rK,'^gency Industrial Environmental Research Laboratory EPA-600/7-76*010
top'ment Research Triangle Park, N.C. 277II
September 1976
SYMPOSIUM ON
PARTICULATE CONTROL IN
ENERGY PROCESSES
Interagency
Energy-Environment
Research and Development
Program Report
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RESEARCH REPORTING SERIES
Research reports of the Office of Research and Development, U.S.
Environmental Protection Agency, have been grouped into seven series.
These seven broad categories were established to facilitate further
development and application of environmental technology. Elimination
of traditional grouping was consciously planned to foster technology
transfer and a maximum interface in related fields. The seven series
are:
1. Environmental Health Effects Research
2. Environmental Protection Technology
3. Ecological Research
4. Environmental Monitoring
5. Socioeconomic Environmental Studies
6. Scientific and Technical Assessment Reports (STAR)
7. Interagency Energy-Environment Research and Development
This report has been assigned to the INTERAGENCY ENERGY-ENVIRONMENT
RESEARCH AND DEVELOPMENT series. Reports in this series result from
the effort funded under the 17-agency Federal Energy/Environment
Research and Development Program. These studies relate to EPA's
mission to protect the public health and welfare from adverse effects
of pollutants associated with energy systems. The goal of the Program
is to assure the rapid development of domestic energy supplies in an
environmentally—compatible manner by providing the necessary
environmental data and control technology. Investigations include
analyses of the transport of energy-related pollutants and their health
and ecological effects; assessments of, and development of, control
technologies for energy systems; and integrated assessments of a wide
range of energy-related environmental issues.
REVIEW NOTICE
This report has been reviewed by the participating Federal
Agencies, and approved for publication. Approval does not
signify that the contents necessarily reflect the views and
policies of the Government, nor does mention of trade names
or commercial products constitute endorsement or recommen-
dation for use.
This document is available to the public through the National Technical
Information Service, Springfield, Virginia 22161.
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EPA-600/7-76-010
September 1976
SYMPOSIUM
ON PARTICULATE CONTROL
IN ENERGY PROCESSES
David E. Blake, Editor
Aerotherm Divlsion/Acurex Corporation
485 Clyde Avenue
Mountain View, California 94042
Contract No. 68-02-1318, Task 22
Program Element No. EHE624
EPA Task Officer: Dennis C. Drehmel
Industrial Environmental Research Laboratory
Office of Energy, Minerals, and Industry
Research Triangle Park, NC 27711
Prepared for
U.S. ENVIRONMENTAL PROTECTION AGENCY
Office of Research and Development
Washington, DC 20460
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INTRODUCTION
Methods of control of particulate emission from conventional electric
power generation facilities are well developed and reasonably effective.
However, several recent developments in combustion technology have significantly
increased the difficulty of particulate collection. There is an immediate pro-
blem with the rapidly increasing use of low-sulfur, high-ash Western coals.
Because of the high mineral content of these coals, ash is produced at higher
rates than for high-sulfur (Eastern) anthracite or bituminous coals. Com-
pounding the problem, the low sulfur content decreases the effectiveness of
electrostatic precipitators. The most severe problems arise when attempting
to switch existing power plants (with precipitators sized for high-sulfur coal)
to low-sulfur coals.
A longer-range problem with particulate collection exists with several
high temperature and pressure energy processes now being developed. In order
to obtain maximum efficiency of power generation from coal gasifier or fluid-
ized bed combustion facilities, it is necessary to remove particulates from
hot, high pressure streams upstream of the turbines.
The "Particulate Control in Energy Processes" symposium, held in San Fran-
cisco on May 11-13, 1976, devoted six sessions, two panel discussions, and some
22 papers to consideration of these problems and their solutions. This volume
is a collection of the papers presented at the symposium.
The symposium was sponsored by the Particulate Technology Branch of the
U.S. Environmental Protection Agency (Research Triangle Park, North Carolina),
and by the Electric Power Research Institute (Palo Alto, California).
The symposium was hosted and the proceedings produced by the Aerotherm
Division of Acurex Corporation (Mountain View, California).
I wish to express my appreciation for the help and support of the spon-
soring organizations, and most particularly to Dr. Dennis C. Drehmel of the
EPA and Mr. Walter Piulle of EPRI for the many hours they spent working with
me to assure the symposium's success.
David E. Blake
Symposium General Chairman
Aerotherm Division
Acurex Corporation
iii
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TABLE OF CONTENTS
Keynote Address; LOW SULFUR COAL AND THE PARTICULATES PROBLEM
(Talley, Laska, Abbott)
Page
Section 1 — Use of Low Sulfur Coals
LOCATION OF LOW SULFUR COALS AND WESTERN POWER
GENERATION NEEDS (Schmidt) 13
PROBLEMS AND CONTROL OPTIONS USING LOW SULFUR COAL
IN UTILITY BOILERS (Green) 53
WESTERN COAL USE IN INDUSTRIAL BOILERS (Maloney). . 75
Section 2 — Electrostatic Precipitators and Low Sulfur Coals
FACTORS AFFECTING ESP PERFORMANCE ON WESTERN COALS
AND EXPERIENCE WITH NORTH DAKOTA LIGNITES (Selle,
Hess) 105
ELECTROSTATIC PRECIPITATOR OPTIONS FOR COLLECTION
OF HIGH RESISTIVITY FLY ASH (Sparks) 127
ELECTROSTATIC PRECIPITATOR OPERATING EXPERIENCE
WITH LOW SULFUR COAL - JIM BRIDGER POWER PROJECT
(Ashton) 143
OPERATING EXPERIENCE WITH ESP CONDITIONING IN RELA-
TION TO AN ELECTROSTATIC PRECIPITATOR UPGRADING
PROGRAM (Cragle) 171
AUSTRALIAN EXPERIENCE WITH FLUE GAS CONDITIONING
(Watson) 189
Section 3 — Recent Technology Developments in the Use of Elec-
trostatic Precipitators
MATHEMATICAL MODELLING OF FINE PARTICLE COLLECTION
BY ELECTROSTATIC PRECIPITATION (Gooch, McDonald). . 217
IMPROVED PRECIPITATOR TECHNOLOGY BY PILOT PLANT
TESTING AND EVALUATION OF COAL CORE-CORES
(Tassicker, Sproull) 267
TECHNIQUES FOR CONDITIONING FLY ASH (Dismukes) . . 297
RAPPING REENTRAINMENT STUDIES (Spencer) 317
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TABLE OF CONTENTS (Concluded)
Section 4 — Alternatives to ESP's for Low Sulfur Coal
FABRIC FILTRATION PERFORMANCE MODEL (Cooper, Hampl). . 343
PERFORMANCE AND ENGINEERING EVALUATION OF THE NUCLA
BAGHOUSE (Ensor, Hooper, Scheck, Carr) 377
OPERATING EXPERIENCE AND PERFORMANCE AT THE SUNDBURY
BAGHOUSE (Spagnola, Turner) 401
PARTICIPATE REMOVAL CHARACTERISTICS OF THE 170 MW
MOHAVE SCRUBBERS (Jones, Weir, Johnson, Martin,
Calvert) 429
EVALUATION OF A PARTICULATE SCRUBBER ON A FULL SCALE
UTILITY BOILER (Sparks, Ensor, Calvert, Wallon) . . . 467
Section 5 — High Temperature and Pressure Energy Processes
COLLECTION MECHANISMS AT HIGH TEMPERATURE AND PRESSURE
(Calvert, Parker) 491
MEASUREMENT METHODS AT HIGH TEMPERATURE AND PRESSURE
(Blann) 523
FBC PARTICULATE CONTROL PRACTICE AND FUTURE NEEDS:
EXXON MINIPLANT (Hoke) 551
PARTICULATE CONTROL REQUIREMENTS FOR GASIFIED COAL-
COMBINED-CYCLE POWER PLANTS (Robson, Blecher) .... 563
vi
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KEYNOTE ADDRESS
LOW SULFUR COAL AND THE PARTICULATES PROBLEM
Wilson K. Talley
Assistant Administrator
Office of Research and Development
Environmental Protection Agency
Richard Laska
Special Assistant to the Deputy Assistant Administrator
Office of Energy, Minerals and Industry
Environmental Protection Agency
James H. Abbott
Chief, Particulate Technology Branch
Utilities and Industrial Power Division
Industrial Enviornmental Research Laboratory - RTP
Environmental Protection Agency
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LOW SULFUR COAL AND THE PARTICULATES PROBLEM
All of the participants in this symposium are involved, in some manner, in a
classic environmental problem. In essence, the problem we are addressing is that,
in controlling one form of pollution (in this case sulfur oxides emissions), we have
found increased difficulties in controlling another form of pollution (particulates).
This is a classic environmental problem in that it brings to mind all of the buzz-
words which have embellished environmental debate in the past. The problem
involves trade-off analysis. It is based on health effects studies and on emission
standards. And, above all, it requires a comprehensive, holistic and integrated
solution with appropriate control technology development. As I said, it1 s a classic
environmental problem.
DDT BROUGHT THE HOUSE DOWN
In its complexity, this problem reminds me somewhat of another classic
environmental problem which came to light several years ago on the other side of
the globe. The story, which is said to have taken place in Borneo, involved a
government effort to rid isolated native tribes and theirx thatched houses of lice and
malaria. To do so, the government sent teams to the villages to spray the huts with
DDT. The effort was a stunning success. For the first time in history, the natives
stopped itching and malaria was curbed.
In a short time, however, an unexpected problem developed. It seems that the
lizards that shared the thatched huts with the natives were eating the poisoned
insects and concentrating the DDT. And the village cats were eating the sick
lizards and then expiring in droves. It didn't take long before the jungle rats,
realizing their opportunity, invaded the villages and terrorized the natives. In
response, the government' s Royal Air Force airlifted hundreds of cats into the
villages to drive back the rats. Unfortunately, that wasn't the last of the problems.
It seems that, without their natural lizard predators, the worms that dwelt in the
thatched roofs of the huts proliferated and literally ate the natives out of house and
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home. Entire villages collapsed. Thus the natives were left without homes to call
their own and surrounded by a bunch of strange cats. I1 m sure you all know the
feeling.
The moral of my story is that, when environmental problems, such as we are
addressing in these proceedings, arise, our solutions must be as far-sighted as
possible. Rather than just resolving today' s crisis, we should try as hard as
possible to develop the kind of solutions and control technologies which will also
hold the greatest potential for solving tomorrow1 s problems as well.
The particular problem that this symposium is addressing is the need for ways
in which to utilize low-sulfur coal in an environmentally compatible manner. Some
of the more far-sighted discussions of solutions to this problem are presented in the
papers on high-temperature, high-pressure particulate control technology — a
capability which is essential to the successful development of promising second-
generation energy technologies.
EPA' S PARTICULATE CONTROL R&D EFFORTS
The remainder of this paper will be devoted to placing in perspective EPA1 s
research and development role in particulate control, and to presenting some of our
significant problems and identifying the major programs which are attempting to
find solutions to these problems.
As the nation strives to reduce dependence on foreign energy sources, increased
consideration is being given to the energy-related environmental problems arising
from domestic energy resource development. Accelerated efforts to expand existing
energy supplies and to develop new energy sources are accompanied by accelerated
efforts to anticipate, plan for, and control the more serious adverse health and
ecological effects resulting from such energy development.
Since its establishment in 1970, the Environmental Protection Agency has been
heavily involved in energy-related environmental research efforts, including the
development of pollution control technology, supportive of the EPA' s statutory
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responsibilities. Since Fiscal Year 1975, the EPA has also had the added responsi-
bility of planning and coordinating the entire 17-Agency Federal Energy/Environment
Research and Development Program. Through the Interagency Program to date, the
EPA has planned and implemented more than a quarter of a billion dollars of energy-
related environmental research and control technology development. Responsible
for this centralized coordination of the Interagency Program is the Office of Energy,
Minerals and Industry, within EPA1 s Office of Research and Development. This
organization pursues two basic goals — to provide a focus for EPA' s own Environ-
ment/Energy and Industry-related R&D efforts, and to serve as the coordinator of
the entire Federal Interagency Program.
EPA1 S GOAL — CLEAN USE OF COAL
EPA1 s mission in this area recognizes that timing is critical in the overall
relationship between energy development and environmental protection. Efforts to
expand use of domestic supplies will focus on the nation' s two primary fuel
resources — coal and nuclear energy. Through the 1970' s and early I9601 s, coal
use in particular will be increased. This will occur chiefly through direct combus-
tion in power plant and industrial boilers, but only if technologies to control
emissions of particulates and other hazardous pollutants are developed and applied.
Thus, Environment/Energy R&D on near-term coal use centers on our continuing
program to advance environmentally acceptable ways to extract and utilize coal.
To the degree that these control technologies can be implemented rapidly, there will
be early environmental, economic, and social benefits from such R&D efforts. For
this reason, funds have been weighted heavily in the FY 1975 and FY 1976 budgets
towards facilitating near-term coal use.
The Clean Air Act is the basis for our research into and development of flue
gas cleaning technologies. This R&D is focused on identifying a series of standards
which can be implemented within the constraints of sound engineering practices and
economic viability.
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We are aggressively pursuing research on particulate control because of the
potential widespread impacts of accelerated coal resource development and usage.
Particulate control needs and required technologies are, to a large extent, dependent
upon the type of coal to be used, the combustion process involved and the location
of such usage. The bulk of coal usage is in the Eastern States, but the bulk of low
sulfur coal is in the Western States. See Charts One and Two for the distribution
of domestic reserves now available. There is almost equal distribution of coal
between Eastern and Western states, but there is five times as much low-sulfur coal
in the West. There is also a greater proportion of strip-mineable coal in the West.
For Eastern users the coal may be obtained by deep mining low-sulfur Eastern coals,
by desulfurizing strip-mineable, high-sulfur coal, or by shipping Western coals to
the East. These coals, whether of high or low sulfur content, will be available for
utility and industrial use for many decades.
THE SULFUR/PARTICULATE TRADEOFF
Typical of the problems facing the EPA and industry is the effect on electro-
static precipitator operations of conversion to low-sulfur coal usage. The electro-
static precipitator (ESP) is the most common device for controlling the particulate
emissions from large coal-fired utility boilers. Those familiar with the power
industry know that in an electrostatic precipitator, dust suspended in the gas stream
is electrically charged and passed through an electric field which causes the
particles to migrate toward a collection electrode.
The particle collection efficiency of an ESP is, in part, a function of the
resistivity and particle size distribution of dust being collected.
The fact that fly ash resistivity, and thus precipitator performance, are related
to coal sulfur content has long been recognized. Correlations have been developed
between ESP performance and coal sulfur content. It is clear that a given electro-
static precipitator will have higher collection efficiencies when collecting fly ash
from the combustion of high-sulfur coal than when collecting fly ash from the
combustion of low-sulfur coal.
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CHARTS ONE & TWO: COAL RESERVES IN THE UNITED STATES
C/5
UJ
-B
H W
DC _,
** or
o £
C/3
I^J
UJ
C/5
UJ
H
cvi (-
|_ C/J
52
$£
O H
C/)
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^
RESERVES
%
BILLION-TONS
RESERVES
< 1% SULFUR
1.1 TO 3.0%
>3%
UNKNOWN
RESERVES
%
BILLION-TONS
RESERVES
< 1% SULFUR
1.1 TO 3.0%
> 3%
UNKNOWN
DEEP-MINABLE
83
169
33 BILLION TONS
33 "
55 "
82 "
56
131
167 BILLION TONS
38
11 "
18
STRIP-MINABLE TOTALS
17 100
33 202
16
27
40
16
44 100
103 234
71
16.0
5
8
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The switch to a low-sulfur coal, as necessitated by efforts to achieve or
maintain sulfur oxides emissions standards, can increase particulate emissions by
more than a factor of 10. As can be seen from Chart Three, such an increase in
particulate emissions on a large scale basis would result in unacceptable air quality
impacts. Major programs to identify and evaluate alternatives to overcome this
problem, or to establish the viability of other control strategies, are being imple-
mented by our research program (see Chart Four).
In parallel with our efforts to develop the necessary control technologies, an
Interagency Program effort is being pursued to identify health effects, ecological
fate and effects, and atmospheric effects of particulates. In addition to the EPA,
major Interagency Program efforts in this area involve NIOSH, ERDA and NIEHS.
Through our role in coordinating such interagency efforts, we are becoming increas-
ingly aware of the need for continuing research into the effects of particulates as
well as into the development of control technology options.
FINE PARTICULATES ~ A PERVASIVE PROBLEM
Specific program focus is on the range of smaller particulates (see Chart Five).
Fine particulates appear to pose a far greater health hazard than do the more easily
controlled large particulates because, in contrast to coarse particles, they can
bypass the body' s respiratory filters and penetrate deeply into the lungs. Fine
particles released into the atmosphere remain airborne for extended periods of time,
obstruct light, cause the limited visibility typical of air pollution, haze and smog
and are added to the planetary particulate loading which may, according to recent
studies, have a significant climatic cooling effect.
Such fine particles have also been identified as transport vehicles for gaseous
pollutants . The health hazards of fine particulates are intensified by the tendency
of emissions from pyrometallurgical and combustion processes to condense as
chemically active fine particles. Many toxic and potentially hazardous compounds
are also emitted as fine particulates. Such particulate matter, formed in the
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CHART THREE
PARTICULATE EMISSION ESTIMATES
EMISSIONS (1000 TONS PER YEAR)
1970 1975
SOURCE CATEGORY
o COAL-FIRED STEAM ELECTRIC
POWER PLANTS
o INTEGRATED IRON AND STEEL
MILLS AND COKE PLANTS
o PORTLAND CEMENT PLANTS
o OTHER
o ALL SOURCES
POTENTIAL
34533
15259
10643
29017
89452
ACTUAL
4188
2800
906
19606
27500
POTENTIAL
47530
18532
12291
34824
113177
FULL COMPLIANCE
914
521
69
10896
12400
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CHART FOUR
FINE PARTICULATE PROGRAM
o MEASUREMENT TECHNIQUE DEVELOPMENT
GOAL IS TO PRODUCE DEVICE WHICH IS CONTINUOUS AND ACCURATELY MEASURES
FRACTIONAL EFFICIENCY
o ESP PROGRAM
CHARACTERIZATION FOR SEVEN SOURCES
MATH MODEL IS DEVELOPED
CAPABLE OF HIGH EFFICIENCY. IF NO RESISTIVITY PROBLEM
CONDITIONING BEING EVALUATED
o SCRUBBERS
EIGHT DEVICES TESTED ON A VARIETY OF SOURCES
PERFORMANCE DEPENDENT ON PRESSURE DROP
TCA SCRUBBER UNCHARACTERISTICALLY EFFICIENT
FORCE/CONDENSATION SCRUBBERS UNDER DEVELOPMENT
o FABRIC FILTERS
TESTING FOR THREE SOURCES COMPLETED
EFFICIENT DOWN TO 0.3/u
PROGRAM AIMED AT INCREASING APPLICABILITY AND OPERABILITY
o NEW IDEAS/NOVEL DEVICES
40 NEW CONCEPTS EVALUATED; 9 APPEAR PROMISING
30 NEW DEVICES EVALUATED; 10 APPEAR PROMISING
o HIGH TEMP./HIGH PRESSURE CT
FBC AND LOW-BTU GASIFICATION NEED CT
PROGRAM AIMED AT DEFINING REQUIREMENTS AND EVALUATING CONCEPTS AND DEVICES
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CHART FIVE
FINE PARTICULATE CONTROL
PROGRAM-BACKGROUND
-FINE PARTICULATES ARE HEAL TH HAZARDS BECAUSE:
o AIRBORNE FOR EXTENDED TIME PERIODS
o PENETRATE DEEPLY IN LUNG
o ACT AS TRANSPORT AGENTS FOR OTHER POLLUTANTS
-EPA R,D & D PROGRAM INCLUDES
o MEASUREMENT TECHNIQUE DEVELOPMENT
o IMPROVEMENT AND CHARACTERIZATION OF PRESENT TECHNOLOGY
ESPs
SCRUBBERS
FABRIC FILTERS
o NEW IDEAS/NOVEL DEVICES
o HIGH TEMP./ HIGH PRESSURE CT
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atmosphere from chemical reaction and condensation, is called secondary participates
Such secondary particulates raise added problems and issues. The phenomena
associated with the formation and transport of secondary particulates makes it
difficult to relate atmospheric particulate pollution levels to specific sources. This
difficulty has hampered the development of effective fine particulate control strate-
gies and the establishment of meaningful fine particulate emission standards. The
control of these secondary forms of particulates must be through control of their
precursors, and it is thought that primary emitted particulates play an important role
in the formation cycle.
Hence, the information which is being developed through our health and ecologi-
cal effects research efforts continues to add support to the need for a control tech-
nology capable of controlling fine particulates while, at the same time, allowing for
effective control of sulfur emissions as well. The more we learn about the pathways
and hazards of particulate pollution, the more our need for an effective set of control
technologies is manifest. Like the Borneo natives, we require control technologies
which are relevant to, and compatible with, the entire range of environmental condi-
tions that we have identified.
CONCLUSION
In conclusion, there are several points which I would like to emphasize:
First, in the near term there will be no energy panaceas. Our new sources of
energy will look a lot like our old sources of energy and will employ technologies
with which we are already familiar.
Second, the protection of public health and welfare requires that the emission
of both sulfur oxides and of particulate matter be controlled at the source.
And finally, in those instances where low-sulfur coal is to be used as an energy
source in order to control sulfur emissions, there is a technological conflict between
the requirements imposed by sulfur oxide standards and the ability of the traditional
particulate control devices to meet particulate control needs.
11
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With these points in mind it is obvious that particulate control in energy
processes is a topic deserving of immediate and intensive study. Solutions to the
existing problems, especially in regard to low-sulfur coal usage, require concentrated
effort not only on the part of governmental agencies but also in cooperation with
utilities and other consumers of the fuel. It is fortunate that the natural and under-
standable reluctance of these industries to give rapid trial and endorsement to
changing technology has begun to abate. Such tempering of attitude is beneficial and
will assuredly lead to quicker and probably better solutions to the present problems.
We have a common problem. We should work together to develop a common solution.
12
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LOCATION OF LOW SULFUR COALS AND
WESTERN POWER GENERATION NEEDS
Richard A. Schmidt
Electric Power Research Institute
Palo Alto, California
13
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INTRODUCTION
Assessments of the coal supply in the ground have been published by the
dozens in the last years. These assessments come in all shapes and
varieties, some reporting dire shortages within a few years to no likely
shortage within a few million years. When one analyzes these forecasts,
one can come to the conclusion that each may be correct. The variations
depend entirely on the different assessments of the coal availability
measured with geologic, technological, economic and political parameters.
Within the last few years, since the enactment of the Clean Air Act, a
new parameter has been added, the quantitative amount of qualitative
sulfur content of the unmined or in-situ coal. The first set of para-
meters are difficult enough to comprehend and make judgment upon to
arrive at a conclusion. The last parameter concerning the sulfur con-
tent of coal upset the whole trend of reasoning.
To most of us, a resource is a thing or action we can turn to in time of
need or emergency. To a geologist the word indicates a.concentration of
raw inplace coal in such form that economic extraction is currently or
potentially feasible. But the common man's concept of resource, when it
comes to coal, includes only "reserves." Reserves are assumed to be
ready for extraction and consumption, and can be drawn upon to take care
of a need or an emergency. Undiscovered seams of coal included in
resource are not available for use, they are only presumed to be present
in the earth's surface.
Coal Resources and Reserves; Quantity
The most serious problem in determining both coal resources and recover-
able reserves is that much of the basic technical field work in the
public domain was performed many years ago. Many estimates are founded
on work done in the early part of this century without the advantage of
modern techniques and instrumentation and thus are often lacking the
degree of refinement and specificity necessary in assessment of potentials
14
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for the large scale, costly developments of the present. An intensive
effort in this regard is to be carried out by the present programs of
the U. S. Geological Survey and the U. S. Bureau of Mines. The Geologi-
cal Survey is primarily responsible for determination of the nation's
resources, while the Bureau of Mines provides estimates of recoverable
reserves. The data on coal resources and reserves resulting from the
work of the Geological Survey and the Bureau of Mines represent the only
authoritative accounting of the nation's coal deposits, and these data
will be used to describe the situation apparent at present.
The United States Geological Survey estimates the total remaining coal
resource in the United States with certain specifications, namely beds
of bituminous coal and anthracite 14 inches or more thick and beds of
subbituminous coal and lignite 2-1/2 feet or more thick with over burden
depths of 3,000 and 6,000 feet.
The resulting estimate of the amount of coal resources — 3,968 X 10
(billion) tons — is a theoretical number, some of this coal is mineable
under economic conditions, but most of the coal cannot be extracted with
present mining techniques.
The total remaining coal resources of the United States to a depth of
3,000 feet as determined by mapping and exploration is about 1,580
billion tons. The total comprises, in rough terms, the following:
686 billion tons bituminous coal
424 billion tons subbituminous coal
. 450 billion tons lignite
Clearly this is a substantial amount. However, as will be shown in this
section, not all the coal resource can be recovered, thereby reducing
the quantity of usable reserves. Additionally, as shown in the next
section, the different heating values and quality of coals further
reduce the magnitude of the available coal reserves.
15
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In order to be considered targets for development, coal seams must be of
relatively high quality (ideally, at a minimum, high heating value and
low sulfur content) occurring in relatively thick, uniform, and exten-
sive beds that are located near enough to the earth's surface to permit
application of economic mining technology. Thickness of coal seams is a
most important parameter, because it controls the lateral extent of
operations and strongly influences operating costs. Figure 1 shows the
relationship of coal seams thickness to area required to account for 1
million tons of reserves. It may be seen that a 4 foot coal seam (common
in certain eastern coal fields) requires 140 acres for 1 million tons,
whereas a 12 foot seam (common in certain western coal fields) requires
only 50 acres. Clearly, other things such as coal character and quality
being equal, it would be preferable to work in the more concentrated
deposit. This is not always possible, however, because of consumer
requirements for coal properties and quality.
As a part of planning for Project Independence, a new evaluation of the
nation's coal deposits were made b'y the IJ.S. Bureau of Mines and Geologi-
cal Survey. A new term was coined; the "reserve base", defined as "the
quantity of in-place coals calculated under specified depth and thickness
criteria." This new term is intended to be intermediate between resources
(the total stock of coal in the ground) and reserves (the quantity of
coal actually recovered).
The coal reserve base uses more realistic criteria for estimating. It
includes only coal reserves in the measured and indicated resource
category; in beds 28 inches or more thick for bituminous and anthracite
and 60 inches or more for subbituminous and lignite; in the 0 to 120
feet overburden category for lignite, which is considered suitable only
for strip mining, and in the 0 to 1,000 foot overburden category, which
are deemed to be suitable for strip, auger and underground methods.
The "demonstrated coal reserve base", includes the measured and indi-
cated reserve categories, as in the coal reserve base, but based on a
high degree of identification and engineering evaluation with emphasis
16
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320
280
240
"5 200
s
5 160
5
l_
CO
< 120
80
40
1800 Tons/Acre-Feet
100 Percent of Coal in Place
8 10 12 14
Seam Thickness, (feet)
16
18
20
22
Figure 1 Relationship of Seam Thickness to Area per Million Tons
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given to a high degree of certainty that the coal is technically and
economically mineable at some time. No consideration was given the
marketability or competitiveness of one seam of coal against another.
In effect it is assumed that all of the coal can eventually be mined if
the mining technology is perfected to extract it and there is a customer
demand great enough that it is willing to pay the price.
The "reserve base" data represent the most recent estimate of the quantity
of U.S. coal deposits which may experience development in the future.
Table 1 presents "reserve base" data together with the authors' estimate
of recoverable reserves for both underground and surface mining.
Tables 2 through 5 show, for each of the four principal coal ranks, the
following information by state:
(a) total estimated identified resources
(b) indicated reserve base, by method of mining
(c) recoverable reserves, by method of mining, and tota.1
(d) recoverable reserves as percent of identified resources
The tables are, with one exception, compiled from available information.
The exception is that an estimate of recoverable reserves by underground
mining was prepared for each rank of coal. Note that the reserve base
represents 100% of the coal in-placej no allowance for incomplete recovery
has been made. Actually, the amount of coal that can be recovered from
a given deposit varies from about 25% to 90% of the coal in-place.
Therefore, in order to arrive at a reasonable estimate of the amount of
recoverable coal, the "reserve base" data must be reduced by an appropri-
ate amount to allow for incomplete recovery.
All of the schedules of coal resources and reserves presented so far are
estimates of the coal in the ground and not necessarily mineable under
unspecified circumstances.
In general terms, the Bureau of Mines uses a round recoverability factor
of 50%. The recoverability factor is defined as the percentage in the
18
-------
reserve base that can be recovered by established mining practices. The
50% recovery factor was established from past mining records for under-
ground mines in the eastern states. In the East, most of the mining to
date has been accomplished in areas where multiple seams of coal are
present. For economical reasons, it can be conceded that the prime seam
of coal was or is being mined first. Ideally, this procedure leaves the
seams above and below undisturbed. The Bureau of Mines counts undisturbed
seams as mineable, which may not be necessarily true. The seams above a
mined out seam, if the interval between the two is not sufficiently
thick, may be fractured and subsided to such an extent that the seam is
not mineable under any conditions. A similar situation may develop with
coal seams below the mined out seam. Water may seep through fractures
in the interval cover material to such an extent that the roof is not
supportable, again the seam is lost for mining, but still considered as
a coal reserve.
Seams of coal under populated areas, Federal and state owned forests,
parks, reservations, navigable rivers and streams, etc., are not legally
mineable areas, but these areas are considered in the coal reserve
count. Neither is the land around oil and gas wells mineable, large
blocks of coal have to be left standing to prevent the hazard of oil and
gas seepage.
The various geological surveys that have worked many years in compiling
geological data because of limited funds they have had to use limited
data, coal seam outcrop information, water well drilling records and any
other available data to make their judgement on the continuity of a seam
of coal. Usually, this type of available data is very sparse, and has
to be extrapolated for very large areas. Such data, many times, does
not show the thinning out of a seam of coal, seam splitting, geological
faults and voids and many other conditions, but the factors are not
considered when appraising an area.
Very large tracts of land were assembled in the past for a mining opera-
tion which may have been successful in mining all of the available
19
-------
recoverable prime seam coal and leaving the remaining seams in place.
In many cases, these large tracts were later subdivided and sold off as
farms, building lots, etc., making the reassembling of these lands into
a mineable tract impossible; again this coal is lost as a mineable
reserve. In the West, where very thick seams of coal are found, Federal
and state laws restricted the mining of coal which may disturb dry river
beds, alluvial plains, ceremonial grounds, etc., which eliminates many
more tons of coal reserves. The following analysis represents an
initial attempt to take these factors into account in arriving at an
estimated actual recoverable coal reserves.
Recoverable reserves for underground mining were estimated by applying a
recovery factor to the reserve base. Although the recovery of coal
reserves in a_ given deep mine may be about 50%, there are substantial
areas between mines, under cities, in pockets too small for economic
recovery, too badly faulted, etc., where mining is either impossible or
otherwise precluded. The effect of these and other conditions is to
lower the recoverable reserves. To allow for the influence of the above
factors, it is assumed in this analysis that coal recovery by underground
mining from the "reserve base" will be, at a maximum, only about 30% of
the total coal in place. Clearly, this recoverability factor is uncertain;
it could go as low as about 20% or as high as about 35% of the total
coal in place. It is considered unlikely, however, as suggested by the
U.S. Bureau of Mines, that recoverability from the "reserve base" would
be as high as 50%. Table 1 shows that the total estimated recoverable
coal through underground mining is nearly 85 billion tons, with about
two-thirds of the total comprised by eastern and midwestern bituminous
coal, about one-third subbituminous coal, and minor amounts of anthracite.
This estimate, recognized to be conservative, could be increased slightly
by applying a different recoverability factor. In view of the several
uncertainties regarding estimating coal reserves, it is believed prudent
to be conservative at the present so that developments of finite resources
of coal can be carefully planned and evaluated in advance.
For surface minable reserves, the estimates of the U.S. Bureau of Mines2
have been adopted (Table 1). These data show that the estimated recoverable
20
-------
coal through surface mining is roughly 45 billion tons. More than half
of the total is subbituminous coal, with nearly one-third bituminous
coal and the remainder lignite.
The recovery of coal from reserves through surface mining may also be
less than the commonly quoted estimates of 85% or 90% or more.3 While
these magnitudes may apply to a given mine, they may not apply to an
entire coal field. Data in Table 2 through 5 permit an assessment of
the apparent resource recovery percentage in surface mining. Assuming
that the strippable reserve base includes all coals that may be ex-
tracted through surface mining, and assuming further that the separately
estimated recoverable strippable reserves represent the portion of the
reserve base that can be actually produced, then it is possible to
calculate an indicated recovery percentage for surface mining for each
4
rank of coal, by state. Bearing in mind the shortcomings of the approach
used, the data have been employed to calculate an "indicated recovery
percentage" for strippable reserves of different rank coals, by state
(Table 6). The table shows that there are few states where the indicated
recovery percentage is even close to the commonly stated "norm" of 85%
to 90% or greater recovery. In most major coal-producing states, only
about one-quarter of the reserve base is estimated to be recoverable
(West Virginia, Illinois, Kentucky). Stated another way, if the reserve
base figure had been employed together with the "normal" 85% to 90%
recovery factor to derive recoverable strippable reserves, values would
have resulted which would be much larger than those determined by Bureau
of Mines measurements. The actual recoverable strippable reserves,
which largely are only about 45% of the reserve base, and will clearly
require careful management and conservation if they are to satisfy
projected demands.
i
The foregoing suggests that the coal reserves recoverable from the
reserve base are:
Commonly Estimated Indicated Recovery
Recoverable Reserves, Percentage, Percent
Mining Method Percent of Reserve Base of Reserve Base
Underground 50 30
Surface 90 45
21
-------
In addition, it will be essential to extend the reserve base by more
complete geological knowledge and through the perfection of new and
improved technology.
The total estimated recoverable coal resources as indicated by the above
data is 130 billion tons, or about 8% of the total U.S. coal resources.
Clearly, the quantity of recoverable reserves is still substantial.
However, as will be discussed next, the coal quality factors limit the
potential use of recoverable coal reserves.
Coal Resources and Reserves; Quality
To an electric utility coal consumer, the heating value of coal (ex-
pressed as the number of Btus per pound or per ton) is the important
parameter, not merely tonnage alone. In producing a given amount of
heat for production of electricity, a utility would have to use signi-
ficantly larger amounts of low Btu coal than of high Btu coal. Unfor-
tunately, in consuming the additional tonnage of low Btu coal to make up
the Btu difference, the sulfur content of the additional tonnage is also
emitted, and consequently the effective sulfur content of the coals is
increased. Employing a standardized heating value for coals used by
electric utilities, Rieber^ found that "conventional estimates of both
known resources and recoverable reserves of low sulfur coal are grossly
overstated." In particular, Rieber's analysis shows a small increase in
the estimates of bituminous coal resources/reserves, and to a large
reduction in the estimates of subbituminous coal and lignite. A signi-
ficant portion of U.S. coal resources and reserves which are normally
considered to be low in sulfur content are reclassified to higher sulfur
categories. For example, known recoverable reserves in the lowest
sulfur category (<_ 0.7 wt. percent sulfur) are reduced from a conven-
tional estimate of 68.2 billion tons to 16.4 billion tons on a con-
sistent Btu-sulfur adjusted basis (Table 7). Should coal demand be such
as to require a doubling over present levels by 1985, the cumulative
production would be on the order of 20 billion tons. With conventional
reserve estimates there would be close to 50 billion tons of coal re-
maining, even after such expanded production. However, under the revised,
22
-------
standardized reserve estimates, known recoverable reserves of lowest
sulfur coal would be completely exhausted before 1985, and there would
even be a deficit!
The reduction in conventional reserve estimates is concentrated in the
western states (Table 7), where nearly 85% of these reserves are shifted
to higher sulfur categories through the standarization procedures. The
reserves most affected are the lower heating value coals (subbituminous
and lignite) which are present in the largest tonnages.
These data are of exceptional significance for long-range policy planning.
There seems to be no practical alternative to use of bituminous coals of
high heating value and high sulfur content, simply because of the scarcity
of low sulfur coal of equivalent heating value. Revision of air quality
standards would be the most immediate way to achieve continued use of
certain high sulfur coals and expanded use of others, with or without
flue gas desulfurization devices. Alternatively, high sulfur coals
could be gasified or liquefied to produce "clean fuels;" it remains to
be seen whether such substitute fuels would be restricted to markets
other than electric utilities because of the inherent properties of the
converted fuels and shortages of natural fluid hydrocarbons.
Table 8 through 11 presents data on recoverable reserves of the several
coal ranks by state. The average heating value and sulfur contents as
reported by the U.S. Bureau of Mines were used to calculate the sulfur
contents per million Btus. The sulfur/heating value data were compared
with recoverable reserve data to arrive at an assessment of potential
recoverable reserves. Only those states having relatively large recover-
able reserves of low sulfur coal per unit of heating value were judged
to have significant development potential. These estimates are upper
limits, with the actual production realized being determined by indivi-
dual coal seams and probably less than these amounts.
Table 12 summarizes the potential recoverable reserves for each rank by
sulfur content per million Btu. Roughly half of the total recoverable
reserves are bituminous coal, and roughly two-thirds of this has more
23
-------
than 1.5 Ibs S/MMBtu. The next largest amount is subbituminous coal,
all of which is less than 1.5 Ibs S/MMBtu. The relatively smaller
recoverable reserves of anthracite and lignite are all less than 1.5 Ibs
S/MMBtu.
While about one-third of the total potential recoverable reserves have
less than 0.7 Ibs S/MMBtu, nearly three-quarters of this lowest sulfur
coal occurs in the western states. The relatively remote location of
these reserves and the uncertainty about their development resulting
from pending litigation could delay the pace of their development.
About one-quarter of the total potential recoverable reserves are in the
next lower sulfur category; again roughly three-quarters of these re-
serves occur in the West and their early development is not at all
assured. The remaining reserves, while occurring in the East, are
greater than 1.5 Ibs S/MMBtu. Although many of these reserves have been
developed, it is not at all apparent that the coal could be used without
treatment of some sort in electric utility plants- under present environ-
mental control requirements.
Tables 13 and 14 present a comparison of assessed recoverable reserves
against projected production. The data show an optimistic estimate of
future coal production to 1985 (assuming that new production is in
addition to present production which is depleted by about one-quarter
because of mine retirement or exhaustion). Only about one-quarter of
the 1985 production would require at least some treatment prior to use.
The remaining half of total 1985 production is quite high in sulfur
content and would require rather extensive treatment in order to permit
use of these coals in power generation.
The estimated 1985 production of bituminous and subbituminous coal is
distributed among broad geographic regions as follows:
Estimated 1985 Production (million tons)
(Ibs S/MMBtu)
Region <0.7 0.7 - 1.5 > 1.5 Total
East
West
Total 249.0 230.5 497.3 976.8
24
-------
The data indicate that two-thirds of total 1985 production is expected
to come from east of the Mississippi River; about three-quarters of this
eastern coal has greater than 1.5 Ib S/MMBtu and some sort of treatment
will be required to achieve environmental quality standards at coal-
fired power plants.
The remaining third of total 1985 production is estimated to come from
the West. About two-thirds of this western coal has less than 0.7 Ibs
S/MMBtu and could be used directly in power plants. Note, however, that
the development outlook for this lowest sulfur coal is clouded by the
litigation in the Sierra Club_ vs_ Morton (Kleppe) suit, and there is real
doubt whether the amounts estimated here' can be realized in the next
decade. Should there be shortfalls in production from western deposits,
then it seems likely that expanded eastern production (and treatment of
the higher sulfur coal involved) would be necessary to meet electric
utility industry requirements.
Coal Resources and Reserves: Utilization
Up to now, the discussion has been primarily on the quantity and quality
of coal deposits in the ground. Certain bituminous coals have chemical
and physical properties which make them valuable for special purposes,
especially for manufacture of coke for reduction of iron ore to pig
iron, chemical uses, and export. These coals are considered too valu-
able to be used for production of process steam or electric power gen-
eration.
Table 15 shows data on resources, reserve base, and recoverable reserves
of bituminous coals in the United States with respect to coking proper-
ties. "Premium-and marginal-grade" coking coals^ are distinguished from
total bituminous coals in an attempt to arrive at an estimate of the
magnitude of noncoking coals that would be of principal interest for
steam-raising applications. The table shows that only about 11 billion
tons of estimated recoverable reserves (16% of the total) are coking
coals, and about 56 billion tons are estimated "steam" coals. Only the
latter would appear to be available to meet electric utility industry
25
-------
requirements; because of the inability to substitute for coking coals in
much of the steel production industry, it appears that most coking coals
will have to be dedicated to metallurgical use. Because coking coals
are low in sulfur content, the remaining coals for steam-raising must
necessarily be higher in sulfur content. If these higher sulfur coals
are to be used for power generation in compliance with air quality
control standards, some sort of coal treatment, emission controls, or
fuel/load switching (or combination of all approaches) would appear to
be required. However, the fact that positive emission control technology
remains to be proven reliable and economical in routine, continuous
usage probably means that it will be essentially unavailable to meet
power generation requirements much before the end of the century. This
will place further emphasis on the search for coals with inherently low
sulfur content or which can be simply processed or cleaned to satisfy
emission control requirements.
Comparison of Projected Coal Supply and
Demand for Electric Utility Use_
Estimated future coal requirements for electric utilities in the conti-
guous United States were taken from data published by the National
Electric Reliability Council (NERC). These data give estimated total
coal requirements by year to 1984; eastern coal requirements were calcu-
lated by the writers (Table 16).
Estimated future coal production for electric utility use was derived
Q
from data published by the. National Coal Association (NCA). These data
give estimated total coal production, in addition to current production:
It is anticipated that the end uses of this new production is expected
to be similar to current production; about two-thirds of eastern coal
will be dedicated to steam uses and about one-third to metallurgical
uses, while essentially all western will be for steam. To derive esti-
mated coal production for electric utility requirements, these factors
were employed in re-calculation of the totals published by NCA.
The following assumptions were used-in estimating future coal supply:
26
-------
(a) The supply base of eastern steam coal was assumed to be
340 million tons. To allow for depletion of established
production operations, this amount was reduced by 3% per year
(the average historical depletion rate).
(b) The supply base of western steam coal was assumed to be
60 million tons. This amount was not reduced for depeletion/
because it is "new" production and depletion is not anticipated
in the next ten years.
(c) Planned or additional tonnages compiled by the NCA were re-
duced by 10 percent to allow for actual capacity utilization.
Because of labor, maintenance, transportation, or other problems,
a 90% capacity utilization factor is experienced.
The resulting estimates of new coal production were added to the supply
base to obtain total estimated coal supply for electric utilities in the
next decade (Table 16). The data are illustrated in Figure 2.
The table shows that deficits in coal supply are projected for eastern
electric utilities beginning in 1979. These deficits increase to more
than 130 million tons by 1984.
Although a surplus of supply over demand is shown for western utilities,
the magnitude,, of the eastern deficit is such that an overall supply
deficit results beginning in 1983.
To be sure, it is important to note NCA's caution that estimated coal
supply additions are rather uncertain past 1981. This could reduce the
size of estimated supply deficits, but it is not at all sure that the
deficits would be easily eliminated. Note further that the situation
indicated by the data in Table 1 is optimistic: projected coal supply
from the East is not reduced because of sulfur content restrictions and
western supplies are not curtailed by environmental litigation. In
short, Table 1 presents minimum coal supply deficits. Actions that
adversely impact supply conditions would both increase to size of the
supply deficit and cause the deficit to occur earlier.
27
-------
NJ
00
2000
1800
1600
1400
1200
c 1000
o
800
600
400
200
• HISTORICAL: National Coal Association. Coal Facts, 1974-1975.
• PROJECTED PRODUCTION: National Coal Association, June 1975.
A PROJECTED DEMAND: National Electric Reliability Council, July 1975.
Western
> Projected
Supply
Deficit
-»—•-«.-«__
Eastern
I
1
1965
1970 1975 1980 1985 1990 1995
Figure 1 Comparison of Projected Coal Supply and Demand for Electric Utility Use
2000
-------
Also of concern is the apparent surplus of western coal supply over
demand, which amounts to 60-70 million tons by the early 1980's. Western
coal is of lower heating value and markedly different combustion condi-
tions in contrast to eastern coal, and cannot be simply used as a replace-
ment for shortfalls in eastern production. Therefore, while fable 1
suggests that the quantity of projected western coal supplies may help
to meet indicated coal demand, the lower quality will lessen its contri-
bution to reduction of the indicated supply deficit in the East. In
this situation, the electric utilities would have the worst of both
possible worlds — a supply deficit in the East and overcapacity in coal
production in the West that cannot really be employed effectively to
eliminate the deficit.
Clearly, great care is required in matching coal production operations
with electric utility demand so as to avoid the potential supply paradox
indicated above.
SUMMARY
Recoverable reserves of coal are of the order of 8 percent of the total
remaining resources. Natural geological conditions, as well as disturbances
to strata from previous mining operations, may combine to make unminable
coal seams which otherwise might be considered as potential reserves.
The quality of coal exerts an important influence on its ultimate use.
The reserves of coking-type coals must be subtracted from the overall
coal reserve figures, because of their dedication to the needs of the
steel industry. Electric utility coal consumers are concerned about
reliable supplies of coals having (at a minimum) satisfactory heating
value and sulfur content. Uncertainties about factors critical to the
development of coal production operations could have a serious adverse
impact upon the availability of future coal supplies to meet projected
demands.
29
-------
REFERENCES
1. This estimate includes coals to a depth of about 3,000 feet.
Additional coals are known to occur down to about 6,000 feet, but
these are not considered further in this treatment because their
extraction prospects are regarded as speculative.
2. U.S. Bureau of Mines, "Strippable Reserves of Bituminous Coal and
Lignite in the United States," Information Circular 8531, 1971.
3. See, for example, Phelps, E.R. "Modern Mining Methods-Surface" in
Elements of Practical Coal Mining," Cassidy, Editor, Am. Institute
Mining, Metallurgical, and Petroleum Engineers, New York, 1973,
page 377.
4. It is recognized that the reserve base estimates and the recoverable
strippable reserves data were prepared at different times and using
similar but somewhat different criteria. Nevertheless, the data
are not significantly different, lending a measure of confidence
that they could be employed together. This approach is intended to
provide a rough approximation only.
5. Reiber, M., "Low Sulfur Coal: A Revision of Reserve and Supply
Estimates," University of Illinois, Center for Advanced Computation,
CAC Document No. 88, November 30, 1973.
6. "Premium-grade" coking coals have sulfur content of less than 1.3%
and ash content of less than 8.1%. "Marginal-grade" coking coals
have from 1.3 to 1.8% sulfur and ash from 8.1 to 12.0%.
7. National Electric Reliability Council, "Estimated Fossil Fuel
Requirements, Projected Generating Capacity, and Electric Energy
Production for the Electric Utility Industry (Contiguous U.S.)
1975-1984," Manuscript, July 1975.
30
-------
National Coal Association, "New Coal Mines and Major Expansions ci
Existing Mines Planned, Announced, or Under Construction in the
United States; 1975-1985," Manuscript, June 1975.
31
-------
TABLE 1
SUMMARY OF RECOVERABLE COAL RESERVES BY RANK (MILLION TONS)
Percent of
Rank Strippable Deep Total Total Resources
Anthracite
Bituminous
Subb ituminous
Lignite
TOTAL 45,031 88,084 133,115
45
13,597
24,318
7,071
2,186 2,231
56,616 70,213
29,282 53,600
7,071
10
10
13
2
32
-------
TABLE 'I
(I)
(2)
(3)
(4)
(5)
(6)
(7)
ANTHRACITE Total Esti- Total Mea- Total Strip- Total Under- Projected Re- Projected Re- Total Pro- Recoverable
jected Re- Resources as
coverable Percent of
Reserves Identified
Resources
(5) + (6) (7) i (1)
(million tons)
ARKANSAS
COLORADO
NEW MEXICO
PENNSYLVANIA
VIRGINIA
WASHINGTON
TOTAL
mated Iden-
tified Re-
sources
Jan. 1,1972
0-3000 ft.
430
78
4
20,510
335
5
21,362
sured and
Indicated
"Reserve
Base"
Jan. 1,1974
0-1000 ft.
96
28
2
7,120
138
—
7,384
pable Reserve ground Reserve
Base Base
0-150 ft. (2) - (3)
96
28
2
90 7,030
138
—
90 7,294
coverable
Strippable
Reserves
(4) x 50%
—
—
45
—
—
45
coverable
Underground
Reserves
(4) x 30%
28
8
netj
2,109
41
—
2,186
28
8
ne
-------
TABLE 3
BITUMINOUS
(million tons)
ALABAMA
ALASKA
ARIZONA
ARKANSAS
COLORADO
GEORGIA
ILLINOIS
INDIANA
IOWA
KANSAS
KENTUCKY
MARYLAND
MICHIGAN
MISSOURI
MONTANA
NEW MEXICO
N. CAROLINA
OHIO
OKLAHOMA
OREGON
PENNSYLVANIA
(1)
Total Esti-
mated Iden-
tified Re-
sources
Jan. 1,1972
0-3000 ft.
13,342
19,413
21,246
1,638
62,339
24
139,124
34,573
6,509
18,674
64,842
1,158
205
31,014
2,299
10,752
110
41,358
3,281
50
56,759
(2)
Total Mea-
sured and
Indicated
"Reserve
Jan. 1,1974
0-1000 ft.
1,955
1,201
—
537
10,097
1
65,665
10,623
2,885
1,388
25,541
1,048
119
9,487
1,384
1,777
31
21,077
1,294
—
23,880
(3)
Total Strip-
pable Reserve
Base
0-150 ft.
157
1,201
—
231
870
— •
12,223
1,674
1,000
1,388
7,354
146
1
3,414
—
250
—
3,654
434
—
1,091
(4)
Total Under-
ground Reserve
Base
(2) - (3)
1,798
—
~
306
9,227
--
53,442
8,949
2,885
—
18,187
902
118
6,073
1,384
1,527
31
17,423
860
—
22,789
(5)
Estimated
strippable
Reserves
(USBM 1C
8531)
134
480
—
149
500
—
3,247
1,096
180
375
1,758
21
1
1,160
—
--
—
1,033
111
—
752
(6) (7) (8)
Projected Re- Total Pro- Recoverable
coverable jected Re- Resources as
Reserves coverable Percent of
Identified
(4) x 30% (5) + (6) Resources
(7) i (1)
539
--
~
92
2,768
__
16,032
2,685
865
~
5,456
271
35
1,822
415
458
9
5,277
258
__
6,837
673
480
—
241
3,268
—
19,279
3,781
1,045
375
7,214
292
36
2,982
415
458
9
6,260
369
—
7,589
5
2
—
15
5
—
14
11
16
2
11
25
18
10
18
4
8
15
11
—
13
-------
01
BITUMINOUS
(million tons)
TENNESSEE
TEXAS
UTAH
VIRGINIA
WASHINGTON
WEST VIRGINIA
WYOMING
OTHER
(1)
Total Esti-
mated Iden-
tified Re-
sources
Jan. 1,1972
0-3000 ft.
2,572
6,048
23,541
9,352
1,867
100,628
12,705
610
(2)
Total Mea-
sured and
Indicated
"Reserve
Jan. 1,1974
0-1000 ft.
987
—
4,043
3,513
251
39,590
4,524
—
(3)
Total Strip-
pable Reserve
Base
0-150 ft.
320
—
262
679
—
8,984
—
—
TABLE 3
(Concluded)
(4)
Total Under-
ground Reserve
Base
(2) - (3)
667
—
3,781
2,834
251
30,606
4,524
—
(5)
Estimated
strippable
Reserves
(USBM 1C
8531)
74
—
150
258
—
2,118
—
~
(6)
Projected Re-
coverable
Reserves
(4) x 30%
200
—
1,134
850
75
9,181
1,357
~
(7)
Total Pro-
jected Re-
coverable
(5) + (6)
274
—
1,284
1,108
75
11,299
1,357
—
(8)
Recoverable
Resources as
Percent of
Identified
Resources
(7) * (1)
11
~
5
12
4
11
11
—
TOTAL
686,033
232,858
45,333
188,564
13,597
56,616
70,213
10
OTHERS = California, Idaho, Nebraska, Nevada
* No coal shown in strippable reserve base for Iowa in Project Independence. Value from 1C 8531 used here.
+ Value shown in strippable reserve base for West Virginia in Project Independence is less than 1C 8531 recoverable
reserves. 1C 8531 value of "recoverable strippable resource" (roughly equivalent to reserve base) used here
instead.
-------
TABLE 4
OJ
ON
SUBBITUMINOUS
ALASKA
ARIZONA
COLORADO
MONTANA
NEW MEXICO
OREGON
UTAH
WASHINGTON
WYOMING
OTHERS
(2)
Total Esti-
mated Iden-
tified Re-
sources
Jan. 1,1972
0-3000 ft
110,668
—
18,242
131,855
50,671
284
180
4,190
107,951
32
(3)
Total Mea-
sured and
Indicated
"Reserve
Base"
Jan. 1,1974
0-1000 ft
10,148
350
4,745
99,212
2,615
1
—
1,695
46,704
—
, (4)
Total Strip-
pable Reserve
Base
0-150 ft
5,902
400*
—
35,431
2,646+
neg
—
500
23,674
—
(5)
Total Under-
ground Reserve
Base
(3) - (4)
4,246
—
4,745
63,781
607
—
—
1,195
• 23,030
—
(6)
Estimated
Recoverable
Strippable
Reserves
(USBM 1C
8531)
3,926
387
—
3,400
2,474
—
—
135
13,971
—
(7)
Projected Re-
coverable
Underground
Reserves
(5) x 30%
1,274
—
1,424
19,134
182
—
—
359
6,909
—
(8)
Total Pro-
jected Re-
coverable
Reserves
(6) + (7)
5,200
387
1,424
22,534
2,656
—
—
494
20,880
—
<9)
Recoverable
Resources as
Percent of
Identified
(8) •* (2)
5
a
17
5
—
—
12
19
—
TOTAL
424,073
165,470
67,865
97,604
24,318
29,282
53,575
13
OTHERS- California, Idaho
'Project Independence reserve base for Arizona less than 1C 8531 recoverable reserves} 1C 8531 data used here.
+IC 8531 data used here.
-------
TABLE 5
LO
LIGNITE'
ALABAMA
ALASKA
ARKANSAS
MONTANA
N. DAKOTA
S. DAKOTA
TEXAS
WASHINGTON
OTHERS
(2)
Total Esti-
mated Iden-
tified
Resources
Jan. 1,1972
0-3000 ft.
2,000
(?)
350
87,521
350,630
2,031
6,824
117
46
(3)
Total Mea-
sured and
Indicated
"Reserve
Base"
Jan. 1,1974
0-1000 ft.
1,027
296
32
7,131
16,003
428
3,272
8
—
(4)
Total Strip-
pable Reserve
Base
0-250 ft.
1,027
—
—
7,131
16,003
428
3,272
8
—
(5) (6)
Total Under- Estimated Re-
ground Reserve coverable
Base Strippable
Reserves
(3) - (4) (USBM 1C
8531)
__ __
__
25
3,497
2,075
160
1,309
—
__
(7) (8)
Projected Re- Total Pro-
cover able jected Re-
Underground coverable
Reserves Reserves
(5) x 30% (6) + (7)
— —
—
25
3,497
2,075
160
1,309
—
—
(9)
Recoverable
Reserves as
Percent of
Identified
Resources
(8) * (2)
—
—
7
4
1
8
19
—
—
TOTAL
449,519
28,197
27,901.
7,071
7,071
OTHERS = California, Idaho, Louisiana, Mississippi
-------
Notes to Tables 2-5
(1) Source is U.S. Bureau of Mines, "Demonstrated Coal Reserve Base of
the United States on January 1, 1974," Mineral Industry Survey, June
1974.
Note that the maximum depth for all ranks of coal except lignite, is
1000 feet. Only the lignite beds that can be mined by surface methods
are included; depths are less than about 120 feet. Seams greater than
the following thicknesses are included: bituminous coal and anthracite,
28 inches or more; subbituminous coal and lignite, 60 inches or more.
Note that these data include 100% of the coal in-place, and no recoverability
factor has been included.
(2) Source is EPRI, using data from the U.S. Bureau of Mines and U.S.
Geological Survey.
(3) Source is U.S. Bureau of Mines, "Strippable Reserves of Coal and
Lignite in the United States," Information Circular 8531, 1971.
38
-------
TABLE 6
INDICATED RECOVERY PERCENTAGE FOR STRIPPABLE DEPOSITS (RESERVE BASE -f RECOVERABLE RESERVES)
STATE
ALABAMA
ALASKA
ARIZONA
ARKANSAS
COLORADO
GEORGIA
ILLINOIS
INDIANA
IOWA
KANSAS
KENTUCKY
MARYLAND
MICHIGAN
MISSOURI
MONTANA
NEW MEXICO
NO. CAROLINA
NO*. DAKOTA
OHIO
OKLAHOMA
OREGON
PENNSYLVANIA
SO. DAKOTA
TENNESSEE
TEXAS
UTAH
VIRGINIA
WASHINGTON
WEST VIRGINIA
WYOMING
OTHERS
ANTHRACITE
76
BITUMINOUS
85
40
65
57
27
65
18
27
24
14
100
34
28
26
69
23
57
38
24
SUBBITUHINOUS
67
97
LIGNITE
78
10
93
49
13
37
40
27
59
MEAN
41
49
43
-------
TABLE 7
SUMMARY COMPARISON OF ESTIMATES OF COAL RESOURCES
AND RESERVES {£0.7% SULFUR) JANUARY 1, 1965
(Million Short Tons)
Conventional Estimates
(Tonnage Only)
Resources
Reserves
Standardized Estimates
(Normalized for Btu/Sulfur)
Resources
Reserves
Appalachian
Bituminous 37,320
Anthracite 12,550
Interior
Bituminous 445
Rockies
Bituminous
Subbituminous
Lignite
West Coast
Bituminous
Subbituminous
Total 626,500
4,105
1,630
70
44,784
14,056
472
4,926
1,826
74
45,215
181,670
344,620
4,585
19,470
37,905
48,581
46,329
0
4,925
4,632
0
900
3,780
80
340
855
0
76
0
68,185
155,077
16,459
40
-------
TABLE 8
POTENTIAL RECOVERABLE RESERVES OF ANTHRACITE (MILLION TONS)
(1)
RECOVERABLE
STRIPPABLE
RESERVES
ARKANSAS
COLORADO
NEW MEXICO
PENNSYLVANIA 45
VIRGINIA
WASHINGTON
TOTAL 45
ASSESSED RESERVES
(2) (3) (4) (5) (6) (7)
RECOVERABLE TOTAL HEATING Lbs S ASSESSMENT OF
UNDERGROUND RECOVERABLE VALUE MM btu POTENTIAL RECOV-
RESERVES RESERVES MM Btu/ton REMARKS ERABLE RESERVES
28 28 28.0 1.64 S too high; reserves low;
development unlikely
8 8 27.8 0.4J S satisfactory; reserves low;
development unlikely
neg neg 27.2 0.66 S marginal; reserves low;
development unlikely
2,109 2,154 26.2 0.54 S satisfactory; reserves OK
development likely
41 41 23.8 0.59 S marginal; reserves low;
development unlikely
—
2,186 2,231
< 0.7 Ib S/MM Btu =• 2,154 million tons
> 0.7 Ib S/MM Btu = O
O
O
O
2,154
O
—
2,154
TOTAL 2,154
-------
TABLE 9
POTENTIAL RECOVERABLE RESERVES OF BITUMINOUS COAL (MILLION TONS)
ALABAMA
ALASKA
ARIZONA
ARKANSAS
COLORADO
GEORGIA
ILLINOIS
INDIANA
IOWA
KANSAS
KENTUCKY
MARYLAND
MICHIGAN
MISSOURI
MONTANA
NEW MEXICO
(1)
RECOVERABLE
STRIPPABLE
RESERVES
134
480
—
149
500
__
3,247
1,096
180
375
1,758
21
1
1,160
—
—
(2) (3)
RECOVERABLE TOTAL
UNDERGROUND RECOVERABLE
RESERVES RESERVES
539 673
480
—
92 241
3,768 3,268
—
16,032 19,279
2,685 3,781
865 1,045
375
5,456 7,214
271 292
35 36
1,822 2,982
415 415
458 458
(4)
HEATING
VALUE
MM Btu/ton
27.6
24.6
23.2
28.2
25.6
—
25.2
1
26.0
23.4
25.4
26.8
28.2
26.2
24.6
24.4
25.8
(5)
Lbs S
MM BtU
0.80
0.16
0.78
1.60
0.55
~
2.54
2.31
4.70
2.38
1.87
1.00
1.54
3.58
1.07
0.54
(6) (7)
ASSESSMENT OF
POTENTIAL RECOV-
REMARKS ERABLE RESERVES
S medium; reserves adequate;
development likely
S low; reserves adequate;
development likely
S medium; reserves low;
development unlikely
S high; reserves marginal;
development unlikely
S satisfactory; reserves high;
development likely
~
S high; reserves large;
development possible
S high; reserves large;
development possible
S high; reserves adequate;
development unlikely
S high; reserves marginal;
development possible
S high; reserves large;
development possible
S medium; reserves marginal;
development likely
S high; reserves small;
development unlikely
S high; reserves large;
development unlikely
S medium; reserves adequate;
development likely
S low; reserve adequate;
673
480
0
0
3,268
0
19,279
3,781
0
375
7,214
292
0
0
415
458
development likely
-------
NO. CAROLINA
OHIO
OKLAHOMA
OREGON
PENNSYLVANIA
TENNESSEE
TEXAS
UTAH
VIRGINIA
WASHINGTON
WEST VIRGINIA
WYOMING
OTHER
TOTAL
ASSESSED
(1)
RECOVERABLE
STRIPPABLE
RESERVES
—
1,033
111
—
752
74
—
150
258
—
2,118
—
__
13,597
RESERVES <
>
(2)
RECOVERABLE
UNDERGROUND
RESERVES
__
5,227
258
—
6,837
200
—
1,134
850
75
9,181
1,357
__
56,616
0.7 Ib S/MM Btu =
0.7 < 1.5
(3)
TOTAL
RECOVERABLE
RESERVES
6,260
369
—
7,589
274
—
1,284
1,108
75
11,299
1,357
—
70,213
8,030 million
9,238
TMlI.r 0
(Concluded)
(4) (5)
HEATING Lbs S
VALUE MM BtU
MM Btu/ ton
^_
25.8 2.25
27.6 1.52
22.4 0.63
27.6 1.31
27.2 1.03
25.8 1.78
26.8 0.37
28.0 0.57
24.8 0.48
27.4 1.82
25.8 0.54
—
tons
(6)
REMARKS
S high; reserves large;
development possible
S high; reserves marginal;
development possible
S low; reserves negligible;
development unlikely
S medium; reserves large;
development likely
S medium; reserves marginal;
development likely
S high; reserves negligible;
development unlikely
S low; reserves adequate;
development likely
S low; reserves adequate;
development likely
S low; reserves small;
development possible
S high; reserves large;
development possible
S low; reserves adequate;
development likely >
—
Likely = 17,193
Possible = 45,032
Unlikely = 4,304
(7)
ASSESSMENT OF
POTENTIAL RECOV-
ERABLE RESERVES
0
6,260
369
0
7,589
274
0
1,284
1,108
75
11,299
1,357
_
65,850
> 1.6 < 2.1 =15,262
> 2.2 =29,695
-------
TABLE 10
POTENTIAL RECOVERABLE RESERVES OF SUBBITUMINOUS COAL (MILLION TONS)
ALASKA
ARIZONA
COLORADO
MONTANA
NEW MEXICO
OREGON
UTAH
WASHINGTON
WYOMING
OTHERS
TOTAL
(1)
RECOVERABLE
STRIPPABLE
RESERVES
3,926
387
—
3,400
2,474
—
~
135
13,971
__
24,318
(2)
RECOVERABLE
UNDERGROUND
RESERVES
1,274
__
1.424
19,134
182
__
--
359
6,909
—
29,282
(3)
TOTAL
RECOVERABLE
RESERVES
5,200
387
1,424
22,534
2,656
__
—
494
20,880
—
53,575
(4) (5)
HEATING Lbs S
VALUE MM BTU
MM Btu/ton
22.4 0.18
23.2 0.78
24.6 0.33
23.4 0.95
23.0 1.05
22.4 0.63
__
22.2 0.72
24.6 0.49
{6)
REMARKS
S low} reserves large}
development possible
S marginal} reserves adequate;
development possible
S lowj reserves adequate}
development likely
S marginal} reserves large:
development possible
S medium} reserves adequate}
development possible
S low} reserves negligible}
development unlikely
--
S medium} reserves adequate}
development possible
S low} reserves large}
development likely
—
ASSESSMENT OF
POTENTIAL RECOV-
ERABLE RESERVES
5,200
387
1,424
22,534
2,656
0
0
494
20,880
~
53,575
Likely •= 22,304
Possible = 31,271
ASSESSED RESERVES < 0.7 Ib S/MM Btu - 27,504 million tons
> 0.7 < 1.5 - 26,071
-------
TOTAL
TABLE 11
POTENTIAL RECOVERABLE RESERVES OF LIGNITE (MILLION TOMS)
(1)
RECOVERABLE
STRIPPABLE
RESERVES
ALABAMA
ARKANSAS 25
MONTANA 3,497
N. DAKOTA 2,075
S. DAKOTA 160
TEXAS 1,309
WASHINGTON
OTHERS
(2) (3) (4) (5) (6) (7)
RECOVERABLE TOTAL HEATINQ Lbs S ASSESSMENT OF
UNDERGROUND RECOVERABLE VALUE MM Btu REMARKS POTENTIAL RECOV-
RESERVES RESERVES MM Btu/ton ERABLE RESERVES
WHO — — »» __ MM
25
3,497 22.4 0.54 S low; reserves large;
development likely
2,075 21.8 0.64 S low; reserves large;
development likely
160 21.8 0.92 S marginal; reserves small;
development unlikely
1,309 22.0 1.27 S medium; reserves adequate;
development possible
—
—
0
0
3,497
2,075
0
1,309
0
0
7,071
ASSESSED RESERVES < 0.7 Ib S/MM Btu
> 0.7 < 1.5
7,071
5,572 million tons
1,309
Likely = 5,572
Possible" 1,309
6,881
-------
TABLE 12
SUMMARY OF POTENTIAL RECOVERABLE RESERVES (MILLION TONS)
>1.5 <2.1 Ibs/MMBtu >2.1 Ibs/MMBtu Total
— — 2,154
15,262 33,320 65,850
— — 53,440
— — 6,881
TOTAL 43,260 36,618 15,262 33,320 128,325
RANK
ANTHRACITE
BITUMINOUS
SUBBITUMINOUS
LIGNITE
<0.7 Ibs/MMBtu
2,154
8,030
27,504
5,572
>0.7 <1.5 Ibs/MMBtu
—
9,238
26,071
1,309
-------
TABLE 13
COMPARISON OP ASSESSED RECOVERABLE RESERVES OF BITUMINOUS AND SUBBITUMINOUS COAL
AGAINST PROJECTED PRODUCTION (MILLION TONS)
ALABAMA
ALASKA
ARI ZONA
ARKANSAS
COLORADO
GEORGIA
ILLINOIS
INDIANA
IOWA
KANSAS
KENTUCKY
MARYLAND
MICHIGAN
MISSOURI
MONTANA
NEW MEXICO
N. CAROLINA
OHIO
OKLAHOMA
OREGON
(1) (2) (3)
ASSESSED RECOVERABLE RESERVES
LIKELY POSSIBLE UNLIKELY
673
480 5,200(s)
387 neg
neg
l,424(s)
3,268
neg
19,279
3,781
1,045
375
7,214
292
36
2,982
415 22,534(s)
458 2,656{s)
neg
6,260
369
neg
(4)
LBS S
MM BTU
0.80
0.16
0.78
. 1.60
0.55
2.54
2.31
4.70
2.38
1.87
1.00
1.54
3.58
1.07
0.54
2.25
1.52
0.63
(5)
PRESENT
PRODUCTION
(1973)
19.2
0.7
3.3
0.4
6.2
61.6
25.3
0.6
1.1
127.7
1.8
4.7
10.4
9.1
45.8
2.2
(6)
FORECAST
NEW
CAPACITY
(by 1985)
19.6
27.0
2.9
40.6
11.0
69.0
57.3
13.6
14.2
(7)
OPTIMISTIC
PRODUCTION
IN 1985
(5) + (6)
34.0
1.0
29.5
0.5
7.6
86.8
30.0
1.0
1.0
164.8
2.0
5.0
65.1
20.4
48.6
2.0
OPTIMISTIC PRODUCTION BY SULFUR
CONTENT (Lbs/MM Btu)
<0.7 >0.7<1.5- >1.5
34.0
1.0 1.0
29.5
0.5
7.6
86.8
30.0
1.0
1.0
164.8
2.0
5.0
65.1
20.4
48.6
2.0
-------
00
(1) (2) (3)
ASSESSED RECOVERABLE RESERVES
LIKELY POSSIBLE UNLIKELY
PENNSYLVANIA 7,589
TENNESSEE 274
TEXAS neg
UTAH 1,284 neg(s)
VIRGINIA 1,108
WASHINGTON 494 (s)
75
WEST VIRGINIA 11, 299
WYOMING 20, 880 (s)
1,357
OTHER neg
TABLE 13
(Concluded)
(4) (5) (6) (7)
FORECAST OPTIMISTIC
PRESENT NEW PRODUCTION OPTIMISTIC PRODUCTION BY SULFUR
LBS S PRODUCTION CAPACITY IN 1985 CONTENT (Lbs/MM Btu)
MM BTU (1973) (by 1985) (5) + (6) <0.7 >0.7<1.5 >1.5
1.31 76.4 31.5 88.8 88.8
1.03 8.2 4.9 11.1 11.1
1.78
0.37 5.5 36.6 40.7 40.7
0.57 34.0 13.7 39.2 39.2
0.48 3.3 3.0 5.5 5.5
1.82 115.5 t 71.0 157.6 157.6
0.54 14.9 123.4 134.6 134.6
TOTAL 17,193 48,657 4,304
22,034(s) 31,271(8) neg(s)
577.9
539.3
976.8
249.0
230.5
497.3
-------
TABLE 14
COMPARISON OF ASSESSED RECOVERABLE RESERVES OF LIGNITE AGAINST PROJECTED PRODUCTION (MILLION TONS)
ALABAMA
ARKANSAS
MONTANA
N. DAKOTA
S. DAKOTA
TEXAS
WASHINGTON
OTHERS
TOTAL
ASSESSED RECOVERABLE RESERVES
LIKELY
3,497
2,075
5,572
POSSIBLE
UNLIKELY
neg
neg
160
1,309
LBS S
MM BTU
0.54
0.64
0.92
1.27
FORECAST
PRESENT NEW
PRODUCTION CAPACITY
(1973) (to 1985)
0.3
6.9
6.9
1,309
160
14.1
21.9
16.7
38.6
OPTIMISTIC OPTIMISTIC PRODUCTION BY SULFtR
PRODUCTION CONTENT (Lbs/MM Btu)
IN 1985* <0.7 >0.7 <1.5 >1.5
0.3 0.3
27.1 27.1
21.9
49.3
27.4
21.9
21. 9
* See note on Table 13.
-------
TABLE 15
Resources, Reserve Base, And Recoverable Reserves of
Bituminous Coals in the United States with Respect to Coking Properties
(billion tons)
Item
1. Remaining
Identified
Resources
2 . Demonstrated
Reserve Base
3. Estimated
Recoverable
Reserves
Notes
(a) Averitt, P.
1975.
