for Low-Pressure Reductive Drconposit ion
Process - 10 Percent S02,
Process Section
Re Ktrni;rat ion
Sorbent Circulation
Sulfur Recovery
(Sulfuric Acid Recovery)
Total
Process Sulfur Load
0.06
11.4
31.9
27.7
(16.4)
71.0 (61.7)
1 ฐ-
7
31
18
(12
57 .2
03
.8
.1
.3
.2)
(51.1)
| 0.01
3.'
30.1
8.8
(5.9)
42.0 (39. I)
Energy Costs
Energy costs associated with each of the regeneration processes have been
projected using the following basis:
Interest during construction included at 7-1/27,/yr. 3-1/2 yr construction
t ime
Mid-1977
Capital charges of 157,/yr
Operating and maintenance cost of 57, of investment per year
707, plant capacity factor
Sulfuric acid recovery not considered
Mo credit for sulfur produced
Coal at SO.80 per GJ
Methane at Sl.O per GJ
Dolomite at $10.0 per MR (purchase plus disposal)
Sorbent Ca/S ratio of 1.0 for all three process sulfur loads.
For a once-through sorbent operation with dolomite, the required Ca/S ratios as a
function of the process sulfur load are given as follows, based on a once-through
sorbent utilization of 50 percent:
Process Sulfur Load Once-Through Ca/S
0.06 1.7
0.03 1.5
0.01 1.2
Tables XI through XIII give the projected energy costs for the regeneration
processes and compare them to the once-through operation energy cost.
The energy costs of the regeneration processes are co isiderably greater than
the energy costs of once-through sorbent operation on the oasis applied in this study.
For the optimistic assumption that the regenerative processes ma/ be operated with a
Ca/S ratio of C.5, the cost to which dolomite must rise in order to result in a once-
through energy cost identical with the regenerative energy cost is shown in Table XIV.
Economic Comparison with Limestone V.'e'-Scrubbing
The pressurized f luidized-bed combustion po.-er plant with regenerative sorbent
operation must compete economically with commercial power generation systems such as a
conventional coal-fired power plant with limestone vet-scrubbing of the plant stack
gases. The investment costs and energy costs ot regenerative pressurized fluidized-bed
combustion (with elemental sulfur recovery) are cocpared with a conventional power
plant in Tcble XV based on a process sulfur load of 0.03 (4.0 w/o sulfur coal).
The only regenerative PFBC power generation system that compares favorably with
Che conventional power plant with limestone wet-scruobing if the system based on the
low-pressure reductive decomposition.
824
-------
Table XI. Energy Cost for Pressurized Reductive
Deconposition - 1 Percent S02 , mills/kWh
1
i
Capital Charges
Operating and Maintenance
Coal
Methane
Dolomite
Total
Cost Relative to Once-through
Oreration
Process Sulfur Load
0
:i
I
5
0
I
11
9
06
57
19
14
45
64
99
20
i 0.
j
2.
0.
2.
0.
0.
7.
5.
01
50
83
57
23
90
03
68
j 0
1
0
0
0
0
3
2
01
42
47
77
07
35
07
65
Tsble XII. Energy Cost for Pressurised Reductive
Decomposition - 2 Percent S02 , mills/kWh
Capital Charges
Opcrat inf.; and Maintenance
Coal
Methane
Do 1 om i t e
Total
Cost Relative to Once-through
Operation
i
i
i 0.06
2.74
0.91
2.54
0.45
1.64
8.28
5.49
Process Sulfur Load
I 0.03 j
1.91
0.63
1.27
0.23
0.90
4.94
3.59
0.01
1.15
0.39
0.38
0.07
0.35
2.34
1.92
Environmental Comparison
The environmental performance of the regeneration processes for PFBC is com-
pared with once-through PFBC and conventional coal-fired power plants with lines tone
wet-scrubbing in Table XVI. All of the power generation systems are assumed to satisfy
the E'.'A emissions standards (SO2. NOX , part iculat es) for coal-fired plants.
The low-pressure reductive decomposition process is the most environment ally
satisfactory of the regeneration processes. The once-throuph PKBC operation is
environmentally superior to the regeneration processes in all aspects except that of
spent sorbent production. The environmental impact of the regenerative spent sorbent.
versus the once-through spent sorbent due to differences in chemical nature is not
known. The conventional power plant with limestone wet-scrubbing requires coal con-
sumption at a greater rate than do al'. of the PTSC power plants except for the pres-
surized reductive decomposition with 1 v/o S02.
825
-------
Table XIII. Energy Cost for Low-Pressure Reductive
Decomposi-. ion - 10 Percent S(>2 , mills/kUh
I Process Sulfur Load
Capit.il Charges
Operating and Maintenance
Coal
Methane
Dolomite
Total
Cost Relative to Once-through
Ope rat ion
! 0.06
2.04
0.68
0.47
0.45
1.64
5.28
2.49
] 0.03
1.64
0.55
0.23
0.23
0.90
3.55
2.20
i 0.01
1.21
0.40
0.07
0.07
0.35
2.10
1.68
Table XIV. Cost of Dolomite Required to Givซ> Eoual
Once-through and Regenerative Costs. S/Mg
Regeneration Process
Reduce ive
17. S02
Reductive
27. SO,
Reduct. ivซ.-
107. S02
Decomposition with
Decomposition with
Decomposition with
0.06
57
38
23
Process Sulfur Load
i 0.03 I
i i
73
50
34
0.01
118
88
79
Basis: Regenerative Ca/S =0.5
Table XV. Comparison of Regenerative Pressurized Fluid-Bed Combustion
with Conventional Coal-Fired Power Generation
Conventional Plant
Once-through PFBC
Regenerative PFBC
Reductive decomposition with 17. S00
Reductive decomposition with 27. S02
Reductive decomposition with
107. S02
Capital Investment.
$/kW
570
424
526
502
491
Energy Cost.
mills/kWh
23.7
19.8
25.4
23.3
22.0
826
-------
Table XVI, Comparison of Environmental Impacts for PFBC*3
Pressurised
Decompo
17. S02
Plane Heat Rate,c KJ/kWh 12.100
Raw Materials
Coal input ,d Mg/hr 263
Sorbcnt input,6 Mg/day 588-1175
Methane input . 10
kJ/hr 136
Plant Exports fe,
Spent sorbont, Mg/day 435-870
AshE, Mg/day 631
Sulfur. h Mg/day 141
Reduce ive
sit ion Low-Pressure
27. S02 107, S02
10.800 9.600
228 201
588-1175 588-1175
136 130
435-870 435-870
547 482
141 141
Once-through
Operation
9.040
195
1,763
0
1,900
468
0
Conventional
Power Plant"
>.11.000
237
840
0
1.850
569
0
00
ro
Basis: 635 MWe power plant capacity. 4 w/o sulfur coal, emission standards for SOy, NOX and particulates
satisfied.
ฐNow plant with limestone wet-scrubbing.
'Includes auxiliary coal and methane inpuf .
dlncludcs coal for regeneration reductant.
ฐCa/S (dolomite) of 0.5-1.0 for reductive decomposition, 1.0-2.0 for two-step regeneration, 1.5 for once-
pthrough PFBC and 1.2 (limestone) for limestone wet-scrubber.
-Dry, granular for PFBC. limestone sludge for wot-scrubber.
RlO w/o ash in coal.
"Sulfur in auxiliary coal is neglected.
-------
Assessment of Scgcnerative PFHC
An integrated PKBC regeneration system has yet to be denonstrated. Most per-
formance data have he-en generated on small-scale. batch, and serai -continuous apparatus.
and reliable information concerning the critical performance factors for commercial
operation is not available.
The technical performance ol the two PFBC sorSenn regeneration schemes evaluated
is uncertain. The pressurised reductive decomposition will result in such low S02
concentrations (1-2 v/o) that huge amounts of coal for reductant will be required, and
complex energy recovery and sulfur recovery systems will be necessary. The low-
pressure reductive decon;posir ion appears technically favorable except for major
uncertainties in the solids transport system. All the regeneration processes are com-
plex, and th- ir operability and reliability are major concerns.
The overall environmental performance of the low-pressure reductive decomposi-
tion is superior to the other regeneration processes. Both the pressurized and the
low-prossure redactive decomposition processes require the consumption of clean fuels
such as ne'.hano. The once-through sorbent operation is superior to the regenerative
operations in all environmental aspects except the quantity of spent sorbenr produced.
Th: only regenerative 1'FBC power plant t.h'at is ocononically attractive when
compared to a conventional power plant with limestone wet-scrubbing is basi.-d on the
low-pres.iure reductive decomposition. The once-through PFRC power plant has a con-
siderably lower energy cost than iio any of the regenerative power plants, based on a
dolonite cost of SlO/Mg.
ACKNO\;LEDCME:.TS
This work was performed under Contract No. 68-02-2132 for the Industrial
Environmental Research Laboratory of the Environmental Protection Agency. We
acknowledge Mr. D. B. Henschel of EPA for his contribution to this study as project
officer.
REFERENCES
1. Realms, D. L.. et al., Evaluation o^ the Fluidized-Bed Combustion Process
Vol. II, Office of Research and Development, EPA Westinnhouse Research Labora-
tories. Pi'tsburgh. Pa. 15235. December 1973. EPA-650/2-73-048b. NTIS number
PB231-163.
2. Kcairns. D. L., et al., "Fluidized Bed Combustion Process Evaluation - Phase II-
Pressurized Fluidized Bed Coal Combustion Development," Westinghouse Research
Laboratories contract report to EPA - 650/2-75-027c. September, 1975, NTIS
PB 246-116.
3. D. H. Archer, et al.. "Evaluation of the Fluidized-Bed Combustion Process."
Vol. II. submitted to the Office of Air Program, EPA by Westinghouse Research
and Development Center, Pittsburgh. PA, Nov. 1971. Contract 70-9. NTIS
Number PB 212-916.
4. Pope. Evans and Rot/bins R&D Program. Proceedings of a workshop on Regeneration
of Sulfated Limestone/Dolomite for Fluidized Bed Combustion, ERDA. Washington,
DC, March 1975.
5. Monthly Progress Reports - A Development Program on Pressurized, Fluidized-Bed
Coal Combustion, EPA, Argonne National Laboratory. Argonne, Illinois, August and
October 1976.
6. J. Y. Shang and R. A. Chronowski, Comparison of AFBC with PFBC, Proceedings of
the Fourth Intern. Conf. on Fluidized-Bed Combustion. December. 1975.
828
-------
QUESTIONS/RESPONSES/COMMENTS
MR. FABER: Thank you, Dr. Newby. Are there any quick questions.
MR. SMYK: Gene Smyk, Argonne. I'd like to ask you what calcium
to sulfur mole ratio you used for AFBC and the once-through?
DR. NEWBY: I think it was about three and a half.
MR. FABER: Is there another question?
DR. NEWBY: This is fron David Henzel from Dravo Corporation.
He had two questions. "What effect on calcium utilization and
overall cost increase would an appreciable calcium sulfate formation
in the regeneration step cost?" The important parameter here is the
ratio of the number of moles of calcium sulfide generated divided by
the number of moles of S02 generated in the regenerator. As this
ratio increases, you're consuming more reductant from the coal to
generate calcium sulfide so you're not utilizing this for the SO?
generation. Therefore, you're reducing the SC>2 concentration,
increasing the rate of coal consumption in the regenerator and
depending on how high that parameter would get could have a very
significant impact on the cost of the process. I'm not sure what
effect it would have on the utilization of calcium, since it's
normally assumed any calcium sulfide generated would be oxidized back
to calcium oxide in the combustor.
The second question is, "Is there any idea of the savings in
cost between once-through and regenerative waste disposal for AFBC?"
It seems to me that those materials could potentially be so much
different in particle size distribution, in calcium sulfide content,
calcium oxide content, in the ratio of fly-ash to coarse sorbent
material, that handling systems may have to be largely different for
those. We haven't really looked at the details. We normally assume
something like 3 to 5 dollars a ton for an offsite disposal.
MR. HARVEY: All right. Did you have a question? Oh.
DR. SMYKE: Dr. Smyk from Argonne National Laboratory asks,
"How would you propose to control the feed of fly-ash to the kiln if
the carbon content were variable, and it probably would be i.i a real
life situation?"
DR. NEWBY: Well, that's a very good question. I've oeen asking
that myself. For one mole to regender one mole of calcium sulfate,
you need half a mole of carbon. But this reactor is e:x!othennic.
You need .7 moles of carbon to supply the heat, if you want to burn
829
-------
Cdrbon to supply the heat. Now, this is a control problem. I don't
know how much the carbon content in the tly-ash varies fron a fluid-
ized bed combustor. If it varies from 2 percent to 60 percent in the
next minute, then we're in trouble. If it varies from 15 percent to
20 percent on an hourly basis, it's not bad at all. We're talking
about time average basis. So ideally you would like to know the
carbon content from the fly-ash. Say you sample the content twice a
day or three times a day. That would be the ideal case. So again
it, is a control problem.
830
-------
INTRODUCTION
MP. FARFR, CHAIRWI: r>ur third speaker this afternoon is
Mr. Jerome V. Morton. '!r. "orton is with Rurns and Roe Industrial
Services Corporation. He has been manager for the past three years
of the process enq^neerinq and previously held project nanaqer and
supervising mechanical engineer. Receiving his r'F degree, Rf'F rieqree
fron City College of New York and heing licensed in several states,
he has been employed with Purns and Roe for something like five years.
Previous employment include ?5 years of diversified experience in in-
dustry and consulting engineering, fir. Morton.
831
-------
s
An Engineering Study on the Regeneration of
Sulfated Additive from a FluMzed-Bed
Coal-Fired Power Plant
J. H. Bianco. D. A. Huber. J. WLMorton.
and R. M Costeiio
Burns and Roe Industrial ServicesCbrporation
Paramus. New Jersey
ABSTRACT
I'r.Jer DOE sponsorship (Contract No. EX-76-C-ra-2371) , an engineering study of
the regeneration o: sulfated additives from a coal-lired fluidized-bed power plani was
performed.
