-------
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
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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
2
n
AT 730ฐ C
_
P*1
H hi
lei ^ i
;;
;;
/'
1
'
';
;
'
.
:-'
I
\x
^
1
0
is
AT 795ฐ C
-.
_
B
il
ILI
tm
:
x
'.
'
.
i
^.
-
-
ซ>
1 "
|
|
v
'..
V ,
V
ง
\'
s
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
Whcclock and D. R
Wheclock and D. R
Hanson. G. F
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X. Boylan. Chr". Eng. Pro^r.. 6^. AIChE. New York. Nov.
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T. U. Wheelock and D.
D.C., March 1960, p. 21i.
T. D. Wheelock and i). R. Boylan.
1960, p. 590.
T. D. Wheelock and D.
1968.
T. D.
T. D.
T. D.
A. M.
National Meeting of Aa. Cher.. Soc.
W. M. Swift and'T. D. Wheelock. Ind. Eng. Chcra. . Process DCS. Dev.. lt>. Am. Chen.
Soc.. '..'ashin^ton. D.C., July 1975. p. 323.
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Aug. 27. 1969. A. Skopp. J. T. Sears, and R, R. Bertrand. Linden. S..'.. 1969.
Esso Research and Engineering Co.. "A Regenerative Limestone Process for Fluldized
Bed Coal Combustion ar.d Dcsulfurizat ion," Final Report. C. A. Mammons and
A. Skopp. Linden. N.J.. Feb. 26. 1971.
Boylan. U.S. Patent 3.037.790. April 30. 1963.
Boylan. U.S. Patent 3.260.031. July 12. 1966.
Boylan. U.S. Patent 3.607.0i5. Sept. -!1 . 1971.
Rotter. W. R. Brade. and T. D. Wheelock. Fri-print. 158th
:c-.. York. Sept. 7-12. 1969.
753
-------
II'. !.. A R'Uh. I'r.-ijriri! . f'i'ir'.h In'.<:rn:i! i'u<;t ir,n iT'.ces!,." hep-iri No. Ki'A-'.OV 7 - 77 - i 07 . k. f; iio><-. K R .
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
-------
I'll'-. '.llll/l'il I lillitl/iซ| HIM | ป .ii.il ( .iillilili'.lKHI
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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
-------
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
-------
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
-------
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
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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
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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
-------
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
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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
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ments". Chen. Er.:;. Prog. Tech. Manual, AIChE, N. Y. (1971).
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RMป. Linden. N. J., EPA-hOO/7-76-01I (1976).
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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).
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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 V,
Figure 7. Direct Operating Coit Seniitivity
Figure 8. Total Operating Colt Seraitivity
> 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
-------
:. .'': . .'.;., '.' '; '*:.'; '.'" *'. \ '"".'.*'.'..
-->.: :'. , :-,:. :..:ซ'.:.:.. <'.'.-.:. '.:...: :....-.t.:.., 1: *;.: :..-..;:.. -. . .. ..;;
:.:;:!.:.:. .:v. '.-.-r. ':-..:'. '.'.! ':,-::. ]':-. va I . .
''ii.i'.a: :..-.'. :.::.-.[ '.:.:: V- .:.' ;...:'. a!. :.:.: :.; r.<- '.','',
:.i'":t.:-j-- :ป. v '!'-.: I r.v. :-r.a ! .::!;'.:/ :'.::t, Ir.j- .' r.I 1 ::/K*'f,
; :', 1 I'.w ; r;.- ;,:i:"ir:>-'.--:-.: :ii->: v:u- !.-! 1 1.
/.ri'.r-t ' .-!
i-.-i-..,,...-:,! :,.., j.lu.: '.|.i'i-ai. iii,' !:>..', '..:'. ~ ;'av t:..;.: ~ '.': J i'-
"'ti ':: <;u:-v.; i;. i i :'>'.' tii.'it. mri.viit (.r'lc:-- i.; [... r.;::'. '.r.;. -.:!. ar.t : :i:\-ir."*.'.T, f.'Jl
Lli.":t, 'j:i;. !V.i) ':v:'.t '"t. i t-M- t::ii.:ur I ly I'.'i'.-V'-'r :jr": ai::'. Sil'." t~.; ort-'ii.'- . ''-.'! i .:(..- :i.vj
:;ulf'jr ;-r<'l'. i -n1'.- >i!i ;^.; '/rfi.-.t. . In r.o::'. i-...:i a:;;j iy:;i-:: tjy -,t.;.--i-:;, -:. r.iiruv :)'', I :;
r.ot. i ti'.- ;-i j"-i. All >!' '-r," :'! i owlnr. :J:IM |y. .ซ.-.; ;ir'.- 'i^ri" ir^js-lrn" 't ' - :.'!;.'! i' y l'-'i'.-t.',r,
no :"jl!'ui' crr-ii'.t., ari'i a .'1:1 i |I:MI:-- f>!' J.'^/tori vX':< |.t. W:M. r-.- otti'.-:'w I :T- n'.ivj.
:".,-,'; I--.-:; , ttir-j'ic.f: :- l"::.or.:;t I'ut'.- t >,. .-rr'-f:'. '>f v:(r!-it lo-n: .:' C'lj'.-.ii! co::t ,
(-;j;..-|'.-!ty .'"actlu:-, :;',rt,--!.t. ::-;i;..-, r.uM'ur crซ-l Iป. , ci.;il i-rl'.-f, .'ill 1 rr. .n tr." !;<. f-iy.MT-
;iLi ,ii uy=tt.-Ei:; ci>::t..
Fl.-.ur-i- Si r.how:: tปiซ> t.-lT-.-'.-t of capital 'itT.t IrsrrcT.-o on t!if n"t .:y^t"n cjr.t.
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
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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
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-------
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
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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
-------
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
-------
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
-------
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
-------
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
C-12
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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
C-13
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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
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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
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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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