(b) Sheridan, E
(1)
Total
Bituminous
Coal
747
67(£)
"Coal Resources of
. T. "United States
(2)
Premium-Grade
101
20 W)
6
the United States,
(3)
Coking Coals
Marginal Grade
82
5
-------
TABLE 16
COMPARISON OF PROJECTED COAL SUPPLY AND DEMAND FOR ELECTRIC UTILITY USE
1975
1976
1977
1978
1979
1980
1981
1982
1983
340.0
17.9
357.9
345.2
12.7
329.8
45.7
375.5
365.7
9.8
319.6
71.0
390.6
381.0
9.6
309.4
96.7
406.1
405.2
0.9
299.2
116.5
415.7
419.2
(3'. 5)
289.0
129.3
418.3
444.7
(26.4)
278.8
136.6
415.4
461.2
(45.8)
268.6
141.6
410.2
479.1
(68.9)
258.4
145.2
403.6
499.8
(96.2)
248.2
148.6
396.8
528.1
(131.3)
60.0
19.5
7571
74.4
5.1
60.0
49.5
104.5
88.0
21.5
60.0
80.2
14~072
106.8
33.4
60.0
119.4
179.4
135.5
43.9
60.0
171.5
171.6
59.9
60.0
208.5
268.5
196.3
72.2
60.0
226.7
286.7
213.2
73.5
60.0
242.8
302.8
231.8
71.0
60.0
250.0
310.0
240.8
69.2
60.0
253.6
313.6
252.7
60.9
EAST
Supply: (a)
(b)
Total Supply
Demand: (NERC)
Surplus of Supply
WEST
Supply: (c)
(b)
Total Supply
Demand: (NERC)
Surplus of Supply
TOTAL SUPPLY: (d)
(b)
DEMAND: (NERC)
Surplus of Supply
(a) Supply Base - 340 million tons discounted 3%/year
(b) Source - NCA study on additional mine capacity, at 90% capacity factor
(c) Supply Base - 60 million tons
(d) Supply Base - addition of eastern and western
1984
400.0
37.4
437.4
419.6
17.8
389.8
95.2
485.0
453.7
31.3
379.6
151.2
530.8
487.8
43.0
369.4
216.1
585.5
540.7
44.8
359.2
288.0
64772
590.8
56.4
349.0
337.8
686.8
641.0
45.8
338.8
363.3
702.1
674.4
27.7
328.6
384.4
713.0
710.9
2.1
318.4
395.2
713.6
740.6
(27.0)
308.2
402.2
, 710.4
780.8
(70.4)
-------
PROBLEMS AND CONTROL OPTIONS
USING LOW SULFUR COAL IN UTILITY BOILERS
George P. Green
Environmental Affairs and Planning
Public Service Company of Colorado
Denver, Colorado
53
-------
PROBLEMS AND CONTROL OPTIONS
USING LOW SULFUR COAL IN UTILITY BOILERS
Presented by:
George P. Green
Manager, Environmental Affairs and Planning
Public Service Company of Colorado
Denver, Colorado
This paper describes briefly the problems associated with the handling and
burning of low sulfur fuels and the control options available for flue gas
emissions.
Aside from the question of emission control options, one must also consider the
other problems associated with low sulfur, low rank western coals such as the
handling of the coal, i.e., railroad cars and unloading facilities at the power plant.
The paper will also point out the considerations to be given in design for pulver-
izers, increased primary air for pulverizer drying, and boiler thermal efficiencies.
Additional considerations must be given to the boiler design to account for the
particular flame characteristics of the low rank fuels along with the high fusion
ash and the basic composition of the ash.
The paper will also present in chronological order the efforts of Public Service
Company of Colorado (PSCo) in installing control equipment for the low sulfur western
coals. A brief history of the PSCo experience with electrostatic precipitators to
include the installation of cold side precipitators, the addition of gas conditioning,
and, finally, the application of hot side electrostatic precipitators. It will also
discuss the use of wet stack gas scrubbers for particulate collection and briefly
review the use of bag filters.
Adequate particulate removal efficiencies can be obtained with all the control
devices mentioned above, however, high operational and maintenance cost and, in some
cases, a lower than satisfactory availability record for the devices must be consi-
dered in a case-by-case choice for each specific coal and location.
There is no one specific device that is recommended over and above the other
control devices; it is a matter of judgement and good engineering practices on behalf
of the owner and operator.
54
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PROBLEMS AND CONTROL OPTIONS
USING LOW SULFUR COAL IN UTILITY BOILERS
INTRODUCTION
With the advent of the Clean Air Act and subsequent regulations promulgated by
the Environmental Protection Agency, a great deal of emphasis has been placed on
burning low sulfur western sub-bituminous coal. The object of this paper is to
attempt to shed some light on the problems incurred with burning such coals. It will
be of particular interest to note these problems when considering the conversion of
present boilers now burning an eastern coal to a configuration to burn western low
sulfur coals. Not only are there serious problems choosing the proper particulate
control device, but there are problems to be considered in the handling of the
western coals and boiler designs.
It is not my attempt to layout guidelines for the design and production of a new
power generation facility but to dwell more on the problems that have occurred and
that should be understood when considering a switch in fuel supply to the western
low sulfur sub-bituminous coals. Since the promulgation of the Environmental Protec-
tion Agency regulations, several companies have attempted to convert present boilers,
all of which have experienced some difficulty not only in the operation of the boiler
but, probably more importantly, in the operation of the control devices associated
with the boiler.
The Public Service Company of Colorado has been burning low sulfur coal since
the 1920s and instituted the use of electrostatic precipitators in the early 1960s.
The information contained herein is related to the experiences of the Public Service
Company of Colorado and other companies which have installed systems using western
coal as the main source of fuel.
WESTERN COAL
Western coal reserves, which are predominantly lower rank coals and include
lignite, sub-bituminous and bituminous, have been measured by the United States
Bureau of Mines to be 216 billion tons, as indicated in Table 1. The distribution
of this coal is primarily in Montana, Wyoming, North Dakota and Colorado. During
1974 the estimated rate of production was greater than 2 million tons per year in
each of the eleven western states, with Wyoming producing the greatest tonnage. It
55
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TABLE 1 COAL IN THE WESTERN U.
State
Arizona
Colorado
Kansas
Missouri
Montana
New Mexico
North Dakota
Oklahoma
South Dakota
Texas
Utah
Washington
Wyoming
RESERVES AND PRODUCTION BY STATE
In Million Tons
In-place coal
reserve
350
14,870
1,388
9,488
107,727
4,394
16,003
1,294
428
3,272
4,042
1,954
51,228
Estimated 1974
production
3.2
6.9
.8
4.3
13.6
9.5
7.2
2.4
.0
6.0
6.5
3.9
20.5
New mine*
capacity
by 1983
8
** NA
.5
NA
21
5
NA
NA
NA
.1
1.3
1
82.5
216,439
84.8
119.4
* New capacities reported by 1983 are representative of firm plans
announced by major producing firm. The values given do not
reflect plans for coal gasification.
** NA indicates that no value was given in the reference. Some of
the states affected are known to be experiencing major expansions
in coal production.
Sondreal and Tufte, 1975.
56
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is anticipated that both Wyoming and Montana will experience the greatest increase in
production of the western coals. As indicated in Table 1, plans for expansion of
mines in the West will increase the production to over 200 million tons per year by
1983.
Important properties of western coals compared to eastern and midwestern coals
include a lower sulfur content, generally averaging approximately 0.7 percent, a
lower heating value, a higher moisture content, and higher alkali oxides such as
Na-O, MgO and CaO. "Alkali content tends to be the highest in the lowest rank coal,
lignite, and progressively less prevalent in the sub-bituminous and bituminous coals.
This trend is related to the greater ion exchange capacity of low-rank coals compared
with the higher rank. Variations in alkali content are also influenced by the
minerology of the overburden and the course of ground water movement. Alkali content
in western coal ash varies from under 10 to over 50 pet, with important variations
occurring within individual mines.
"A guideline for assessing the importance of the amount of alkali in western
coal is the ratio of the alkali to coal sulfur. For a coal containing 7.5 pet ash
and 20 pet alkali in the ash, the total alkali is chemically equivalent to slightly
more than 120 pet of a 0.7 pet sulfur content. For some lignites, the alkali/sulfur
ratio can be several hundred percent. Thus there is ample alkali to interact
importantly with sulfur oxides in a wet scrubber in burning many western coals."
(Sondreal and Tufte, 1975.)
For relative comparison purposes only, Table 2 shows typical analyses of
western and midwestern coals. These do not represent any specific mines.
Transportation. At present, the majority of the western coals are strip mined
and used at mine mouth-plant operations or transported to load centers by unit trains.
In the location of the more recent large units there seems to be a preference for the
use of the unit train concept.
In considering unit train operation, there are problems associated with the
transportation of western coals which should be considered. For example, the lower
rank coals are generally much higher in moisture content, so built into transporta-
tion cost is the added expense of hauling more moisture from the mine to the power
plant. More tonnage is hauled because of the lower BTU, thus adding the expense of
additional railroad cars.
In unit train operation, the cars travel approximately 100 thousand miles per
year. In many cases, coal is loaded at the mine on the fly and unloaded at the
power plant, also on the fly. Therefore, the cars virtually never come to a
standstill except for momentary stoppage of the trains to transfer crews. This
57
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TABLE 2 TYPICAL COAL AND ASH ANALYSIS
Colorado
Coal
Wyom i ng
Coal
Midwestern
Steam Coal''
Proximate Analysis
as Received:
Moisture
Volatile Matter
Fixed Carbon
Ash
Sulfur
HHV, BTU/lb.
Hardgrove Grindability
Pounds of Sulfur/million Bill1;
9.8%
35-3%
45-5%
9-4*
0.7?
11,010
44
0.64
29.0%
33.4%
32.4%
5.2%
0.6%
8,250
55
0.73
2.2%
35-6%
53.3%
8.9%
2.2%
13,280
1.66
Ash Fusion Temperatures:
Initial Deformation
Softening (H=W)
Fluid
2560°F.
2665°F.
2690°F.
2010°F.
2150°F.
2210°F.
2090°F.
2210°F.
2330°F-
Ash Composition:
Silica
Alumina
Iron Oxide
Phosphorus Pentoxide
Titanium Oxide (Ti02)
Calcium Oxide (CaOJ
Magnesium Oxide (MgO)
Sodium Oxide (N
Potassium Oxide
Sulfur Trioxide
54.1%
28.3%
.0%
3-
1
1%
0.7%
0%
5%
5%
0%
1.8%
34.5%
20.6%
6.8%
0.2%
Trace
20.0%
2.7%
0,2%
4,3%
10.7%
21
47.2%
23.2%
9%
2%
0%
4%
.6%
.4%
1.6%
1.6%
1.
2.
I/ Pittsburgh No. 8 steam coal from Jefferson County, Ohio.
Green, G. P. and W. S. Landers, 1974.
58
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factor becomes very important when considering that an average box car will travel
approximately 17 thousand miles per year. Maintenance of rolling stock now takes on
a new perspective; and when considering historical maintenance cost on an annual
basis, it must now be increased several fold. In the past, a typical piece of
rolling stock, traveling approximately 17 thousand miles per year, had an average
life of twenty years, now it covers the equivalent mileage in three or four years.
Not only is there now greater annual wear in the moving parts of the car such
as the wheels and truck assemblies, but the very nature of the western coal tends
to aggrevate the problems of spontaneous combustion, thus necessitating additional
care in providing tight railroad cars. Also, western coals are less dense than an
eastern coal, so that when a car is fully loaded, the center of gravity will tend
to shift upward and, in some cases, will induce a serious swaying in the train
necessitating lower tonnage hauled per car than originally contemplated.
Since most unit trains are designed with special features such as swivel
couplers or automatic dump hoppers, it is imperative that when entering into the
unit train operation the utility company give serious consideration to how many spare
railroad cars should be purchased to complement the original unit train. Actual prac-
tice has shown that a greater number of spare cars are necessary than were initially
projected during the design of the unit train operation.
Handling. There appears to be a greater number of fines in the western coal.
While in transit the fines tend to dry out rapidly causing a dust problem at un-
loading facilities. The dust problem can be minimized with the proper design and
utilization of dust supression systems and wetting agents.
A greater percentage of fines also increases dust problems associated with long-
term storage. In spite of efforts to compact coal in the storage piles, the inherent
strong winds in the western portion of the United States tend to cause fugitive dust
problems in the coal storage area. When designing long-term coal storage areas,
special consideration has to be given to such items as lowering wells and bulk
handling equipment. When handling the coal in storage, coal should not be allowed
to fall any appreciable distance in the open air as it aggrevates the fugitive dust
problem. It is highly recommended that such items as telescoping spouts be used to
minimize dust associated with the fines.
Spontaneous combustion in coal piles tends to be greater with the western coal.
Extensive surveillance by plant operation personnel as well as good practices in
compacting stored coal are necessary. There is an economic consideration to the
amount of compaction that can be obtained in the coal storage pile. This cost of
compaction must offset against the increased surveillance for coal pile fires, and
59
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the individual utility company must determine which course of action is to its
particular economic advantage.
BOILER OPERATING CONSIDERATION
Because of the higher inherent moisture and lower BTU of western coal, special
consideration must be given to the design of the boiler and associated pulverizers
and air preheaters. Larger air heaters are required, particularly for primary air
heating, to vaporize the inherent moisture in the coal. By the same token, pre-heated
air to the coal pulverizer cannot be of such a temperature to cause ignition of the
coal in the pulverizer. In many cases, there is a fine point in determining the
optimum mill temperature in relation to fires and explosions within the mills.
Since western coals have a lower heating value, greater tonnages are required,
thereby increasing the size of the mills and increasing the capacity of the primary
air fans. Lower boiler thermal efficiencies must be expected due to the loss asso-
ciated with water vaporization, and this has a tendency to lower the fire box
temperatures and requires greater heat absorption surfaces. In nexj designs, these
factors are taken into consideration and can be compensated for. Greater problems
exist in attempting to substitute the western coal in an existing installation
which was designed for use with higher ranking coals.
Burning characteristics are quite different between the higher and lower rank
coals. The lower rank coal (after mill drying) will ignite closer to the burner and
at a faster rate, thus effecting the pattern of heat release and the location of
additional heat absorption surface within the radiant section of the boiler.
Perhaps the most serious consideration to be given to the use of the western
coals is their characteristic fusion temperatures. The high fusion temperature will
generally not permit their use in wet bottom or slag tap boilers. As mentioned
previously, the higher alkalinity in the ash content of the coal will also have an
effect on the ash collection systems. Many boilers designed for the higher rank
coals will encounter plugging in reheat and super-heat sections of the boiler when
burning western coals.
CONTROL OPTIONS
The most difficult problem to overcome in the burning of low sulfur western
fuels is the collection of particulate matter in the flue gas.
Mechanical Collectors. Initially, only mechanical collectors were utilized for
the collection of particulate matter. Mechanical collectors which were properly
designed and maintained exhibited an efficiency of approximately 80-85 percent
removal, which by today's standards would be totally unacceptable.
Electrostatic Precipitators. The next step came with the installation of
electrostatic precipitators. Precipitators were generally designed and predicated
60
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on factors evolved from eastern higher sulfur fuels and, all too often, purchased on
price only. As a result of these earlier efforts in electrostatic precipitators,
many units were found to be entirely too small, reaching, in some instances, only
one-half the efficiency that was originally guaranteed hy the manufacturer .
Resistivity. Another problem encountered was the complete lack of understanding
of the effects of resistivity. Figure 1 is a typical resistivity curve depicting
the variance of resistivity with temperature.
To overcome the problem of high resistivity of low sulfur western fuels, two
courses were taken. One was to install larger electrostatic precipitators that
operate in the traditional 270-300 F range, and the other course was to install
hot side electrostatic precipitators generally operating above 700°F. Both approaches
have their merits. A well constructed cold side precipitator with sufficient surfaces
and low gas velocity can be operated well within design efficiencies. Another
approach tried on a very limited basis is to reduce the operating temperature
within the electrostatic precipitator to approximately 240 F. The resistivity is
reduced at this lower temperature, however you now encounter a very serious dew point
problem. Particulates in combination with S0? tend to raise the dew point temper-
ature and can cause serious corrosion problems within the ductwork.
On the other hand, the hot gas precipitators attempted to overcome the effect
of resistivity by operating at a temperature at which the sulfur content no longer
affected the resistivity. It was thought that if one could overcome the resistivity
difficulties, a precipitator could then be designed to perform at near clear-stack
status. The concept of the hot gas precipitator was encouraging even though this
device encountered some difficulties, particularly in the physical design of the
structure. At these elevated temperatures, there was, obviously, greater movement
in the structural members of the precipitator thus causing warpage and cracking of
the precipitator boxes. Although this problem was serious, with re-evaluation of
the structural design and greater emphasis placed on the thermal expansion, this
problem can and has been overcome. There are some indications that a hot side gas
precipitator, under some conditions, seems to develop a coating on the wire emitters
which reduces the current emitted and, in some cases, enhances back corona.
Additional work needs to be done in this area to determine the extent of this
phenomenon.
Gas Conditioning Agents. Gas conditioning agents have proven to be effective
in improving collection efficiencies of electrostatic precipitators. Gas conditioning
can be applied only with the cold precipitator (270-300 F), since the gas conditioning
agent has an effect only on the surface resistivity of the dust particle. The basic
61
-------
.12
u
l
X
i-l
•H
>
VI
01
oi
10
10
10 -
Normal operating
range for cold
side precipitators
Normal operating
range for hot
side precipl tatorsi
10
200
300
400
500
600
700
800
900
Figure 1. ESP Resistivity Effects
62
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concept of gas conditioning is to inject within the flue gas stream a material such
as liquid SO^ or vaporized sulfuric acid to artificially modify or change the charac-
teristics of the ash. This has the result of allowing the electrostatic precipitator
to act as though it were really cleaning an ash from a higher sulfur eastern fuel.
Therefore, if the design criteria of the electrostatic precipitator were correctly
based on an eastern or higher sulfur fuel, then the limit attained by the addition
of the gas conditioning agent will be to approach the limit of operation of the
precipitator utilizing the aforementioned fuel.
Gas conditioning has aided in bringing many electrostatic precipitators into
compliance and can be an effective method for improving precipitator operation. Gas
conditioning was first attempted on a large scale with western coals in 1970 and
found to have a relatively good success. If the intent is to convert an existing
unit to burn a low sulfur western sub-bituminous coal, gas conditioning can be
considered as an aid to an existing precipitator provided that the precipitator is
at least large enough to provide marginal efficiency.
In designing new plants, the option between the use of a cold precipitator
verus a hot precipitator is up to the individual design considerations of the company
and the specific fuel to be used. If it is planned to utilize a large cold
precipitator on a new installation, it is highly recommended that the initial design
and configuration be predicated on meeting the desired efficiencies without the use
of a gas conditioning agent. I strongly urge that gas conditioning be considered
only as an aid in helping improve the efficiency of the precipitator if the actual
operating efficiency is below the design specifications.
Fabric Filter Baghouse. The fabric filter baghouse can be used for the collection
of particulate matter as an alternate to a precipitator. If one forgets the initial
baghouse installation at the Southern California Edison station and considers the
more recent baghouse installations at Nucla (Colorado) Station of Colorado-Ute and
the Sunberry Station of Pennsylvania Power and Light, the success of these recent
baghouse operations is attractive. Extensive tests performed at the Nucla Station
show overall collection efficiencies greater than 99.9 percent with exit grain
3
loadings of less than 0.005 gr/ft .
The real proof of operation of the baghouse is the visual inspections of the
stack, and these have been basically clear. The Electric Power Research Institute
in conjunction with Meteorology Research Institute and Stearns Roger Inc. recently
completed an in-depth study of the Nucla baghouse installation and will publish
their results in the near future. I believe it is sufficient to state at this time
that the preliminary collection efficiencies of the baghouse exceed the collection
63
-------
efficiencies of the electrostatic precipitator; and I can personally say from
observation on my part that during full operation the three stacks of the Nucla
Station were absolutely clear. I am not suggesting that a baghouse or any other
device is the total answer for participate control. I do, however, suggest that
serious consideration should be given to the fabric filter collection devices not
only for new installations but also for retrofitting older units, particularly
smaller size units.
Scrubber. The final option to be considered for particulate collection of low
sulfur fuels is the use of a wet particulate scrubber. My own company, Public
Service Company of Colorado, has five full size particulate scrubbers in operation
on units that range from 100 megawatts up to 350 megawatts. In addition to Public
Service Company of Colorado, particulate wet scrubbers have been installed at Dave
Johnston unit number 4 of Pacific Power and Light, at Clay Boswell and Aurora
Stations of Minnesota Power and Light Company, the Four Corners Station of Arizona
Public Service Company, and the Holtwood Station of Pennsylvania Power and Light.
Table 3 is a list of some of the design parameters associated with these scrubbers.
The aforementioned scrubber installations \jere designed and installed specifically
as particulate scrubbers. This is not intended as a complete list of scrubbers
that have been installed in the United States during the last four or five years.
A more complete list of all scrubbers in operation in the United States is provided
by Pedco under contract with the Environmental Protection Agency. Figures 2 through 7
are included to depict flow diagrams of the various particulate scrubbers mentioned
previously.
All of the scrubbers at the Public Service Company of Colorado are TCA scrubbers.
You will note the difference between Figure 2, Cherokee Station, and Figure 3,
Valmont Station, is the attempt to operate one module of the Valmont Station as a
SO- scrubber. The Valmont scrubber was our first particulate scrubber and was
constructed in two separate modules. After several years of operation and facing
a more stringent S0« standard, one module was converted to a S0~ mode for experi-
Z, £
mental purposes. Several things have been determined in this short test period,
none of which are particularly surprising. For example, the scaling problem was
intensified when the scrubber liquid pH was controlled. Secondly, the maintenance
and operation costs were extensively increased in the SO- mode. And, finally the
question of how to dispose of the tons of sludge produced by the addition of the
limestone to the scrubber slurry arose.
From a purely operation and maintenance cost standpoint, the operation of a
particulate scrubber far exceeds the cost of operating a precipitator, particularly
64
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1C
a
w
Ui TABLE 3
SUMMARY OF OPERATING PARTICIPATE SCRUBBERS
Public Pacific Power Minnesota
Service Company of Colorado And l.ipht Power And Light
Arapahoe
' Station
No. of Equipped Boilers
Scrubber Capacity, MW
No. of Modules
Vendor
Type
Start Up Date
Si
Availability, PCT.
Particulate Removal, PCT.
SO? Renoval, PCT.
L/C, gal/1000 acf
Pj in. H2O
Treated Gas Flow
1
100
2
UOP
3 stage
TCA
11/73
86
96
45
54.2
10-18
299,000 SCFM
Cherokee
Station
3
600
9
UOP
3 stage
TCA
Unit 1 9/73
Unit 3' 10/72
Unit 4 9/74
Unit 1 55
Unit 3 95
Unit 4 82
97
20
55.9
10-18
Unit 1
302,000 SCFM
Unit 3
360,000 SCFM
Unit 4
897,000 SCFM
Valraont
Station
1
100
2
UOP
3 stage
TCA
11/71
76
96
45
58.3
10-18
299,000 SCFM
Dave
Johnston
Station
1
330
3
CHEMICO
Venturi
4/72
KA
99
40
13
15
1,500,000
ACFM <3
270°F
Clay
Boswell
Station
1
350
1
KREBS
High Pres-
sure Spray
5/73
KA
99
20
8
4
1,300,000
ACFM @
254°F
Aurora
Station
2
116
2
KREBS
High Pres-
sure Spray
6?71
KA
98
20
8
4
291,160
ACFM (3
340°F ea.
boiler
Arizona
Public Service
Four Corners
Station
3
575
6
CHEMICO
Venturi
12/71
80
99.2
30
9
28
Unit 1
814,000 ACFM @
340°F
Unit 2
814,000 ACFM
-------
Lime Addition
Polymer Addition
Gas Flow to Stack
Outlet Flow
Control, Damper
' Steom Inlgt
Reheater Finned
Tube Section
Reheater Plain
Tube Section
Condensate
Tank
Make-up Water
Pumps ,
Wash Nozzles
A A
emister
ecirc. Nozzles
Neutralization Flocculation
Isolation
Damper
Exhaust Gases
From Boiler
Glorification
.Recirc. Water
Underflow to
Ash Ponds
Prespturator
Section
Scrubber
Booster
Fan
Recirculation
Pump
Slurry
Pumps
Slowdown Slurry
Figure 2. Simplified Cherokee Station Scrubber Flow Diagram
Green, 1975.
-------
Flue Gas to Reheater
Lease Equipment
X"~ Demister Wosh
Scrubber Bolls
(3) Sections
Presoturator'
Section i
Solid Waste Disposal Pond
Make-up Water
Water
Figure 3 Valmont Station Modified Scrubber Flow Diagram
Green, 1975.
-------
00
Flue gas from
air heaters
Mist eliminators
Lime —i
I—Cooling tower blowdown
Clear
Fly ash pond
Figure//•. — Simplified flow diagram for the Dave Johnston fly ash scrubbers
Green, 1975.
-------
Make up water
Quench
spray
scrubber
To stack
Make up water
I
Wet ID fans
(2)
humidification
spray
Mist eliminator
Punch plate
i — ±r^~
ter
(r ,
— • jr-x
1
r
Figure5. - Simplified flow diagram for the particulate
scrubber at the Clay Bosvyell station.
Green, 1975.
69
-------
To stack
Flue gas from
air heaters
Wafer
Spray
Mist eliminators
and reheater
Mist eliminators —'
Fly ash transfer
tank
Liquid transfer
tank
Figure 6. Simplified Flow Diagram for the Four Corners
Fly Ash Scrubbers.
70
-------
Inlet Damper
Clean Gas-
Outlet Damper
Exis
Precif
20%
ting
Jitotor
*" *^
^
fc-J Chemico
Venfuri
Scrubber
v WV
^-Existing Air
Heater
Existing
Stack
4
-i
-Recycle to Scrubber
-Make-up Water
Lime-Hopper
Feeder
Make-up Water
•»»To Ash Pond
Fly Ash Disposal
Sump
Figure 7. Holtwood Station Scrubbing System
General Arrangement
Green, 1975.
-------
in terms of increased maintenance. The particulate scrubbers have proven to be
efficient devices in collecting fly ash. They are not subject to resistivity diffi-
culties, wire breakage, plate warpage and many of the other concerns of precipitators.
On the other hand, particulate scrubbers introduce a whole host of problems never
conceived when operating precipitators. I will not dwell on specific problems
associated with scrubbers as there have been numerous papers and discussions on the
merits of various scrubbers. I believe it is sufficient to note that when consi-
dering the scrubber as a control device, the problems associated with flue gas
desulfurization scrubbers are also basically inherent in the particulate scrubber.
CONTROL OPTION CONCLUSIONS
The use of the wet scrubbers for control of particulate matter has proven to be
successful, particularly in cases where a low sulfur coal is used as the primary
boiler fuel; however, these wet scrubbers have required numerous modifications to
improve on-line availability and also require a high degree of maintenance as com-
pared to other particulate control devices. The wet scrubber has also introduced
the added dimension of water pollution control for the scrubber effluent.
The operation and maintenance cost associated with the wet scrubber is consi-
derably higher than an electrostatic precipitator to achieve the same collection
efficiency. It is for this reason that most utility companies place less consider-
ation on the use of the wet scrubber for particulate control. Many new generating
installations will utilize electrostatic precipitators for particulate control, and
consider using scrubbers only for meeting future S0_ emission requirements. It is
also surprising to note the renewed interest in fabric filters for particulate
control.
New electrostatic precipitator designs for both hot and cold gas precipitators
are now proven to be more successful in collecting high resistivity fly ash than were
earlier designs. It is my belief that for the collection of particulate matter a
precipitator may be the preferred device, particularly when considering operational
cost, water pollution and availability problems associated with the wet scrubbers.
During the past twenty years we have seen tremendous improvements in the
state of the technology. We have by no means found all the answers to the many
problems associated with the control of particulates. Future projects planned by
such organizations as the Electric Power Research Institute and many individual
utility companies will add further insight into the problem of particulate collection,
and, hopefully, we will continue to improve upon these devices. It seems that in
this day and age the most important decision to be made in the construction of any
new power plant is the determination of the air pollution control equipment to be
72
-------
used. I believe that when burning the lower sulfur western fuels, the decision of
the proper control equipment becomes even more difficult to resolve.
73
-------
REFERENCES
Green, G. P. Operating Experience withWet Scrubbers for Particulate Matter
Removal. October, 1975.
Green, G. P. and R. J. Blatnik. Lessons from 15 Years Experience with Electrostatic
Precipitators and Low Sulfur Coal. Presented at The International Energy
Engineering Congress, Chicago, Illinois, November, 1975.
Green, G. P. and W. S. Landers. Operating Experience with Gas Conditioned
Electrostatic Precipitators. Presented at the Joint U.S./U.S.S.R. Symposium-
1974, San Francisco, California, January, 1975.
Sondreal, E. A. and P. H. Tufte. Scrubber Developments in the West. Presented at
the 1975 Lignite Symposium, Grand Forks, North Dakota, May, 1975.
74
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WESTERN COAL USE IN INDUSTRIAL BOILERS
Kenneth L. Maloney, Ph.D.
KVB, Incorporated
Tustin, California
75
-------
This program has been funded at least in part with Federal funds
from the Environmental Protection Agency under Contract No. 68-02-1863.
The content of this publication does not necessarily reflect the views
or policies of the U. S. Environmental Protection Agency, nor does
mention of trade names, commercial products, or organizations imply
endorsement by the U. S. Government.
76
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KVB
P-157
WESTERN COAL USE IN INDUSTRIAL BOILERS
Prepared by
Kenneth L. Maloney, Ph.D.
KVB, Inc.
Tustin, CA 92680
ABSTRACT
Six small and intermediate-sized (1Q,000 to 250,000 Ib/hr steam)
coal-fired boilers in the upper Midwest have each been tested on both
a bituminous eastern coal and a subbituminous western coal.
The purpose of this study was to determine the feasibility of
substituting western subbituminous coal for eastern bituminous coal
as a means of reducing the SOx emissions from this class of boiler and
to demonstrate the feasibility of greatly expanded western coal utiliza-
tion as a means of reducing the use of oil and gas.
The scope of this study was such that the representative boiler
types were tested on both eastern and western coal for a period of time
sufficient to completely characterize their individual emission and
operational characteristics.
ACKNOWLEDGEMENT
The author would like to express his appreciation to Mr- Dave
Lachapelle, EPA, for his continued interest in the use of western coal.
This research was supported under Environmental Protection Agency
under Contract No. 68-02 1863.
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1.0 INTRODUCTION
Faced with the problem of complying with sulfur dioxide control
regulations, electric utilities and industries in the Midwest have been
increasing their use of low-sulfur western coal. The extent to which
Midwestern demand for western coal will continue to increase depends on
a number of factors. Foremost among these are: (1) the evolution of
federal, state, and local sulfur dioxide control regulations, (2) the
growth of coal as a boiler fuel, and (3) the cost of western coal rela-
tive to the costs of alternate fuels and control technologies.
The upper Midwest region (Minnesota, Wisconsin, Iowa, Nebraska,
and Illinois) is presently the only area where low-sulfur western sub-
bituminous coal is cost competitive with midwestern and eastern coals.
Within this region, there is considerable variation with regard to
western coal use versus the traditional eastern supply. This variabi-
lity is due in part to equipment limitations which dictate that a certain
coal be burned.
For this reason it is necessary to determine the operational
compatibility of western coal with existing industrial coal-fired
equipment, if fuel substitution is to be considered a viable sulfur
oxides control strategy.
The purpose of this program, the test results of which are
detailed in this paper, was to assess the effectiveness of the use of
lower sulfur western coals as a means of reducing sulfur oxides emissions
from industrial-sized boilers in the size range 10,000 to 250,000 pounds
of steam per hour. The impact on SOx, NOx, CO, particulates, and
unburned hydrocarbons emissions has been assessed as a consequence of
this fuel conversion.
The scope of the testing program included testing six represen-
tative types of coal-fired industrial boilers for a period of one month
each on eastern and western coal. During this testing period, the
78
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pollutant emissions listed above were measured both in a baseline con-
figuration and in an optimized firing mode. Operational problems of
the unit were characterized for each coal. Potential reductions of
pollutant emissions have been estimated for each unit type and each
coal tested.
2.0 PROPERTIES OF WESTERN SUBBITUMINOUS COALS
A large supply of low sulfur, subbituminous coal exists in the
Powder River region of Wyoming and the Fort Union region of southeast
Montana. This coal is being mined at a rapidly increasing rate. One
mine in Wyoming, for example, increased production from 0.89 million
tons per year in 1973 to 3.3 million tons per year in 1974, a factor of
3.7 in only one year. However, the most impressive statistics are the
reserve capacity of these western coal fields. That same mine in Wyoming
whose production increased so dramatically in 1974 has a reserve capacity
of 18.5 billion tons. This translates to a lifetime of 50 years at
current production rates. The large reserves, coupled with the relative
ease of strip mining, point to a ready supply of coal for fuel if other
constraints are met. One of these constraints is the subject of this
paper.
The compatibility of these western subbituminous coals with
existing industrial boilers could be a hinderance to their wide accep-
tance as a boiler fuel. The compatibility of coal and boiler are deter-
mined both by coal properties and by boiler design. Since the boiler
designs are fixed in existing units, the coal properties are the varia-
bles of interest.
Western coal characteristics are those of a typical subbituminous
coal: an ash-free higher heating value of 8,200 to 10,500 moist Btu/lb,
and a high moisture content of 20% to 30%. The ash content of most of
these coals is less than 10% by weight. The western subbituminous coals
exhibit high volatile to fixed carbon ratios, typically approaching a
value of one. A typical, as-received, analysis of a Montana subbituminous
coal is given in Table I.
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TABLE I
TYPICAL WESTERN COAL CHARACTERISTICS
(Colstrip, Montana)
PROXIMATE ANALYSIS
As Received Dry Basis
% Moisture
% Ash
% Volatile
% Fixed Carbon
24.36
8.86
32.53
34.25
xxxxx
11.71
43.00
45.29
100.00
100.00
FUSION TEMPERATURE OF ASH
Reducing
Initial Deformation 2130°F
Softening (H=W) 2180°F
Softening (H= 1/2 W) 2205°F
Fluid 2240°F
H is Cone Height
W is Cone Width
SILICA VALUE = 0.09
T250 = 62.42°F
ESTIMATED VISCOSITY
at Critical Viscosity
Temperature of = 2380°F
ULTIMATE ANALYSIS
% Weight
As Received Dry Basis
Moisture
Carbon
Hydrogen
Nitrogen
Chlorine
Sulfur
Ash
Oxygen (diff)
24.36
51.19
3.36
0.77
0.01
0.88
8.86
10.57
XX XXX
67.67
4.44
1.02
0.01
1.17
11.71
13.98
100.00
MINERAL ANALYSIS
100.00
Phos. pentoxide,
Silica, SiO
Ferric oxide,
Alumina, Al O..
Titania, Tio
Lime , CaO
Magne s ia , MgO
Sulfur trioxide,
Potassium oxide,
Fe-O_
SO.
K 0
Sodium oxide, Na 0
Unde te rmined
% Weight
Ignited Basis
0.20
40.85
9.33
17.36
0.83
11.50
3.76
15.13
0.48
0.42
0.14
100.00
P-157
80
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The western subbituminous coals are also classed as "free-burning"
coals. In the free-burning coals, the pieces do not fuse together, but
burn separately or, after fusion, the mass breaks up quickly into frag-
ments. This characteristic causes problems in certain types of stokers
where there is inadequate control of the undergrate air distribution.
Some specific western coal problems for five types of combustion
devices, and the property that causes them, are presented in Table II.
The high moisture content of the western coals causes the great-
est combustion difficulty in industrial-sized equipment. In most units
with superheaters, it leads to high steam superheat temperatures. It
also causes flame stability problems in pulverized coal combustion and
ignition problems in stoker-fired units. It is clear that in order to
recover the lost steam capacity, some pre-drying of western coals will
be necessary, for firing in units designed for eastern coal.
The second major problem with western coal is the size distribu-
tion of the delivered coal. Most western coals do not travel or weather
well. The coal has a tendency to break into fine sizes while in transit.
Therefore, even if the coal has been sized before shipment, the as-
received coal will exhibit a shift in size distribution toward the
smaller sizes. This shift becomes more severe with longer transit and/
or storage periods. The effect of this coal property on stoker unit
performance is discussed below.
This paper is divided into a discussion of pulverized coal fir-
ing and stoker firing of both eastern and western coal. A general
overview of boiler performance is presented in Table III. Here, the
units tested are rated in terms of emissions, efficiency, and overall
ease of operation. Comments are presented where appropriate. The type
and source of the coals tested are also given for each boiler.
3.0 PULVERIZED COAL COMBUSTION
The pulverized coal-fired boiler tested was Unit No. 3 at Dairy-
land Power Cooperative at Alma, Wisconsin generating station. This four
burner face-fired unit manufactured by Riley Stoker Corporation is rated
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TABLE II
SOME SPECIFIC WESTERN COAL PROBLEMS
J
g
u
ffi
<
PROPERTIES
Low Heating
Value
High Moisture
Low Sulfur
Dustiness/
Fineness
-Friability
Free Burning
Slaking/
Weathering
Na and Ca
VIBRATING
GRATE STOKER
o Ignition
problems
o Uneven
fuel bed
o Uncovered
grate
<
SPREADER
STOKER
o Flame
stability
o Reduced
capacity
o High
superheat
temperature
o Poor grate
coverage
o Overheating
of grates
o Ash pit
fires
o Increased
fouling
:OMBUSTION DEVIC
PULVERIZER
o Reduced
capacity
o High
maintenance
o Flame
stability
o High
superheat
o Poor
grinding
o Feeder
plugging
o Fouling
o Sintering
ES
UNDERFED
STOKER
o Reduced
capacity
o Poor
ignition
o Nonuniform
bed
o Uneven fuel
bed with
uncovered
grate areas
o Ash pit
fires
o Uncovered
grate
OVERFED
STOKER
o Reduced
capacity
o Poor
ignition
o Carbon
carryover
o Undergrate
air res-
triction
o Ash pit
fires
O Uncovered
grate
TRANSPORTATION
o Lower Btu' s/
ton-mile
o Freezing
o Coal loss
o Fugitive
dust
o Coal size
degradation
STORAGE
o Increased
equipment
loading
o Fires
o Fires
o- Fugitive
dust
o Coal size
degradation
ASH COLL/
DISPOSAL
o More ash/Btu
o More ash/Btu
o Inefficient
orecipitators
o Caking
00
K>
P-157
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TABLE
DESIGN TYPE OF UNITS TESTED AND
OVERALL PERFORMANCE ON EASTERN AND WESTERN COALS
TYPE OF UNIT TESTED
OVERALL PERFORMANCE RATING
PER COAL
Fair
Unacceptable
PULVERIZED COAL (Riley)
o 230,000 Ib/hr steam
o Four Burner Face-Fire'd
o Two Ball Tube Mill Pulverizers
o UOP ESP
Reduced maximum capacity
Western Kentucky
(River King)
Eastern
and
Western
Montana
Sarpy Creek
(Westmoreland)
VIBRATING GRATE STOKER (Detroit)
o Water-cooled Grate
o 45,000 Ib/hr Steam
o FD Fan/Natural ID
o Cinder Trap Partic. Removal
Improved coal sizing
would improve performance
Western Kentucky
(Vogue)
Wyoming
(Big Horn)
TRAVELING GRATE STOKER (LaClede)
o 60,000 Ib/hr Steam
o FD Fan/Natural ID
o No Particulate Controls
Severely affected by coal
size
Western Kentucky
(Vogue)
Eastern
Western
Wyoming
(Big Horn)
UNDERFED STOKER (Westinghouse)
c- Multiple Retort
o 100,000 Ib/hr Steam
o Cyclone Dust Collector
Specially sized western
coal was used for the
test, however, the unit
would not respond to load
demand. Modifications
are necessary to under-
grate air system in order
to burn western coal.
Kentucky and
Illinois
Wyoming
(Big Horn)
SPREADER STOKER (Detroit)
o 160,000 Ib/hr Steam
o Traveling Grate
o Multiclone Cyclone
o FD and ID Fans
o Superheat, Economizer,
and Air Heater
Maximum load reduced to
130,000 Ib/hr steam on
western coal due to high
superheat temperatures.
Large carbon losses on
eastern coal—smoking.
Southern
Illinois
Montana
(Colstrip)
SPREADER STOKER (Westinghouse)
o 100,000 Ib/hr Steam
o Traveling Grate
o FD and ID Fans
o Superheat, Economizer
Able to maintain full
load on western coal
Kentucky
(Vogue)
Western
and
Eastern
Montana
(Colstrip)
P-157
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at 230,000 Ib/hr steam flow. The coal is pulverized with two ball tube
mills, one mill for the upper two burners and one for the lower two
burners. The unit is equipped with a steam spray attemperator. Fly
ash collection is accomplished with a UOP designed cold side electro-
static precipitator (ESP).
The two fuels used during the testing were:
Western Kentucky Coal
o 4% sulfur
o 16% ash
o 10,500 Btu/lb
o 18% volatiles
and
Montana Coal
o 0.77% sulfur
o 12% ash
o 8,400 Btu/lb
o 37% volatiles.
3.1 Boiler Performance - Alma Unit No. 3
The boiler performed well on both coals, although the unit was
limited in maximum load due to excessive superheat steam temperature.
The steam attemperation system was not adequate to reduce the tempera-
ture to the desired 900°F level at loads above 174,000 Ib/hr steam on
western coal. This compares to a maximum load of 204,000 Ib/hr steam
on eastern coal. The boiler is design rated at 230,000 Ib/hr steam,
however this load is no longer achieved.
The primary factor causing the excessive steam temperature is
the high moisture content of the coal. This water reduces the flame
temperature which in turn reduces the radiant heat flux to the water
walls resulting in lower steam generation. This lower heat transfer (a
function of temperature to the fourth power) removes less heat in the
radiant section, however, the gas still contains a large enthalpy which
then acts on a decreased amount of steam in the convective section
resulting in increased steam temperatures. The water in the fuel also
results in greater gas flows which increase heat transfer rates in the
convective pass.
The excessive steam temperature problem is a function of boiler
design. For example, a boiler designed for western coal might not be
able to make design steam temperature on eastern coal.
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Increased steam attemperation would result in full capacity
operation on western coal.
3-2 Pulverizing Mill Performance. Eastern coking coals, when exposed
to furnace temperatures, will swell and form lightweight, porous coke par-
ticles. These may float out of the furnace before they are completely
burned. As a result, carbon loss will be high unless pulverization is
very fine. Free-burning (western) coals, on the other hand, do not
require the same degree of fineness because the swelling characteristic
is absent.
High-volatile (western) coals ignite more readily than those
with a low volatile content. Therefore, they do not require the same
degree of fine pulverization. With the exception of anthracite, however,
the low-volatile coals are softer, and may be said to have a higher grind-
ability. As a result, mill capacity is greater at increased fineness
than with high-volatile coals.
Table IV shows the screen analyses and the loads of the coal
burned in tests 9, 16, 57, 63, 75, and 78. Tests 9 and 16 were on
eastern coal. Test 16 was with one mill, a total of two, operating so
the load in the mill was the same as it would have been with both mills
operating at 104,000 Ib/hr steam. The screen analyses of tests 16 and
78 may then be compared. It is seen that the western coal did not grind
quite as well as the eastern coal. However in the opinion of Reference 1,
free-burning coals need not be ground as fine as coking coals, and this
was not thought to be a severe problem. An equally important factor in
mill grinding capacity is moisture. From Reference 1, frequently too
much emphasis is placed on grindability, while other factors such as
moisture, which also affects mill capacity, are almost entirely overlooked.
The capacity of a pulverizer is not directly proportional to the grinda-
bility of a coal. Correction must be made for variation in fineness and
moisture content.
1. Combustion Engineering, O. de Lorenzi, editor, Combustion Engineering
Co., Inc. 1947, p. 7, 8.
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TABLE IV
SCREEN ANALYSES OF PULVERIZED COAL
Test No.
Load , thousands Ib/hr
steam
- 80 mesh, %
- 80 +100 mesh, %
-100 +140 mesh, %
-140 +200 mesh, %
-200 mesh, %
Moisture, %
EASTERN
9 16
125 52
0.65 1-00
0.75 1.00
2.75 3.35
6.95 8.20
88.90 86.45
5.15 2.20
WESTERN
57
170
2.90
2.30
20.65
34.60
39.55
22.05
63
131
1.40
1.30
4.05
8.30
84.95
12.75
63
94
0.65
2.32
8.91
32.26
55.66
17.94
75
160
8.75
7.70
21.65
16.75
45.15
19.33
78
110
1.53
2.23
6.60
17.46
72.18
17.75
Without quantitative analysis, it can be seen in Table IV that
the moisture content of test 78 is a factor of 8 times higher than test
16. The grindability of the other tests tend to follow the moisture
content. Test 57 with the highest moisture content exhibited the poorest
grindability, followed by tests 75, 73, 78, and 63 in order of increasing
grindability.
The poorly pulverized coal burns more slowly resulting in lowered
heat transfer in the near-flame region (radiant section) and increased heat
transfer to the convective section. At high loads (tests 57 and 75), the
poor grind probably contributed to the excessive superheat steam tempera-
ture problem.
3.3
Emissions from Alma Unit No. 3
A coal performance comparison for Alma Unit No. 3 is presented in
Table V. In this table, western coal, test 66, is compared to the nearly
identical eastern coal, test 9.
Significant differences in coal performance are noted for:
o SOx emissions
o NO emissions
o Carbon carryover
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TABLE V
COAL PERFORMANCE COMPARISON, ALMA UNIT NO. 3
Test No.
Load, Klb steam/hr
Excess 02, %
SOx at 3% O2/ ppm
NO, dry at 3% 02, ppm
CO, at 3% O2, ppm
Particulate, lb/106 Btu
ESP Efficiency, %
Carbon Carryover, % by wt
Unburned HC, at 3% O2, ppm
Boiler Efficiency, %
Western
66 ESP Inlet
130
3.4
996
372*
31
6.77
99.6
0.55
25
85
Eastern
9 ESP Inlet
130
3.4
3283
490**
21
7.84
99.6
4.13
31
•—•—
* 6 p-157
0.52 lb/10 Btu
** 6
0.69 lb/10 Btu
In each case the western coal performed better than the eastern coal.
Sulfur oxides emissions were reduced by a factor of 3 by substituting
western coal. At the same time nitric oxide emissions were reduced 24%
and carbon carryover was virtually eliminated. The electrostatic pre-
cipitator's performance was not affected by the fuel switch. It con-
tinued to operate at 99+% efficiency. Carbon monoxide and unburned
hydrocarbon emissions were generally less than 100 ppm each. In the
optimum furnace configuration, these emissions are controlled by excess
air. Below three percent (3%) excess O,, in the flue gas, these emiss-
ions became significant. Soot formation, resulting in a black stack,
was also a problem below 3% excess O2-
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3.4 Discussion
Figure 1 is a plot of nitric oxide as a function of excess O
in the flue for western coal at four loads. Figure 2 contains the same
type data for the base eastern coal. Both figures show increasing NO
with increasing O at a constant load, however, the absolute magnitude
of NO emissions from western coal is less at any given load and 0 .
Most of the NO data on Figure 1 fall below the EPA limit for new coal-
fired units of 0.7 Ib of NOx as NO per million Btu (about 500 ppm).
Attempts to reduce the NO emissions of the eastern coal to these same
(less than 500 ppm) levels resulted in high CO emissions.
Included in the factors that influence NO emissions are:
o Flame temperature
o Fuel nitrogen
o Excess oxygen
In order to control CO emissions from the eastern coal it was necessary
to operate at higher O levels; this led to higher NO emissions. For
western coal firing, it has been shown that the furnace can operate at
lower excess O , thus lower NO. Western coal also contains less bound
fuel nitrogen than eastern coal. This fuel nitrogen can be as little
as half the amount found in typical eastern coals. Assuming complete
conversion of this fuel nitrogen to NO, it can be seen that western coal
has a distinct advantage over eastern coal.
The third factor affecting NO emissions is flame temperature.
The high moisture content of western coal causes the temperature of the
western coal flame to be lower than the eastern coal flame. This lower
flame temperature lowers both the fixation of molecular nitrogen in the
combustion air as well as reducing the fuel nitrogen conversion to NO.
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00
800
700
e eoo
•M
n)
500
Q
*t
i 400
300
200
(High CO)
O 90 K #/Hr
Q 60 K #/Hr
A 130 K #/Hr
• 170 K #/Hr
10
12
°
Figure 1. Nitric Oxide vs. Oxygen - Alma #3, Western Coal.
-------
VD
O
700
600
500
dP
n
O
300
200
49
BOOS 41
(Smoke)
D21
17
16
D
17119
O 200 K #/hr
& 130 K #/hr
Q 60 K #/hr
I
I
10
12
20Q
14
Figure 2. Nitric Oxide vs. Oxygen - Alma #3, Eastern coal.
P-157
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Sulfur oxides emissions are largely a function of sulfur in the
fuel. There has, however, been some work that indicates that coal ash
composition may affect the amount of sulfur oxides emitted (Ref. 2).
The comparison of the eastern and western coals on this unit (see Table
VI) show the benefit of fuel substitution in the control of SOx emissions.
4.0 STOKER-FIRED BOILERS
Coal firing of industrial boilers can be separated into two broad
classes—suspension firing and grate firing.
Suspension firing is normally applied in larger sized units,
however, units as small as 35,000 Ib/hr steam have been built for pul-
verized coal firing. Current economics would indicate a break-even point
in the 200,000 to 250,000 Ib/hr-steam flow range. Suspension firing
includes both pulverized coal firing (70% through a 200 mesh screen) and
cyclone firing (crushed to 1/4" with about 10% through a 200 mesh screen).
Grate firing comprises three general stoker types:
o Underfed
o Overfed
o Spreader
Within these three types, there are a number of variations in feed methods
and grate design. Stoker-fired boilers have been built covering the entire
capacity range of this study, 10,000 to 250,000 Ib/hr steam. The present
stoker-fired boiler population represents a highly individualized array
of equipment.
Table III, presented previously, lists stoker types tested in this
study. From this assortment of units, the emissions and operating charac-
teristics of western coal firing have been determined.
2. Grouhoud, G. H., Tufte, P. H., and Selle, S. J., "Some Studies on
Stack Emissions from Lignite-Fired Power Plants," presented at the
1973 Lignite Symposium, Grand Forks, ND, May 9-10, 1973.
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4.1 Operational Characteristics of Western Coals in Stokers
A chronological display of the development of the various stoker
types would appear as follows:
Chronology of Stokers
1850 1900 1950
Single Retort Underfed Stokers 1
Multiple Retort
Underfed Stokers
Chain or Traveling .
Grate Stokers
I Spreader Stokers-
Vibrat-^
ing Grate
Two properties of western subbituminous coals result in operational
problems for stoker-fired units. They are:
o Coal weathering - resulting in size reduction
o "Free burning" characteristic - resulting in an
uncovered grate
Many older underfed and traveling grate stokers were manufactured with
insufficient control of the undergrate air to use western coal as a fuel.
The problem manifests itself when a dark spot of unburned coal develops
on the grate. This patch of coal can grow into a large clinker if special
measures are not taken to remove it. The "black patches" occur because
there is insufficient local air pressure under the patch to maintain
vigorous burning. The loss of local air pressure occurs because some
other portion of the grate, in the same plenum control area, has become
thin or bare and allows the combustion air to pass through easily- These
units were designed for an eastern coal that formed some coke while burn-
ing and in turn maintained even coverage of the grate. The free-burning
western coals, on the other hand, tend to form a fine powdery ash which
either blows off or falls through the grate, leaving it bare. This prob-
lem is compounded by the serious size reduction that occurs while in
transit. The small coal particles burn more rapidly when there is avail-
able air, however, when there is insufficient air they tend to plug the
grate and fuel bed openings and form dark patches which turn into clinkers.
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The older underfed and overfed stokers designed for eastern coal
will have to have modifications to the undergrate air chamber to allow
finer control of the air distribution, if western coals are to be used.
Spreader stokers are affected by the same coal properties but
to a lesser extent since approximately half of the combustion takes place
in suspension. This suspension burning reduces the number of "fines"
that reach the grate. However, the fines in the coal tend to burn close
to the spreader, sometimes flashing back into the feeder opening. This
flash back mode can be dangerous since there is the possibility of a
fire in the coal feeder. Coke and slag also have a tendency to build up
on the spill plates and rotor blades if the flash back is allowed to
persist. This problem can be alleviated somewhat by rotor speed and
spill plate adjustments.
The western coal performed well in the spreader stoker units.
In some units designed for eastern coal, the maximum attainable load was
about 80% while on western coal. This was due to insufficient induced
draft fan capacity and as in pulverized units, high superheat steam
temperatures. Removing the major part of the moisture from the coal
prior to combustion would alleviate both of these problems.
4.2 Emissions from Stoker-Fired Units
4.2.1 Sulfur Oxides. The emissions of sulfur oxides from stokers is
as expected, governed by the sulfur in the fuel. These emissions are
independent of load and excess air in the flue gas. Table VI contains
the results of a SOx emission comparison for all the coals tested in
this study. The overall average SOx emissions from actual operating
industrial type boilers decreases from 5.19 lb/10 Btu on eastern coal
to 2.34 lb/10 Btu ,on western coal, or 55%. This is to be compared to
the reduction as estimated from the fuel analysis from 6.43 lb/10 Btu
on eastern to 1.76 lb/10 Btu on western coal, or 73%. The sulfur con-
tent of the fuel was calculated from published values and not from
analyses of actual fuels burned. When the results of the fuel analyses
are obtained, the calculation will be corrected to those values.
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TABLE VI
SOx EMISSION COMPARISON FOR EASTERN AND WESTERN COALS
WESTERN COAL
Montana (Sarpy Creek)
Wyoming (Bighorn)
Montana (Col strip)
Montana (Colstrip)
Wyoming (Bighorn)
Overall Average
EASTERN COAL
Kentucky (River King)
Kentucky (Vogue, Seam II)
So. Illinois
So. Illinois (Stonefort)
W. Kentucky (Vogue)
Overall Average
Average SOx Reduction (Flue
Average SOx Reduction (Est.
Avg Fuel*
Sulfur, %
0.80
0.60
0.85
0.88
0.60
0.746
3.96
3.22
1.9
3.1
3.22
3.375
Emissions) =
*
from Fuel) =
Avg SO2 Emissions
lb/106 Btu
1.86
1.18
1.8
2.56
1.28
2.34
6.46
4.73
—
4.92
4.66
4.19
2.85 lb/106 Btu (55%)
4.67 lb/106 Btu (73%)
Based on published values, not on actual analyses of fuels
used in the tests.
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4-2-2 Nitric Oxides. Nitric oxide emissions from stokers exhibited a
similar dependence on excess C>2 in the flue gas as the pulverized coal
firing. At constant load, nitric oxide emissions increased with increas-
ing excess C>2, as well as increasing slightly with increasing load.
However, the slope of the NO vs. O curve is less for stoker-
fired units than for the higher intensity combustion devices. Figure 3
shows an interesting NO vs. O2 result for a water-cooled vibrograte
stoker. The western coal (Wyoming Bighorn) has a slope of 12- (ppm NO/
% 0 ) compared to the eastern coal (Kentucky Vogue) which has a slope
of 35 (ppm NO/% O2). Figure 4 gives the NO vs. O plot for the same
two coals on an overfed traveling grate stoker without a water-cooled
grate. On this unit, both coals exhibit the same NO vs. O dependence.
In fact, of the boilers tested, the water-cooled grate was the only unit
having different NO vs. O2 slopes for the two coals tested. It is spec-
ulative as to whether the additional cooling of the grate affects the
conversion of fuel bound nitrogen to NO.
Stokers have overall lower NO emissions than pulverized coal
units since the stokers operate in a "staged combustion" configuration.
The stokers that have little or no suspension burning such as underfed
and overfed stokers have a greater degree of staging than do the spreader
stokers. In the stoker, the fuel devolatizes in the fuel bed under
reducing conditions, then is mixed with the combustion air above the bed.
Mixing is provided by overfire air jets or by front or rear arches in
the furnace. Clinkering in the fuel bed establishes a limit to the
degree of staging that can be achieved on stokers. Figure 5 shows these
limits for a 100,000 Ib/hr steam spreader stoker.
Carbon Monoxide and Unburned Hydrocarbons. Carbon monoxide (CO)
and unburned hydrocarbons (UHC) emissions from stokers as with all com-
bustion systems, can be controlled by providing adequate excess air and
proper mixing to insure complete combustion. High excess air conditions
can cause CO and UHC emissions as well as too low excess air. Figure 6
gives the results of CO emission measurements on a 100 Klb/hr steam
spreader stoker as a function of excess air for both eastern and western
95
-------
250
225
O
to
to
Hi
o
X
w
<*>
200
175
a
i
150
125
O
oL
0
Eastern Coal
(Kentucky Vogue)
14-15 Klb/hr
25 Klb/hr
40 Klb/hr
Western Coal
(Wyoming Bighorn)
25 Klb/hr
15-18 Klb/hr
I I
10
11
12
Excess O
2'
Figure 3.
Water-C^Ued Vibrograte Stoker, 45,000 Ib/hr Steam. Comparison of
Western and Eastern Coal Nitric Oxide Emissions (University of
Wisconsin-Stout).
P-157
-------
350
300 -
(A
8 25°
X
w
I
O
z
200
150
V
I
I
19.
12
Western Coal (Wyoming Bighorn)
Load
O 30-35 klb/hr steam
• 15-18 klb/hr steam
Eastern Coal (Kentucky Vogue)
Load
A 23-27 klb/hr steam
^ 15-18 klb/hr steam
10 12
Excess Oxygen, %
14
16
18
Figure 4.
Overfed Traveling Grate Stoker, 60,000 Klb/hr Steam (University of
Wisconsin-Eau Claire).
P-157
-------
12.0
11.0
10.0
1
u
4J
9.0
w
-------
2000
vo
vo
<*>
n
•P
R]
8
1000
-------
coal. At high load (90 Klb/hr steam), CO emissions increase with decreas-
ing excess air; however, at low and intermediate loads, a point is reached
where increasing excess air results in rapidly increasing CO emissions.
This behavior was observed for both coals. At low excess air, CO results
from inadequate mixing of fuel and air. At high excess air settings, the
fuel bed is thin even to the extent of some uncovered grate area which is
thought to lead to local quenching of the flame by the combustion air and
incomplete oxidation of CO to CO . The western coal can be fired at 2%
lower excess O at high load while producing comparable levels of CO
emissions. This means higher unit efficiency because of lower dry gas
and combustible losses.
Table VII contains unburned hydrocarbon emission data for both an
eastern and a western coal on the same spreader stoker unit described
above. Unburned hydrocarbon emissions were higher at low load and high
excess air than at high load and normal excess air, thus following the
same trends as the CO emissions.
TABLE VII
A COMPARISON OF HC EMISSIONS
FROM EASTERN AND WESTERN COALS
Eastern Coal
Load
(klb/hr)
40
60
90
90
o2(%)
15.
12.7
9.7
8.7
HC
(cor ppm)
114
54
48
44
Western
o2(%)
13.8
11.3
—
8.8
Coal
HC
(cor ppm)
125
18
—
44
No appreciable differences in UHC emissions were noted between
eastern and western coals.
Carbon monoxide emissions are a much more sensitive measure of
incomplete combustion than are unburned hydrocarbons. A comparison of
CO emissions and carbon carryover can be made. Figure 8 is a plot of
% carbon in the outlet flyash of a 160 Klb/steam spreader stoker firing
western (Montana) coal. This unit exhibited rather high carbon losses
100
-------
30
25
20
I
rfl
15
14
10
I
70
80
90 100
Unit Load, Klb/hr
110
120
130
Figure 8. Percent Carbon in Outlet Flyash, 160 Klb/hr Steam
Spreader Stoker (Willmar Unit 3), Western Coal.
P-157
101
-------
which increased with unit load. The carbon losses on eastern coal were
even larger than for the western coal. However, the point to be made
here is that by measuring the carbon monoxide emissions, an indication
of the other combustible losses can be gained. The CO emissions for the
same tests are shown in Figure 8.
4.2.4 Particulates. For a given ash content in the coal, the quantity
of particulate matter in the flue gas from stoker-fired boilers depends
primarily upon the amount of burning that takes place in suspension or
on the grate. Table VIII gives an average flue gas particulate loading
from both eastern and western coals as measured before the control device
for the stoker types tested.
TABLE VIII
AVERAGE FLUE GAS PARTICULATE
LOADING FOR EASTERN AND WESTERN COALS
Stoker Type
Spreader
Traveling Grate
Traveling Grate
Vibrograte
Klb/hr
Capacity
100
60
45
Average
Particulate Loading, lb/10 Btu
Eastern Coal Western Coal
1.643
0.783
0.649
1.435
0.428
0.488
Comments
13% lower
than eastern
45% lower
than eastern
25% lower
than eastern
Spreader stokers with the greater suspension burning have from two
to three times the particulate loading of the traveling grate and the vibro-
grate stokers. The western coal test results showed a lower particulate
loading than the eastern coal.
102
-------
1800
1600
80
90
100 110
Unit Load, Klb/hr
120
130
Figure 8. Carbon Monoxide Emissions versus Load (Willmar Unit 3),
Western Coal, 160 Klb/hr Steam Spreader Stoker.
P-157
103
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5.0 CONCLUSIONS
This study has shown that western subbituminous coals can be
substituted for eastern bituminous coals as an industrial boiler fuel.
The western coals are compatible with industrial coal-fired units of
current design. Two unit types of older design (underfed and traveling
grate stokers) were found to have difficulty burning western coal. Some
cases have been noted where the maximum load capacity of the boiler had
to be limited. This problem can be eliminated by predrying the coal.
Western subbituminous coals were found to be superior to eastern
coals in terms of SOx, NOx, particulate, and unburned hydrocarbon emiss-
ions. The western coals could be fired at lower excess air and exhibited
substantially lower combustible losses than eastern coals.
The size of delivered western coal proved to be a problem in
most of the stoker-fired units tested. The coal generally had too large
a percentage of fine coal which resulted from the poor weathering charac-
teristics of western coals.
Stoker performance on western coal could be improved if the coal
were sized local to the point of use so that delivery distances could be
reduced to about 200 miles.
This study has defined the operational parameters that must be
followed in order to successfully burn western coal in industrial-sized
stokers and pulverized coal units. Excess O and carbon monoxide moni-
tors for combustion control would improve overall industrial boiler per-
formance on both eastern or western coal. These controls are necessary
since many times the margin of success can be as small as +_ 0.5% excess
O in the flue. For the most part, present instrumentation does not
provide sufficient precision in combustion control. Operator training
and education must go hand-in-hand with improved controls for successful
western coal firing.
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FACTORS AFFECTING ESP PERFORMANCE
ON WESTERN COALS AND EXPERIENCE
WITH NORTH DAKOTA LIGNITES
Stanley J. Selle
and
Lindsey L. Hess
Grand Forks Energy Research Center
Energy Research and Devleopment Administration
Grand Forks, North Dakota
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FACTORS AFFECTING ESP PERFORMANCE ON WESTERN COALS
AND EXPERIENCE WITH NORTH DAKOTA LIGNITES
by
Stanley J. Selle
Mechanical Engineer
and
Lindsay L. Hess
Research Physicist
Grand Forks Energy Research Center
Energy Research and Development Administration
Grand Forks, North Dakota
ABSTRACT
This paper attempts to delineate some of the special characteristics of
Western coals that affect the performance of powerplant ESPs. The absence
of significant quantities of SOg to act as a natural conditioning agent in
the flue gas from the low-sulfur Western coals is somewhat offset by their
generally higher moisture contents. The alkali constituents, major components
of fly ashes from Western coals as contrasted with Eastern or Australian
coals, act in various ways to influence ESP performance.
Considerable research into the effect of chemical components on fly ash
resistivity has shown sodium to play a significant role in determining resis-
tivity and, hence ESP performance. The ramifications of sodium for ESPs,
both as a naturally occurring resistivity reducer and as a possible conditioning
agent, are considered.
The Grand Forks Energy Research Center of the Energy Research and
Development Administration is designing and constructing a 75-lb/hr pulverized
coal-fired furnace and an associated pilot-scale ESP. The objective of the
project will be to burn small quantities of a given coal to generate a repre-
sentative fly ash and flue gas for comparative studies of resistivity and
pilot-scale ESP performance. The design and development of the equipment,
and its possible usefulness as a design and research tool, are discussed.
A compilation and discussion of ESP design and operating experiences
on commercial ESP units at plants burning North Dakota lignites is presented.
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FACTORS AFFECTING ESP PERFORMANCE ON WESTERN COALS
AND EXPERIENCE WITH NORTH DAKOA LIGNITES
INTRODUCTION
With the increased use of the low-sulfur Western coals for power produc-
tion has come the realization that previously reliable ESP design parameters
do not work well on such coals. The success of ESPs installed on powerplants
burning these coals has been mixed, resulting in a wide range of bidder
design specifications for a given application. The need for adequate criteria
for ESP design on such coals has spurred extensive research toward that end,
with some success. This paper will attempt to review the characteristics of
Western coals that have been shown to affect ESP performance and discuss their
possible usefulness as design indicators. Experience with ESPs on North
Dakota lignites will also be described.
PROPERTIES OF WESTERN COALS AFFECTING ESP-PERFORMANCE
Coal mined in the Western U.S. generally tends to differ from Eastern
coals by having lower sulfur content, higher moisture content, and ash containing
a significantly higher proportion of the major alkali and alkaline-earth
elements (Na, Mg, and Ca). The coal constituents which are generally accepted
as acting as "natural" conditioning agents for ESP operation are moisture and
sulfur. The ash chemical analysis is mainly reflected in the resistivity of
the collected fly ash. The complicated nature of the ESP/flue gas/fly ash
system certainly allows for interactions between effects.
Sulfur
The low sulfur content of most Western coals precludes any significant
conditioning effect due to SOg in the flue gas. Based on 51 determinations of
flue gas 803 in four plants burning lignite, only traces of 803 could be
detected.i/ It appears that for coals with significant quantities of alkali
constituents, any 803 that is formed reacts quickly with the alkali in the fly
ash to produce sulfaies. These results indicate a possible area of concern
for those considering 803 conditioning for fly ashes containing large quan-
tities of calcium. Dismukes reported that a greater increment of 803 addition
was required to lower measured resistivity to about lO-1-1-1 ohm-cm for a highly
basic ash than for a neutral ash.—' This is most likely due to the formation
of calcium sulfate, a high resistivity material.
Moisture
The moisture content of Western coals ranges from around 5% for bitu-
minous coals to over 35$ for lignites. A great deal of interest .in the
effect of coal moisture content on ESP performance has been generated by the
excellent operating experience with ESPs on North Dakota lignites. All ESPs
constructed on lignite-burning powerplants to date have met or exceeded
design efficiencies. A fairly wide range of fly ash chemical analysis and
measured resistivity is represented, with little apparent operational difficulty
107
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experienced. The extent to which the high moisture content is responsible
presents an interesting area for study. Aside from the conditioning of the
fly ash to lower resistivity in the low temperature region, flue gas water
vapor probably functions like any electronegative gas to enhance ESP perfor-
mance by allowing a higher operating voltage before sparking becomes a problem.
Flue gas moisture could also enhance the effectiveness of a given level of SO-^
addition for ESP conditioning.
The results of a series of three preliminary pilot-scale ESP tests run
recently at the Grand Forks Energy Research Center (GFERC) are shown in
Figure 1. The tests were run by feeding a high resistivity fly ash into the
flue gas at a point upstream of the pilot-scale ESP. The flue gas was generated
by the combustion of a Western bituminous coal, with the base moisture level
taken as that from the moisture in the coal. Water was injected into the flue
gas to provide higher increments of flue gas moisture corresponding to those
from subbituminous and lignite coals. The flue gas SC>2 was 270 ppm and the
ESP operating temperature 300° F. Due to problems with in-situ resistivity
measurements, no data was obtained. The resistivity of this fly ash normally
falls in the 1012 to lO1^ ohm-cm range.
As can be seen from Figure 1, the increase in flue gas moisture content
from k to 8 pet by volume resulted in an increase in effective migration
velocity, w ,from 0.75 to 0.96 fps. These levels correspond to those for
bituminous and subbituminous coals, respectively. The increase from 8 to 13
pet flue gas moisture increased ui to 1.06 fps, with the high moisture repre-
senting that from a lignite. The effectiveness of an incremental increase in
flue gas moisture content would appear to be greater in going from a bitumi-
nous to a subbituminous coal than from subbituminous to lignite. The effective-
ness of the flue gas moisture change at the lower level is corroborated by
results reported by Busby and Darby.3/ They reported the most effective 1^0
conditioning with add rates in the range of 0.5-1.0 Ib H^O/IOOO ft3 gas. The
levels tested in the GFERC pilot plant were 1.3, 2.7, and k.3 Ib H20/1000 ft3
gas. The coal studied by Busby and Darby was a New South Wales coal which
appears to range in moisture contents from 1 to 7.5 pet H20 in the coal,
similar to a Western bituminous coal.—/
Figure 1 also shows the relationship observed between the maximum pilot
ESP operating voltage that could be maintained and the flue gas moisture con-
tent. A similar set of points was determined to those for o> versus flue gas
moisture content. This would seem to indicate that a significant portion of
the improvement could be due to the electronegativity of water vapor. Voltage/
current curves generated before fly ash was injected verified that the maximum
applied voltage possible increased as the flue gas moisture content increased.
Ash Composition
The fly ash chemical composition is the third major factor influencing
ESP performance on Western coals. Chemical composition affects are reflected
mainly in fly ash resistivities, although reentrainment, hopper ash removal
and conditioning effectiveness can also be related to composition. Work by
BickelhauptZiH./ the authors ,1^J and others have related measured fly ash
108
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resistivity to chemical composition. The sodium content of the fly ash is
generally accepted as the best indicator of resistivity, with resistivity
decreasing with increasing sodium as illustrated in Figure 2.—/ Figure 3
shows a predictor for fly ash resistivity as a function of fly ash chemical
analysis generated by computer regression analysis of data obtained on 35 fly
ashes from Western coals.—'
The conduction mechanism for fly ash is currently described in terms of
sodium (or Li or K) ion migration through the outer or surface layers of the
.glassy fly ash particles. The development of this picture of charge transfer
through the fly ash layer by Bickelhaupt2i§/ provides an excellent means of
understanding conduction for most fly ashes. However, experience with high-
sodium fly ashes and with conditioning by admixture of pure sodium compounds
by means of injection at low temperatures indicate that very fine particles of
sodium compounds serve to enhance conduction without entering into the glassy
phase. Resistivity measurements on such materials indicate that the effect
holds up at temperatures up to 800° F as illustrated by the data in Figure h.
This effect of sodium may well be twofold, with the sodium compounds enhancing
surface conditioning, and also providing lower resistivity conduction paths
through the collected fly ash layer via the very fine sodium materials.
In high-sodium coals, a significant proportion of the sodium upon vola-
tilzing in the flame does not interact with the glassforming elements as a
flux. Rather, it reacts with 803 to form very fine NagSOl). particles.£/ This
formation of very fine particles is partially verified by experience with ESPs
on North Dakota lignites, where downstream hopper catches are usually signi-
ficantly enriched in sodium content, as illustrated in Table 1.
CONDITIONING BY ADDITION OF SODIUM MATERIALS
The only extensive study of sodium conditioning made to date on a full-
scale boiler involved injecting sodium into the burner with the coal. White
reported on tests utilizing this technique with Na2C03, Na2SOij, and NaCl.—/
Each of the three proved to be effective as conditioners, with economics and
material availability the major factors in determining feasibility. This
method has two drawbacks from an operational standpoint. First there is the
potential for increased ash fouling in the superheater region due to the
fluxing action of sodium. The problem is well described by the work of
Tufte, et.al., 9»11/ and could become especially serious where coals with high
ash contents are burned. Secondly, a significant portion of the injected
sodium may not reach the ESP due to deposition upstream in the heat transfer
areas, ductwork and air heater, with the extent of the loss dependent on
boiler design. This approach does lend itself to powerplant operation, since
mixing the material with the coal during pulverization is reasonably simple.
Tests at GFERC have indicated a conditioning effect for sodium materials
injected in dry powder form ahead of the pilot-scale ESP. Previously reported
work indicated a significant reduction in fly ash resistivity and improvement
in pilot-scale ESP performance when 1 pet or less of previously collected
high sodium fly ash (j.6 pet Na20) was injected into the flue gas stream ahead
of the ESP—' . Subsequent work showed similar results when comparable" quantities
of pulverized Nahcolite, a natural-occurring form of NaHC03, were injected
ahead of the ESP. In both cases the flue gas temperature was ^00° F at the
109
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point of injection. The results are shown in Figures 5 and 6. In both series
of tests, the base high resistivity ash and the additive were premixed and
then injected into the flue gas ahead of the ESP by means of a solids feeder.
All materials were screened to minus 100 mesh. The percentages of additive
shown are the percent by weight of the total ash load injected. Total inlet
particulate concentrations were in the range of 2.5 to 3-grains/acf for the
tests on addition of high sodium ash and 1.8 to 2.0 grains/acf for the Nahcolite
work. All ESP tests were run at nominally 300° F with a flue gas flow rate
of about 190 acfm. Table 2 lists the addition rates in terms of the total
flue gas, illustrating the small quantities of material required as indicated
by these tests. These values are probably lower than the addition rates
required for a full-scale application, since the pilot tests approach an ideal
delivery and mixing situation.