The- work involved a review of the literature, selection of a viable process
to be .:sc^, preparation of conceptual flow Jiaorar.s, identification of required eciuip-
-t-nt ar.J order-of-rsaqni tude cost estimates for the complete sul fated sorbent pro-
Cfssir.:: ar.j hand; inn system. The system was si?ed tar service a 600 megawatt power
plant.
ซ Several alternative arrangements of the or.c-.ss.cp re7eneration process were
studied and compared to a once-throu<-h sorbent systerv
REGENERATION Or ADDITIVE
When coal is burned in a fluidized bed contaising an additive such as lire-
stone or dolomite, r-O, fron the conbustion of sulfur in the crjal reacts with the bed
material and forms Ca?O^ which is retained in the bei..
The additive ruterial nay bo osthor reqeneraizd to a form suitable for rouse
in the fluid bed system, or disposed of in its partuilly utilized form in a once-
throuqh systen.
rwo reaeneration processes were selected for study. Both consisted of heaiina
the spor.t additi' - in the presence of reducinci oases at relatively high temperatures
to produce caseous sulfur conji-ounds and either calcizn oxide or calciur. carbonate.
One-Step Regeneration Process
The one-step dolomite or limestone regeneration process consists of a sinale
f luidi.-.od bed reactor in which spent additive contaiaina CaSO. from the coal-fired
fluid-bed system is reacted with a rea-cing <;as, suiii as H, aSd/or CO, to produce
CaO and SO,. The cndothcrmic reaction at. 2000 F (125>3ฐt:) and 1 atmosphere
pressure i~s:
c.-.so. *
cf
CO
CaO + SO,
K,0
of
CO,
(U
The rate cf reaction between Ca.^O. and either H, or CO is quite high at 2000 F.
It is desirable to produce hioh concentrations (10 ta> 15% by Kt.) of SO,. The SO,
equilibria concentration is favored by reduced pressure, being inversely proper-"
tional to the total pressure. The reductian of CafE., to CaO is favored by hich
temperatures arid mildly reducina conditions (one no^5 of either H, or CO for every
mole of CoS04).
At lower temperatures,,1650ฐF (699ฐC) and nore highly reducing conditions, the
following reaction is favored :
832
-------
CaSO,
CO
or
H,
CaS + 4
CO,
H20
(2)
The formation of large amounts of CaS is undesirable since it prevents the
reductive decomposition of CaSO^ to CaO. This will require careful control of pro-
cess conditions. If some CaS is formed along the way, it would eventually be
eliminated (to some extent) by the following reaction at 2000ฐF (1093ฐC):
CaS -t-JCaSO.
4 CaO + 4 SO,
(3)
To limit the forr-.ation of CaS, the concentration of reducing gases must be
carefully controlled. Also, advantage can be tukon from the fact that CO, and H,0
and high temperatures suppress the formation of CaS. '
Two-step Process Regeneration
The two-step dolomite or limestone regeneration process involves, first, the
reduction of CaSO, to CaS, and second, the Reaction of the CaS with CO, and H,O to
form CaCO, and !i,S. The first step at 1650 F (699 C) and 1 atmosphere "pressure
CO
CaSO. ป 4 or
H,
CaS + 4
co2
or
II2O
(4)
The second step at 1COO F (538 C) and 1C atmosphere pressure is:
Cap + !!,O
CO,
CaCO, + H,S
(S)
In the first step, the reaction starts out reasonably fast, but then slews
down quickly due to the tendency of the CaS to cover the pores of the remaining
crystals of CaSO,, thereby decreasing the available contact surface.
The regenerated additive from this two-step process must be recalcined tc CaO.
Furthermore, four times as much reducing gas must be used for the two-step ?roc> :s
compared witn the one-step process.
Process Variables
For cither of the two processes discussed, the most important rrocoss variables
from the standpoint of regeneration performance are temperature, partial press-re
of the reducing gases, and space velocity or contact time of the Ca?0. particle;
and gases . An appropriate system for carrying out the additive regeneration pro-
cess is a fluidized-bed reactor which will provide the necessary temperature unifor-
mity as well as efficient contact between the oases and solids involved.
When operating in the range of 2000ฐF (1093ฐC) deactivation of the CaSO^
particles by sintering or deadburning occurs, especially when the solicis r.ust under-
go repeated cycles of sulfur absorption and regeneration. In addition, a decree of
solids attrition can be expected in the regeneration step. Additive rccirculation
rate through the reoenerator and fresh additive make-up rates to the ccr.bustor arc
determined by the amounts of deactivation and attrition that occur in the overall
process. Therefore, the overall economics of regeneration will be greatly depen-
dent upon the additives resistance to these parameters.
833
-------
";< ':on|.o:;iMor: of .ป.: ..:::. ':O.T.:.O:.ซT.*. .-. t :.. :::..:. .,): tiv.- will
of t!.<- p.i -t I'.-l--:; in t h-- f i .ii ! j /.<*. !.!.
With hi'ih'.-r r--'i--r.- r.it ion ':-;. --r.iM.ir--:: , *h--r'.- :>.
.ow.irds !.i'|h'-r :;o;> -:o:i-:-Tii r-itior::; i :i t h" r<-'!":."r-i '.or
v--r:;ior. to '.'.iO.
mol" f r .i--t . lor. of TXr. i r, t.h*- r"-;ซTiซ-r.iป.or off-').i:; i :.fr<-.is<-.-. , ! !' 1 co::?.::
f>,r rซ"jซ-n<'r.iป-.ion rซ':;surซ-B l,'-c.iijr.'.- th" nol" f r-i'M ion nl .'"/>^ .-it ) MI ! ;.-t th" .'jM.-ict tim" - - x:. -1
throii-ihout 1 1'." l.ซ.-'j, with si :ni I I'.-.mt |-ro;iort lor.-; 'if C-i.r-'> I." in'! fornn.-fj in::t".i'i of
C'jO. Thi:; [iroi>l--rp e-.in ljซ- sh-irf'ly r---lu-_-"ii l"'l'. Tin:: ''r--.it-':; .nl j.ir:ซ-nt rซ-'l-j':i r.-| .i:i'i oxi:. iriitin.it" yi- I'i-'-
hyilro.|.-n Kulfiil-- lor r'-c-.v-ry . T.il.l-- I li::t:: |,ro':- ss t-or.tl i t ions .ir.-.l .-ml j-rociu-.-t .
The- 1'illowiri'l t.. ilml.it. ion li:;t:: t h" .,'ivjnt .i'lซ-r; .ind -ii -I'lv.int-i-l'-:: to th" t wซj
r--li-niT.it ion |iroc<-::::"::' :
a. l-'or th" t.i-;<- \-T'J'-> .;.;
- l:x|"T IRI"!lt*jl 'l.lt.'l IS lV.lllll.1"
- C.i') is fornn.-d dir--';tly
Di s.idvant-'i'l'.'S .irr:
- H--*f.TJtiirt.-s, 2000 K (10'n"C),
uro required to ..ivoid CjS for
mition .'ind d---.ictiv.it > on of the
.idditivc. Also.clor.c t"ni'.-.-r,-iturc
control is needed to avoid aq<)lo:ner-
ation of the cool ash i :i the bed.
- At equilibrium, SO-, concentration
decreases with pressure.
b. For the two-stop regeneration process:
Advnntaies are - Thermodynamics favored by low
temperature 1600ฐF (871ฐC)
- No thermodynamic disadvantage
due to pressure
- Low temperature avoids solid
sinter ing
- Pressure favors HjS production
Disadvantages arc
- Two stages required to form CaCOj
- Second stop requires high pressure
CO2 and HjO
834
-------
TAB!-!: I. kKCKNKRATION PKGCKSS CONDITIONS
Ono-Stop bo-j'-Tn:- rat i on
Condition:;: 2uOOฐI', ono Ate1., press..
Ono [Milt: reducing qascs
Knd Products: CaO for ritcyclo to coribustoi.
SC>2 for sulfur recovery
Tv.o-5t'-p fro.-iofu.-r.'
i i r r;t Sf.-p - Conditions: lbOOฐF, or.o Atn. press..
Four n>oles reducing ;jascs
I f.-qu i r c-d
Knd Products: CaS lor use in second step
S'.-cond Stf.-p - Conditions: 1100 F, ton At*., press.
CO2 and f^O ijascs required
End Products: CaCO^ for recycle to
comlnistor
II2S for sulfur recovery
835
-------
- Little experimental data available
or publicized
- Competing reactions reduce sulfate
to yield SO,
- Produces carbonate- rather than oxide
that must be recalcined for recycle
- Reaction rate of CaS conversion slows
down drastically.
While it is recognized that no firm conclusion can be drawn from an evaluation
of the above, the one-step process was selected for economic evaluation in this study.
SELECTION OF REGENERATION SYSTEM ALTERNATIVES
Regeneration Systems With Claus Sulfur Recovery Plant
After having selected the process to bo used for regeneratinq the spent addi-
tive, the support processes required to achieve a conplete integrated system were
then selected. The philosophy adopted to guide the system design was to utilize com-
mercially proven processes where available in order to limit the time and cost re-
quired to commercial i/.e the plant. This criteria led to the initial selection of a
Claus plant for the separation and recovery of elemental sulfur.
Prelimi.--ry calculations indicated that it would not be economical to pur-
chase H,S for the Claus plant. In order to produce H-S in-piant, the amount of
reducing gas required would be four times that required for additive regeneration.
This factor led to the decision to use a separate reducing gas plant to provide the
raw qas needs for both a II2 producing plant and for the sorbent regeneratiors. It
was further decided that. In the interest of completeness, an estimate would still
be prepared for the case of purchased II2S. However, to facilitate design and cost
estimating efforts, the production of reducing gas for the regeneration process was
still accomplished in a separate process rather than directly in the regenerator
vessel itself. While probably not the most economical approach .or this case, the
incremental costs involved would not significantly affect the conclusion regarding
overall economics between Case I (Purchased HjS) and Case II (In-plant II,S manufacture),
Figures 1 and 2 show block diagrams of processes selected for Cases I and II
respectively. Reducing gases are separately generated in a Koppers-Totzok Coal
Casifier Package Unit. A standard Claus sulfur recovery unit and tail-gas treat-
ment plant arc also shown. It should be noted that the Claus tai:-gas clean-up
plant would only be required if the recycle of tail gas to the fluid bed ccra-
bustors proved technically or economically impractical. While this is considered un-
likely, the clean-up system is included here as a conservative measure. Again, the
cost of this plant does not significantly affect the final conclusions. Other pro-
cesses used in Case II include a conventional water-gas shift reaction for production
of II2 and the catalytic reaction of \\2 and sulfur vapors to produce ll^S gas.
The regenerator itself would be operated as a fluidizcd bed reactor designed
for the following conditions:
Temperature
Pressure - 1-2 Atre.
Regeneration Efficiency - 65%
Residence Time - 5 to 7 Minutes
Fluidiiing Velocity - 4 to 7 Feet per Second
Reactivity - 72%
On the b?sis of the block diagrams shown in Figures 1 and 2, a conceptual flow
diagram and approximate heat and material balances were prepared for each case.
Figure 3 shows the flow diagram for both cases, with Case II the more complex of the
two cases. Approximate sizes for the equipment and piping indicated on the flow
diagrams were then established for both cases.
836
-------
CtMnhotgoM
Wttoon ptoal
lnซrtป to otmeซpปiซrป
1
^M. llall |
COMB
Co* OR
1 ,
\
Sport *orปMl
Rt*i
r
^""....K* TOT7FK
GAS1FIER
Cool PLANT
UST-f< -l REGEN-
S ERATOH
r
i * i
J-Rtgviorot*
idafloo***
J 1
' 1
STO. IT AIL- CAS ,
CLAUS 'CLEANUP!
UNIT t UNIT I
Tl L J
To Mlot or
ซMXbOBt MeHontulfur *ipoปol
PURCHASE!
M2S
Figure 1. Sorbent Regeneration with Claui Plant and Purchand H2S
Casel
iMrti to otmotalwr*
Figure 2. Sorbent Regeneration with Claut Plant and Manufactured H.S
One II
837
-------
00
IO
00
?^^irv^^
ra ii
i -'" Li r~7l ^ V - r-^-il |r\
ฃp t T'i'irp"]^^r\y^ i
I ' ** I ,-.f.i r-.ta-
rfj r>
N-1 h*-
^/vnww**/
"" Mte/f
'**ซ*.'
Cf
"i-'-'-'i
\*^*'
"Vr
l4i
Figur* 3. Sorbent Rejsneration with Clauj Plant
Cam I & II
-------
Regeneration System With RF.SOX ' Plant
When it became apparent that the costs associated with the Claus Sulfur re-
covery system represented a major portion of the total ann'jal operation costs for
Cases I and II, a decision war- made to investigate Foster h^eolcr's RE3OX process
for this application. To our knowledge, there is no RKSOX process in corrjr.ercial
operation as yet, and there is very little technical information available for study
and evaluation. However, it appeared that overall system costs with a RKSOX unit
could be lower than for an equivalent Claus hased system. Therefore, it was de-
cided that for comparison purposes, a third case incorporating a RESG.: system should
be studied on the same basis as Cases I and II. The block diagram on Figure 4 and
the conceptual flow diagram on Figure 5 wore prepared based on our interpretation of
the sparse information available on RESOX systems.
Case III also differs from Cases I and II in that the RESOX tail-gas is re-
cycled back to the AFBC boiler and no tail-gas clean-up system is unc
-------
Owtotoom
lo
RttOl toii QCT
Coal
COMBUST-
r
t
SORBENT
REGEN-
ERATOR
1 RE SOX
REACTORS
SOL
COND
El
Antnracin coal
Sewt HrbMt ood aid
naaanซro1ป< tarittat
IMOHM wlfar ta Mlnar
Figure 4. Sorbent Regeneration with Retox Sulfur Recovery Plant
Can III
I
*.
1U
r
u- r
-.v=i-ii-
?
_^
z. \
J
Figura S. Sorbent Regeneration with Reiox Sulfur Recovery
Gate III
840
-------
TABLE II. BASIS FOR ECONOMIC EVALUATION
Materials
Coal
Labor
12,450 BTU/lb. HHV
4.5% S
$19.50 per ton del'd (Bituminous Coal)
S25.30 per ton dcl'd (Anthracite Coal)
Limestone fc Dolomite SV/ton delivered
H-S S240/ton
0, S40/ton
Amines S0.77/lb.
Electri-
city S0.0225/KKH
Water S.15/1000 gal avg. for all types
Waste Dis-
posal S3/tcn (spent stone & sulfur)
Sulfur S50/ton fob plant (sales)
Operating labor at $20.000/man yr. incl. fringes
Operating superv. at $2S,000/man yr. incl. fringes
Chemist, engineer, etc.
Maintenance
Including labor, supervision, supplies, materials and
parts at 5% of capital cost.
Capital Charges
At 181 of capital cost
Admin, t Overhead
At 40% of labor and maintenance
Cases Studied
Base Case - Once-through Sorbent System
Case I - One-Step Sorbent Pegeneration System Buying H,S
'Over the Fence"
Case II - One-Step Sorbent Regeneration System Making H,S
In-Plant
Case III - One-Step Sorbent Regeneration with Resox Sulfur
Recovery System
841
-------
TA.-,L=: in. CO;;T rac-Aaison FOH 600 we
COMBINED i-ov.'ER J-LAOT
Capital Cost
orieratin/> Cost: -f/yr.
Direct Costs:
(a) Haw Materials -
Additive
Bituminous Coal
Anthracite Coal
Oxyr.en
Liquid I!?S
(Use. Chemicals
Subtota I :
(b) Utilities -
Electricity
Water
Subtotal:
(c) Stone >, Ash Disposal
(d) Maintenance, Etc.
(e) Operating Later
TOTAL DIRECT COSTS:
Indirect Costs
(a) Capitol Cfian;es
(b) Admin. '/ Overhead
TOTAL DiDIKKCT COCTS:
TOTAL ALL COSTS:
Without Sulfur Disposal
Credit for Sulfur Sales
Net Cost With Sulfur Sales
;!et Cost Without Sulfur Sales
SO2 Renoval Cost Per KWH:
With Sulfur Sales
Without Sulfur Sales
Once
Through
SO, 500,000
$1ป, 1*61,000
$1.
$
$
$2
$
$7
$1
$8
$
$8
ฃ8
,1*61,
31*,
'3,
73,
,633,
275,
i?a,
,625,
990,
181,
,171,
,796,
ปV?6,
,796,
000
500
000
000
000
000
ooo
000
000
ooo
000
000
0
000
ooo
$ 2.3 ail.
Regeneration
Case
$26,0}0
$ 953
1,258
28,912
53
$31,177
* 1,671
221
$ 1,392
' $ 901*
$ 1,302
t Ull
S5.686
$ "*,665
685
$ 5,370
$1*1,056
$ 7,390
$33,658
$1*1,500
$ 8.9
$ 11.0
I
,000
,000
,000
,000
,600
,000
,000
,000
,000
,000
,000
,uoo
,000
,000
,000
,000
,000
,000
,000
,000
mil.
mil.
$51
$
2
3
$ 7
$ 1
$ 2
$
$ 2
f
$13
$ 9
$10
$21*
$ 2
$21
$21*
$
$
Ci^e
."/stems
ฃ r
,310,000
953,000
,606,000
,522,000
160,000
,21*1
,911
376
,288
901*
,6ir,
57')
,628
J278
,69"*
,322
,1*57
,865
,1*69
5.8
6.5
,000
,000
,000
,000
,000
,000
,000
,000
looo
,000
,000
,000
,000
,000
mil.
nil.
Case
$1^,200
$ 1,%6
1,307
805
$ 3
t.
$
$ 1
t
$ 7
t 3
$ 3
$10
$ 2
$ 8
$11
$
$
VT'
III
,000
,000
,000
.000
6J5.000
,096,000
510,000
,11'*
,276
,810
,92-*
,158
,766
,051*
2.3
2.9
,000
,000
,000
,000
,oco
,000
,000
nil.
nil.
842
-------
17 T
s
x
>s
CO
_1
x
to
o
u
5
o
2
1U
K
O
CO
WITHOUT SULFUR SALES
WITH SULFUR SALES
0 5 10 15 20 25 30 35 40 45 50 55 60 65 70
CAPITAL COST $ ป I06
Figure 6. Capital Cott Sensitivity
843
-------
In comparing direct operating cost, Figure 7 indicates that oven if the- actual
value wore 40V. less than the estimated value on Table III, Cases I and II would bo
uneconomical compared with the- once-through cose. However, within tho accuracy of
this estimate. Case III cost when credited with sulfur sales is comparable to costs
for a once-throu-jh system.
Fiqure 8 shows tho affect of total opera-ting cost on SO, removal costs for
the various cases.
Figure 9 shows tho additive- cost that would be required in order for the var-
ious sorbent regeneration cases to be equivalent economically to the once-through
operation case. When taking credit for sulfur sales, those breakeven sorbent costs
are as follows: Case I - S^B/ton; Case II - 534/ton; and Case III - S7/ton.
ADDITIONAL REGENERATION I'I MIT DATA
Table IV details tho chomi-.il complex manning requirements for the various
cases studied. With proper training all personnel listed are considered interchange-
able with power complex personnel.
Table V show:; a summary of overall material flows for the various cases ;n tons
per day.
CONCLUSIONS
The results of this study indicate tho following:
1 - From a technical viewpoint, sorbent regeneration appears feasible if f-
<|uirod to reduce the environmental impact of fluid bed solid v.iste dis-
posal and utilization. .More oxper imenta 1 data is required if a commercial
;.!"!* is to be designed with significant confidence levels.
2 - Sorbent regeneration utilizing sulfur recovery processes with commerci-.il
operating experience, such as the Claus system, cannot bo economically
justifii-d unless sorbent costs approach S"50 per ton.
3 - Additional development efforts are required in order to achieve an
economical ;:dditive reqenoration nystvn. These efforts must be focused
on the development of an economical sulfur recovery system, such as
KKSOX, as well as on tho regenerator itself. Development of o.iซ- with-
out the other will be of no use economically.
4 - If the currently projected costs of a RKSOX system provo realistic, sorlx-nt
regeneration jtilining this system for sulfur recovery may be more econom-
ical than a once-through sorbent system based r.n a sorbent cost of over
$7 per ton. However, to our knowledge, there is no RKSOX system in com-
mercial operation today.
ACKNOWLEDGEMENT
This study was carried out on Contract EX-76-C-01-2371 for the U.S.Department
of Kneny. The authors grcatfully acknowledge the assistance of Mr. George Weth of
D.O.E. in tho publication of this paper. We wish to acknowledge tho assistance and
cooperation of Argonne National Laboratory in developing the input information for
Case III.
REFERENCES
1. Evaluation of the Fluidized-Bed Ccubustion Process, Volume 1, Pressurized-Bed
Combustion Process Development and Evaluation, Kestinghouse Research Labora-
tories, Prepared for EPA, December, 1973. P. 165-167. 149.
2. Iloke, R.C., et.al., "Combustion and Desulfurization of Coal in a FluicJized
Bed of Limestone", Combustion, January, 1975, P. 6-11.
3. Montagna, J.C., ct. al., "Bench Scalp Regeneration of Sulfated Dolomite and
Limestone by Reductive Decomposition", Presented at the Fourth International
Conference on Fluidized-Bcd Combustion, December, 1975, P.393-423.
4. K.M. Guthric, W.R. Grice s Co., "Data and Techniques for Preliminary Capital
Cost Estimating", Chemical Engineering Magazine, March 24, 1969. P. 114-142.
844
-------
/ /
^ ซ. '\ f. I* ซ. * ซ -. ft
Figure 9. Additive Coit Sensitivity
845
-------
TABLE IV. REGENERATION PLANT OVERALL MANNirJG REQUIREMENTS
1
Basis: 7 Days/Me.
Once Regeneration Systems
Through Case I Case II
Overall Supv.
Operation
Operation Supv.
Shirt Supv.
Operators
Helpers
Chemists
Clerks
Total
Maintenance
Maintenance Supv.
Electricians
Holoors
Millwrights
Helpers
Pipe Fitters
Helpers
Machinists
Instrument Tech.
Engineers
Laborers
Total
Overall Total:
_ _
S l
5
3 6
3 6
2
- 1
6S 21
S I
1
1
1
1 1
1 1
1 1
1
1
1
2
_ii iL.
II 33
Note: It is our considered opinion that
personnel listed
plex personnel.
2
6
t
e
2
2
28
1
2
2
2
2
2
2
2
2
2
3.
22
50
after proper
above are interchangeable with
Case III Remarks
From Power
Complex
1
5
5
5
2
1
19
1
1
1
1
1
1
1
1
1
1
2
12
31 ,
training all
power com-
846
-------
TABLE V. SUMMARY OT OVERALL MATERIAL FLOWS
FOR 600 MWe COMBINED POWER PLANT
Basis: Tons/Day
Once
Through
Regeneration Systems
Case I Case II Case III
Entering;
Coal: For Power 5,000
For Regeneration
For Sulfur Recovery -
Oxygen
Liquid H2S
Misc. Chemicals:
Nitrogen
Amines
Additwc 2425
Leavingซ
Spent Additive t Ash 3340
Recovered Sulfur o
5.000
245
458
10
0.1
518
1147
563(1)
5.000
509(2)
335
20
0.1
518
1147
167(2)
5.390
255
121
1025(5)
1390
164(4)
Notes: (1) High sulfur recovery due to purchase of H-.S
(2) Includes sulfur in coal to regenerator an3 qasifior
(3) Total coal used for regenerator and qasificr for
hydrogen generation
(4) Stoichiomctric recovery at 901
(5) As optimized by Argonnc National Laboratory. How-
ever, according to Argonnc data the total costs arc
quite insensitive to additive feed rate over the
range of these studies.
847
-------
QUESTIONS/RESPONSES/COMMENTS
MR. FAPFP, CHAIP^AN: Yes, do we have a question or two' <-'ould
you qo to the r.ike, so we don't lose our speaker off the end of tho
honch here.
MP. COrilPAN: Mank Cochran, Oak Pidqe. Would it he possible to
provide the hydrogen sulfide requirenents of the Klaus (?) plant by
qasifyinq tho hiqh sulfur fraction fron the coal heneficiation, and
providinq the hydrogen sulfide in that way'
rซT. MORTON: '-'ell, that's a very interestinq Question. It
sounds like a good idea hut obviously without thinking about it and
studying it a little hit I really couldn't answer it. It would have
to he looked into.
fp. STFINRFRf-: Steinberg fron Rrookhaven "ational Laboratory.
When comparing your reqenerativo FPC with once-through, have you
tried to nake any estimates on tho cost, including the cost of
disposal, of the once-through system, as you mentioned possibly a
landfill or sone other means and what are the results?
MP. MORTON: Yes. Actually, we -1id include the cost of disposal
in all of these estimates. l*e used the value of 3 dollars per ton
for the landfill costs. In the case of the baseline system v/hich was
the once-through system, the total amount of limestone ash was being
disposed of at that cost and in the other systems, some percentage, I
think it came out to approximately ?o percent, which was our makeup
rate, was charged at thdt 3 dollars a ton. So, we did include dis-
posal costs.
HP. STFINRFPfi: Put there may be a break-even cost going up in
land disposal in certain areas where the cost of land disposal comes
pretty high.
MR. MORTON: Yes. ^e use three dollars a ton and, obviously, if
that ninher got much higher, it could change all the economics.
f9>. POPTFR: .Hn Porter, Fnergy Resources. In you calculation
of the break-even prices of limestone, what did you assume for the
number of times the limestone could be regenerated before it spent
itself out?
MR. MORTON: Okay, here we used the data that Argonne has de-
veloped. And using some of their data it appeared that by replacing
the limestone that was lost by attrition and by sizing our regenerator
848
-------
large enouqh to get an adeouate anount cf reci reflation, we could got
hy with approximately twenty....! think the exact nunher cane out at
?? percent nakeup rate. Approximately that.
This, hy the way, is sonewhat of a tradeoff between the size of
the regenerator and the anount of naterial you want to nake up. YOU
have sone leeway there.
MR. McnAM|_Fv: 'ty nane's Mc^auley, Pope, Tvans and Pobhins. I'n
particularly interested in whether these sulfur plants you're examining
have turn-down. That is, can they follow the power plant. Or not.
Or what does it cost to put the turn-down on?
MR. WPTW': Okay. We actually didn't get into that stage of
detail to worry ahout that sort o^ thing, odiously, it's a question
that would have to he exanined if this were carried further. Rut we
didn't do it.
MR. HAPVEY: All right, are there any further questions here? I
apologize for naking this rather disjointed in ny anxiety to hear fron
peon's e like Hick Miller. I overlooked tha fact that there were two
questions left fron the first session and to when were they addressed?
SPEAKER: Mr. Morton.
MP. MAP"Fv: Allright. You have. ..where is he? You going to
answer? V'ould you use that nicrophone, please?
MP. WPTOP: Mr. nijk Newhy asks the followir.g question of Terry
Morton of Burns and Roe. Hhat SO? fraction was assuned to he pro-
duced fron the regenerator? In all three cases.. .or nay I say that
1n case one and two the volume fraction was 7.ฐ. In case three it was
&.<>. The next question was "What pressure was the regenerator oper-
ated at?" Essentially atnospheric. Third question vas, "What was the
source of your Pesox cost data7" Fron a nodular cost estinating of
each conponent in the envisioned systen and sone general infomation
we ohtained fron conversations with Argonne National Lahoratories.
That's how we arrived at the estinated capital cost.
349
-------
INTRODUCTION
MR. FABER, CHAIRMAN: Our next, lecture is on the subject of
economic feasibility of regenerating sulfated limestones, presented
by Eugene B. Snyk He's with the Chemical Engineering Division of
Argonne National Laboratory, receiving his BS in chemical engineering
from Notre Dane, his ML from the University of Florida. He was
employed by the Commonwealth Edison Company as a process engineer in
charge of operating a wet SOp scrubbing facility from '71 to '75.