The results of the pilot-scale tests indicate that dry injection of high
sodium materials in a powder form may be useful for conditioning ESPs operating
on high resistivity fly ash. Application of this technique on a large scale
would require the development of an adequate delivery system. The effective-
ness of such a conditioning system would depend on the degree of dispersion of
the material in the gas stream and on the fineness of the added sodium material.
The economics of sodium conditioning will depend on the cost and avail-
ability of the high sodium materials. It would appear that any material with
roughly 10 pet or more sodium (reported as Na20) would be a candidate for use.
High sodium fly ash is a waste material, and there may be other materials
containing significant quantities of sodium which are normally waste products.
An additional incentive for using Nahcolite is its ability to remove signifi-
cant quantities of S02 from the flue gas, as shown in the same series of
tests. Nahcolite levels greater than required for ESP conditioning are neces-
sary for significant S02 removal. It is possible that the two operations
could be combined under the proper conditions.
TEST FURNACE AND PILOT-SCALE ESP
The Grand Forks Energy Research Center has been actively involved with
projects related to the utilization and combustion of Western coals since
about 1963. In 1967, a 75-lb/hr pulverized coal-fired furnace was constructed
to study ash fouling on boiler tubes. This facility played an important role
in determining that sodium content was the chief factor affecting ash fouling
rate."?11/ A flow diagram of this system is shown in Figure 7.
In 1968 a plate-type pilot ESP was retrofited to the furnace to study ESP
performance on fly ash from Western coals. Several aspects of ESP operation
on such coals have been studied, most notably the affect of chemical compo-
sition on fly ash resistivity, and the sodium conditioning work mentioned
earlier.—'
A major problem with the existing furnace has been its inability to
provide a fly ash at the pilot ESP inlet representative of full-scale units in
terms of particle concentration, size distribution and chemical composition.
To overcome this shortcoming and retain the advantage of convenient size, a
new 75-lb/hr furnace is presently under construction at GFERC. A cross-
sectional view of the unit is shown in Figure 8.
110
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The main combustion section of the new furnace (labeled A) has an inside
diameter of two feet and a length of 10 1/2 feet. This section and the
following 12 1/2 feet of transitions and cooling zone (B), will be lined with
a castable insulating refractory. All but the transitions and furnace bottom
will be water cooled by means of an external shell. Six water-jacketed heat
exchange sections (single shell and tube) follow, each six feet in length
(C). The heat exchangers are designed to allow outlet temperatures from 300
to TOO0 F. It is believed that maintaining a velocity greater than 50 fps (—
3 fps in the combustion zone) while minimizing bends and surface area should
minimize fly ash loss in the combustion train. An additional design improve-
ment for maintaining the fly ash in the gas stream is the substitution of
controlled distribution of tertiary combustion air through cross-axial ports
in place of mixing by means of swirling the combustion gases. This will
hopefully eliminate much of the fly ash loss from deposition on the walls of
the combustion zone due to centrifugal force.
A pilot-scale ESP will be attached to the new furnace, as shown in
Figure 8. A tube-type ESP has been selected, primarily because it allows for
the study of a greater number of variables in ranges approaching those for
full-scale units. For example, the maximum flue gas velocity possible at 300° F,
with 2 foot high plates and six inch plate-to-plate spacing as in the
present unit,.-is 3.33 fps with the 200 acfm of flue gas available. An 8-inch
ID tube, however, allows the range to be varied from h to 8 fps by simply
bypassing some of the flue gas. Specific collecting areas will still be
uncharacteristically low, but in a unit of this size that appears to be unavoid-
able. It is felt that a small plate-type ESP is probably no more representative
of full-scale units than a tube-type in this size range. Also, our primary
goal is not to be able to provide dimensional design data, per se, but to
provide a ranking of coals on the pilot-scale ESP relative to other coals on
which experience has been gained previously by the ESP industry and power
companies. The facility will also allow for studies of conditioning methods
and basic research on ESP phenomena, such as fly ash conduction mechanism.
EARLY EXPERIENCE ON WESTERN SUBBITUMINOUS COALS
The need to reevaluate the ESP design criteria normally used for fly ash
from Eastern coals when considering Western coals became apparent early in the
1960's. Many of the first ESPs installed on boilers burning subbituminous
coals failed to meet guaranteed efficiencies. These were low-temperature
units designed with specific collecting areas of 90 to 215 ft2/1000 actual ft3
for removals in the range of 87 to 98 pet. Gas conditioning by injection of
about 20 ppm S02 into the flue gas increased removal efficiencies, but not to
the level of the guarantee for all units. One utility was able to overcome a
high resistivity prpblem by lowering ESP operating temperatures by diluting
the flue.gas with fresh air.i£/
EXPERIENCE ON NORTH DAKOTA LIGNITES
All ESPs now in operation on powerplants burning North Dakota lignites
have met or exceeded design removal efficiencies. None of the operating or
planned units appear to have any unusual degree of overdesign, as can be seen
from the data shown in Table 3. Eight of the eleven installations are retrofit
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units, with four pc-fired, four cyclone-fired, and three spreader-stoker units
shown. These units range from 21 to Wo MW capacity. Flue gas temperatures
range from 288 to Ia8° F, with SCAs from 236 to 375 ft2/1000 acfm. As the
table indicates, all units tested to date have operated at 99 pet efficiencies
or better, except for the Heskett No. 2 unit designed for 97 pet removal.
The powerplants listed typically produce fly ashes with Na20 contents
within the range of 1 to 10 pet. None of these fly ashes represent an extremely
difficult removal problem for an ESP. It has been suggested that the main
reason for the high level of success for ESPs on Worth Dakota lignites is the
high moisture content (roughly 35 pet) of the coal. The variations in flue
gas moisture from a "bituminous to subbituminous to lignite coal would certainly
be of a magnitude sufficient to affect ESP performance as demonstrated by
Busby and Darby.I/ In fact the typical differences in flue gas moisture
"between grades of coal are greater than the 0.5 to 1.0 Ib H20/1000 ft3 flue
gas range required for effective moisture conditioning as shown by their work.
However, based on both laboratory and in-situ resistivity measurements,
sodium still appears to significantly affect the resistivity of lignite fly
ash in spite of the high flue gas moisture. Reports from lignite burning
plants also show measureable effects on ESP performance as a result of changes
in coal sodium levels. In spite of the current degree of success, caution
should be exercised in making generalizations relative to ESP design on any of
the Western coals.
Aside from minor wire breakage and TR set inadequacies, the only major
problem experienced on lignite-burning powerplants so far has been pneumatic
hopper ash removal. For pc- and cyclone-fired units this problem has arisen
where a high-sodium fly ash is being collected, and particularly in the
downstream hoppers where the collected ash is generally enriched in sodium.
The high-sodium fly ashes tend to be sticky and to hold up in the hoppers.
Also, high sodium fly ashes tend to have a low bulk density and to absorb
moisture at the lower temperature of the outlet hopper, especially if heaters
and insulation are not adequate to compensate for the cold North Dakota
winters. The solution to the problem has been to increase the rappers,
heaters, and insulation on the collection hoppers and conveying ducts. The
problem has thus been solved wherever it has occurred so far. An additional
problem has occured in one spreader-stoker installation, where carbon carry-
over has burned in the hoppers to form clinkers, when the hoppers were allowed
to get too hot. Extremely high sodium fly ash has a very low bulk density and
is very light and fluffy in appearance. Collection of such a material may
result in rapping reentrainment problems.
SUMMARY
The three main factors affecting ESP performance on Western coals are the
coal moisture content, coal sulfur content, and fly ash chemical composition.
The relatively high moisture content of some Western coals may compensate to
some degree for the absence of flue gas 803, due to the low sulfur and high
alkali content. Sodium content serves as the best indicator of fly ash resis-
tivity for fly ash from such coals. Any proposed ESP installation for a
powerplant burning a Western coal should be analyzed on an individual basis,
112
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rather than based on generalizations relative to the grade of coal to be
burned. The goal of the ESP program at GFERC is to develope a meaningful set
of criteria and test equipment to serve as a basis for ESP design on such
units. In spite of a high degree of general success with ESPs on powerplants
burning Western coals, there is still research to be done to minimize the
uncertainties involved in sizing and selecting an ESP for a new installation.
In our experience, the cooperation between the ESP industry, the power indus-
try, and the research facilities working on ESP problems has been very good.
Continued cooperation and pursuance of the various approaches being taken
will answer many of the remaining questions.
REFERENCES
1. Gronhovd, G.H., P.H. Tufte, and S.J. Selle. "Some Studies on Stack
Emissions from Lignite-Fired Powerpiants," pp. 83-102: Technology and
use of Lignite—Proceedings, Bureau of Mines-University of North
Dakota Symposium, Grand Forks, N. Dak., May 9-10, 1973. BuMines
1C 8650, 197^, 262 pp.
2. Dismukes, E.B. "Conditioning of Fly Ash With Sulfur Trioxide and
Ammonia." Southern Research Institute, Birmingham, Ala. TVA Research
Agreement TV36921A, EPA Contract No. 68-02-1303, Environmental
Protection Agency, Research Triangle Park, N.C., August 1975. Publi-
cation No. EPA-600/2-75-015 (TVA-F75 PRS-5), pp. 121-2.
3. Busby, H.G. Trevor, and K. Darby. "Efficiency of Electrostatic Pre-
cipitators as Affected by the Properties and Combustion of Coal,"
journal of the Institute of Fuel, May 1963, pp. 18U-197-
k. Durie, R.A., et. al. Investigation of the Electrostatic Precipitation
of Fly Ashes from Coals to be Supplied to the Liddell Power Station,
Part 1, CSIRO Div. Min. Chem., Invest. Rept. 68, July 1967.
5. Bickelhaupt, R.E.. "Surface Resistivity and the Chemical Composition
of Fly Ash," Journal of the Air Pollution Control Association, 25:1^8
(February 19757^
6. Bickelhaupt, R.E. "Electrical Volume Conduction in Fly Ash," Journal
of the Air Pollution Control Association, 2^:251 (March 1971*T
7. Selle, S.J., P.H. Tufte, and G.H. Gronhovd. "A Study of the Electrical
Resistivity of Fly Ashes from Low-Sulfur Western Coals Using Various
Methods," paper 72-107 presented at the 65th Annual Meeting of the
Air Pollution Control Association, Miami Beach, Florida (1972).
8. Selle, S.J., L.L. Hess, and E.A. Sondreal. "Western Fly Ash Composition
as an Indicator of Resistivity and Pilot ESP Removal Efficiency,"
paper 75-02.5 presented at the 68th Annual Meeting of the Air
Pollution Control Association, Boston, Mass., (1975).
113
-------
9. Tufte, P.H. and W. Beckering. "A Proposed Mechanism for Ash Fouling
Burning Northern Great Plains Lignite," paper 7^-WA/CD-3 presented at
the ASME Winter Annual Meeting, New York, NY, November 1971*-
10. White, H.J. "Resistivity Problems in Electrostatic Precipitation,"
Journal of the Air Pollution Control Association, Vol. 2k, No. h,
pp. 313-338.
11. Tufte, P.H., G.H. Gronhovd, E.A. Sondreal, and S.J. Selle. "Ash
Fouling Potentials of Western Subbituminous Coal as Determined in a
Pilot Plant Test Furnace." presented at the American Power Conference,
Chicago, IL, April 1976.
12. Gronhovd, G.H. and E.A. Sondreal. "Technology and Use of Low-Rank Coals
in the U.S.A." Presented at Seminar on Technologies for the Utilization
of Low Calorie Fuels, Varna, Bulgaria, April 20-22, 1976.
114
-------
TABLE 1. - Hopper location versus fly ash sodium content,
ESPs on North Dakota lignites
Hopper Fly Ash NagO,
Powerplant
Leland Olds #1
Leland Olds #2
Hoot Lake #3
Big Stone
Date
1/28/75
5/9/75
11/20/7U
V8/76
Coal Ash Na20, percent Inlet Row
5,U 7.2
6.1 9. it
U.I 6.6
5-7
Hopper Location
2nd Row 3rd Row
8.3
10.7
9.0 7.8
percent
Outlet Row
Q.h
15. i*
8.6
10.6
-------
TABLE 2. - Rate of addition of high sodium materials as ESP
conditioners based on pilot-scale ESP tests
Composition of injected
fly ash/conditioner
mixture
100% Colstrip
.75$ High Na Ash
10$ High Na Ash
30% High Na Ash
1$ Nahcolite
5$ Nahcolite
50$ Nahcolite
Ib conditioner/
106 cf,
flue gas
0
3.03
39. 3U
122.01
2.6k
1U.U5
131*. 70
Ib Na/10^ cf,
flue gas
0
0.17
2.22
6.88
0.6l
3.35
31.19
Dust loading
(grains/acf )
—
2.8
2.7
2.8
1.8
2.0
1.9
In-situ
Resistivity
(ohm-cm)
3.3 x 1013
1.5 x 1012
2.7 x 1011
6.0 x 1010
2.3 x 1011
U.I x 1010
6.2 x 109
CJ
(ft /sec )
0.50
0.77
0.91
0.92
0.88
0.99
1.07
-------
TABLE 3. - ESP design and teat data for powerplanta burning North Dakota lignites
Basin Electric
Utility Company Power Cooperative
Leland Leland
Station Olds II Olds 12
Stanton,
Location North Dakota
Boiler capacity MW.. 215 liliO
Firing method , pc cyclone
ESP vendor Research Western
Cottrell
New or retrofit
installation Retrofit New
Completion date 11/T>> 9/75
Flue gas
Temperature, °F.. 360 . 373
volume, 1,000 acfm.. 1,000 2,100
velocity, Fps.. 5.01 5.00
Specific collection area*
ft2/!,000 acfn.. 320 267
Number of TR sets 16 Uo
Collecting surface/TR set,
ft8.. 19,970 lU.OltO
Collecting surface/Rapper,
ft2.. 2,500 1,755
Mlnnkota
Power Cooperative
Milton R. Milton R.
Young II Young 12
Center,
North Dakota
Otter Tail Power Company
Hoot Hoot
Lake *2 Lake 13
Fergus Falls,
Minnesota
Ortonville
Ortonville,
Minnesota
Big Stone
Milbank,
South Dakota
Montana Dakota Utilities
Heskett II Heskett 12
Mandan,
North Dakota
United Power
Association
UFA - Stanton
Stanton,
Horth Dakota
I-1
235
cyclone
Research
Cottrell
Retrofit
6/75
1(38
cyclone
Wheel-
abrator
New
5/77
61
Research
Cottrell
79
pc
Research
Cottrell
Retrofit Retrofit
5/T2 i/72
21
spreader-
stoker
Research
Cottrell
Retrofit
6/72
1*1.0
cyclone
Wheel-
abrator
New
5/75
25
spreader-
stoker
Research
Cottrell
Retrofit
6/75
spreader-
stoker
Research
Cottrell
Retrofit
6/75
160
pc
Research
Cottrell
Retrofit
5/76
385
1,170
5-55
288
16
21,050
2,630
380
2,200
5.00
375
32
25,800
1,120
330
280
It.23
252
U
310
390
5.07
236
It
17,650 23,075
3,530 2,560
133
It.25
280
k
9,310
2,070
288
2,330
5.25
355
21)
1,120
189
3.80
352
6
11,100
1,850
333
It52
It.28
280
10
12,650
2,090
350
85lt
5.17
235
12
16,720
2,785
Inlet loading gr/acF.. 2.30 1.30
Outlet loading gr/acF.. 0.0125 0.0125
Design efficiency pet.. 99.50 99.05
Measured efficiency., .pet.. 99A5 HA
Migration velocity,
cm/sec.. 8.26 NA
1.00 1.0 to 2.7
0.01 0.006
99.00 99.'tO
99.82 NA
11.15
NA
1.87
0.015
98.50
99.00
9.28
2.09
0.015
98.50
99*
9.9
0.97
0.00li2
98.90
99*
8.It
1.17
0.01 It
98.80
99.63
8.01
2.5 to U.I
0.0225 0.021
99.U5
97.00
0.01 gr/ft3 0.01 gr/ft3"
NA
98.00
NA
NA
I/ ESP downstream of mechanical collector.
2J Data not available.
3/ Only outlet loading has been measured to
-------
00
Ld
0. O
CO >
UJ
CO
Q.
U_
28
26
24
22
20
I.I
I
0.9
0.8
0.7
°-
8
10
12
FLUE GAS MOISTURE, % by volume
8
10
FIGURE 1. - Effect of flue gas moisture on pilot-plant ESP performance.
-------
I01'
10"
10"
10'
,10
I .0'
en
CO
I01
I01
10'
1 I I I I I
o o
oo
I I
300 °F
log f (300) = -0.390Na20+ 11.409
R=-0.838
o
o o
o
10'
10''
10'
E 10"
E
-C
o
^ 10s
en
LU
cc
0 I 23456789 10
Na20 IN FLY ASH, percent
10 "
10'
10'
i—i—i—r
600°F
log/) (600) "0.404 Na20 "10.846
R=-0,806
o
01 23456789 10
Na20 IN FLY ASH, percent
Figure 2 - Laboratory resistivity vs. NazO in fly ash at 300 and 600°F.
FIGURE 2. - Effect of fly ash sodium content on measured laboratory resistivity at 300 and 600° F.
-------
10
13
10
E
o
10 '
(300) = 0.011 Si02 - 0.035 CaO
-0.338 Na20 + 11.594
R = 0.886
10
10' 10'° 10"
PREDICTED RESISTIVITY, ohm-cm
10
12
Figure 3 - Measured vs. predicted resistivity at 300°F based on fly ash
chemical composition.
120
-------
10*
9
I01
E
o
I
E
10'
o
<
_J
I01
10*
•
• 100% OOL3TRIF
A 0.79% HI0H HO ASH ADDED
• 1% NAHOOLITI ADDED
HUMIDITY 12%
VOLTAGE GRADIENT 1.5 Kv/cm
JL
200
FIGURE 1*.
400
• 00
•00
TfMP. »F
- Effect on laboratory resistivity of small
quantities of high sodium materials.
121
-------
1.0
o
rO
.8
\L.
NAHCOLITE
HIGH Na FLY ASH
-.7ft-
.6
.1—
40
10 20 30
PERCENTAGE OF HIGH SODIUM
MATERIAL ADDED , WEIGHT %
FIGURE 5- - Effect of sodium conditioner on pilot-plant ESP performance.
-------
10
14
I- 10
CO
CO
UJ
CK
H
CO
z
E
o
I
£
10'
10
12
HIGH No ASH
10 30 50
PERCENTAGE OF HIGH SODIUM MATERIAL ADDED, WEIGHT
FIGURE 6. - Effect of sodium conditioner on in-situ resistivity.
-------
To stack
Point-to-plane
resistivity
apparatus
Electrostatic
precipitator
FIGURE !• - Schematic of pilot plant system used
for ESP work to date.
-------
c ..
B
ELECTRIC
AIR
HEATER
TERTIARY
AIR
COAL L_
FEEDER^
IN-SITU
RESIST
ESP
FIGURE 8. - Schematic of pilot plant system
currently under construction.
125
-------
ELECTROSTATIC PRECIPITATQR OPTIONS FOR
COLLECTION OF HIGH RESISTIVITY FLY ASH
L. E. Sparks
Industrial Environmental Research Laboratory
Environmental Protection Agency
Research Triangle Park, North Carolina
127
-------
ABSTRACT
ELECTROSTATIC PRECIPITATOR OPTIONS FOR COLLECTION
OF HIGH RESISTIVITY FLY ASH
by
L. E. Sparks
Industrial Environmental Research Laboratory
Environmental Protection Agency
Research Triangle Park, North Carolina 27711
An analysis "of the technical options for collecting high resistivity
fly ash is presented. The options are a large specific collector area
electrostatic precipitator operating at normal temperatures (about
1500°C), a wet electrostatic precipitator, a moderate to large specific
collector area electrostatic precipitator operating at either high (about
370 C) or low temperature (about 110°C), a moderate specific collector
area normal electrostatic precipitator plus fly ash conditioning. The
advantages and disadvantages of each option are discussed. The potential
for adverse environmental impact of the conditioning option is also
ussed.
128
-------
ELECTROSTATIC J'RECIPITATOK OPTIONS FOR COLLECTION OF
HIGH RESISTIVITY KLY ASH
INTRODUCTION
Combustion of low sulfur coal generally produces fly ash with higher
resistivity at 150°C than does combustion of high sulfur coal. Thus, fly ash
from low sulfur coal is more difficult to collect in electrostatic precipitators
than is fly ash from high sulfur coal.
There are several methods of overcoming the influence of high resistivity
caused by low sulfur coal. The options are:
1. Provide additional precipitator capacity; i.e..add collector
plate area to overcome the poorer performance.
2. Use a wet electrostatic precipitator.
3. Change the flue-gas temperature to achieve a more favorable
resistivity.
4. Add fly ash conditioning agents to change the resistivity or
otherwise modify electrical conditions in the precipitator.
In deciding which of these options is optimum for a given situation, one
must recognize that coals, even from the same mine, vary. This means that
the design coal will not be burned all the time. The variation in coal compo-
sition, especially of the sodium, iron and potassium content of the coal,
Bickelhaupt (1975), will result in variations in fly ash resistivity with re-
sultant variations in electrostatic precipitator performance. Available
data indicate that coal composition variations can easily result in day to
day resistivity variations of plus or minus a factor of two. The fact that
coal composition and thus fly ash resistivity will vary from day to day must
be considered in selecting ESP options. The method and costs of dealing
with the variation roust be considered. Questions such as, "will plate area
be added to handle the expected resistivity extremes?"-or "will the boiler be
operated at reduced load?" must be answered. Regardless of which option is
picked, it is- essential that the precipitator be well designed, built, operated
and maintained. Gas flow distribution must be good to excellent; sneakage and
rapping reentrainment losses must be minimized; wire alignment must be good;
ash handling must be properly designed; etc.
129
-------
COLD SinF, KU'CTROSTATIC J'RI'Cl): ITATOR
Cold side electrostatic precipitators (operating at ^ 150°C) can collect
fly ash from low sulfur coal if .sufficient collector plate area is provided.
The required collector plate area depends on the resistivity of the fly ash
23 2
and can be as high as 200 m /Am /sec ( ^ 1015 ft /AMCFM). Specific collector
areas in the range of 80 m?"/Am3/sec to 120 M /A.m3/sec (400 ft /AMCFM to 600
2
ft /AMCFM) are common. High efficiency precipitators with specific collector
areas in this range have been bui.lt for low sulfur applications.
If the required specific collector area is near the low end of the range
cited above, a normal cold side precipitator is likely to be the most econom-
ical method of collecting high resistivity fly ash. Some designers feel that
in general cold side units are less expensive than hot side units for
collection of fly ash from low sulfur, Matts (1975). Other designers believe
the opposite, Walker and.Fri.tsch (3.976) .
The disadvantages of cold side electrostatic precipitators for collect-
ing high resistivity all center around the sensitivity of precipitator
performance to small changes in resistivity. This sensitivity to resistivity
makes initial design of the precipitator somewhat uncertain, since fly ash
resistivity is likely to vary somewhat during the life of the boiler. Once
the precipitator is built, normal variations in resistivity can result in
performance degradation such that violations of air pollution regulations occur.
If the resistivity variation is of a short term nature, emission limits can
be met by reducing boiler load. However, if the resistivity change is of long
terra,such as might be expected if a new coal supply is used to fuel the boiler,
long term derating of the boiler or retrofit of additional particulate
collection equipment may be required.
WET ELECTROSTATIC PRKCIPITATOR
Because the dust layer is continuously washed from the electrodes and the
gas is saturated with water vapor, dust resistivity is not. a factor in the
performance of wet electrostatic precipitators. Thus, wet electrostatic
precipitators may be an economically attractive method of collecting high
resistivity fly ash.
130
-------
Whether or not wet precipitators are economically feasible for fly ash
collection depends on the cost of purchasing and operating the precipitators.
Because wet .precipitators must be built of corrosion resistant metaterials,
they cost more per square meter of plate area than do dry precipitators.
Presently available information indicates that for utility boiler applications
wet precipitators are likely to cost 2-1/2 times or more per square meter
of collection plate area. If the total collection plate area of the wet
precipitator is small enough, the added cost per unit of collector area can
be overcome. Performance data, Gooch and McCain (1975), on large plate type
wet electrostatic precipitators indicate that specific collection plate area
2 3
of about 60 m /Am /sec may be required for high efficiency fly ash collection.
Wet electrostatic precipitators have all the potential water pollution
problems that wet scrubbers have. The problems associated with meeting water
pollution requirements are likely to increase the cost of wet electrostatic
precipitators.
RESISTIVITY MODIFICATION
As discussed in the Cold Side ESP section, precipitator performance becomes
extremely sensitive to slight changes in resistivity when the resistivity
is high. Under such conditions relatively small change in ash resistivity
can lead to excessive particulate emission. Compliance with emission
standards can require operation at reduced boiler load with consequent loss
of power generation. Therefore, ac some point it becomes necessary to consider
steps to reduce the resistivity to a more acceptable lavel.
Resistivity can be changed by operating the precipitator at higher or
lower temperature or by adding chemical agents either to the fuel or to the
flue gas to condition the fly ash and reduce its resistivity. Although resist-
ivity modification may appear to be more expensive than living with the
problem, the advantages of improved confidence in design and lessened sens-
itivity to small changes in resistivity must also be considered.
131
-------
HI (Til TEMPERATURE OPERATION
Location of the electrostatic precipitator before the air heater allows
operation at about 370°C. At this temperature resistivity of most fly ash
is in the range 2-4 x 10 ohm-cm or lower. Thus, except for rare problem
ashes, hot side pre.cipitators can be. designed for relatively high current
densities and with confidence that changes in fuel chemical composition are
unlikely to affect performance.
Calculations based on the electrostatic precipitator performance model
described by Gooch et al (1975) and data reported by Walker andFritsch (1976)
2 3
indicate that a specific collector area of at least 67 m /Am /sec is required
for 99.5% particle collection efficiency in a hot side precipitator.
The primary disadvantage of hot side operation is the increase in gas
volume, due to elevated temperatures. Thus the reduction in SCA for high
temperature operation is somewhat offset by the increased gas flow. Table 1
compares calculated collector plate areas for 99.5% particulate removal
for various current densities at 150°C with collector plate area required
2,3 23
at 370°C on the basis of m /Am /sec and m /Nm /sec. Note that the hot side
2 3
precipitator requires a specific collector area of 158 m /Nin /sec.
Problems associated with thermal expansion must be given special attention
during the design and construction of hot side precipitators. There have
been cases where sufficient care was not taken in dealing with thermal
expansion which resulted in poor electrostatic precipitator performance.
Hot si.de precipitators must be well insulated to reduce heat losses.
Insulation and unavoidable heat losses are likely to increase the cost of
hot side operation. Walker and Fritscb (1976) present a lucid discussion of
design construction and operation of hot side electrostatic precipitators.
LOW TEMPERATURE ELECTROSTATIC PRECIPITATORS
For some cases a 20-50°C reduction in electrostatic precipitator operating
temperature rmy result in resistivities in the acceptable range. The exact
nature of the resistivity versus temperature curve at low temperatures is
variable and depends on the particular situation. A tenfold or greater varia-
tion in low temperature resistivity is common. Thus the. option of reducing gas
132
-------
Table 1. ELECTROSTATIC PRECTPITATOR SPECIFIC COLLECTOR
AREA REQUIRED FOR 99.5% PARTICLE COLLECTION
FOR VARIOUS CURRENT DENSITIES
Current^Density
na/cm"
5
20
40
30
Temperature
°C
150
150
150
370
Specif ic» Coll ec to
m /Am /sec in
122
53.1
39.4
67
r Area
/Nm /sec
189
82.3
61
158
Normal conditions 0°C at 1 atmosphere pressure
Calculations based on Gooch et al,(1975)
133
-------
temperature is subject to verification that the resistivity docs indeed decrease
witli decreasing temperature Lo a .satisfactory level at achievable temperatures.
This option is also subject to verification that the rate of corrosion is
acceptable; this verification may be possible if the corrosive affect of SO is
minimized by the CaO content, of the ash. Several precipitator installations have
performed satisfactorily at temperatures in the 95-120°C range without undue
problems, when sufficient attention was given to the ash-removal system to keep
the collected ash hot and to maintain the insulators free of condensed acid.
Achieving a flue gas temperature low enough to overcome the resistivity
problem presents problems that must be dealt with on an individual plant basis.
If there is sufficient capacity in the air heater, the problem can be handled
simply; however, this is not often the case. Whitehead (1971) reported that
120°C is about the minimum temperature that can be achieved even if the air
heater has sufficient capacity. An alternative approach to reaching low flue
gas temperatures in an existing boiler with insufficient capacity in the air
heater or if temperatures below 120°C are required is to increase the flow of
secondary air through the air heater and by-pass so'ine of the excess air to the
stack. This approach requires additional fan capacity and imposes an energy
penalty due to reduction in boiler efficiency, and therefore in most cases
is not too attractive.
Berube (1971) presents an excellent discussion of low temperature operation
at. one plant. For this plant a 0.8% reduction in boiler efficiency was
necessary to obtain the required temperature.
In general low temperature operation is sensitive to coal composition.
In a situation where coal composition varies over a wide range low temperature
may not be practical.
CONDITIONING
The addition of a chemical agent to the fuel or to the flue gas to
reduce fly ash resistivity has been successful in some cases. Whitehead (1971)
and others report that flue, gas conditioning should be considered as an option
for most high resistivity applications.
A partial list of presently available or proposed conditioning agents is
shown in TabJo 2.
134
-------
Table 2. PARTIAL LIST OF CONDITIONING AGENTS
Aluminum Sulfate
Ammonia
Ammonium Bisulfate
Ammonium Sulfate
Hydrogen. Chloride
Iron Oxide
Iron Sulfate
Organic Amines
Sodium Carbonate
Sulfamic Acid
Sulfur Trioxide
SuIfuric Acid
Vanadium Oxide
135
-------
The most commonly used conditioning agent is sulfur trioxide, SO,,. S0_
is injected into the i7lue gas either as SO or as sulfuric acid, H SO . Many
of the other agents shown in Table 2 are sold as proprietary agents which
may be blends of several of the compounds listed. At the temperatures existing
in the flue gas many of the compounds in Table 2 for example, ammonia and sulfur
trioxide disassociate into other compounds. Except for SO,, there is little
industrial experience with these conditioning agents some of which have not
been demonstrated as being effective.
With the exception of sodium and SO conditioning, the mechanism by which
many of the conditioning agents work are unclear. Both sodium conditioning
and SO conditioning reduce the resistivity of the fly ash, Bickelhaupt (197^)
and Disraukes (1975). Some of the sulfate conditioning agents appear to reduce
resistivity probably because SO is one of the products of disassociation,
Dismukes (1.97k) . Vanadium and iron oxides are claimed to catalyze the
reaction of S09 and 0 to SO, and thus increase the quantify of SO present
in the flue gas.
The effectiveness of conditioning with some of the agents listed in
Table 2 is summarized in Table 3. Based on the available data it appears
that properly engineered, operated and maintained conditioning systems using
S0_ as the active agent can improve resistivity limited performance of
electros catic. precipitators. The data on most other agents are insufficient
to make a general judgment of the effectiveness possible.
There are few data on the fate of conditioning agents after they are
injected into the flue gas. It is generally assumed that all the conditioning
agent is deposited on the fly ash and then collected in the precipitator,
Hall (1973). The available data, although limited, cast doubt on this
assumption.
There are more data on the effects of SO than other conditioning agents.
Whitehead (1971) reported that in Great Britain conditioning with up to 50 ppm
SO resulted in no increase in SO emissions. Dismukes (1975) reported data
for two plants that showed from 28 to 80% of the injected conditioning agent
was emitted. Published data on SO.., emissions with and without conditioning
for many plants with conditioning systems are lacking. Hall (1973) attributes
any increase in SO to poor system design and operation.
136
-------
Table 3. EFFECT OF CONDITIONING ON EFFICIENCY
ESP Efficiency or Part:Lculat_e Eirdssions
Without Injection With Injection
94% 95.2% Disraukes (1975)
37.5% 51.4%
67.3% 77.3%
69.7% 93.5%
Sulfamic Acid No statistically significant difference, Dismukes (1974)
Apollo LPA40 0.349/10bJ 0.139/106J, McNinch (1976)
Apollo LPA40 90%' 97% Harris (1976)
SO 80-88% 95-96% Klipstein (1975)
SO 85% 94-95% Cook (1975)
137
-------
Because SO, emit tod due to conditioning will form par L J.cula te sulfuric
acid or par ti culate sulfates essentially instantaneously upon reaching the
atmosphere, it seems reasonable to count the excess SO as participate sulfurir
3 3
acid. 1 ppin of sulfuri.c acid is 5.3 rag/Nm (0.0023 gr/scf" ) of particulate.
It is clear that the amount of SO that can escape the stack, if SO is counted
as particulate, must be kept at a minimum in order to reduce ambient particulate
loading.
Except for data taken by Harris (1976), there are no published data
on emissions from other agents. Harris reported that the injection of
Apollo Chemical's proprietary conditioning agent LPA40 appeared to result in
increased emissions of SO , ammonia, surface active agent and other organic
compounds. Many of the other proprietary conditioning agents and proposed
conditioning agents probably disassociate at flue gas temperatures to produce
SO and/or Nil — some of which may be emitted to the atmosphere.
3 3
The agents that are claimed to catalyze the oxidation of S0_ to SO are
likely to catalyze S0? to SO in the plume as well as in the boiler. Thus,
the emission of such an agent to the ambient air may increase the ambient
particulate concentration above acceptable limits due to formation of sulfates.
It should be noted that man)' of the conditioning agents listed in Table 2
are known to be toxic. Thus, an appreciable increase-in the emission of these
substances may result in adverse health effects — at least near the power
plant.
As stated above there are few data to support or contradict the discussion
above. It is clear, however, that toxic substances are being used as con-
ditioning agents and at least some of these agents are emitted to the atmosphere
The impact of these emissions on ambient air quality is unclear.
Conditioning does appear to be an acceptable ontiori for collecting high
resistivity fly ash in electrostatic precipitators under some conditions and
with some agents. Whenever conditioning is used, careful measurements of all
emissions are required to ensure that emissions of toxic substances are not
increased. EPA is presently conducting research to determine the. overall
environmental impact of conditioning agents.
138
-------
Disraukes (1975) reviewed the capital and operating cost of SO condition-
ing. His? data show that the cost of SO conditioning is sensitive to the
method used to produce the SO and the concentration of SO injected.
McNish (1976) reports data on costs of Apollo conditioning agent.
Dismukes (1974) presents some data on costs of sulfamic acid conditioning.
CONCLUSIONS
The ideal general solution to the problems associated with collecting high
resistivity fly ash from combustion of low sulfur coals does not exist.
Utilities who design plants initially for operation using low sulfur coal have
several options, one of which is likely to be best for their application.
In situations where fuel source and chemical composition is uncertain, hot side
electrostatic precipitators should certainly be considered. In cases where coal
supply is certain and coal chemical composition Is well defined, large cold
side electrostatic precipitators may be attractive,
Conditioning is likely to be the least expensive retrofit solution.
However, extreme care must be taken in all steps of design, construction and
operation of conditioning systems to avoid unacceptable emissions due to the
conditioning agents.
139
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1. Berube, D. T. (1971) "Low fias Temperature Solution 1.0 High Resistivity
Ash Problems." Proceedings of the Electrostatic Prcci pita tor Symposium, 197]..
2. Bickclhaupt, R. E. (1974n) Sodium Conditioning to Reduce- Fly Ash
Resistivity. EPA Report EPA-65'0/2-74-092, NTIS PB 236-922/AS.
3. (1975) Effect of Chemical Composition on Surface
Resistivity of Fly Ash. EPA Report EPA-600/2-75-017, NTIS PB 244-88.5/AS.
4. Cook, R. E. (1975) "Sulfur Trioxicle Conditioning" i.n Proceedings
of Symposium on Electrostatic Precipitators for Control of Fine Particles.
EPA Report EPA-650/2-75-016, NTT.S PB 240-440/AS.
5- Dismukes, E. B. (1974) Conditioning of Fly Ash with Sulfamic Acid,
Ammonium Sulfate, and Ammonium Bisulfate. EPA Report EPA-650/2-74-114,
NTIS PB 238-922/AS.
6- (1975) Conditioning of Fly Ash with Sulfur Trioxide
and Ammonia. EPA Report EPA-600/2-75-015, NTIS PB 247-231/AS.
7. Gooch, J. P. and McCain, J. D. (1975) Particulate Collection
Efficiency Measurements on a Wet Electrostatic Precipitator EPA Report
EPA-650/2-75-033, NTIS PB 244-173/AS.
8. Gooch, J. P., McDonald, J. R. and Oglesby, S. (1975^ A Mathematical
Model of Electrostatic Precipitation. EPA Report EPA-650/2-75-037,
NTIS PB 246-188/AS.
9. Hall, H. J. (1973) SO /H SO Gas Conditioning for Control of High
Resistivity Ash Problems in Electrostatic Precipitators. Technical Report
HAR 73-35.
10. Harris, D. B. (1976) Personal communication.
11. Klipstein, D. H. (1975) "Improved Precipitator Performance by Gas
Conditioning." Combustion, pg. 43, October.
12. Matts, S. (1975s) " 'Cold Side' Electric Precipitators for
High-ResiGtivity Fly Ash Require Different Design Philosophy" in Proceedings
of Symposium on Electrostatic Precipitators for Control of Fine Particles.
EPA Report !iPA~650/2--75~G16, HITS PB 240-440/A3..
140
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13. McNinch, M. 0. (1976) "Chemotherapy Tames Fly Ash from Low
Sulfur Coal" Electrical World, 32, January 15.
14. Walker, A. B. and Frisch, N. W. (1976) U. S. Design and
Operating Experience with Hot Electrostatic Precipitators on Coal-Fired
Utility Boilers. Paper presented at Joint US/USSR Symposium on Cortrol
of Particulate Emissions and Related Problems.
15. Whitehead, C. (1971) "Gas Conditioning." Proceedings of the
Electrostatic Precipitator Symposium, 1971.
NOTE: EPA Reports are available from National Technical Information Service
(NTIS), U. S. Dapartment of Commerce, Springfield, Va. 22151
141
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ELECTROSTATIC PRECIPITATOR OPERATING EXPERIENCE
WITH LOW SULFUR COAL - JIM BRIDGER POWER PROJECT
Thomas M. Ashton
Pacific Power & Light Company
Portland, Oregon
143
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ELECTROSTATIC PRECIPITATQR OPERATING EXPERIENCE
WITH LOW SULPHUR COAL - JIM BRIDGER POWER PROJECT
THOMAS M, ASHTON
PACIFIC POWER & LIGHT COMPANY
PORTLAND,, OREGON
ABSTRACT
BROAD SCALE REVIEWS OF THE ELECTROSTATIC PRECIPITATOR
TECHNOLOGY WERE APPLIED TO THE SELECTION AND DESIGN OF THE JlM
BRIDGER POWER PROJECT ELECTROSTATIC PRECIPITATORS, INCLUDED
WERE INVESTIGATIONS OF MULTI-YEAR EXPERIENCE WITH ELECTROSTATIC
PRECIPITATORS BY AUSTRALIAN ELECTRIC UTILITIES, EXTENSIVE PILOT
PLANT TESTING, CONSULTATION WITH RECOGNIZED UNITED STATES EX-
PERTS AND THOROUGH LABORATORY ANALYSES OF COAL AND ASH CHARAC-
TERISTICS INFLUENTIAL IN PRECIPITATOR DESIGNS,
FROM THIS COMPREHENSIVE APPROACH TO ENGINEERING OF AN
ELECTROSTATIC PRECIPITATOR INSTALLATION, PACIFIC POWER & LlGHT
COMPANY AND IDAHO POWER COMPANY ARE REPORTING ON THE SUCCESS OF
ONE OF THE FIRST MAJOR NEW "COLD SIDE" ELECTROSTATIC PRECIPI-
TATORS APPLIED TO A LARGE WESTERN, LOW SULPHUR COAL-FIRED ELEC-
TRIC UTILITY BOILER, THE PRECIPITATORS ARE CAPABLE OF ACHIEVING
PARTICULATE REMOVAL EFFICIENCIES IN EXCESS OF 99,3% WITH OUTLET
GRAIN LOADINGS LESS THAN 0,02 GRAINS PER ACTUAL CUBIC FOOT PER
MINUTE,
144
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ELECTROSTATIC PRECIPITATQR OPERATING EXPERIENCE
WITH LOW SULPHUR COAL - JIM BRIDGER POWER PROJECT
THOMAS M, ASHTON
PACIFIC POWER & LIGHT COMPANY
PORTLAND, OREGON
JTH BRIDGER POWER PROJECT
THIS PAPER WILL REVIEW THE ENGINEERING DEVELOPMENT AND
EARLY OPERATING EXPERIENCE OF THE ELECTROSTATIC PRECIPITATORS AS
APPLIED TO THE JlM BRIDGER POWER PROJECT, INCLUDED WILL BE DIS-
CUSSIONS OF BROAD SCALE INVESTIGATIONS OF THE U, S, AND FOREIGN
TECHNOLOGY AND EXPERIENCE, EXTENSIVE PILOT TESTING, SELECTION OF
FINAL DESIGN CRITERIA AND RESULTS OF FIRST PERFORMANCE TESTS,
FIRST, AN INTRODUCTION TO THE JlM BRIDGER POWER PROJECT
is APPROPRIATE:
1, LOCATION - WYOMING, 35 MILES ALMOST DUE EAST OF ROCK
SPRINGS IN THE SOUTHWESTERN WYOMING COUNTY OF SWEET-
WATER, THE PLANT IS APPROXIMATELY SIX MILES NORTH OF
INTERSTATE HIGHWAY 80 (FORMERLY U, S, 30), THE AREA
IS HIGH, 6,800 FEET, THE CONTINENTAL DIVIDE IS NEARBY,
THE LAND IS DRY, DESERT SHRUBS AND DESERT GRASSES
PREVAIL, THIS IS DESOLATE COUNTRY, ROCK SPRINGS,
GREEN RIVER, RELIANCE AND SUPERIOR ARE THOSE NEARBY
TOWNS THAT STAND OUT IN THIS WIDE EXPANSE OF THE HIS-
TORIC WEST, THE OLD OVERLAND TRAIL CROSSES THE AREA
AS DOES THE MAIN LINE OF THE UNION PACIFIC RAILROAD,
145
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2, OWNERS - PACIFIC POWER & LIGHT COMPANY AND IDAHO
POWER COMPANY, SHARING TWO-THIRDS AND ONE-THIRD,
RESPECTIVELY,
3, SIZE - FOUR UNITS, EACH 500 MEGAWATTS, A TOTAL OF
2,000 MEGAWATTS,
4, EilEL - WESTERN COAL, LOW BTU CONTENT (9,600 BTU/LB),
LOW SULPHUR CONTENT (0,6%), MODERATE ASH. (10%),
COAL FEED AT MAXIMUM LOAD IS 275 TONS PER HOUR EACH
UNIT,
5, COAL MINE - A STRIP MINE LOCATED APPROXIMATELY
FOUR MILES NORTHEAST OF THE PLANT SITE, OWNED AND
OPERATED BY A SUBSIDIARY OF THE OWNERS, COAL LANDS
UNDER LEASE TO THE OWNERS FROM THE UNION PACIFIC
RAILROAD AND BUREAU OF LAND MANAGEMENT TOTAL 18,000
ACRES, A FUEL RESOURCE OF OVER 130 MILLION TONS
WITH OVERBURDEN OF 120 FEET OR LESS EXISTS AT THIS
MINE, APPROXIMATELY 125 ACRES PER YEAR WILL BE MINED
WITH THE OWNERS DEDICATED TO A FULL SCALE PROGRAM OF
LAND REHABILITATION,
6, WATER SUPPLY - FROM THE GREEN RIVER BY WAY OF A
42-MILE PIPELINE, UTILIZING CONTROLLED RELEASE
OF PURCHASED WATER FROM THE FONTENELLE RESERVOIR,
WATER CONSUMPTION WILL BE 18,700 GPM MAXIMUM FOR
FOUR POWER PLANT UNITS,
7. TRANSMISSION FACILITIES - THREE 345 KV LINES FROM
THE JIM BRIDGER SWITCHYARD TO THREE MAJOR LOAD
POINTS IN SOUTHEASTERN IDAHO,
146
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8. PROJECT STATUS;
A) UNIT 1 WAS PLACED IN COMMERCIAL OPERATION
NOVEMBER 30, 1974,
B) UNIT 2 IN COMMERCIAL OPERATION DECEMBER 1, 1975,
c) UNIT 3 is UNDERGOING START-UP NOW AND is SCHEDULED
FOR COMMERCIAL OPERATION SEPTEMBER 1, 1976,
D) UNIT 4 IS IN THE INITIAL STAGES OF CONSTRUCTION
WITH COMMERCIAL OPERATION SCHEDULED FOR DECEMBER,
1979,
THE EARLY SITUATION AND INITIAL TECHNICAL REVIEWS
RELATIVE TO PARTICULATE EMISSION CONTROL, QUALITY OF THE
AIR AND THE JlM BRIDGER POWER PROJECT, THESE WERE THE CIRCUMSTANCES
PREVAILING AT THE BEGINNING, THE PERIOD IN TIME IS 1970-1971, PUB'
LIC CONCERNS FOR PROTECTION OF THE ENVIRONMENT, PARTICULARLY AIR
QUALITY, WERE MOUNTING, ACROSS THE COUNTRY, STRINGENT STATE AND
FEDERAL AIR QUALITY CONTROL REGULATIONS WERE IN VARIOUS STAGES OF
"DRAFT," PROMULGATION AND ENFORCEMENT, WYOMING HAD PARTICULATE
EMISSION STANDARDS OF 0,18 LB, PER MILLION BTU INPUT (0,05 GRAINS
PER ACF) AND A RlNGLEMANN #1 OPACITY LIMITATION, PACIFIC POWER
AND IDAHO POWER SAW A NEED FOR A NEW POWER PLANT WITH "CLEAR"
CHIMNEYS — CHIMNEY PLUMES THAT WOULD NOT VIOLATE THE BLUE SKY,
PRISTINE CONDITION OF THE WIDE OPEN ATMOSPHERIC SPACES OF WYOMING,
147
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A TECHNICAL APPRAISAL AT THIS STAGE EQUATED A "CLEAR" CHIMNEY
TO PARTICULATE EMISSIONS NOT EXCEEDING 0,02 GRAINS PER ACTUAL
CUBIC FOOT OF FLUE GAS,
WHERE DID THE PARTICULATE COLLECTION TECHNOLOGY STAND
AT THIS POINT IN TIME TO SUPPORT A CLEAR CHIMNEY DESIGN? DlSMAL
AT BEST WAS THE TECHNICAL SITUATION, THERE WERE NO HIGH EFFI-
CIENCY PRECIPITATORS OPERATING SUCCESSFULLY IN THE U, S, IN THE
PROCESS OF COLLECTING LOW SULFUR COAL ASH, THOSE PRECIPITATORS
IN OPERATION WERE EITHER NOT OPERATING SUCCESSFULLY OR NOT
DESIGNED FOR COLLECTION EFFICIENCIES IN THE 99%+ RANGE — OR CLEAR
CHIMNEY RANGE,
TECHNICAL UNDERSTANDINGS OF ASH RESISTIVITY, ASH CON-
STITUENTS, MOISTURE, GAS CONDITIONING, TRANSVERSE RAPPING AC-
CELERATIONS, RIGID VS SUSPENDED WIRE ELECTRODES, "HOT" VS "COLD"
WERE STILL DEVELOPING, AT LEAST AS APPLIED TO HIGH EFFICIENCY
COLLECTION OF LOW SULFUR, WESTERN COAL ASH PARTICLES, THIS IS ALL
TRUE WHEN VIEWED FROM THE BACKGROUND OF UNITED STATES EXPERIENCE,
IN EUROPE, INCLUDING GREAT BRITAIN, AND IN AUSTRAILIA, MULTI-YEAR
EXPERIENCE HAD BEEN RECORDED IN THE SUCCESSFUL HIGH EFFICIENCY
COLLECTION OF ASH PARTICLES FROM COALS SIMILAR TO THOSE EXISTING
IN THE WESTERN UNITED STATES,
PACIFIC POWER, IDAHO POWER AND THEIR ENGINEERS, THEREFORE,
UNDERTOOK EXTENSIVE INVESTIGATIONS OF PRECIPITATOR EXPERIENCE IN
AUSTRALIA (NEW SOUTH WALES AND QUEENSLAND) AND GREAT BRITAIN, A
SUMMARY OF IMPORTANT OBSERVATIONS AND CONCLUSIONS RESULTING FROM
THESE INVESTIGATIONS FOLLOWS;
148
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1. AUSTRALIA - WITH THE VERY THOROUGH AND EXPERT AS-
SISTANCE OF SEVERAL ENGINEERS OF THE ELECTRICITY COMMISSION OF
NEW SOUTH WALES, PLANT SITE INVESTIGATIONS WERE UNDERTAKEN AT
SEVERAL OF THE COMMISSION'S POWER PLANTS — M.UNMORAH, VALES POINT,
LlDDELL AND TALLAWARA IN NEW SOUTH WALES, AS WELL AS THE SWANBANK
POWER STATION IN QUEENSLAND (SOUTHERN ELECTRICITY AUTHORITY),
HERE WERE SEVERAL LARGE SCALE, COLD SIDE PRECIPITATORS OPERATING
WITH COLLECTION EFFICIENCIES IN EXCESS OF 99% AND COLLECTING ASH
PARTICLES FROM LOW SULFUR COAL (1/2% OR LESS), CHIMNEY PLUMES
WERE NOT CLEAR, INLET GRAIN LOADINGS WERE HIGH, TWO TO THREE
TIMES GREATER THAN EXPECTED LOADINGS AT JlM BRIDGER, THEREFORE,
99%+ COLLECTION STILL RESULTED IN OUTLET GRAIN LOADINGS OF AP-
PROXIMATELY 0,1 GRAIN PER CUBIC FOOT, WELL WITHIN THE VISUAL PLUME
RANGE, ALSO, WELL WITHIN THE AUSTRALIAN EMISSION STANDARD OF 0,2
GRAINS/CU, FT, EXISTING AT THAT TIME, THE IMPORTANT POINT HERE IS
THAT LOW SULFUR COAL ASH WAS BEING COLLECTED AT HIGH EFFICIENCY
LEVELS ON A SUSTAINED BASIS,
THE ELECTRICITY COMMISSION OF NEW SOUTH WALES, IN 1971,
HAD A WELL DEFINED APPROACH TO THE SIZING AND SPECIFICATION OF NEW
PRECIPITATOR EQUIPMENT, THE PRINCIPAL ELEMENT IN THIS APPROACH WAS
THOROUGH PILOT TESTING OF THE COAL TO BE BURNED IN THE NEW UNIT,
NEED FOR SUCH TESTING HAD BEEN DEMONSTRATED IN THE ECNSW EXPERIENCE
BY AN INABILITY TO PROPERLY SIZE PRECIPITATOR EQUIPMENT BASED SOLELY
ON EXAMINATION AND COMPARISON OF COAL AND ASH LABORATORY ANALYSES,
149
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FURTHER EXPERIENCE OF THE ECNSW ESTABLISHED THE NEED FOR
VERY HEAVY RAPPING FORCES, NOT ONLY IS THE AUSTRALIAN ASH DIFFICULT
TO PRECIPITATE, IT IS ALSO DIFFICULT TO REMOVE FROM THE COLLECTING
PLATES, AUSTRALIAN SPECIFICATIONS THEREFORE CONTAIN THE REQUIRE-
MENT FOR A MINIMUM TRANSVERSE ACCELERATION OF 50 G'S AT ANY POINT
ON THE PLATE, THE GENERAL THEORY IS HEAVY, INFREQUENT RAPPING AS
OPPOSED TO LIGHT, FREQUENT RAPPING, AND SUCH A HEAVY RAPPING
SPECIFICATION DICTATES A STRONG, RUGGED STRUCTURAL DESIGN OF PLATES
AND SUPPORTING COMPONENTS,
ONE OTHER OBSERVATION FROM THE AUSTRALIAN INVESTIGATIONS,
TO ASSIST THE ECNSW IN RESOLUTION OF THE SEVERAL BAFFLING ISSUES
EXISTING WITHIN THE ELECTROSTATIC PRECIPITATOR TECHNOLOGY, SER-
VICES OF THREE LABORATORIES WERE ACTIVE IN DEVELOPMENTAL AND RE-
SEARCH WORK, COMMONWEALTH SCIENTIFIC INDUSTRIAL RESEARCH ORGANIZA-
TION, CISRO, HAS BEEN INVOLVED IN COAL ASH PRECIPITABILITY AND PAR-
TICLE SIZE RESEARCH, AUSTRALIAN COAL INDUSTRY RESEARCH LABORATORY,
ACIRL, HAS CONDUCTED EXTENSIVE RESEARCH IN THOSE AREAS OF PARTICLE
SHAPE AND DISTRIBUTION EMPLOYING PHOTOMICROSCOPY, WOLLONGONG
UNIVERSITY COLLEGE HAS CONCENTRATED EFFORTS IN THOSE AREAS OF A
BETTER UNDERSTANDING OF ASH RESISTIVITY AND THE ADHESIVE FORCES OF
ASH DEPOSITED ON COLLECTING SURFACES,
FROM AUSTRALIA THEN, THESE PERTINENT POINTS:
A) LOW SULFUR COAL ASH WAS BEING COLLECTED AT
HIGH LEVELS OF EFFICIENCY, COLD SIDE;
B) PILOT TESTING HAD BEEN DETERMINED TO BE A
REQUIREMENT TO CONFIRM A SATISFACTORY DESIGN
FOR AN UNKNOWN COALj
150
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C) HEAVY RAPPING FORCES AND A RUGGED STRUCTURAL
DESIGN TO WITHSTAND THESE FORCES WERE CON-
SIDERED NECESSARY TO KEEP COLLECTING PLATES
CLEAN AND PRECIPITATOR PERFORMANCE AT A CON-
TINUING SATISFACTORY LEVEL,
2, GREAT BRITAIN - ONE OF THE OWNER'S ENGINEERS VISITED
THE STAYTHORPE "B" POWER STATION OF THE CENTRAL ELECTRICITY GEN-
ERATING BOARD IN ENGLAND, AGAIN. THE PRINCIPAL INPUTS TO THE JIM
BRIDGER PRECIPITATOR DECISION FROM THIS VISIT WERE:
A) STAYTHORPE PRECIPITATORS WERE OPERATING SATIS-
FACTORILY AT COLLECTION EFFICIENCIES IN EXCESS
OF 99%. TREATING ASH FROM LOW SULFUR COALS,
B) THE SPECIFIC PRECIPITATOR DESIGN INCLUDED HEAVY
RAPPING,
MANUFACTURER PREQUALIFICATIQN
IN MiD-1971, THE JIM BRIDGER OWNERS ISSUED TO CANDIDATE
ELECTROSTATIC PRECIPITATOR MANUFACTURERS A LETTER REQUESTING
INFORMATION TO BE USED IN A VENDOR PRE-QUALIFI CATION SURVEY, THIS
PRE-QUALIFICATION LETTER REQUESTED!