That in itself would get him all the credentials he needs to get into
heaven, I'm sure. From '75 through '76, he was employed by Air
Correction Division of UOP Incorporated as a project engineer. He
joined Argonne National Laboratory in April of 1976 and has worked on
pressurized fluidized bed combustion and, more recently, in bench
scale fluidized bed regeneration. Eugene.
350
-------
Economic Feasibility of Regenerating Sulfated
Limestones
E. B. Smyk. J. C. Montagna. G. J. Vogel. and A. A. Jonke
Argonne National Laboratory
ABSTRACT
Rop.encr.it ior. of t hi- C.iSO. in the spent sorbc-nt of a f lui ili xod-hed cor.huslor ami
re-cycle of the- rcsuit in:' CaO offers a roans of reducing the waste disnosal burden of
this t echr.oloi".-. A f luidi::c-d-bed reductive decomposition limestone re-.-enerat ion pro-
cess has been developed in PDl'-scalo equipment.. Predict ions fron a model based on
these experiments combined vith cost data developed by Vest inp.house allows not costs
to be determined for various opor.it in" and economic parameters. Economic feasibility
will denend on the values selected for these factors A framework is provided for
analy;-. in;- specific cases.
i::~RODrcTio::
Fluidi:'ed-hซ-d combustion is beini* developed as an environmentally acceptable
ir.ethod of utili::ir.p our nation's vast resources of coal to rent-rate electricity anil/or
steam. In addition to fly ash waste, vhich is also produced hy conventional boilers.
f luidi/ed-bc-d cor.bustors produce significant amounts of spent sorbent vhich mist be
disposed of, Reri-nerat ion of C.isO. in the spi nt sorbent followed by recycle of the
sorbent to the comhtistor (boiler) offers a means of substantially reducing this addi-
t ional waste disposal burden.
A fluidi7.ed-bi.-d reductive-decomposition 1 iniestone-rei'er->rat ion process (in
which both the heat of 'iction and '.he reducinr. >hr>iisc: nerforned a cost st udv on a sorbent rcp.cnerat ton system for a
6 "J 'j .'-!W AFBC coT^bustor. The system considered was an atmospher ic-oresst'.re , fluidixcd-
bed reduct i ve-Oecompnsit iop. process, followed by a RKSOX Sulfur Recovery System with
tail ras incineration. Capital costs were presented for x-arious process sulfur load-
ings (i'SL. pounds of sulfur to the rcc.enerator versus pounds of con I to the conbustor)
ami for different rei*onor.'itor off-pas SO. concentrations. In addition, data on
capital cost versus solids ciruclafion rate between the combustor and the rcpenerator
was presented for the sorbent recirculation system.
However, in their cost analysis. MostInr.houso did not specifically include
different regenerator off-pas SO concentrations and fresh sorbent feed mole ratios
but instead used fixed values. (The fresh sorbent feed mole ratio, hereinafter refer-
red to as FR. is the cole fraction of CaO as fresh feed to the corabustor. compared
with the total CaO feed to the combustor including regenerated sncnt sorbent.) ANL
cyclic experiments with Creer limestone and Tymochtee dolomite- "'u show that the sor-
bent. feed rate, recirculat ion rate, off-p.as SO;, concentration, and regenerator coal
feed rate can all be predicted for specific operating conditions."" The ANL results
also demonstrate that an economic analysis is necessary to determine the most favor-
able value of Fh.
RESULTS AND DISCUSSIONS
The A"L model and the PDU-data were used to relate the values of various para-
meters f.o FR. For example, a low FR (C.R., 0.05-0.10) would translate to low sorbent
consumption, hir.h, coal consumption (in the rer.cnerator). a larp.e rcp.encrator. lower
SO.- concentration in t*H> regenerator off-p.as, and hip,h solids recirculat ion. A hip.h
FR (C.R., 0.30-0.&0) woulu translate to medium sorbcnt consumption, lower ccal con-
suription in the regenerator, a smaller regenerator, hip.hcr SO. concentration in the
-,\l\ results froT these AM. experiments were not available to Westinp.house when they
made their economic study
851
-------
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Ki|-.J"<- V :-.liou:: tin- '.rrvct of plant. c-apa.:!ty fact. )i- nn rift systv.-a .-c:>*>.. r I.Vir-.- 8 r.Sown
ttif >.!'.":I- c.f cti?i I i.rSco oil iu.-v r,y:;t<-n co:;t.. Kl^ur'.- *> rr.owj tr." vlTvct. cf r.ulrui-
credit on net :;yct'-m O'jr.t.
At a KR sil' i . 0 ' corre:;::'jriUI ru* to a o:ici'-thr'C.u>;h sys'-oiTi wit,;. t It art.- fairly fl'it !n the rr region L^tween rj. '<.' ari'l O.;i.
Th'i:;, t->ie .conoralcs oT a cyuti.-Ri Juct hf-aklns t"/en (i.e., a net coct tT ^ero) iio i:ct
ilfpcnsi str'jnj^ly on trio vaiuซt ol" KH chocen for o:?-rrat Ion wri^n r'H 1^ between '".'.ซ".' ar-d
O.-'i. liOBCV'-T, a:; the net ir,ein.-rat Ion system c-j;-,t cie-j;-ซ2Ees furt.'.'-r, t r.e cftir.u.-n FR
'i'.'ercasns until at atout -l.o nlll/kWh, thf O|.tinum r'R CCC-JTS at 0.10 to 0.1'.j. Mlnl-
muD breakeven sorbent price (Kii:>P) Including disposal K.IG chOoen as the variable In
whicr. to express all recultr. and was aeflnecl as the nlniaua total r.orbent price Tor
which a ฃlven nystem Ic economically feasible. Since a regeneration system conserves
sorb^nt , a high UP :;orbent i;rlce tenas to nalce the systen core viable, while a decrease
In sorci.-nt price nakes It less viable. Therefore, a total sorbent price greater than
K3SP makes the system economically viable.
852
-------
25
d 15
in
O
u
10
05
50%
00
0 10
020 03O
FEED RAT!0
040
050
Figure 1. Plant Coil (Including Operating. Maintenance. Contingency.
Contractor Fee. and Capital Chargt) ซ. Feed Ratio at
Variout Capacity Factor*
853
-------
0 40 '
0 30 I
Figure 2. Utility COM w Feed Ratio
10
0 8
06
Z
i~*
O
04
02
010
O 20 0.30
FEED RATIO
040
050
Figure 3. Net Coal Coit n. Feed Ratio at Various
Cod Pricet (S/Ton)
854
-------
00
en
en
80
7.0
60
3.0
*
M
^
in
en
O
u
4.0
30
20
10
010
0 20 0 30
FEED RATIO
040
050
0 010 020
FI.CO HAHO
Figure 5. Net Sulfur Credit vs. Feed Ratio at Various Sulfur Prices (S/Ton)
Figure 4. Net Sorbenl Savings vt. Faed Ratio at Variout Sorbent Prices
(S/Ton. including disposal)
-------
0.10
O.20 0.30 O.40
FEED RATIO
Figur* & Nซt Cod of Regeneration Syrttm v*. Feed Ratio for Varioui
Sorfaent Prica ond Capital Coil Escalations. Capacity Factor
70K Coal Prica S25/Ton. No Sulfur Credit
856
-------
-0.5
0.4
FEED RATIO
Fioir* 7. Net Colt of Regeneration System vs. Feed Ratio for Various
Sorbent Prices and Capacity Factor*. Coal Price $2S/ton. No
Sulfur Credit
25
20
10
o
o
05
05
10
SORBENT
5/TON
SORBENT
AT MO/TON
ISORBENT
/ AT MS/TON
J
I
0 01 02 03 0.4
FEED RATIO
Figure 8. Net Cost of Regeneration System vs. Feed Ratio for Various
Sorbent and Coal Prices. Capacity Factor 70%. No Sulfur
Credit
-------
O.I
0.3
FEED RATIO
0.4
Figure 9. Net Cent o! Regeneration Syitem w Feed Ratio for Various
Sorbent Prices and Sulfur Credits. Capacity Factor 70%.
Coal Price S2S/Ton
858
-------
859
-------
25
20
a.
in
m
10
I
I
I
I
COST CAPACITY SULFUR
ESCALATION FACTOR CREDIT
% % $/TON
50 50 0
50
50
70
50
70
50
70
50
70
25
50_
0
0
25
25
50.
50
0
70 25
70 50_
10 20 30 4O
COAL PRICE. $/TON
50
Figure 10. MBSP vi. Coal Prica at Various Capacity Factors. Cost
Escalations, and Sulfur Credits
860
-------
The second computer program can predict the effects of the following variables
cr. net plant cost:
1. Capital cost
2. Capital cost escalation
3. Plant size dXVJ)
^. Plant efficiency (heat rate)
.5. Coal heating value (iiHV)
6. Coal sulfur content
7. Sorbent calcium content
3. CaO/S feed ratio
9. Regenerator off-gas S02 concentration
10. Contingency
11. Contractor's fee
12. Capital charge
13. Operating and maintenance charge
IH. Utility cost
15. Coal price
16. Sulfur credit
17. Sorbent Price
13. Feed ratio (FFO
REFERENCES
1. Fluldized Bed Combustion Development, Volune II,Calcium Based Sorbent
Regeneration, VIestingtiouso Research and Development Center, Contract ilo. 68-02-
213Jr, USKi'A, February lyYY.
2. G. J. Vogel ct al. , "Supportive Studies in Flu;-.11 zed-bed Combustion," Argonne
liatlonal" LAboratory, July-Septcriber 1976, A.'JL/hlS-CEN-lOl? and FE-1780-5.
3. S. J. Vo?;el ':t al., "Supportive Studies In Fluidi7.ed-Bed Combustion," Argonne
National Laboratory, January-Xiarch 1977, A!IL/K::-CK!;-1019 and FK-1780-7.
U. r,. J. Vogel et al., "Supportive Studies in Fluidized-Bed Combustion," Argonne
I;atloi;al Laboratory, October-Uecenber 1976, ANL/ES-CiCN-lOlS and FE-1730-6.
ACKNOVfLEDGMENTS
This work was performed under the auspices of the U.S. Energy Research atid
Development Administration and Environmental Protection Agency. We thank D. Webster
and Les Burris for their support and guidance. We also thank J. Simmons the technical
editor.
861
-------
QUESTIONS/RESPONSES/COMMENTS
MR. FABLR: Would you pass your written questions to the center
aisle, please. Bill will pick them up.
SPEAKER: I guess I have some misunderstanding. I think you
said the fresh feed to recycle ratio is determined by economics and
if one has to have a bed with fixed reactivity, given a coal through-
put and there sulfur throughput through that bed, then I think the
fresh feed to recycle ratio in the bed is fixed by the loss of
reactivity of the stone as it goes through the bed. There's a direct
correlation between the two. Maybe I misunderstood what you said.
MR. SMYK: Let me r,o through our experimental technique first.
Correct tne if I'm wrong on this, John (Vogel). We ran cycle
combustion-regeneration experiments (i.e., with no fresh make-up
sorbent) with both Tymochtee dolomite and Greer limestone results
only. Since the combustion experiments were run with Sewickley coal
(containing 4.3 percent S), 83 percent S0ฃ retention was necessary
to comply with EPA emission regulations. In combustion cycle #1, a
Cal/S mole ratio necessary to obtain 83 percent SC>2 retention was
utilized. In combustion cycle #2 (i.e., the stone had been regener-
ated once), the same CaO/S mole ratio was used as before and the
S02 retention was measured. The S0ฃ retention in subsequent
cycles was measured and used to correlate the loss of reactivity of
the stone as a function of combust-regeneration cycle.
Our mathematical model allows us to predict the age distribution
of the bed sorbent material as a function of the ratio of fresh
sorbent feed to the bed versus the recycled sorbent feed from the
regenerator. For example, at a low fresh feed to regenerated feed
ratio one would have an "old" and fairly low reactivity bed. Conver-
sely, at a high fresh feed to regenerated feed ratio one would have a
"young" and highly reactive bed.
Now, if a bed is highly reactive it can attain 83 percent
retention at a fairly low total CaO/S mole ratio; but if it has a
fairly low reactivity, a higher total CaO/S mole ratio must be util-
ized to attain 83 percent S02 retention in either case; the trade-
off is this: At a low fresh feed to total feed (including regenerated
feed) ratio one uses less fresh sorbent but must recycle large amounts
of regenerated stone therby necessitating a large regenerator and
transport system. At a high fresh feed to total feed ratio one used
more fresh sorbent but does not have to recycle as much regenerated
sorbent therby requiring a small regenerator and transport system.
Thus, the decision is a balance between operating and capital costs -
a question which can only be answerod by an economic evaluation which
is what we have attempted to do.
862
-------
MR. FABER: Before we take a break, if you've got any more
questions written down, pass them to the center. Bill Harvey
and I have a little announcement to make. Most of you are not
aware that our light man is one of the most intellectual and educated
and highest paid light men in the world, Dr. John Minnick, a friend
of ours. And we think that since he's done such a good job in the
first half, that we will be cble to go on through the whole session
without retraining him this afternoon.
863
-------
Appendices
365
-------
Appendix A
Table of Contents
Volume I
PREFACE 1-il
Welcoming Remarks: Richard S. Greeley, 3
The MITRE Corporation/Metrek Division
McLean, Virginia
KEYNOTE ADDRESS 5
Introduction: George Fumich, U.S. Department of 7
Energy, Washington, D. C.
Keynote Address: S. David Freeman, Tennessee Valley 8
Authority, Knoxville, Tennessee
Questions/Responses/Comments ^
DINNER ADDRESS 17
Introduction: Charles A. Zraket, The MITRE 18
Corporation, Bedford, Massachusetts
Dinner Address: John A. Bel ding, U.S. Department 19
of Energy, Washington, D. C.
Questions/Responses/Comments 23
OVERVIEW OF U.S. AND INTERNATIONAL PROGRAMS 31
Introduction of Plenary Session Chairman: 33
V. Ovcharenko, United Nations Center for
Natural Resources, Energy and Transport,
United Nations, New York
An Overview of the Progress in Fit.id Bed 35
Combustion in the United Kingdom: W. G.