1, AN INSTALLATION LIST OF ELECTROSTATIC PRECIPITATORS
SUCCESSFULLY OPERATING ON UNITS WTH LOW SULFUR COAL
(1% OR LESS) AND EFFICIENCIES OF 99% OR GREATER,
151
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2, A STATEMENT OF APPROPRIATE PRECIPITATOR TESTS
AND ANY EVIDENCE OF SUSTAINED SATISFACTORY PER-
FORMANCE,
3, THE APPROACH USED IN DETERMINING THE SIZE FOR
THE ABOVE UNITS AND A STATEMENT RELATIVE TO MANU-
FACTURERS EXPERIENCE AND USE OF PILOT PRECIPITATORS,
4, A STATEMENT OF MANUFACTURERS PROPOSED METHOD OF
SIZING THE JIM BRIDGER PRECIPITATORS,
5, INFORMATION AS TO THE AVAILABILITY OF A THREE-
STAGE PILOT PRECIPITATOK,
THE VENDOR QUALIFICATION CRITERIA HERE WAS TO LIMIT PRO-
POSALS TO THOSE PRECIPITATOR MANUFACTURERS WITH DIRECT EXPERIENCE
AND SUCCESS IN THE DESIGN AND FURNISHING OF EQUIPMENT TO OPERATE
ON LESS THAN 1% SULFUR COAL ASH AND AT A COLLECTION EFFICIENCY
OF 99% OR GREATER, ADDITIONALLY, THE CRITERIA SEEKS TO ESTABLISH
AN UNDERSTANDING OF PRECIPITATOR SIZING TECHNIQUES AND THE MANU-
FACTURERS POSITION WITH REGARD TO THE APPLICATION AND USE OF PILOT
TESTING, PILOT TESTING WAS GOING TO BE A MAJOR PART OF THE JlM
BRIDGER PRECIPITATOR DEVELOPMENT PROGRAM,
ELECTROSTATIC PRECIPITATQR SPECIFICATION
FROM THESE EARLY INVESTIGATIONS COME SEVERAL UNIQUE FEA-
TURES OF THE BIDDING SPECIFICATION FOR THE JlM BRIDGER ELECTRO-
STATIC PRECIPITATORS:
1, PROPOSALS WERE TO BE BASED UPON A COMPLETE DE-
SIGN FOR PRECIPITATORS WITH A BASE SPECIFIC
COLLECTING AREA OF 400 SQ, FT,/I,000 ACFM,
152
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2, UNIT PRICES WERE TO BE QUOTED FOR SUBSEQUENT ADJUST-
MENTS IN SPECIFIC COLLECTING AREA, NUMBER OF ELEC-
TRICAL SETS AND CHANGES IN DUCT CONFIGURATION,
3, SEPARATE GAS PATHS WERE SPECIFIED WITH THE SPECIFIC
REQUIREMENT THAT ANY ONE GAS PATH COULD BE REMOVED
FROM SERVICE, WlTH ONE GAS PATH OUT OF SERVICE,
THE PRECIPITATOR WAS SPECIFIED TO TREAT FULL GAS
FLOW AND NOT EXCEED GUARANTEED MAXIMUM PARTICULATE
EMISSION RATES, COMPLETE ISOLATION CAPABILITY OF
EACH GAS PATH WAS CALLED FOR, ELECTRICALLY AND
MECHANICALLY, SO THAT MAINTENANCE WITHIN AND UPON
THE OUT OF SERVICE GAS PATH COULD BE ACCOMPLISHED,
4, GUARANTEED PERFORMANCE AT AN EMISSION RATE OF 0,02
GRAINS/ACF (EQUIVALENT TO 99,33% COLLECTION EFFI-
CIENCY BASED ON MAXIMUM CALCULATED INLET GRAIN LOAD-
ING), THE MAXIMUM PARTICULATE EMISSION RATE WAS TO
BE GUARANTEED WHEN FIRING COALS WITHIN THE RANGE OF
33 SEPARATE CORE HOLE SAMPLE ANALYSES REPRESENTING
THE ENTIRE COAL FIELD,
5, A MINIMUM COLLECTING PLATE RAPPING ACCELERATION OF
50 G MEASURED NORMAL TO THE SURFACE WAS SPECIFIED,
6, PROVISION FOR A SPARE EMPTY SECTION (20% ADDITIONAL
SURFACE) WAS SPECIFIED,
7, FLUE GAS CONDITIONING WOULD NOT BE PERMITTED,
153
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8, A COMPLETE DESCRIPTION OF MODIFICATIONS, ADJUST-
MENTS AND/OR ADDITIONS CONTEMPLATED TO CORRECT
PERFORMANCE DEFICIENCIES IN THE EVENT GUARANTEES
WERE NOT MET WAS A SPECIFIED REQUIREMENT,
9, GAS FLOW DISTRIBUTION MODEL TESTING WAS ALSO A
SPECIFIED REQUIREMENT,
10, TO ESTABLISH FINAL DESIGN PARAMETERS, SPECIFIC
COLLECTING PLATE AREA, 'ELECTRICAL CHARACTERISTICS,
ETC.j PILOT PRECIPITATOR TESTS WERE SPECIFIED,
THE SUCCESSFUL VENDOR WAS TO SUPPLY:
A) PILOT ELECTROSTATIC PRECIPITATOR, COMPLETE,
WITH AT LEAST THREE ELECTRICAL SECTIONS TO
BE USED AT A POWER STATION (AT THIS TIME NOT
YET IDENTIFIED) FIRING JIM BRIDGER COAL,
B) TECHNIQUES FOR CORRELATION OF PILOT TEST RE-
SULTS WITH JIM BRIDGER DESIGN AND PERFORMANCE
ELEMENTS,
C) ALL TEST INSTRUMENTATION,
D) TEST SUPERVISION AND PERSONNEL,
E) DESIGN OF CONNECTING INLET AND OUTLET DUCTWORK,
F) CHEMICAL ANALYSES, SIZE DISTRIBUTION TESTS
AND RESISTIVITY TESTS FOR BOTH PARTICULATES
AND COAL AS APPROPRIATE,
154
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G) TEST PROCEDURES, INCLUDING MEASUREMENT TECH-
NIQUES WERE TO BE DEVELOPED JOINTLY, OWNERS
AND MANUFACTURER,
11, PERFORMANCE TESTS OF THE JIM BRIDGER UNITS WERE
SPECIFIED AS FOLLOWS:
A) PREOPERATIONAL TESTS - A PHYSICAL INSPECTION
AND COLD AIR FLOW DISTRIBUTION TEST TO VERIFY
PROPER CONSTRUCTION,
B) PERIOD No, 1 - STARTING UPON INITIAL OPERA-
TION AND ENDING WITH WRITTEN NOTIFICATION
THAT EQUIPMENT IS READY FOR SATISFACTORY
OPERATION, A PERIOD OF 1,500 HOURS OR LESS
OF ON LOAD OPERATION, ALLOWING FOR MANUFACTUR-
ERS OPERATIONAL TESTS AND ADJUSTMENTS AS RE-
QUIRED,
c) PERIOD No, 2 - STARTING UPON RECEIPT OF NOTI-
FICATION THAT EQUIPMENT IS READY FOR SATIS-
FACTORY OPERATION AND EXTENDING FOR 2,000 HOURS
OF ON-LOAD OPERATION, No ALTERATION OF CON-
TROL SETTINGS, NO ADJUSTMENTS, NO MODIFICATIONS,
NO ADDITIONS WOULD BE ALLOWED,
«A «
D) PERFORMANCE TEST "A,
E) PERIOD No, 3 - STARTING AT THE SUCCESSFUL COM-
PLETION OF TEST "A" AND EXTENDING FOR A PERIOD
OF 4,000 HOURS OF ON-LOAD OPERATION, AGAIN,
NO CONTROL SETTING ALTERATIONS, NO ADJUSTMENTS,
NO MODIFICATIONS, NO ADDITIONS WOULD BE ALLOWED,
155
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F) PERFORMANCE TEST "B" - SUCCESSFUL COMPLETION
OF TEST "B" WILL CONSTITUTE THE DATE OF PLAC-
ING THE EQUIPMENT INTO COMMERCIAL OPERATION,
DURING THIS PERIOD OF TIME IN THE JIM BRIDGER PRECIPITATOR
DEVELOPMENT PROGRAM, THERE WERE CONTINUING INVESTIGATIONS AND TECH-
NICAL INPUTS, THE AUSTRALIAN LABORATORIES CISRO, ACRIL AND WOLLONGONG
UNIVERSITY COLLEGE HAD COMPLETED ANALYSES AND STUDY OF JIM BRIDGER
COAL SAMPLES, MANUFACTURERS ACTIVE IN THE PROGRAM WERE CONTRIBUTING
FURTHER TECHNICAL UNDERSTANDINGS, IN ADDITION TO BECHTEL CORPORATION,
ENGINEERS FOR THE PROJECT, CONSULTATIONS WITH SEVERAL TECHNICAL EX-
PERTS WERE CONTINUING, DR, HARRY J, WHITE, RETIRED PORTLAND STATE
UNIVERSITY PROFFESSOR,- INDIVIDUALS FROM SOUTHERN RESEARCH INSTITUTE
AND MR, KEN WATSON, ELECTRICITY COMMISSION OF NEW SOUTH WALES, ALL
WERE INVOLVED DEEPLY IN THE PROGRAM DEVELOPMENT AT THIS STAGE, TWO
INTERESTING CHANGES WERE MADE IN THE ELECTROSTATIC PRECIPITATOR SPEC-
IFICATION AS A RESULT OF THESE FURTHER INVESTIGATIONS,
THE COLLECTION EFFICIENCY GUARANTEE OF 99,33% AND THE EMIS-
SION RATE GUARANTEE OF 0,02 GRAINS/ACF WERE TIED TO THOSE COAL FIELD
CORE SAMPLES WHOSE ANALYSES INDICATED AN ALKALI-SULFATE INDEX (ASI)
EQUAL TO OR GREATER THAN THE ALKALI-SULFATE INDEX OF THE PILOT
TEST COAL (ASIp), THE ALKALI-SULFATE INDEX is AN EMPIRICAL TECH-
NIQUE IN THE ASSESSMENT OF ASH RESISTIVITY AND THEREFORE ASH PRE~
CIPITABILITY, IT IS NOT THE PURPOSE OF THIS PAPER TO DISCUSS THE
TECHNICAL VALIDITY OR DETAILS OF THE ALKALI-SULFATE INDEX OUTSIDE
THE INFLUENCE ON THE FINAL JlM BRIDGER ELECTROSTATIC PRECIPITATOR
SPECIFICATION, IN THE SPECIFIC CASE OF JIM BRIDGER, THE RESULT WAS
THAT THE EMISSION GUARANTEES (0,02 GRAINS/ACF AND 99,33%) WOULD
APPLY ONLY DURING THE FIRING OF COALS WITH AN ASI GREATER THAN ASIp,
INCLUDING COALS REPRESENTED BY 25 OF THE 33 CORE SAMPLES,
156
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FURTHER ECONOMIC EVALUATIONS DICTATED A PRUDENT MODIFICA-
TION TO THE ORIGINAL SPECIFICATION THAT GUARANTEED PERFORMANCE BE
MAINTAINED WITH FULL GAS FLOW AND ONE GAS PATH OUT OF SERVICE, THE
FINAL SPECIFICATION LANGUAGE READS, "SELLER GUARANTEES A COLLECTION
EFFICIENCY OF 98,73% WITH ONE GAS PATH REMOVED FROM SERVICE WITH
FLUE GAS FLOW LIMITED TO 95% OF DESIGN, "
BASED ON THESE PREQUALIFICATION CRITERIA AND SPECIFICATIONS,
A JOINT VENTURE OF A, B, SVENSKA FLAKTFABRIKEN AND CARBORUNDUM IN-
CORPORATED WAS SELECTED AS THE SUPPLIER OF THE JlM BRIDGER UNIT 1,
2 AND 3 ELECTROSTATIC PRECIPITATORS, THE DATE, OCTOBER, 1971, THE
TOTAL COST - $32,000,000 FOR THREE UNITS, EQUAL TO $21+ PER KILOWATT,
PILOT PRECIPITATQR UNIT
CONCURRENT WITH THE SELECTION OF THE JIM BRIDGER ELECTRO-
STATIC PRECIPITATOR MANUFACTURER, SELECTION OF A POWER PLANT SITE FOR
THE PILOT TESTING WORK, OPENING OF THE JlM BRIDGER COAL MINE AS A
FUEL SOURCE AND ARRANGEMENTS FOR SHIPMENT OF THE COAL TO THE TEST
PLANT WERE ALL UNDERWAY, PERTINENT DETAILS OF THESE THREE PHASES
OF THE PROGRAM FOLLOW:
1. TEST SITE - PACIFIC POWER & LIGHT COMPANY'S
DAVE JOHNSTON PLANT, APPROXIMATELY 30 MILES DUE
EAST OF CASPER, WYOMING WAS SELECTED AS THE
TEST SITE, UNIT 1 HAS THESE STATISTICS; 100 MWj
B&W, 1550 PSIG, 1000°F, SINGLE REHEAT, FRONT FIRED
BOILER; B&W EL76 BALL MILLS,
2, TEST COAL - FROM A REVIEW OF THE JIM BRIDGER COAL
MINE CORE HOLE ANALYSES, AN AREA REPRESENTATIVE
OF THE ENTIRE COAL FIELD WAS SELECTED, THERE
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WERE, OF COURSE, PRACTICAL MINING LIMITATIONS IM-
POSED ON THIS SELECTION PROCESS, THIRTY THOUSAND
TONS WERE MINED FROM THIS AREA, THE FIRST COAL FROM
THE JIM BRIDGER FIELD,
3, COAL SHIPMENT AND HANDLING - THE TEST COAL WAS MINED,
PROCESSED THROUGH AN INITIAL CRUSHING OPERATION AND
LOADED ONTO RAILROAD CARS, AT THE DAVE JOHNSTON PLANT,
THE COAL WAS UNLOADED AND STOCKPILED IN A CONICAL
READY PILE, ISOLATED FROM THE OTHER DAVE JOHNSTON
OPERATIONS TO PRECLUDED CONTAMINATION, SIMILARLY, THE
FEED AND SILO SYSTEMS OF UNIT 1 WERE CLEARED OF DAVE
JOHNSTON COAL PRIOR TO INTRODUCTION OF THE JIM BRIDGER
COAL,
APPROPRIATE HERE is A DESCRIPTION OF THE PILOT PRECIPITATOR:
1. OWNERSHIP - SF HAD JUST COMPLETED THE ASSEMBLY OF A
NEW PILOT PRECIPITATOR IN CANADA, SIMILAR TO THE
UNIT USED FOR AUSTRALIAN TESTS FOR VALES POINT, LlDDELL
AND WALLERAWANG POWER STATIONS,
2, SECTIONALIZATIQN - THREE IN LINE FIELDS, EACH EQUIPPED
WITH A HIGH VOLTAGE ELECTRICAL SUPPLY AND RAPPING
MECHANISM FOR COLLECTING PLATES AND FOR DISCHARGE
ELECTRODES,
3, PHYSICAL SIZE - APPROXIMATELY 31' LONG x 12' WIDE x
20' HIGH, CAPABLE OF TREATING A NOMINAL 2,500 ACFM
OF FLUE GAS,
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4. GAS FLOW DISTRIBUTION - TWO PERFORATED PLATES PRE-
CEDED THE INLET FIELD, ADDITIONALLY, FIELD ADJUST-
MENT OF SEVERAL GUIDE VANES AND BAFFLES ATTACHED TO
THE SECOND SCREEN ACCOMPLISHED A UNIFORM GAS DISTRI-
BUTION ACROSS THE FACE OF THE INLET FIELD,
5, INLET DUCTS - TWO INLET DUCTS, ONE FOR COLD GAS TESTS
AND ONE FOR HOT GAS TESTS, WERE DESIGNED SPECIFICALLY
TO ACHIEVE ISOKINETIC SAMPLING, DUCTS WERE EQUIPPED
WITH A HEAT EXCHANGER USED AS A COOLER (AIR AND WATER)
AND ELECTRIC HEATERS TO CONTROL GAS TEMPERATURES, PlLOT
UNIT CASING WAS FITTED WITH ELECTRIC HEATERS TO MAIN-
TAIN NEAR CONSTANT TEMPERATURES THROUGH THE PRECIPITATOR,
6, ELECTRODES - FOUR PARALLEL GAS PATHS, SPACED 9-3/4"
AND FORMED BY COLLECTING ELECTRODES 6'-6'3/4" HIGH,
TOTAL COLLECTING SURFACE 826,7 SQ, FT, DISCHARGE
WIRES WERE SPIRAL WOUND, ROUND SECTION, 0,1" DIAMETER,
7, RAPPERS - TUMBLING HAMMER TYPE, FOR BOTH COLLECTING
AND DISCHARGE ELECTRODES,
8, ELECTRICAL - FULL-WAVE SELENIUM RECTIFIERS, GOKV EACH
AT 45 MA RATING,
9, ASH REMOVAL - A CHAIN CONVEYOR DISCHARGES COLLECTED
ASH INTO THREE STEEL CANS UNDER EACH COLLECTING HOPPER,
THROUGH AN AIR-TIGHT VALVE,
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THE ENTIRE PILOT UNIT WAS HOUSED IN A TEMPORARY, QUONSET
TYPE BUILDING ADJACENT TO DAVE JOHNSTON UNIT 1, THE BUILDING WAS
PROVIDED WITH HEAT, LIGHTS AND WORK BENCHES, OFFICE SPACE AND
TELEPHONE SERVICE, INCLUDING A COMPUTER TERMINAL, WERE PROVIDED
IN A NEARBY TRAILER, CALCULATIONS TO VERIFY THE VALIDITY OF A
TEST RUN WERE QUICKLY MADE PRIOR TO THE COMMENCEMENT OF THE
NEXT TEST,
TESTING PROGRAM
FOLLOWING A SERIES OF PRETESTS TO CHECK EQUIPMENT OPERA-
TION AND TO TRAIN PERSONNEL, THE TEST PROGRAM MOVED TO THE FIRST
TEST RUN ON NOVEMBER 29, 1971, DURING A THREE-MONTH PERIOD, NOVEMBER,
1971 THROUGH JANUARY, 1972, OVER 175 TEST RUNS WERE MADE, ENDING WITH
THE FINAL TEST JANUARY 26, 1973, NOT ALL OF THESE TESTS WERE RUN
WHILE FIRING JlM BRIDGER COAL, SEVERAL TEST SERIES WERE UNDERTAKEN
WITH DAVE JOHNSTON COAL IN ORDER TO CONFIRM VALID PERFORMANCE OF THE
TEST UNIT, TO MAKE CORRECTIONS IN OBSERVED TEST UNIT "SNEAKAGE" AND
TO CONSERVE JlM BRIDGER COAL FOR VITAL TESTS, Op THE 175 TEST RUNS,
128 WERE CONDUCTED WHILE UNIT 1 WAS FIRING JlM BRIDGER COAL, ALL
TESTS WERE CONSIDERED GERMANE TO THE ISSUE TO THE JlM BRIDGER ELECTRO-
STATIC PRECIPITATOR DEVELOPMENT PROGRAM,
IN CONDENSED FORM, THE COMPLETE PROGRAM EVALUATED THESE
ISSUES:
1, FLUE GAS TEMPERATURE EFFECTS - DEVELOPMENT OF PILOT
MIGRATION VELOCITIES,
2, MOISTURE CONDITIONING
3, AMMONIA CONDITIONING
4, VOLT-AMPERE RELATIONSHIPS
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5, COAL ANALYSES RELATIONSHIPS
6, DUST ANALYSES RELATIONSHIPS
THE PRIMARY ISSUE TO BE RESOLVED WAS THE DETERMINATION
OF THE REQUIRED SPECIFIC COLLECTION AREA TO ACHIEVE THE SPECIFIED
0,02 GRAINS/ACFM EMISSION GUARANTEE FOR THE FULL SCALE JlM BRIDGER
ELECTROSTATIC PRECIPITATORS,
SOMETHING NOW ABOUT THE TEST CREW AND ITS MAKE UP OF PER-
SONNEL, SF-CARBORUNDUM WAS REQUIRED BY CONTRACT TO BE RESPONSIBLE
FOR THE CONDUCT OF THE TESTING PROGRAM, PACIFIC POWER, IDAHO POWER,
BECHTEL AND OTHER OWNER SELECTED CONSULTANTS WERE INVOLVED,
A TYPICAL TEST CREW WAS MADE UP OF APPROXIMATELY 10 PEOPLE;
TWO TO TEST FOR PARTICULATES, ONE TO TAKE ORSTAT READINGS, TWO TO
WEIGH AND HANDLE THE DUST DISCHARGE. ONE LABORATORY TECHNICIAN,
TWO FOR COAL SAMPLING, ONE FOR CALCULATIONS AND ONE TO ACT AS THE
TEST CREW SUPERVISOR, NOT ALL TESTING WAS CONDUCTED ON AN AROUND-
THE-CLOCK BASIS, THE EARLY TEST RUNS AND PARTICULARLY THOSE CON-
DUCTED WHILE FIRING DAVE JOHNSTON COAL WERE UNDERTAKEN WITH ONE
SHIFT DURING THE DAY, IN THE LATER STAGES, WITH THE JlM BRIDGER
COAL SUPPLY SUBSTANTIALLY REDUCED, A 24-HOUR TESTING EFFORT PRE-
VAILED,
AT THE PEAK INTENSITY OF THIS TESTING PROGRAM, IN EXCESS
OF 30 TECHNICAL INDIVIDUALS WERE INVOLVED AT THE SITE, NOT IN-
CLUDED IN THIS COUNT WERE A NUMBER OF MAINTENANCE AND CONSTRUCTION
INDIVIDUALS, A SPECIAL REPRESENTATIVE OF THE PLANT OPERATING
STAFF MAINTAINED APPROPRIATE LIASON AND COORDINATION WITH DAILY
PLANT OPERATIONS WHICH BY NECESSITY COMMANDED THE HIGHEST PRIORITY,
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IT HAS BEEN ESTIMATED THAT THE ENTIRE TEST PROGRAM COST
IN EXCESS OF $600,000 INCLUDING THE SUPPLY AND SHIPMENT OF THE JlM
BRIDGER TEST COAL,
PILOT TEST r.ORRFIATIQN TECHNIQUE
THE SF TECHNIQUE FOR SIZING AN ELECTROSTATIC PRECIPITATOR
FROM PILOT TEST DATA IS UNIQUE,, WAS REQUIRED TO BE SPECIFIED IN AD-
VANCE OF THE PILOT TESTING EFFORT AND PERHAPS CONTROVERSIAL IN SOME
ASPECTS, THIS PRESENTATION WILL NOT DISCUSS IN DEPTH THE NUMERICAL
OR THEORETICAL DETAILS OF THE SF APPROACH TO PRECIPITATOR SIZING BUT
WILL DESCRIBE THE BASIC ELEMENTS AS APPLIED TO THE JlM BRIDGER PRE-
CIPITATOR DEVELOPMENT PROGRAM,
FOR JIM BRIDGER. PILOT UNIT TO FULL-SIZE PRECIPITATOR COR-
RELATIONS WERE MADE BY APPLICATION OF A MODIFIED FORM OF THE CONVEN-
TIONAL DEUTSCH COLLECTION EFFICIENCY FORMULA,
Q
EFF = 1 - E
WHERE: EFF = COLLECTION EFFICIENCY
A = COLLECTING SURFACE AREA
Q = FLUE GAS VOLUME FLOW RATE
WK = MIGRATION VELOCITY. CORRECTED
K = CORRECTION FACTOR
E = BASE OF NATURAL LOGARITHMS
THE UNIT K IS A CORRECTION FACTOR, DEVELOPED BY EXPERI
ENCE AND APPLIED TO ADJUST THE CALCULATION FOR THEORETICAL VS
ACTUAL PERFORMANCE DIFFERENCES, WK IS AN EFFECTIVE MIGRATION VE
LOCITY DEVELOPED FROM MIGRATION VELOCITIES MEASURED IN THE PILOT
TEST WORK AND CORRECTED BY A NUMBER OF FACTORS EVALUATING THE IN
FLUENCE OF DUST, GAS AND PHYSICAL PROPERTIES ON FULL-SIZE PRE-
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CIPITATOR PERFORMANCE,
MIGRATION VELOCITY CORRECTION FACTORS UTILIZED IN THE
JIM BRIDGER CALCULATION WERE:
1- TEMPERATURE CORRECTION - TO DEVELOP A CORRELATION
BETWEEN FLUE GAS TEMPERATURE AND REQUIRED COLLECT-
ING AREA,
2, GEOMETRIC CORRECTION - LEAKAGE AND SNEAKAGE AREA
IN THE PILOT UNIT IS LARGE IN COMPARISON TO SIMI-
LAR BOUNDARY AREAS OF THE FULL-SIZE UNIT REQUIRING
A CORRECTION FACTOR TO COMPENSATE FOR DIFFERENCES,
THIS FACTOR DEVELOPED FROM EXPERIENCE,
3, UNCERTAINTY CORRECTION - A SAFETY FACTOR DEVELOPED
BY EXPERIENCE,
4, DUST CONCENTRATION CORRECTION - A FUNCTION OF DUST
CONCENTRATION, PILOT UNIT VS FULL-SIZE UNIT, AGAIN,
A RELATIONSHIP DEVELOPED BY EXPERIENCE,
5, MOISTURE CORRECTION - A FUNCTION OF PILOT UNIT
FLUE GAS MOISTURE VS FULL-SIZE UNIT GAS MOISTURE,
RELATIONSHIP DEVELOPED IN THE PILOT TEST WORK,
PARTICLE MIGRATION VELOCITY is DEPENDENT ON CURRENT DEN-
SITY WITHIN THE PRECIPITATOR, To LIMIT LARGE CORRECTIONS IN MIGRA-
TION VELOCITY, SF ELECTS TO OPERATE THE PILOT UNIT AT THE SAME
CURRENT DENSITY AS WILL BE THE CURRENT DENSITY OF THE FULL-SIZE
UNIT, FROM EXPERIENCE, A CURRENT DENSITY FOR JlM BRIDGER WAS
ESTABLISHED AS 14,5 MICRO AMPERES PER SQUARE FOOT OF COLLECTING
SURFACE, PERTINENT PILOT TESTS WERE CONDUCTED AT THIS SAME CUR-
RENT DENSITY,
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PILOT TEST RESULTS
A CONDENSATION OF THE RESULTS OF THE PILOT TEST WORK IS
IN ORDER, THE PRINCIPAL EFFORT IS TO DESCRIBE THE SELECTION OF
THE FINAL SIZE FOR THE FULL-SCALE JlM BRIDGER PRECIPITATORS,
1. JIM BRIDGER PRECIPITATQR SIZE - THE ORIGINAL INTENT
WAS TO SET THE DAVE JOHNSTON UNIT 1 PULVERIZER COAL
GRIND TO DUPLICATE THE JlM BRIDGER SPECIFIED GRIND
OF 70% THROUGH A 200 MESH SCREEN, As THE PROGRAM
PROGRESSED, THIS TECHNIQUE PROVED NOT ONLY DIFFI-
CULT TO MAINTAIN, BUT THERE WAS EVIDENCE THAT COAL
GRIND SIGNIFICANTLY INFLUENCED PILOT PRECIPITATOR
PERFORMANCE, As THE COAL GRIND BECAME MORE COARSE,
PILOT PRECIPITATOR PERFORMANCE INCREASED,
EVEN AFTER A CAREFUL ANALYSIS OF ALL TEST DATA
THE ISSUE OF COAL GRIND VS ASH PARTICLE SIZE WAS NOT
CLEARLY DEFINED, HOWEVER, THERE WAS CLEAR EVIDENCE
THAT CORRELATION OF A VERY FINE COAL GRIND WOULD PRE-
DICT A VERY LARGE FULL-SIZE JlM BRIDGER PRECIPITATOR,
CORRELATION OF THOSE TESTS WITH A COARSER COAL GRIND,
APPROXIMATING THE JlM BRIDGER SPECIFIED 70% THROUGH
A 200 MESH SCREEN, SUGGESTED A PRECIPITATOR SIZED WITH
A SPECIFIC COLLECTING AREA OF 375 (375 SQ. FT./1000
ACFM),
THE SERIES OF PILOT TEST RUNS IDENTIFIED WITH
THE EXTREMELY FINE COAL PULVERIZER GRIND WAS DESIG-
NATED AS SERIES A, THOSE TEST RUNS CONDUCTED WITH A
PULVERIZER GRIND APPROXIMATING THE 70% THROUGH A 200
MESH SCREEN (JlM BRIDGER SPECIFICATION) WERE DESIGNATED
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AS SERIES C, A JIM BRIDGER PRECIPITATOR SIZED ON
SERIES A DATA WOULD EQUATE TO AN SCA OF 586, As
PREVIOUSLY MENTIONED, A JlM BRIDGER PRECIPITATOR
SIZED ON THE SERIES C TESTS WOULD EQUATE TO AN SCA
OF 375,
SF DEVELOPED THEIR RECOMMENDATION BASED ON THIS
RATIONAL;
A) A SIZE BASED ON SERIES C TESTS IS THE
MOST REPRESENTATIVE OF FULL-SCALE JlM
BRIDGER CIRCUMSTANCE,
B) IT IS POSSIBLE, BUT NOT LIKELY, THAT
CONDITIONS SUPPORTING THE FINE GRIND,
SERIES A TESTS COULD PREVAIL AT SOME
TIME IN THE OPERATION OF THE JlM
BRIDGER PLANT,
C) A RESERVE, EMPTY COLLECTING SPACE OF
20% HAS BEEN SPECIFIED IN THE DESIGN,
D) ALLOCATION OF THIS 20% RESERVE SPACE
TO PERFORMANCE PROTECTION SHOULD THE
EXTREME CIRCUMSTANCES OF SERIES A CON-
DITIONS PREVAIL, SUGGESTS THAT THE PRU-
DUENT FULL-SIZE DESIGN BE 586 SCA (SERIES
A, FINE GRIND REQUIREMENT) LESS THE 20%
RESERVE SPACE MARGIN,
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E) THIS CALCULATION RESULTS IN A RECOMMENDED
JIM BRIDGER PRECIPITATOR SIZE OF 488 SCA,
F) THE SIZE OF 488 SCA INCLUDES A MARGIN
SUFFICIENT TO COMPENSATE FOR THE "BAND"
OF RESULTS THAT WOULD BE THE COMPLETE
SERIES C TESTS (TESTS REPRESENTING THE
MOST LIKELY JlM BRIDGER CIRCUMSTANCES),
THE OWNERS AND THEIR ENGINEERS CONCLUDED THAT
THIS WAS A CONSERVATIVE. THOUGH NOT EXTRAVAGANT
CONCLUSION AND THE JlM BRIDGER UNIT 1, 2 AND 3
PRECIPITATORS WERE CAST INTO FINAL MANUFACTURE
BASED ON A SIZE EQUAL TO 488 SQ, FT, OF COLLECTING
SURFACE FOR EACH LOGO ACFM,
2, MOISTURE CONDITIONING - RESULTS INDICATED PRACTICALLY
NO SUSCEPTIBILITY TO MOISTURE CONDITIONING,
3, AMMONIA CONDITIONING - RESULTS SUGGESTED A VERY
SMALL INCREASE IN PRECIPITATOR PERFORMANCE WITH
AMMONIA CONDITIONING IS POSSIBLE,
4, HIGH TEMPERATURE RESULTS - THE TEST WORK EVALUATING
HIGH TEMPERATURE EFFECTS WAS LIMITED, THIS DATA,
PLUS DATA FROM AUSTRALIAN TEST WORK DID NOT SUGGEST
AN ADVANTAGE TO CONSIDERATION OF A "HOT" PRECIPITATOR
DESIGN,
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FIRST PERFORMANCE TESTS
UNIT 1 WAS THE FIRST TO UNDERGO TESTING FOR PERFORMANCE,
TEST A WAS CONDUCTED DURING THE PERIOD OF APRIL 14-30, 1975, UNIT 1
HAD BEEN IN OPERATION IN EXCESS OF 4,000 HOURS WITH AT LEAST THE
REQUIRED MINIMUM OF 2,000 HOURS WITHOUT ADJUSTMENT OR MODIFICATION,
TESTING PRODUCED COLLECTION EFFICIENCIES RANGING FROM 99,5 TO 99,6%
(GUARANTEE - 99,33%) WITH ALL GAS PATHS IN SERVICE, WITH ONE GAS
PATH OUT OF SERVICE EFFICIENCIES OF 99,4% WERE RECORDED, WELL WITH-
IN THE SPECIFIED 98,73% UNDER THIS CONDITION, PARTICULATE EMISSIONS
IN EVERY CASE WERE WELL WITHIN THE GUARANTEED 0,02 GRAINS/ACF,
TEST B FOR UNIT 1 AND TEST A FOR UNIT 2 WERE SCHEDULED FOR
FEBRUARY, 1976 AND MARCH, 1976, IN FEBRUARY, 1976, PRELIMINARY
TESTS OF UNIT 1 INDICATED PERFORMANCE OF QUESTIONABLE QUALITY, As
A RESULT, FURTHER OFFICIAL PERFORMANCE TESTS OF UNIT 1 AND 2 WERE
DEFERRED PENDING TECHNICAL INVESTIGATION TO DETERMINE THE CAUSE OF
THE REDUCED PRECIPITATOR CAPABILITY, THESE INVESTIGATIONS ARE CON-
TINUING AS OF THIS DATE,
BASED ON EARLIER EXPERIENCE, IT WAS SUSPECTED THAT THE
POOR PRECIPITATOR PERFORMANCE, INDICATED BY UNSATISFACTORY ELECTRICAL
READINGS AND BACK CORONA, WAS CAUSED BY PHENOMENA LINKED TO ELECTRODE
DEPOSITS FROM OIL FIRING DURING UNIT START-UP, SATISFACTORY PER-
FORMANCE WOULD RETURN WITHIN A FEW DAYS FOLLOWING START-UP,
CORRECTIVE AND INVESTIGATIVE MEASURES NOW ACTIVE FOR SOLU-
TION OF THIS PROBLEM ARE:
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A) REMOVAL OF A REPRESENTATIVE SAMPLE OF ELECTRODES
FROM UNIT 1 AND UNIT 2 AND SUBJECT THEM TO LAB-
ORATORY EXAMINATION TO DETERMINE THE NATURE OF
THE COATING,
B) CLEAN EMITTING ELECTRODES IN ONE ELECTRICAL SECTION,
CLEAN COLLECTING ELECTRODES IN A SECOND ELECTRICAL
SECTION, AND CLEAN BOTH EMITTING AND COLLECTING ELEC-
TRODES IN A THIRD SECTION, OBSERVE ELECTRICAL AND
OPERATING CHARACTERISTICS OF EACH SECTION BEFORE AND
AFTER CLEANING,
C) AN IMPROVED OPERATING PROCEDURE HAS BEEN ISSUED TO
COVER UNIT START-UP, SHUTDOWN AND LOW-LOAD OPERATION,
SF-CARBORUNDUM ARE CONFIDENT THAT A COMPLETE AND LASTING
SOLUTION WILL BE DEVELOPED, As OF THIS VERY DATE, UNIT 1 AND UNIT 2
ARE AVAILABLE FOR FULL LOAD OPERATION WITHOUT PRECIPITATOR IMPOSED
LIMITATIONS,
CONCLUSION
PACIFIC POWER, IDAHO POWER, THEIR ENGINEERS AND SF-CAR-
BORUNDUM CONSIDER THE ENGINEERING EFFORT OF THE JlM BRIDGER ELECTRO-
STATIC PRECIPITATOR DEVELOPMENT PROGRAM TO BE THE MOST THOROUGH AND
TECHNICALLY SOUND OF ANY UNITED STATES PRECIPITATOR INSTALLATION,
HUNDREDS OF INDIVIDUALS HAVE BEEN INVOLVED OVER A PERIOD IN EXCESS OF
SEVEN YEARS, THE JlM BRIDGER ELECTROSTATIC PRECIPITATORS PROVE
THAT HIGH EFFICIENCY COLLECTION OF LOW SULFUR, WESTERN COAL ASH IS
POSSIBLE AND PRACTICAL, "COLD SIDE,"
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ACKNOWLEDGEMENTS
MANY INDIVIDUALS FROM SEVERAL ORGANIZATIONS CONTRIBUTED
SIGNIFICANTLY TO THE SUCCESS OF THE JlM BRIDGER ELECTROSTATIC ?RE-
CIPITATOR DEVELOPMENT PROGRAM, THE AUTHOR HAS CHOSEN TO LIST, FOR
PURPOSES OF PERSONAL RECOGNITION, THE NAMES OF MOST OF THOSE RE-
SPONSIBLE FOR THIS MAJOR ENGINEERING ACCOMPLISHMENT, OMISSIONS
THERE ARE, FOR THIS HUMAN DEFICIENCY THE AUTHOR APOLOGIZES,
THERE HAS BEEN NO CONSCIOUS INTENT TO OMIT APPROPRIATE RECOGNITION,
PACIFIC POWER & LIGHT COMPANY
G, L, BEARD
P, G, HUMPHREYS
A, H, SEEKAMP
BECHTEL CORPORATION
A, L, CAHN
VIVIAN ESTCOURT
ROBERT STEINER
WILLIAM LANE
TONY HENRY
SOUTHERN RESEARCH INSTITUTE
SABERT OGLESBY
GRADY NICHOLS
IDAHO POWER COMPANY
G, J, HALL
GOMER CONDIT
RALPH CLEMENTS
RUSSELL STEWART
AB SVENSKA FLAKTFABRIKEN
SIGVARD MATTS
BENGT HALENE
BENGT LINDHOLM
CONSULTANTS
HARRY J, WHITE
KENNETH WATSON
169
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OPERATING EXPERIENCE WITH ESP CONDITIONING IN RELATION TO
AN ELECTROSTATIC PRECIPITATOR UPGRADING PROGRAM
Scott H. Cragle
Pennsylvania Power & Light Company
Allentown, Pennsylvania
171
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OPERATING EXPERIENCE WITH ESP CONDITIONING IN RELATION TO AN
ELECTROSTATIC PRECIPITATOR UPGRADING PROGRAM
SCOTT H. ORACLE. PENNSYLVANIA POWER & LIGHT CO.
Abstract
This report summarizes the major areas of an ongoing electrostatic
precipitator (ESP) upgrading program at Pennsylvania Power and Light Co.
Particular emphasis is placed on the area of flue gas conditioning
including eighteen months of operating experience with Apollo Chemical
Corp. products, a planned 803 injection trial, and other potential
agents.
Results of ongoing trials of Apollo products at Montour SES on two 750
MW bituminous coal fired units are presented. Performance data, operating
problems and cost information are covered. Eight chemical formulations
have been injected into various combinations of four locations in
efforts to improve product performance, solve air heater plugging problems
and improve distribution of the product on the fly ash. A significant
reduction of emissions on a difficult fly ash has been noted although
only half as great a reduction as originally expected based on tests run
with a good collecting fly ash. Air heater plugging problems have been
eliminated.
The report also covers some of the practical considerations and potential
problems that can occur in evaluating proprietary additives.
Also, the report briefly summarizes projects on six of the Company's
ESP's including work in the following areas: fuel quality effects; flue
gas conditioning, rappers; voltage controls; sectionalizing; gas flow
distribution; operating and maintenance and additional dust collector
capacity.
May 11, 1976
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Pennsylvania Power & Light Company (PL) began a series of efforts to
upgrade its existing electrostatic precipitator (ESP) capacity in mid-
1973. At that time it had become apparent that the particulate emissions
from the two new boilers at PL's Montour station did not meet consistently
acceptable levels. The Montour station has two 750 Mw units with pulver-
ized bituminous coal boilers. The ESP's were built in a chevron configu-
ration and designed for a 99.4% efficiency for coals with a 1.5% to 3.0%
sulfur and an efficiency of 99% for coals with a 1.0% to 1.5% sulfur
content. More detailed information is shown in Table 1.
Initial Performance Tests (9/73-11/73)
An initial series of efficiency tests on Unit 1 resulted in efficiencies
from 90.9% to 95.4%. A compliance test by the Pennsylvania Department
of Environmental Resources showed emissions 9 to 10 times the emission
limit of 0.1 Ib./lO^ BTU. Another series of outlet tests was run to
determine the effect of the sulfur content of the different fuels
burned at the plant and also to determine the feasibility of using the
two recently installed Lear Siegler opacity monitors as indicators of
ESP performance.
The tests showed a relation between optical density and dust flow from
the precipitator (Figure 1). As a result of this test, the monitors
have become an integral tool in monitoring precipitator performance.
The tests showed that the ESP's performed within regulation with certain
high sulfur fuels, but performed poorly with the low sulfur fuel from
our Greenwich Mine, a major fuel source for the plant. The test points
are numbered on Figure 1 in the order that the tests were performed.
Tests 1 through 4 show that we didn't observe the full effect of a low
sulfur coal until it had been burned for several consecutive days. This
was due to the residual fly ash in the ESP. Tests 5 through 8 indicate
that the good effects from a high sulfur coal were not noticed on the
first day after burning a low sulfur coal, but only on the second and
third day. Test 8 was a lower sulfur coal, but high levels of 863 were
reported during the test, possibly due to a high vanadium content in the
ash.
EXPERIENCE WITH THE APOLLO FLY ASH CONDITIONING AGENT
Based on the test results, we viewed flue gas treatment to be the major
solution to the ESP performance problem at Montour. Ongoing work done
in other areas is covered briefly later in the paper. After a short
study (1/74) of conditioning alternatives, we chose Apollo (2/74) on the
basis of a low first cost and a short three month lead time for installa-
tion. We thought that Apollo would be a quick, low cost demonstration
of the effectiveness of fly ash treatment on our problem fuels. Other
alternatives required large capital investments and nearly a one year
lead time to set up any demonstration.
Our initial plans were for the trial to last three months on one 750 MW
unit. The cost projection was $150,000 (20<: Kw) to install the system
and $90,000 for three months of chemical supply. The system has now
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been expanded to both Montour units for a cumulative capital cost of
about $400,000 ($200,000 per unit or 27<:/Kw). The cost of chemicals to
treat both units is $650,000 per year (0.0067C/KW HR).
Performance LPA-40 <•& (9/74-1/75)
All conclusions about ESP performance in this report are based on data
taken by PL personnel.
The Apollo product used in the initial testing was LPA-40. This was a
sulfamic acid (PL analysis) based multicomponent product that is injected
into the flue gas upstream of the boiler economizer. Through breakdown
to NH3 and 863 it is intended to affect flyash resistivity, agglomeration,
and space charge in the flue gas.
Because of a 3 month strike at PL, initial operation of the Apollo
system was delayed until September, 1974. We quickly discovered that we
were limited to injecting less than 0.1 Gal/Ton of coal (25 GPH) because
of increased pressure drop across the Lungstrom air heaters. A series
of emission tests were run at this feed rate, and two opacity monitors
made continuous observations. Table 2 shows the improvement that occurred
with this Apollo additive using our regulated emission limit (0.1 LB/10
BTU) as a base line (Ix). Emissions were reduced by 30-45%.
Internal inspection of the air heater showed no adverse effects related
to the high pressure drop. Whenever treatment was shut off for a few
days the pressure drop would return to normal. Most likely, the plugging
was due to unvaporized product on the air heater or recombination of
vaporized product to ammonium bisulfate.
During the next phase of the test program, our efforts were centered on
reducing the air heater plugging to allow higher treatment rates. These
five approaches were taken (in chronological order):
1. Product Atomization (10/74) - air atomized nozzles were added to
the system to assist in faster breakdown of the product. No major
improvement resulted from this change.
2. Flow Distribution Tests (11/74) - We suspected uneven treatment of
the flue gases might be a significant problem because of the location
of the injection ports (Figure 3, Phase I). Tests run at the inlet
of the air heater showed evidence that 80% of the gas stream was
getting uniform treatment and 20% of the gas stream was receiving
lower than normal concentrations of the product. We are not certain
if this situation improves upstream of the precipitator. Even if
it is a significant problem, no ready solution is available to us
to allow more even treatment of the flue gas.
3. LPA - 40, Neutral Formulation (11/74) - Apollo changed the major
component of their product from sulfamic acid to ammonium sulfate
(results of PL analysis) indicating that the new product decomposed
faster. No significant improvement in performance resulted from
the change of product.
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4. Flue Gas Temperature at the Point of Injection (1/75-4/75) - Flue
gas temperature at the point of injection is apparently critical to
the application of LPA - 40. Injecting the product into the flue
gas stream at too high a temperature results in the chemical combina-
tion of the products with the fly ash. Also, production of S03 at
high temperatures is not useful because the 863 will convert to SC>2
by the nature of the S02 - 803 chemical equilibrium. Injecting at
too low a temperature, the product does not decompose as completely
and as rapidly as necessary and air heater plugging results. Our
problem was low temperature at the injection point.
At the start of the program, Apollo indicated that the ideal tempera-
ture of injection was 800-1000° F. Following their change to the
neutral formulation, Apollo indicated that they had seen good
results at injection temperatures above 1200° F and that temperatures
as high as 1500° F would be optimum for their product. Currently
they believe the optimum temperature is 1250-1300° F.
Figure 3, a boiler cross section, shows the locations we have
injected LPA-40 with the associated flue gas temperatures for both
full boiler load and half boiler load. Note how the temperature
varies with load.
Phase I (950-750° F) - This is the original penthouse location
where we have had the best results so far. Injection here at the
rate of 0.1 gal/ton resulted in air heater plugging.
Phase II (2300-2200° F) - Injection above the firing zone at high
rates (0.4 gal/ton) had no effect. No effect was seen on precipi-
tator performance and no air heater plugging occurred.
Phase III (1650-1250° F, upper ports; 1750-1350° F, lower ports) -
Injection into the side of the boiler into the superheat area
showed no effect at full load. No effect was observed on precipi-
tator performance and no air heater plugging occurred.
Phase IV - The problem of finding an injection point with the right
temperature caused the next phase of testing. Product was injected
through the Phase I location during periods of high boiler load and
injected into the Phase III location during periods of low boiler
load in an attempt to keep the injection temperatures acceptable.
This approach was intended to reduce the chance of air heater
plugging while providing continuous treatment of the fly ash at all
boiler loads. No improvement in precipitator performance or in the
plugging problem resulted from this approach.
5. Revised Apollo Chemical Formulations (5/75 - Present)
LPA 401, Catalytic Formulation - This is the LPA-40 formulation
with a metal added to cause faster breakdown. This formulation
caused more air heater plugging problems than the normal LPA 40.
175
-------
LPA 402, Agglomerator Formulation - This formulation contains an
organic compound in trace quantities. Results with this have been
good. Air heater plugging has been eliminated and ESP performance
has been equivalent to LPA-40.
LPA 403, Combination of LPA 401 and 402 - No trial was run due to
the problems with 401.
LPA 404, Non-Ammoniated formulation 803 - release agent - This was
designed to produce 863 only. Results with this chemical showed no
improvements in 803 concentration in the flue gas or improvement of
the in-situ resistivity of the fly ash.
LPA 405, Another Non-Ammoniated formulation - This is similar to
404, but has not been tried due to handling concerns resulting from
its low pH of 1.0.
LPA 402A, Agglomerator Formulation - this is similar to 402 only
with a different organic compound. It is more volatile than 402.
Results with this chemical seem identical to those with 402 and it
is the formulation presently being used.
PERFORMANCE - LPA 402 AND 402A
Table 3 shows results for Unit 2 with LPA 402A in the same format as
Table 2 shows results for LPA 40 on Unit 1. The Unit 2 precipitator in
general performs better than Unit 1 due to some variations in physical
flow arrangement. The results indicate that 402 and 402A are similar
performers to LPA 40 in the Montour boilers when injected at the penthouse
location.
During April and May, 1976, a series of ESP efficiency tests are being
run on Montour 2. The tests are being run with high sulfur coal, low
sulfur coal, and low sulfur coal conditioned with LPA-402A.
Resistivity
Figure 4 plots in-situ resistivity taken by Apollo and Southern Research
Institute using a point to plane probe. These were taken at various
times during our test program. The scatter of the data is evident in
this graph and, as a result, it is difficult to make conclusive statements
about it. By looking at selected points, we have noticed certain trends,
however. Increased sulfur content of the fuel has the expected effect
of lowering the resistivity. Lower temperatures seem to produce lower
resistivities as expected. The school of thought at PL, which has not
been fully investigated, is that lower flue gas temperatures may help
the performance of the Apollo flue gas treatment by lowering the resis-
tivity. In general, for a low sulfur fuel, it seems that the Apollo LPA
402 treatment lowers the resistivity about one half of an order of
magnitude from the low IQH OHM-CM to the high 1010 OHM-CM.
176
-------
At this point our data on available S03 in the flue gas is incomplete.
Apollo picks up most of the sulfur compounds in hopper samples of fly
ash.
It is difficult to draw conclusions from either resistivity analysis or
303 analysis because of the variable sulfur content of fuel burned at
Montour. Monitoring minute-to-minute changes in the sulfur content of
the coal is difficult.
General Observations
These are some suggestions that will be helpful to a company considering
a test program with a proprietary chemical flue gas conditioning additive.
1. PL has found that trials planned to be quick and inexpensive can
turn out to be more involved than initially anticipated. If you
are using expediency for justification of a choice of a certain
alternative, be certain that your trial program will be expedient.
2. Prior to the start of a test program, define the ground rules for
the evaluation program.
a. What is the expected mechanism or action of the additive?
b. What parameters will be used to measure the performance of the
additive (resistivity, 803 concentration, NH3 concentration,
fly ash analysis, etc.)? Ideally you will set up a program to
measure the additives effect in more ways than just precipi-
tator performance. In this way you can determine if changes
in ESP operation are due to the additive or other parameters.
Also it will be easier to troubleshoot any lack of performance.
3. Get a committment from the vendor prior to the program as to the
"ideal" temperature and location for injection of the product. If
possible, investigate the temperature (and its variation with
boiler load) at proposed injection locations before installing
ports and starting the trial.
4. Be sure your injection system is capable of adequate distribution
of the additive in the flue gas. Cutting corners in this area may
well limit the success of your trial.
5. Use great care in installing the injection system and, if possible,
observe its operation (injecting water) from inside the boiler
during an outage. An unfortunate error in our installation resulted
in 16 economizer leaks and a six day unit outage.
6. To demonstrate the capabilities of the ESP, arrange to burn some
higher sulfur fuels and test the units under ideal resistivity
conditions.
177
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7. If your system is installed during an annual outage, allow the
boiler to become dirty before testing emissions and drawing conclu-
sions on the performance of an additive product. PL's Chem Lab
feels that a clean boiler after an outage acts as a catalytic bed
in the conversion of S02 to 803. During the period following an
outage the higher levels of 803 result in somewhat better precipi-
tator performance. Also the precipitator is usually cleaner follow-
ing an outage.
Also, note any modifications that are made during an outage to the
boiler such as addition of economizer of other tube surface to the
boiler. These may affect the amount of S03 formed by catalytic
action or they may affect the temperature (thus resistivity) of the
fly ash.
8. Often, ESP performance improvements attributed to additives are
due, in part, to changes in other parameters. Data in this report
was taken by PL under carefully controlled conditions to minimize
this problem.
a. During both baseline and additive testing, maintain uniform
operating conditions in the boiler (coal quality, excess air,
gas temperature, mill classifier setting, etc.). Variations
in these parameters can cloud your results.
Continuous reading S02 and opacity monitors have been a useful
tool in quickly determining relative changes in these parameters.
Absolute values from these instruments depend on monitor
reliability and maintenance.
b. Familiarize yourself with the operation of the ESP. Use
consistent set points for rappers and automatic voltage controls
(at sparking threshold) during all testing. Observe the power
levels during emission testing to detect any abnormal conditions
that may invalidate results of that particular test.
9. At Montour we have found that changes in the type of coal burned in
the boiler do not have the full effect on the precipitator for a
number of days. Be careful not to draw positive or negative conclu-
sions too soon after adjusting any operating parameter.
10. By following these guidelines and being open-minded towards alternate
chemical formulations and injection locations, recommended by the
supplier, PL has benefited with improved performance. Try to
identify all possible approaches before the start of the trial.
178
-------
Summary
The original plan for the Apollo trial was a quick, simple, and inexpen-
sive demonstration with hopes for impressive results. Although the
trial turned out to be long, involved, and costly with less than spec-
tacular results, day to day emissions at the Montour plant have been
reduced significantly through the use of Apollo.
PL'S PARTICULATE COMPLIANCE PROGRAM (1/75-Present)
PL's particulate compliance program is a series of trials in 8 investi-
gative areas on six of PL's precipitators.
Starting with the formation of a 5 man Air Quality Project Team in
January, 1975, PL started a two year investigative program to bring its
units into full day-to-day compliance with Pennsylvania air quality
regulations. In the past, PL had attempted to have ESP's comply by
making modifications that seemed most promising. Often, too many changes
were made to a unit at one time. This made it difficult to determine
what, if any, improvement was gained, and if the changes could be applied
to other units with similar problems. Previous work done on Montour No.
1 is a prime example. This work included:
A. Modification of rapper shafts and the rapping program.
B. Adjustment of electrode control set points.
C. Trials with updated high voltage controls on sections of the
precipitator.
D. Installation of turning vanes in an attempt to improve gas
flow distribution.
E. Experimentation with adding weight to the electrode wire
tensioning weights.
F. Collecting plate realignment and wire support frame stabiliza-
tion.
G. Blending of high sulfur coal with low sulfur coal at the mine
in order to raise the overall sulfur level of the coal.
H. Adjustments to operating parameters such as excess oxygen and
flue gas temperature.
I. Apollo fly ash conditioning trial.
179
-------
The combination of concurrent changes made it extremely difficult, if
not impossible, to reasonably evaluate the effect of any single change.
The intent of the present program is to produce as conclusive a result
as possible about the effects of various modifications that will be
made- Table 1 lists the design parameters of the six precipitators.
This is a brief summary of the work in each investigative area.
1. Fuel Quality Effect
The root of the problem with precipitators is often the type of
fuel burned. PL uses a wide variety of coals ranging from 10% to
30% ash and from 0.8% to 2.5% sulfur. A study is being done on all
of the six units to determine the effects of the various fuels on
the dust collector performance. At the end of the study PL will
consider the economics and practicality of achieving compliance by
controlling the quality of fuel shipped to our plants. This approach
is possible by using the capabilities of cleaning plants at our
mines and coal blending.
2. Flue Gas Conditioning
A. Apollo - Use of the Apollo product will continue unless a more
effective gas conditioning system is demonstrated or additional
collecting capacity is installed at Montour. PL has an agree-
ment with Pennsylvania regulatory authorities to continue any
reasonable activity that lowers emissions at Montour.
B. 803 injection, Brunner Island Unit 1 - a Wahlco 863 injection
utilizing a liquid SC>2 vaporizer system will go into trial
service on this unit during August, 1976. A series of tests
with the 803 injection will be run to get data similar to that
in Tables 2 and 3. With this data, we hope to assess how
effective such a system would be at Montour in comparison to
the Apollo system. The total capital cost of the trial system
on this 350 Mw boiler is $445,000 ($1.28/Kw) and the expected
cost for liquid S02 during the six month trial period will be
about $70,000 (.0058<:/KwHr). Any permanent installation of
this system will include a sulfur burner resulting in
$400,000 additional capital cost and a reduction of
chemical cost by a factor of 5 or 6. We expect to have results
available from this trial by October 1, 1976.
C. Trimex, Martins Creek 2 - PL's Research and Development group
is sponsoring this trial in an effort to demonstrate the
feasibility of reducing S02 emissions from the boiler. An
offshoot of this trial may be conditioning of the fly ash.
After almost five months of testing our Chem Lab reports that
the Trimex chemical has shown no affect on the S02 emissions
or on the performance of the ESP's.
180
-------
D. Other Additives - Although no additional trials are planned,
PL is investigating the use of triethylamine, concentrated
atnmoniacal liquor, Betz proprietary additives, and Nalco
proprietary additives. A trial with Nalco's experimental
catalytic additive was cancelled after we had unsuccessful
laboratory trials.
3. Rappers
Internal inspection of many of PL's precipitators equipped with
vibrating rappers have shown heavy dust buildups on the plates. On
one of the two Montour Unit 2 precipitators, the vibrating rappers
have been replaced with impact rappers. The new rappers'
performance is being evaluated.
4. Voltage Controls
Because of the wide variety of fuels burned at PL's plants, we have
found that the automatic voltage controls on many of the units are
not capable of following the varying conditions in the EPS's.
Eight thyristor controls manufactured by L.L. Little are being
installed on one of the two precipitators at Brunner Island Unit 3.
We hope to evaluate any improvement in the efficiency of the pre-
clpitator, any reduction in burnout of wires, and the general
ability of the control to follow changes in operating conditions.
5. Sectionalizlng
The precipitators at Martins Creek have a relatively high degree of
sectionalization and offer a certain amount of flexibility in
allowing quick modification of their sectionalizing arrangement.
Experiments are being run with varying amounts of plate area per TR
set to determine what possible benefit may be gained by increasing
the sectionalization of the precipitators on other PL units and
future units. This work is in progress at present.
6. Gas Flow Distribution
The Brunner Island 3 precipitator and Montour 1 and 2 precipitators
are built with a chevron inlet configuration. There is evidence
that the flow distribution is not uniform over the face of the
precipitator. Field tests have been run to determine the flow
distribution and a model study is being conducted to determine if
improvement is possible. If modifications are made, efficiency
tests before and after the modifications will be run to determine
the effect of improvement in flow distribution.
7* Operating and Maintenance
A complete review of operating and maintenance procedures is being
conducted throughout the PL system. Since the precipitators are
181
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not a megawatt producing piece of equipment, historically they have
received less attention than other equipment. We are reviewing the
impact of operating parameters and maintenance on the performance
of the precipitators with our plant personnel.
8. Additional Collector Capacity
At the conclusion in early 1977, should the efforts in the seven
previously listed investigational areas fail to bring Brunner
Island 1 and 3 and Montour 1 and 2 into compliance (Martins Creek 1
and 2 are in compliance) with air quality regulation, additional
collector capacity will be installed. The ground work in this area
is being done at present to allow for the shortest lead time in
purchasing and constructing any additional capacity that may be
necessary. The alternatives that are being studied are hot precipi-
tators, cold precipitators, fabric filters, and particulate scrubbers.
Scott Cragle, Project Engineer
Air Quality Project Team
Pennsylvania Power & Light Co.
SHC:NGD
069616
182
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TABLE #1 DESIGN DATA ON PRECIPITATORS INCLUDED IN
PL'S PARTICULATE COMPLIANCE PROGRAM
CD
OJ
Unit
Brunner Island #1
Brunner Island #3
Montour #1
Montour #2
Martins Creek #1
Martins Creek #2
Ratine
MW
350
750
750
750
160
160
Research
Cottrell (1961)
Buell (1965)
Western (1969)
Western (1972)
Western (1973)
Buell (1968)
Buell (1968)
Flow
ACFMx 103
550
550
2,300
2,300
2,300
600
600
SCA
Ft2
103 ACFM
180
180
204
204
204
274
274
Face
Velocity
Ft
Sec
4.2
5.2
5.3
5.3
5.3
4.3
4.3
Treatment
Time
Sec
4.5
3.5
4.6
4.6
4.6
6.2
6.2
Typical
%S
1.5-2.5
1.5-2.5
1.5-2.5
0.8-2.5
0.8-2.5
2.0-2.5
2.0-2.5
Fuels
%Ash
12-25
12-25
12-25
12-30
12-30
12-30
12-30
-------
TABLE #2
MONTOUR
H-
00
Fuel
Greenwich
Greenwich
Oneida
Oneida
Ash
18%
18%
18%
14%
LPA-40 PERFORMANCE
Sulfur LPA-40
1.0%
1.0% 0.1 gal/ton
2.5%
2.5%
Emissions
6x - 9x
4x - 5x
Ix - 2x
Ix (Base)
-------
TABLE #3 MONT OUR »2
LPA-402A PERFORMANCE
Fuel Ash Sulfur LPA-402A Emissions
Greenwich 20% 1.3% 4x - 12x
Greenwich 0.1 gal/ton 1.5x - 2x
Oneida 12% 2.4% 0.6x
-------
00
H
tu
Q
u
a.
O
FIGURE I
OPTICAL DENSITY VS. GRAIN LOADING
MQNTOUR UNIT 1, NOV. 1973
.7
.6
.5
A
.3
.2
.1
o.o
ALLOWABLE EMliSSION RATE
Ato
• 3
0.05
QIO
0.15
0.20
0.25
0.30
GRAIN LOAOING(GRAiNs/Fi*)
HLOW SULFUR A-MEDIUM SULFUR B-H1GH SULFUR
0.35
-------
FIGURE 2.
APOLLO INJECTION LOCATIONS
(TEMP. FULL LOAD - TEMP. HALF LCW?)
PHASE III
(1750-1250F)
PHASE II
(2300-2200F)—
PHASE I
(950-750F)
TO
PRECIP.
187
-------
FIGURE 3.
SELECTED RESISTIVITY DATA
WITH AND WITHOUT LPA- 4O2 CONDITIONING
IOB
10"
10
10
X
0
260 270 280 290 300 310
TEMPERATURE
(°F AT MEASURMENT POINT)
•-HIGH SULFUR FUEL - NO CONDITIONING
V + X- LOW SULFUR FUEL- NO CONDITIONING
® $ ©- LOW SULFUR FUEL - WITH LPA 402
320
188
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AUSTRALIAN EXPERIENCE WITH FLUE GAS CONDITIONING
Kenneth S. Watson
New South Wales, Australia
189
-------
SUMMARY
Australian black coal fired in power stations usually contains less
than 1% sulphur and has 15 to 30% ash. Target emission levels for new
plant are often better than 0.1 gms/m3 at NTP requiring collection efficien-
cies of 99.5%. Highly resistive fly ash predicates specific collecting areas
in excess of 80 m2/m3/s (400 ft2/1000 cfm) for cold precipitators. Early
precipitators frequently had specific collecting areas of less than 50 m2/
m3/s and performed badly. Gas conditioning has been regularly used to improve
collection efficiencies since the middle 60's and to date has been provided
on 26 precipitators on boilers from 25 to 500 MW capacity and totalling over
4,000 MW.
Steam and/or water temperature modification, P205, NH3 and $63 were
used experimentally from 1955 onwards. Sulphuric acid was first used opera-
tionally in 1964 and anhydrous ammonia or an ammoniacal steelworks waste
product have been used since 1965. Triethyl amine was first used in 1975 and
is currently being evaluated.
The effect of reagents varies from station to station and is also
dependent on precipitator type and reagent concentration. Considerable care
must be given to the injection systems to allow proper mixing and sufficient
residence time. Improvements of migration velocity in excess of 300% have
been attained with S03/H2S04 and in excess of 200% with ammonia, which is
the preferred reagent because of relative safety and convenience. Triethyl-
amine appears to be even more effective than SO.,.
Comparatively few operational problems have been reported and most can
be traced to over-conditioning or to poor distribution. Capital costs range
from 0.05 to 0.12 $A/KW and operating costs from 0.01 to 0.05 mils/KW-hr
depending on the reagent and injection concentration.
190
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AUSTRALIAH BXPERIgtjCE VflTH FLUB GAS COKDITIOHDIG
Kenneth S. Watson *
1.0 Introduction
Australia is approximately the same area as the continental United States
and comprises six States and the Northern Territory (Figure 1). Of the
total population of 13 million, 90$ occupy the south eastern corner of
the continent and the island State of Tasmania. In all States (except
Tasmania) coal fired steam power stations provide the majority of elec-
trical energy and will continue to do so in the foreseeable future as
there are considerable proven reserves of suitable coal.
Hew South Wales has tha largest population of any State (<3. 500,000) and
approximately 9Gf° of -its energy is generated in thermal stations j the
annual coal consumption beinj; 9 million tonnes (1975), Relatively high
population densities and the degree of industrialisation combined to make
Hew South Wales and Victoria the 'two States where effective Clean Air
legislation was first introduced about 1 5 years ago.
Although Australian coals are usually low in sulphur (often less than 1?°)
only the black coals of New South Wales and Queensland are known to produce
fly ashes which can be difficult to collect in cold precipitators. The
fly ash from Victorian lignites and the sub bituminous South Australian
coals can be effectively handled by precipitators having specific collect-
ing areas of 40 to 60 mS/m-Vs whereas the black coal plants require up to
or more additional plate area.
This paper reviews the application of flue gas conditioning as an aid to
precipitator performance in Hew South Wales with a passing reference to
Queensland where, to date, it has been used operationally at one plant.
2.0 Coal quality
The locations of coal sources and principal power stations are shown in
figure 1. At this time, generation in the Sydney area is negligible, the
comparatively small quantity of coal used there coming from within the
northern or western coalfields whereas 20 years ago the Metropolitan
stations imported coal from a large number of mines located in the several
coal bearing areas. Table 1 summarises the characteristics of coal
supplied to the major stations.
3.0 Fly Ash Characteristics
Table 2 gives typical chemical analyses of some Australian fly ashes, each
ash being related to its parent coal by name. Considerable variability of
constituents is found and therefore the stated values should bo regarded
as typical rather than absolute. Unless otherwise stated, any resistivit-
ies quoted are determined by Kevatron equipment, at, or about, tha normal
precipitator operating temperature and at a typical value of n.oisture in
flue gas. All particle size determinations are by Banco equipment.
The nature of most Hew South Vales coals is such that the fineness of
fly ash more or less reflects the fineness of coal grind in the boiler
milling equipment
* Supervising 2ngineer, Power Design Group, Electricity Commission of tl.S.W.
191
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Table 1
Characteristics of
Power Station Black Coal
Coal seam -
Area
Stations -
Air dried basis
Moisture
Ash
Volatiles
Fixed carbon
Heat Value KJ/Kg
Sulphur total
pyritic
sulphate
•organic
(Ait f . )
fixed .
Carbon
Hydrogen
Nitrogen
Sulphur (organic)
Oxygen
Hargrove Index
Ash fusion
temperatures °C
Reducing: atmosphere
Deformation
Hemisphere
Flow
Oxidising atmos-
phere
Deformation
Hemisphere
Plow
Great
Northern
Central
Coast
Vales Pt.
Munmorah
3.1
15.6
30.6
50.7
27.80
0.36
0.00
NIL
0.36
0.06
67.7
4.3
1.5
0.4
7.6
47
1270
>1400
>1400
•
1320
>1400
>1400 '
Tongarra/
Vangawilli
Southern
Tallawarra
1.1
32.8
22.8
43.3
22.6
0.42
-
-
-
-
55.3
3.5
1.2
0.4
5.2
65
> 1400
> 1400
> 1400
> 1400
> 1400
>1400
Bayswater
Hunt.er
Valley
Liddell
3.4
30.0
23.2
43.1
21.24
0.34
0.11
ra
0.23
0.16
53.7
3.3
1.2
0.2
8.2
51*
1350
> 1400
> 1400
> 1400
> 1400
> 1400
Lithgow
Western
Wallerawang
3.0
18.0
31.2
47.0
25.84
0.58
-
-
-
-
65.6
4.3
1.5
0.7
6.9
54
>1600
>1600
Callide
Callide (Qld^
Gladstone
9.0 (1)
15.8
26.7
57.7
25.74
0.2
0.12
NIL
0.08
0.13
67.2
3.3
•1.0
0.1
12.6
834
1390
1550
>1600
1520
1600
>1600
(1) As Fired.
192
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Table 2
Analyses of Typical
Black Coal Fly Ashes
Coal Seam
(Station)
Si 02
A1203
Fe2°3
Ti 02
Mn304
Ca 0
Mg o
Na_ 0
K20
P2°5
SO,
Great
Northern
(Vales Pt. )
55.6
29.6
4.53
1.41
0.07
2.27
1.35
0.65
2.98
0.07
1.00
Vangawilli
(Tallawarra)
65.0
25.0
7.24
0.60
0.14
0.33
0.48
0.07
1.01
0.05
0.20
Bayswater
(Liddell)
57.8
28.9
4.71
1.10
0.04
1.76
1.70
0.50
1.27
0.40
1.35
Llthgow
(Wallerawang)
. 63.1
26.8
0.59
0.88
Nil
0.51
0.30
0.12
" 2.52
0.18
0.30
Callide
41.4
38.8
10.15
1.94
0.20
2.80
1.70
0.14
0.08
1.04
0.16
(1) Queensland Alumina Ltd., Gladstone, Queensland.
,.4.0 Precipitator Installations.
During the Second World War and the years immediately after, shortages of
coal and of generating plant diverted attention from deficiencies of dust
collecting plants and it took some time to accept the fact the early precip-
itators were quite unable to meet guaranteed performance. From about 1955
on, there has been a consistent effort to improve the collecting efficiency
of existing equipment and to ensure that new plant performs adequately.
Table 3 lists N.S.W. precipitator installations ordered prior to 1958.
The lack of understanding of the problems- of fly ash collection from low
sulphur coals, general at those times, is regrettably obvious when the
specific collecting areas provided are related to the guaranteed efficiencies,
As a consequence emission levels were up to 40 times the expected level.
193
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Table 3
Precipitatora Ordered Before 1958
Year in
Service
1 939/41
1951/56
1 952/53
1953
1953/55
1957/59
1 958/60
1 960/61
1961
Station/
Plant No.
Bunnerong
Nos. 19-22
White Bay
Nos. 1-4
Pyrmont
Nos. 1-2
Bunnerong.
Nos. 23-24
Pyrmont
Nos. 3-4
Vallerawang
Nos. 1-4
¥angi
Nos. 4-6
Tallavrarra
Nos. 5-6
Vallerawang
Hos. 5-6
Total
Equivalent
Htf
140
100
100
60
100
120
180
200
120
Typical
Operating
Temperature
C
180
180
180
150
180
360
150
140
150
Nominal
S C.A.
m /m'/s
8
N.A.
40
40
40
16
52
54
40
Guaranteed
Efficiency
r'ia
N.A.
98.08(1) (4)
96.0
97.0
97.5
99.0(1 )(2)(3)
98.3(1 )(3)(4)
99.37(0(3X4)
98.5(1) (4)
TOTAL = 25 plants (1100 MW)
Hotes 1) Includes precollectors
2) Hot precipitators
3) Now retired
4) Gas conditioning available
Precipitators purchased for N.S.W. power stations since 1958 have,
directly or indirectly, been sized from extensive pilot plant tests with the
result that the performance of the new installations is significantly better
than previously. (See Table 4).
194
-------
Table 4
Preoipitators Ordered after 1958
Year in
Service
1963
1 964/65
1966
1967/68
1969
1970/73
1976
1976
1977/78
1981
Station
(Unit)
Vales Point (No. 1)
Vales Point (Nos.2,3)
Vales Point (No. 4)
Munmorah (Nos.1,2)
Munmorah (Nos.3»4)
Liddell (Nos.1-4)
Wallerawang (No. 7)
Wallerawang (Nos.SZ,
6X)
Vales Point (Nos.5,6)
Wallerawang (No. 8)
Equivalent
MW
200
400
275
700
700
2000
500
240
1320
500
Nominal
Temp °C
132
132
132
138
138
Pri: 180
Sec: 115
115
150
104
115
Design
S.C.A.
nr/m3/s
53
64
74
96
78
73
112
86
87
112
Guaranteed
Efficiency
%
99.0 (4)
99.0 (4)
99.0 (4)
99.0
99.0 (4)
98.5 (4)
98.5 (0
99.0(2)(3)(4)
99.5
98.5 (0
TOTAL = 18 plants (6835
Notes: (1) At maximum gas temperature
(2; With gas conditioning at H.C.R.
(3) Replacement units
(4) Gas conditioning available
With the exception of one plant where there was a small shortfall, all
the operating plants listed in Table 4 have met or exceeded guarantee, some by
substantial margins.
The replacement precipitators at Wallerawang (5X, 6X) are of special
interest as they are pressurised units, installed between the I.D. fans and the
stack. The existing collectors (mechanical on boilers 1 to'4,"precipitators on
boilers 5 and 6) will remain in service to protect the I.D. fans. It will be
noted from figure 2 that the nominal gaa temperature of 150°C is particularly
adverse at Wallerawang. (Curve G). Unfortunately, intermittent operation of
the station precludes use of water sprays for temperature-modification.
195
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5.0 Flue Gas Conditioning: - Historical Background
The beneficial effects of flue gas conditioning were first demonstrated
in Australia at Wangi and Pyrmont in the late 1950's. 'Whitehead (1)
has reported experience with steam and/or water at Pyrmont in 1956/57.
For reasons not fully understood at the time, the results were both variable
and inconclusive. In addition, there was continual trouble with sprays
blocking, hoppers plugging and fallout in ducts. These considerations led
to water sprays being abandoned as the primary means of improving precip-
itator performance although some steam (or water) conditioning has had a
secondary role up to the present time.
In 1960/61 30-j from a sulphur burning plant at Tallawarra dramatically
improved the collection efficiency of the precipitators. Unfortunately,
operating problems with the sulphur burner led to the trials being ter-
minated after a abort time. At about the same time, several precipitator
manufacturers were operating large pilot plants at Wallerawang, Wangi and
Pyrmont and some testing time was given to conditioning trials.
Because stabilised SOj is not commercially available in Australia, an
experimental sulphuric acid plant was built and operated at Pyrmont in
1964. Modifications of this equipment were later used in other stations
with varied success. In that year pilot plant studies with ammonia and
with temperature modification were made at Tallawarra where exceptionally
high resistivities ( >• 10 ohm metres) make cold precipitation most
difficult. Operational conditioning with ammonia commenced at Tallawarra
in 1965/66.
From 1965 on, gas conditioning has been increasingly used to boost the
performance of undersized plants or plants where changed circumstances
have resulted in excessive emissions. By February, 1976 more than 4,000
MW of boiler plant has been provided with gas conditioning facilities to
aid precipitator performance when, and as, required.
6.0 Gas Conditioning Options
Flue gas conditioning can provide a means of raising the performance of
a sub-standard precipitator to a satisfactory level or reducing the size
of a new. precipitator installation. A number of conditioning agents are
distinctly unpleasant substances and all involve problems of installation,
control, maintenance and add to operating charges.
The techniques tried in Australian power stations are listed below:
6.1 Gas Temperature Modification
While the broad relationship between temperature and resistivity is a
matter of topical interest and has been referred to in many recent papers,
the effect of gas temperature on the precipitability of fly ashes is of more
direct concern than resistivity to operators and designers. Figure 2 shows
the temperature sensitivity of several fly ashes from which it is seen
that not all of these conform to the 'vee' form or something approximating
a mirror image of a resistivity-temperature curve. Further, in those cases
where a definite vee occurs, minimum collection efficiency is found at
perhaps 60°C above the point of maximum resistivity.
In cases where temperature sensitivity is favourable, performance can be
improved by suitably modifying gas temperature without other conditioning.
This was proposed for use at Tallawarra but later abandoned as a primary
corrective measure in favour of ammonia.
196
-------
6.2 Water or Steam Conditioning
Extensive trials have been made with water and steam at several stations.
Although stack emission was reduced in most cases the improvement was not
sufficient and trials were discontinued after several months. The tech-
nique is worthwhile where a precipitator is marginally undersized or
where gas temperatures are initially considerably above 100°C.
6.3 Phosphorous Pentoxide
Although ?2 05 has been found to improve collection efficiency in pilot
precipitators, its cost and nature has discouraged its use in full sized
plants.
6.4 SOx Conditioning
Although sulphur trioxide is not commercially available in Australia, snail
quantities have been imported for pilot plant tests. Performance improved
as injection rates were increased to about 60 p.p.m. ^/V, however with most
fly ash a plateau is reached at about 30 p.p.m. and no improvement occurs
with higher concentrations.
6.5 Sulphuric Acid
Figure 3 shows diagrammatically the sulphuric acid vapourising system
employed at Pyrnont and other statiors . Concentrated sulphuric acid
passes into tubular Hastalloy grids carried above a bank of electric
heating elements in a furnace of firebrick, the resulting vapour being
injected into the boiler after the superheater. The system is comparative-
ly simple and has a high availability with close attention. However,
concentrated acid is not an attractive commodity in power stations and
other techniques are preferred wherever possible.
In the developmental stage of acid vapourisers there was a problem with an
occasional discharge of acidic agglomerates from station stacks which was
countered by providing for neutralising by ammonia injection after the
precipitator.
6.6 Ammonia
This is the most commonly used conditioning reagent in Australian applica-
tions although it is by no means universally (or equally) effective.
Furthermore it has not been found possible to predict its effect with any
certainty.
Figure 4 shows (in a simplified form) the anhydrous ammonia injection system
at Liddell. Road tankers transfer liquid ammonia to the bulk storage tank
from which it is drawn in the liquid phase. Passing through a water warmed
vapouriser (drawing. water from the C.W. discharge canal), the gas passes
from the receiving area to the station by a graded line and is delivered
to each of the four precipitators by a bus main. Each precipitator is
provided with remotely operated valves arranged to allow the plant operators
to select manually either '1CO^', '50$« or 'OFF1 by push button operation
from the main plant control rooms. Pre-set needle valves control the
injection rate which gives a nominal concentration of 40 ppm /Vat iull load.
Below approximately 605^ load the 50$ setting is used. Ko injection is
needed at loads less than at
197
-------
The Liddell boilers have a divided flue gas system (provided to bum high
moisture coal in the later life of the stations). Up to the 30$ by weight
of flue gas at temperature from 125°C to 200°C are handled by two of five
parallel precipitator paths. The balance, at temperatures between 90 and
125°C, is treated by three separate parallel paths. Injection of ammonia
takes place at a point between the four air heater outlets and the point of
duct division into the fire precipitator paths. Figure 5 shows the arrange-
ment of sparge pipes in relation to the ducts and also the details of these
within the duct. When ammonia is not being injected,the vent valves (see
Figure 4) allow a small flow of atmospheric air to be sucked into the
ducts to inhibit jet plugging.
An interesting alternative to anhydrous ammonia is a steelworks by product,
ammoniacal liquor (C.A.I.) which was readily available at low cost and was
used at two stations. Almost the only constant thing
about C.A.L. was the ammonia content, which ran at approximately 1 7$, and
its general unpleasantness. V'ith proper attention to storage and filtering
it was found to be quite satisfactory and was extensively used at both
stations.
Figure 6 shows the basic concepts of the C.A.L. system used on the 100 1W
units of Tallawarra "B1 where its effectiveness was judged to be even better
than anhydrous ammonia.
It is to be noted that all stations using ammonia inject it after the air
heater and therefore the problem of air heater plugging does not arise
unless ammonia leaks occur after boiler shut down. The only cases reported
of build up on precipitator internals, or of difficult dust extraction from
hoppers, are believed to haire occurred when injection rates have been in-
creased in attempts to offset the effect of zones or paths being out of
service or when load reductions have not been followed by reduced injection
rates.
6.7 Triethylamine
The discovery (or rediscovery) of triethylaiaine has been reported by
Paulson (2) and is perhaps the most significant development in the field
for many years. An expensive product of the petrochemical industry,
the comparatively low concentration required can make it competitive with
H2 304 when the cost of post conditioning with ammonia is included.
However, in the Australian price structure,: it is rather more expensive
than ammonia where that reagent is adequately effective.
Triethylamine has been used regularly at Wallerawang since October, 1975
after being tested there in a large pilot plant earlier in the year. It
has also been used experimentally at a cement plant in New South Wales,
at two stations in Western Australia with very satisfactory results.
At the time of writing (February, 1976) certain environmental aspects are
still being studied find the extent of its future use depends on the outcome
of these and on tests being carried out on other ashes.
198
-------
7.0 Effectiveness of Conditioning
Reference to figure 2 shows the relative effect of temperature modification
on a number of fly ashes. It should be noted that the curves demonstrate
the effect of variations of flue gas temperature on precrpitability
(expressed as migration velocity) and are not some reciprocal function of
resistivity. It is readily seen that prior knowledge of temperature response
is helpful in determining whether an increase or decrease of temperature is
likely to be helpful. (See Appendix for an explanation of V/k).
As earlier mentioned, steam or water injection may modify temperature as
well as increasing the moisture content of flue gas. If both of these act
in the same direction i.e. if the lowered temperature results in operation
at a more favourable part of the temperature curve as well as reducing gas
volume AND the added moisture further reduces resistivity, maximum benefit
results. While there is a general tendency for added moisture (in flue gas)
to aid precipitation, this response does not appear in some cases, probably
because of the particular chemistry of the surface layer of dust particles (3).
The effectiveness of ammonia, sulphuric acid and triethylamine as conditioning
agents are shown in fig. 7 fron which it is at once apparent that sulphuric
acid (or SOj) is generally more effective than ammonia and that triethylamine
appears to be superior to both. However, it will also be noted that the
responses of fly ashes to all three reagents are quite variable. While there
is little doubt that SO-j (or H^ SO^) reduces resistivity and that therefore it
might be argued that a competently designed and well maintained precipitator,
sized to achieve high collection efficiencies with high sulphur coal should be able
to achieve the same efficiency if highly resistive fly ash if conditioned with
SO-j, Australian data suggest that other factors may be involved.
Some evidence of the possible significance of pa.rti.cle size has recently
emerged at Wallerawang where pilot precipitator tests during 1975 gave an
improvement of 30$ in ¥ with dust of median sise 13-15/xm from the B station,
whereas A station dust, following mechanical collectors and having a median
size of 6-8/xm, showed little response. V/hile one explanation is that
the mechanical collectors removed the chemically reactive portion of the dust,
a review of earlier data showed a similar lack of sensitivity in dusts below
about 10^um (median). Further study is indicated.
Notwithstanding the marked improvement in electrical characteristics with
Tallawarra fly ash, as earlier reported (4), there was no measurable
change in resistivity. This finding is consistent with that of Dismukes (5)
at several T.V.A. stations. However, although the Tallawarra ash showed a
quite spectacular response to ammonia, this was less rapid than reported by
Dismukes at Gallatin.
8.0 Conditioning Rates
The optimum injection rate for any reagent is obviously the minimum quantity
capable of producing the desired improvement. Unfortunately, there is consid-
erable variation in the optimum rate of injection of most conditioning agents
and, further, as mentioned earlier, it is possible to over condition.
With S03 and E2.S04, the preferred rate has quite often been 20-30 ppm '/V.
Figure 8 shows the response to sulphuric acid vapour at Uangi.
199
-------
The optimum rate for anhydrous ammonia- and GAL (Section 6.6) is very variable;
at one time Tallawarra was conditioning at the rate of 200 to 300 ppm whereas
the Gladstone plant of Queensland Alumina operates at only 22 ppm (6) Pie. 8
also shows the effect of varying the rate of anhydrous ammonia injection at
Wangi.
Reference was earlier made to the potency of triethylamine. Although insuff-
icient work has been done to determine whether this also varies in its effect
from ash to ash, it appears that the optimum rate is less than 10 ppm by weight.
At the risk of stating the obvious, the need to achieve effective nixing of
any conditioning agent with flue gas sufficiently far upstream of the precip-
itator to provide adequate residence time cannot be too strongly emphasised.
In retrospect, it is probable that some of the plants requiring high injection
rates were suffering from the effects of poor distribution, a situation which
arises if nozzles are blocked or if the basic design is poor.
Where a carrier gas is used (e.g. hot air) to introduce a conditioning agent
into flue gas ducts, mixing will be assisted if close attention is given to
achieving densimetric similarity between the carrier gas plus the gaseous
reagent and flue gas0 If a reasonable match is achieved at K.C.R. and satis-
factory results obtained, performance at lower loads is also likely to be good
without modifying the carrier gas temperature.
9.0 Economic Considerations
The capital costs of conditioning plant reflect the basic differences between
the storage and injection systems appropriate to each reagent. The properties
of H2 S04 permit use of carbon steel in some parts of the installation but also
require the use of more exotic materials to handle spills and leaks when the
acid may be diluted. These considerations and the need for post neutralising
ammonia injection make the capital cost of sulphuric acid plants the highest of
those so far used in Australia, an indicative cost being $A 0.12/KW of boiler
plant. The comparative simplicity of anhydrous ammonia systems reduces the
prime cost to about SA 0.05/KW and it is probable that triethylamine installa-
tions may be cheaper still.
The approximate cost of gas conditioning, based on February 1976 prices is
given in Table 5.
Conditioning Costs
(Feb. 1976 prices)
Conditioning Agent Price Cost Injection Rate
" $A tonne cents/Mwhr
™3 240 1.0 20 ppm V/V
*2 S04 ^ 5'° 20 ppm V/V plus
neutralising
ammonia
N (C2 H5)3 1200 . 5.0 ' 10 ppm
200
-------
10.0 Precipitator Capital Costs
Reference was earlier made to the fact that in New South Wales gas conditioning
has been provided on more than 4000 MW of precipitators. Although at first
sight, this would appear to "be an indictment of cold precipitators, this is not so,
as 6^> of conditioninginstallations are on the older plants listed in Table 3 and
of the plants listed in Table 4, four installations are conditioned because
emission limits were tightened between ordering and commissioning and one
because it has a percentage of plate discharge electrodes which have not met
^performance expectations.
Some comments on the estimated relative Australian costs of hot and cold, precip-
itators (conditioned and unconditioned) may be of interest. 1976 prices for a
large (87,000 ni2) cold precipitator of 20 fields are approximately 3A 12/KW,
including ductwork from air heater to stack. Assuming that a hot precipitator
costs of the order of 205* more than a cold unit of the same plate area and with
an allowance of 70/» for additional .gas volume, a hot precipitator becomes com-
petitive when the total migration velocity (Wjj-) approaches twice the cold value.
This corresponds to a V?. or 40/£ increase in e.m.v. (w). On the Australian
scene, only two fly ashes are known to offer the possibility of lower cost hot
precipitators and although the specifications admitted hot plants, they were
not offered.
V/ith the exception of those few cases where fly ash is impossible to collect
in any reasonably sized cold precipitator, the three precipitator options are
cold, cold conditioned and hot. The relative plate areas and relative, cost of
each of these are summarised in Table 6.
Table 6
Comparison of Precipitator Alternatives
For 99.5^ Collection Efficiency .
Wk Specific Collecting Ratio of Ratio of
Ratio Area Plate Areas Capital Costs (4)
Min Max Min Max Hin Max
Hot (1) 2 40 80 1.0 2.0 1.0 1.8
Cold (2) 1 80 160 1.2 2.4 1.0 1.8
Cold (2)(3) 2 40 80 0.6 1.2 Q.5 0.9
(Conditioned)
Note: (1) Performance data from U.S. plants,
(2; Performance data from Australian plants.
m Assume We increased by not less than 100^ with conditioning.
(4) Hot including cost of conditioning plant.
It is seen from Table 6 that conditioning offers the prospect of reduced capital
costs in cases -where there is a high level of confidence in the response of fly
ash to conditioning and that use of hot precipitators does not necessarily confer
a cost advantage. It is readily conceded that other considerations such as
availability and 0 and M charges can have considerable impact in the long term.
201
-------
11.0 Conclusions.
Gas conditioning in one form or another can be considered a reliable means of
improving precipitator performance in some circumstances. Prediction of optimum
concentration or the extent of improvement is seldom, if ever, possible without
some form of prior testing. Care in positioning and servicing the injection
nozzles is necessary as is attention to controlling injection rate, although
load modulation should not be required.
Installation of undersized precipitators with gas conditioning in conjunction
with new boiler plant is not recommended without adequate supporting evidence of
the likely effectiveness of conditioning with all possible coals. However, the
order of saving in capital expenditure possible, where an adequate confidence
level can be established, supgests that investment in appropriate prior research
and/or pilot plant testing may be worthwhile. >
Triethylamine offers considerable promise as a highly effective reagent although
its chemical similarity to ammonia may be shown to make it unsuitable for some
fly ashes.
Gas conditioning is comparatively inexpensive in cases where it produces a
reasonable improvement at low concentrations; even where concentrations of 50
to 100 ppm are required to effect the desired improvement in collection effic-
iency, conditioning may well be cheaper than alternative actions.
Although resistivity is an important factor in determining precipitator
performance, it is not the only consideration 'and therefore recourse to use
of appropriately sized pilot plants is recommended whever possible.
Size for size, hot precipitators are more expensive than cold and, further,
attract additional cost penalties unless the boiler layout is suitably modified
at the design stage. On the evidence available, it is doubtful if all claims
as to their merits in low sulphur coals can be sustained. In some instances,
a significantly cheaper conventional unit will perform as well or better. There
are also cases where hot plants will not only be cheaper than unconditioned cold
units but where, in fact, they provide the only technically acceptable alter-
native.
202
-------
References
1. DARK K., WKITEHEAD C.; "High Efficiency Precipitator Performance
on Modern Power Stations Firing Fuel Oil and Low Sulphur Coals".
Proc. Second International Clean Air Conf. - Washington, B.C.
December 1970.
2. PAULSON C.A.J. et al; "New Ideas on Precipitation Technology from
the CSIRO Combustion Rig". Proc. Institute of'Fuel Symposium,
Adelaide, November 1974.
3. COLLIH P.J.; "Some Aspects of the Surface Chemistry of Fly ashes".
Proc. Institute of Fuel Symposium, Adelaide, November 1974.
4. WATSON K.S., BLECHER K.J.; "Further Investigation of Electrostatic
Precipitators for Large Pulverised Fuel Fired Boilers". Proc.
Australasian Clean Air Conf., Sydney, August 1965.
5. DISMUKES B.B.; "Conditioning of Fly Ash with Ammonia". A.P.C.A.
' Journal, February 1975.
6. HORTH J.C.; "Problems Encountered in Application of Coldside
Precipitators to Callide Coal". Joy Manufacturing Symposium,
Sydney, 1976.
Acknowledgements
Thanks are due to the Electricity Commission of New South Wales
for permission to present this paper and to the many people who assisted
in its preparation.
203
-------
APPENDIX
fatal Average Mi ^ration Velocity - Wk
Extensive use is made of "total average migration velocity" (omega K
or Wk) as distinct from "effective migration velocity "(e.m.v.) as used in the
classical Deutsch formula.
E = 1xf|w - (0
where E = collection efficiency
A = total active surface of
collecting electrodes
Q = gas quantity in unit time
e = 2.7102
(Note that -r = specific collecting area or S.O.A.)