Kaye, National Coal Board, Coal Research
Establishment, Stoke Orchard, England
A-l
-------
Paye
The International Energy Agency Program: 43
David H. Broadbent, NCB (IEA Services)
Limited, London, England
Fluidized-Bed Combustion - Overview of the 47
Program of the Federal Republic of
Germany: Rolf Holiqhaus, Kernforschunqs-
anlaqe ("FA), Juelich, Federal Republic
of Germany
The U. S. Department of Energy Program: 57
Steven I. Freedman, U. S. Department
of Energy, Washington, D. C.
The EPA Fluidized-Bed Combustion Program - 62
An Update- D. Bruce Henschel, Industrial
Environmental Research Laboratory, U. S.
Environmental Protection Agency, Research
Triangle Park, North Carolina
The Ohio Energy and Resource Development 77
Authority Program: Eric K. Johnson,
Ohio State Department of Energy, Columbus,
Ohio
Overview of New York State's Fluidized Bed 81
Combustion Program: Richard H. Tourin,
Energy ".esearch an* Development Authority,
New York, New York
The Fluidized-Bed Combustion Proqram of the 87
Tennessee Valley Authority: Harold L. Folken-
berry, Tennessee Valley Authority,
Chattanooga, Tennessee
Overview of EPRI FBC Program: Terry E. Lund, 94
Electric Power Research Institute, Palo
Alto, California
COMMENT by James J. Markowsky 100
American Electric Power Service Corporation,
New York, New York
A-2
-------
Page
APPENDICES 101
Appendix A - Conference Program A-l
Appendix B - Table of Contents - Volume II B-l
Appendix C - Table of Contents - Volume III C-l
Appendix D - Attendees - Alphabetical Listing D-l
Appendix E - Attendees - Alphabetical Organizational E-l
Affiliation
Appendix F - Author Index F-l
A-3
-------
Appendix B
Table of Contents
Volume II
Page
INTRODUCTION
OVERVIEW OF COMMERCIALIZATION ACTIVITIES 1
A State of the Art Resume: Alan A. Smith, 4
Babcock and Wilcox, Ltd., England
The Practical and Commercial Application of 14
Fluid Bed Combustion for Use in Industrial
Boiler Plants: P. B. Caplin, The Energy
Equipment Company, Limited, Energy House,
Olney, Buckinghamshire, England
Status of Fluidized Bed Combustion in Norway 20
and Sweden: Frode Pedersen, 0. Mustad and
Sons A/S, Gjovik, Norway
Industrial Coal Fired Fluidized Bed Boilers 22
and Waste Heat Boilers: Michael J. Virr,
Stone-Platt Fluidfire Limited, Netherton,
West Midlands, England
INDUSTRIAL APPLICATIONS 41
A Comparison of Industrial and Utility Fluidized 44
Bed Combustion Boiler Design Considerations:
J. William Smith, The Babcock & Wilcox
Company, Barberton, Ohio
Industrial Application - Fluidized Bed Combus- 61
tion - Georgetown University: Robert Tracey,
Fluidized Combustion Company; Frederick
Wachtler, Foster Wheeler Energy Corporation,
Livingston, New Jersey; Victor Buck, Pope,
Evans and Robbins, Inc., New York, New York
B-l
-------
-------
Page
Fluidized Bed Combustor For Small Industrial 91
Applications: H. A. Hanson, D. G. DeCoursin,
and D. D. Kinzler FluiDyne Enqineerinq C^rpora-
tion, Minneapolis, Minnesota
The Anthracite Culm/Anthracite Combustion Pro- 107
ijram: John Geffken, U.S. Department of Energy,
Washington, 0. C.
A Comparison of Conventional Oil-Fired and ll"/
Fluidized Bed Coal-Fired Petroleum Refinery
Atmospheric Crude Furnaces: Charles Bliss,
The MITRE Corporation, METREK Division,
McLean, Virginia
A Fluidized Bed Mot Gas Generator for Conversion 145
of Oil-Fired Boilers into Coal-Firing: Prabir
Basu, K. L. Das, and M. K. Chanda, Central
Mechanical Engineering Research Institute,
Durgapur, India
UTILITY APPLICATIONS - Atmospheric Mode 157
The Department of Energy Atmospheric Fluidized 160
Bed Combustion Utility Demonstration Program:
Edward Trexler, U.S. Department of Energy,
Washington, D. C.
Technological Development Priorities of Atmos- 169
pheric Fluidized-Bed Combustion: John E.
Mesko, Pope, Evans and Robbins, Inc., New
York, New York
The Rivesville Installation from the View of the 187
Monongahela Power Company: Homer T. McCarthy,
Allegheny Power Service Corporation, Greens-
burg, Pennsylvania
First Performance Results from the Rivesvillo 191
Multi-Cell Fluidized Bed Boiler: G. Claypoole,
D. Hill, and R. Mineo, Pope, Evans and Robbins,
Inc., Rivesville, West Virginia
B-2
-------
Page
Startup and Initial Operation of the Rives- 203
ville 30 MWe Fluid Bed Boiler: Thomas E.
Stringfellow, Pope, Evans and Robbins, Inc.,
Rivesville, West Virginia; John G. Branam,
The MITRE Corporation, toetrek Division, McLean,
Virginia
Lightoff ^ ' Multicell Fluidized Bed Boilers by 221
Hot Bed Transfer: L. Gasner, 9. Turek, Uni-
versity of Maryland, College Park, Maryland;
C. Aulisio and Ouane Mill, Pope, Evans and
Robbins, Inc., Alexandria, Virginia
Material Handling Systems For FBC'S - Extrapo- 239
lating Rivesville to 600 Megawatts: James J.
Murphy, Allen-Shennan-Hoff Company, Malvern,
Pennsylvania
An Investigation of Alternative Feed Systems For 248
Utility-Scale Fluidized-Bed Steam Generators/
Combustors (200 MWe or Larger Units): B. K. Bis-
was, Foster Wheeier Development Corporation;
J. U. Baley, Foster Wheeler Energy Corporation,
Livingston, New Jersey
The Application of Atmospheric Fluidized Bed Combus- 267
tion for Electrical Power Generation: Thomas W.
Becker, Babcock and Wilcox Company, Barberton,
Ohio
Conceptual Design of a Foster Wheeler Energy 285
Corporation Atmospheric Fluidized T-ed Steam
Generator for Stone and Webster Engineering
Corporation and Tennessee Valley Authority:
Kenneth A. Reed and George Cervenka, Foster
Wheeler Energy Corporation, Livingston, New
Jersey
Conceptual Design and Preliminary Evaluation of a 313
570 MW Electric Power Generating Plant Using a
Babcock & Wilcox Company or a Foster Wheeler
Energy Corporation Atmospheric Fluidized-Bed
Boiler: P. F. Lipari am1 T. C. Wells, Jr.,
Stone & Webster Engineering Corporation, Boston,
Massachusetts
B-3
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Conceptual Design of a 570-MW Combustion Engi- 326
neering, Inc. Atmospheric Fluidized-Bed Steam
Generator: Russell B. Covell, Combustion
Engineering, Inc., Windsor, Connecticut
The Conceptual Design of an AFBC Electric Power 343
Generating Plant: William J. Bradley, Burns
and Roe, Inc., Woodbury, New York
A Comparison of Selected Design Aspects of Three 355
Atmospheric Fluidized Bed Combustion Conceptual
Power Plant Designs: D. N. Garner, W. C. Howe,
and P. S. Dzierlenga, Radian Corporation, McLean,
Virginia
Experiences of Fluidized-Bed Combustion of Peat 375
in Finland: A. Jahkola, Helsinki University of
Technology, Helsinki, Finland
UTILITY APPLICATIONS - Pressurized Mode 385
Design of a Gas Turbine Plant with a Pressuri- 388
zed Fluidized Bed Combustor: H. D. Schilling and
H. Schreckenberg, Bergbau-Forschung GmbH, Essen
Wied, Vereinigte Kesselwerke A. G., Dussel-
dorf. Federal Republic of Germany
The Curtiss-Wright Pressurized Fluidized Bed 401
Pilot Electric Plant: Seymour Moskowitz, Cur-
tiss-Wright Corporation, Wood-Ridge, New Jersey
General Electric Pressurized Fluidized Bed Power 413
Plant Status: R. D. Brooks and J. R. Peterson,
General Electric Company, Schenectady, New York
Conceptual Design of a Coal Fueled, Fluid Bed 434
Combined Cycle Power Plant: D. A. Huber and
R. M. Costello, Burns and Roe Industrial Services
Corporation, Paramus, New Jersey; J. J. Morgan and
A. J. Giramonti, United Technologies Research Center,
Hartford, Connecticut; J. W. Smith, Babcock and
Wilcox Company, Barberton, Ohio
B-4
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ENVIRONMENTAL ASPECTS
Monitoring 459
Development of Environmental Objectives Based 462
on Health and Ecological Effects: B. W.
Cornaby, K. S. Murthy, D.A. Savitz, and
L. Pomerantz, Battelle Columbus Laboratories,
Columbus, Ohio
Design of Ambient Monitoring Programs For Pro- 478
totype Fluidized-Bed Combustion Facilities:
J. M. Allen, D. Ambrose, R. Clark, and P. R.
Sticksel, Battelle Columbus Laboratories,
Columbus, Ohio
Plans and Studies on Flue Gas Cleaning and 493
Particulate Monitoring in PFBC: W. M.
Swift, S. Lee, J. C. Montagna, G. W. Smith,
I. Johnson, G. J., Vogel, and A. A. Jonke,
Argonne National Laboratory, Argonne, Illinois
Emission Characterization and Control 523
Thermodynamic Projections of Trace Element 526
Release in Fluidized-Bed Combustion Systems:
M. A. Alvin, E. P. O'Neill, and D. L.
Keairns, Westinghouse R&D Center, Pittsburgh,
Pennsylvania
Multimedia Pollutant Emissions Data for Fluid- 544
ized Bed Combustion of Coal: K. S. Kurthy,
D. A. Sharp, K. M. Duke, and J. M. Allen,
Battelle Columbus Laboratories, Columbus, Ohio
Effluent Characterization from a Conical Pres- 560
surized Fluid Bed: R. J. Priem, R. J. Rollbuhler,
and R. W. Patch, NASA-Lewis Research Center,
Cleveland, Ohio
NO Reduction by Char in Fluidized Combustion: 577
Janos M. Beer, Adel F. Sarofim, Lisa K. Chan,
and Alice M. Sprouse, Massachusetts Institute
of Technology, Cambridge, Massachusetts
B-5
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Page
Control of Nitric Oxide and Carbon Monoxide 594
Emissions in Fluidized Bed Combustion:
Koya Sakamota, Babcock Hitachi Company,
Hiroshima, Japan
A Model Study for the Development of Low Nox 605
Fluidized-Bed Coal Combustors: Masayuki Horio,
and Iwao Muchi, Nagoya University, Nagoya,
Japan; Shigekatsu Mori, Nagoya Institute
of Technology, Nagoya, Japan
Particulate Control for Pressurized Fluidized- 626
Bed Combustion Processes: D. F. Ciliberti,
D. H. Archer and D. L. Keairns, Westinghouse
R&D Center, Pittsburgh, Pennsylvania
Particle Size in Pressurised Combustors: K. K. 642
Pallai and W. V. Battcock, (Speaker, H. R. Hoy),
National Coal Board Utilisation Research
Laboratory, Leatherhead, England
Characterization of Efflux from a Pressurized 655
Fluidized Bed Combustor: K. L. Bekofske,
C. M. Thoennes, and W. G. Giles, General
Electric Company, Schenectady, New York
SOX Sorbent - Selection 677
Initial Assessment of Alternative S0ฃ Sorbents btfO
for Fluidized-Bed Combustion Power Plants:
R. A. Newby and D.L. Keairns, Westinghouse
R&D Center, Pittsburgh, Pennsylvania
Regenerative Iron Bearing Sorbents for Use in 701
Fluidized Bed Combustion: P. J. hcGauley,
Pope, Evans and Robbins, Inc., New York,
New York; A. A. Dor and F. A. Bumje, The
Hanna Mining Company, Cleveland, Ohio
B-6
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Page
Effects of Coal Composition and Ash Reinjection 729
on Sulfur Retention Burning Lignite and Wes-
tern Subbituminous Coals: Gerald M. Goblirsch,
Richard W. Fehr, and Everett A. Sondreal,
Grand Forks Energy Research Center, Grand Forks,
North Dakota
SOX Sorbent - Utilization 745
The Prediction of Limestone Requirements for 748
SC>2 Emission Control in Atmospheric Pres-
sure Fluidized-Bed Combustion: Robert B.
Snyder, W. Ira Wilson, and Irving Johnson,
Argonne National Laboratory, Argonne, Illinois
Limestone Utilization Optimization in Fluidized 763
Bed Boilers: Larry L. Gasner and Scott E.