It is frequently found in commercial precipitators that the ~:1 in the
Deutsch formula decreases as the gas volume decreases and increases as the gas
volume increases. Because W is not independent of 77 its use to transpose pre-
cipitator performance will give grossly misleading results. ?or example, if
the specific collecting area of a precipitator operating at 9ft> is increased
'by 2Cfi' the emission, if computed by the Deutsch Wf will be halved; .in reality
a 25/» reduction is about all that can be expected. Conversely, use of the
Deutsch formula to compute the effect of a 10;i increase in gas volume indicates
a 4O?o increase in envirsion whereas a properly designed and correctly operating
precipitator should limit the increase to about 15>.
Several variants of the Deutr.ch formula have been put forward over
the years. The one almost universally used in the Australian power induntry
is th'it developed by the Swedish company of Svenska Plaktfabriken and published
by Matts and Ohnfeldt *.
E=1_e-Wk>k -(2)
where Wk = total average effective
migration velocity
K = 0.4< K <0.6
(usually 0.5)
The total average effective migration velocity (l'/k) has the advantage
that it is independent of specific collecting area and therefore serves very
well to predict the effect of adding (or removing) internal hardware or of
varying gas velocity. It has been found that well designed and maintained
precipitators invariably conform closely to a V/k relationship.
V/k serves as a performance index or figure of merit and is also useful who
when comparing flyashes collected in precipitators of similar design or when
evaluating precipitators of different designs,
Comparison of equations (1) and (2) at once shows that
, .. /~Wk
and -W = —
A^
Q
* Matts 3., Ohnfeldt P.Q.: "Efficient Gas Cleaning with S.F., Electrostatic
precipitators" S.P. Review, 1964.
204
-------
from which it is seen that
ffisl _ (21) and II _ M.
Wk2 ~ (W2) n ¥2 ~ V Wk2
Figure 9 shows a standard precipitator performance plotting sheet. This
is one of several sizes ranging from 380 ram x 340 mm down and scaled in either metric
or English units as desired. The vertical scale is efficiency plotted as ln(-ln(l-S))
and the horizontal scale is 10510 (specific collecting area). Superinpor.ed are two
further scales, the more steeply inclined bein.t the Deutsch 'v/1 and the less steeply
inclined Vfk, The format is different from thit proposed by Hatts and Ohnfeldt
(figure 10) and is a development of the still commonly used Efficiency (lo?) vs SCA
(linear) presentation (see figure 11) which has the disadvantage of representing
precipitator performance by a curve.
By way of illustration, actual test points from an extensive series of
pilot tests at Vfalleravang in 1969 are plotted on each of figures 9, 10 and 11,
together with the related performance bands. The versatility of the presentation
shown in figure 9 is self evident, as is the validity of the Wk concept.
205
-------
• Wollerowong
(Lithgowj
Lidddl
(Myixttr)
B
lodHonc
ollidt)
ISBANt
t^
4W
Figure 1. Location of Power Stations and Coal Fields
(Coal fields in brackets).
206
-------
99-9
100
150 200 250
QAS TEMPERATURE - °C
300
350
Figure 2. Temperature Sensitivity of Some Flyashes.
207
-------
NElfTRALISINC PLANT
Liquid Ammonia
Displocement
Metering Pump
Electrically, heated
Acid Vapouri*er
Bqsement
Figure 3. Sulphuric Acid Injection System Pyrmont.
208
-------
4,S«XOOO |b»/hr
1,560; OOP a elm.
Am
VENT
( 5 ptr fcoikr)
wv
10 WR ATOM
.eow^Roc
mvca
}KJSCUN<30 (M RATl)
VALVC8 (LO ftATI]
WATt*
IQ FAMJ
n
VtNT
STACK
AMMONIA
CW CISCHAMQt
Figure 4. Liddell Ammonia Conditioning System
(diagrammatic only)
209
-------
DENOTES INJECTION POINTS.
PRIMARY
A/H
OUTLET
SECONDARY
A/H
OUTLET
SECONDARY
A/H
OUTLET
ho
H1
O
PRECIPITATOR SECTIONS
PRIMARY
A/H
OUTLET
CLEANING
PLUG
DUCT TOP
36 HOLES 8 DtA.
EQUALLY PITCHED
IN
N. B.
TUBE
z N,B.
GUARD TUBE
(SLOTTED ON
DOWNSTREAM
CENTRELINE)
DUCT FLOOR
SPARGE PIPE
DETAIL
Figure 5. Liddell: Location of Sparge Pipes,
-------
ECTRIC
HEATING ELEMENTS
'AINER
GAUGE
GLASS
•-C.AL CUT-OFF
LEVEL
-WATER
ADMISSION
LEVEL
Figure 6. Tallawarra Power Station Cal. Gas
Conditioning Plant (diagrammatic
arrangement).
211
-------
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213
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215
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216
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MATHEFIATICAL MODELLING OF FINE PARTICLE COLLECTION
BY ELECTROSTATIC PRECIPITATION
John P. Gooch
and
Jack R. McDonald
Southern Research Institute
2000 Ninth Avenue South
Birmingham, Alabama
217
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MATHEMATICAL MODELLING OF FINE PARTICLE COLLECTION
BY ELECTROSTATIC PRECIPITATION
John P. Gooch and Jack R. McDonald
Southern Research Institute
2000 Ninth Avenue South
Birmingham, Alabama 35205
ABSTRACT
A mathematical model is described which calculates collection efficiency
in an electrostatic precipitator as a function of particle size and operating
conditions. The model takes into account the dynamics of particle collection by
determining the electric field, particle charge, and removal efficiency as func-
tions of position along the length of the precipitator. The model contains pro-
cedures for estimating collection efficiency losses caused by non-uniform gas
velocity distributions, gas bypassing the electrified regions, and particle re-
entrainment. Comparisons of the model predictions with experimental data obtained
under essentially idealized conditions in a laboratory scale precipitator indicate
that the model provides a reasonable basis for describing the precipitation process.
Model predictions are compared with field data obtained from precipitators used
to collect fly ash and the results are discussed.
218
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MATHEMATICAL MODELLING OF FINE PARTICLE COLLECTION
BY ELECTROSTATIC PRECIPITATION
INTRODUCTION
The techniques used in the sizing of electrostatic precipitators for
various applications differ with the manufacturer and the application area.
These techniques may be classified into two general categories:
1) The development of statistically-based relationships between
precipitator performance and the properties of the dust or
fuel, or the operating parameters of the precipitator. Usually
this approach involves a correlation of a precipitation rate
parameter which is used as an empirical factor in the Deutsch-
Andersen equation.
2) The use of a pilot plant and prior experience for determining the
collection area required for the design efficiency at a given
application.
Although these techniques can be effective, they can also result in
serious design errors if improperly used. Furthermore, in order to design a
precipitator for a new installation, it is necessary to either have extensive
data for similar applications, or to conduct a test program with a pilot unit.
A preferable approach is to develop theoretical relationships to the point that
precipitator design can be performed by calculation from theoretical principles
and the properties of the dust to be collected. This approach offers the poten-
tial for increased confidence in design and in cost savings by preventing under-
sizing on the one hand and over-sizing on the other. The reliability of predictions
obtained from such an approach is subject to the extent to which certain funda-
mental parameters are known, the degree to which the theoretical relationships
describe precipitator operation, and the accuracy with which the factors that
correct for non-ideal conditions can be modelled and determined. At present,
efficiency losses due to non-ideal conditions are accounted for only by estimation
procedures in which assumed values of the descriptive parameters are normally used.
In order to model particulate collection by the electrostatic precipitation
process from first principles, it is necessary to (1) calculate electric fields
as a function of position and operating conditions, (2) calculate particle charge
219
-------
as a function of time, electrical conditions, and particle size, and (3) mathe-
matically represent the process by which charged particulate are transported
from the bulk gas stream to the collection electrode. The mathematical model
described in this paper includes these operations and predicts collection effi-
ciency as a function of particle size and operating conditions in electrostatic
precipitators.
Comparisons of the model predictions with experimental fractional effi-
ciency data obtained in a laboratory precipitator indicate that the model provides
a reasonable basis for describing the precipitation process. Application of the
model to simulate conditions existing in full-scale and pilot-scale precipitators
used to collect fly ash leads to results which are in reasonable agreement with
experimental field measurements. In certain cases, however, part of the ideally
calculated fractional efficiency curve lies below the measured curve. This type
of discrepancy can be attributed in part to certain unmodelled and approximately
modelled processes which are important under certain conditions in precipitators.
DESCRIPTION OF THE MATHEMATICAL MODEL
Ideal Calculation of Particle Collection Efficiency
The details of a mathematical model for electrostatic precipitation and
a computer program which performs the various operations are described elsewhere
in the literature.1'2'3 In the following, a brief description of the physical
processes is given, and the mathematical framework of the model is developed.
The first fundamental step in the model is to determine the electric
field existing in the precipitator. An accurate calculation of the electric
field is important because of its influence on particle charging and particle
collection. Although the electric field in a precipitator is a function of posi-
tion, the value of the electric field used in the particle charging calculations
is approximated by the average value between the discharge and collection electrodes.
In order to calculate the velocity of charged particles near the collecting elec-
trode, however, it is necessary to compute the local electric field values in this
region of space.
In the model the electric field is determined as a function of position
by the numerical relaxation technique suggested by Leutert and Bohlen.1* Using
this technique, Poisson's Equation and the current continuity equation are solved
simultaneously and subject to the existing boundary conditions. These two equa-
tions can be written in two dimensions in discrete form as:
220
-------
A2V A2v _ _p_
Ax2 Ay2 eQ
(1)
and
Ax Ax Ay Ay
where
V = potential, volts
p = space charge , coul/m
x = distance perpendicular to gas flow from wire to plate, m
y = distance parallel to gas flow from wire to wire, m.
The numerical solutions to Equations (1) and (2) are obtained by iterating
on a grid of potential and space charge density until convergence is obtained.
In the iteration procedure, the space charge density at the wire is adjusted until
solutions are found that satisfy experimentally measured boundary conditions of
applied voltage and average current density at the plate. The major approximation
in the method is the assumption that the motion of all charge carriers can, on the
average, be described by a single "effective mobility." In addition, the "effective
mobility" is assumed to be independent of position in formulating Equation (2) . In
previous work, it has been shown that this technique yields theoretical electric
field distributions which agree well with existing experimental data in the litera-
ture.
The introduction of a significant number of fine dust particles or a heavy
concentration of large particles into an electrostatic precipitator significantly
influences the voltage-current characteristics of the interelectrode space. Quali-
tatively, the effect is seen by an increased voltage for a given current compared to
the dust-free situation. The increased voltage results from the lowered "effective
mobility" of the charge carriers which occurs as the highly mobile gas ions are
bound to the relatively slow dust particles, thus creating a "space charge". It is
desirable to determine the space charge resulting from dust particles because this
quantity influences the electric field distribution, especially near the collecting
plate where, for the same current, it is raised above the dust free situation. Also,
the "space charge" is a function of position along the length of the precipitator.
221
-------
It has been shown3 that the effect of "space charge" on the electrical con-
ditions can be estimated by using an "effective mobility" given by:
= b. { ^ ] , (3)
200 jp + JT
where
b = effective mobility, m2/ (volt-sec)
e
b = ionic mobility, m2/( volt-sec)
i
j = total current density at the plate, amp/m2
T
j = particulate current density at the plate, amp/m2.
The derivation of Equation (3) is based on the assumptions that (1) free
electron current can be ignored, (2) the current carried by the particulate is
insignificant, (3) particle mobility can be treated as independent of size and
on the average, b.^ = 200 bp, and (4) the particles instantaneously acquire a
limiting charge given by the "saturation charge" from field charging theory.
The quantity j^ is measured experimentally and jp can be determined for a given
incremental length of the precipitator from an estimated or calculated overall
mass efficiency.
In a given incremental length of the precipitator, the effect of "space
charge" on the electrical conditions can be determined by solving Equations (1)
and (2) subject to the condition that:
where PpEp is the product of the space charge density and electric field at the plate
averaged over the number of grid points used in the direction of gas flow.
The second fundamental step in the model involves the calculation of particle
charge as a function of particle size, residence time, and precipitator operating
conditions. Accurate values of particle charge are necessary because the electrical
force which causes the particle to migrate towards the collection electrode is
directly proportional to the charge on the particle. In the model particle charge
is calculated using the charging rate equation developed by Smith and McDonald5
given by:
222
-------
dq N0eb
-------
Equation (5) represents the sum of three charging rates corresponding
to three different charging regions on the particle surface. The first term
corresponds to a region of the particle surface where field charging occurs;
the second, to a region where field-enhanced diffusion occurs; and the third,
to a region where diffusion occurs with the effects of the applied electric
field neglected as an approximation. The charging rate given by Equation (5)
approaches the results obtained from the classical field charging equation for
large particles and high electric fields and reduces to the classical diffusion
charging equation in the absence of an applied electric field.
The agreement between the results predicted by Equation (5) and
Hewitt's6 experimental data for dioctyl phthalate droplets is within 25%
over the entire range of data that is available and is within 15% for
practical charging times in precipitators. The agreement between the
theory and the experiment over the particle size range 0.1-1.3 ym and
a wide range of electric field strengths indicates that Equation (5)
should predict closely particle charge as a function of time in the pre-
cipitator model.
The third fundamental step is the calculation of particle collection effi-
ciency. Once the particle charge and the electric field adjacent to the collect-
ing electrode are calculated, the electrical drift velocity, or migration velocity,
resulting from the coulomb and viscous drag forces acting upon a suspended
particle can be obtained. For particle sizes in the size range of interest,
the time required for the particle to achieve the steady-state value of velocity
is negligible, and the migration velocity is given by:
where
(6)
6iray
w = migration velocity of a particle of radius a, m/sec
E = electric field adjacent to the collecting electrode, volt/m
C = Cunningham correction factor
y = gas viscosity, kg/(m-sec).
224
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Gas flow velocities in most cases of partical interest are between
0.60 and 1.8 m/sec, while theoretical migration velocities for particles smaller
than 6.0 ym are usually less than 0.3 m/sec. The path of these smaller particles
therefore tends to be dominated by the turbulent motion of the gas stream in
the interelectrode region. The classical equation for describing particle collec-
tion in electrostatic precipitators under turbulent flow conditions was derived
by Deutsch and gives collection efficiency as a function of gas volume flow,
collection area, and migration velocity in the form:
n = 100 [1 - exp (-Apw/Q)] , (7)
where
f) = collection efficiency of a particle of radius a, %
A^ = collecting -area, m2
Q 5= gas volume flow, m3/sec.
The assumptions on which the derivation of Equation (7) is based are
discussed in detail elsewhere in the literature and will not be restated here.1'8
Of greatest concern in the modelling of the precipitation process is the assump-
tion that gas turbulence provides sufficient mixing to establish a uniform
particle concentration at any cross section of the precipitator. The validity
of this assumption is most important in determining theoretical collection
efficiencies for fine particles.
The mathematical model uses the Deutsch Equation to predict the
collection fraction, n. , , for the i-th particle size in the j-th incremental
1 '3
length of the precipitator. Thus, the Deutsch Equation is applied in the form:
-W. . A./Q
n. . - 1- e ^ 3 (8)
1,D
where W. . (m/sec) is the migration velocity of the i-th particle size in the
j-th increment, and Aj (m2) is the collection plate area in the j-th increment.
Since the Deutsch Equation is based on the assumption that the migration velocity
is constant over the collection area of the precipitator, it is necessary to
make the incremental lengths sufficiently small so that the electric field at
the plate and the charge accumulated by a given particle size remain essentially
constant over the increment.
225
-------
The collection fraction (fractional efficiency) r\. for a. given
particle size over the entire length , of the precipitator is .determined from:
y n. . N. .
4" i,3 1,3
n. = - , '(9)
1 N- i
1,1
where N is the number of particles of the i-th particle size' per cubic
i'3
meter of gas entering the j-th increment. The quantity N, . can be written
i,3
in the form:
-W. . A /Q
N. . = N. . , e 1/D"1 3~1 , (10)
i,3 i/3-l
where N. = N. , the number of particles of the i-th particle size per cubic
i,l i,0
meter of gas in the inlet size distribution.
The overall mass collection efficiency r) for the entire polydisperse
aerosol is obtained from:
n = niPi ,
i
where P. is the percentage by mass of the i-th particle size in the inlet size
distribution.
Methods for Representing Non- Ideal Effects
In the preceding section, a basis for calculating ideal collection.
efficiencies has been developed. This section will discuss the non-idealities
which exist in full-scale electrostatic precipitators and describe calculational
procedures for estimating the effects on predicted collection efficiencies. The
factors of major importance are: (1) gas velocity distribution, (2) gas sneakage,
and (3) particle reentrainment due to rapping.
Non- idealities will reduce the collection efficiency that may be achieved
for a precipitator operating with a given specific collecting area. Since the
model is structured around the Deutsch Equation for individual particle sizes,
it is convenient to represent the effect , of the non-idealities in the model as
correction factors which apply to the exponential argument of the Deutsch Equation.
In the subsequent discussions, these correction factors will be used as divisors
226
-------
for the theoretical migration velocities. The resulting "apparent" migration
velocities are empirical quantities only and should not be thought of as an
actual reduction in the migration velocity in the region of space adjacent to
the collecting electrode,
Although it is widely known that a poor velocity distribution gives
a lower than anticipated efficiency, it is difficult to apply a numerical
description for gas flow quality. White8 discusses non-uniform gas flow and
suggests corrective actions. Preszler and Lajos9 assign a figure-of-merit
based upon the relative kinetic energy of the actual velocity distribution
compared to the kinetic energy of a uniform velocity. This figure-of-merit
will be a measure of how difficult it may be to rectify the velocity distribu-
tion but not necessarily a measure of how much the precipitator performance
would be degraded.
It is possible to develop an approach to estimating the degradation of
performance due to a non-uniform velocity distribution based upon the velocity
distribution, the ideal collection efficiencies, and the Deutsch Equation. 2
It will be assumed that the Deutsch Equation applies to each particle size with
a known migration velocity and that the specific collection area and size of
the precipitator are fixed. A constant k can be determined from the Deutsch
Equation such that:
AW , . .
P / _i_\
k = = u In ( ±_ } , (12)
Al
where A is the total inlet cross-sectional area (m ) and ua is the average
inlet velocity (m/sec). Now, the mean corrected penetration p for a given
particle size can be written as:
N _k.
P = £- £ u± e " »1 , (13)
cl 1—1
where N is the number of points in a velocity traverse and U;L are the point
values of velocity (m/sec).
227
-------
For any practical velocity distribution' and efficiency, the mean penetra-
tion obtained by summation over the velocity traverse will be higher than the
calculated penetration based on an average velocity. If an apparent migration
velocity for a given particle size is computed based upon the mean penetration
and the Deutsch Equation, the result will be a value lower than the value used
for calculation of the single point values of penetration. The ratio of the
original migration velocity to the reduced migration velocity is a numerical
measure of the performance degradation caused by a non-uniform velocity distribu-
tion. An expression for this ratio may be obtained by setting the penetration
based on the average velocity equal to the corrected penetration obtained from
a summation of the point values of penetration, and solving for the required
correction factor, which will be a divisor for the migration velocity.
The correction factor "F" may be obtained from:
N
exp -k/u = p •
Therefore,
Whether the quantity F correlates reasonably well with statistical measures
of velocity non- uniformity is yet to be established. A limited number of
traverse calculations seem to indicate a correlation between the factor F and
the normalized standard deviation of the velocity traverse. Figure 1 shows
F as a function of the ideal efficiency for several values of gas velocity
standard deviation. These curves were obtained by computer evaluation of Equa-
tion 15, and the data on which the calculations are based were obtained from
Preszler and Lajos. The standard deviations have been normalized to represent
a fraction of the mean. The overlapping of the curves for standard deviations
of 1.01 and 0.98 indicates that the standard deviation alone does not completely
determine the relationship between F and collection efficiency.
The data in Figure 1 were used to obtain the following empirical rela-
tionship between F, the normalized standard deviation of the gas velocity dis-
tribution, and the ideal collection predicted for the particle size under con-
sideration:
228
-------
1.786
.86 X
F = 1 + 0.766 r]0 + 0.0755 a_ In - , (16)
' y \i_n/
where
\ N £ < W
" i=l
0g = • (17)
u
a
This relationship is based on a pilot plant study, and should be
regarded as an estimating technique only. If it is desirable to simulate the
performance of a particular precipitator, the preferred procedure would be to
obtain the relationship between F, n and a for the conditions to be simulated
from a velocity traverse at the entrance to the unit.
Gas sneakage occurs when gas by-passes the electrified areas of an
electrostatic precipitator by flowing through the hoppers or through the high
voltage insulation space. Sneakage can be reduced by frequent baffles which
force the gas to return to the main gas passages between the collection plates.
If there were no baffles, the percent sneakage would establish the minimum
possible penetration because it would be the percent volume having zero collec-
tion efficiency. With baffles, the sneakage re-mixes with part of the main flow
and then re-bypasses in the next unbaffled area. The limiting penetration due
to sneakage will therefore depend on the amount of sneakage gas per section,
the degree of re-mixing, and the number of sections.
If the simplifying assumption is made that perfect mixing occurs
following each baffled section, an expression for the effect of gas sneakage
may be derived. Let:
S = fractional amount of gas sneakage per section,
r) = collection fraction of a given size particle obtained with
no sneakage for total collection area,
ru = collection fraction per section of a given particle size =
i - (i - n)1/Ns,
N = number of baffled sections, and
p. = penetration from section j.
229
-------
Then the penetration from section one is given by:
Pl = S + (1 - Tij) (1 - S)
and from section two,
p2 = Spj + (1 - Tij) (1 - S)p1
= p [s + (i - n.) (i - s)J
1 3
= [S + (1 - Tlj) (1 - S)]2
and from section N (the last section) ,
p = [S + (1 - m) (1 - S)]Ns
= [s + (i - s) (i -n )
1/Ns] s . d8)
Figure 2 shows a plot of the degradation of efficiency from 99.9%
design efficiency versus percent sneakage with number of baffled sections as
a parameter. For high efficiencies, the number of baffled sections should be
at least four and the amount of sneakage should be held to a low percentage.
With a high percentage of sneakage, even a large number of baffled sections fails
to help significantly.
A gas bypass or sneakage factor B can be defined in the form of a
divisor for the effective, or length averaged, migration velocity in the ex-
ponential argument of the Deutsch Equation. The factor B is obtained by taking
the ratio of the effective migration velocity w under ideal conditions to the
apparent value of the effective migration velocity w ' under conditions of gas
sneakage so that:
we In(l-n) In (1-n)
B =
w
e
In P N In [S + (1-S) (l-n)1/Ns] . (19)
Figure 3 shows a plot of the factor B versus sneakage for a family of ideal
efficiency curves for five baffled sections. Similar curves can easily be
constructed for different numbers of sections.
The foregoing estimation of the effects of sneakage is a simplification
in that the sneakage air passing the baffles will not necessarily mix perfectly
with the main gas flow, and the flow pattern of the air in the bypass zone
will not be uniform and constant. The formula is derived to help in designing
and analyzing precipitators by establishing the order of magnitude of the prob-
lem. Considerable experimental data will be required to evaluate the method and
establish numerical values of actual sneakage rates.
230
-------
Rapping reentrainment is defined as the amount of material that is
recaptured by the gas stream after being knocked from the collection plates
by rapping or vibration. With perfect rapping, the sheet of collected material
would not reentrain, but would migrate down the collection plate in a stick-slip
mode, sticking by the electrical holding forces and slipping when released by
the rapping forces. However, the rapping forces are necessarily large to over-
come adhesion forces, and much of the material is released into the gas stream
as sheets, agglomerates, and individual particles. Most of the material is
recharged and recollected at a later stage in the precipitator.
Based upon the simplifying assumptions that a fixed fraction of the
collected material of a given particle size is reentrained, and that the frac-
tion does not vary with length through the precipitator, an expression can be
derived identical in form to that obtained for gas sneakage:1'2
PR = [R + (1-R) (l-n)1/NR]NR (20)
where
PR = penetration corrected for reentrainment
R = fraction of material reentrain
NR = number of stages over which reentrainment is assumed
to occur
H = collection fraction of a given particle size obtained
with no reentrainment.
Since Equations (18) and (20) are of the same form, the effect of re-
entrainment can be expected to be similar to the effect of sneakage, provided
that a constant fraction of the material is always reentrained. It is doubtful
that such a condition exists, since precipitators frequently use different rapping
programs on different sections, agglomeration occurs during collection, and
different holding forces exist in different sections. However, until sufficient
data on rapping losses PER SECTION as a function of particle size can be
accumulated, the relationship may be used to estimate the effect of rapping re-
entrainment on precipitator performance.
Figure 4 shows the effect on resultant efficiency for a given size par-
ticle of various degrees of reentrainment for a four-section precipitator with
the indicated values of no-reentrainment efficiency.
231
-------
Since reentrainment and sneakage effects are estimated with identical
mathematical expressions, a combined correction factor B1 is used in the mathe-
matical model. From input values of the fraction of material assumed to be
lost by reentrainment and sneakage, and the number of stages over which losses
are assumed to occur, B1 is determined from the ideal collection fraction for
each particle size.
In summary, the mathematical model takes into account the non-ideal
effects of non-uniform gas velocity distribution, gas sneakage, and rapping
reentrainment by reducing the ideally calculated migration velocities we by the
correction factors F and B' . An "apparent" migration velocity we" is determined
for each particle size from:
w
e F-B'
(21)
Using w ' , the corrected fractional collection efficiencies are
calculated.
EXPERIMENTAL LABORATORY MEASUREMENTS AND COMPARISON WITH MODEL PREDICTIONS
Laboratory-Scale Precipitator
A laboratory-scale precipitator was constructed for the purpose of
studying collection of fine particulate under idealized conditions. The test
unit has an electrode irrigation capability and provisions for humidifying flue
gas produced from a gas- fired burner. For the experiments described here,
however, the carrier gas was ambient air, and the particulate source was an
atomizer which produces a polydisperse aerosol of dioctyl phthalate (DOP) .
Electrode irrigation and humidification were not used, and the collected oil
droplets drained from the collection electrodes by gravity.
Figure 5 is a schematic drawing of the laboratory precipitator assembly.
wire-plate geometry is employed, with plate-to-plate spacing and wire-to-wire
spacing of 12.7 cm (5 in.). Discharge electrodes consist of 0.269 cm (0.106 in.)
diameter wires. The collecting plate area is 2.32 m2 (25 ft2) which results
in a specific collecting area of 35.0 m2/(m3/sec) (178 ft2/1000 cfm) at a gas
velocity of 1.37 m/sec (4.5 ft/sec). The precipitator is divided into four
length sections separated by baffles to limit bypassage, or sneakage, through
the hoppers and the region above the collecting electrode. Three power supplies
are used, the first two of which each energize a 76.2 cm (30 in.) length section.
The third power supply is connected to the last two sections.
232
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Gas bypassage was estimated by measuring gas velocity with a thermal
anemometer in the hoppers and in the region above the collection electrodes.
The estimated amount of bypassage, expressed as a fraction of the total gas flow
through the unit, was between 8% and 10% for the two values of gas flow reported
here. Gas flow uniformity was determined by conducting a traverse with the
thermal anemometer at the precipitator inlet. The standard deviation of the
velocity distribution, expressed as a percentage of the average velocity, was
11.2% for an average gas velocity of 1.09 m/sec.
Measurement Technique
Fractional efficiency measurements were conducted using a five-stage
Brink impactor at the precipitator inlet and outlet. Aluminum foil substrates
were used to facilitate weighing of the small masses collected on the individual
stages. The mass determinations were made with a Cahn Electrobalance. For the
experiments conducted at the lower gas velocity, a sampling time of six hours
was required at the precipitator outlet to obtain weighable quantities on the
impactor substrates. Six and four point traverses were conducted with the impactor
at the inlet and outlet, respectively, in order to insure that a representative
size distribution was obtained at each location.
Since the particulate exiting from an electrostatic precipitator will
have an electrical charge, it was necessary to employ a charge neutralizer
at the inlet of the sampling line for all measurements at the precipitator outlet.
Measurements with an optical particle counter indicated that, if metal substrates
were used in the impactor, the charge acquired by the particulate substantially
influenced the size distribution obtained with the impactor. Since the impactor
was calibrated with an uncharged aerosol, the use of the charge neutralizer
minimized sizing errors due to electrostatic attraction.
Table 1 gives a typical set of inlet and outlet data obtained from
duplicate experiments at the indicated conditions. Good reproducibility was
obtained for all stages at the inlet, and for stages 2 through 5 at the outlet.
Isokinetic sampling was not employed, since the effect of anisokinetic conditions
on particles with diameters less than 2.0 urn is expected to be negligible. The
effect of changing the gas flow rate through the impactor on the indicated size
distribution was checked by reducing the flow rate from 56.6 cm3/sec to 28.3 cm /sec.
The higher flow rate, which was used for all of the fractional efficiency
233
-------
determinations, resulted in a gas velocity of 0.76 m/sec at the sample point.
Figure 6 shows the inlet size distributions from Table 1 and the distribution
obtained at the same location with the reduced impactor flow rate. These data
indicate that the same size distribution is obtained at either flow rate. The
higher flow rate is more desirable because of the smaller cut points and the
increased amount of gas that may be sampled per unit time.
Experimental and Theoretical Results
Fractional efficiency measurements were performed with the Brink impactor
and DOP aerosol at current densities of 26.9, 53.8, and 107.5 nA/cm2 at a gas
velocity of 1.37 m/sec, and at current densities of 26.9 and 53.8 nA/cm2 at a
gas velocity of 0.72 m/sec. The results obtained from these experiments, ex-
pressed as collection efficiency as a function of particle size, are presented
in Figures 7, 8, and 9. Computed collection efficiencies obtained with the mathe-
matical model are also shown. In view of the difficulties in making these types
of measurements, the agreement between measured and calculated efficiencies shown
in Figure 7 is considered good. Figures 8 and 9 suggest that the agreement ob-
tained between computed and measured results decreases with decreasing gas
velocity and current density. The assumption of 8% gas sneakage over four stages
improves agreement between computed and measured results at a gas velocity of
0.72 m/sec. However, the data obtained at this lower gas velocity with 26.9 nA/cm
are considerably below the computer projections. Possible causes of this lack
of agreement are unmodelled effects such as non-uniform current density and
electric field, and particle concentration gradients in the inter-electrode
space.
Another series of experiments is being conducted at the present time
in which fractional efficiencies are being measured for different wire^-to-plate
spacings and wire sizes. Figure 10 shows experimental and theoretical fractional
collection efficiencies and migration velocities for a plate^to^-plate spacing of
19.05 cm (3.75 in), wire radius of 0.1588 cm (0.125 in), and gas velocity of
1.46 m/sec (4.79 ft/sec). These results indicate that the model predictions
give good agreement with experimental data for current densities of 26.9 na/cm2
or less. However, for current densities greater than 26.9 na/cm2 it appears
that the agreement worsens and, in fact, the ideally calculated fractional
efficiency curve lies below the measured curve.
The disagreement in Figure 10^between theory and experiment for current
densities greater than 26.9 na/cm2 occurs at a plate spacing and gas velocity
234
-------
which are not too different from those encountered in the field. Since space
charge effects are not significant in these experiments, it might be sus-
pected that particle concentration gradients exist under certain conditions
and in these cases the collection mechanism in the model is not strictly
applicable.
Measurements have been made at the outlet of the laboratory scale
precipitator under several operating conditions to determine whether or not
concentration gradients existed. A "Clitnet" optical particle counter was used
to count 1.0, 0.5, and 0.3 ym diameter particle sizes at several positions
located between the collecting plate and the wire. The measurements at the
12.7 cm plate spacing under all conditions presented here showed no build-up
of particles near the collecting electrode. These results are consistent with
the fact that the model yields good agreement with experiment under all conditions
at the 12.7 cm plate spacing. The measurements at the 19.05 cm plate spacing
showed definite concentration gradients existing for all three particle sizes
under several conditions- In these cases, the concentration was greatest near
the collecting plate and decreased towards the discharge electrode. This type
of behavior might be expected due to the directed coulomb motion of a particle
which is superimposed on the turbulent motion of the gas stream. The ratios
of the concentration near the collecting plate to that near the wire for 1.0,
0.5 and 0.3 ym diameter particles were 1.7, 1.3, and 1.3, respectively, at
approximately 43.0 na/cm . The causes for such gradients, the conditions under
which they would exist, and their effect on predicted collection efficiency of
fine particles need further investigation.
FIELD MEASUREMENTS AND COMPARISON WITH MODEL PREDICTIONS
In this section the predictions of the mathematical model are compared
with experimental field data for several differing precipitators. These pre-
cipitators were involved in the collection of coal flyash under various operating
conditions. Also, "typical" performance curves are generated using the model
for "cold" precipitators with operating current densities from 5 to 40 na/cm
and for a "hot" precipitator operating at a current density of 30 na/cm .
Figure 11 gives a comparison of fractional effective migration velocities
obtained from the model with those obtained from inertially-determined fractional
efficiency measurements on a coal-fired power boiler. This "cold" full-scale
precipitator has a low specific collection area (SCA) and moderate current
density. The particulate under collection had a mass median diameter (HMD)
of approximately 25 ym and a low value of resistivity. In addition to the
235
-------
ideal predictions, Figure 11 shows the effect of correcting for a gas
velocity standard deviation (0g) of 0.25. A comparison of the we values in
Figure 11 indicates that the theoretical predictions are. low for the particle
diameters below 2.0 ym diameter and high for larger particles. Due to limita-
tions in the measurement technique, no data are available for particle dia-
meters larger than 5.0 ym. Figure 12 shows the fractional efficiency data
from which the we values of Figure 11 were calculated. The overall mass
efficiency predicted by the model was 99.8% as compared to a measured value
of 99.6%.
Figures 13 through 18 give results predicted by the model in terms
of overall mass efficiency as a function of SCA for current densities of 5,
10, 15, 20, and 40 na/cm2 at 150°C, and for a current density of 30 na/cm2 at
370°C. The relationships presented in these figures are based on a 22.86 cm
(9 in) plate spacing, and secondary volt-amp curves and particle size dis-
tribution (HMD = 25 ym) considered to be typical for a coal-fired power boiler.
Also given are test results obtained under conditions approximating the
electrical conditions represented by the given values of current density.
A comparison of the limited amount of applicable test data with the computed
results indicates that the ideally-predicted overall mass efficiencies are
generally higher than those obtained from the field measurements. The mea-
surements were taken with sampling techaniques which insured that essentially
all of the particulate mass larger than 0.3 ym diameter was captured by the
sampling device.
The use of the empirical correction factors reduces the computed
values of overall mass collection efficiency to the range of values obtained
from the field measurements. Reentrainment and sneakage effects were combined
by assuming that the indicated fractional losses per stage occurred over four
effective stages. Note that high efficiency (greater than 99.5%) precipitators
operating on the low temperature side of the air heater fall reasonably close
to the computed line obtained with a~ = 0.25 and S = 0.1. A gas velocity
O
distribution with a standard deviation of 0.25 is generally considered to be
a good distribution for a full-scale unit. The computed results show that a
poor distribution (ag =0.5) seriously degrades performance. The detrimental
effects of sneakage and reentrainment are also indicated in the decrease of
computed performance resulting from variations in the parameter S.
Figures 19 and 20 show results obtained from the analysis of a "cold"
236
-------
full-scale precipitator having a large SCA [560 ft2/1000 ft3/min)] and low
current density (12 na/cm2). The experimental results were obtained from
inlet and outlet size distribution measurements obtained using inertial,
optical, and electrical sizing techniques. The precipitator is preceded by
a mechanical collector so that the inlet size distribution consisted of finer
particles with an HMD of approximately 3 ym. The fly ash under collection
had a high resistivity resulting in limitations on the current density. The
measurements were performed with no rapping so that reentrainment should be
minimized. The ideally predicted fractional efficiency and migration velocity
curves lie below the experimental data for particle sizes in the 0.06-2.5 ym
range. Over this size range the theoretical curve has the same shape as the
experimental data with the minimum occurring at about 0.4 ym in both cases.
The overall mass efficiency predicted by the model is 99.7% as compared to
the measured value of 99.9+%.
Figures 21 and 22 show results obtained from a study of a "cold" pilot
precipitator operating at various SCA's and current densities ranging from 40
to 7 na/ft2. The experimental results were obtained from inlet and outlet
measurements with mass trains using in-stack filters. Particle size distri-
bution measurements made at the pilot precipitator inlet with inertial im-
pactors indicated that a typical fly ash distribution was obtained and that for
particle diameters below about 5 ym, the cumulative mass loadings obtained in
the main duct were within the range of those obtained at the pilot plant inlet.
Both Figures 21 and 22 indicate that in order to match the test data
with the predicted model results at comparable current densities, it is
necessary to either assume a poor gas velocity distribution (0 = 0.50) or
relatively large losses due to reentrainment and gas bypassage. The bypassage
and reentrainment losses are assumed to occur over three effective stages,
since the pilot unit consisted of three sections. Gas flow measurements with
a thermal aneomometer indicated that at this installation gas flow quality was
poor (a = 0.46 or greater), and that gas sneakage was estimated to be no smaller
than 8% and no greater than 30% of the total gas flow. Resistivity data from
in-situ measurements and voltage current curves from the pilot unit power
supplies indicated that the current densities of 11 and 32 na/cm could be
sustained without experiencing electrical breakdown of the dust layer deposited
on the collection electrodes under the test conditions.
Figure 23 shows results obtained from the analysis of a "hot" full-scale
precipitator having an SCA of 430 ft2/(1000 ft3/min) and current density of
237
-------
35 na/OTi2. The experimental results were obtained from inlet and outlet size
distribution measurements obtained using inertial, optical, and diffusional
sizing techniques. For this installation, the fractional efficiency curve
obtained from the measurements is lower than the theoretical curve obtained
from the mathematical model. Although the average electrical operating con-
ditions for this unit are used in the model, some of the electrical readings
for individual sets were characterized by anomalously low current and voltage
levels, suggesting the possibility of electrode misalignment. This is a
contributing factor in the disagreement between theoretical and measured
results since the average electrical conditions will not adequately represent
the effects of low voltages and current in individual sets on collection
efficiency. The theoretically predicted overall mass efficiency was 99.9%
at the indicated condition, and the average measured collection efficiency
was 99.3%. If a gas velocity distribution standard deviation of 0.25 is
assumed, sneakage and reentrainment losses of 10 to 20% over three stages
are required to reconcile the computed and measured mass efficiency.
CONCLUSIONS
Comparisons of theoretically calculated collection efficiencies with
laboratory data obtained under essentially idealized conditions indicate that
the mathematical model gives good agreement with experiments for particle sizes
from 0.55 to 2.5 pm diameter over a wide range of current densities at gas
velocities of 0.72 and 1.37 in/see with a 12.7 cm plate spacing. For a 19.05
cm plate spacing and gas velocity of 1.46 m/sec, the model predicts ideal
fractional efficiency curves which lie below the measured data for current
densities greater than 25 na/cm2. This discrepancy may at the present be
attributed to unmodelled concentration gradients which were measured for
0.3, 0.5, and 1.0 ym diameter particles under these conditions.
Comparisons of model predictions with field data obtained from full-
scale and pilot precipitators indicate that the theory generally gives a good
prediction of overall mass collection efficiency although, in certain cases,
the fine portion of the fractional efficiency curve lies below the measured
data. The laboratory measurements obtained to date indicate that the cases
in which the model underpredicts fine particle collection may be due to con-
centration gradients which exist under certain conditions.
Considering the broad range of laboratory and field data with which
the theory has been compared, it can be concluded that the mathematical model
provides a basis for indicating performance trends caused by changes in specific
238
-------
collecting area, electrical conditions,, and particle size distribution, pro-
vided that back corona does not exist, Current density, applied voltage,
and the particle size distribution are the most important variables in the
calculation of overall mass collection efficiency for a given specific
collection area. The theoretical calculation of ideal overall collection
efficiency of polydisperse particulate in an electrostatic precipitator
generally gives results higher than those obtained from performance measure-
ments on coal-'-fired power boilers. Corrections to the idealized or theoretical
collection efficiency to estimate the effects of non-uniform gas flow, rapping
reentrainment, and gas bypassing the electrified sections reduce the overall
values of calculated efficiency to the range of values obtained from field
measurements. These calculations suggest that the theoretical model may be
used as a basis for quantifying performance under field conditions if
sufficient data on the major non-<-idealities become available.
239
-------
ACKNOWLEDGMENTS
The work described in this paper was supported under contracts
with the Environmental Protection Agency, the Electric Power Research
Institute, and the Tennessee Valley Authority.
240
-------
REFERENCES
1. J. P. Gooch and N. L. Francis, "A Theoretically Based Mathematical Model
For Calculation of Electrostatic Precipitator Performance." Journ. APCA,
Vol. 25, No. 2 (February 1975).
2. J. P. Gooch, J. R. McDonald, and S. Oglesby, Jr., "A Mathematical Model of
Electrostatic Precipitation." Environmental Protection Technology Series,
Publication No. EPA-650/2-75-037 (April 1975).
3. J. P. Gooch and J. R. McDonald, "Mathematical Modelling of Fine Particle
Collection by Electrostatic Precipitation." AICHE 1975 Air Symposium
Volume (to be published).
4. G. Leutert and B. Bohlen, "The Spatial Trend of Electric Field Strength
and Space Charge Density in Plate Type Electrostatic Precipitators."
Staub., 32(7) (July 1972).
5. W. B. Smith and J. R. McDonald, "Calculation of the Charging Rate of Fine
Particles by Unipolar Ions." Journ. APCA, Vol. 25, No. 2 (February 1975).
6. G. H. Hewitt, "The Charging of Small Particles for Electrostatic Pre-
cipitation." AIEE, 76:300-306 (July 1957).
7. W. Deutsch, Ann. der Physik. 68:335 (1922).
8. H. J. White, Industrial Electrostatic Precipitation. Addison-Wesley
Publishing Co., Inc., Reading, Mass. 1963.
9. L. Preszler and T. Lajos, "Uniformity of the Velocity Distribution Upon
Entry into an Electrostatic Precipitator of a Flowing Gas." Staub. 32
(ll):l-7 (November 1972).
241
-------
1.01 0.98
99 99
99.9
g
h-
go
tr
i—
en
o
O
O
_l
Ld
•5.
cc
o
i
i-
is
o
z
UJ
a
u.
UJ
2 3
CORRECTION FACTOR F
Figure 1. "F" as a function of ideal efficiency and gas flow
standard deviation.
242
-------
999
N= NUMBER OF
BAFFLED SECTIONS
.001
1/10%
S% SNEAKAGE PER SECTION
Figure 2. Degradation from 99-9% efficiency with sneakage.
243
-------
ho
4>
-IS
TO
C
i-l
0)
o
o
ro
o
H-
o
n
rt
o
t-h
O
cn
03
cn
3
m
tu
tw
-------
99,9
o
2
UJ
CJ
Lu
LL.
Lul
LU
_J
O
o
CO
01
tr
% REENTRAINMENT PER SECTION
80-
50 —
0
20 30 40 50 60 70 80 90
% OF COLLECTED DUST REACHING HOPPER
100
Figure 4. Effect of reentrainment on the efficiency of a
four-section precipitator designed for a no
reentrainment efficiency as indicated for a
monodisperse particulate.
245
-------
MIXING CHAMBER
(TO
l-f
ro
fD
rt
H-
o
H
I
H-
3
00
s-
o
H
(B
rt
O
l-i
id
H
•o
H-
O
l-t
POWER SUPPLIES
SPRAY
TOWERS
BLOWER
-------
DIAMETER ,um
=> _ c
-be
•
A
cfo
Dn^
1
AK 1
D 1
\ 0 c i
OdJ
O e
/
A
FROM
REDUC
'fe
TABLE 1
ED FLOW RATE
AC$
5
as
»
^
/
D
O
1
0.01
O.I
10 50
% SMALLER THAN INDICATED SIZE
90
Figure 6. Inlet particle size distributions from Brink data.
247
-------
o
z
UJ
98.0
UJ
z
o
UJ 95.0
o
o
90.0
80.0
70.0
60.0
99.9
99.8i .,
j = 26.9 no/cm , v =1.37 m/sec
j =53.8 na/cm^ ,v = 1.37m/sec
j = 107.5 no/cm2,v = 1.37 m/sec
99.51 " / *
99,0
J 1 1 I I I II I I I I i I i I i
o.i i.o 10.0
PARTICLE DIAMETER, Micrometers
Figure 7. Experimentally measured and ideally calculated
fractional collection efficiencies in the laboratory
precipitator.
248
-------
99.99
o
o
U.
u_
LJ
O
LU
o
o
99.95
99.9
99.8
99.5
99.(
I I I
j=53.8na/cm2
v=0.72m/sec
THEORETICAL
'
CORRECTED
/ FOR 8%
SNEAKAGE
O EXPERIMENTAL
10.0
PARTICLE DIAMETER, Micrometers
Figure 8. Experimentally measured, ideally calculated, and
sneakage-corrected fractional collection efficiencies
for a current density of 53.8 nA/cm2 and a gas
velocity of 0.72 m/sec in the laboratory precipitator.
249
-------
99.99
99.98
99.95
99.9
o
5 99.8
o
u.
LL
LU
O
o
LJ
o
99.5
99.0
98.0
95.0
T
T
Mil
j = 26.9 no/cm2
v =0.72 m/sec
THEORETICAL
90.01-—
O.I
CORRECTED FOR
8% SNEAKAGE
I
EXPERIMENTAL
1.0
PARTICLE DIAMETER, Micrometers
10.0
Figure 9. Experimentally measured, ideally calculated, and
sneakage-corrected fractional collection efficiencies
for a current density of 26.9 nA/cm2 and a gas
velocity of 0.72 m/sec in the laboratory precipitator.
250
-------
99.9
99.8
o
Z
LU
o
UL
u.
O
LU
O
u
99
95
90
80
60
30
0
NEGATIVE CORONA
'GAS VELOCITY = 1.46 m/sec
O EXPERIMENTAL,
j = 107.5fxA/m2
O EXPERIMENTAL,
j = 268.8^(A/m2
. ^EXPERIMENTAL,
j = 430fzA/m2
••••» THEORETICAL
•«•• THEORETICAL
8%SNEAKAGE
j I
28.0
24.0
20.0
16.0
.1 1.0 10.0
GEOMETRIC MEAN DIAMETER, .urn
E
u
O
O
m
12.0 y
z
o
cc
8.0 i
4.0
Figure 10. Experimental and theoretical data for a wire of radius
1.488 x 10 3m.
251
-------
ro
Ln
28.0
Z 24.0
O
< J8 20.0
11
2 - 16.0
UJ H
P O 12-°
O _J
LU LU
LU
8.0
4.0
'CALCULATED FROM INERTIALLY
DETERMINED FRACTIONAL
EFFICIENCY MEASUREMENTS
COMPUTED AT
20 na/cm2
0.4 1.0 2.0 4.0
PARTICLE DIAMETER, «m
10.0
Figure 11. Effective migration velocities for a full-scale precipitator
on a coal-fired boiler (SCA = 55.7 m2/(m3/sec).
-------
99.99
o
o
0.1
COMPUTED AT.
20 na/cm2
S = O, = 0.25
1.0
PARTICLE DIAMETER,
10.0
Figure 12. Fractional collection efficiencies for a full-scale
precipitator on a coal-fired power boiler.
-------
100 200 300 400 500 600 700 800
SPECIFIC COLLECTING AREA, ft2/(IOOO ft3/min)
Figure 13. Computed performance curves at 5 nA/cm2
254
-------
100 200 300 400 500 600 700 800
SPECIFIC COLLECTING AREA,
ft2/{1000 ft3/min)
Figure 14. Computed performance curves at 10 nA/cm .
255
-------
90
100 200 300 400 500 600 700
SPECIFIC COLLECTING AREA,
ft2/(1000 ft3/min)
800
Figure 15. Computed performance curves at 15 nA/cm*
256
-------
99.99
100 200 300 400 500
SPECIFIC COLLECTING AREA,
600
Figure 16. Computed performance curves at 20 nA/cm2
257
-------
99.99
100 200 300 400 500
SPECIFIC COLLECTING AREA,
ft2/(IOOO ft3/min)
600
Figure 17. Computed performance curves at 40 nA/cm2,
258
-------
99.99
100 200 300 400 500
SPECIFIC COLLECTING AREA,
ft2/1000 ft3/min)
600
Figure 18. Computed performance curves for "hot" precipitator.
259
-------
55
O
UJ
z
UJ
CL
o
0.01
0.05
1
10
30
60
90
o o
THEORETICAL
0.01
• IMPACTORS (W/O RAPPING PUFFS)
A OPTICAL
O EAA
NO RAP DATA
0.1 1
PARTICLE DIAMETER, Mm
99.99
99.9
99
90
60
40
10
10
o
UJ
O
LL
LL
UJ
U
LU
O
O
Figure 19. Measured and theoretical fractional efficiency.
-------
E
u
>
H
O
O
m 10
O
5
DC
0.01
1 I I I | •
• IMPACTORS (W/0 RAPPING PUFFS)
A OPTICAL \ NQ RAP DATA
OEAA >
THEORETICAL
00*0
I
I
0.1 1.0
PARTICLE DIAMETER,
10.0
Figure 20. Effective migration velocities vs particle diameter.
-------
O
z
UJ
O
LJJ
z
O
u
UJ
O
O
I I I I
COMPUTED
CD = 30 uA/ft2
I I 1 ! I
200 500 800
SPECIFIC COLLECTING AREA, ft2/1000 cfm
Figure 21. Computed and measured data from pilot precipitator.
262
-------
UJ
o
u.
u_
UJ
99.9
99.5
O 99
HI
O
O
95
90
T
T
COMPUTED
CD = 10 uA/ft2
I I
100 200
500
700
SPECIFIC COLLECTING AREA, ft2/1000 cfm
Figure 22. Computed and measured data from pilot precipitator.
263
-------
ro
a^
-P-
3?
O
2
LU
O
LL
LL
LU
2
O
P
O
LU
_J
O
O
99.98
99.9
99.5
98
95
90
60
30
— I II S \ —
— <^jT * *-
*" * ^ — ^x^ *
— — ^ o ~~"
_ 0^0° _
MEASUREMENT METHOD:
— A CASCADE IMPACTORS —
O OPTICAL PARTICLE COUNTERS
• DIFFUSIONAL
PRECIPITATOR CHARACTERISTICS:
TEMPERATURE — 335°C
SCA — 85 M2/(M3/sec)
CURRENT DENSITY — 35 nA/CM2
I II I
0.05 0.1 0.5 1.0 5.0 10.0
PARTICLE DIAMETER, Wm
Figure 23. Comparison of measured and computed fractional
collection efficiencies.
-------
Table 1
TYPICAL INLET AND OUTLET PARTICLE SIZE DATA FOR WET ESP EXPERIMENTS
(OOP Sprayer, Gas Velocity = 1.37 m/sec. Current Density = 55 na/cm2)
Mass Loading, mg/am3
Inlet
Test No.
Lower Size
Stage Limit, um
1 3.04
2 1.80
3 1.24
4 0.65
5 0.46
2ai
26.
18.
7.
7.
2.
341
659
249
448
954
2bi
28.
13.
6.
6.
1.
9S6
526
285
984
775
2ci
24.
14.
7.
6.
2.
596
128
238
743
514
24
15
8
7
2
2di
.088
.226
.226
.419
.654
Average
25.
15.
7.
7.
2.
995
385
250
148
474
2ao
0.
0.
0.
0.
0.
0414
1494
2575
5724
3862
Outlet
2bo
*.
0.
0.
0.
0.
0
1103
2115
4782
3770
Average
0.
0.
0.
0.
0.
207
1299
2345
5253
.1816
Geometric
Mean
Diameter,
urn
2
1
0
0
-
.34
.49
.90
.55
Collection
Efficiency,
99
96
92
84
-
.16
. 77
.65
.58
-------
IMPROVED PRECIPITATOR TECHNOLOGY
BY PILOT PLANT TESTING AND EVALUATION
OF COAL CORE-CORES
Owen J. Tassicker
Electric Power'Research Institute
Palo Alto, California
and
Wayne T. Sproull
Consultant
Glendale, California
267
-------
IMPROVED PRECrPITATOR TECHNOLOGY
BY PILOT PLANT TESTING AND EVALUATION OP COAL BORE-CORES
by
Owen J. Tassicker Electric Power Research Institute Palo Alto, California
and
Wayne T. Sproull Consultant Glendale, California
Abstract
When coal from a new mine field is to be used in a proposed pulverized-coal-
fired boiler, it is important to be able accurately to size the precipitator need-
ed to collect the flyash. Especially if the unknown coal is from the Western
United States the ash resistivity may be high, which increases the size of the
precipitator required for a specified collection efficiency. As much information
as possible on this point may be derived from laboratory tests of coal samples
from an adequate number of bore-cores from the mine field, early in the precipita-
tor planning and design. Small-scale combustors, burning 2 to 75 pounds of pulver-
ized coal per hour, can produce flyash representative of that eventually produced
by the planned full-scale boiler. Thus the ash resistivity and particle size dis-
tribution may be determined in advance, facilitating more accurate precipitator
size selection.
Pilot precipitator tests, properly interpreted, permit a correlation to be
evolved between effective migration velocity (a measure of precipitator perfor-
mance) and the ash resistivity and particle size distribution. Thus the correct
precipitator size may be more accurately predicted by a combination of coal bore-
core analysis and pilot precipitator testing. Such testing is to be elaborated
soon by construction of a new experimental facility at Arapahoe Station, Denver,
Colorado.
268
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IMPROVED PRECIPITATOR TECHNOLOGY
BY PILOT PLANT TESTIHG AND EVALUATION OP COAL BORE-CORES
1, Introduction
men a power company ("Utility") plans to build a new generating plant, to be
coal-fired, it is often necessary to open a new mine field to provide the required
coal. To minimize air pollution and meet the legal standards for particulate emis-
sion, the flyash produced by the coal burning is usually collected from the com-
bustion gases by an electrical precipitator. Such precipitators may allow only %tf>
of the ash to escape up the stack if it comes from "coal A", but, say, 2%, or four
times as much, to escape if it comes from another "coal B".
When the utility asks several precipitator manufacturers to bid, given the gas
flow rate, temperature, required collection efficiency, etc, the bids it receives
may vary by a factor as great as four, in some instances. It is obviously desira-
ble to reduce this spread to a lower maximum such as, say 2 to 1, or hopefully a
maximum such as, say, 1.1 to 1. Technical methods for doing this are the subject
of this paper.
2. New Coal Mine Field Survey
The new coal mine field may have an area of several square kilometers or square
miles. A survey of the field in the early planning stages for the new power sta-
tion results in a geological map. This should include a series of isopleths, each
connecting points of a chosen overburden depth, for example. Likewise for coal
seam depth.
Tests of the coal from each of the dozens of bore-cores for such properties
as heat of combustion, for example, may permit one to draw isopleths for this on
the field map, also. Ashing part of each of these coal samples, followed by chem-
ical analyses of the resulting ash, may permit one to add isopleths for a "boiler-
fouling index", for example.
Since there is some evidence of a relationship between ash (chemical) compo-
sition and ash resistivity (to be discussed later), one might further attempt to
add approximate isopleths for ash resistivity to be expected from coal mined from
various parts of the field.
Finally, as described later, ash resistivity and ash precipitability are re-
lated, so one might even attempt rather speculative isopleths for ash precipita-
biltty. Examination of these isopleths may permit the precipitator designer more
accurately to size the new precipitator to operate satisfactorily, even on ash
from the unfavorable parts of the new coal field.
269
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3. Combustors
Since it is the ash that the precipitator collects, not the coal, it is very
desirable to find a way to produce an ash sample from the coal that may be avail-
able from the new mine field, so that the ash may be tested in the laboratory to
estimate how easy or how difficult it may be to precipitate. To do this, it is
necessary to burn the coal in such a way that its ash will very closely resemble
the ash produced later when that coal is burned in the full-scale completed boi-
ler. The device designed to burn the pulverized coal in the laboratory in this
special way is usually called a "combustor".
Several such combustors already exist. As far as we are aware, the first one
built primarily to help predict precipitator performance on flyash is the "ACIRL
combustor" operated by the Australian Coal Industry Research Laboratories . Some
of the results obtained to date with this combustor have been described by Tassi-
2
cker and Sullivan . Figure 1 shows this ACIRL "micro-combustor" which burns only
a few grams of coal per hour.
Figure 2 shows a larger combustor recently built by Exxon Research and Engi-
neering Co. in Baytown, Texas, their chief investigator being E. C. Winegartner.
It burns about 5 kilograms of coal per hour. Figure 3 shows a comparable combus-
tor in the Babcock & Wilcox Laboratories in Alliance, Ohio, known as their "Labor-
atory ashing furnace" or "LAP"; however, we shall call it the "B & ¥ combustor".
We understand that their chief investigator, C. L. Wagoner, plans to modify this
combustor in four respects; (1) insert a liner into the furnace zone; (2) sec-
tionalize the boiler portion of the equipment; (3) add transmissometer capability;
and (4) change the coal feeder so as to obtain coal rates of 1 or 2 kilograms per
hour, from the present 2 to 9» while retaining the same volumetric heat input
rate.
At this (San Francisco) symposium, Selle and Hess have described some of the
recent work with the pulverized coal combustor at the Energy Research Center,
U. S. Bureau of Mines, Grand Forks, North Dakota. Figure 4 shows this combustor
and the experimental precipitator connected so as to collect the flyash from the
coal which is burned at a rate around 35 kilograms per hour. There is also a com-
bustor of comparable size in Australia at the Commonwealth Scientific and Indus-
trial Research Organization, or "CSIRO". Its coal-burning rate is about 25 kilo-
grams per hour, and some of the work with it is mentioned by Tassicker .
The Aerotherm Division of Acurex Corporation in Mountain View, California has
built and is now operating a "Multi-fuel furnace" rated at 3/4 million kilocalo-
ries/hour (3 million BTU/hour). With coal as a fuel, this amounts to about 100
kilograms or 250 pounds per hour. This combustor was designed to burn natural gas,
270
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distillate oil, No. 6 oil, or pulverized coal. It was designed originally to study
methods of reducing N0x emission. Those attending the (San Francisco) symposium
had an opportunity to inspect this combustor in operation, at Mountain View. It
is shown in Figure 5 and was described for the visitors by R. Brown. The future
flyash studies are to be conducted by J. Steiner.
To summarize these corabustors in order of nominal coal burning rates, we have:
(1) ACIRL, a few grams per hour; (2) B & W, 2 kg/hr; (j) Exxon, 5 kg/hr; (4) CSIRO,
25 kg/hr; (5) Grand Forks, 35 kg/hr; and (6) Acurex, 100 kg/hr. There may be others.
4. Combustor Technology
If the combuator is to produce ash like that from the same coal burned in a
full-scale boiler, the coal must be pulverized before it is fed into the combus-
tor, much as it is for the boiler. A jet of air, flowing at the proper rate to
yield the desired percentage of excess air (above stoichlometric), is usually
used to carry the pulverized coal, fed at the desired rate, into the combustion
chamber. V/hen the combustor first starts up, there must be provision for igniting
the mixed coal dust and air. Some laboratories use a "pilot" butane burner to do
this. Others heat the walls of the combustion chamber with electrical heaters to
such a high temperature that the air-coal mixture ignites spontaneously. The com-
bustion chamber walls are made of a refractory material capable of containing the
burning coal at temperatures between 1100 and 1650 C.
The hot, ash-laden gases from the combustion chamber may then (a) go to a
small test precipitator for collecting the ash, or (b) go to a fiber-glass bag
for ash collection, the ash to be re-dispersed later in a synthetic gas stream
for precipitation, or more usually the bagged ash can be tested for resistivity,
particle size distribution, etc, using separate apparatus. In any of these cases,
the ash-laden gas from the combustor must be cooled from, say, 1400 to, say,
1500 C before it is led to the test precipitator or glass bag. A heat exchanger
is required to do this.
While going from the combustor through the cooling exchanger to the test pre-
cipitator or glass bag the gas velocity throughout should be kept in the vicini-
ty of 15 to 20 meters per second. Otherwise a large fraction of the ash will set-
tle out before it reaches the precipitator or bag. Thus the diameter of the duct
leaving the combustor might be, say, 8 centimeters, where the gas temperature is,
say, 1400 C, but the duct should gradually taper down to, say 4 centimeters as
the gas enters the precipitator or bag, to compensate for the contraction of the
gases as they cool while moving along.
To quote from a B & ¥ report about their "LAF" combustor: "Our experience has
indicated that the residence time for combustion in the LAF of about 0.9 second
is a good model of typical full-scale values in the range of 0.8 to 1.0 second.
271
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However, a heat rate of about 700,000 kilocalories/meterVhour (80,000 BTU/ft*/hr)
in the LAP is required to simulate full-scale ash sintering characteristics pro-
duced at firing rates between 90,000 and 1 JO,000 kcal/m5/hr (11,000 and 15,000
BTU/ft^/hr). This is at least in part due to the high temperature of the furnace
walls in the LAF and the resulting reduction in apparent furnace volume".
The foregoing discussion points up the need for due care in designing and oper-
ating a laboratory combustor to produce flyash practically identical to that from
a full-scale boiler burning the same coal. Can such a combustor follow the differ-
ence observed between ash from a cyclone-fired boiler and ash from other types of
boiler burning the same coal?
A few of the combustors, such as the one at Grand Forks, for example, are di-
rectly connected to a test precipitator, and can be so operated that the gas from
the combustor, after going through the heat exchanger, can be made to enter the
precipitator at a chosen temperature such as, say, 150 C if that is the tempera-
ture anticipated in the future full-scale precipitator. With such a laboratory
system, one may run "coal A", say, and then later, "coal B". If the former yields
an "effective migration velocity" (w) of 10 cm/sec, for example, and the latter
yields w = 5 cm/sec, this indicates that the final full-scale precipitator should
be about twice as big and twice as costly if "coal B" is to be burned as would
be necessary if "coal A" were burned, other things (gas flow rate, collection ef-
ficiency, gas temperature, etc.) being equal.
5. Correlation Between Pilot Precipitator Migration Velocity (w), Ash Resistivity,
and Particle Size Distribution
If the value of w can not be obtained directly by connecting a laboratory
precipitator to the combustor as just described, the alternative is to collect
the ash in a glass bag and later, using separate equipment, measure the bulk re-
sistivity of the ash, its particle size distribution, etc, and from these data
estimate the value of w to be anticipated in the new full-scale precipitator, for
that coal and that ash, under the expected conditions such as temperature, etc.
Wolcott was the first to publish quantitative data relating the electrical
resistivity of a dust to its precipitability, in 1918. Since then, apparatus and
procedures for determining bulk resistivity of dust samples have been described
£ rj
by a number of investigators , and official standards, such as the one by APCA'
have been established. A typical recent resistivity graph showing the variation
of resistivity of an Australian flyash with temperature and humidity may be
O
seen in Figure 6 from an earlier reference .
On the left or cool side of the peaks in the curves in Figure 6, a typical
dust, in bulk, conducts electricity mainly via surface layers on each particle,
these adsorbed layers consisting typically of moisture, and in the case of fly-
272
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ash, sulfur trioxide, and other salts. On the right or hot side of the peaks in
Figure 6, most of these adsorbed surface layers evaporate and the conduction is
mainly via the interior of the dust particles. The volume conductivity increases
with rising temperature, as v/ith a typical semi-conductor, but the surface conduc-
tivity decreases as the temperature rises because the rise evaporates more and
more of the adsorbed conducting surface layer. Thus the resistivity (the inverse
of conductivity) reaches a maximum typically around 110 to 150° C, depending upon
the humidity and dust particle composition.
O
In 1955 Sproull published the graph in Figure 7, showing in a general way
how the effective migration velocity (w) attainable in precipitating limestone
dust varies with its resistivity. Note that a resistivity of 2 X 1012 ohm-cm cor-
responds to a value of w around jj- cm/sec, which is very poor. In addition to
resistivity, migration velocity is also affected by many other factors, including
gas density, viscosity, sparkover voltage, and by particle loading. Despite such
ambiguities, Figure 7 presents a general picture of the way the performance of a
single-stage (conventional) precipitator passes through a transition from good
below 10 ohm-cm to poor above 10 ohm-cm.
Recent resistivity graphs for some different ashes from Western coals, in a
Q
paper by Walker are shown in Figure 8. He notes that although these coals range
in sulfur content only from 0.17 "to 0.91%, at temperatures below 205° C (400° P)
the ash resistivity ranges over five orders of magnitude. Thus the sulfur content
of these low-sulfur coals does not help one predict their ash resistivities.
Table 1, from the same paper by Walker, lists pertinent data from four stations
operating on Wyoming sub-bituminous coal, which he found to be more readily in-
terpreted by using the equation
F=1-e^f)r (1)
where F = collection efficiency (as a decimal, not a percentage), w has already
been defined, f = "specific collecting area" = collecting electrode area/gas flow
rate, and r is an arbitrary exponent selected to give the best fit with the data.
If r = 1, the above equation becomes simply the Deutsch equation, but several
precipitation investigators in recent years have found that a better fit with
experimental data is often attained if r is assigned values other than 1.
In equation (1), the specific collecting area (often called "S C A"), f,
should be expressed in seconds/cm if w is in cm/sec, as it usually is. The folks
who use the metric system often put f in seconds/meter, with w in meters/second.
Precipitation engineers in the U. 5. commonly express f, or "S C A" in square feet
per 1000 cfm, as Walker has done in Table 1. Multiply these values by 0.1968 to
convert to seconds/meter.
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Table 1.
Summary of Precipitator Performance Tests on Western Coals
Precipitator (Prom paper by A. B. Walker, expanded) In situ Drift
Case Test
No. Ho.
1 1
2
3
4
5
2 1
2
3
3 1
2
4 1
2
3
4
5
6
7
temperature '.
°F °C
330
330
300
300
300
330
330
330
885
885
345
309
300
292
301
305
293
166
166
149
149
149
166
166
166
474
474
174
154
149
144
150
152
145
Efficiency 9 S C .
% ft /1 000
99.2
99.1
99.5
98.8
99.9
94.1
91.6
94.8
99.2
98.5
87.0
89.3
89.1
78.9
82.9
81.3
88.8
253
253
664
664
664
184
204
204
235
235
99
111
120
103
107
109
125
A 5 C A re
acfm. sec/meter
.4
.4
.1
.1
.1
.3
.9
.4
.8
.8
.1
.4
.4
.9
.3
.5
.2
50.0
50.0
131.
131.
131.
36.3
40.2
40.2
46.4
46.4
19.5
21.95
23.7
20.5
21.2
21.5
24.6
1
1
6
6
6
9
7
7
1
7
2
3
1
2
1
sistivity vei, w
ohm— cm cm/sec
.98
.98
.5
.5
.5
.7
.9
.4
.6
.3
.6
.4
.9
.3
.6
X
X
X
X
X
X
X
X
—
—
X
X
X
X
X
X
X
8
Q
Q
10*
10*
9
Q
Q
10*
10ii
10
10
10
10
10
10' '
9.66
9.42
4.03
3.30
5.26
7.80
6, f.
.15
7.35
10.4
9.08
10.45
10,15
9.35
7.60
8.35
7.80
8.90
Using the Deutsch equation (equation (1 ) with r = 1), we have calculated the
drift velocity, w, from the values of efficiency (F) and SCA (f) given in Table 1.
These calculated values of w are in the last column, next to the ash resistivity
values in the adjacent column. If resistivity were the only factor to be considered,
w should always increase as resistivity decreases, but the table shows that it does
not. Among other factors to be considered is the ash particle size distribution.
The Deutsch equation is based on the assumption that the particles are of uniform
size; they are not. A prime reason for introducing the arbitrary exponent r into
equation (1 ) was to adjust the equation to fit the experimental data better in
... ,10
this respect .
Particle size distribution is commonly measured with an instrument like the
"Bahco analyzer". For flyash, the size distribution is close to "log normal", which
means that ordinarily it can be approximated by a straight line on the proper type
q
of logarithmic graph paper, as in Figure 9, from Walker's paper .
Figure 10 by Hewitt shows how the drift velocity, w, reaches a minimum at a
particle size ("diameter") near 0.3 micron, both theoretically and experimentally,
for positive corona. Industrial precipitators use negative corona, but the corre-
sponding curves closely resemble those in Figure 10, as seen in Figure 11, from
1 P
a recent paper by McCain, Gooch, and Smith . The performance of an industrial
precipitator is not as sensitive to particle size as Figure 10 indicates, because
considerable agglomeration occurs in all precipitators. Examination of the parti-
cle size distributions of the ash collected in the "A", "B", and "C" fields (in
series) in a precipitator reveals a surprising uniformity.
274
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In general, other things being equal (and they rarely are), a larger precipi-
tator will be needed for unusually fine ash than for ash of ordinary coarseness.
6. Ash Resistivity Related to Its Chemical Composition
Investigators are devoting substantial effort to finding a correlation between
the resistivity of flyash and its chemical composition, but more work is needed
in this area to establish a sufficient confidence level. Elements such as sodium,
potassium, and iron in the coal are also found in its ash, and these appear to
some extent to take over the role of sulfur (in high-sulfur coals) in reducing the
surface resistivity of the ash particles when sulfur and SO becomes scarce, as in
1 -7 A M X
most Western coals. Selle and Hess and Bickelhaupt are among the pioneers in
studying this effect, and preliminary indications are that more sodium, potassium,
and iron in the coal and its ash tends to reduce its surface resistivity. For a
given required collection efficiency, this resistivity reduction tends to in-
crease w, thus allowing a reduction in f or SCA, and hence in precipitator size
and cost,
These elements do not occur in the ash as such, but rather, in such forms as
la-O, KpO, and Pe-O,. These compounds, when present in the ash, obviously got
there from the coal when it burned, so they must have been in the coal. Na_0
sublimes at 1275° C or 2327° P and Fe20, melts at 1565° C or 2849° P, so these
compounds simply pass through the combustion process and emerge in the ash, de-
spite the high furnace temperature. Predicting ash (surface) resistivity from
elemental analysis of the coal (and its ash) from bore cores thus provides a
complementary approach to that from producing flyash from a combustor.
7. Correlation-Regression Equations
2
Prom data contained in the paper by Tassicker and Sullivan , one may write a
regression equation such as
w = aQ + &1 lne f + a2 log1Q p -f a? p + &^ (T/100) + a.^ (T/100)2
+ ag (1/d) ,+ a z + aa log1Q z + ag (T/100)5 (2)
where w = effective migration velocity, f = specific collecting area, as already
defined, p = resistivity of the ash, T = temperature in C, d = particle size
in microns, and z = ash content of the coal, in percent.
Using pilot precipitator test data from 161 tests, with coal (ash) from 5
mines, 3 seams, with ash contents (z) from 16.9 to 22.1%, f values for the preci-
pitators from 30 to 120 seconds/meter (9 to 36 seconds/foot), average particle
8 12
sizes ranging from 9.6 to 14»4 microns, resistivities from 5X10 to 5X10
ohm-meters (5 X 1010 to 5 X 10 4 ohm-cm), and gas temperatures from 85 to 350° C,
the following results were obtained from application of equation (2):
275
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Value Reduction in
variance
+1.06146
19.6%
-0.25149
-0.02077
+4.0436 x io~15
-0.12856
+0.03037
+0.09565
+0.02099
-0.74618
33.2%
20.7%
1 %
7 %
Coefficient Value Reduction in Comments
variance
Coefficient is a kind of "dump".
Steady coefficient in all tests.
Most significant reduction in variance.
Steady coefficient in all tests.
Slight statistical significance.
Low statistical significance.
a Significant above 250 C in Figure 1J,
Miltiple correlation coefficient R = 0.899
Standard error SE = 0.0085 meters/second.
The above data, as well as Figures 12 and 13 are from a paper by Tassicker .
The meaning of terms such as "Variance", "Correlation coefficient", and "Stan-
dard error" may be found in appropriate textbooks .
Figure 12 illustrates how some of the data from the above-mentioned 161 tests
fit the regression equation, as a function of f or SCA. Figure 13 shows the fit
for temperature. Similar analytical approaches have been published by Coy and
17
Frisch .
8. "Exact" Mathematical Analytical Models
Instead of assigning r a value differing from 1 in equation (1), one can stay
with the basic Deutsch equation (r = 1 ) and apply it repeatedly to dust particles
in narrow size ranges such as, say, 0.2 to 0.4 micron, then 0.3 to 0.4 micron, etc,
up to, for example, the fraction larger than 50 microns. Then the results can be
added properly to arrive at a weighted average value of f, say, if appropriate
values of w have been chosen for each size fraction. This is too tedious unless
1 ft
one uses a computer, as Gooch and Francis have done. A simplified outline of
their computer program is shown in Figure 14. This program, to date, has been ap-
plied to flyash only.
Inspection of Figure 14 does hot reveal how such information as expected gas
temperature and moisture content, expected resistivity of the ash (at that temper-
ature and humidity), etc, can be fed into the program. It was probably designed
for use in cases where no data are available from bore cores from a new mine field,
so that an experimental approach could be used, as already outlined.
276
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Bi fact, Reynolds, Theodore, and Marino19, who have developed a mathematical
model along similar lines, state in their paper: "A model for calculating collec-
tion efficiencies—one based solely on physical principles and, as a result, one
requiring no prior information from pilot plant or field testing—has been em-
ployed to generate performance curves for twelve industries presently using elec-
trostatic precipitators". The twelve industries are: (1) Flyash (electric power),
(2) Municipal incinerators, (3) Gypsum, (4) Cement, (5) Pulp (and paper) salt
cake, (6) Sintering, (?) Steel open hearth, (8) Blast furnaces, (9) Coke oven de-
tarring, (10) Sulfuric acid mist, (11) Phosphorus mist, and (12) Petroleum crack-
ing.
Several precipitator manufacturers are reported to have such computer pro-
grams either in use or under development. Just how successful and dependable these
may be is not known. Coy and Frisch'? state that they use "various approaches,
including statistical ones, to improve the reliability of the sizing under condi-
tions of minimal information".
9, EPRI Pilot Precipitator Experimental Station to be Built Arapahoe Station
Experimenting with full-size precipitators in an effort to improve their per-
formance, or to try out new techniques oe unorthodox designs in an effort to ad-
vance the technology, is prohibitively expensive because of their great size and
cost. Then too, power plant owners and operators do not wish to vary the opera-
tion of thei plant (load factor, excess air, etc.) just to accommodate the preci-
pitator testing. 1 relatively small pilot precipitator is therefore used for such
work, for the gas flow rate through it may be varied by merely adjusting a dam-
per, or a change in the electrode system may be made quickly and relatively eas-
ily without disturbing the station in any way. There is also a great saving in
cost and manpower, the work proceeds faster, and there are other reasons.
After the desired tests are finished, the pilot precipitator is ordinarily re-
moved from the plant, later to be shipped to some other plant for other tests.
Test results obtained at the second plant may not be compared accurately with
those at the first plant because a different coal may be used, or the boilers may
differ in type and size. Thus there is an advantage in establishing a permanent
pilot precipitator experimental station at a selected power plant, so that many
ideas and designs may be tested and compared with each other, and experiments. such
as those outlined in this paper may be conducted routinely.
For these reasons, EPRI has made arrangements to establish a permanent pilot
precipitator experimental station at Arapahoe Station on the outskirts of Denver,
this station being owned and operated by the Public Service Co, of Colorado. The
plant burns Western coal routinely. Construction of the new experimental precipi-
tation facility is expected to begin this year (19?6).
277
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It is anticipated that new ideas will be tested, such as, for example, improved
ways of charging the ash or dust, ways to agglomerate the ash more effectively
than it is now in a conventional precipitator, etc. Moreover, if the above-de-
scribed combustor tests with small coal samples from bore cores are insufficient
in some cases, perhaps a coal shipment from the new mine could be burned and the
ash tested at Arapahoe.
Figure 15 is a schematic plan of the new facility, showing two pilot precipi-
tators nominally rated at 8,500 meters / hour (5»000 cfm) each, and two others nom-
inally rated at 60,000 nr/hr (35»000 cfm) each. These are ducted so that the two
small ones can be operated either in series or in parallel, on gas from either the
hot or the cold side of the air preheater. Likewise the two large pilot precipita-
tors can be operated either in series or in parallel, but only on gas from the
cold side of the heater. The laboratory building is to be located in the center,
as shown. It will contain all of the controls and instruments, which will be much
more elaborate than those normally accompanying the usual movable pilot precipita-
tor.
10. Conclusions
The task of the precipitator designer is to select a precipitator for a given
installation, which will be as small as possible, as inexpensive as possible, and
which will perform dependably at the required collection efficiency when operat-
ing at the specified full-load gas flow rate, burning the specified coal in boil-
ers of the chosen type and size,
If only a few kilograms of the coal is available, the designing task is diffi-
cult, as this paper has shown. Even if the coal is available in large quantities,
the task is still complex. The developments described here represent significant
progress along these lines, and when some or all of these new techniques are oper-
able, the designing of new full-scale precipitators will be greatly facilitated.
No longer should a power company receive bids for a new precipitator differing by
a factor such as 2 or more. A factor such as 1.1 should be expected eventually.
Certainly more work along these lines is needed.
Finally, the new test facility at Arapahoe Station should lead to significant
basic developments and technical "break-throughs" in electrical precipitation, in
the years to come.
278
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References
1. P. Dugan, R. E. Guyot, and V. J. Moran, "Laboratory Techniques for the Examin-
ation of the Combustion Characteristics of Coal", ACIRL Report PR 69-4, Part
II, July, 1969, page 32.
2. 0. J. Tassicker and K. M. Sullivan, "Estimation of Precipitator Performance
for Collection of Flyash by Examination of Low-Sulfur Coal Bore Cores". APCA
Conference Paper 73-311, Chicago, 1973.
See also 0. J. Tassicker, "Some Aspects of Electrostatic Precipitator Research
in Australia", APCA Journal 2£, 122, Feb. 1975; ibid. 26, 60, Jan. 1976.
3. S. J. Selle and L. L. Hess, "Factors Affecting ESP Performance on Western Coals
and Experience with North Dakota Lignites", (this) San Francisco Symposium,
May 11, 1976.
4. See reference 2, above; also R. A. Durie, C. A. J. Paulson, and E. C. Potter,
"Mitigating Flyash Emissions in Australia: The Role of CSIRO Testing Facility1,'
Australian Chemical Process Engineering, 25_ (7), 24 (1972); C. A. J. Paulson,
E. C. Potter, and R. Kahane, "New Ideas on Precipitation Technology from the
CSIRO Combustion Rig", Symposium on ESP, Adelaide, Australia, Nov. 1974; E.
C. Potter and C. A. J. Paulson, "What Size Precipitator? A New Basis for De-
signing Plant to Specified Dust-Extraction Performance", Proc. International
Clean Air Conference, Rotorua, New Zealand, Feb. 1975» PP« 159-175? E. C.
Potter and C. A. J. Paulson, "Improvement of ESP Performance by Carrier-gas
Additives", Chem. Ind. July 6, 1974, p. 532.
5. E. R. Wolcott, "Effects of Dielectrics on the Sparking Voltage", Physical Re-
view, 22., 284 (1918).
6. For example, W. T. Sproull and Y. Nakada, "Operation of Cottrell Precipita-
tors—Effects of Moisture and Temperature", Ind. & Engrg. Chemistry, 4^t 1350,
June, 1951; or H. White, "Industrial Electrostatic Precipitation", Addison-
Wesley Pub. Co., Reading, Mass, or Palo Alto, Calif., 1963» Chapter 9.
7. (Committee), APCA Journ. J£, 257, June, 19655 see also ASME Power Test Code 28.
8. W. T. Sproull, "Collecting High-Resistivity Dusts and Fumes", Ind. & Eng.
Chem., 47_, 940, May, 1955.
9. A. B. Walker, "Characteristics and Electrostatic Collection of Particulate
Emissions from Combustion of Low-Sulfur Western Coalft", Paper No. 74-11, APCA
Meeting, Denver, Colo. June, 1974.
10. C. Allander and S. Matts, "Einwirkung der KorngrBssenverteilung auf den Ab-
scheidegrad von Elektrofiltern", Staub-Heinhaltung der Luft, ^, 738 (1957).
11. G. W. Hewitt, "The Charging of Small Particles for Electrostatic Precipita-
tion", AIEE Conference Paper 56-353, New York, N. Y. Jan. 30-Feb. 3, 1956.
279
-------
12. J. D. McCain, J. P. Gooch, and W. B. Smith, "Results of Field Measurements of
Industrial Particulate Sources and ESP Performance", APCA Journ. 25, 117,
Feb. 1975.
13. Reference 3» above.
14, R. E. Bickelhaupt, "Surface Resistivity and the Chemical Composition of Fly-
ash", APCA Journal, 2£, 148, Feb. 1975.
15. 0. J. Tassicker, "Performance of Cold Side and Hot Side ESP's Treating High-
Resistivity Flyash", The Institute of Fuel, Symposium on the Changing Technol-
ogy of ESP, Adelaide, South Australia, Nov. 8, 1974.
16. For example, L. G. Parrat, "Probability and Experimental Errors in Science",
John Wiley & Sons, Inc., Hew York, 1961.
17. D. W. Coy and N, W. Frisch, "Specifying Precipitators for High Reliability",
APCA Journal, 25_, 128, Feb. 1975.
18. J. P. Gooch and N. L. Francis, "A Theoretically Based I%thematical Model for
Calculation of ISP Performance", APCA Journal, 25, 108, Feb. 1975.
19. J. P. Reynolds, L. Theodore, and J. Marino, "Calculating Collection Efficien-
cies for ESP's", APCA Journal, 2_5_, 610, June, 1975.
280
-------
Scale in Centimeters
A Hopper With Level
B Stirrer
C Screw Feed
D "Y" Junction
E Nozzle Feed
F Mixer
G Back Plate
H Support Arm
J Support Arm
K Bracket
L Bracket
M Electric Motor
N Electric Motor
P Glass Tube Insert
O Wheel
R Bracket
Figure 1 The ACIRL micro-combustor.
281
-------
OO
Air
Test Fuel Feed
w\
1
r
r~
To Stack
Bag House
Blower
Gas Analyzer
Radiation
Pyrometer
Thermocouple
V
Test combustor
Figure 2 The Exxon combustor.
-------
Figure 3 The B&W combustor (recently modified). The vertical segmented column is a water-cooled
heat exchanger to cool the hot gases from the combustor. A test precipitator now planned
will be installed in one of the segments where the gas temperature is at the desired level. B&W
is also building a laboratory precipitator resembling the Australian "Kevatron".
283
-------
To Stack
N3
00
.p-
Air-cooled Test Probe
Figure 4 The Grand Forks combustor (rated 35 kg/hr coal) and electrostatic precipitator test system.
-------
C. E.- type corner
fired burners
375 megacalorie/hr
(1.5 million Btu/hr)
I FRF burner
Staged injection ports
Heat exchange sections
Drawer assemblies
Combustion chamber
(1 meter cube)
1760°C (3200° F)
refractory
Observation ports
Ignition and flame
safeguard
r-1
•a
I
«£
=^=^
L-I
"^T1" " "
> ' • '-'•"'[*
L.-T ^--. - .T-. .^f- '
•r-1
— """"""T!
N
K
T~ * * 1
-M-i
r~-
—
1
-1
Figure 5 The Aerotherm-Acurex multi-fuel combustor.
285
-------
10007(273 + temperature "O
2.8 2.7 2.6 2.5 2.4 2.3 2.2 2.1
2.0
10
Water by Weight in Flue Gas
Bayswater Ash
stress 4 kV/cm
80 100 120 140 160 180 200 220
Resistivity of Fly-ash as a Function of Temperature and Humidity fC)
Figure 6 Typical resistivity curves for flyash from low-sulfur coal.
286
-------
20
15
5
> 10
Q
101
Resistivity of Precipitated Dust Layer (ohm - cm)
Figure 7 Curve showing general trend of conventional precipitator performance (w) as
a function of dust resistivity.
287
-------
10'
10'
S io1
—
'in
QJ
cc
"ro
o
u
109
108
107
Reference: Walker
38 93 149 204 260 315
Temperature (°C)
Figure 8 Laboratory resistivity curves for flyash samples from several Western coals
Stress 1 kv across 3 mm ash layer; 10 to 12% moisture by volume except 6%
tor one hot precipitator sample.
288
-------
9.99
99
Reference: Walker
TJ
C
CO
C
.C
90
80
70
60
50
40
30
20
10
Range for P.C. Boilers
Range for Cyclone Boilers
0.1
5 10 20 50 100 200
Equivalent Particle Diameter (microns BAHCO)
500
1000
Figure 9 Ash particle size distribution curves approximate straight lines when
plotted as shown here.
289
-------
13.5
12.0
10.5
_ 9.0
o
o
A. Experimental Data for 3.6 kV/cm Field
B. Theoretical, Field Charging Only,
Same Strength
C. Theoretical, Thermal Charging Only
D. Theoretical, Field Charging plus
Thermal Charging
Reference: Hewitt
0.03 0.05
0.1 0.2 0.3
Particle Radius (microns)
0.5 0.7 1.0
Figure 10 Drift velocity (w) as a function of particle size.
290
-------
99.98
99.9
99.8
99.5
99
98
95
90
c
o
cS
60
PRECIPITATOR CHARACTERISTICS:
Temperature - 160°C
SCA- 67m2/(m3/sec)
Current Density — 15.6 nA/cm2
Reference: McCain, Gooch, and Smith
30
• Cascade Impactors
T Optical Particle Counters
• Diffusional
0.05
0.1
0.5 1.0
Particle Diameter (microns)
5.0
10.0
Figure 11 ("Cold") precipitator efficiency for varying flyash particle sizes.
291
-------
0.12
0.08
0.06 EMV (m/sec)
Theoretical Performance
According to Deutsch
Temperature 115 C
Low Sulfur Newcom Ash
40.0
0.04
0.02
60 80
Specific Collecting Area (m/sec)
Figure 12 Collection efficiency as a function of f or SCA. Illustrating the fit of the
regression equation to some experimental data.
292
-------
0.12
ho
VD
• Regression Equation
• Test Results
100
150
200
Temperature (°C)
250
300
Figure 13 Precipitator performance as a function of temperature, illustrating the fit of the regression
equation to the test data.
-------
Caic. no. of part, in each size band
Calc. space charge due to participate based on TJQ estimate
Calc. reduced free ion density for particle charging calc.
Compute average field for charging
Call E. field, compute field at plate
Call charge, calc. charge on each size part.
Calc. 77 for each size from W for each size
Calc. no. of part, removed in each size
Sum weight of particles removed
Caic. size distribution to next section
Repeat for each part, size
Calc. MMD and weight collected for this increment
Repeat for each increment
Repeat till converges ±0,05%
Calc. effective We for each size
Calc. precipitation rate parameter
Calc. correction factor for gas velocity -*
Calc. correction factor for reentrainment-sneakage
Calc. reduced effective We
Calc. reduced efficiency
Repeat for each part, size
Calc. reduced overall efficiency
Calc. reduced precipitation rate parameter
Reference: Gooch and Francis
Figure 14 Simplified flow diagram of precipitator model computer program.
294
-------
38.2m
-47.5m
Two Nominal 60,000 m3/hr Precipitators (unit 1
Pilot Plant
Operation & Support
Building
Existing Electrical Trench
Existing Fence
Arapahoe Power
Plant
Two Nominal 8,500 m3/hr Precipitators (unit 2)..
34.2m
Figure 15 Preliminary plan of pilot precipitator experimental facility to be built at Arapahoe Station near Denver.
-------
TECHNIQUES FOR CONDITIONING FLY ASH
Edward B. Dismukes
Southern research Institute
2000 Ninth Avenue South
Birmingham, Alabama
297
-------
TECHNIQUES FOR CONDITIONING FLY ASH
Edward B. Dismukes
Southern Research Institute
2000 Ninth Avenue South
Birmingham, Alabama 35205
ABSTRACT
Conditioning of fly ash is discussed in connection with electrostatic pre-
cipitators from several points of view: the purposes served, the types of chemicals
used, and their mechanisms of action. Major emphasis is given to resistivity modi-
fication with sulfur trioxide and other agents. Comments are also made on alterna-
tive conditioning mechanisms: increasing the cohesiveness of deposited fly ash to
minimize reentrainment and improving the electrical properties of flue gas to
increase the efficiency of collection.
298
-------
TECHNIQUES FOR CONDITIONING FLY ASH
INTRODUCTION
Conditioning of fly ash usually involves the injection of a chemical (most
often sulfur trioxide) into flue gas to control the electrical resistivity of the ash
and improve its collection in an electrostatic precipitator. Currently, however, the
term is assuming a much broader meaning, for conditioning involves processes other
than resistivity modification. This paper discusses conditioning to achieve resis-
tivity modification, but it also deals with two other purposes of conditioning: to
increase the cohesiveness of fly-ash particles deposited in a precipitator and to
improve the electrical properties of the gas stream flowing between discharge wires
and collection electrodes.
RESISTIVITY MODIFICATION: TREATMENT OF ASH FROM LOW-SULFUR COAL
Conditioning by Sulfur Trioxide
Origin of High Resistivity. The production of fly ash with a high electrical
resistivity is the usual—but not inevitable—consequence of the combustion of a low-
sulfur coal. The problem of high electrical resistivity is clearly associated with
the combustion of Western coals containing around 0.5% of sulfur. It is widely
assumed to occur when Eastern coals of relatively low-sulfur content (often around 1%)
are substituted for the more prevalent Eastern coals containing 3 to 4% sulfur. How-
ever, in terms of the frequently quoted upper limit of acceptable resistivity values
(about 1 x 1010 ohm cm), ash from Eastern coals presents less of a difficulty than
ash from Western coals. There are at least two factors that control resistivity:
one is the sulfur content of the coal, and the other is the overall elemental compo-
sition of the fly ash.
Sulfur occurs in coal as organic compounds and as inorganic compounds—notably
pyrite and sulfate salts. Regardless of the distribution of the sulfur in these dif-
ferent forms, however, the predominant product of combustion in the boiler is sulfur
dioxide. From a thermodynamic point of view, sulfur dioxide would be the only pro-
duct found in the boiler, and sulfur trioxide (or, more exactly, sulfuric acid) would
be the only product present when the flue gas reaches a temperature around 150°C
where electrostatic precipitators normally operate,1 as shown in Table 1. Thermo-
dynamics evidently is the process-limiting factor in the boiler but not toward the
end of the flue-gas train, for the kinetics of oxidation does not allow more than a
small fraction of the sulfur oxides to appear as sulfur trioxide.
299
-------
Table 1 EQUILIBRIUM DISTRIBUTION OF OXIDIZED FORMS
OF SULFUR AT VARIOUS GAS TEMPERATURES3
Temperature, Relative concentrations, %, of sulfur compounds
°C
1400
1000
800
600
500
400
300
200
100
SO 2
99.7
97.6
88.7
42.2
12.3
1.6
0.1
0.0
0.0
SO 3
0.3
2.4
11.3
57.7
87.3
94.8
59.3
6.5
0.0
HaSO^b
0.0
0.0
0.0
0.1
0.4
3.6
40.6
93.5
100.0
a. Calculated from the data in JANAF Tables,1 assuming con-
centrations of oxygen and water vapor equal to 4% and 10%
by volume, respectively.
b. The maximum absolute concentration of each compound is
sharply limited below 300°C as a result of the condensa-
tion of the predominant compound, HzSOi*, in a binary
HaSO^-HaO liquid mixture.
Even so, the small fraction of the total of sulfur oxides occurring as sulfur
trioxide at 150°C can be sufficient to lower the resistivity of ash to an acceptable
range. Obvious factors of importance are the available concentration of sulfur
dioxide (roughly proportional to the amount of sulfur in the coal) and the extent of
oxidation to sulfur trioxide. A less obvious factor is the nature of the interaction
of sulfur trioxide with the fly ash. We are only beginning to gain answers to this
last question, but this much seems clear: sulfur trioxide and water vapor are
jointly adsorbed or condensed on the surfaces of fly-ash particles, and either vapor
assists in the deposition of the other. The resulting acid may remain essentially
intact in a chemical sense, producing a conductive surface layer on an otherwise
poorly conducting substrate. The acid may, however, react with basic constituents of
the ash and thus undergo conversion to a nonconducting layer of sulfate salts. Cal-
cium oxide, a fairly abundant component of ash from Western coals, is a probable
cause of acid neutralization and its nullification as a conductor. Thus, two coals
similar in sulfur content may ultimately produce similar concentrations of sulfur
trioxide, but the one producing a more alkaline ash is likely to have a substantially
higher resistivity.
Conditioning by Naturally Occurring Sulfur Trioxide. Some of the concepts
outlined above were verified2 in a study of the natural conditioning of fly ash in a
plant burning an Eastern coal containing about 2% of sulfur and producing an ash in
300
-------
which the total percentage of alkaline oxides was low, about 7% by weight (4.2% as
the alkali-metal oxides Li20, Na20, and K20, and 2.4% as the alkaline-earth oxides MgO
and CaO). Experimental data showing fly-ash resistivity in a range of temperatures
at the precipitator inlet are shown in Figure 1. The upper curve shows the expected
maximum—in this instance, around 175°C—that occurs as the result of the transition
between volume conduction at high temperatures and surface conduction at low tempera-
tures. The lower segment of a different curve indicates the effect produced by
injecting water vapor in an amount sufficient to double the normal level—i.e.,
increasing the concentration of water vapor from 7% to 14% by volume.
Other experimental data of value in explaining the observed changes in resis-
tivity are given in Figures 2 and 3. Measured concentrations of sulfur trioxide in
gas samples taken upstream and downstream from the air preheater indicate the occur-
rence of two effects: (1) the thermodynamically predicted increase in the oxidation
of sulfur dioxide to sulfur trioxide as the gas temperature falls across the air
heater (but short of the extent predicted for equilibrium) and (2) the loss of sulfur
trioxide from the gas phase, a phenomenon aided by either the lowering of temperature
or the injection of water vapor. Measurements of slurries of fly ash in water show
corresponding increases in the acidity of the ash (decreases in pH) and in the sul-
fate content of the ash.
Conditioning by Injected Sulfuric Trioxide. In plants where the naturally
available concentration of sulfur trioxide is too low—in an absolute sense, as the
result of a low sulfur concentration in the coal or, in a relative sense, as the
result of reaction with a highly alkaline ash—the conditioning agent can be intro-
duced by an artificial process. Sources of sulfur trioxide are vapor generated by
vaporizing the liquid form of this compound, vaporizing sulfuric acid, or oxidizing
sulfur dioxide gas in a vanadium pentoxide catalyst. In the last-mentioned type of
process, sulfur dioxide may be purchased as the liquid compound under pressure and
evaporated prior to conversion, or it may be generated at the site in a burner for
elemental sulfur. The molecular composition of vapors injected in the various
systems depends not only on the source material but on the temperature and water
vapor concentration in the dilution air. The vapor may be sulfur trioxide per se, or
it may be sulfuric acid. It is of little consequence which substance is injected
for, once injected, either compound will occur only as sulfuric acid vapor in the gas
stream. This statement stems from the rapid rate of the gas phase reaction
S03 + H20 —> HaSOi, under typical flue-gas conditions (temperatures around 150°C and
water-vapor concentrations around 7% to 10% by volume), to approach equilibrium con-
ditions as indicated in Table 1.
301
-------
10
1 "*
10
s
u
£ 1012
H
E-i
10
1 1
10
A. WITHOUT H20 INJECTION
B. WITH H20 INJECTION
140 150 160 170 180
TEMPERATURE, °C
190
200
210
Figure I Electrical resistivity of fly ash as a function
of temperature or water vapor concentration.
302
-------
o
CO
CM
O
2
O
H
E-i
pa
u
s
o
u
UPSTREAM FROM
PRECIPITATOR
UPSTREAM
FROM AIR
PREHEATER
O WITHOUT H20 INJECTION
WITH H2O INJECTION
140
150
160 170 180
TEMPERATURE, °C
190
330
340
Figure 2 Concentration of sulfur trioxide as a function
of sampling location or temperature.
303
-------
o
-P-
tc
5.0
5.5
6.0
6.5
7.0
H
\O ®
o \ WITH H2O
O ^INJECTION
\ ^
WITHOUT \O O °
__H20 INJECTION\^^ _
O ^ _^
O
— —
1 1 1
10 150 160 170 180 19
0.7
0.6
<*>
EH
§ 0.5
EH
13
O
U
g 0.4
h
D
w
0.3
0.2
0 1^:
1
— >»Xt ® ® —
"^•V. WITH H2O
® ">%'>% INJECTION
O
— ^""^^m 0 —
o o ^^--^^^^^
J»
WITHOUT H20
INJECTION
1
10 150 160 170 180 19
TEMPERATURE, °C
TEMPERATURE, °C
Figure 3 Acidity and sulfate content of fly ash as functions
of temperature and water vapor concentration.
-------
Investigators at the Institute have studied the results achieved by the
injection of sulfuric trioxide or sulfuric acid in a number of full-scale power
plants.2 We have considered both practical and theoretical questions and obtained
answers that, to us, seem generally satisfactory, subject to the limitations imposed
by the characteristics of plants available for study.
One question considered is whether there is any high-resistivity ash that
cannot be suitably conditioned with sulfur trioxide. We believe that the answer is
no, but we do recognize that the quantity of conditioning agent required varies with
such factors as the chemical composition of the ash and the gas temperature. The
importance of these factors is illustrated in Figure 4, which gives the results
obtained with ashes characterized broadly as acidic, neutral, or alkaline at temper-
atures varying from 110 to 160°C. If a resistivity of about 1 x 1010 ohm cm is
required, more sulfur trioxide is required if the ash is alkaline or if the gas
temperature is high.
Another question considered is whether the efficiency of conditioning is
markedly affected by the choice of location in the flue-gas train where the agent is
added. Our studies have included plants with and without mechanical collectors to
aid precipitators in the collection of fly ash; they have included plants with injec-
tion ahead of the mechanical collectors, between the mechanical and electrostatic
collectors, and ahead of precipitators operating without mechanical collectors. No
important change in the efficiency of conditioning can be attributed to these varia-
tions in plant operation. We have not had experience with injection ahead of the air
preheater, which would be attractive from the point of view of ensuring uniform intro-
duction of sulfur trioxide or sulfuric acid as a vapor.
Another question was whether a choice is to be made among the different types
of injection systems (based on injection of sulfur trioxide directly, as catalyti-
cally oxidized sulfur dioxide, or as sulfuric acid). We have not recognized any
change in the efficiency of conditioning that we believe to be fundamentally attrib-
utable to the type of system. We did have experience with one acid-injection system
that was inefficient as the apparent result of excessive condensation near the point
of injection.
A fourth matter of practical interest was whether a significant fraction of
the injected sulfur trioxide is lost through the stack. A close examination of this
point was made in only one plant where conditions favored a stack loss, as the result
of the acidic character of the ash and incomplete removal of the conditioning agent.
Figure 5 compares the measured concentrations of sulfur trioxide with and without
injection as a function of distance across the outlet duct at the precipitator or, in
305
-------
t 1 3
LO
O
PLANT 2
PLANT 3
PLANT 5
PLANT 6
PLANT 7
PLANT 5 (135°C, BASIC ASH)
PLANT 3 (110°C, BASIC ASH)
PLANT 6 (160°C, NEUTRAL ASH)
PLANT 7 (125°C, ACIDIC
ASH) '
PLANT 2 (143°C, T
(WEAKLY BASIC ASH)
5 10 15 20 25 30
CONCENTRATION OF SO3 INJECTED, PPM
Figure 4 Resistivity as a. function of the concentration
of injected sulfur trioxide.
-------
u
o
w
D
W
s
E-i
150
140
130
120
110
1 i I i i r
SO3 WITH
INJECTION
SO3 WITHOUT
INJECTION
10
U
2
4 8
O
Cfl
0
0 0.2 0.4 0.6 0.8 1.0
FRACTION OF DISTANCE ACROSS GAS DUCT
Figure 5 Concentration of sulfur trioxide as a function
of gas temperature at the outlet of a precipitator.
other words, as a function of the gas temperature (a variable across the duct as a
result of the influence of the Ljungstrom air heater). The difference in the areas
under the two concentration curves corresponds to an average concentration of 5 ppm,
about 30% of the 14 ppm injected. An increase in the sulfur trioxide concentration
by 5 ppm would produce a concentration of condensed sulfuric acid of 20 yg/m3 (about
0.01 gr/ft3) in the relatively cool plume from the stack—a significant level in
terms of "clear stack" concentration levels for fly ash. We recognize that, under
typical conditions with an alkaline ash from low-sulfur Western coals, the stack loss
of sulfur trioxide will be lower and perhaps insignificant. But it is not a matter
that can be ignored.
A theoretical question of some importance is: What is the mechanism of condi-
tioning? One facet of this question is whether deposition of the agent into ash
occurs by adsorption or condensation. Our conclusion is that adsorption of vapor
above the acid dew point can certainly occur, although condensation may also occur if
the gas is below the dew point. This conclusion is based on a prediction of dew
points from thermodynamic data that are applicable only if the condensed vapor occurs
as a binary liquid mixture of sulfuric acid and water. The possibility that reaction
of vapors, with alkaline fly-ash constituents may take precedence over condensation
makes prediction of dew points on this basis uncertain; if such a reaction occurs,
however, the deposition of vapors would be even more likely to occur as an adsorption
process.
307
-------
Conditioning by Other Agents
No discussion of conditioning of high-resistivity fly ash would be complete
without mention of the role of ammonia, although the subject must be approached with
caution. It appears that this agent sometimes is effective by other mechanisms of
conditioning, as discussed subsequently, and that its ability to alter resistivity is
not clear-cut.
The value of ammonia as a conditioning ash for one type of suspended particu-
late matter—catalyst dust in a petroleum refinery, rather than fly ash from a coal-
burning boiler—has been known for many years. J. F. Chittum of the former Western
Precipitation Company was one of the pioneers in the use of ammonia. He believed
that ammonia was of greater value than sulfur trioxide in treating an acidic dust,
such as the catalyst material in a refinery.
We may very well ask: What can be expected of ammonia as a conditioning
agent for fly ash? Suppose we accept Chittum1s premise that the use of ammonia and
not sulfur trioxide is indicated if the ash is acidic. We would then have to imagine
a fly ash that is at once acidic and high in resistivity. The experience and inter-
pretation of ash chemistry leads to the viewpoint that these may be mutually exclu-
sive factors. None of the constituents of fly ash except adsorbed sulfur trioxide
is sufficiently acidic to have a strong affinity for ammonia, which is a rather
weakly basic compound. But, if adsorbed sulfur trioxide is present, how can resis-
tivity be high?
These are ideas that will have to change if enough clear-cut evidence of a
resistivity effect by ammonia becomes available. There are data available either
directly indicating that ammonia lowers resistivity or indirectly indicating that
ammonia has this effect, such as changes in precipitator voltages and currents that
are most easily explained as a result of lowered resistivity. It does not appear,
however, that conclusive evidence of lowered resistivity with ammonia conditioning is
now in existence.
What about other agents? Sodium is an example that we believe to have clear-
cut significance in lowering resistivity. The importance of sodium first became evi-
dent in a comparison of resistivity values of ashes varying naturally in percentages
of sodium oxide. It was later confirmed in trials of various sodium compounds
(notably the carbonate) as additives to coal being fed to the boiler of a power plant.
Our experience on the use of sodium is entirely restricted to boiler additions (not
flue-gas treatment), "* but others have reported the use of sodium as a flue-gas
additive.5
308
-------
What about some of the proprietary agents now on the market? For some of
these agents, the claim is made that they function as resistivity modifiers, but the
data upon which such claims are made are not often, if ever, given. From what is
known about the compounds present in the proprietary agents and what has been pub-
lished from laboratory studies of the compounds, their functioning as resistivity
modifiers is not an unreasonable assumption.6 One of the compounds identified is
sulfamic acid; another is ammonium sulfate. It is perhaps noteworthy that each of
these compounds is, in theory at least, derived from sulfur trioxide and ammonia.
From a constitutional point of view, sulfamic acid is equivalent to these constitu-
ents in a 1:1 mole ratio (HO-S02-NH2 = S03 + NH3). Similarly, ammonium sulfate or
(NHiOzSOi* is equivalent to sulfur trioxide, ammonia, and water in the molar propor-
tions SOsiNHatHjO = 1:2:1. Thermal decomposition of either compound can thus produce
both sulfur trioxide and ammonia as gases along with further breakdown products, such
as SOz and Na.
RESISTIVITY MODIFICATION OF ANOTHER TYPE: TREATMENT OF ASH FROM HIGH-SULFUR COAL
It is important to realize that the problem of excessively low resistivity
occasionally exists, although it is less widely recognized than the problem of high
resistivity. It occurs in the burning of a high-sulfur coal, and it conceivably may
occur if ash from a low-sulfur coal is overtreated with sulfur trioxide. If resis-
tivity is low—below 1 x 10 ohm cm—the practical consequence may be excessive
reentrainment, owing to the inadequacy of the electrical force across the ash deposit
for maintaining the physical integrity of the deposit under the influence of scouring
or electrode rapping.
In view of the likelihood that low resistivity can be attributed to an exces-
sive amount of sulfur trioxide, it is reasonable to try ammonia as an agent for chemi-
cally neutralizing the acid. From a theoretical point of view, one must be troubled
over the prospects of increasing resistivity sufficiently by simply converting sul-
furic acid to ammonium sulfate. The conductivity of ammonium sulfate in water is
lower than that of sulfuric acid, but the ratio of conductivities is roughly 1:4 and
thus not as great as one order of magnitude. A study of the mechanism of ammonia
conditioning under circumstances where it apparently gave the desired resistivity
change failed to confirm the mechanism assumed.
309
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REENTRAINMENT CONTROL: INCREASING THE COHESIVENESS OF PRECIPITATED ASH
If, as just discussed, the resistivity of ash from a high-sulfur coal can be
increased, one has a means for minimizing reentrainment. If one is unable to make
any significant change in the resistivity and thus in the electrical force needed to
maintain physical integrity of deposited ash, he has the option of increasing mechan-
ical forces that bind individual particles together.
Dalmon and Tidy7 recognized the value of sulfur trioxide as an agent for
increasing the binding forces between particles of fly ash and unburned carbon. Car-
bon is a highly conductive material (relatively speaking) and is especially difficult
to retain after it is electrostatically precipitated, except by means of cohesive
forces. The method used by Dalmon and Tidy for demonstrating the reported activity
of sulfur trioxide involved packing a bed of ash and carbon in a bed with one end
detachable from the other and determining the mechanical force needed to rupture the
bed of particles.
Investigators at the Institute2 obtained evidence in one plant for a similar
effect of sulfur trioxide, although in this instance there was no indication of exces-
sive carbon being the cause of excessive reentrainment. In this plant, conditioning
by sulfur trioxide had been adopted as a measure for restoring the original effi-
ciency of a precipitator after a change had been made to burn coal with a lower sul-
fur content. Our data failed to indicate that resistivity was a problem; the
resistivity of untreated ash was found to be about 1 x 1010 ohm cm, and the value
with 14 ppm of sulfur trioxide injected was about 1 x 109 ohm cm. One might have
expected increased reentrainment as a result of the change in resistivity. In actu-
ality, however, we found that injection of sulfur trioxide increased the precipitator
efficiency markedly. Data in Table 2 show a parallel between the effects achieved by
injecting sulfur trioxide or discontinuing electrode rapping. Our conclusion, there-
fore, is that the principal mechanism of conditioning in the circumstances investi-
gated was increased cohesiveness of the fly ash and decreased rapping reentrainment.
Table 2 REDUCTION OF RAPPING REENTRAINMENT
BY SULFUR TRIOXIDE
S03 Injected, Precipitator
ppm Rapping
0
14
Normal
Off
Normal
Off
69.7
82.0
93.5
96.8
310
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We also found evidence that ammonia injection suppressed rapping losses in
plants burning high-sulfur coal.2 A reproduction of a chart recording the signal
from a transmissometer at the precipitator outlet in one of these plants is given in
Figure 6. After ammonia injection was started, the intensity of rapping puffs was
slowly suppressed. After ammonia injection was discontinued, the intensity of rap-
ping puffs gradually increased again. The findings at the plant in question would
not be complete without mention of other experimental results: no measurable effect
of ammonia on the resistivity of the ash (about 1 x 109 ohm cm) was evident, but
significant effects of ammonia on the electrical properties of the gas stream were
discerned, as discussed later.
It appears that triethylamine, which has recently been found a promising
agent in Australia,8 may act in the manner described for ammonia but even more effec-
tively. Both agents are basic, but the basic strength of triethylamine is greater
and the reported stronger effect of triethylamine seems logical. It must be recog-
nized that the Australian workers have concluded that triethylamine acts as an
"agglomerating" agent, which we interpret to mean an agent that increases particle
size before precipitation occurs. However, if this mechanism of conditioning does
occur, specific evidence for the process does not appear to have been obtained.
ALTERATION OF ELECTRICAL PROPERTIES OF THE GAS STREAM: THE SPACE-CHARGE EFFECT
Conditioning through this mechanism has been observed in several power plants
where ammonia was used for conditioning fly ash.2'9 Two of the plants investigated
o
burned high-sulfur coals and produced fly ash with resistivities around 1 x 10 to
1 x 109 ohm cm. Another two plants burned so-called low-sulfur coals from Eastern
mines, but in these instances the sulfur level was around 1.0% and measured resistiv-
ities were about 1 x 1010 to 1 x 1011 ohm cm. Moreover, significant concentrations
of sulfur trioxide were found in the flue gas—around 2 to 5 ppm, in contrast to
usually undetectable levels (less than 1 ppm) in plants burning 0.5%-sulfur Western
coals.
No measurable effect of ammonia on resistivity could be detected. However,
the gradual change in the intensity of rapping puffs, indicating an increase in the
cohesiveness of deposited ash, was detected clearly in one plant. Moreover, there
were other changes that were virtually instantaneous. One effect was a change in the
precipitator electrical conditions in the direction consistent with a decrease in the
mobility of charge carriers in the gas phase, as shown in Figure 7. Another phenome-
non that occurred rapidly was a marked enhancement in the concentration of suspended
311
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1300
1400
1200
1100
1000
0900
0800
100 80 60 40 20
1500
1600
1700
1800
1900
0 0
RELATIVE VALUE OF
LIGHT OBSCURATION
20 40
8O 100
Figure 6 Reduction of rapping reentrainment by ammonia.
-------
w
u
50
40
30
20
10
NH3 ON
(20 PPM)
NH3 OFF
TR 4C1
1000
1200
1100
HOUR
Figure 7 Rapidity of the effect of ammonia on
the voltage of a precipitator.
fine particles at the precipitator inlet. Coupled with these effects were pronounced
decreases in the concentration of sulfur trioxide as a constituent of the gas stream.
Taken together, the observations suggested a chemical reaction between the
injected ammonia and the normally present sulfur trioxide (actually, the vapor of
sulfuric acid) to produce a fume of fine particles of ammonium sulfate or perhaps the
bisulfate, as shown by, the following equations:
2NH3(gas) + H2S<\(gas)
(NHl»)2SOi» (solid)
NH3(gas) + H2SOi,(gas) —> NHi^HSOi, (solid or liquid)
The bisulfate in the second equation is shown as either solid or liquid; the transi-
tion point is about 144°C.
What is the likelihood of a similar space-charge effect in a plant burning a
typical low-sulfur Western coal, and what would be the value of such an effect? The
313
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reaction of injected ammonia with normally occurring sulfur trioxide would generally
be less likely to occur in such a plant because of the insignificant quantity of sul-
fur trioxide present. However, much higher concentrations of ammonia than are gener-
ally used for conditioning—say, 100 to 200 ppm—could react with trace concentrations
of sulfur trioxide (less than 1 ppm). Concentrations of ammonia in this range, it
may be noted, are used to advantage with high-resistivity Australian ashes. Such
concentrations of ammonia may also react with sulfur dioxide to produce fine parti-
cles of ammonium sulfite, although this does not appear likely from the thermodynamic
data now available.
What about the possible value of the space-charge effect if it does occur?
Possibly this effect could suppress the abnormal currents that are encountered in
instances of back corona. Data published by Watson and Blecher in Australia some
years ago10 gave evidence of back corona suppression by ammonia, at a rate rapid
enough to be attributed to a space-charge effect.
PERSPECTIVE ON THE FUTURE OF CONDITIONING
Chemical conditioning is only one of the alternatives open to the power
industry in dealing with difficult fly ash. The technology of conditioning has been
tried far more often on an experimental basis than it has been adopted for full-time
continuous use. When permanent installations for conditioning have been made,
usually they have been intended to upgrade the performance of an existing precipita-
tor, not to ensure adequate performance by a new collector.
A major factor leading to this policy has been the unpopularity of condition-
ing with sulfur trioxide, mainly as a result of operating difficulties with injection
systems for this agent and hazards associated with the compound. Another factor has
been the frequent lack of success with ammonia and other flue-gas additives as alter-
native agents and the absence of an understanding of the mechanisms by which they
operate.
Some of the needs that must be met if conditioning is to have a continuing or
expanding application are: (1) development of new agents that are attractive on the
basis of cost and convenience and (2) research on the mechanisms of action by various
agents that will give a more dependable basis for predicting performance. Coupled
with these developments, there must be a careful diagnosis of the specific problems
in individual power plants that cause inefficient precipitation and a recognition of
the fact that conditioning cannot rectify all difficulties. In other words, chemical
314
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companies, equipment vendors, and utility companies must all accept their responsi-
bilities if, in the long range, conditioning is to serve the immediate interests of
all concerned.
One final point to be made is that research must be done to ensure that con-
ditioning has a favorable overall impact on the environment. Regulatory agencies
will not, in the long run, tolerate conditioning as a means of reducing fly-ash
emissions if it leads to other undesirable effects that can be avoided through the
adoption of some other corrective measure.
315
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REFERENCES
1. Stull, D. R., and H. Prophet (ed.). JANAF Therrnochemical Tables. Washington,
National Bureau of Standards, 1971. Unnumbered pages listed in this alpha-
betical order: H20, H2O^S, 02S, and 03S.
2. Dismukes, E. B. Conditioning of Fly Ash with Sulfur Trioxide and Ammonia.
Southern Research Institute, Birmingham, Alabama. Contract TV36921A,
Tennessee Valley Authority, Chattanooga, Tennessee, and Contract 68-02-1303,
Environmental Protection Agency, Research Triangle Park, North Carolina.
August 1975. Publication TVA-F75 PRS-5 or EPA-600/2-75-015. 157 p.
3. Chittum, J. F. Western Precipitation Corporation, Los Angeles, California.
Unpublished data from studies in 1942-1945.
4. Bickelhaupt, R. E. Electrical Volume Conduction in Fly Ash. J. Air Pollut.
Contr. Assoc. 24:251-255. March 1974.
5. Selle, S. J., and L. L. Hess. Factors Affecting ESP Performance on Western
Coals and Experience with North Dakota Lignites. Grand Forks Energy Research
Center. (Presented at the Symposium on Particulate Control in Energy Pro-
cesses. San Francisco. May 11-13, 1976.)
6. Dismukes, E. B. Conditioning of Fly Ash with Sulfamic Acid, Ammonium Sul-
fate, and Ammonium Bisulfate. Southern Research Institute, Birmingham,
Alabama. Contract 68-02-1303. Environmental Protection Agency, Research
Triangle Park, North Carolina. October 1974. Publication No. EPA-650/2-
74-114. 51 p.
7. Dalmon, J., and D. Tidy. The Cohesive Properties of Fly Ash in Electrostatic
Precipitation. Atmos. Environ. (Oxford, England). 6:18-92, February 1972.
8. Potter, E. C., and C. A. J. Paulson. Improvement of Electrostatic Precipita-
tor Performance by Carrier Gas Additives. Chem. Ind. 1974:532-533. July 6,
1974.
9. Dismukes, E. B. Conditioning of Fly Ash with Ammonia. J. Air Pollut. Contr.
Assoc. 25:152-156. February 1975.
10. Watson, K. S., and K. J. Blecher. Further Investigation of Electrostatic
Precipitators for Large Pulverized Fuel-Fired Boilers. Air Water Pollut.
Int. J. (Oxford, England). 10:573-583, September 1966.
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RAPPING REENTRAINMENT STUDIES
Herbert W. Spencer, III
Southern Research Institute
2000 Ninth Avenue, South
Birmingham, Alabama
317
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RAPPING KEENTRAINMENT STUDIES
Herbert W. Spencer, III
Southern Research Institute
2000 Ninth Avenue, South
Birmingham, Alabama 35205
ABSTRACT
The results of experimental investigations of rapping reentrainment in
four electrostatic precipitators connected to coal-fired boilers and in a large
pilot electrostatic precipitator are reported. The objective of the studies
was quantification of rapping reentrainment in terms of the percentage and
particle size distribution of the reentrained particulate.
Impactors and mass trains were used to measure separately rap and nonrap
emissions. An extraction real time optical particle sizing system was also used
to determine the percentage of rapping emissions in five particle size bands.
This system also supplied data on the temporal variations of the emissions due
to rapping. During the pilot-scale test and during one of the field tests, the
rapping emissions were observed to depend on the time interval between raps.
Rapping emissions decreased from 53% to 18% of total emissions for the pilot
test as the time interval between raps was increased. The percentage contribu-
tion of rapping reentrainment to total emissions for all tests ranged from a
high of 53% to a low of 6.5%. Data on the particle-size distribution of the rap-
ping puffs indicated that the mass median diameters of the particles were on the
order of 10-20 ym. The large particles emitted during the rapping puffs appeared
to be agglomerates. Rapping emissions were observed during the pilot-scale test
to have a significantly higher concentration in the lower half of the precipitator.
One of the major causes for this appeared to be hopper "boil-up". It was photo-
graphically observed that most of the particulate dropped into the hoppers; then
a portion rebounded and slowly escaped over the hopper baffles to exit the precipi-
tator.
318
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RAPPING REENTRAINMENT STUDIES
INTRODUCTION
Rapping reentrainment accounts for significant losses from many
electrostatic precipitators. The losses occur when the collected dust is
removed from precipitator collection plates by rapping or shaking. The
dust is lost by direct reentrainment in the gas stream and by dispersion
when the dust falls into the collection hoppers.
The rapping losses are in addition to losses from uncollected ma-
terial and to losses from other sources of reentrainment. Other sources of
reentrainment can include the direct scouring action of the gas on the collected
dust, scouring during sparking, and sweepage of dust directly from the hoppers,
caused by poor gas-flow conditions or by air inleakage to the hoppers.1
This paper discusses the results of experimental investigations of
rapping reentrainment of fly ash in four electrostatic precipitators connected
to coal-fired boilers,2'3 and the results of an experimental investigation of
rapping reentrainment of fly ash in a large pilot precipitator at FluiDyne
Engineering Corporation's Rosemount Laboratory.1*
The main objective of the studies was quantification of rapping re-
entrainment in terms of the percentage of total emissions and in terms of the
particle size distribution of the reentrained particulate with the intent of
providing data for inclusion of losses due to rapping reentrainment in a systems
model of electrostatic precipitation. In addition, during the studies using
the pilot precipitator, the basic mechanics of removal of dry dust by rapping
and the variations in the removal mechanisms with changes in dust properties
were investigated.
Background
The removal of particulate from a gas stream by a dry electrostatic
precipitator involves four separate processes. The first is the precipitation
of the particulate from the gas stream onto the collection plate. A Southern
Research Institute report discusses the physical mechanisms involved and des-
cribes a computer model for the dust collection from the gas stream.5 The second
process is the removal of the particulate from the collection plate, and the
third is the transfer of this particulate to a hopper. The fourth process is
the removal of the collected particulate from the hopper to a disposal area.
In present-day precipitator applications, two approaches are prevalent
with regard to the removal and transfer of the particulate from the collecting
plates. One approach is to rap often and to provide maximum rapping acceleration
to these plates during each rap in an attempt to minimize the thickness of the
residual dust layer. The other approach is to vary the intensity and frequency
of rapping in an attempt to minimize the quantity of material reentrained. A
determination of the best rapping technique for a specific application depends
on an understanding of the method by which dust is actually removed and trans-
ferred from the collection plates during a rapping sequence and of the effects
of residual dust layers.
319
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The mechanics of the dust removal process vary with the properties
of the dust, precipitator operating conditions, and rapping parameters. Dust
properties and precipitator operating conditions affect the adhesion and co-
hesion of the dust layer. The adhesion and cohesion of dust layers depend
upon particle-to-particle forces. According to Tassicker,6 the component
forces are: London-van der Waals, triboelectric, capillary, surface dipole,
and electric-field corona forces. These component forces are influenced by
the following: particle diameter, porosity and compaction of the layer, com-
plex dielectric constant, humidity in the gas, adsorbed surface dipolar
molecules, work-function interfaces on the material, and the electric field
and current density in the dust layer.
Theory of the Mechanics of Dust Removal. An elementary theory of dust
removal which considers only the tensile strength (P) of the dust layer and the
acceleration (a) normal to the plate has been developed by Tassicker. The
theory predicts that the dust layer is removed only when
where 6 is the bulk density, & is the dust layer thickness, and M/A the mass
per unit area. According to this equation, for a given dust thickness, the
rapping intensity must be of sufficient magnitude to produce an acceleration
greater than the ratio of the tensile strength of the ash layer and the mass
per unit area. For a given acceleration, dust is removed only when
M/A > P/a; (2)
that is, when the mass per unit area (dust surface density) is greater than the
ratio of dust layer tensile strength to the normal plate acceleration. Since
the mass per unit area depends on the dust layer thickness, which in turn is
related to collection time between raps, the time interval between the raps is
directly related to the efficiency of dust removal from the plates.
Experimental data obtained by Sproull8 and by Penney and Klingler9 show
that the requirements for removal of a precipitated dust layer are in basic
agreement with Tassicker"s elementary theory for dust removal.
Emissions Due to Rapping. Emissions due to rapping and their dependence
on rapping parameters have been reported previously by Sproull,10 Plato,11
Sanayev and Reshidov,12 Schwartz and Lieberstein,13 and Nichols, Spencer, and
McCain. Sproull found that reducing the intensities of the raps led to a re-
duction in rapping emissions. He also observed a vertical stratification of
the emissions during rapping, with higher concentrations in the lower portion
of the precipitator. The work by the others showed improvements in performance
of full-scale precipitators when the time intervals between raps were increased.
Although these previous studies have added to our understanding of rapping
reentrainment, they have not provided quantitative data on the amounts of emis-
sions due to rapping reentrainment.
320
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EXPERIMENTAL PROCEDURES
A complete characterization of rapping reentrainment requires the measure-
ment of a large variety of variables. A block diagram of an experimental layout
for our pilot rapping reentrainment study is shown in Figure 1. In addition to
the data that is obtained with this arrangement, a complete characterization
utilizes the precipitator design data.
The field experiments included a similar set of measurements to those
made during the pilot studies. However, sampling viewing ports for photograph-
ing rapping emissions and for determining the vertical stratification of the
rapping emissions were not available in the full scale units nor were load cells
for measuring the quantity of fly ash collected on the collection plates. Hence
these measurements were not included in the field tests.
The adhesion and cohesion properties of the collected dust layers were
also not determined during each test since suitable field test equipment was
not available. However, some information about the tensile strength of the
collected ash layers was obtained during the pilot studies by using the ele-
mentary theory of adhesion given in the introduction, the load cell data, and
plate acceleration data.
Quantification of rapping reentrainment in terms of the percentage of
total emissions and in terms of the particle size distribution of the reentrained
particulate required determining the amount of emissions with and without rapping
reentrainment. To define the mechanisms by which the rapping losses occurred
required time-resolved data on the particulate concentrations and size dis-
tributions across typical portions of the precipitator exit plane.
During the investigations the above requirements were met by using
real-time systems and integrating systems. A real-time system based on an
optical single-particle counter was developed which included a five-channel
analog ratemeter, which provided parallel monitoring of the instantaneous
concentrations of particles in five pre-selected size intervals ranging from
0.6 to 20 urn.
To conform to instrumental limitations on the total concentration of
aerosol particles in the sample gas stream arriving at the sensor, the aerosol
sample from the flue was diluted before measurement. Because of the very steep
gradient in the size distribution, on ^a number basis, at the exit of a pre-
cipitator on a power boiler the diluter was made as a size-selective device
which, under ideal conditions, dilutes small particles in the sample gas stream
by fairly large factors while passing a relatively confined and undiluted stream
of the lower concentrations of large particles directly to the particle sensor.
Figure 2 illustrates the operating system for the particle dilution train. Be-
cause of the potentially large probe losses that can result from impaction and
settling in the probe, the system was intended to be used only from below a
duct with a minimum number of bends (one) between the sampling point and the
particle sensing volume. The probe was oriented in the vertical direction so
that settling was in the direction of the probe axis and consequently did not
result in a loss of particles.
321
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For those circumstances in which it was not possible to sample from
below the duct, a second sample extraction system was constructed. These
samples were removed at high flow rates, 0.002-0.02 m3/sec (5-40 cfm) through
a large bore probe (4 cm diameter) and conveyed to a suitable location beside
or on the top of the duct, at which point a secondary sample was extracted
into the diluter and counter as illustrated in Figure 3. This sampling method
provided information on relative concentrations of particles of various sizes
during and between puffs, but did not provide quantitative concentration data
because of the uncertainties in the probe losses and in the degree to which
the secondary sample represents the average concentration in the high flow
rate probe.
Integrated measurements of rapping emissions were made with separate
sets of inertial impactors and mass trains during the pilot studies to in-
dependently quantify rapping reentrainment losses and non-rapping reentrainment
losses. One set of impactors and mass trains measured losses between raps and
the other set measured rapping losses. The dust feed was turned off during
rapping. During two of the field tests, losses were measured with separate sets
of impactors and mass trains during alternating periods with rapping and without
rapping. The weighted time average of the measurements was compared with data
obtained during normal operation to determine if losses were affected by use
of the non-rapping intervals. During the other two field tests, measurements
were made during non-rapping periods and compared with measurements during
either normal operation or operation with a modified rapping frequency.
RESULTS OF EXPERIMENTAL STUDIES
In the tests on the five electrostatic precipitators, the following
data were obtained: the percent of total emissions due to rapping reentrain-
ment, the percent of emissions for a given particle size due to rapping, the
particle size distribution of the particulate emitted due to rapping, and the
effects of rapping reentrainment on fractional collection efficiencies. Selected
examples of the test results are contained in this paper.
The percentage contribution of rapping reentrainment to total emissions
at each of the test locations and for various test conditions are tabulated in
Table 1. The percentage contributions ranged from a high of 53% to a low of
6.5%.
The percentages of emissions for a specific particle size that are
due to rapping are given in Figure 4. This data indicates that rapping re-
entrainment accounts for only 2-5% of the emissions of 1 pm diameter particles,
and that the percentage of emissions due to rapping increases significantly
with increasing particle size, ranging from 24% to 48% for 4 urn diameter particles
and from 50% to 88% for 12 ym diameter particles.
Averaged particle size distributions for rapping puffs are plotted in
Figures 5 and 6. The particle size distribution data shows that the rapping
puffs consisted of relatively large particles when compared with non-rapping
emissions.
322
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A selected example of the effect of rapping reentrainment on precipitator
fractional efficiency is shown in Figure 7. Fractional efficiencies for alternating
periods with the rappers off and with the rappers on are given along with fractional
efficiencies determined during normal rapper operation. These data show that
collection efficiency is relatively independent of rapper operation for particles
with diameters from 0.1 ym to 2 ym with the possible exception in the region of
0.3 to 8 ym in which there is some ambiguity in the data. Above 2 ym, rapping
significantly reduced collection efficiency.
Discussion of Experimental Results
The percentage contribution of rapping reentrainment to total emissions
varied significantly for changes in test conditions at the same location and
for the different precipitators. The lowest rapping emissions were apparently
obtained for a precipitator with high plate accelerations, i.e., accelerations
greater than 150 G's (1.5 x 10s cm/sec ) and with a rap interval of 30 minutes.
There are several possibilities for the apparent low rapping emissions at this
location that are of interest and which merit further investigation. One is
the use of large and relatively deep hoppers (depth 6 meters). It is suspected
that these aided in reducing hopper "boil-up", which is discussed in the section
on reentrainment mechanisms. Another possibility is the effect of having a
relatively fine inlet dust. There also is the possibility that other sources of
reentrainment may have contributed to non-rap emissions, resulting in the apparently
low rapping emissions.
The highest percentage rapping emissions were obtained during the pilot
studies with a rapping interval of 12 minutes and with normal plate accelerations
on the order of 11 G's (1 x 1011 cm/sec2). The rapping emissions during the
pilot study (test location number 2) decreased with increasing time between raps.
The effect on overall efficiency is shown in Figure 8. The percentage of the
collected dust removed from the precipitator plates also increased with increased
time between raps, as shown in Figure 9.
The collecting of a larger mass per unit area (dust surface density)
produces .the two effects mentioned above. According to the elementary theory
of dust removal, the product of the normal plate acceleration and dust surface
density has to be greater than the tensile strength of the layer, typically
0.5 x 103 to 3 x 103 dynes/cm2, if the dust layer is to be removed. This theory
indicates that dust removal efficiency should increase with increased time between
raps, as illustrated by the data shown in Figure 9. Laboratory measurements by
Sproull8 produced similar data. Figure 9 also illustrates the build-up of a
residual dust layer that was not removed with normal plate accelerations on the
order of 11 G's. There are several possible causes for the development of the
residual layer. For one, the dust layer directly in contact with the collection
plates has a much higher tensile strength than the remainder of the layer. Estimates
for removal of the layer called for accelerations greater than 103 G's
(9.8 x 105 cm/sec2).7 Consolidation of the dust that remains on the plate after
a rap also aids in producing residual layers. The vibrations during a rap can
have the effect of compacting the dust layer if it is not removed making it more
difficult to remove. A third possible cause of the residual layer is the removal
of patches of dust only from selected locations on the collection plates where
the removal criteria are met. Dust can be removed from one location during one
rap and from another location on the next rap due to changes in distribution
of the dust surface density. This results in a nonuniform dust layer and the
323
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presence of a residual layer. This is often the result of nonuniform plate
accelerations. At one location where plate accelerations are on the order of
only 4 to 5 G's (3.4-4.9 x 103 cm/sec2) residual dust layers as thick as 2 cm
were observed in the vicinity of plate baffles where the plate accelerations
are dampened. Between the baffles, the residual layers were only 1 to 2 mm thick.
The decrease in rapping emissions observed during some of the tests
with an increase in the time between raps or correspondingly an increase in
dust surface density appears to be the result of a shift in the size distribution
of the reentrained particulate to larger particle sizes. Thicker dust layers
appear to produce larger particles. In this case the large particles that
are emitted are not the primary particles that entered the precipitator but
agglomerates consisting of smaller particles. Inspection of impactor substrates
after collecting dust at the outlet of test locations number 2 and 3 indicated
that the majority of the large particles in the rapping puffs were agglomerates.
The shift to larger particles in the rapping puffs observed during the pilot
studies as the time between raps was increased is illustrated by the shift of
the cumulative percent particle size distribution of the rapping puffs shown
in Figure 5.
This effect was previously observed by Plato,11 who found that with a
dust layer surface density greater than 1 kg/m dust was removed in cakes and
for values less than this it formed clouds. Producing agglomerates instead of
individual particles when the plates are rapped reduces the quantity of material
that escapes from the precipitator due to rapping. This is because the large
agglomerates are recollected faster by gravitational settling and by electrostatic
precipitation than discrete single particles or smaller agglomerates.
The above provides an explanation for the results that are given in
Figures 8 and 10. Sanayev and Reshidov have also used the above explanation for
the decrease in the percentage of dust collected on the plates that is emitted
due to rapping with an increase in dust surface density. A comparison of results
obtained during the pilot studies'* and by Sanayev and Reshidov12 is shown in
Figure 11.
Both sets of data indicate that for the conditions used during the tests,
minimum reentrainment loss occurred with a dust surface density on the order of
1 kg/m or larger. The Sanayev and Reshidov data also show the expected dependence
on gas velocity. The results given in Figure 11 do not include the effects of
collecting thicker layers on overall performance.
The 50 to 88% emissions due to rapping reentrainment observed for large
particles (>10 ym) which are shown in Figure 4 can be explained either by a con-
tinuing transfer of collected large particles from the inlet to the outlet by
rapping or by the above mentioned agglomeration of the collected particles. Both
effects probably occur, with agglomeration being the dominant effect, as indicated
by the agglomerates observed on the upper stage of inertial impactors.
During collection by an electrostatic precipitator, the mass median
diameter of the fly ash from a coal-fired boiler typically decreases from the
inlet of the precipitator to the outlet. This is illustrated by a shift in the
particle size distributions of the inlet and outlet dust concentrations. In
324
-------
Figure 6, the dm/dlogD particle size distributions are plotted for the inlet
dust at test location number 4 and for the outlet dust with the rappers off and
with the rappers on. The expected shift in particle size which can be theoretically
predicted for the inlet and outlet dust concentrations with the rappers off is
clearly evident. When the plate rappers are on, the mass median diameter of the
particle size distribution shifts back towards larger size particles as shown
by the dm/dlogD plot in Figure 6 with the rappers on and by the effective average
dm/dlogD for the rapping puffs. The large particle size of the particulate in
the rapping puffs is also illustrated by the data given in Figure 5. The rapping
emission percentages in Figure 4 reflect the shift in the outlet particle size
distribution produced by plate rapping. At locations such as at test location
number 3 where the precipitation is very large and preceded by a mechanical collector
almost all of the large particles (£lO ym) are collected before reaching the
outlet. Thus the percentage of emission due to rapping can approach 100% for
the large particles.
Rapping reentrainment has the effect of transferring particles from
particle size intervals consisting of small single particles to particle size
intervals consisting of larger particles.
At test location number 4, data were obtained at normal operating current
densities (see Table 1) and at one-half the normal operating current density.
This resulted not only in a decrease in efficiency between raps but also, as
shown by the data given in Figure 4, in an increase in the percentage contribution
of rapping reentrainment for large particles. This implies that a change in
particulate properties such as an increase in resistivity that reduces operating
current density can result not only in a decrease in efficiency between raps but
also an increase in the percentage emissions due to rapping reentrainment.
However, since the forces holding the dust layer together theoretically increase
with resistivity, the effects may not be the same as those obtained during the
test at location number 4, where dust resistivity was held constant.
Reentrainment Mechanisms
The amount of rapping reentrainment that occurs depends on the processes
by which the dust is reentrained into the gas stream. Sproull10 has suggested
that the rapping processes occur in a sequence in which the collected dust is
removed by what is termed a slip-stick process. The concept is that the dust
layer builds up to a sufficient thickness so that the acceleration of the plate
causes a shearing force between the dust layer and the plate. The dust then
slides along the surface of the plate until it is recollected on the plate. The
process is repeated with each successive rap until the dust finally falls into
the hopper.
Another concept of rapping is that the collected dust is dislodged in
large agglomerates or clumps and falls directly into the hopper without being
recollected.
Neither concept adequately explains the rapping process as observed in
the FluiDyne pilot precipitator or in the Southern Research Institute small
scale precipitator. Motion pictures (32 frames/sec) of the removal of a dust
layer (2-3 mm thick) by rapping in the SRI pilot unit show the dust layer frac-
turing along lines of discontinuity in the dust surface. The resulting fractured
325
-------
sheet of dust starts to fall as separate sheets which break up as they en-
counter other falling sheets and patches of unremoved dust. The dust appears
to fall without being recollected and to become turbulently mixed as it falls.
The motion pictures show the majority of the dust dropping into the hoppers
from which a portion boils up and becomes reentrained into the gas stream.
Motion pictures taken in the large pilot precipitator at the FluiDyne Engineering
Rosemount Laboratory showed similar behavior.
Measurements of the vertical distribution of the rapping loss at the exit
of the FluiDyne unit indicated that 82% of the rapping emission occurred in
the lower 32% of the precipitator. This effect, evidently due to hopper
"boil-up" combined with gravitational settling of the reentrained particulate
produced a significant vertical stratification of the reentrained emissions
from the large pilot precipitator as shown in Figure 12.
Rapping puffs observed in the lower portion of the precipitator occurred
in two bursts for both upstream and downstream raps as shown in Figure 13. The
first burst lasted 2-4 seconds. This burst was interpreted as being the result
of particulate reentrained directly in the gas stream and being carried out of
the precipitator at the velocity of the gas through the unit. The longer lasting
second burst, which for the larger particles was a series of puffs, can be inter-
preted as resulting from hopper "boil-up". These data indicate that hopper
"boil-up" contributes significantly to rapping reentrainment emissions.
EFFECT OF RAPPING REENTRAINMENT ON COLLECTION EFFICIENCY, PENETRATION, AND
SIZING OF ELECTROSTATIC PRECIPITATORS
Rapping reentrainment can have a considerable effect on collection effi-
ciency, penetration, and sizing of electrostatic precipitators. In this section
two approaches to estimating the effects of rapping reentrainment are discussed.
The examples are based on the results of the FluiDyne pilot tests and computer
calculations by Gooch, et al.5 The results of the example calculations are
tabulated in Tables 2 and 3.
These examples show that the percentage increase in collection efficiency
obtained with no rapping reentrainment decreases with increasing efficiency of
the precipitator. However, due to the exponential dependence of precipitator
efficiency on particle migration velocity, plate collection area, and gas volume
flow rate, the effect of rapping reentrainment on the size of a precipitator
required for a given collection efficiency is substantial. Estimates of the
effect of rapping reentrainment on precipitator size are tabulated in Tables 2
and 3 for various conditions. Two approaches were used to generate the tabulated
values. One approach was to assume that the percentage penetration due to rapping
was independent of the size of the unit. This assumption, although not strictly
true, may fairly represent real situations. According to the data obtained dur-
ing the pilot experiments, the rapping puff concentration ;has a significant
vertical gradient and the rapping puff consists of large particles. Thus, in
a multi-section precipitator the rapping puffs from upstream sections may be
recollected in the lower portion of a downstream section and later transferred
from this lower position into the hoppers without making any significant con-
tribution to outlet emissions. In this case, outlet fields would contribute
the largest portion of the rapping emission, and the percentage emission due to
rapping should be relatively independent of the size of the unit. Estimates
based on this approach are tabulated in Table 2.
326
-------
The second procedure was developed by Francis.^ This approach assumes
that the fraction of collected dust per section reentrained and emitted is the
same for every section, that the between-rap sectional collection efficiency is
constant from one section to another, and that the recollection efficiency for
the rapping puffs is the same as the between rap sectional collection efficiency.
According to Francis' procedure, the penetration of any downstream section is
equal to the penetration of the first section (P^ raised to the power of the
number of preceding sections plus one. Thus, the penetration of the N section
(PN) is given by:
PN = (Pl)N (3)
Using this equation, the collection efficiencies of a precipitator having
1, 2, 3, or 4 sections with a between-rap sectional collection efficiency of
94.8% were calculated for percent penetrations of the first section due to rapping
reentrainments of 52%, 30%, and 15%. These percentages correspond to 5.9%, 2.3%,
and 0.97% of the collected material per section being reentrained and emitted
from a section. These values were chosen to correspond approximately with those
obtained during the FluiDyne experiments. The results of the calculations based
on this procedure are tabulated in Table 3. According to this procedure, for the
conditions corresponding to the worst rapping reentrainment case at FluiDyne, the
contribution of rapping reentrainment to total emissions would be greater than
90% for a three or four section precipitator with 94.8% sectional collection
efficiency. For a case in which reentrainment accounts for only 15% of the
emissions for the first section, a situation slightly better than the best ob-
tained during the FluiDyne experiments, rapping reentrainment, according to
Francis1 procedure, would account for 39% of emissions for a 3-section precipi-
tator and 50% of emissions for a 4-section precipitator, if the precipitator
had a sectional efficiency of 94.8%. These calculations indicate that, as the
size of the precipitator increases, the percentage contribution of rapping re-
entrainment to total emissions increases. As previously discussed, this may not
happen. Investigations of rapping reentrainment in multi-section precipitators
do not appear to show a general increase in the percentage contribution of rap-
ping reentrainment to total emission with precipitator size except for large
particles. To obtain an indication of the effect of rapping reentrainment on
the sizing of a precipitator, estimates of the theoretical specific collection
area needed to obtain a given collection efficiency for a particular set of
conditions were obtained from the model developed by Gooch, et al.5 For the
example chosen, a typical particle size distribution for an ash from a coal-fired
boiler plant was selected and a precipitator operating current density of
20 nA/cm2 was assumed (see the report of Gooch, et al,5 for further details).
The theoretical specific collection area (SCA) required to obtain a de-
sired collection efficiency with rapping reentrainment was compared with a
theoretical SCA that would give the same efficiency if there were no rapping
emissions. This gave an estimate of the increase in SCA needed to counterbalance
the rapping reentrainment emissions. The theoretical SCA needed to obtain the
desired collection efficiency with rapping reentrainment was determined by first
computing the collection efficiency that would be obtained if there were no rapping
reentrainment, by assuming that either the percentage of emissions due to rapping
reentrainment or the fraction of collected dust reentrained and emitted from a
section due to rapping were known. Once this collection efficiency was calculated
and assuming that there were no other nonidealities beside rapping reentrainment,
327
-------
the required theoretical SCA was obtained from data given by Gooch, et al.
Then, again using the data given by Gooch, et al, the SCA was determined that
would give the desired efficiency if there were no nonidealities such as rapping
reentrainment. For the particular cases considered, the increase in required
SCA due to the rapping reentrainment ranged from 6% to 84% (see Tables 2 and 3) .
Thus, although the percentage decrease in collection efficiency due to rapping
reentrainment is relatively small, the effect on the size required for a given
collection efficiency can be significant. The data obtained from the FluiDyne
pilot experiments indicates that even moderate changes in rapping parameters can
significantly affect the size of a precipitator for a given collection efficiency.
SUMMARY
Experimental investigations of rapping reentrainment were conducted
utilizing a large pilot electrostatic precipitator and several electrostatic
precipitators installed at coal-fired power plants. The objectives of the
experimental study included a fundamental study of the mechanics of removal
of dry fly ash by rapping and the quantification of rapping reentrainment in
terms of the percentage and particle size distribution of the reentrained fly
ash.
During the experiments the percentage of fly ash removed from the plates
was observed to depend on the mass per unit area of ash collected on the plates,
as predicted by theory. The build-up of residual dust layers was also observed.
During the tests the contribution of rapping reentrainment to the total
emissions ranged from 6.5 to 53%, depending on the rapping conditions and the
precipitator design. A significant decrease in rapping emissions was obtained
by increasing the time interval between raps. This decrease was due to the
resulting larger mass per unit area collected on the plates when they were rapped.
Particle size distribution measurements showed that the mass median diameter of
the particles emitted during the raps increased with increased time between raps.
As would be expected, this produced lower overall emissions. The increase in
the size of the particles emitted during rapping was ascribed to an increase in
the extent of agglomeration of the particulate removed from the precipitator
plates with the corresponding increase in the mass per unit area collected on
the plates (thicker dust layers) before the plates were rapped.
A major portion of the reentrained material was observed to result
from hopper "boil-up." A small portion of the dust was observed to pass out a
pilot precipitator in a short burst at the velocity of the gas passing through
the unit, while the remaining portion of the material was observed to fall into
the hoppers, then to rebound, and finally to escape slowly over the baffles and
out of the precipitator. This produced a significant vertical gradient in the
concentration of the dust emitted from the precipitator due to rapping reentrain-
ment.
The fractional collection efficiencies of the precipitators both excluding
and including rapping reentrainment emissions were determined. Rapping reentrain-
ment decreased the fractional collection efficiencies for particles larger than
2 ym. Below 2 ym, rapping reentrainment did not appear to contribute significantly
to overall emissions.
328
-------
Two approaches using simplified assumptions were used to estimate the
effect of rapping reentrainment on the size of a precipitator. The first
assumed that a fixed percentage of emissions was emitted from a precipitator
independent of the size of the unit, and the second assumed that the same per-
centage of collected material was reentrained and emitted from each section
and that the recollection efficiency for the reentrained material for all sec-
tions was the same as the efficiency between raps, with all sections having the
same efficiency- The estimates based on the above approaches showed that the
increase in precipitator size needed to recover the rapping reentrainment emis-
sions can account for 6% to greater than 80% of the size of the unit for the
cases considered.
The parameters studied during the experiments were somewhat limited. A
data bank for detailed modelling of rapping should include additional data on
rapping emissions as a function of plate accelerations and gas velocity.
The pilot experiments also indicated a need for a study of the parameters
affecting hopper "boil-up" and the development of methods for reducing it.
Additional studies on the effects of residual dust layers for high resistivity
dust and studies of dust cohesive properties are also needed to aid in determining
optimum rapping conditions.
ACKNOWLEDGMENTS
This work was conducted under various contracts from the Tennessee Valley
Authority, the Environmental Protection Agency, and the Electric Power Research
Institute.