Setesak, University of Maryland, College Park,
Maryland
The Mechanism of the Salt Additive Effect on the 776
S02 Reactivity of Limestone: J. Shearer,
Irving Johnson, and C. Turner, Argonne
National Laboratory, Argonne, Illinois
Modelling Desulfurization Reactions in Fluidized 787
Bed Combustors: C. Georgakis, J. Szekely,
C. W. Chang, J. W. Chrostowski, and T. Trinh,
Massachusetts Institute of Technology,
Cambridge, Massachusetts
SOX Sorbent - Disposal 797
Potential Uses for the Residue from the Fluidized 800
Bed Combustion Process: Richard H. Miller, Sr.,
Valley Forge Laboratories, Inc., and Villanova
University, Devon, Pennsylvania
Leaching Experiments on Soil and Mine Spoil 821
Treated vปith Fluidized Bed Combustion Waste:
R. C. Sidie, W. L. Stout, J. L. Hern, and
0. L. Bennett, U.S. Department of Agriculture,
Morgantown, West Virginia
B-7
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Page
Characterization of Fluidized Bed Combustion 833
Waste Composition and Variability as they
Relate to Disposal on Agricultural Lands:
J. L. Hern, W. L. Stout, R. C. Sidle, and
0. L. Bennett, U.S. Department of Agriculture,
Morgantown, West Virginia
Impact 843
Environmental Impact of the Disposal of Pro- 846
cessed and Unprocessed FBC Bed Material and
Carry-Over: C. C. Sun, C. H. Peterson, and
D. L. Keairns, Westinghouse R&D Center,
Pittsburgh, Pennsylvania
Assessment of the Impact of S02, NOX, and 875
Particulate Emission Standards on Fluidized-Bed
Combustion System Design and Energy Costs:
R. A. Newby, N. Ulerich, E. P. O'Neill, D. F.
Ciliberti, and D. L. Keairns, Westinghouse
R&D Center, Pittsburgh, Pennsylvania
Relative Environmental Impact of Two 570 MW 892
Atmospheric Flซ:idized-Bed Electric Power
Generating Plants Compared to a Pulverized
Coal Fired Plant Equipped with a Wet Limestone
Flue Gas Desulfurization System: R. C. Stone,
Stone & Webster Engineering Corporation, Boston,
Massachusetts
Environmental Analysis of the General Electric PFB 898
Combined Cycle Power Plant Design: V. H. Lucke
and K. R. Murphy, General Electric Company,
Schenectady, New York
APPENDICES 915
Appendix A - Table of Contents - Volume I A-l
Appendix B - Table of Contents - Volume III B-l
Appendix C - Author Index C-l
B-8
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Appendix C
Author Index
ALLEN, J. M.
Design of Ambient Monitoring Pro-
grams for Prototype FTuidized-
Bed Combustion Facilities,
Vol. II, p. 478
Multimedia Pollutant Emissions
Data for Fluidized-Bed Com-
bustion of Coal, Vol. II, p.544
ALVIN, M. A.
Thermodynamic Projections of
Trace Element Release in Fluid-
ized-Bed Combustion Systems,
Vol. II, p. 526
AMBROSE, D.
Design of Ambient Monitoring Pro-
grams for Prototype Fluidized-
Bed Combustion Facilities,
Vol. II, p. 478
ARCHER, D. H.
Particulate Control for Pres-
surized Fluidized-Bed Combus-
tion Processes, Vol. II, p. 626
AULISIO, C.
Lightoff of Multicell
Fluidized Bed Boilers by Hot Bed
Transfer, Vol. II, p. 223
Results of Recent Test Pro-
gram Related to AKB Combustion
Efficiency, Vol. Ill, p. 82
BACHALO, WILLIAM D.
Particle Field Diagnostics
Systems for Fluidized Bed Com-
bustion Facilities, Vol. Ill,
p. 362
BALEY, 0. U.
An Investigation of Alternative
Feed Systems for Utility-Scale
Fluiciized-Bed Steam Generators/
Combustors (*GG Mw'e or Larger
units), vol. li, p. 246
BAft-LUritN, A.
Fluid Dynamic Modelling ot
Huidized bed Combustors,
Vol. Ill, p. 45B
BARON, K. E.
A Model of Coal Combustion
in Fluidized Bed Combustors,
Vol. Ill, p. 437
BASU, PRABIR
A Fluidized Bed Hot Gas Generator
for Conversion of Oil-Fired Boilers
into Coal-Firing, Vol. II, p. 145
BATTCOCK, WHALLEY V.
Particle Size in Pressurized
Combustors, Vol. II, p. 642
BECKER, THOMAS W.
The Application of Atmospheric
Huidized Bed Combustion for
Electrical Power Generation,
Vol. II, p. 267
BEER, JANOS, M.
A Model of Coal Combustion in
Fluidized Bed Conbustors, Vol. Ill,
p. 437
NO Reduction by Char in Fluid-
ized Combustion, Vol. II, p. B77
BEKOFSKE, K. L.
Characterization of Efflux from
A Pressurized Fluidized Bed Com-
bustor, Vol. II, p. 655
C-l
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BELTRAN, A. M.
Turbine Materials Corrosion in
the Coal-Fired Combined Cycle,
Vol. Ill, p. 714
BENNETT, 0. L.
Leaching Experiments on Soil and
Mine Spoil Treated with Fluidized
Bed Combustion Waste, Vol. II,
p. 821
Characterization of Fluidized Bed
Combustion Waste Composition and
Variability as they Relate to Dis-
posal on Agricultural Lands,
Vol. II, p. 833
BERKOWITZ, DAVID A.
Dynamic Modelling, Testing, and
Control of Fluidized Bed Systems,
Vol. Ill, p. 488
BERTRAND, R. R.
Evaluation of a Granular Bed
Filter for Particulate Control
in Fluidized Bed Combustion,
Vol. Ill, p. 504
Pressurized Huidized Bed Coal
Combustion and Sorbent Regen-
eration, Vol. Ill, p. 756
BIANCO, J. H.
An Engineering Study on the Re-
generation of .culfated Additive
from a Fluid-i.ed-Bed Coal-Fired
Power Plant, Vol. Ill, p. 832
BISWAS, D. K.
An Investigation of Alternative
Feed Systems for Utility-Scale
Fluidized-Bed Steam Generators/
Combustors (200 MWe or Larger
Units), Vol. II, p. 248
BLISS, CHARLES
A Comparison of Conventional Oil-
Fired and Fluidized Bed Coal-
Fired Petroleum Refinery
Atmospheric Crude Furnaces,
Vol. II, p. 117
BONK, D. L.
B&W/EPRI's 6' x 6' Fluidized
Bed Combustion Development
Facility: An Overview, Vol. Ill,
p. 24
BORGHI, G.
A Model of Coal Combustion in
Fluidized Bed Combustors, Vol. Ill,
p. 437
BRADLEY, JEFFREY F.
Multiple Jet Particle Collec-
tion in a Cyclone by Reheating
Fluidized Bed Combustion Products,
Vol. Ill, p. 607
BRADLEY, WILLIAM J.
The Conceptual Design of an AFBC
Electric Power Generating Plant,
Vol. II, p. 343
BRANAM, JOHN G.
Startup and Initial Operation of
the Rivesville 30 MWe Fluid Bed
Boiler, Vol. II, p. 203
BROADBENT, DAVID H.
The International Energy
Agency Program, Vol. I, p.
43
A Technical Description of
the Plant Design and Project
Progress Report, Vol. Ill, p.
310
BROOKS, R. D.
General Electric Pressurized
Fluidized Bed Power Plant Status,
Vol. II, p. 413
C-2
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BUCK, VICTOR
Industrial Application - Fluidi/ed
Bed Combustion - Georgetown
University, Vol. II, p. 61
BUNGE, F. H.
Regenerative Iron Bearing Sorbents
for Use in Flnidi zed Bed Combus-
tion, Vol. II, p. 701
BUSH, JOHN
High Temperature, High Pressure
Electrostatic Precipitation,
Vol. Ill, p. 640
BYAM, J. W.
Atmospheric Fluidized Bed Com-
ponent Test and Integration
Facility - An Update, Vol. Ill,
P. 4
CALVERT, S.
Granular Bed Filters for Particu-
late Removal at High Temperature
and Pressure, Vol. Ill, p. 516
CAPLIN, P. B.
The Practical ana Commercial
Application of Fluid Bed Combus-
tion for Us? in Industrial
Boiler Plants, Vol. II, p . 14
CATIPOVIC N.
Solid Tracer Studies in a Tube-
Filled Fluidized Bed, Vol. Ill,
p. 135
CERVENKA, GEORGE G.
Conceptual Design of a Foster
Wheeler Energy Corporation Atmos-
pheric Fluidizsd Bed Steam Gen-
erator for Stone and Webster
Engineering Corporation and
Tennessee Valley Authority,
Vol. II, p. 285
CHAN, LISA K.
NO Reduction by Char in Fluidized
Bed Combustion, Vol. II, p. 577
CHANDA, M. K.
A Fluidized Bed Hot Gas Generator
for Conversion of Oil -Fired Boil-
ers into Coal-Firing, Vol. II,
p. 145
CHANG, C. VI.
Modelling Oesulf urization Reac-
tions in Fluidized Bed Combustors,
Vol. II, p. 787
CHEN, ,). C.
Centrifugal Fluidized Bed
Combustion, Vol. Ill, p. 288
CHEN, JAMES M.
Regeneration of Lime-Based Sor-
bents in a Kiln with Solid
Reductants, Vol. Ill, p. 798
CHERRINGTON, 0. C.
Industrial Application of Fluid-
ized Bed Combustion - Single Tube
Heat Transfer Studies, Vol. Ill,
p. 184
CHROSTOWSKI, J. W.
Modelling Desulfurization Reac-
tions in Fluidized Bed Combus-
tors, Vol. II, p. 787
CILIBERTI, D. F.
Particulate Control for Pressur-
ized Fluidized-Bed Combustion
Processes, Vol. II, p. 626
Assessment of the Impact of
NOX, and Particulate Emission
Standards on Fluidized-Bed Comlus-
tion System Desiqn and Energy
Costs, Vol. II, p. 875
CLARK, R.
Design of Ambient Monitoring Pro-
grams for Prototype Fluidized-Bed
Combustion Facilities, Vol. II,
p. 478
CLAYPOOLE, GEORGE
First Performance Results from
the Ri.vesville Multi-Cell
Fluidized Bed Boiler, Vol. II, p. 191
C-3
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CORNABY, B. W.
Development of Environmental
Objectives Based on Health and
Ecological Effects, Vol. II,
p. 462
COSTELLO, R. M.
Conceptual Design of a Coal
Fueled, Fluid Bed Combined
Cycle Power Plant, Vol. II,
p. 434
An Engineering Study on the
Regeneration of Sulfated Addi-
tive from a Fluidized-Bed Coal-
Fired Power Plant, Vol. Ill,
p. 832
COVELL, RUSSELL B.
Conceptual Design of a 570-MW
Combustion Engineering, Inc.
Atmospheric Fluidized-Bed Steam
Generator, Vol. II, p. 326
CRAWFORD, R. W.
Pressurized Fluidized-Bed Combus-
tion Component Test and Integra-
tion Unit Design Status, Vol. Ill,
p. 102
DAS, K. L.
A Fluidized Bed Hot Gas Generator
for Conversion of Oil-Fired Boil-
ers into Coal-Firina, Vol. II,
p. 145
DECKER, N.
Thermal Stresses and Fatigue of
Heat Transfer Tubes Immersed in
a Fluidized Bed Combustor,
Vol. Ill, p. 700
DECOURSIN, D. G.
Fluidized Bed Combustor for Small
Industrial Applications, Vol. II,
p. 91
DEGANI, DAVID.
Particulate Removal from Hot
Gases Using the Fluidized Bed
Cross-Flow Filter, Vol. Ill,
P. 551
DIVILIO, R.
Results of Recent Test Program
Related to AFB Combustion
Efficiency, Vol. Ill, p. 82
DOR, A. A.
Regenerative Iron Bearing Sor-
bents for Use in Fluidized Bed
Combustion, Vol. II, p. 701
DOWDY, T. E.
B&W/EPRI's 6' x 6' Fluidized Bed
Combustion Development Facility:
An Overview, Vol. Ill, p. 24
DREHMEL, T. t.
Granular Bed Filters for Particu-
late Removal at High Temperature
and I ressure,. Vol. Ill, p. 516
DUKE, K. M.
Multimedia Pollutant Emissions
Data for Fluidized-Bed Combustior
of Coal, Vol. El, p. 544
DZIERLEN'GA, P'. STANLEY
A Comparison of Selected Design
Aspects of Three Atmospheric
Fluidized Bed Combustion Con-
ceptual Power Plant Designs,
Vol. II, p. 355
FALKENBERRY, HAROLD L.
The Fluidized-Bed Combustion
Program of the Tennessee Valley
Authority, Vol. I, p. 87
FARBER, GERALD
Regeneration of Lime-Based
Sorbents in s. Kiln with Solid
Reductants, ฅol. Ill, p. 798
C-4
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FEHR, RICHARD W.
Effects of Coal Composition
and Ash Reinjection on Sulfur
Retention Burning Lignite and
Western Subbituminous Coals,
Vol. II, p. 729
FELDMAN, PAUL
High Temperature, High Pressure
Electrostatic Precipitation,
Vol. Ill, p. 640
FELTON, G. W.
Pneumatic Solids Injector and
Start-Up Burner for Battelle's
Multisolid Fluidized-Bed Com-
bustion (MS-FBC) Process,
Vol. Ill, p. 241
Battelle's Multisolid Fluidized-
Bed Combustion Process, Vol. Ill,
p. 223
FITZGERALD, T.
Solid Tracer Studies in a Tube-
Filled Fluidized Bed, Vol. Ill,
p. 135
FRAAS, ARTHUR
Atmospheric Fluidized Bed Combus-
tion Technology Test Unit for
Industrial Cogeneration Plants,
Vol. Ill, p. 55
FREEDMAN, STEVEN I.
The U.S. Department of Energy
Program, Vol. I, p. 57
GARNER, DONALD N.
A Comparison of Selected Design
Aspects of Three Atmospheric
Fluidized Bed Combustion Con-
ceptual Power Plant Designs,
Vol. II, p. 355
GASNER, L.
Lightoff of Multicell Fluidized
Bed Boilers by Hot Bed Transfer,
Vol. II, p. 223
Limestone Utilization Optimiza-
tion in Fluidized Bed Boilers,
Vol. II, p. 763
GEFFKEN, JOHN
The Anthracite Culm/Anthracite
Combustion Program, Vol. II,
p. 107
GEORGAKIS, C.
Modelling Desulfurization Reac-
tions in Fluidized Bed Combustors,
Vol. II, p. 787
GIAMMAR, R. D.
Pneumatic Solids Injector and
Start-Up Burner for Battelle's
Multisolid Fluidized-Bed Com-
bustion (MS-FBC) Process, Vol III,
p. 241
GILES, W.
Characterization of Efflux from
a Pressurized Fluidized Bed
Combustor, Vol. II, p. 655
GIRAMONTI, A. J.
Conceptual Design of a Coal
Fueled, Fluid Bed Combined Cycle
Power Plant, Vol. II, p. 434
GLICKSMAN, L.
Thermal Stresses and Fatigue cf
Heat Transfer Tubes Immersed in
a Fluidized Bed Conbustor.