329
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Co
u>
o
Table 1. PERCENTAGE CONTRIBUTION OF RAPPING REENTRAINMENT TO TOTAL EMISSIONS
Avg. Plate
Gas SCA*, Total Penetration
2
Test Type Acceleration Intervals Velocit
Location Comments Rapper G's Min. m/sec
1
2
3
4
5
X,y,x axis
Vibrator 3.7, 1.6, 2.9 4, 6.5, 6.5
Vibrator 3.5, 1.8, 2.6 8, 13, 13
SO 3
injection Vibrator 4, 6.5, 6.5
no rap
S03
in j ection no rap
pilot test Drop hammer 11, 16, 15 12 0.87
pilot test Drop hammer 32
pilot test Drop hammer 52
pilot test Drop hammer 150
pilot test no rap
Drop hammer 10, 10,
20, 20 1.52
two plates 60, 60
Drop hammer 10, 20, 1.25
60
two plates
Rotating 6, 6, 12, 1.83
Drop hammer 12
.y Densities** Temp m
: raA /cm2 °C m3/sec
7.6, 10.4, 137
15.3, 13.9
11.2, 11.2, 137
16.7, 12.5
11.2, 11.2, 132
16.7, 12.5
23.3 122 33
8.4, 9.1,
13.2 154 110
15.1, 12.3
11.1, 17.6, 157 48
22.7
4.6, 7.8, 157
11.3
14, 24.6, 162 52
33.0, 47.7
Penetration Due to Rapping
% Reentrainment , %
32
25
6
17
3
11
7
6
6
5
0
0
0
0
.1 46
.8 33
.2 48
.4
.2
.4 53
.6 32
.1 18
.9 25
.2
.85 30
.40 35
.95 38
.19 6.5
*SCA - specific collection area
**for each electrical field
-------
Table 2. EFFECT OF RAPPING REENTRAINMENT ON EFFICIENCY FOR RAPPING LOSSES
ACCOUNTING FOR 20%, 30%, 40% AND 50% OF EMISSIONS FOR VARIOUS OVERALL EFFICIENCIES
(first approach)
Assumed Penetration
percent of Due to
total pene- rapping
tration due reentrain-
to rapping ment.
reentrainment %
3.75
2.25
50 0.75
0.25
0.05
3.00
1.80
40 0.60
0.20
0.04
2.25
1.35
30 0.45
0.15
0.03
1.50
0.90
20 0.30
0.40
0.08
Without
rapping
reentrain-
ment,
%
3.75
2.25
0.75
0.25
0.05
4.50
2.70
0.90
0.60
0.06
5.25
3.15
1.05
0.35
0.07
6.00
4.60
1.20
0.10
0.02
Total
%
7.50
4.50
1.50
0.50
0.10
7.50
4.50
1.50
0.50
0.10
7.50
4.50
1.50
0.50
0.10
7.50
4.50
1.50
0.50
0.10
Assumed
Overall
Efficiency
n
%
92.50
95.50
98.50
99.50
99.90
92.50
95.50
98.50
99.50
99.90
92.50
95.50
98.50
99.50
99.90
92.50
95.50
98.50
99.50
99.90
Efficiency
without
rapping re-
entrainment ,
TlH, %
96.25
97.75
99.25
99.75
99.95
95.50
97.30
99.10
99.70
99.94
94.75
96.85
98.95
99.65
99.93
94.00
96.40
98.80
99.90
99.92
Increase in
n without
rapping re-
entrainment ,
%
3.90
2.31
0.76
0.25
0.05
3.24
1.88
0.61
0.20
0.04
2.43
1.41
0.46
0.15
0.03
1.62
0.94
0.30
0.10
0.02
SCA* required
For n without For n» or for n
rapping
re- with
rapping
entrainment , reentrainment ,
n,2/
m3/sec
26.2
41.8
68.9
26.2
41.8
68.9
26.2
41.8
68.9
26.2
41.8
68.9
(ft2/
1000 ft3/ m2/
sec) rnVsec
16.2
21.3
(133) 35.5
(212) 53.2
(350) 81.7
18.7
(133) 32.9
(212) 50.2
(350) 78.8
(133) 30.5
(212) 47.3
(350) 75.8
(133) 29.2
(212) 68.9
(350) 73.3
(ft2/
1000 ft3/
sec)
( 82)
(108)
(180)
(270)
(415)
( 95)
(167)
(255)
(400)
(155)
(240)
(385)
(148)
(350)
(372)
Additional SCA
required
because
of reentrainment
(ft2/
m2/ 1000 ft3/
raVsec
9.3
11.4
12.8
6.7
8.5
9.8
4.5
5.5
6.9
3.0
3.5
4.3
sec)
(47)
(58)
(65)
(34)
(43)
(50)
(23)
(28)
(35)
(15)
(18)
(22)
Addi-
tional
SCA,
%
353
27.4
18.6
25.6
20.3
14.3
16.5
13.2
10.0
11.3
8.5
6.3
*Specific Collection Area (SCA) based on data given by Gooch, et al.
(Figure 26 computed performance curves at 20 nA/cm ).
-------
Table 3. EFFECT OF RAPPING REENTRAINMENT ON PERFORMANCE FOR VARIOUS REENTRAINMENT
PERCENTAGES AS A FUNCTION OF NUMBER OF SECTIONS AND A CONSTANT COLLECTION EFFICIENCY PER SECTION.
OJ
U)
ro
Assumed
percent of
collected
material
per sec-
tion re-
entrained
5.9
2.3
0 97
SCA* Needed
Penetration
Number
of
Sections
1
2
3
4
1
2
3
4
1
2
3
4
Due to
rapping
reentrain-
ment.
%
6.2
1.02
0.136
0.0193
2.23
0.28
0.027
0.0023
0.90
0.10
0.009
0.0007
Without
rapping
Percent of
penetration
due to rap-
reentrain-
ment,
%
5.2
0.27
0.014
0.0007
5.20
0.27
0.014
0.0007
5.20
0.27
0.014
0.0007
ping reen-
Total tn
11
1
0
0
7
0
0
0
6
0
0
0
%
.4
.29
.150
.020
.43
.55
.041
.0030
.1
.37
.023
.0014
ainment.
%
52
79
91
97
30
51
66
77
15
27
39
50
Efficiency
Assumed
efficiency
Tl without
rapping re-
T] , entrainntent ,
'.
88.
98,
99,
99.
92.
99.
99.
99.
93.
99,
99.
99.
I
.6
.7
.85
.98
.57
.45
.959
.997
.9
,63
.97
,9986
94
99
99
99
94
99
99
99
94
99
99
99
%
.8
.73
.986
.9993
.80
.73
.986
.9993
.8
.73
.986
.9993
Increase
in n with
no rapping
Without rapping
reentrainment to
obtain T)
reentrain-
ment , nt /
;
6.
1.
0.
0.
2.
0.
0.
0.
0.
0.
0,
0.
I.
.54
.03
.14
.02
.35
.28
.03
,002
.95
.10
.016
,001
m3/sec
28.0
62.1
98.5
40.4
83.7
108.5
46.3
91.0
108.3
(ft2/
1000 ft3/
sec)
(142)
(315)
(500)
(205)
(425)
(550)
(235)
(462)
(550)
With
rapping
reentrainment
m2/
ma/sec
51.6
105.4
108.3
51.6
105.4
108.3
51.6
105.4
108.3
(ft2/
1000 ft3/
sec)
(262)
(535)
(550)
(262)
(535)
(550)
(262)
(535)
(550)
Additional SCA
m2/
m3/sec
23.6
43.3
11.2
21.7
5.3
14.4
(ft2/
1000 ft3/
sec)
(120)
(220)
( 57)
(110)
( 27)
( 73)
Addi-
tional
SCA,
%
84
69
28
25
11
16
*Specific Collection Area {SCA) based on data given by Gooch, et al.
(Figure 26 computed performance curves at 20 nA/cm2).
-------
GAS ANALYSIS
503, s°2- H2°
TEMPERATURE
TIME INTEGRATED
PARTICLE SIZE
MEASUREMENTS
RESISTIVITY
MEASUREMENT
MASS LOADING
VELOCITY
DISTRIBUTION
' TWO SETS AT 3 LOCATIONS: ONE TO MEASURE DURING
RAPS AND ONE TO MEASURE BETWEEN RAPS.
' TWO SEPARATE UNITS: ONE TO LOOK AT LOWER HALF OF THE
PRECIPITATOR OUTLET AND ONE TO LOOK AT UPPER HALF OF
THE PRECIPITATOR OUTLET,
HOPPER
SAMPLES
" REALTIME
PARTICLE SIZE
MEASUREMENTS
Figure 1. Block diagram of experimental layout for a rapping reentrainment study.
333
-------
SAMPLE
FLOW/RATE
MANOMETER
'X /j/
7 <
/
V*
<
•»
^
N>
/
^
GAS FLOW
\
PROCESS EXHAl
> PRORP LINE
\
HEATER
DILUTION AIR
MANOMETER
AND ORIFICE
LARGE PARTICLE
COUNTER
MAIN FRAME
VERTICAL
Figure 2. Schematic of diluter for real time particle sizing system.
BLOWER
EXHAUST
FLOW
REGULATOR
/
/
DILUTER
AND COUNTER
&^
DUCT TOP
EXTRACTION PROBE BE
GAS
' FLOW
Figure 3. Extractive sampling system for real time system
334
-------
100
oc
o
to
O
u
tr
LOCATION 4
NORMAL CURRENT
40
20
8 12
PARTICLE DIAMETER,
Figure 4, Percent emissions due to rapping for particles with diameters at I to 20 fj.m.
0.01 0.1 1 10 20 40 60 80
PERCENT LESS THAN INDICATED SIZE, by mass
Figure 5. Cumulative percent distribution for rapping puffs, rapping intervals
of 12, 32, and 52 minutes, pilot test.
335
-------
10
10'
GEOMETRIC MEAN DIAMETER,
Figure 6. Particle size distributions for inlet particulate, outlet particulate with rapping,
outlet particulate without rapping, and average effective distribution for
rapping puffs.
U.U 1
0.1
0.2
1
2
RATION, %
o ui
t 20
2
£ 30
40
60
80
I I 1 I 1 1 III 1 ll|
«
— —
— -w
- § . . .
_ w —
• g
8/S ^ n
• O a a °
i a o o ^
: •• , • • :
o o
— —
— * __
O RAPPERS ON TEST (8)
~ • RAPPERS OFF TEST (8) ~
Q NORMAL RAP INTERVAL TEST (6,7,9)
~ —
1 1 1 1 — 1 1 1 1 1 1 1 1 1 L_
»a.as
99.9
99.8
99
98
o\*
95 |
90 |
so a
70
60
40
20
10-'
ID"1 10°
GEOMETRIC MEAN DIAMETER, jjm
10'
Figure 7. Fractional efficiencies as function of particle diameter for half normal
current density test. Test Location Number 3.
336
-------
100
UJ
O
LL
95
90
85
WITHOUT RAPPING
20
60 100 140
TIME INTERVAL BETWEEN RAPS, min
180
Figure 8. Average efficiencies for FluiDyne pilot precipitator for various rapping intervals.
COLLECTED BETWEEN RAPS
0 % OF TOTAL PARTICULATE
COLLECTED ON THE PLATES _|
40 60 80 100 120
TIME INTERVAL BETWEEN RAPS, minutes
0.26 0.78 1.3 1.8 2.3 2.9 3.4
MASS/AREA GAINED BETWEEN RAPS, kg/m2
140
160
Figure 9. Dust removal efficiency versus the time interval between raps.
337
-------
100
95
90
I 85
80
75
70
65
NO RAP
NO RAP
2N
INTERVAL BETWEEN RAPS
Figure 10. Precipitator efficiency at Test Location Number 1 as a function of the
rapping interval.
o
Q
UJ
O
uLJ
O
o
fe
D
Q
U-
O
Ol
o
a:
SPENCER4 GAS VELOCITY 1 m/sec
SANAYEV & RESHIDOV15 GAS VELOCITY 2 m/sec
SANAYEV & RESHIDOV15 GAS VELOCITY 2 m/sec
DUST SURFACE DENSITY, kg/m2
Figure 11. Percent of dust collected on precipitator plates emitted from pilot
precipitators due to rapping.
338
-------
106
z
S 105
z
h-
z
0
u
111
o
W -,
N £
55
_j ^
24 um
^
^^
^ • •••
~~ * • * *"
_
A
^
A *
A ^
^
I I I I I t, I I I I
B 20 40 60 80 100' ' 330 350 370 390 41 (
A
F
p DISTANCE FROM BOTTOM BAFFLE, cm
L
E
Figure 12. Spatial distribution of particles in rapping puff.
339
-------
50
40
24 Hm DIAMETER PARTICLES
o
o
z
o
u
20
10
500
400
300
200
100
Li
12 - 24 urn DIAMETER PARTICLES
5x10J
| 4
o
c- 3
O
O
6 12 urn DIAMETER PARTICLES
3-6,nm DIAMETER PARTICLES
UPSTREAM
1
O
5x10
I
D 2
O
O
1
DOWNSTREAM
1.5 - 3 jjm DIAMETER PARTICLES
I I I IJ I I IVJ I I i
TIME, 12 sec/div
Figure 13. Rapping puffs at the exit plane of the pilot precipitator, upstream and
downstream raps.
340
-------
REFERENCES
1. White, H. J. Industrial Electrostatic Precipitation. Addison-Wesley
Publishing Company, Inc., Reading, Mass., 1963, p 331.
2. Nichols, G. B., H. W. Spencer, and J. D. McCain. Rapping Reentrainment
Study. Report SORI-EAS-75-307 to Tennessee Valley Authority, TVA Agree-
ment TV36921A, November 1975.
3. Gooch, J. P-, H. W. Spencer, J. D. McCain, G. H. Marchant. A Study of
the Performance of a Full Scale Electrostatic Precipitator with a High
Specific Collection Area. Report to Electric Power Research Institute,
EPRI Project RP 413-1, December 1975. (draft)
4. Spencer, H. W. A Study of Rapping Reentrainment in a Nearly Full Scale
Pilot Electrostatic Precipitator. Environmental Protection Agency
Publication No. EPA-600/2-76-140. May 1976.
5. Gooch, J. P., J. R. McDonald, and S. Oglesby, Jr. A Mathematical Model
of Electrostatic Precipitation. Environmental Protection Agency,
Washington, D.C. Publication No. EPA-650/2-75-016. April 1975.
6. Tassicker, O. J. Some Aspects of Electrostatic Precipitator Research in
Australia. J. Air Pollut. Control Assoc. 25:122-128, February 1975.
7. Tassicker, O. J. Aspects of Forces on Charged Particles in Electrostatic
Precipitators. Dissertation, Wollongong University College, University
of New South Wales, Australia, 1972.
8. Sproull, W. T. Fundamentals of Electrode Rapping in Industrial Electrical
Precipitators. J. Air Pollut. Control Assoc. 15:50-55, February 1965.
9. Penney, G. W. and E. H. Klingler. Contact Potentials and Adhesion of Dust.
Trans. Amer. Inst. Elec. Bug. Part I 81:200-204, 1962.
10. Sproull, W. T. Minimizing Rapping Loss in Precipitators at a Megawatt Coal-
Fired Power Station. Paper No. 71-125, APCA Meeting, Atlantic City, N.J.,
1971.
11. Plato, H. Rapping of Collecting Plates in Electrostatic Precipitators.
Staub-Reinhalt. Luft (in English) 29:22-30, August 1969.
12. Sanayev, Yu. I. and I. K. Reshidov. Study of Dust Reentrainment Phenomena
and Their Influence on Efficiency of Industrial Electrostatic Precipitators.
Promyshlennaya Sanitarnaya Ochistka Gazov, Industrial and Sanitary Gas
Cleaning Inst. (Moscow USSR) 1974(1):l-5.
13. Schwartz, L. B. and M. Lieberstein. Effect of Rapping Frequency on the
Efficiency of an Electrostatic Precipitator at Municipal Incinerator.
Proceedings of the Fourth Annual Environmental Engineering and Science
Conference, March 4-5, 1975, University of Louisville, Kentucky-
14. Gooch, J.P. and N. L. Francis. A Theoretically Based Mathematical Model for
Calculation of Electrostatic Precipitator Performance. J. Air Pollut.
Control Assoc. 25:108-113, February 1975.
341
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FABRIC FILTRATION PERFORF1ANCE MODEL
Douglas W. Cooper
and
Vladimir Hampl
GCA Corporation
GCA/Technology Division
Burlington Road
Bedford, Massachusetts
343
-------
FABRIC FILTRATION PERFORMANCE MODEL
by
Douglas W. Cooper
Vladimir Hampl
GCA CORPORATION
GCA/TECHNOLOGY DIVISION
Burlington Road
Bedford, Massachusetts 01730
ABSTRACT
The goals of this program were to formulate and develop mathematical methods for
predicting and analyzing fabric filter drag (pressure drop per unit face velocity)
and fabric filter collection efficiency as a function of particle size. The liter-
ature with respect to filtration is vast, but the efforts to model fabric filtration
have been few. Most attempts have used a linear approximation to the increase of
fabric drag with added weight per area, and those who attempted to model collection
efficiency did so with models based upon arrays of cylinders. The woven fabric struc-
ture is much more like an array of pores (before the cake is formed) and this repre-
sentation has given rise to a new, nonlinear model of fabric drag and a new model of
fabric collection efficiency, both presented here. The nonlinear drag model fits the
experimental data much better than does the linear. The efficiency model appears to
give results which are both reasonable and informative. These models have been com-
bined to form a preliminary computer program for an entire multi-bag filtration
installation, and results of such modeling are presented for both flow and efficiency
variables. Also presented are design equations which are readily adaptable for a
programmable calculator„ Finally, we indicate several areas in which further in-
vestigation seems warranted in analyzing cleaning and seepage.
For presentation at the EPA/EPRI symposium, "Particulate Control in Energy
Processes," San Francisco, 11-13 May 1976
344
-------
FABRIC FILTRATION PERFORMANCE MODEL
Introduction
This presentation indicates our progress toward development of a predictive model for
industrial fabric filtration, with special emphasis on the use of fabric filtration to
control the particulate emissions of coal-fired power plants. Primary goals are the
prediction of collection efficiency versus particle size and the relationship between
the flow per unit area and the pressure drop for a fabric filtration installation.
The following factors can be expected to influence efficiency and the pressure/flow
relationship:
1. Dust — chemical composition, particle size distribution, particle
shape, particle phase, particle concentration.
2- Gas - chemical composition (especially moisture content), temperature,
pressure.
3. Fabric — material, weave (including fill and warp counts), finish,
history (especially accumulated dust).
4. Cleaning operations - cleaning type, intensity, frequency, duration.
Ideally, we would like to be able to predict fabric filter drag as well as efficiency
for all combinations of the above, but this is a very ambitious task. In forming
a useful model, one treads a narrow path between untractable complexity and imprac-
tical simplicity — this seems particularly difficult in modeling fabric filtration.
Description of a Filtration System. Figure 1 shows a very simple schematic of
a fabric filter installation. A dust-laden flow of gas enters the filter installa-
tion with a volumetric flow rate Q and a concentration c. The flow is divided among
n compartments, the bags within these compartments having average dust weight per
,t area of fabric given by W
;s is given by the equation
unit area of fabric given by W.o The pressure drop across the i compartment's
Ap. = Q^/A^ (1)
in which Q. is the volume flow rate, A. is the cloth face area and S. is the drag
caused by the fabric and any accumulated dust. The ratio S./A. is analogous to elec-
trical resistance (with Q. analogous to current and Ap analogous to voltage). The
total flow, Q, is the sum of the individual flows, Q. (as long as temperature and
pressure corrections are made). Usually the compartments are in parallel and so
constructed that the pressure drop at any given time is the same across all of them,
so that Ap. just becomes Ap.
345
-------
N
^p
/
3
c
w
Si
QI
w
S2
2
Q2
W
= 3
3
0,
W
S4
4
Q4
w
».
n
0,
Figure 1. Schematic of n-compartment baghouse.
The relationship between the pressure drop and the volume flow for a particular
installation will depend upon the locus of the intersections of the system fan curve
and the system resistance curve, each of which can be expressed as volume rate of
flow versus pressure drop. Usually, one of the following conditions holds approx-
imately for the installation:
1. The fan produces a constant volume of flow, and pressure drop
changes with system resistance.
2. The fan produces a constant pressure drop, and the volume flow
changes with system resistance.
Objectives. The cost of the installation will depend upon the fabric chosen and
its total area, among other things, and one of the operating costs will be the power
consumption due to the pressure drop across the system. Often the available choice
of fabrics will be limited, so that the major question becomes the area of cloth
needed to handle a volume flow rate of gas.
The cloth area required has as its criterion "operation at an acceptable pressure
2
drop across the cloth for a predetermined cycle." One might add that this assumes
that the collection efficiency is adequate under these conditions. Thus one impor-
tant goal is to be able to predict the pressure drop for a particular dust and fabric
combination at a given ratio of air flow to cloth area, the face velocity. The drag
will depend upon how much dust is on the filter surface, how it is distributed, its
geometrical nature in cake form, the geometry of the fabric, and the viscosity of
the gas. A second goal is to predict the emissions from such a system.
346
-------
Outline of Model. A procedure for calculating the pressure/flow relationship
and the efficiency of an installation can be developed by subdividing the fabric area
into smaller homogeneous sub-units (compartments, bags, or areas on bags) and then
doing the following:
!• Calculate the drag (the pressure drop per unit face velocity)
for the sub-unit.
2. Determine the flow from the drag and the instantaneous
pressure drop.
3. Determine the penetration, the fraction of the particulate
concentration reaching the sub-unit which then penetrates
to the clean air side.
4« Calculate the emissions rate from the sub-unit (penetration
times inlet concentration times volume flow rate).
5« Calculate the new dust loading of the sub-unit.
6. Determine the new pressure drop or the new total flow
rate by combining the resistances of the sub-units
according to the law for the addition of parallel
resistances:
n
Q = Apf = Ap 2 A./S, (2)
S 1-1 L 1
S =
n
(3)
2 Vsi
To develop a time profile of the performance, this procedure must be done iteratively,
with any cleaning taken into account as well.
Developing this kind of model was our goal. With such a model one should be able to
predict the collection efficiency and the relationship between flow and pressure drop
for fabric installations for reasonable combinations of variables relating to dust,
gas, fabric, and cleaning. Another part of the work at GCA has been laboratory in-
3
vestigations by Dennis et al., the results of which have influenced aspects of this
model.
Models for Drag - Background
Much of the work in filtration has been done on systems other than fabric filters.
Of the fabric filtration work, most of the predictive modeling has been directed at
forecasting the pressure drop given a fabric and dust combination.
347
-------
Robinson et al. , for example, produced a three-compartment model. For the increase
of drag in any one compartment with increased weight per unit area on the filter,
they used the linear drag expression:
- s • SE + K2 w w
in which the term S is the effective drag, the drag on the fabric obtained by
JLj
extrapolating the linear portion of the actual drag versus weight curve (S versus
W) back to W = 0. By solving the equations for the total drag at constant pressure
4 th
drop for the three-bag system, Robinson et al. fit the flow rate of the i bag
Q. = a.txi (5)
where t. is the time since last cleaning. They used this flow to obtain the loadings
on each bag as a function of time. They combined the individual compartment drags
to get the total drag for the system, developing a complicated approximate expres-
sion that is only appropriate for a three-compartment system, although the technique
can clearly be extended to many-compartment systems.
Solbach also used the linear drag versus weight model, Equation (4). For the case
for which the pressure drop, Ap, is fixed but the flow is changing, Solbach expressed
the increase of drags due to the increase of the dust loading as
dS = K2 c V dt (6)
and integrated this to give the velocity of the gas flowing through a single compart-
ment at any time (t = 0 right after cleaning):
Solbaclf extended this to a multicompartment system by averaging V to obtain the expressioi
(8)
This expression is derived assuming constant pressure drop and infinitely many com-
partments. It is approximately correct (assuming linear drag versus weight) for
systems with much more than a few compartments and even for constant flow in such a
system rather than for constant pressure drop. (The reason it is nearly correct for
a many-compartment system is that the parallel flow arrangement tends to even out the
348
-------
pressure drop changes.) One obtains the required total filter area from the total
volume flow rate and the average face velocity
A = Q/V_ (9)
av
6
Dennis and Wilder derived an equation for the average drag for a changing pressure
drop:
Sav = av/Vav = SE + K2Wc/2
Wc = cVav fcc
It can be shown that this result and Solbach's agree: in both cases, the average drag
is just the drag at half the final cloth loading (W ).
This is not an exhaustive list of the investigators who worked with linear models, but
it is meant to present the linear drag multi-compartment model. We note that the linear
models cannot describe the nonlinear part of the performance curves and there remains
the problem of obtaining both parameters K? and S_.
To predict K9, one approach is to use the Kozeny-Carman equation (or variants), as
7
did Williams et al., a version of which is
K2 = (25/6)M(l-e)Sp2/Ppe3 (12)
where K2 = cake resistivity, sec
fi = gas viscosity, poise
€ = filter cake porosity (fractional void volume)
S = surface-to-volume ratio of particulate material, cm
P 3
Pp = particle density, g/cm .
K. is quite dependent upon particle size (increasing for smaller particles) and the
porosity of the cake deposit. (Unfortunately, K~ is often measured in such English
2
units as (inches WC/fpm)/(Ib/ft ).)
Fraser and Foley took a somewhat different approach, keeping the K2 particle size
dependence (and integrating over the particle size distribution), while using correc-
tion factors for particle shape, gas velocity and fabric-surface interaction, follow-
ing Billings and Wilder.
The effective drag, S , is really just an extrapolated value, artificial in that such
E «
a drag does not occur at that (W = 0) condition, so it is not surprising that there
is no theory for predicting it, only correlations. Often one must measure it for the
dust and fabric of immediate interest.
349
-------
The only nonlinear model of drag versus weight of which we are aware was presented
by Davis et al. which we learned about after beginning our rather different approach.
Davis et al. fit equations of the form
S = a Wb (13)
to data obtained for the nonlinear part of the drag versus weight curve. They found
correlation coefficients of greater than 0.90 for 104 of 106 data sets. Davis et al.
2 2
evaluated this at a loading of 0.0244 g/cm (0.05 Ib/ft ), where their drag data
generally became linear with W, to obtain S . They found that both a and b could be
lit
correlated with 1C, but an analysis of variance did not show that K~ was different
for different fabrics for a given dust, which is as expected, even though contradicting
12 11
the conclusion of Draemel. The model for drag presented by Davis et al. is useful
because it includes the nonlinear, cake-repair zone. The model we present also has
this feature, with the advantage that the parameters seem more amenable to theoretical
prediction rather than just empirical correlation.
Nonlinear Drag Model
The linear models of drag versus weight are incorrect for the low dust loading portion
of the filtration cycle, during which the flow and the emissions through the bag are
the greatest. The new fabric drag is S . The conditioned, cleaned fabric drag is S
O R
at residual loading W • Typically, the drag of the filter increases with added weight
R
per area at one rate, K , as the conditioned bag collects dust right after cleaning.
This rate (which is the slope of S versus W) changes until it reaches a value K»,
after which the rate remains constant (the S versus W curve becomes linear). The
linear part of the performance curve starts at loading W . An improvement on the
linear model would be one that starts with S and K , approximates the transition to
the linear zone, and matches the linear part of the drag curve.
Viewing the filter as an assemblage of pores (rather than cylinders), we assumed
that pores which were unclogged ("open") had a different dust resistivity (K ) than
pores which were caked (closed), assumed to have resistivity K , as has the cake.
We would expect KQ and K2 to be different because the velocity in the pores is much
higher than through the cake but the particles in the "open" pores are not packed
together the way they are in the cake. (Depending upon the magnitudes of such
velocity gradient and dust packing effects, K0 may be greater than K or vice versa.)
z o
By assuming that pores are either caked or open and that the decrease in the popula-
tion of open pores is proportional to the number of pores still open and to the in-
crease in fabric loading, dW, we obtained the expression:
350
-------
dN = -(1W*) N dW (!4)
where N = number of open pores per face area of filter
W* = fabric loading that is characteristic of pore-caking.
The fraction of pores still open at loading W becomes:
-W/W*
where NQ is the number of pores per area in the fabric when W = 0.
If the caked pores and open pores contribute to the drag in proportion to their
resistivities and their numbers, then it can be shown:
|| = KQ f (W) + K2 g(W) (16)
where f(W) = exp^ ' (the fraction of pores which are open) and
g(W) = 1 - f(W) (the fraction of pores which are caked).
By integrating Equation 16 we obtain:
( 1 - e~W/W*j
S = SQ + K2W + (KQ - K2) W* 1 - e~ (17a)
in which S is measured new cloth drag; alternately
H. - e~W'/W*J
S = SR + K2W + (KR - K2) W* H. - e~ (17b)
where W' = W - W_, the difference between the loading and the residual loading,
K
W at which S = S.,.
K. K
We hope to be able to predict K_ and W* from dust and fabric characteristics, but
thus far we too have had to rely on correlations. To show how well this form of
equation can fit data from an actual fabric, we present Figure 2, for a glass fabric
filtering mica dust and cleaned by shaking, and Figure 3, for Dacron filtering fly
ash and cleaned by a combination of shaking and reverse air. Both are fabrics
appropriate for high-temperature operations. The optimal choice of W* produces
excellent agreement with the measured values of drag versus weight. The three other
such comparisons we made were similarly successful; Tables 1 and 2 give more informa-
tion about the systems against which the model was compared in our Interim Report
to EPA.16
To use this model for a variety of filtration systems, we need theoretical equations
°r predictive correlations for the parameters KR, SR, and W* . . , assuming we have
the usual information about the fabric (permeability, from which comes SQ) and dust
(resistivity, K2).
351
-------
w, ib/ft.z
Figure 2. Theoretical and experimental curves
of glass fiber filter.
Figure 3. Theoretical and experimental curves
of Dacron-type filter.
352
-------
Table 1. TECHNICAL DATA — FABRIC FILTERS SELECTED FOR VERIFICATION
Type of fabric
Glass fiber
Polypropylene
Dacron
Cotton
Polyacrylester
Weight,
oz/yd2
9.06
4.30
10.0
10.0
9.8
Weave and yarn count
3/1 crowfoot, filament
55 x 58
3x1 twill, filament
74 x 33
Plain, staple
30 x 28
Unnapped sateen
95 x 58
2x2 twill, spun
39 x 35
Frasier
permeability,
ft3/min
@ 0.5 "H20
7.9
15.0
55.0
13.0
60.0
Reference
Spaite and
Durham
Dennis and
Dennis and
14
Durham
Walsh13
Wilder6
Wilder6
UJ
Ul
Table 2. EXPERIMENTAL DATA — FABRIC FILTERS SELECTED FOR VERIFICATION
16
Fabric type
Glass fiber
Polypropylene
DacrotL
Cotton
Polyacrylester
«2
KR
" WC-ft-min/lb
16.0
6.1
12.4
15.0
4.6
400.0
135.4
93.8
72.0
26.3
SR
" WC /f pro
0.84
0.35
0.08
0.5
0.05
WI
lb/ft2
0.0036
0.0135
0.0066
0.0075
0.030
Type of
cleaning
Shaking
Shaking
Shaking and
reverse air
Shaking and
reverse air
and shaking
only
Shaking
Dust type
Wet ground
mica
Fly ash
Fly ash
Fly ash
Fly ash
Remarks
Conditioned
velocity 2 ft/rain
Conditioned
velocity 4 ft/min
Used fabric
velocity 3 ft/rain
Velocity 3 ft/min
Conditioned
velocity 4 ft/min
Ref.
Spaite and
Walsh13
, 14
Durham
Dennis and
Wilder6
Dennis and
Wilder6
14
Durham
-------
The value of KR is expected to depend upon the dust, the fabric, and the cleaning, so
we thought it might correlate with K~ for the dust and So for the fabric for cleaning
by shaking. We calculated (K?/K ) for data on glass, ' Dacron, ' ' and Nomex
fabrics. (Aerosols of mica and fly ash were used with the glass and the Dacron
filters; the Nomex tests were with fly ash.) The correlation was
^ = 515 K2 S' (18)
where S is in the units inches WC/fpm and K,,, K0 are in inches WC-ft-min/lb.
O K ^
Similarly, correlations for the same five fabric-duct combinations gave:
(!„
(residual drag increases with fabric drag and dust resistivity) , and
W* = 7.68 x 10-4/K2°'913 So°'945 (20)
(pore caking requires more material for rather open fabrics and more coarse dusts) .
2
Again, the parameters are in English units (W in. Ib/ft and K in the units of S/W).
It should be mentioned, at least in passing, that data from the reported measurements
of K- we used were better correlated with the inverse of particle mass median diameter
rather than the inverse squared (the latter is predicted by the Carman-Kozeny equa-
tion ) and showed a face velocity effect (as has been reported before ).
To sum up: the form of our nonlinear expression fit quite well that data for which
we evaluated it. The parameters in the model should be theoretically predictable, a
problem on which we are working, but until they are, one can use correlations such as
those presented here. Such a model promises to be an improvement on the linear model
because it can describe that part of the filter performance for which flow and emis-
sions are greatest, the low- loading, nonlinear part.
Efficiency Models - Background
Before presenting our model for woven fabric filtration efficiency as a function of
particle size, we will discuss several recent efforts in that direction.
A theory of particle collection by clean cylindrical fibers was adopted by Fraser
9
and Foley for the evaluation of the collection filter efficiency, expressed as a
•I Q
function of three collection mechanisms. They used an equation derived by Davies
for low gas velocity (Re ~ 0.2) and single cylinders, incorporating impaction,
interception and diffusion. Then, they applied a correction factor to the single
fiber efficiency Y to account for the effect of the neighboring fibers
YM = (1 + 4.5/3) (21)
354
-------
to obtain the total collection efficiency. The penetration for a fiber was expressed
as:
Pn = exp
(22)
Due to the difficulties related to the precise measurement of the fiber diameter, they
18
used an effective fiber diameter, derived by Davies:
d
f
70 T) 01>5L(1 + 52 01'5)
Ap
1/2
(23)
(The filter packing density is B, its thickness is L, and the gas viscosity is n •)
O
That model of the penetration of the particles through the clean fiber filter was
extended to the dusted fabric filter under the assumption that the main effect of
the dust on the fabric itself is to change the effective fiber diameter. An equation
for the change of the thickness of the filtering layer during the filtration cycle
was developed by assuming that the dust fills up first the free volume of the filter
(at the packing density of the dust) before building up the dust cake. The residual
dust loading was expressed as a function of a cleaning parameter and permeability.
The collection efficiency used in that model was based on the calculation of particle
18
trajectories around a single fiber, derived by Davies for a clean fiber with a very
small diameter. It is doubtful that the same equation will hold for fiber filters
having fiber diameters ~10 um. If we assume the average face velocity of 2 cm/s
(4 ft/min), with fly ash particles having the diameter of 10 |itn, fiber diameters
approximately 800 urn (which is an average yarn size) and filter packing density
around 0.3, then the Stokes number equals 0.02, a lower value than the critical
Stokes number, below which the impaction mechanism does not contribute to collection
efficiency. (Also under such conditions the diffusion parameter is negligible.) Then
the penetration for the particles would be 0.99, considering the interference effect
of neighboring fibers to be (1 + 4.5 /3). Modeling the fabric filter as an array of
cylinders does not seem successful thus far.
19
Leith and First Model. By using tagged fly ash aerosols, Leith and First were
able to distinguish between those fly ash particles which, under laboratory condi-
tions, penetrated a needled felt fabric filter immediately and those particles which
exhibited a delay in their penetration. These researchers postulated three types of
penetration mechanisms: direct penetration, gradual seepage of the dust, and the
breakage and penetration of plugs of material in the vicinity of pinholes.
355
-------
The experiments were done at face velocities from 5 ctn/s (10 fpm) to 15 cm/s (30 fpm)
for dust cakes up to 60 |_im thick. These velocities are higher than those in normal
use for the filtration of fly ash (~1 cm/s) and these cake thicknesses are rather
low. The efficiency as a function of particle size was such that the penetration was
found to "remain relatively constant for particles from 0.3 to 4.0 micrometers in
diameter," The penetration increased with face velocity. From their figures for
mass flux versus time (at 10 cm/sec) we conclude:
1. Straight through mass flux seemed to decrease (roughly exponen-
tially) with deposit thickness, but may have gone through a
minimum near 20 um thickness.
2. Seepage mass flux remained fairly constant with time and
deposit thickness.
3. Pinhole plug mass flux decreased with increasing deposit
thickness.
The fraction of the total penetration which was due to pinhole plugs and seepage was
greater than the direct (straight through) contribution (at 60 urn thickness) for
15 cm/s but substantially less than the direct for 10 cm/s, which suggests the emis-
sions due to seepage and pinhole plugs might be very much less than the direct at
1 cm/s (2 fpm).
19
Leith and First postulated the following functional forms for the three types of
penetration:
-aXb
1. Direct penetration proportional to e , where X is the
deposit thickness
2. Constant seepage mass flux versus thickness of deposit
3. Pinhole plug mass flux proportional to Xe~C (in which
a,b,c, are proportionality factors).
The correlations they found in using these equations ranged from 0.86 to 0.91 corre-
lation coefficients, for mechanism-by-mechanism comparison, adding support for their
proposed mechanisms.
19
Leith and First found only a weak particle size dependence for efficiency, with the
efficiency decreasing slightly as particle size increased. The particle size depen-
dence, the velocity dependence, and their general appraisal of the filtration process
led them to conclude:
"Because media filtration theory does not describe the trends in
penetration found in a fabric filter, and was not developed for
the operational conditions found there, it should not be used to
predict or interpret the penetration characteristics of fabric
filters."
356
-------
This conclusion underscored the need for a new approach to the prediction of fabric
filter efficiencies, especially for woven fabrics.
A New Model for Fabric Filter Collection Efficiency
As indicated by our review of former models, most treatments of filter collection
efficiency were not for fabric filters but for other kinds of filters. The syn-
thesis attempted for fabric filters by Fraser and Foley used as its basis formulas
for collection on cylinders. Because of the overlapping caused by weaving and the
closeness of the yarns, the flow in a fabric filter is very unlike that of a cyl-
inder in a uniform flow stream, the starting point for models of filtration based
on collection by cylinders. Rather, the fabric produces a flow which is directed
through the interstices of the weave at and near the intersections of the yarns.
This much more resembles flow to an array of apertures than it does flow around
single or grouped cylinders. Because the nature of the flow has a strong bearing
on the collection efficiency, one expects a pore model to be .inherently better
than models which start from a cylinder or an array of cylinders.
Our analysis fits into a baghouse simulation model outlined in the flow chart in
Figure 4. Once the starting data are put into the simulation, the program cal-
culates flow, pressure drop, fabric loading, penetration, and emissions as a func-
tion of time for the bags in the simulated baghouse. This section discusses the
equations and analysis which went into calculating the emissions.
Collection Mechanisms. The model can calculate the fraction of incoming parti-
cles which are captured in the cake, in the fabric, and in the pinholes which
go through cake and fabric. The collection mechanisms modelled are impaction,
interception, and diffusion.
Collection by Impaction, Interception, and Diffusion in the Cake. The particu-
late material collected by the fabric filter builds up first in the vicinity of the
pores. As it builds, this deposit starts to form a cake through which a substantial
portion of the flow passes, and particles are captured by the particulate matter of
the cake. We seek to model the capture of particles by the filter cake. The
geometry is quite complex where polydisperse aerosols are involved. The uncertain-
ties are so great with regard to geometry that the available information warrants
only an approximate model. We have formed such a model on the basis of analysis of
characteristic time scales, characteristic lengths, etc.
An idealized view of the filter cake is shown in Figure 5. The cake is viewed as
a series of layers, with a total thickness L, in cm. The cake thickness is given by
357
-------
MAIN
1 • V
SUBROUTINE DATA
MODEL INITIALIZE
" 1
SUBROUTINE SUBROUTINE
REAOIM READIT
4
tf LOOP OH A SUBROUTINE
— >V TIME J PLOTIN
LOOP ON }
BAGS J
DRAG ^/*
,, $
SUBROUTINE SUBROUTINE
FABDRG CAKDRG
4
4
f END BAG "\
V LOOP J
I
f LOO
^ a«
1
C LOOf
V PARTIC
*
P ON A,,
fis y
> (TM \j
LE SIZE J"
t
< CALCULATE ^"\
PENETRATION ^S
. \
,114
SUBROUTINE SUBROUTINE SUBROUTINE SUBROUTINE
IMPINT DIFUSN CAKIMP CAKOIF
,
S' CALCULATE ^\
4
SUBROUTINE i
PINDRG ^ CAt_cu
X. WEIGHT
i
f t*
^ PARTICl
1
f E
^ BAG
i
LATE ^S^^
GAIN .>
D >
.t LOOt' _/
•JD ~\
LOOP j
^^PRWf 8 PLOT^V.
< PRESSURE, EMISSIONS^
\^ WEIGHT ^^
•\
/O) fc/" END TIME "\
VC/ 'V LOOP J
i
T END ")
t
EXECUTE
PLOT
PROGRAM
Figure 4. Baghouse simulation program flow chart.
358
-------
Figure 5. Fabric filter cake, resting on fabric.
L = W/P (1 - e)
P
(24)
In which W = particle loading, g/cm
o
p m particle material density, g/cm
P
e = void volume fraction, porosity„
By taking the number of particles and dividing it into the volume of the cake, we
can get the mean volume of cake per particle, thus the mean thickness of each layer:
L* = [ir/6(l - e)]1/3 VMD
in which VMD = particle volume mean diameter, cm.
(25)
Assuming that the number of void volumes in the cake is equal to the number of par-
ticles, we have the following void volume mean diameter, d :
1/3
VMD
d - re/(l - e)]
v L J
The mean velocity in the open portions of the cake is given by:
U = vf/e
(26)
(27)
in which vf = face velocity, cm/s.
A characteristic velocity for flow through the cake is the face velocity divided by
the porosity, U. A characteristic dimension over which the flow changes direction
is the mean diameter of a void, d .
v
The inertial parameter (Stk, the Stokes number) can be defined from these equations
as:
Stk = UT/d
(28)
359
-------
This parameter determines how large inertial impaction may be. We make the simple
20
assumption (e.g., see Fuchs ) that the impaction efficiency of a single cake
layer is approximately equal to the Stokes parameter at each void volume (up to
Stk = 1):
EIi ~ 1J (29)
The likelihood that a particle will penetrate through one void volume is just one
minus the efficiency:
Pnli = 1 - Ui/dv (30)
and the likelihood of penetrating M such volumes (i.e., through M layers) is:
Pn-j. = (1 - ^Ii)li> (31)
(This assumes that the collection efficiency of each void volume in the cake through
which the aerosol flows is independent of the earlier portions of the flow path.)
The number of such layers is approximately the thickness of the cake, L, divided by
the mean distance between particles forming the cake, L*, the mean layer thickness:
M = L/L*» (32)
Interception is the collection contribution due to solely the size of the particles
compared with the dimensions of the flow channels. Even if the particles follow
the streamlines of the flow, those with centers on streamlines which come within a
particle radius of an obstacle will strike the obstacle and be captured (it is
assumed) . We can approximate the likelihood of penetration in passing through the
^th layer as that for passage through a cylindrical channel of diameter d :
Pnc. = (1 - dp/dv)2 (33)
and from this we derive an approximate expression for the penetration of a particle
through the dust cake of thickness L by:
Pnc - (1 - dp/dv)2 . (34)
The Brownian motion of particles causes them to diffuse in a concentration gradient
so as to transport particles from higher concentration regions to lower. In the
presence of a surface to which the particles adhere, there will be a net diffusive
flux of particles to the surface due to this mechanism,, The equation governing
this diffusion for a flowing aerosol is given by:8
+ v . vn . D V 2n (35)
360
-------
in which n » number concentration of particles of diameter d
v « flow velocity
D = particle diffusivity
yn = gradient operator
""2
V = Laplacian operator
For systems of interest it is usually true that the convective term, "v"0 A^T, is
negligible. Within a void volume this equation becomes approximately:
n = D t/(0.5 d )2 (36)
for the capture due to diffusion. The time, it, is the characteristic flow length
divided by a characteristic flow:
At = d /U (37)
As before, for M (=L/L*) layers in the cake, the penetration becomes:
PnD = (1 - 4 D/U dv)L/L*. (38)
20
As Fuchs has noted, there is no simple way for combining correctly the collection
efficiency terms (or, equivalently, the penetration expressions). Commonly, the
mechanisms are treated as operating independently, in which case the penetration
becomes the product of the penetrations:
Pn - Pnx Pnc Pn^ . (39)
We have used this approach in our model.
Impaction and Interception at a Pore or Pinhole. Near the intersections of
the fabric yarns are openings which we call pores, the geometries for which are dis-
3
cussed in the Interim Report by Dennis et al. and, to an extent, in Appendix A of
our report. It has been observed that often a fraction of such pores, and/or
pores caused by other means, stay open during the filtration cycle. We denote
them "plnholes." Our model simplifies both pores and pinholes into circular
cylindrical channels in the fabric (and in the cake as well, for pinholes). The
task then becomes to calculate the collection efficiency of such an array of
channels.
21
Spurny et al. presented calculations for a somewhat similar situation in their
analysis of membrane and nuclepore filters, modeled as a highly porous array of
parallel circular cylindrical channels. It can be shown that their equations for
361
-------
impaction do not go to the correct limits for low-porosity structures, which is
our situation for fabric filter having a few percent or less free area.
22
Smith and Phillips found numerical solutions to the problem of the collection
efficiency of an aperture due to impaction and interception, but they included only
capture at the face of the material having the orifice and not along the walls of
the channel, which can be shown to be substantial in our context,
Thus, although we could use equations available in the literature for collection
due to diffusion in a channel, we did not have applicable equations for impaction
and interception of particles by the pores and pinholes and had to derive approxi-
mate expressions, discussed next,,
Figure 6 shows the geometry of the model. The pore (channel, orifice, etc.) has
a radius (or hydraulic radius) R* and a length L „ The opening is one of several
2 P
and the area irR is the total face area divided by the number of pores, distributed
approximately uniformly,, The face velocity, vf, far away from the pore is just the
flow through the pore, Q , divided by the area per pore:
v, = Q /TTR 2 (40)
i p o
The flow is simplified by dividing it into three regions (see Figure 6):
I. In this region (radial distance, r, greater than Ro) there is a
uniform face velocity toward the plane of the opening, and the
face velocity is given by:
Vf = V^o2
II. In this region (radial distances: R* <_ r < RQ) the flow is the
same as it would be in potential flow to a point sink, being
directed radially toward the center of the pore face, with the
velocity at distance r governed by:
v(r) = v*(R*2/r2) (42)
in which
v* = Q /2irR*2 (43)
P
III. In the third region (r < R*) and within the channel there is
assumed a uniform flow profile.
Of course, this set of flow profile assumptions is somewhat different from what
actually would take place. Still, far away from the pore opening the flow would
generally be uniform; there would be an intermediate range for which the flow
362
-------
Figure 6. Schematic of flow model for analysis of
impaction and interception .
363
-------
would be approximately that toward a point sink at the center of the orifice; and
converging flow at the orifice would be expected to produce a rather uniform flow
profile near the entrance before the typical parabolic flow was established for
laminar flow conditions. Thus the flow model preserves many of the features of a
more rigorous solution of the flow velocity field,
A particle with an initial velocity of v, will travel a distance L ' in a station-
20
ary medium, with L' T given by:
S JL
Lsl(1 •
where
L8I - * vf (45)
T = Cp d 2/18M. (46)
in which C is the Cunningham slip correction, p is the particle density, and d is
P P
the particle diameter. The variables L and T are the particle stopping distance
20 S
and particle relaxation time, and t is the time.
We will not give more details of the rather complex method used by us to obtain
efficiency due to impaction. Essentially it involved comparing the stopping dis-
tances (L ', including mean residence times in regions II and III) with the dis-
s
tance a particle would have to travel to strike the face of the pore or the inner
wall, keeping track of particles which were captured in one region so as not to
double-count them in capture in subsequent regions.
The other mechanism considered was diffusion to pore and pinhole walls. The formulas
for the decrease in particle concentration in traversing a tube while diffusing have
20
been presented by Fuchs. The concentration decreases approximately exponentially
with the parameter u as the argument:
u = D L /V R*2 (47)
in which the particle diffusivity is D. The approximate formula for u much less
20
than 1 is:
Pn-l-E-1- 2,56 u2/3 + 1.2 n + 0.177 u4/3 (48)
364
-------
Summary of Model Equations. The essential aspects are the collection mechanisms
of impaction, interception, and diffusion as applied to the filter fabric, the cake,
and pinholes in fabric and/or cake.
Results of Modeling of Fabric Filtration Systems
Introduction. This portion will present some of the results we have obtained
in modeling fabric filtration systems with an initial emphasis on coal-fired power
plants using fabric filtration, such as the Sunbury plant that was one of the sub-
3
jects of the Interim Report by Dennis et al. The order is as follows:
1. Q and Ap versus time for 15-bag system, using a linear drag model.
2. Collection efficiency versus particle size and time for a single
bag with linear drag.
3. Collection efficiency versus particle size and time for a single
bag with nonlinear drag.
4. Collection efficiency versus particle size and time for a three-
bag system, using nonlinear drag.
Flow Modeling. In reference 16 we presented the computer program for our model.
The drag equation used was linear:
S = SE + K2 W (49)
W = cVT (50)
where S£ = 005 in. WC/fpm = 2.5 cm WC/(cm/s) = 250(N/m2)/(cm/s)
K2 = 11 in. WC/fpm/lb/ft2
-1.1 (N/m2)/(cm/s)(g/m2)
c = 5 gr/ft3 =11.5 g/m3
t = time since cleaning, minutes.
Either V was constant ( 2 fpm = Icm/s face velocity) or it was obtained from the pres-
sure drop and the drag. Each bag was cleaned instantaneously (to the residual weight
level) every 30 minutes, one bag cleaned per 2 minutes. The results are shown in
Figures 7 and 8, where the information is plotted every minute so that we are getting an
instantaneous reading at two points in the cycle, just after cleaning and midway between
cleanings. From Figure 7, we see that after about two complete cycles (60 minutes), the
pressure drop has come into dynamic equilibrium, with a minimum near 1.4, a midpoint
2
about 1.45, and an (inferred) maximum about 1.5 in. WC (375 N/m ). For 15 bags,
the range of pressure drop values is much less than for one bag, as we expected.
365
-------
"b.oo
20.00 HO.00 60.00 80.00 100.00 120.00
TIME (MINUTES)
Figure 7. Test run No. 5: 15-bag simulation pressure versus time graph.
"0.00 20.00 iJO.OQ 60.00 80.00 100.00 120.00
TIME (MINUTES)
Figure 8. Test run No. 5: baghouse simulation individual flow rate graph.
366
-------
Because the many-bag system greatly reduces the pressure drop variations with time,
we can approximate this new system as follows: even though the total system is one
with a constant total flow, it behaves, bag-by-bag, almost as though each bag is
kept at a constant pressure drop (and has a variable flow). This is significant,
in turn, because we have available to us the general expression governing the relation-
ship between drag and time (thus flow and time) for a bag with constant pressure drop
and variable flow, which we presented as Equation 8, from Solbach.
Figure 8 shows the flow rate for bags No. 5 to No. 1 (the cleaning order was from No. 15
to No. 1). The total flow rate remained constant, but the individual flow rates declined
for these bags as they accumulated more weight than the other bags in the system, which
were being cleaned before them. When it became the turn of bag No. 5 to be cleaned, its
flow rate increased, while the other continued to drop. This process repeated itself
with the other bags, and it is interesting to note that their flow rate versus time curves
all look very much the same except for being displaced in time.
Efficiency Modeling. Figure 9 shows the results of modeling the collection efficiency
of a single bag with the program discussed above and shown in Figure 4.
en
-------
Table 3. TEST RUN NO. 13A BAGHOUSE SIMULATION INPUT DATA
Number of bags: 1 bag
Cycle time: 60 min.
Number of increments per cycle: 15 increments
Velocity, Q/A: 1 cm/s
Concentration: 1.142 x 10~ g/cm
o o
Initial drag resistance, S • 2.5 x 10 dyn/cm /cm/s
r O 9
Cake resistance coefficient, K: 1.12 x 10 dyn/cm /cm/s/g/cm
^o
Pore sphere of influence radius: 4.78 x 10~ cm
-3
Pore radius: 2.5 x 10 cm
_2
Pore length: 4.0 x 10 cm
_2
Yarn radius: 2.0 x 10 cm
3
Particle density: 1.0 g/cm
-4
Gas viscosity: 2.08 x 10 poise
Cake porosity: 0.5
Temperature: 350°K
Caked area (uncleaned fraction): 0.0
-2
Pinhole radius of influence: 4.78 x 10 cm
Pinhole radius: 5.0 x 10~3 cm
Particle volume mean diameter: 5.0 x 10 cm
2
Number of pinholes per cm : 0.31
Particle information:
Diameter (cm)
1.0 x 10~5
3.0 x 10~5
1.0 x 10~4
3.0 x 10~4
1.0 x 10~3
3.0 x 10~3
Distribution
0.00
0.01
0.12
0.32
0.38
0.17
Slip correction
2.87
1.56
1.16
1.05
1.02
1.01
Diffusion constant
7.02 x 10~6
1.27 x 10~6
2.85 x 10~7
8.51 x 10~8
2.49 x 10~8
8.21 x 10~9
Tau
7.79 x 10~8
3.82 x 10"7
3.16 x 10~6
2.58 x 10"5
2.75 x 10"4
2.46 x 10~3
chosen to be one (the density of fly ash would have been a better choice). S and
K were not used, but rather S and K0 were the same values as above. Gas viscosity
O Ji £.
was that appropriate for temperature of 350°K (= 77°C = 171°F) ; a higher temperature
would probably be more appropriate for coal-fired systems. The caked area (fraction
of filter from which the cake is not removed by cleaning) was set to zero. The
pinholes, which never close during the filtration cycle, were given the same radius
of influence as the pores, with a radius estimated as twice the size of a typical
368
-------
pore, and with a number per area corresponding to about 2 per square inch (see
Dennis et al. ) The cake dust was assumed to have a volume mean diameter of 5 ^m
and a porosity of 0.5. In Table 4 we give the approximate size distribution for
the Sunbury fly ash (Dennis et al. ) used to approximate the inlet particle size
distribution; this is log normal with a mass median aerodynamic diameter of 6[im and
a geometric standard deviation of 3.0. The values 0.1, 0.3, 1.0, 3.0, 10, 30 urn
have been used as the midpoints of equal logarithmic intervals (geometric means of
the intervals) for the size distribution. The penetration results in Figure 9 are
plotted on semilogarithmic scale. The penetrations are weighted by the mass frac-
tion values in Table 4 to give the curve marked "all particles."
Table 4. APPROXIMATE SUNBURY FLY ASH SIZE DISTRIBUTION
(log normal, d = 6, a =3)
aeSO e
Size interval
(urn)
17 -52
5.2 -17
1.7 -5.2
0.52 -1.7
0.17 -0.52
0.052-0.17
Logarithmic midpoint,
geometric mean
(um)
30
10
3.0
1.0
0.3
0.1
Aerosol
mas s in
size interval
(%)
17
38
32
12
1
« 1
The particle size with the greatest penetration (of those tested) is 0.3 (im diameter,
which is in accord with experimental results from many different aerosol filtration
studies. The initially linear slope on this scale (semilogarithmic) is consistent with
an exponential decay as the cake gets thicker (recall this is constant velocity, thus
the rate of cake build-up is constant). Once the region is reached where much of
the flow and virtually all of the penetration is that through the pinholes which
do not cake, the penetration becomes nearly constant with cake thickness. A sub-
stantial fraction of the 0.3 and 1.0 um particles penetrate the clean fabric, but
then each of these decreases toward an asymptote as the pinhole penetration pre-
dominates. The 3 um particles apparently penetrate only the pinholes and this
penetration does not change appreciably during the cycle. The 10 and 30 um par-
ticles do not even penetrate the fabric or the pinholes at all (Pn < 10 ) so that
the cake buildup does not affect them; they are caught by inertial impaction and
by interception by fabric pores and by pinholes. The mass penetration, "all
particles," starts at about 10 percent and rapidly declines to about 1 percent. The
leveling off of the curve for total mass penetration indicates that this penetration
369
-------
is almost wholly due to pinholes, and it will thus be quite sensitive to the assump-
tions made about the pinhole population. (The pressure drop versus time, which
follows the linear model, went from 2.5 cm WC (1 in. WC) to about 7 cm WC (about
2.7 in. WC), reasonable values.)
Figure 10 has penetration versus time curves for the same situation except that the
nonlinear model of drag was used. We started with an initial drag which was roughly
that of the Sunbury fabric (Sfi = 0.05 in. WC/fpm), then used the correlations developed
above to get the ratio of KZ to ^ from SQ, ^/I^ = 12.5, and to get W* = 0.004 lb/ft2
= 0.002 g/cm . The efficiency curves are very similar to those for the linear drag
model, except that the initial rapid change in pressure drop somewhat changes the dis-
tribution of the flow between the fabric plus cake and the pinholes, shifting the flow
more rapidly to the pinholes, thus producing a bit of increase in the penetration of
the 3 ^m particles initially, before their penetration becomes effectively indepen-
dent of time. Again, major features include the maximum penetration by the 0.3 jam
diameter particles, the initial exponential decrease for those particles being fil-
tered out primarily by the cake, and the important role of the pinholes as the cake
becomes thick.
10. MICRON
Figure 10. Test run No. 13b: baghouse simulation penetration versus time
graph; single bag, nonlinear drag, with pinholes.
370
-------
To demonstrate further the impact of the pinholes, we repeated the last test (non-
linear drag, etc.) but removed the pinholes, so that all the flow is through fabric
pores or cake only. Figure 11 is the penetration without pinholes. The most marked
difference is that collection efficiency for all particle sizes now continues to
decrease as the cake thickens, and this efficiency becomes better than 99 percent
by mass for the particles within 10 minutes. The 0.3 um penetration is still
greatest. The 0.1, 0.3, and 1.0 um particles all show an exponential decrease with
cake thickness now. The 3 um particles show more rapid changes in collection at the
beginning part of the cycle, where pressure drop is changing most rapidly, and this
is typical for other collector where inertial impaction is the predominant mech-
anism, as it is for these 3 um particles.
The final figure in this series is Figure 12. Here we have used three bags rather
than one, the same nonlinear drag parameters as before, and we have restored the
pinholes. First, the bags all have the same flow rate, then after one is cleaned,
only two have the same flow rate, and the third is much higher in flow. Finally,
all three have been cleaned once, and a pattern develops which will repeat itself.
Figure 12 shows the changing pattern for particle penetration versus time with par-
ticle size as a parameter. The relative order of penetration with respect to size
is the same, and we also still see the change of the penetration for the 3 um par-
ticles as pressure drop changes. Finally, the asymptote provided by the pinhole
mechanism is evident. The three-bag system with nonlinear drag produces nearly the
same total mass penetration as the one-bag system after 15 minutes, but the clean-
ing process then returns one of the bags to low efficiency and high flow rate, pro-
ducing a major increase in penetration for the total system.
Calculator Model for Drag. In some cases, a quick and approximate estimation of
an average air-to-cloth ratio of a baghouse unit is necessary or satisfactory.
Therefore, we tried to derive a simple equation for calculating an average air-to-
cloth ratio. For our derivation we used an expression for the average face velocity
(air-to-cloth ratio), V, which was derived by Solbach and independently also by
/:
Dennis and Wilder. This expression is based on the linear drag model. The average
face velocity is given by Equation (8):
Co ,ct ,Ap + S_ - S_ (51)
*• \ C\ Ei Ji
371
-------
MICRON
MICRON
MICRON
MICRON
MICRON
O.OQ 15.00 30.00 45.00
TIME (MIMJTES)
60.
O
Z O
O —i_
CC
cc.
a- -
o
b
-------
where K2 = dust resistivity, in. WC-min-ft/lb
o
c = dust concentration, Ib/ft
t = total filtration time, min
Ap = pressure drop across the baghouse, in. WC
SE = effective drag, in. WC/fptn.
If we express the effective drag, S_, as a function of the clean fabric drag S
16 E c
by:
SE = 7*4 So
0.672
and K2 as a function of particle mass median diameter, d , by:
16
K0 = 142.4 d
-1.06
the Equation (51) becomes:
- 7.4 S
0.672
V =
54.76 S
1.344
+ 284.8 d
-1.06
AP ct
142.4 d "1'°6 ct
(52)
(53)
(54)
Table 5. PARAMETERS FOR FIVE BOILERS USING FABRIC FILTRATION
Pl«nt
P'T,n*ylv-inia Power and Li^ht Co.,
Surbury Station, Pa.
CoIo^acJo CT£ Electric rssoc.,
Sucla Power Station, Colo.
Sorg Pajier Co. , Middletown, Ohio
Crisp Co'jrf*iry Power Comsi&sion,
Cordate, Georgia
Pennsylvania Fover and Light Co.,
Hottwood Station, P*.
Clean cloth
Frasier
permeability,
fom & 0.5" R20
54.3
86.2
55.0
55.
45 - 65
(50)
Dust
concentration,
gr/ft3
2.
1.93
3.2
2.6
7.7
Total filtration
cycle time, min
33
29.4
30
60
33
Pressure
drop,
" H20
2.5
4.2
4-5
4.
6.5
Particle
mass median
diameter, ^m
8
8
12
12
10*
Operating
velocity,
fpm
1.92
2.79
2.07
2.2
2.26
Predicted
velocity,
V, fpm
3.90
6.50
6.10
4.69
4.66
Aisuaed valuefl.
Then, using industrial data for Ap, d , S , c and t , shown in Table 5, we calculated
the face velocity (air-to-cloth ratio) for each baghouse. These values are shown in
Table 5 along with the operating industrial values of the face velocity. (Sources of
the data are listed elsewhere. ) The ratios of operating to predicted velocities are
very close. The average of all values of that ratio was 2.29 so that we have a cor-
rected estimate, V', of the average face velocity:
V1 = V/2.29
(55)
The differences between corrected predicted and operating values were within + 15 per-
cent. More data would be needed to improve the reliability of our last equation,
obviously.
373
-------
Final Comments
What is presented here seems the most detailed model yet of a fabric filtration
system, and it has sufficient flexibility to have added to it a number of refine-
ments as our information warrants.
The areas which seem most in need of further study are cleaning and seepage,, Re-
garding cleaning, we would lime to know the amount and spatial distribution of the
residual dust as a function of cleaning type, intensity, duration, and frequency for
several dust/fabric combinations, especially those of applicability to controlling
fossil-fuel boiler emissions. Regarding seepage, we would like to know what variables
affect the rate at which it occurs and what the relationships are. Of course, the
application of fabric filtration does not require such information, but the optimi-
zation of fabric filtration does,,
Accurate methods for prediction of fabric filter drag and collection efficiency would
greatly aid the design of optimal fabric filter systems,. By viewing the fabric as
an array of apertures rather than an array of cylinders, we have been able to derive
expressions describing the nonlinear dependence of filter drag on fabric dust loading
and the collection efficiency of the fabric and the dust cake as functions of par-
ticle size,, Much work remains to be done, especially regarding cleaning of fabrics,
but we believe that this new approach can provide a useful framework for future de-
velopment as well as yielding useful insights concerning the factors which determine
drag and collection efficiency,, As the model is improved and as more data from
operating systems are incorporated, it is hoped we will have a method which will
facilitate the application of fabric filtration to the problem of obtaining power
economically with minimal environmental impact„
Acknowled gment s
We appreciate the help we received from our colleagues at GGA, especially the com-
puter programming done by Roger Stern and the comments and information given us by
Richard Dennis. This program has been funded through EPA Contract No. 68-02-1438.
The project officer is Dr. James H. Turner, Industrial Environmental Research
Laboratory - Research Triangle Park, N.C.
References
(1) Jorgensen, R0 (1970), Fan Engineering, 7th ed, Buffalo, N. ¥„,
Buffalo Forge Co0
(2) Snyder, C. A, (1955), Ind0 Eng0 Chem, Process Des, Dev. 47, 960.
374
-------
(3) Dennis, R., R. W. Cass, and J. E. Langley, Development of Data Base
and Fabric Filtration Models for Design of Participate Control Systems.
Environmental Protection Agency, Research Triangle Park, N.C. Contract
No. 68-02-1438, Interim Report for Task 5.1, Dec 1975. 119 p.
(4) Robinson, J. W., R. E. Harrington, and P. W. Spaite (1967), Atmoa. Environ.
1, 499-508,
(5) Solbach, W. (1969), Staub (English) 29(D. 28-33.
(6) Dennis, R. and J. Wilder. Fabric Filter Cleaning Studies - Appendix J.
GCA/Technology Division, Bedford, Massachusetts. For: U. S. Environ-
mental Protection Agency. Publication No. EPA-650/2-75-009 (Environ-
mental Protection Techntiogy Research Report Series), Jan 1975.
(7) Williams, C. E., T. Hatch, and L. Greenberg (1940), Heat/Piping/Air Cond.
12, 259-263.
(8) Bird, R. B., W. E, Stewart, and E. N. Lightfoot (1962), Transport
Phenomena, New York, N. Y., J. Wiley and Sons, Inc.
(9) Fraser, M. D. and G. J. Foley. A Predictive Performance Model for
Fabric Filter Systems: I - Intermittently Cleaned Single - Compartment
Systems. Presented at 67th Annual Meeting of APCA, Denver, 1974.
(10) Billings, C. E. and J. Wilder, Handbook of Fabric Filter Technology.
Vol. I. Fabric Filter Systems Study. National Technical Information
Service. Publication No. PB 200-648, 1970.
(11) Davis, W. T., K, E. Noll, and P. J. LaRosa, A Predictive Performance
Model for Fabric Filtration Based on Pilot Plant Studies. Proceedings
of the User and Fabric Filtration Equipment II, Specialty Conference
APCA, Niagara Falls, N. Y,, 1975.
(12) Draemel, D. C., Relationship Between Fabric Structure and Filtration
Performance in Dust Filtration. Environmental Protection Technology
Series. Report EPA-R2-73-288, 1973.
(13) Spaite, P. W. and G. W. Walsh (1963), Amer. Ind. Hyg. Assoc. J. 24.
357-365.
(14) Durham, J. R., Filtration Characteristics of Fabric Filter Media NAPCA,
PHS, USDHEW, Interim Report, 1969.
(15) Dennis, R», Reported Experimental Measurements, GCA Tests, 1975.
Personal Communication.
(16) Cooper, D. W., V. Hampl, and R. Stern, Development of Fabric Filtration Model,
Interim Report for Task 5.3, January 1976.
(17) Borgwardt, R. H., R. E. Harrington, and P. W. Spaite (1968), J. Air Pollu.
Control Assoc. 18(6), 387-390.
(18) Davies, C. N. (1952), Proc. Inst. Mech. Eng. 1B(5), 185-213.
(19) Leith, D. and M. W. First. Particle Collection by Pulse-Jet Fabric
Filter. Presented at 68th Annual APCA Meeting. Boston. 1975,
375
-------
(20) Fuchs, N. A. (1964), Mechanics of Aerosols, New York, N. Y., Pergamon
Press.
(21) Spurny, K. R., J. R. Lodge, Jr., E. R. Frank, and D. C. Sheesley (1969),
Environ. Sci. and Technol. 3, 453-469.
(22) Smith, T. N. and C. R. Phillips (1975), Environ. Sci. and Technol. 9,
564-568.
376
-------
PERFORMANCE AND ENGINEERING EVALUATION
OF THE NUCLA BAGHOUSE
D. S. Ensor
and
)
R. Hooper
Meteorology Research, Inc.
Altadena, California
R. W. Scheck
Stearns-Roger, Inc.
Denver, Colorado
R. C. Carr
Electric Power Research Institute
Palo Alto, California
377
-------
PERFORMANCE AND ENGINEERING
EVALUATION OF THE
NUCLA BAGHOUSE
D.S. Ensor
R. Hooper
Meteorology Research, Inc.
Altadena, California 91001
R.W. Scheck
Stearns-Roger, Inc.
Denver, Colorado 80217
R.C. Carr
Electric Power Research Institute
Palo Alto, California 94304
For presentation at the EPA-EPRI Symposium
on Particulate Control in Energy Processes,
May 11-13, 1976, San Francisco, California.
378
-------
Acknowledgments
The Nucla evaluation was funded by Electrical Power Research Institute
under contract number RP 534-1. The assistance of Mr. Bill Carlson and
Mr. Don Dove and other staff members of the Colorado Ute Electric Associa-
tion is greatly appreciated. A special thanks to Mr. Gil Sem of Thermosys-
tems, Inc. , for assistance in applying the Electrical Aerosol Size Analyzer
to source testing.
Meteorology Research, Inc., internal research funds were used to write
this paper.
379
-------
I. Introduction
A. Objective
The objective was the evaluation of the baghouses at the Nucla Station
of the Colorado Ute Electric Association with a field performance test and
engineering study. This paper is a summary of the study. For more details,
the final report by Ensor et al. [1976] should be consulted.
B. Site Description
The Nucla Station is located five miles southeast of Nucla, Colorado, a
small town one hundred miles southeast of Grand Junction. The coal, pur-
chased from Peabody, Inc. , is mined about ten miles away and is trucked to
the plant. The coal has a heating value of about 12, 000 Btu/lb, 14-20 percent
ash, 45 percent fixed carbon and 0.5 to 0.7 percent sulfur.
The station has three generators, each with a nameplate rating of
12,650 kw. The three Springfield spreader stoker boilers are rated at
139, 000 Ib/hr of steam at 825T and 625 psi. Gas from each of the boilers
passes through an economizer, air preheater, baghouse, induced draft fan,
and out the stack. The gas streams of each of the three boilers are identical
and separate. A diagram of the system is shown in Figure 1.
Each boiler has a Wheelabrator-Frye Size 814, Model 264, Series 8,
six Module Dustube Dust Collector-. The design parameters are summarized
in Table 1. Each Collector has six compartments which are independent and
can be cleaned or maintained without interference to the other operating com-
partments. The bags are suspended from the shaker mechanism by a hook
and a spring to maintain the desired tension and are connected to the "cell
plate" at the bottom by a snapring.
The compartments are cleaned in sequence when the pressure drop
across the baghouse reaches about 4. 5 in HgO. The cycle consists of a 1
minute delay, 15 seconds reverse air, 1 minute delay, 10 second shake, 45
second delay, another 15 seconds of reverse air, and 45 second delay for
each compartment. The reverse air is obtained from the discharge side of
the baghouse by a 30 hp centrifugal blower. The reverse air flow is about
5500 cfm for an air-to-cloth ratio of 1. 1 acfm/sq ft. The frequency of the
shakers is 4 cycles per second.
II. Test Methods
A. Size Distribution
Cascade Impactors. The impactor used in this study was designed at
Meteorology Research, Inc. , (MRI) to facilitate sampling of particulate mat-
ter in stacks. The design is based on a simple annular arrangement of jet
and collectors reported by Cohen and Montan [ 1967]. The assembly has
been constructed using quick disconnect rings to increase the flexibility of
380
-------
U!
00
Air Preheater
Gas Sampling
Point
Ash
Clinkers
Forced Draft
Fan
Ash
YYV
Baghouac
(Six Compartments)
Ash Induced
Draft
Fan
Outlet Test
Point
•Stack
Figure 1. Flow Diagram of Major Gas Streams, one of three identical systems
-------
Table 1. Baghouse Design Parameters
Gas Flow Rate
C ompartment s
Bags /Compartment
Bag Manufacturer
Model
Diameter
Length
Active Area/Bag
Fabric
Finish
Permeability
Weight
Air to Cloth Ratio,
6 Compartments
Air to Cloth Ratio,
5 Compartments
{cleaning or maintenance
on one compartment)
Pressure Drop (normal)
Pressure Drop (cleaning)
86,240 acfm
6
112
W. W. Criswell Division
Wheelabrator-Frye
No. 445-04 "Dustube"
8 in.
22 ft
46.0 ft3
Fiberglass
66 x 30 thread count, 3x1 twill
Silic one /G r aphite
45 to 65 acfm/ft2 @ 1/2 in.H3O
10.5 Oz/yds
2. 8 acfm
ft-2
3.35 acfm
ft3
4. 5 in.H2O
up to 6 in.HsO
382
-------
application. "O" rings are used under direct compression for a positive gas
seal of the plates containing the jets.
The particulate matter was collected on lightweight stainless steel col-
lection discs coated with Apiezon. The detailed procedures are discussed by
Ensor et al. [ 1976]. The material lost to the walls of the impactor was
brushed onto the appropriate substrate surface. A Cahn 4100 balance was
used to weigh the discs to 0.01 mg.
Submicron Particles. The submicron particles were measured with an
extractive sampling system consisting of a
• Impactor in-stack precutter
• Three-stage dilution system outside of the stack
• Diffusion battery/condensation nuclei counter
(Gardner Associates, Inc. )
• Model 3030 Electrical Aerosol Size Analyzer
(EASA) (Thermo-Systems, Inc.).
The system is described in detail by Ensor et al. [1976] .
Submicron particle testing was a separate and complex task at the inlet
and outlet at ports adjacent to the ports used for the cascade impactor tests.