Vol. Ill, p. 700
Fluid Dynamic Modelliuq of
Fluidized Bed Combustors.
Vol. Ill, D. 458
GLUKHOMANYUK. A. M.
Research of Gas Combustion in
Fluidized Bed Plants. Vol. III.
n. 268
C-5
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GOBLIRSCH, GERALD M.
Effects of Coal Composition
and Ash Reinjection on Sulfur
Retention Burning Lignite and
Western Subbituminous Coal,
Vol. II, p. 729
GOLDMAN, J.
FBC-Modelling and Data Base,
Vol. Ill, p. 406
GOLAN, L. P.
Industrial Application of Fluid-
ized Bed Combustion-Single Tube
Heat Transfer Studies, Vol. Ill,
p. 184
GREGORY, M. W.
Evaluation of a Granular Bed
Filter for Particulate Control
in Fluidized Bed Combustion,
Vol. Ill, p. 504
GRIMM, U.
Fluidized-Bed Combustion of
Lignite and Lignite Refuse,
Vol. Ill, p. 211
GUILLORY, J. L.
Filtration Performance of a
Moving Bed Granular Filter:
Experimental Cold Flow Data,
Vol. Ill, p. 567
GUTFINGER, CHAIM
Particulate Removal from Hot
Gases Using the Fluiriized Bed
Cross-Flow Filter, Vol. Ill,
p. 551
HALOW, J. S.
Fluidized-Bed Combustion of
Lignite and Lignite Refuse,
Vol. Ill, p. 211
HAMMITT, F. G.
Industrial Application of Fluid-
ized Bed Combustion-Sinole Tube
Heat Transfer Studies, Vol. Ill,
p. 184
HANSON, H. A.
Fluidized Bed Combustor for Small
Industrial Applications, Vol. II,
p. 91
HAZARD, H. R.
Pneumatic Soliils Injector and
Start-Up Burner for Battelle's
Multisolid Fluidized-Bed Com-
bustion (MS-FBC) Process, Vol. Ill,
p. 241
HENSCHEL, D. BRUCE
The EPA Fluidized-Bed Combustion
Program - An Update, Vol. I, p. 62
HERN, J. L.
Characterization of Fluidized Bed
Combustion Waste Composition and
Variability as They Relate to
Disposal on Agricultural Lands,
Vol. II, p. 833
Leaching Experin -nts on Soil and
Mine Spoil Trea'.ad with Fluidized
Bed Combustion Waste, Vol. II,
p. 821
HILL, DUANE
First Performance Results From
the Rivesville Multi-Cell
Fluidized Bed Boiler, Vol. II,
p. 191
Lightoff of Multicell Fluidized
Bed Boilers by Hot Bed Transfer,
Vol. II, p. 223
HODGES, J.
A Model of Coal Combustion in
Fluidized Bed Combustors, Vol. Ill,
p. 437
HOKE, R. C.
Evaluation of a Granular Bed
Filter for Particulate Control
in Fluidized Bed Combustion,
Vol. Ill, p. 504
C-6
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HOKE, RONALD C.
Pressurized Fluidized Bed Coal
Co:nbustion and Sorbent Regenera-
tion, Vol. Ill, p. 756
HOLCOMB, R. S.
Atmospheric Fluidized Bed Com-
bustion Technology Test Unit for
Industrial Cogeneration Plants,
Vol. Ill, p. 55
HOLIGHAUS, ROLF
Fluidized-Bed Combustion -
Overview of the Program of the
Federal Republic of Germany,
Vol. I, p. 47
MORGAN, J. J.
Conceptual Design of a Coal
Fueled, Fluid Bed Combined Cycle
Power Plant, Vol. II, p. 434
HORIO, MASAYUKI
A Model Study for the Develop-
ment of Low NOX Fluidized-Bed
Coal Combustors, Vol. II, p. 605
HOWARD, J. R.
Combustion Experiments Within a
Rotating Fluidized Bed, Vol. Ill,
p. 275
HOWE, WILLIAM C.
A Comparison of Selected Design
Aspects of Three Atmospheric
Fluidized Bed Combustion Concei.
tual Power Plant Designs,
Vol. II, p. 355
HOY, H. R.
Further Experiments on the Pilot-
Scale Pressurized Combustor at
Leatherhead, Vol. Ill, p. 123
HUBER, D. A.
Conceptual Design of a Coal
Fueled, Fluid Bed Combined
Cycle Power Plant, Vol. il,
p. 434
HUBER, D. A.
An Engineering Study on the
Regeneration of Sulfated Additive
from a Fluidized-Bed Coal-Fired
Power Plant, Vol. Ill, p. 832
HUGHES, R.
Fluid Dynamic Modelling of
Fluidized Bed Combustors, Vol. Ill,
p. 458
JAHKOLA, ANTERO
Experiences of Fluidized-Bed
Combustion of Peat in Finland,
Vol. II, p. 375
JOHNSON, ERIC K.
The Ohio Energy and Resource
Development Authority Program,
Vol. I, p. 77
JOHNSON, IRVING
Plans and Studies on Flue Gas
Cleaning and Particulate Monitor-
ing in PFBC, Vol. II, p. 493
The Prediction of Limestone
Requirements for S02 Emission
Control in Atmospheric Pressure
Fluidized-Bed Combustion, Vol. II,
p. 748
The Mechanism of the Salt Additive
Effect on the $03 Reactivity of
Limestone, Vol. II, p. 776
JONKE, A.A.
Plans and Studies on Flue Gas
Cleaning and Particulate Moni-
toring in PFBC, Vol. II, p. 493
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
Economic Feasibility of Regen-
erating Sulfated Limestones,
Vol. Ill, p. 851
C-7
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JOVANOVIC, G.
Solid Tracer Studies in a Tube-
Filled Fluidized Bed, Vol. Ill,
p. 135
KATTA, S.
Evaluation of Sorbent Regenera-
tion Processes for AFBC and
PFBC, Vol. Ill, p. 811
KAYE, W. G.
An Overview of the Progress in
Fluid Bed Combustion in the
United Kingdom, Vol. I, p. 35
KEAIRNS, D. L.
Particulate Control for Pressur-
ized Fluidized-Bed Combustion
Processes, Vol. II, p. 626
Initial Assessment of Alterna-
tive S(>2 Scrbents for Fluid-
ized-Bed Combustion Po^er
Plants, Vol. II, p. 680
Environmental Impact of the
Disposal of Processed and Unpro-
cessed F3C Bed Material and
Carry-over, Vol. II, p. 846
Assessment of the Impact of S02,
NOX, and Particulate Emission
Standards on Fluidized-Bed Com-
bustion System Design and Energy
Costs, Vol. II, p. 875
Thermodynamic Projections of
Trace Element Release in
Fluidized-Bed Combustion Systems,
Vol. II, p. 526
Evaluation of Sorbent Regenera-
tion Processes for AFBC and
PFBC, Vol. Ill, p. 811
KINZLER, D. D.
Fluidized Bed Combustor for
Small Industrial Applications,
Vol. II, p. 91
KIVIAT, G.
The Effects of Finned Tubing
on Fluidized Bed Performance,
Vol. Ill, p. 156
LaNAUZE, R. D.
High Temperature Corrosion of
Metals and Alloys in Fluidized
Bed Combustion Systems, Vol. Ill,
p. 682
LARKIN, ROBERT
A Particulate Sampling System
for Pressurized Fluidized Bed
Combustors, Vol. Ill, p. 379
LEE, S. H. D.
Plans and Studies on Flue Gas
Cleaning and Particulate Monitor-
ing in PFBC, Vol. II, p. 493
LEVY, E. K.
Centrifugal Fluidized Bed
Combustion, Vol. Ill, p. 288
LIPARI, P. F.
Conceptual Design and Preliminary
Evaluation of a 570 MW Electric
Power Generating Plant Using a
Babcock & Wilcox Company or a
Foster Wheeler Energy Corporation
Atmospheric Fluidized-Bed Boiler,
Vol. II, p. 313
LIU, K. T.
Battelle's Multisolid Fluidized-
Bed Combustion Process, Vol. Ill,
p. 223
LOUIS, J. F.
FBC-Modelling and Data Base,
Vol. Ill, p. 406
LOWELL, CARL E.
Erosion/Corrosion of Turbine
Airfoil Materials in the High-
Velocity Effluent of a Pres-
surized Fluidized Coal
Combustor, Vol. Ill, p. 660
C-8
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LUCKE, V. H.
Environmental Analysis of the
General Electric PFB Combined
Cycle Power Plant Design,
Vol. II, p. 898
LUND, TERRY E.
B&W/EPRI's 6' x 6' Fluidized Bed
Combustion Development Facility:
An Overview, Vol. Ill, p. 24
Overview of EPRI FBC Program,
Vol. I, p. 94
LUTHRA, K. L.
Turbine Materials Corrosion in
the Coal-Fired Combined Cycle,
Vol. Ill, p. 714
MAKHORIH, K. YE
Research of Gas Combustion in
Fluidized Bed Plants, Vol. Ill,
p. 268
MARTIN, N. W.
Centrifugal Fluidized Bed
Combustion, Vol. Ill, p. 288
MASTERS, WILLIAM
A Particulate Sampling System
for Pressurized Fluidized Bed
Combustors. Vol. III. p. 379
MCCARRON. R. L.
Turbine Materials Corrosion in
the Coal-Fired Combined Cycle.
Vol. Ill, p. 714
MCCARTHY, HOMES
The Rivesville Installation
from the View of the Mnnonnahelป
Power Company. Vol. II. p. 187
MCGAULEY, P. J.
Regenerative Iron Bearing Sorbents
for Use in Fluidized Bed Combus-
tion, Vol. II, p. 701
MEI, J. S.
Fluidiied-Bed Combustion of
Lignite and Lignite Refuse,
Vol. Ill, p. 211
MESKO, JOHN E.
Technological Development Priori-
ties of Atmospheric Fluidized-
~"ed Combustion, Vol. II, p. 169
METCALFE, C. I.
Combustion Experiments Within a
Rotating Fluidized Bed, Vol. Ill,
p. 275
MILLER, GABRIEL
The Effects of Finned Tubing on
Fluidized Bed Performance,
Vol. Ill, p. 156
MILLER, KICHAHD H.
Potential Uses for the Residue
from the Fluidized Bed
Combustion Process, Vol. II,
p. 800
MINEU, RONALD
First Performance Results from
the Rivesville Multi-Cell
Fluidized Bed Boiler, Vol. II,
p. 191
MONTAGNA, J. C.
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, P. 776
Plans and Studies on F^ue Gas
Cleaning and Particulate Monitoring
in PFBC, Vol. II, p. 493
Economic Feasibility of ฐegen-
erating Sulfated Limestones,
III, p.
C-9
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MORI, SHIGEKATSU
A Model Study for the Develop-
ment of Low NOX Fluidized-Bed
Coal Conibustors, Vol. II, p. 605
MORTON, J. W.
An Engineering Study on the Re-
generation of Sulfated Additive
from a Fluidized-Bed Coal-Fired
Power Plant, Vol. Ill, p. 832
MOSKOWITZ, SEYMOUR
The Curtiss-Wright Pressurized
Fluidized-Bed Pilot Electric
Plant, Vol. II, p. 401
MOSS, G.
Progress in the Development of
the Desulfurizing Gasifier,
Vol. Ill, p. 300
MUCH I, Iwao
A Model Study for the Develop-
ment of Low NOX Fluidized-Bed
Coal Combustors, Vol. II, p. 605
MULY, E. C.
Particulate Analysis Instrumenta-
tion for Advanced Combustion
Systems, Vol. Ill, p. 347
MURPHY, JAMES J.
Material Handling Systems for
FBC'S - Extrapolating Rivesville
to 600 Megawatts, Vol. II, p. 239
MURPHY, K. R.
Environmental Analysis of the
General Electric PFB Combined
Cycle Power Plant Design,
Vol. II, p. 898
^
MURTHY, K. S.
Multimedia Pollutant Emissions
Data for Fluidized-Bed Combus-
tion of Coal, Vol. II, p. 544
Development of Environmental
Objectives Based on Healtii and
Ecological Effects, Vol. II,
p. 462
MACK, H.
Battelle's Multisolid Fluidized-
Bed Combustion Process, Vol. Ill,
p. 223
NEWBY, R. A.
Assessment of the Impact of SO? ,
NOX, and Particulate Emission
Standards on Fluidized-Bed Combus-
tion System Design and Energy
Costs, Vol. II, p. 875
Initial Assessment of Alternative
S0ฃ Sorbents for Fluidized-Bed
Combustion Power Plants,
Vol. II, p. 680
Evaluation of Sorbent Regeneration
Processes for AFBC and PFBC,
Vol. Ill, p. 811
NORCROSS, WILLIAM R.
Industrial Fluidized-Bed Program:
A Status Review, Vol. Ill, p. 33
NUNES, F. F.
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
NUTKIS, M. S.
Evaluation of a Granular Bed Fil-
ter for Particulate Control in
Fluidized Bed Combustion, Vol. Ill,
p. 504
Pressurized Fluidized Bed Coal
Combustion and Sorbent Regenera-
tion, Vol. Ill, p. 756
C-10
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O'NEILL, E. P.
Assessment of the Impact of SC>2,
NOX, and Particulate Emission
Standards on Fluidized-Eod Com-
bustion System Design ana Energy
Costs, Vol. II, p. 875.
Thermodynamic Projections of
Trace Element Release ir. Fluid-
ized-Bed Combustion, Vol. II,
p. 526
PALLAI, K. K.
Particle Size in Pressurized
Combustors, Vol. II, p. 642
PARKER, RICHARD D.
Granular Bed Filters for Partic-
ulate Removal at High Tempera-
ture and Pressure, Vol. Ill,
p. 516
PATCH, R. W.
Effluent Characterization from
a Conical Pressurized Fluid Bed,
Vol. II, p. 560
PATTERSON, RONALD G.
Granular Bed Filters for Partic-
ulate Removal at High Tempera-
ture and Pressure, Vol. Ill,
p. 516
PEDERSEN, FRODE
Status of Fluidized Bed Combus-
tion in Norway and Sweden,
Vol. II, p. 20
PELLOUX, R.
Thermal Stresses and Fatigue of
Heat Transfer Tubes Immersed in
a Fluidized Bed Combustor,
Vol. Ill, p. 700
PETERSON, C. H.
Environmental Impact of the Dis-
posal of Processed and Unpro-
cessed FBC Bed Material and
Carry-over, Vol. II, p. 846
PETERSON, J. R.
General Electric Pressurized
Fluidized Bed Power Plant
Status, Vol. II, p. 413
PODOLSKI, W. F.
Pressurized Fluidized Bed Combus-
tion Component Test and Integra-
tion Unit (CTIU): Design Status,
Vol. Ill, p. 102
POMERANTZ, L.
Development of Environmental
Objectives Based on Health and
Ecological Effects, Vol. II, p. 462
PORTER, JAMES H.
ERCO's Fluid-Bed Combustion
Development Facility, Vol. Ill,
p. 73
PRIEM, R. J.
Effluent Characterization from a
Conical Pressurized Fluid Bed,
Vol. II, p. 560
RASSIWALLA, F. M.
Thermodynamics of Regenerating
Sulfated Lime, Vol. Ill, p. 740
RAVEN, P.
Further Experiments on the Pilot-
Scale Pressurized Combustor at
Leatherhead, Vol. Ill, p. 123
RAY, ASOK
Dynamic Modeling, Testing, and
Control of Fluidized Bed Systems,
Vol. Ill, p. 488
REED, KENNETH A.
Conceptual Design of a Foster
Wheeler Energy Corporation Atmos-
pheric Fluidized Bed Steam Gener-
ator for Stone and Webster Engi-
neering Corporation and Tennessee
Valley Authority, Vol. II, p. 285
C-ll
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REtD, R.
Results of Recent Test Program
Related to AFB Combustion
Efficiency, Vol. Ill, p. 82
REHMAT, A
A Mechanistic Model to Explain
Ash Agglomeration in Fluidized
Bed Combustors and Gasifiers,
Vol. Ill, p. 475
RICE, R. L.
Fluidized-Bed Combustion of
Lignite and Lignite Refuse,
Vol. Ill, p. 475
RINGWALL, CARL G.
A High-Temperature High-Pres-
sure Isokinetic/Isothermal
Sampling System for Pressur-
ized Fluidized Bed Applications,
Vol. Ill, p. 326
ROBERTS, A. G.
Further Experiments on the
Pilot-Scale Pressurized Com-
bustor at Leatherhead, Vol. Ill,
p. 123
ROB INSCN, MYRON
High Temperature, High Pressure
Electrostatic Precipitation,
Vol. Ill, p. 640
ROGERS, E.A.
High Temperature Corrosion of
Metals and Alloys in Fluidized
Bed Combustion Systems,
Vol. Ill, p. 682
ROLLBUHLER, R. J.
Effluent Characterization from
a Conical Pressurized Fluid Bed,
Vol. II, p. 560
ROME, ANNE P.
Erosion/Corrosion of Turbine
Airfoil Materials in the High-
Velocity Effluent of a Pres-
surized Fluidized Coal
Combustor, Vol. Ill, p. 660
RUTH, L. A.
Pressurized Fluidized Bed Coal
Combustion and Sorbent
Regeneration, Vol. Ill, p. 756
SAKAMOTO, KOYA
Control of Nitric Oxide and
Carbon Monoxide Emissions in
Fluidized Bed Combustion,
Vol. II, p. 594
SAROFIM, ADEL F.
NO Reduction by Char in Fluid-
ized Combustion, Vol. II, p. 577
A Model of Coal Combustion in
Fluidized Bed Combustnrs,
Vol. Ill, p. 437
SAVITZ, D. A.
Development of Environmental
Objectives Based on Health and
Ecological Effects, Vol. II,
p. 462
SAXENA, S. C.
A Mechanistic Model to Explain
Ash Agglomeration in Fluidized
Bed Combustors and Gasifiers,
Vol. Ill, p. 475
SCHILLING, H. D.
Design of a Gas Turbine Plant
with a Pressurized Fluidized Bed
Conbustor, Vol. II, p. 388
SCHRECKENBERG, HEINZ
Design of a Gas Turbine Plant
with a Pressurized Fluidized
Bed Combustor, Vol. II, p. 338
SETESAK, SCOTT E.
Limestone Utilization Optimization
in Fluidized Bed Boilers, Vol. II,
p. 763
SHACKLETON, MICHAEL A.
Feasibility of Barrier Filtration
Using Ceramic Fibers, Vol. Ill,
p. 620
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SHARP, D. A.
Multimedia Pollutant emissions
Data for Fluidized-Bed Combus-
tion of Coal, Vol. II, p. 544
SHEARER, J.
The Mechanism of the Salt Addi-
tive Effect on the S02 Reac-
tivity of Limestone, Vol. II,
p. 776
SHEN, T.
Thermal Stresses and Fatigue of
Heat Transfer Tubes Immersed
in a Fluidized Bed Combustor,
Vol III, p. 700
SHEN, MING-SHING
Regeneration of Lime-Based
Sorbent in a Kiln with Solid
Reductants, Vol. Ill, p. 798
SIDLE, R. C.
Leaching Experiments on Soil
and Mine Spoil Treated with
Fluidized Bed Combustion Waste,
Vol. II, p. 821
Characterization of Fluidized
Bed Combustion Waste Composition
and Variability as They Relate
to Disposal on Agricultural
Lands, Vol. II, p 833
SMITH, ALAN A.
A State of the Art Resume,
Vol. II, p. 4
SMITH, G. W.
Plans and Studies on Flue Gas
Cleaning and Particulate Moni-
toring in PFBC, Vol. II, p. 493
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
SMITH, J. WILLIAM
A Comparison of Industrial and
Utility Fluidized Bed Combustion
Boiler Design Considerations,
Vol. II, p. 44
Conceptual Design of a Coal
Fueled, Fluid Bed combined
Cycle Power Plant, Vol. II,
p. 434
SMYK, E. B.
Economic Feasibility of Regen-
erating Sulfated Limestones,
Vol. Ill, p. 851
Development of a Process
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
SNYDER, ROBERT B.
Th? Prediction of Limestone
Requirements for SO? Emission
Control in Atmospheric Pressure
Fluidized-Bed Combustion, Vol. II,
p. 748
SONDREAL, EVERETT A.
Effects of Coal Composition and
Ash Reinjection on Sulfur Retention
Burning Lignite and Western Sub-
bituminous Coals, Vol. II, p. 729
SPACE, C. C.
Atmospheric Fluidized Bed Com-
ponent Test and Integration
Facility - An ipdate, Vol. Ill,
P. 4
SPACIL, H. S.
T-irbine Materials Corrosion in
the Coal-Fired Combined Cycle,
Vol. Ill, p. 714
SPROUSE, ALICE M.
NO Reduction by Char in Fluidized
Combustion, Vol. II, p. 577
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STEINBERG, MEYER
Regeneration of Lime-Based
Sorbents in a Kiln With Solid
Reductants, Vol. Ill, p. 798
STICKSEL, P.
Design of Ambient Monitoring
Programs for Prototype for
Fluidized-Bed Combustion Facil-
ities, Vol. II, p. 478
STONE, R. C.
Relative Environmental Impact of
Two 570 MW Atmospheric Fluidized-
Bed Electric Power Generating
Plants Compared to a Pulverized
Coal Fired Plant Equipped with a
Wet Limestone Flue Gas Desulfuri-
zation System, Vol. II, p. 892
STOUT, W. L.
Leaching Experiments on Soil and
Mine Spoil Treated with Fluidized
Bed Combustion Waste, Vol. II,
p. 821
Characterization of Fluidized
Bed Combustion Waste Composition
and Variability as they Relate
to Disposal on Agricultural
Lands, Vol. II, p. 833
STRINGER, JOHN
High Temperature Corrosion of
Metals and Alloys in Fluidized
Bed Combustion Systems, Vol. Ill,
p. 682
STRINGFELLOW, THOMAS E.
Startup and Initial Operation
of the Rivesville 30 MWe Fluid
Bed Boiler, Vol. II, p. 203
SUMARIA, V.
Dynamic Modelling, Testing, and
Control of Fluidized Bed Systems,
Vol. Ill, p. 488
SUN, C. C.
Environmental Impact of the
Disposal of Processed and Unpro-
cessed FBC Bed Material and
Carry-over, Vol. !I, p. 846
SWIFT, W. M.
Plans and Studies on Flue Gas
Cleaning and Particulate Moni-
toring in PFBC, Vol. II, p. 493
SZEKELY, J.
Modelling Desulfurization Reac-
tions in Fluidized Bed Combustors,
Vol. II, p. 787
TARDOS, G. I.
Particulate Removal from Hot
Gases Using the Fluidized Bed
Cross-Flow Filter, Vol. Ill, p. 551
TAYLOR, D. F.
Pneumatic Solids Injector and
Start-up Burner for Battelle's
Multisolid Fluidized-Bed Com-
bustion (MS-FBC) Process, Vol. Ill,
p. 241
TEATS, F. G.
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
TERADA, HIROSHI
Particulate Removal from Pressur-
ized Hot Gas, Vol. Ill, p. 538
THOENNES, C.
Characterization of Efflux from a
Pressurized Fluidized Bed
Combustor, Vol. II, p. 655
A High-Temperature High-Pressure
Isokinetic/Isothermal Sampling
for Pressurized Fluidized Bed
Applications, Vol. Ill, p. 326
C-14
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TOUR IN, RICHARD H.
Overview of New York State's
Fluidized Bed Combustion
Program, Vol. I, p. 81
TRACEY, ROBERT
Industrial Application-
Fluidized Bed Combustion -
Georgetown University, Vol. II,
p. 61
TREXLER, EDWARD
The Department of Energy Atmos-
pheric Fluidized Bed Combustion
Utility Demonstration Program,
Vol. II, p. 160
TRINH, T.
Modelling Desulfurization Reac-
tions in Fluidized Bed Com-
bustors. Vol. II, p. 787
TSAO, KEH C.
Multiple Jet Particle Collection
in a Cyclone by Reheating Fluid-
ized Bed Combustion Products,
Vol. Ill, p. 607
TUNG, S. E.
FBC-Kodelling and Data Base,
Vol. Ill, p. 406
TUREK, DAVID
Lightoff of Multicell Fluidized
Bed Boilers by Hot Bed Transfer,
Vol. II, p. 223
TURNER, C.
The Mechanism of the Salt Addi-
tive Effect on the S02 Reac-
tivity of Limestone, Vol. II,
p. 776
ULERICH, N.
Assessment of the Impact of S02,
NOX, and Particulate Emission
Standards on Fluidized-Bed Com-
bustion System Design and Energy
Costs, Vol. II, p. 875
VAN VALKEN8URG, E. S.
Particulate Analysis Instrumenta-
tion for Advanced Combustion
Systems, Vol. Ill, p. 347
VIRR, MICHAtL
Industrial Coal Fired Fl'jidized
Bed Bo-Uers and W*ne Hea1
Boilers, Vol. II, . . 22
VOGEL, G. J.
Plans and Studies on Flue Gas
Cleaning and Particulate
Monitoring in PFBC, Vol. II,
p. 493
Development of a Process for
Regenerating Partially Sulfated
Limestone from FBC Boilers,
Vol. Ill, p. 776
Economic Feasibility of Re-
gener^ting Sulfated Limestones,
Vol. Ill, p. 851
WACHTLER, FREDERICK
Industrial Application -
Fluidized Bed Combustion -
Georgetown University, Vol. II,
p. 61
WANG, JAMES C. F.
A High-Temperature High-Pres-
sure Isokinetic/Isothermal
Sampling System for Pressurized
Fluidized Bed Applications,
Vol. Ill, p. 326
WELLS, T. G.
Conceptual Design and Preliminary
Evaluation of a 570 MM Electric
Power Generating Plant Using a
Babcock & Wilcox Company or a
Foster Wheeler Energy Corporation
Atmospheric Fluidized-Bed Boiler,
Vol. II, p. 313
WHEELOCK, T. D.
Thermodynamics of Regenerating
Sulfated Lime, Vol. Ill, p. 740
C-15
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WIED, ERWIN
Design of a Gas Turbine Plant
with a Pressurized Fluidized
Bed Coinbustor, Vol. II, p. 388
WIGTON, H. F.
Mathematical Model cf a Cross-
Flow Moving Bed Granular Filter,
Vol. Ill, P. 583
WILSON, J. S.
Atmospheric Fluidized Bed Com-
ponent Test and Integration
Facility - An Update, Vol. Ill,
P. 4
WILSON, M.
Dynamic Modelling, Testing, and
Control of Fluidized Bed Systems,
Vol. Ill, p. 438
WILSON, W. IRA
The Prediction of Limestone
Requirements for SO^ Emission
Control in Atmospheric Pressure
Fluidized-Bed Combustion,
Vol. II, p. 748
WRIGHT, S. J.
A Technical Description of the
Plant Design and Project Progress
Report, Vol. Ill, p. 310
YAMAMURA, R.
Particulate Removal From Pres-
surized Hot Gas, Vol. Ill,
p. 538
YANG, RALPH T.
Regeneration of Lime-Based
Sorbents in a Kiln with Solid
Reductants, Vol. Ill, p. 798
YOUNG, D. T.
Fluidized Combustion of Beds of
Large, Dense Particles in Re-
processing HTGR Fuel, Vol. Ill,
p. 254
YUNG, KUANG T.
Multiple Jet Particle Collection in
a Cyclone by Reheating Fluidized
Bed Combustion Products, Vol. Ill,
p. 607
YUNG, SHUI-CHOW
Granular Bed Filters for Particu-
late Removal at High Temperature
and Pressure, Vol. Ill, p. 516
ZAKKAY, VICTOR
The Effects of Finned Tubing on
Fluidized Bed Performance,
Vol. Ill, p. 156
ZELLARS, GLENN R.
Erosion/Corrosion of Turbine
Airfoil Materials in the High-
Velocity Effluent of a Pressurized
Fluidized Coal Combustor, Vol. Ill,
p. 660
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