The diffusion battery was used at a given location for an entire day. The
EASA was moved between the inlet and outlet sampling port as many times
daily as feasible.
Dilution and drying of the inlet sample were necessary; dilution ratios
ranged from 2OX to 100X. Low particle concentrations at the baghouse outlet
prevented sample dilution, which meant the dilution system was required only
at the inlet sampling site. However, without dilution at the outlet, condensa-
tion in the sampling train became a major concern. To reduce the moisture
in the sample, two in-line diffusional dryers were used in the sampling train.
An impactor precutter was used at both the inlet and outlet sampling
locations. The impactor offered the advantage of a well-defined cut-off at
2. 7 microns.
B. Gas Sampling
Gas samples were taken at the inlet and outlet of the baghouse. The
water was condensed at the stack and the dry gas piped to the trailer with a
diaphragm pump.
A Theta Sensors, Inc., three-gas analyzer (O2, SO2, and NO) and a
Horiba CO analyzer were used in the test. The inlet and outlet points were
alternately sampled every 1-1/2 hours. The instruments were zeroed and
calibrated with span gas twice daily.
383
-------
C. Opacity
The Plant Process Visiometer (PPV) is a light scattering instrument
developed at MRI for real-time monitoring of stack opacity. Flue gas is
isokinetically withdrawn from the stack with a stainless steel probe and
transported into the measurement chamber. The aerosol particles in the
chamber were illuminated by a flash lamp with an opal glass filter. The
scattering light was detected by a photomultiplier tube at approximately right
angles to the flash lamp. The optics have been designed so that the output of
the photomultiplier tube is proportional to the extinction coefficient due to
scattered light. The instrument is a physical analog of the following equation:
b = 2nf1Tp(9) sin 9d9 (1)
scat Jo
where
b = the scattering coefficient due to scattered light
scat &
(3 (0) = volume scattering function
Q = scattering angle
If there is no light absorption, the scattering coefficient is identical to the
extinction coefficient. The extinction coefficient is related to plume opacity
with the Bouguer Law.
Opacity (percent) = [~1 - exp (-b ,L)1 100 (2)
l» €JXt J
where
b , = extinction coefficient, m
ext
L = stack diameter, m
The internal calibrator is an opal glass lens of known scattering coefficient.
The lens was mechanically placed in the view of the detector for calibration
and was retracted into a sealed chamber between calibrations. The PPV was
described in detail by Ensor et al. [1974].
The use of the PPV allowed the measurement of opacity at both inlet
and outlet of the baghouse. The large difference between inlet and outlet
particulate concentration required ranges of 0 to 0.2 m"1 and 0 to 5 x 10"3 m"1,
respectively. The sensitivity of the PPV allowed the measurement of opacity
variation of the invisible outlet plume.
D. Test Locations
The inlet sampling location for manual testing was a 6 ft. x 4. 5 ft. rec-
tangular duct. Six 3-inch ports offered access to the duct. This location was
poor because it was only about 1.4 diameters downstream of a bend, and 1.8
diameters upstream of the baghouse entrance.
384
-------
The outlet sampling location was at the 5.5 ft. diameter stack. The
ports were about 8.4 diameters downstream of the ID fan and about 8 diam-
eters upstream of the stack discharge.
III. Results
A. Overall Collection Efficiency
The primary determination of overall collection efficiency was from the
catch in the cascade impactors. Four inlet impactor tests in different ports
and two outlet tests were performed each day. The Phase I tests were con-
ducted from September 12 to 26, 1975, and the Phase II tests were conducted
from November 2 to 15, 1975. No problems with broken bags were experi-
enced during the tests. Only the particulate results will be summarized here.
The details of the test results are reported by Ensor et al. [l976j. The pri-
mary manipulated parameter was the generator load of the power plant. The
effect of load (air to cloth ratio) is shown in Figure 2. The particle penetra-
tion increased rapidly as the air to cloth ratio increased. With increasing
load, the baghouse cleaning cycle increased in frequency, and the pressure
drop also increased. As indicated at 6 Mw, the baghouse operated during a
full day test period without requiring cleaning, and the pressure drop was
nearly constant at 3 in tig O. As the load increased, the cleaning became
more frequent. (The average air to cloth ratio is somewhat larger than in-
dicated because of the frequent cleaning. ) The large variation in the 12 Mw
results is believed to be related to bag cleaning and to the condition of the bag
filter cake.
In Table 2, the overall results of the study are summarized. Of par-
ticular interest is estimation of the emission rate for the facility. The unit
is well within the New Mexico particulate standard of 0. 05 lb/10s Btu total
and 0.02 lb/106 Btu for particles less than 2 y, m.
The stack was clear during both test periods. The average opacity
measured with the PPV was well below the visual threshold of about 2%.
B. Fractional Penetration Results
The major objective of the field test program was to obtain the particle
size dependent penetration through the baghouse at various boiler loads. The
results are presented, in detail by Ensor et al. [1976],
The penetration for an 11 Mw load is shown in Figure 3. The penetration
for particle sizes from 0. 01 to 1.0 ^m were obtained using an Electrical
Aerosol Size Analyzer (EASA), and the penetration for particle sizes from
0.5 to 10 i^m were obtained with cascade impactors. The testing schedules
and process variation are illustrated in Figure 4 with strip charts of the
Plant Process Visiometers and CO analyzer and bar charts of the manual
tests. The submicron tests were conducted between bag cleaning cycles
while the outlet impactor tests included at least three cleaning cycles. The
difference between the EASA penetration curve and impactor penetration
curve may be due to the increased emissions during the cleaning cycle.
385
-------
Pressure Drop
Symbol Load Cleaning Between Cleaning
MW Frequency in H,O
14, 2 T99-
!•
O
X
(n
t-j
o
4-J
u
<"rt
0.
s
X
'^
•o
(0
3
O
GO
rti
ffl
X
bo
3
O
X
c
o
-------
Table 2. Summary of Overall Efficiency as Determined by Cascade Impactors
Load
Mw
6
11
12
Air to Cloth
Ratio
ft^/min/ft2
1.87 ±0.09
2.47 + 0.07
2.74 i_0. 17
Fractional
Penetration
(1 Std Deviation)
(2.07^.85)xlO~4
(2.74+11.38)xlO~4
(8.52^4. 13) x!0"4
Number
of Days
3
9
7
Efficiency
Percent
99.98
99.97
99.92
Outlet
Concentration
gr/fl*'1"
0.00021
0.00087
0.0011
Outlet
Opacity
Percent
~0.02
~0.03
0.6
Total(a)
Emission
Ib/l^Btu
0.007
0.02
0.04
^ . . (a)
Emission
Less than
2 ^m
lb/108 Btu
0.003
0.004
0.006
CX>
—I
(a)-- Preliminary Results
(b) -- Dry Standard, 760 mm Hg 21. 1°C
-------
Sulfuric Acid
Nuclei
10
-2
3
O
•a
frf
00
o
o
n)
10
-3
10
10
0.01
November 12, 1975
11 MW Load
Hourly Cleaning Frequency
Pressure Drop 3-4.5 in. H2O
• Electrical Aerosol Size
Analyzer
Q Cascade Impactor
One Standard Deviation Limits
Particle Density =2.0
0.10 1.00
Particle Diameter, microns
10.0
Figure 3. Fractional Penetration through Baghouse,
388
-------
C'.imorator
Output
Megawatts
Submic ron
ParticU
Tasting
11.0 II.0
11.0
11. a
10.7
10.6
OUTLET
INLET INLET OUTLET
I—I M I I
10. a
INLET
ii. i
C.L,-Inlet Port No.
Concentration
tng/tn I 10
C. I.-Outlet
Concentration
mg/m
ZOdo)"2
-, 16 do)'z
6,
I I2do)-z
01
e
8(10)'
4
i
4.7
&
I
14.7
1.25
1 4
I I
1.6 2.1
I. 11
PLANT PROCESS VISIOMETER
INLET
Optic3 Cleaned
l.84(io)"
?Q92(io)-«
Bag Cleaning
Malfonetioa
PLANT PROCESS V^IOMSTER
OUTLET
f
Baghou** Shaka Cycle
O
u
500
400
300
200
100
0
1200 1300 1400 1500 1600 1700 1800 1900 2000 2100
Local time, hrs
Figure 4. Process Variation During November 12, 1975
389
-------
The flat penetration of particles greater than 1.5 microns in diameter
is illustrative of particle "seepage" through the bags. This "seepage" occurs
after the baghouse is cleaned and is a result of particles sifting through the
newly cleaned bag until a cake is again formed to aid the filtration.
The increased penetration for the 0. 01 ^m particle is believed to be due
to the formation of sulfuric acid nuclei. The stack temperature was at 210°F
which is below the acid dew point. Analysis of cascade impactor final filters
indicated a large enrichment of sulfur. The location of the acid mist forma-
tion is suspected to be in the uninsulated steel stack or in the outlet duct work
in the baghouse. The sulfuric acid nuclei are of very little significance from
a mass emission standpoint.
The penetration results, obtained with cascade impactors by Bradway
and Cass [1975] the previous year for the same baghouse, were about 10
times greater than obtained in the current study. It is suspected that their
data was strongly influenced by bag leakage.
The particle diameter fractional penetration for the half load of 6 Mw
is shown in Figure 5. During the test day, the bags were not cleaned. The
good match of the EASA and cascade impactor penetrations was due to the
lack of bag cleaning cycles.
IV. Engineering Analysis
The objectives of the engineering analysis were to assemble informa-
tion for both capital and operating costs, determine reliability and identify
any major problems. This information is specific for the Nucla installation
only, and careful analysis would be required for extrapolation to other sites
with different boilers and coal. The analysis was performed considering
all three baghouses as a single unit. For detailed cost breakdown, the final
report for the Nucla study by Ensor et al. [ 1976] should be consulted.
A. Capital Costs
The capital costs were estimated from the records of Colorado Ute,
Jelco, Inc. (constructor), and Stearns-Roger (engineer). The Nucla plant
was retrofitted with baghouses and other additional equipment, thus some of
the costs may be unique to this site. A summary of the estimate is shown
in Table 3. When evaluating Table 3, the following aspects of the installation
should be kept in mind:
1. The cost is for a total of three units. A single larger baghouse
of equivalent capacity might be installed for less.
2. The generating units are relatively small, 12.65 megawatts
each. For larger units the installed cost per kilowatt of
generator capacity would be reduced.
3. Booster fans were not required at Nucla. This resulted in a
considerable cost saving.
390
-------
10
-2
4)
M
§
•a
rt
CO
•a
<§
>4
43
It
I*
£
ID
'3
Sulfaric Acid
Nuclei
10
10
-5
November 14,1975
6 MW Load
No Cleaning Cycles
Pressure Drop 3.0 in. H^O
• Electrical Aerosol Size
Analyzer
D Cascade Impactor
One Standard Deviation Limits
Particle Density =2.0 g/cm.3
0.01
0.10
1.00
10.0
Particle Diameter, microns
Figure 5. Fractional Penetration through Baghouse,
391
-------
Table 3, Summary of Capital Cost
Equipment and Installation %
Baghouse and General $1, 740,000 67
Ash Conveyor System 250,000 9
Retrofit Items 210, OOQ __8
TOTAL FIELD COSTa $2,200,000 84
Indirect Owner Costs 120,000 5
Engineering and Fee 300, OOP _U
1973/1974 INSTALLED
SYSTEM COST $2,620,000 100
Estimated Escalation to 1976 680,000 25
1976 INSTALLED SYSTEM
COST $3,300,000
Unit Factors (1976) - $87/kw
. $13/acfm
- $36/sq. ft. filter (gross)
a
Includes material, labor, supervision, field overhead and constructor's fees,
392
-------
4. The cost estimate indicated for the "Retrofit Items" is prob-
ably low. It does not include the many intangible costs in-
curred as a result of the retrofit. These costs are reported
in the other two categories and cannot be separated because
they are very difficult to define. Thus, the fraction of cost
attributable to retrofit is probably low.
5. Nucla is a small, remote mountain community. To a large
extent, craft labor had to be brought in from Montrose.
Nucla has little industry and is not served by rail. These
factors undoubtedly increased field costs.
The cost for 1976 was estimated using the projected figures of the Marshall
and Swift Equipment Cost Index and the Chemical Engineering Plant Cost
Index. The escalation is calculated assuming that the baghouses were in-
stalled in the first half of 1976, two years after the actual date.
B. Operating Costs
The operating cost was defined as any additional costs incurred by the
utility attributable to the operation of the baghouse. The costs were esti-
mated from Colorado Ute records as well as estimates by plant personnel.
The operating costs expected for 1976 are summarized in Table 4.
C. Maintenance
The plant maintenance records were reviewed and summarized in Table
5. The table shows the effort spent for each maintenance category and its
trend over the past two years of operation. The figures given in the table
are maintenance manhours over the number of occurrences. The four times
periods are roughly six months long but have been adjusted somewhat to dis-
tribute the 32, 600 hours of baghouse operation evenly between the periods.
The table shows that during the first three quarters of operation, bag replace-
ment has been the highest maintenance item. In the last quarter, however,
bag replacement was minimal and problems with the control system, the
dampers and the damper actuators occurred more often.
The trend is also shown in Figure 6, a graph of maintenance manhours
required per 1000 hours of baghouse operation. The two lines presented are
simply the averages for the previous three or twelve month period. Aver-
aging in this manner tends to smooth, somewhat, the erratic nature of the
maintenance requirements. The three month curve is much more responsive
and shows a significant drop in March, April and May of 1975. The reduc-
tion corresponds with modifications to the cell plate to extend the bag life.
During the initial months of operation, it was discovered that severe bag
erosion at the inlet of the bags resulted in premature bag failure. During a
six-month period starting in September 1974, gas straighteners called
"thimbles" were installed at the inlet of the bags. The "thimbles" resulted
in a major decline in maintenance.
393
-------
Table 4. Nucla Fabric Filter System Operating Cost Estimate
Direct Costs $/year % mills/kwh
Operation Labor**) $(9,500) (3.3) (0.05)
Maintenance Labor 2,500 0.9 0.01
Maintenance Material 8,500 3.0 0.05
Utilities 31,000 10.8 0.16
Ash Handling 11. OOP 3.8 0.06
Subtotal, $53,000 18.5 0.28
Directs
Indirect Costs
Depreciation 127,000
Interest 81,000
Insurance 3,000
Taxes 23,000
Subtotal, $234,000 81.5 1.25
Indirects
TOTAL $287,000 1.53
Not added since no new costs were incurred.
Based on 188 million kwh/year or 55 percent capacity.
394
-------
Table 5. Baghouse Maintenance Summary
Period
Maintenance
Category
Bag Replacement
Control System
Dampers & Actuators
Reverse Air Fans
Pressure Taps
Hopper Heaters
Miscellaneous
Subtotal
Routine
Total
Dec '73-
July 74
106/24
67/15
19/6
40/7
2/1
1/1
12/4
247/58
6/2
253/60
Aug '74-
Jan 75
99/19
66/11
35/9
80/10
23/6
16/3
7/2
326/60
6/2
332/62
Feb '75-
July 75
46/7
22/4
20/6
10/2
4/1
14/4
0/0
116/24
6/2
122/26
Aug '75-
Dec 75
13/4
42/10
26/7
2/1
2/1
0/0
9/1
94/24
6/2
100/26
Total
264/54
197/40
100/28
132/20
31/9
31/8
28/7
783/166
24/8
807/174
Units: manhours/occurrences
The four periods have the same amount of baghouae operating time.
-------
50
o
OJ
a
o
O
.e
(A
O
60
03
J*
O
O
O
11
O.
U)
1-1
3
O
1C
30
20
10
V Average of
\ Previous
12 Months
Average of
Previous
3 Months
Installation of
Thimbles——*
3
1174
i r
Ml}
Figure 6. Maintenance Labor Requirements
-------
The curve shows a general leveling of maintenance at about 8-man hours
per 1000 baghouse operating hours starting in June. The twelve month curve
also shows the definite trend toward reduced maintenance requirements.
This is a presentation of maintenance requirements over the first two
years of operation. The system is intended to last twenty years, and equip-
ment that has performed well thus far could cause problems in the future.
Projecting maintenance requirements should, therefore, be done with care.
The maintenance requirements for bag replacement have thus far aver-
aged 12 maintenance man-hours to replace 16 bags for each million bag-hours
of operation during the first two years of operation. A total of 18 percent of
the 2, 016 bags were replaced (32, 577 baghouse hours of operation). Most of
the bags were replaced before thimble installation.
Equipment requiring regular maintenance are:
the control system
dampers and actuators
reverse air for drives
plugged pressure taps
hopper heating system
freezing of compressed air piping
E. Reliability
In addition to good collector design and system design, the reliability
of an abatement system is a function of the operating demands and mainten-
ance philosophy of the utility. The units at Nucla are not used for peaking
but rather are operated at near full load. This constancy of operation, com-
bined with the attitude of maintenance personnel toward correcting failures
as quickly as possible, is to a large extent responsible for the high reliability
attained thus far.
Reliability can be expressed in many ways because its definition changes
with one's point of view. From the plant superintendent's viewpoint, the re-
liability is lowered when a situation occurs in which the baghouse either
obstructs, impedes or in any -way influences the on-stream-time of the boiler/
turbine. As of December 1975, no such situation has been encountered at
Nucla; the station's on-stream factor has been limited not by the baghouse,
but by equipment upstream of the baghouse. Thus, the reliability from this
point of view is 100%.
A second definition addresses the question of how dependable the units
are in producing a clear stack effluent:
(total baghouse operating hours) - (non-clear hours)
Reliability = (total baghouse operating hours)
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Thus far, a bag change has been required in an operating baghouse 50 times.
This is used as an estimate of the number of occurrences of a non-clear
stack. The duration of each occurrence is the time between when a bag
breaks and when its compartment is isolated. When a plurne from the stack
is noticed, the compartment is isolated within a few minutes. However, con
sidering that bags break during darkness when the stack appearance cannot
be observed, the average duration for both day and night is about four hours.
Thus, from the standpoint of gas cleaning or stack appearance:
Reliability . 32' = 99.4%
It should be mentioned that the high value of this figure is a direct result of
the attitude of the Nucla plant personnel toward correcting bag failures as
quickly as possible.
Another definition considers the reliability of the six individual com-
partments in each baghouse. Since startup of the system, all 18 of the com
partments have been in service except for 387 hours. This includes approx
imately 134 hours required for installing the thimbles on the cell plates.
From a compartment standpoint:
This may be interpreted to mean that all compartments were in service 99.8%
of the time. The major cause of compartment outage has been bag replacement.
V. Summary and Conclusions
The baghouses at the Colorado Ute Electric Association were evaluated
with both a field study and engineering analysis. The emissions were within
the New Mexico standard of 0. 05 lb/10s Btu total and 0. 02 lb/106 of particulate
matter less than 2 microns. The stack emissions were not visible during the
test. Measurement of opacity with an MRI PPV indicated opacities of less
than 1 percent.
The penetration of particulate matter through the baghouse was a sensi-
tive function of the air to cloth ratio. The cleaning part of the operational
cycle contributes most of the emissions. The baghouse was an efficient col-
lector of submicron particles with less than a factor of ten variation in pene-
tration as a function of particle size.
The engineering analysis indicated the baghouse had not interferred or
limited the operation of the boiler since installation. The maintenance and
bag replacement in the unit has been greatly reduced with the installation of
"thimble" at the bag entrance.
398
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References
1. R. M. Bradway and R. W. Cass, "Fractional efficiency of a utility boiler
baghouse; Nucla Generating Plant," EPA-60012-75-013-a (1975).
2. J. J. Cohen and D. M. Montan, "Theoretical considerations, design,
and evaluation of a cascade impactor, " Am. Ind. Hyg. Assn., 28:95
(1967).
3. D. S. Ensor, L. D. Bevan and G. MarkowsM, "Application of nephel-
ometry to the monitoring of air pollution sources, " 67th Annual
Meeting of the Air Pollution Control Assoc., Denver, Colorado, Paper
No. 74-110 (1974).
4. D. S. Ensor, R. G. Hooper and R. W. Scheck, "Determination of the
fractional efficiency, opacity characteristics, and engineering and
economic aspects of a fabric filter operating on a utility boiler, " MRI
76 FR-1411 (1976).
399
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OPERATING EXPERIENCE AND PERFORMANCE AT
THE SUNDBURY BAGHOUSE
Harry Spagnola
Pennsylvania Power & Light Company
Sundbury, Pennsylvania
James H. Turner
Environmental Protection Agency
Industrial Environmental Research Laboratory
Research Triangle Park, North Carolina
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ABSTRACT
OPERATING EXPERIENCE AND PERFORMANCE AT THE SUNBURY BAGK)USE
Penna. Power & Light Co.'s Sunbury Plant is located on the Susquehanna river
in Snyder Co., Shamokln Dam, Pa. Four of the boilers at the plant are equipped
with baghouses that were placed in service in the spring of 1973- The k boilers
are identical. Each is rated at Ul5,000 Ibs/hour steam flow and has a total
gas flow of 220,000 acfra. The boilers use a combination of fuels comprised of
a mix of 75$ anthracite coal and 25$ petroleum coke. The baghouses on each boiler
are preceded by mechanical collectors, which have a collection efficiency of 70$.
Each of the baghouses is divided into fourteen compartments with 90> thirty feet
long by one foot in diameter bags per compartment. The air-to-cloth ratio is 2 to 1.
The bags are cleaned by the reverse air flow method. The initial installation cost,
including a new fly ash removal system, was $ 5,500,000. Maintenance and operating
cost, excluding complete bag replacement, has averaged $ 30,000. a year.
A total of only 37 bags have been replaced due to failure. Bags in 3 of the k
baghouses were replaced as a precautionary measure after approximately 2 years of
service, at a cost of $ 60,000. per baghouse. The bags in the Uth baghouse are
still in service after 3 years of operation. The draft loss across the bags has
averaged about 3.0" VWC. The draft loss across the 3 year old bags is presently
3.5" VWC.
A test to determine the removal efficiencies at various particle sizes was
performed on 1 of the baghouses in early 1975- Mass efficiencies and efficiencies
by particle size were found using mass trains, impactors and diffusional sizing
equipment. Some difficulties were met with impactor substrates and with the
diffusional apparatus and are described in the paper. Mass efficiencies were 99-88$
and 99-93$ respectively for new bags and for bags with about 2 years of prior service,
Average mass efficiency in the particle size range 1 to 10 m was not lower than 99-5$
and reach a peak of about 99-9$ at ^-M m. Limited diffusion equipment results
indicated an efficiency on the order of 93 to 99$ for all particles in the range of
about 0.003 to O.U,^( m (by number), and that no particles smaller than about 0.015/*\m
were present in the baghouse outlet.
Harry F. Spagnola
Penna. Power & Light Co.
James H. Turner EPA
Power Prod. Dept.
5-18-76
402
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OPERATING EXPERIENCE AND PERFORMANCE AT THE SUNBURY BAGHOUSE
PLANT - GENERAL DESCRIPTION
The Sunbury Steam Electric Station of Pennsylvania Power & Light Company
is located on the west bank of the Susquehanna River in Shamokin Dam, Penn-
sylvania. It generates 402 MW of electricity from six pulverized-coal-fired
steam generators and four turbine-generator sets. Turbines Nos. 1 and 2
are each rated at 87.5 MW. Steam is supplied to these turbines through a
header system from four Foster Wheeler anthracite fired boilers. (A Foster
Wheeler bituminous fired boiler and a 110 MW turbine generator set are also
connected into this header system.) Each of the four boilers is rated at
415,000 Ib/hour steam flow at 955°F 1350 psig and was placed into commercial
service in 1949. The boilers burn a mixture of anthracite (silt and No. 5
buckwheat), petroleum coke, and bituminous coal. (Typical analysis of the
plant coal is given in Table 1.) The mixture varies considerably according to
the availability and the quality of the fuel. The normal mix is 75% anthracite
(consisting of 90% silt and 10% No. 5 buckwheat), 20% coke and 5% bituminous. The
normal fuel consumption is 25 tons/hour/boiler. No. 2 fuel oil is used for
boiler fire-ups and to stabilize ignition.
PARTI CULATE REMOVAL EQUIPMENT
Original particulate removal equipment consisted of a mechanical-electrostatic
precipitator set. Particulate removal efficiency with this equipment was 90%.
In 1973 baghouses were installed on each of the four anthracite fired boilers.
The electrostatic precipitators were gutted due to a weight consideration. The
weight reduction allowed utilizing some of the existing precipitator steel to
support the baghouse. The mechanical collectors were modified by replacing
the original vanes of the multi-cone collectors with new, less efficient vanes.
The effect of this was a reduction of mechanical collector efficiency from 78%
to 70% while the pressure drop decreased from 4.5 in. VWC to 1.5 in. VWC. The mech-
ical collectors were left in to maintain a lower dust loading on the induced draft
fans. It is not certain to what degree, if any, they improve baghouse performance.
The baghouses are installed in series with the mechanical collectors. The
flue gas, after leaving the boiler economizer, passes through regenerative air
heaters, mechanical collectors, the induced draft fans, and the baghouse. The
then clean flue gases discharge to atmosphere through a stack, 300 feet above
the ground. (Figure I)
403
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The baghouses were constructed by Western Precipitator Company. They
were initially put into service on the following dates:
No. 1A Baghouse - February 10, 1973
No. 2B Baghouse - March 5, 1973
No. 2A Baghouse - April 14, 1973
No. IB Baghouse - April 26, 1973
BAGHOUSE - GENERAL DESCRIPTION
Each baghouse is approximately 80 feet long, 40 feet wide and 62 feet high
including the dust hoppers. Each baghouse consists of 14 compartments arranged
in two rows of seven compartments each. Each filter bag compartment is 19' - 9 in.
wide x 10 feet - 1-3/4 in. long x 40 feet high. The bashouse enclosures, includ-
ing the interior partitions, are constructed of 14 gauge mild steel and are of all
welded construction. The 14 gauge partitions and welded construction were decided
upon to insure gas tight construction to permit the safe entering of isolated com-
partments for routine inspections and minor maintenance while the baghouse is
in service.
The entire baghouse including the hoppers, collapsing flues and inlet/outlet
flues up to the point of tie in with the existing breeching are completely heat
insulated and aluminum lagged. The bag filter dust hoppers, located below the
thimble floor, are contained in an enclosure constructed of 0.3:2 in. thick aluminum
siding over steel girts. The floor in the hopper enclosure is checker plate
in order to contain any ash which may spill during maintenance periods or should
malfunctions occur in the ash removal system.
Access for maintaining the filter bags is accomplished in the following
manner. A 2 foot wide x 5 foot high insulated compartment access door is provided
at the thimble floor elevation for each compartment. A checker plate walkway
constructed of 3/16 in. steel plate is provided through the center of the filter
bag rows (three rows of 15 filter bags each are located on each side of the walkway)
to facilitate changing filter bags as well as permit inspection to detect
broken bags. A 2 foot wide x 5 foot high insulated compartment access door and
grating walkway are provided in each compartment for access to the filter bags at
the bag tensioning framework elevation.
An inlet flue and a collapse air flue run the Length of each baghouse
between the rows of hoppers. Side ducts from each flue to the individual com-
partments are provided to the upper portion of the compartment ash hoppers.
404
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The gas inlet damper and the collapse air damper, both timer-operated,
are located in the gas inlet flue and the collapse flue, respectively, to
permit cycling the compartments through the in-service and cleaning cycles.
There is one main collapse fan for each baghouse and two spare collapse fans,
each serving two baghouses.
The outlet flue surmounts the baghouse and ties into the existing stack
breeching. Each compartment has a manually operated outlet flue damper which
permits isolating a compartment during maintenance periods. (See Figure
2 for the baghouse's general arrangement.)
Vent stacks were installed on the top of each baghouse compartment in 1974
and 1975. Originally isolating a baghouse compartment was difficult. Since the
gas inlet damper, a leaf type damper, does not provide a positive shut off, too
much heat and gas leaked into the compartment. With a portable 7,000 CFM fan
mounted on the vent stack enough ventilation is provided to permit entry into
the compartment for maintenance with the remainder of the baghouse in service.
The vents are 2 feet x 2 feet x approximately 4 feet high, constructed of 10 gauge
mild steel and are covered with 2 in. thick insulation and 0.032 in. thick aluminum
siding. A spring clamped cover is mounted on top of each vent stack. (Figure 3)
Filter Bags
The design filter ratio was 1.919:1 CFM/square feet of bag and the active
filter ratio (based on one compartment being out of service at all times) was
2.067:1. Each baghouse compartment contains 90 bags for a total of 1,260 bags
per baghouse giving a total filter area of 115,668 square feet.
Each filter bag is 11-1/2 in. in diameter by 30 feet-4 in. length. The filter
bags, Menardi Southern Company's style No. 601T, are made of glass fiber with
Tuflex finish (Teflon base, specifications are listed in Table 2). The Teflon
coating acts as a lubricant between the glass fiber filaments. It prevents
the glass fiber filaments from abrading one another, and thus lengthens bag
life. Each filter bag contains seven equally spaced anti-deflation rings. With the
rings preventing the bag from collapsing completely during the cleaning cycle,
the dust cake is released freely and drops to the hopper below. A metal
band is sewn into the top and bottom of the bags for installing the bags.
The bags are suspended from a framework in the top of each compartment housing
and approximately 50 Ib. tension is applied.
405
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Cleaning Cycle
The bags are cleaned by reversing the gas flow through a compartment
This partially collapses the bags; some of the dust cake is then released and falls
to"the hopper below. A collapse fan (12,500 CFM @ 300°F and 10 in. WC Buffalo
Forge Company size 60, 1300 rptn ) provides the force for the reverse gas
flow. The collapse air flow is drawn through the collapse duct, into the
collapse fan, and then is discharged into the baghouse inlet flue where any
entrained fly ash is filtered by the bags. (Figure 2)
Each compartment is cleaned in the following manner:
1. The gas inlet damper to the compartment closes, shutting
off the flow of "dirty" flue gas to this compartment.
2. The collapse damper opens, allowing a reverse flow of
"clean" flue gas from the outlet flue to be pulled through
the bags, which partially collapses, and thus cleans the bags.
3. The collapse damper closes.
4. The gas inlet damper opens, returning the compartment to
the filtering mode. (Figure 3 illustrates a compartment in
the filtering mode and a compartment in the cleaning mode.)
This sequence is repeated on each of the 14 compartments. Twice during
the 14 compartment cycle the collapse duct is swept clean. This
is accomplished by opening a sweep valve located at the end of the collapse
duct (Figure 2). This allows ambient air to be drawn through the collapse duct
and fan; and thus, purges any fly ash accumulation.
The time sequence of the cleaning and sweep cycles is detailed in Table 3.
There are null periods between the operation of the gas inlet and collapse
air damper. This is done to prevent reentrainment by allowing for some settling
time for suspended fly ash; and also, to prevent a sudden reversal of air flow
through the bags, which would put undue stress on the bags. The air-operated
dampers are provided with a slow opening and closing device which also prevents
any additional stress on the bags.
Instrumentation
instrumentation
Two instruments, a Bailey bolometer and a differential pressure recorder,
are used to monitor baghouse performance continuously.
406
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The bolometer consists of a light bulb and a receiver located in the
breeching to the stack on each boiler. The receiver output varies as the light
intensity striking it varies, and thus is an indication of smoke density. A
24-hour recorder located in the plant monitors this signal. A reading of zero
of scale indicates no smoke and a reading of 100% of scale indicates 100%
smoke. The bolometers are not calibrated and just give a relative indication
of smoke density.
The differential pressure across each baghouse is recorded on a 24-hour
recorder located in the plant control room. The taps are located in the baghouse
inlet and outlet flues. An audible alarm is sounded if the differential pressure
rises 1 in. VWC above its normal value, giving the operator time to take corrective
i
action.
The differential pressure across each compartment is not monitored. It
was decided that a continuous monitoring system would be too costly and of
questionable value. There are taps across each compartment but no instruments
were installed to indicate differential pressure.
Controls and Alarms
The four baghouse control panels are located in the plant mill room. Each
control panel houses two cleaning cycle timers (one of which is a spare), a
stepping switch, cycle timer selector switch, collapse fan selector switch,
control power "off-on" switch, and the compartment!s status indicating lights.
The cleaning control timers and the stepping switch control the cleaning
cycle operating the air-operated baghouse gas inlet and collapse air dampers
by activating the appropriate solenoid valves. The cleaning cycle control
timers were originally electro-mechanical devices. Solid-state electronic
timers were added in 1975 and 1976 to facilitate adjustment in the cycle time
and increase the reliability.
The compartment status lights indicate the positions of the baghouse gas
inlet and collapse air dampers. These indicate whether a compartment is in
the filtering or the cleaning mode. Indication is accomplished through the use of
limit switches on the dampers.
Adjacent to the baghouse control panels are alarm panels with alarms
to indicate a collapse fan failure or a cleaning cycle timer failure.
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Baghouse Vacuum Cleaning Systems
Individual vacuum cleaning systems are provided for each baghouse.
Vacuum connections have been provided in each bag filter hopper enclosure
as well as outside the compartment access doors at the thimble floor elevation.
Baghouse Ash Removal System
The two fly ash removal systems, one per unit, are United Conveyor Corp.
vacuum type which pneumatically convey fly ash from the ash hoppers. Air flow
required for conveying the ash is produced by an 8 foot Hydroveyor exhauster.
The ash/air mixture is transported by the conveyor to the Hydroveyor exhauster
where the mixture is combined with water to form an ash slurry. This slurry
is discharged into an air separator tank where the conveying air is vented to
atmosphere while the ash slurry discharges by gravity into the Units Hbs.
1-2 slurry tank.
High pressure (300 psi) water at a rate of 750 GPM is provided at the
Hydroveyor for each system by one of three new raw water booster pumps. Plant
ash disposal capabilities were increased by the duplication of the existing
fly ash disposal system and also the installation of two new 2,000 GPM raw
water pumps. The new fly ash disposal system is composed of one ash slurry
tank, an ash slurry pump house, six slurry pumps (in series) with 1600 GPM capacity
and over 2 miles of 10 in. buried steel pipe through which the slurry is pumped
to the fly ash settling basin.
OPERATION
The flue gas enters the baghouse at 350°F and leaves at 310°F. The flue
gas temperature normally stays within this range except during boiler
startups and shutdowns when the temperatures are in a lower range, anywhere
from ambient air temperatures to normal operating temperature. Several times
since 1973 the regenerative air heater drives have failed. This has caused
the baghouse inlet temperature to rise to as high as 510°F. The high temperatures
were sustained for about 1 to 2 hours. There were no discernable effects
on the bags.
The design pressure drop across the baghouse and baghouse duct work was
not to exceed 6 in. VWC, which was dictated by the capacity of the existing induced
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draft fans. Average draft loss across the baghousesls as follows:
Baghouse No. Avg Draft Loss (in, VWC) Bags Installed
IB 3.5 4/26/73
2A 3.5 11/20/74
2B 3.0 2/23/75
1A 2.5 3/20/75
The baghouses are on the discharge side of the induced draft fan and thus,
are under a positive pressure.
The gas flow through each baghouse is 222,000 ACFM at a temperature of 325°F.
Baghouse inlet grain loading is approximately 2 grains/cu ft.
There is no visible stack discharge from any of the baghouses, except for
vapor condensation during the coldest days of the year (approximately 10°F or
lower). When a large bag failure occurs (e.g., a 2 to 3 foot tear in a single bag)
there will sometimes be a very faint stack discharge (approximately 5% opacity
or less) depending on the location of the bag failure.
Boiler Start-Tips and Shutdowns
To prevent blinding of the bags with the oil used during boiler fire-ups,
before a new set of bags is placed in service they are precoated with fly ash
remaining in the boiler gas passes. The boiler is brought on the line and the
baghouse cleaning cycle is not activated until 1 hour after coal is fired. This
allows an additional coating to form on the bags.
Each boiler is shut down and fired-up several times a year. Outside of
the precoating of new bags, no special precautions are taken. During normal
fire-ups, shutdowns and outages the cleaning cycle is left on. The continued
cleaning during outages does not remove the initial fly ash coating but will
prevent the formation of a thick cake, which can accumulate from moisture carry-
over while water lancing boiler slag deposits.
The number of fire-ups for each boiler is given below.
Boiler No, 1973 1974 1975
1A 632
IB 352
2A 335
2B 733
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Operating Problems
The major operational problem has been occasional baghouse high differ-
ential pressure. An alarm sounds to warn the operator of the high differential.
The usual causes are either collapse fan failure, cleaning cycle timer failure,
or gas inlet or collapse air damper failure. Once the problem is corrected
the differential pressure returns to normal in a short time, approximately
30 minutes. Normally the problem is corrected without affecting boiler operation.
However, if the differential increases to approximately 5 or 6 in. VWC the boiler
steam flow must be reduced due to the loading on the induced draft fans.
Baghouse problems infrequently (about 5 or 10 times/year for all baghouses)
have required reductions in boiler load of about 20,000 to 40,nno lb/hr
steam flow, but have never forced a boiler off the line.
At present there are no alarms for collapse fan failure. Flow switches
were installed in the collapse fan duct to indicate fan failure, but the
switches proved to be inoperative. It is planned to install other alarms. At
present collapse fan failure is detected from increased differential pressure
signals.
When a main collapse fan fails the spare collapse fan is put into service
by opening blast gate (butterfly type) dampers. The spare fan is normally filled
with fly ash due to leakage past the blast gate dampers, and normally the fan
must be cleaned out before it is put into service. This can take as much as 2
to 3 hours. Originally the spare fan was isolated by sliding gate dampers.
These dampers provided a tight seal; however, they were difficult to open and
close. It took four men with a chain hoist approximately 4 hours. There
is a study to determine a method to prevent fly ash accumulation in the spare
fan and still allow easy damper operation. As a temporary measure the spare
fans are pressurized with compressed air to prevent in-leakage.
If a cleaning cycle timer fails an alarm is sounded and the operator
switches to the spare timer. Originally the timers failed frequently. This
has been alleviated by adding more reliable, electronic timers.
Damper failures can sometimes be detected by observation of the differ-
ential pressure chart. As the dampers open and close the differential pressure
swings. If a damper fails, the absence of this pressure swing leaves a "gap"
on the differential pressure chart (Figure 4). If a high differential pressure
alarms, the dampers are routinely checked for proper operation. The operator
must go up to the baghouse and visually observe damper operation through the
complete cycle (a total of 32 minutes).
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When a baghouse hopper becomes plugged and fills with fly ash, a high
pressure spike, in sequence with the 32 minute cleaning cycle, is indicated
on the differential pressure recorder. No operational problems are experienced.
The operator must sledge-hammer and/or air-lance the hopper to free the
blockage.
Since the number of bag failures has been small, they have not caused
operational problems. Most failures are single-bag; two-bag failures
are the most amount occurring at a single time. Isolating a compartment
increases the differential pressure by approximately 0.5 in. VWC. Two compartments"
have been taken out of service together with an increase of approximately 1.0 in.
VWC in differential pressure and no reduction in boiler load was necessary.
Any number of compartments may be taken out of service without affecting
boiler load as long as the differential pressure remains below 6.0 in. VWC.
There .have been two major baghouse pressure excursions. On Baghouse No. 2A
in July 1975 after a boiler outage the draft loss increased from 3.0 in. to 6.0 in.
WC over a 3 week period. This was apparently due to moisture carry over
to the bags from heavy water-lancing of boiler slag deposits during the outage.
Several methods were tried to reduce the draft loss: two collapse fans were
placed in service on the baghouse, the collapse dampers were opened and closed
several times during each 2 minute compartment cleaning time (this was
attempted to dislodge any thick dust cake on the bags), all hoppers were
inspected and any blockage was cleared,and the thimble floor was vibrated by pounding
it with sledge hammers to dislodge any buildup on the thimbles. (This thimble
buildup, a fly ash deposit reducing the 12 in. diameter thimble size to 6 in. to 9in.
diameter, has been noted on all baghouses but never has had any noticeable
effect on baghouse performance.) All of these methods proved unsuccessful. A
compartment was isolated and the bags were inspected. The dust cake on the bags
was very thick, from 1/2 in. to 3/4 in. It was decided to manually
shake the bags. Each bag was grasped between the palm of each hand and
vigorously shaken from side to side. The release of the dust cake could be
felt as the dust fell to the hopper below. Alternately, all of the compartments
were taken out of service and the bags were manually shaken. When all com-
partments were returned to service the pressure drop had decreased to 3.5 in.
WC and has remained stable since.
In the fall of 1975 the pressure drop on Boiler No. IB increased from
3.5 to 5.0 in. VWC over a 3 month period. The bags were inspected during
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the boiler outage. The cake was at most about 1/4 in., so the bags were not
manually shaken. All other methods used on No. 2A were tried with no success.
With the new electronic timers it is possible to readily change the cleaning
cycle time. On January 23, 1976 the time was reduced from the 30 minute
cycle to a 15 minute cycle, doubling the number of collapses. In
the last part of February the draft loss began to decrease until it reached
3.5 in. WC at the end of February. The cleaning cycle was returned to the normal
30 minute cycle on March 16, 1976. The draft loss has remained stable at
3.5 in. WC since. Because of the time frame it is uncertain whether the 15 minute
cleaning cycle caused the draft loss reduction. Normally, when a change is
made to reduce draft loss the effect is noticed in several hours. Additional
background information on the Sunbury installation may be found in references
1, 2 and 3.
MAINTENANCE
Most maintenance hours have been spent on bag replacement, collapse fan
repairs and air-operated damper repairs.
Bag Replacement
The procedure for bag replacement is detailed in Figure 5.
During each annual boiler outage all baghouse compartments are inspected.
If an accumulation of dust is found on the compartment floor, each bag in the
compartment is inspected for possible failure and all failed bags are replaced.
To replace a bag with the boiler in service, the associated compartments
must be isolated by closing the gas inlet and outlet dampers. The lower and
upper doors of the compartment are opened to allow ambient air to circulate. The
cover is removed from the vent stack, and a portable 7,000 CFM fan is set on the
vent stack and started to provide forced ventilation. It normally takes 3 to
4 hours to ventilate the compartment sufficiently for men to enter. It is
normally still necessary to wear a malsk to prevent inhalation of sulfur dioxide.
The compartment can then be entered and the failed bag replaced. The entire
procedure (isolating and ventilating the compartment, finding the leak, replacing
the bag, and returning the compartment to service)takes approximately 6 to 8
hours.
Detecting Bag Failures
Bag failures are detected by daily observations of the bolometer charts.
When a bag has failed the bolometer senses the increased particulate emissions;
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generally, there is no visible stack discharge. The bolometer chart will
indicate a periodic spike, in sequence with the cleaning cycle. There will
be a decrease in the bolometer reading when the compartment with the failed
bag is removed from service, a spike when the compartment is returned to service,
and a settling out to a higher than normal reading directly after (Figure 6).
Some spikes are not easily discernable and a careful study of each chart is
necessary. Also some non-periodic spikes occur, but these do not indicate
an abnormal condition.
When it is determined what compartment has the failed bag, the compartment
is taken out of service. The bolometer readings then return to normal. The
compartment is entered, and each of the ninety bags is inspected. An accumulation
of fly ash on the compartment floor is a tell tale sign that there is an actual
bag failure in the compartment. However, bag failures have been found in
compartments with no fly ash accumulations. The bags are inspected by holding
a flashlight at the bottom of the bag and shining it up the side of the bag. Any
tears are illuminated in this manner. Also, a slight tap on the bag will, if
there is a failure, cause a stream of fly ash to flow from the bag which is
illuminated by the light; providing an additional check for failures.
The failed bag is replaced, the compartment floor is cleaned, and the
compartment is returned to service.
Bag Failure Rate
As of May 1, 1976 a total of 37 bags have failed. A list of the failures
is detailed in Table 4.
As a precautionary measure bags in three baghouses were replaced after
approximately 2 years of service (dates are listed in Table 4). The 4th bag-
house, No. IB, still has its original set of bags. The bags will remain
in service until an increased bag failure rate dictates their replacement.
Maintenance Problems
The collapse fans have been a continual problem. The fans were not
designed to handle dirty gas. There is baffling in the hoppers to prevent the
collapse fan from picking up dust. However, this has not been effective. The
fan casings and blades are worn thin from the abrasiveness of the fly ash. This
has caused excessive fan vibration and gas leakage from the fan. Frequent
maintenance is required on the fans.
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The air operated dampers have caused continual problems. They are
located in an environment of high temperatures (150°F to 200°F) and occasional
dusting. This results in increased maintenance and occasional damper failure.
ECONOMICS
The initial installation cost of all four baghouse installations, including
the fly ash removal system, was $5,500,000. Installation cost and yearly
operating and maintenance costs are detailed in Table 5. Increased
mechanical and electrical maintenance cost in 1975 was the result of increased
collapse fan maintenance and the high pressure drop problem encountered on
Baghouse No. 2A.
PERFORMANCE
In order to assess the performance of a fabric filter system as applied
to a utility boiler, EPA askedfor permission to perform efficiency tests after
the Sunbury baghouses had been in service for a few months. Permission was
4
granted and one set of tests was made in August 1973. The results of these
tests were impressive and further testing was scheduled to be done by GCA
Corporation. The results of the GCA testing, performed in January, February
2
and March 1975, are discussed in this paper.
Of greatest interest to EPA was the performance of the baghouse in the
fine particulate range, i.e.,below about 3 ym. Along with mass efficiency
testing there were also measurements made with impactors and diffusion
apparatus. The problems associated with the latter two types of devices led
to some rather large error bands, but with over 30 tests made there is reasonable
confidence in the results.
GCA performed their tests after the baghouses had been in service for
approximately 2 years. Since bag changeover was to coincide with a scheduled
outage, this situation allowed testing of bags at the end of their planned
service life and also (for the new set of bags) at the beginning of their
service life.
Mass Results
For 19 tests made at several combinations of fuel mix and boiler load,
the average outlet concentration was 0.0017 grains/dscf or in terms
of emission rate: 0.0041 lbs/10 BTU. These tests were for the bags which
had been in service for 2 years. Based on an average inlet loading of
2.6634 grains/dscf the average efficiency was 99.94%. For the new
414
-------
bags the average outlet concentration was 0.0024 grains/dscf; emission rate
was 0.00543 lbs/10 BTU; average inlet loading was 2.2853 grains/dscf and
average efficiency was 99.89%. The slightly higher outlet loading was con-
sistent with slightly lower pressure drop through the new bags, and the new
bags would be expected to have performance similar to the old bags after
sufficient stabilization time. The results do indicate that baghouse
efficiency does not deteriorate with time, and tends to get better through the
service life of the bags.
For all 31 tests the range of outlet loadings was from 0.0013 to 0.0033
grains/dscf (0.0028 to 0.0101 lbs/106 BTU).
Impactor Results
Apparent collection efficiency by particle size is shown in Figure 7.
For the size range covered, 10 to 0.1 ym, it is apparent that the
baghouse maintains nearly uniform, high reduction of particles entering the
system. All sizes in the impactor range are collected with better than
99% apparent efficiency. There is some question as to whether or not
particles can change their size as they go through the collection system, but
the result at the baghouse outlet is a clear stack.
Condensation Nuclei Counter Results
A condensation nuclei counter (CNC) was used to count numbers of particles
in the range from about 0.003 to 0.5 ym. Although the data are sketchy they
indicate an apparent collection efficiency of between 93.0 and 99.7%
for particles in this size range. It should be emphasized that this is a number
efficiency and not a mass efficiency.
General Comments
It is obvious that the Sunbury baghouses (as do the Nucla baghouses)
do a fine job at collecting particulate matter. The GCA testing also revealed
that collection efficiency was not significantly affected by changes in fuel
composition, boiler load, filtration velocity or cleaning regime within what
might be normal limits of change. The baghouse seems to operate like a
smoothing device; regardless of inlet variations the outlet showed very little
change. For the changes that did occur, regression equations were formulated
415
-------
which showed particle penetration to be primarily a function of inlet concen-
tration, percent moisture, filtration velocity and pressure drop.
Summary of Baghouse Performance
As far as filtration performance is concerned, the Sunbury baghouses do
an excellent job. Mass efficiency is 99.9 percent, grade efficiency is 99%
or better in the impactor range, and the number efficiency for particles
in the CNC range appears to be between 93 and 99.7%. Changes in process
and collector operating variables have little effect on outlet concentration.
416
-------
REFERENCES
1. Wagner, N.H. and D,C. Housenick. Sunbury S.E. Station - Unit
Nos. 1 & 2 Design and Operation of a Baghouse Dust Collector for a
Pulverized Coal Fired Utility Boiler. Pennsylvania Power & Light
Co., Allentown, Pa. May 1973.
2. Cass, R.W. and R.M. Bradway. Fractional Efficiency of a Utility
Boiler Baghouse - Sunbury S.E. Station - G.C.A. Corporation
NTIS No. (to be assigned). EPA-600/2-76-077a, March 1976.
3. Operating Instruction for Therm-o-flex Filters for Pennsylvania
Power & Light Co., Sunbury, Pa., Joy Manufacturing Company, Western
Precipitation Division, February 1973.
4. Harris, D.B. and J.H. Turner. Particulate and S02/SO-J Measurements
Around an Anthracite Steam Generator Baghouse. Particulate and Chemical
Processes Branch, Control Systems Laboratory, Office of Research &
Development, Environmental Protection Agency, National Environmental
Research Center, Research Triangle Park, N.C. November 8, 1973.
5. Bradway, R.M. and R.W. Cass. Fractional Efficiency of a Utility
Boiler Baghouse — Nucla Generating Plant. GCA Corporation. NTIS No.
PB 246 641/AS. EPA 600/2-75-013a.
417
-------
Table 1. TYPICAL PLANT COAL ANALYSIS
Anthracite Bituminous Petroleum Coke
Heating Value (wet basis) BTU/lb) 8,680 12,159 14,054
% Ash 28.0 15.3 1.7
% Moisture 16.4 5.3 6.4
% Volatile Matter 8.4 27.1 13.3
% Sulfur 0.8 2.6 5.3
Table 2. MENARDI FILTER BAG, STYLE NO. 601T
Fabric Glass Fiber
Average Permeability 38.5
Mullen Burst Strength (Avg. P.S.I.) 595
Weight (oz. per sq. yd.) 10.0
Thread Count 54 X 30
Weave 3X1 Twill
Finish Teflon 10% by weight
Warp Yarn ISO's 1/2
Fill Yarn Bulked 1/4
418
-------
Table 3. CLEANING/SWEEP CYCLE TIME SEQUENCE
Elapsed Time (Sec)
I. Clean compartments (1) through (7) one at a time
Gas Inlet Damper starts to close 0
Gas Inlet Damper closed 2.0
Collapse Air Damper starts to open 16.0
Collapse Air Damper open 17.0
Collapse Air Damper, starts to close 68.0
Collapse Air Damper closed 69.5
Gas Inlet Damper starts to open 81.0
Gas Inlet Damper open 83.0
Next Gas Inlet Damper starts to close 122.0
II. Sweep Collapse Air Duct
Last Compartment Gas Inlet Damper open 0
Sweep Valve open 40
Sweep Valve closed 120
Next Compartment Gas Inlet Damper open 165
III. Clean compartments (8) through (14) one at a time.
Repeat cycle shown under I
IV- Sweep Collapse Air Duct
Repeat cycle shown under II
419
-------
Table 4. BAG FAILURES"
Baghouse
1A
IB
2A
2B
1973
5
0
3
4
1974
1
6
0
7
1975
2
2
0
2
To May
1976
2
2
0
1
TOTAL
10
10
3
14
TOTAL
37
Complete Bag Replacements
1A 3-20-75
IB Bags never replaced
2A 11-20-74 (except for compartment No. 8)
2B 2-23-75
a. Failures include neither those from faulty installation or damage
during maintenance (about 23 bags) nor test bags.
420
-------
Table 5. INSTALLATION, OPERATION AND MAINTENANCE COSTS
Bag Filter Installation Cost
Expenditure Description
Western Precipitator Contract (4 Baghouses)
Structures & Improvements
Land & Land Rights
Boiler Plant Equipment
Ash Removal Equipment
Precipitator Modifications
Accessory Electrical Equipment
Overhead
Total Construction Cost
Total Cost ($)
3,201,700.
372,100.
1,500.
1,190,000.
72,000.
662,800.
5,500,100.
Estimated Operating & Maintenance Cost (@ $12/manhour)
Cost Description 1973
Collapse Fan Power Consumption
Air Compressor Power Consumption
Instrument Department Labor
Mechanical Maintenance Labor
Electrical Maintenance Labor
Construction Department Labor
Complete Bag Replacement
Material ($ 48,000/baghouse)
Labor ($ 11,000/baghouse)
TOTAL COST
Cost ($)
1974 1975
18,600.
3,000.
950.
2,130.
7,410.
3,950.
18,600.
3,000.
950.
5,840.
3,800.
2,350.
48,000.
11,000.
18,600.
3,000.
1,800.
15,678.
6,000.
2,000;
96,000.
22,000.
36,040. 93,540. 165,078.
421
-------
BLANK OFF PLATES
ECONOMIZER
AIR HEATER
MECHANICAL
COLLECTOR
AIR
FROM
FORCED DRAFT
FAN
t INDUCED
jj) DRAFT FAN
Figure 1. Location of dust removal equipment
422
-------
Sweep
Valve
Gas Inlet
Dampers
Outlet Flue
Outlet Dampers
Access Doors
Collapse
Air Damper
Collapse-
Air Fans
Figure 2 - Baghouse General Arrangement
423
-------
I-GAS INLET DAMPER-OPEN
2-GAS INLET DAMPER - CLOSED
3-BAG COLLAPSING DAMPER-OPEN
4-BAG COLLAPSING DAMPER - CLOSED
5-OUTLET DAMPER-OPEN
Vent Stack
FROM
I.D. FAN
Figure 3. Gas flow through baghouse comparUsents during normal
operation and cleaning
424
-------
Figure
Differential Pressure Chart - indi eating a damper
malfunction on 2A, baghouse (notice arrows).
A norzaal differential pressure chart has no gaps.
425
-------
Bag support frame
Bag support frame
INSTRUCTIONS FOR HANDLING AND INSTALLATION OF
CLASS CLOTH FILTER BAGS
Handling
Class cloth is soft, flexible, and durable. How-
ever it is subject to types of damage not always
associated with other fabrics. Care in its hand-
ling will be repaid in greater efficiency and
longer life.
Unpack bags carefully, one by one, as they are
installed. Avoid contact of the fabric with
projections of metal or rough wood which may
cause snags or tears.
Never lay bags on the thimble-plate of the bag-
house where they may be walked upon.
The fabric should never be pulled by grasping
between thumb and fingers, which may cause slip-
page of the yarns. Folded bags should not be
subjected to heavy pressure.
Installation
Step One. After unpacking, place filter Bag 3
in box on thimble plate near position of assembly.
Raise one end of filter bag and insert Cap 2 as
shown in Detail 1. Rotate cap 90° and pull up-
ward until metal band sewn into bag cuff fits
snugly against rolled rim of cap, as shown in
Detail 2.
Step Iwo. Hook onto Cap 2 with hoisting rope and
raise bulk of Bag 3 off thimble plate. Place
lower end of bag over thimble, as shown in Detail
2. Push bag end downward until thimble rim is
completely covered. Then gently and evenly. Move
bag end upward until neta! band sewn into bag cuf:
fits snugly against rolled rim of thimble.
Sfcej> Three. Hoist bag to full height and replace
rope with Hanger 1 as shown in Detail 3. Tighten
hanger to develop 50# tension on bag by compres-
sing hanger spring 1% inches.
Install bag «nd marked "bottom" at hopper floor level
Figure 5. Filter Bag Detail
-------
Figure 6. Bolometer Chart - indicating a bag failure in compartment
No. lit of No. 2B baghouse.
427
-------
J>
K>
00
u
23
w
i—i
u
w
H
O
w
I-J
o
o
H
23
Pn
PM
99.9
99
95
90
50
_FILTRATION VELOCITY = 2.0ft/min.
(1.0 cm/sec)
OUTLET LOADING = 0.0017 grains/DSCF
(0.0039 grams/m3)
10
0.1
0.01
0.01
0.1
1.0
0.01
0.1
5
10
50
z
o
-------
PARTICULATE REMOVAL CHARACTERISTICS OF
THE 170 m MOHAVE SCRUBBERS
Dale G. Oones
Alexander Weir, Or.
John M. Johnson
W. Carl Martin
Southern California Edison Company
and
Seymour Calvert
Air Pollution Technology, Inc.
429
-------
Table of Contents
Page
Number
1.0 ABSTRACT 1
2.0 INTRODUCTION 2
3.0 TEST MODULE DESCRIPTION 2
4.0 OBSERVED SCRUBBER PERFORMANCE 5
4.1 Particulate Removal at Design Operating Conditions 5
4.2 Effect of Turndown 10
4.3 Effect of Liquid Flowrate 12
4.4 Miscellaneous Effects 12
5.0 DISCUSSION 15
5.1 Data Acquisition and Reduction 15
5.2 Characteristics of Mohave Flyash 17
5.3 Performance Predicted by Inertial Impaction 1-7
6.0 CONCLUSIONS 23
7.0 ACKNOWLEDGMENT 24
8.0 REFERENCES 25
9.0 APPENDIX 1 - Particulate size distribution analysis 26
430
-------
PARTICULATE REMOVAL CHARACTERISTICS
OF THE
170 MW MOHAVE SCRUBBERS
1.0 ABSTRACT
This paper presents the particulate removal data obtained
during the 170 MW Test Modules Program at the Mohave Generating
Station. Three scrubber configurations are evaluated: Vertical
TCA, Vertical PPA, and Horizontal Crossflow. The effects of
scrubber variables such as turndown ratio, flue gas pressure
drop, liquid to gas contacting ratio and packing characteristics
on particulate removal are examined.
A method for predicting particulate removal based on the
percentage of electric power consumed by the scrubber and the
inlet grain loading characteristics is developed from impaction
collection theory and Test Program data. At design operating
condition, the Vertical TCA Module required about 357° more elec-
trical power than the Horizontal Module to achieve the same
degree of particulate removal.
431
-------
2.0 INTRODUCTION
Between July, 1971 when pilot plant testing was initiated at
the Mohave Generating Station and July, 1975 when the 170 MW Test
Modules Program was completed, the participants of the Navajo and
Mohave Power Projects funded a scrubber development and evaluation
project which has significantly advanced the understanding of
scrubber technology. The results of this evaluation have been
published,in part, at several EPA Flue Gas Desulfurization
Symposium conferences, 1»2,3,4. The objectives of the 170 MW Test
Modules Program were to evaluate the S02 and particulate removal
performance of more than one full-scale scrubber, as well as to
determine the practical feasibility of scrubbers for commercial
service, including cost, operating flexibility, maintenance
requirements, materials of construction, chemical control problems,
and waste disposal factors.
The purpose of this paper is to describe the particulate
removal characteristics of the 170 MW Mohave Test Modules. During
the test program, a total of 154 inlet/outlet particulate samples
were obtained during tests of three different 170 MW scrubber
configurations. The scrubbers were located downstream from the
cold-side electrostatic precipitators, and were designed to treat
450,000 SCFM of flue gas (equivalent to 170 MW at the Mohave
Generating Station).
3.0 TEST MODULE DESCRIPTION
The 170 MW Vertical Module, located on Unit 1, was tested
in the Turbulent Contacting Absorber (TCA) configuration and the
Polygrid Packed Absorber (PPA) configuration. The TCA configura-
tion, shown in Figure 1, consisted of four stages of thermoplastic
rubber balls supported on stainless steel grids at four foot inter-
vals. The balls were contained in compartments, with 15 compart-
ments at each of the four levels. Although various levels of ball
depths were tested, the compartments were initially filled to the
one-foot level with approximately 1,600,000 balls. The TCA config-
uration was tested with both three and four stages of balls, and
at ball depths of 6, 10, and 12 inches in the four stage configura-
tions. The PPA configuration consisted of two and three stages of
fixed "eggcrate" packing supported at four foot intervals. The
packing consisted of plastic grids 1%" thick with 2" square open-
ings, with the grids stacked to a depth of approximately 17 inches
in each stage.
The 170 MW Horizontal Module, located on Unit 2, was tested
with 1,2,3,4 and 5 spraying stages. The basic four stage configu-
ration, shown in Figure 2, consisted of crossflow spray discharged
from a row of 36 externally-mounted nozzles at each stage. The
Horizontal Module did not contain packing, but consisted of four
stages of cross flow spray. The slurry was cycled through the
scrubber in a countercurrent manner. That is, the fresh lime slurry
432
-------
FIGURE 1
170MW VERTICAL MODULE (4STAGE TCA)
SLURRY SPRAYS
FLUE GAS
FROM FAN
SUMP CHAMBER
SCRUBBED FLUE GAS
TO DEMISTER
.40
THERMO PLASTIC
RUBBER SPHERES
"QUIESCENT" SPHERE
DEPTH OF 6" PER
STAGE
38
433
-------
FIGURE 2
170 MW HORIZONTAL MODULE
(4 STAGE)
FLUE GAS
/FROM FAN
SCRUBBER
FLUE GAS
TO DEMISTER
-------
from the mix tank was first sprayed across the flue gas at the
fourth stage, or discharge end of the scrubbing chamber. The
same liquid was successively collected and pumped to the third,
second, and first stages and successively depleted of alkalinity.
By the time the slurry reached the first stage collection hopper
for return to the lime mix tank, it was almost completely depleted
of any excess alkalinity. The recirculating slurry flow rate was
normally 9,000 gpm, but the slurry was mechanically pumped four
times per circuit for a total installed pump capacity of 36,000 gpm.
Both the Horizontal and Vertical Modules were designed for
identical test conditions, using the same sources of makeup water
and removing particulate matter resulting from the combustion of
low sulfur western coal from the Black Mesa mine near Kayenta,
Arizona.
4.0 OBSERVED SCRUBBER PERFORMANCE
4.1 Particulate Removal at Design Operation Conditions
The observed operating conditions of the three full-
scale scrubbers are indicated in Table 1. The outlet gas flowrate
was greater than the inlet due to water vapor evaporated inside the
scrubber. In addition, the Horizontal Module used an indirect
reheater which injected approximately 80,000 SCFM of hot ambient
air into the outlet flue gas stream. The pressure drop indicated
in Table 1 is the flue gas pressure drop across the scrubbing
chamber itself, not including additional pressure drop due to
ductwork, mist eliminator and reheater restrictions. The electric
power consumption in Table 1 is the total electric power measured
for the entire scrubbing system, and includes control room light-
ing and air conditioning, thickener rake drive, tank agitators,
process water pumps and other miscellaneous power requirements.
However, these miscellaneous factors are estimated to be about 5%
of the total, while fan and slurry pump power requirements accounted
for the remaining 95% of the total.
The observed outlet grain loadings obtained at the design
operating conditions for the Horizontal, Vertical TCA and Vertical
PPA scrubbers are plotted as a function of inlet grain loading in
Figures 3, 4 and 5, respectively. At an inlet grain loading of
0.10 gr/SCF, all three scrubber configurations demonstrated a
relatively high degree of particulate removal between 90% and 93%.
As will be pointed out in the discussion, the percentage particu-
late removal changed with inlet grain loading due to changes in
the average particle size and particle size distribution.
435
-------
TABLE 1
OPERATING CONDITIONS AT DESIGN GAS FLOWRATE
INLET GAS FLOWRATE
OUTLET GAS FLQWRATE
NUMBER OF STAGES
L/G RATIO (GAL/MSCF)
SCRUBBER GAS VELOCITY
PRESSURE DROP
NOZZLE PRESSURE
POWER CONSUMPTION
TYPE OF REHEAT
HORIZONTAL
450,000 SCFM
560,000 SCFM
4
21 (PER STAGE)
22 FT/SEC
1,0 IN, H00
35 PSIG
2,6 MW
INDIRECT
L
VERTICAL TCA
450,000 SCFM
480,000 SCFM
36
12 FT/SEC
14,0 IN, H20
15 PSIG
3,4 MW
DIRECT
VERTICAL PPA
450,000 SCFM
480,000 SCFM
3
60
12 FT/SEC
12,2 IN, H20
30 PSIG
3,9 MW
DIRECT
-------
FIGURE 3 INLET vs, OUTLET GRAIN LOADING
170 MW HORIZONTAL MODULE
1,00
0,50
0,20
0,10
0,05
0,02
0,01
0,005
:+•
-H-
0,001 0,002 0,005 0,01 0,02
0,05
AVFRAGE OPERATING CONDITIONS
INLET FLOWRATE: 450,000 SCFM
OUTLET FLOWRATE; 560,000 SCFM
L/G (PER STAGE); 21 GAL/MSCF
NOZZLE PRESSURE: 35 PSIG
437
NO, OF STAGES: 4
POWER REQ'D: 2,6 MW
AP: 1,0 IN,H20
GAS VELOCITY: 22 FPS
-------
FIGURE 4 INLET vs, OUTLET GRAIN LOADING
170 MW VERTICAL TCA MODULE
1 00 ::i::3:
3,50 --
0,20 ^'-~
Olfl "'"fi
,1U --
0,05 --
0,02 *"::
0 01 ::':=::
U , UJ.
0,005=!!!;;!
1 J L EC
ft
HW
= *
_ —
1
_ — i
1
%
¥
—
-
i
<
-
i
^
i
1
I
1
T§
— (i'd
, — P-.
;!|EEEE"E"
.±.1 ^* 3,j-.
ri
r
!
|
'
;;j, __::
- *--- -1 r- -
_:j.
^* '
• ''T sHL
* i" S f
Bf
;.-:,.5f J_.
H
/ ,,.
FfT
^i
s
"zij±::::::3::i
I
-------- :"|::|'-:
— -j-. --
-i M --
E|EEE|---:EE
1
i|»i|::::::|=:
_ :::::::::! :
T
J_
HI |||i in |
0,001 0,002
0,005 0,01 0,02
0,05
AVERAGE OPERATING CONDITIONS
INLET FLOWRATE: 450/000 SCFM
OUTLET FLOWRATE: 480/000 SCFM
L/G RATIO! 36 GAL/MSCF
NOZZLE PRESSURE: 15 PSIG
NO, OF STAGES: 4
POWER REQ'D: 3,4 MW
AP: 14 IN, H20
GAS VELOCITY! 12 FPS
438
-------
FIGURE 5 INLET vs, OUTLET GRAIN LOADING
170 Fiw VERTICAL PPA MODULE
n'sn
0,10!
0,05
0,1
0,01;
0,1
0,001 0,002 0,005 0,01 0,02
0,05
INLET FLOWRATE: 450,000 SCFM
OUTLET FLOWRATE: 480,000 SCFM
L/G RATIO; 60 GAL/MSCF
NOZZLE PRESSURE; 30 PSIG
NO, OF STAGES: 3
POWER REQ'D; 3,9 MW
AP: 12,2 IN, H2o
GAS VELOCITY: 12 FPS
439
-------
4.2 Effect of Turndown
During the Test Program, it was determined that the
easiest method of operating the scrubbers was to maintain a
constant liquid flowrate and allow the flue gas flowrate to change
under simulated load-following conditions. From an engineering
design standpoint, this simplification avoids the severe erosion
problems encountered with slurry flow control valves. It was
decided to determine the effect on particulate removal of gas
flowrate turndown with the liquid flowrates held at the design
values shown in Table 1.
The observed variation in percentage particulate removal
at a constant inlet grain loading of 0.10 gr/SCF is plotted as a
function of the gas flowrate in Figure 6. Note that the particu-
late removal efficiency of the Horizontal Module increases as a
function of turndown ratio.
One explanation for this is the change in energy input
per unit of flue gas flowrate which occurs with increasing turn-
down ratio. The primary method of shattering slurry droplets and
thus contacting particulate matter differs between the types of
scrubbers tested.
The Vertical TCA and PPA Modules require flue gas pressure
drop to increase the tendency to shatter slurry droplets. The
particulate removal was thus seen to decrease with decreasing
pressure drop. For example, at 450,000 SCFM, the particulate
removal was about 90% at the design L/G ratios. Under these con-
ditions the pressure drop was 14 inches of water in the TCA and
12.2 inches in the PPA. When the flue gas flowrate was decreased
to one-third the design value, the pressure drop decreased to 3.5
inches of water in the TCA and 2.0 inches in the PPA. In spite of
the fact that the L/G ratio was now three times the design value,
the particulate removal decreased to about 7570.
The Horizontal Module utilized nozzle pressure to shatter
the slurry droplets, and particulate removal was thus seen to increase
with decreasing flue gas flowrate. As the flue gas flowrate was
turned down from 450,000 SCFM to 150,000 SCFM, the pressure drop
across the scrubbing chamber decreased from 1.0 inches to water to
0.10 inches. Since the liquid flowrate was constant, a given amount
of flue gas was contacted by three times as many droplets at one-
third load as at full load. The particulate removal efficiency was
observed to increase from 92.5% at full load to 96.570 at one-third
load.
440
-------
441
-------
4.3 Effect of Liquid Flowrate
The effect of increasing the Vertical TCA Module liquid
flowrate beyond the levels indicated in Figure 6 was studied as
shown in Figure 7. It is seen that an increased liquid flowrate
was more effective at a high inlet grain loading than a low inlet
grain loading. For example, the same particulate removal achieved
with 18,000 gpm at 450,000 SCFM (at an L/G ratio of 40) could be
achieved at 350,000 SCFM by doubling the liquid flowrate to 36,000
gpm (at an L/G ratio of 103). The same results could not be
obtained at a reduced inlet grain loading of 0.03 gr/SCF.
With the Horizontal Module, particulate removal is only
improved as the flue gas flowrate is reduced. It was therefore
important to determine how rapidly the particulate removal would
decrease with decreasing liquid flowrate. As shown in Figure 8,
a reduction in liquid flowrate from 9,000 gpm/stage to 4,500 gpm/
stage would result in a decrease in particulate removal from 92%
to 78% at an inlet grain loading of 0.10 gr/SCF and from 84% to 66%
at an inlet grain loading of 0.03 gr/SCF. These tests were conducted
by shutting off individual nozzles to maintain 35 to 40 psig nozzle
pressure, and although the spray coverage patterns were undoubtedly
altered, the slurry droplet size probably did not change. The
required liquid flowrate for a desired degree of particulate removal
can thus be estimated from Figure 8.
4.4 Miscellaneous Effects
Several miscellaneous effects were investigated during the
Test Program. For example, it was found that increasing the static
ball depth from 6" to 12" in the Vertical TCA Module increased the
particulate removal.
The effect of variations in nozzle pressure was studied
with the Horizontal Module. Particulate removal at an inlet grain
loading of 0.032 gr/SCF increased as the nozzle pressure was increased
from 15 psig to 30 psig.
Finally, it was confirmed that increasing the number of
contacting stages improves the particulate removal efficiency. For
example, the particulate removal with the Vertical PPA Module
increased as the number of packed stages was increased from two to
three. Similar results were obtained with the Horizontal Module,
where the effect of four versus five stages was studied in detail.
A comparison of fractional collection efficiencies with four and
five stages of scrubbing indicated that the extra stage increased
the collection of 1.5 micron (aerodynamic) particles. An interest-
ing description of particulate collection in spray tower applica-
tions due to condensation effects can be found in Calvert, et al.^
This may be a partial explanation for the degree of particulate
removal achieved with the Horizontal Module.
442
-------
443
-------
444
-------
5.0 DISCUSSION
5.1 Data Acquisition and Reduction
It must be pointed out that getting good particulate
data by sampling in large ductwork with cross-sections of 200 square
feet or more is not easy. Gas sampling crews of five persons work-
ing under ideal conditions were able to obtain one inlet and one
outlet isokinetic particulate sample by probe traverse in one day.
All testing was subject to unit outages and weather conditions
ranging from 70 mph winds in the winter to 130° F temperatures in
the summer. All testing was conducted by probe traverse, and
electrical discharge from the metal probes to ground were observed
at the inlet duct due to electrical charge carried on the inlet
flyash particles.
Grain loading measurements were conducted using an alundum
thimble with a backup millipore filter, and gas volumes were metered
and corrected for water vapor condensation in ice bath condensers.
Particle size distribution measurements were obtained with both
Anderson and University of Washington cascade impactors. Most of
the particle size distribution measurements on the Vertical TCA
Module were conducted with Anderson impactors without backup filters
while similar measurements on the Horizontal Module were conducted
primarily with University of Washington impactors with backup
filters. It was found that physically identical impactors do not
produce the same results from the same sample and must be indivi-
dually calibrated using particles of a known size and instrumenta-
tion to measure the concentration of particles upstream and
downstream from the impactor.
Because of these differences, the particle penetration
data for the Horizontal and Vertical TCA Modules at design operating
conditions is difficult to compare directly. However, it can be
stated with a reasonable degree of confidence that both the
Horizontal and Vertical TCA Modules achieved approximately 6570
collection of the 1.5 microns (aerodynamic) particles at design
operating conditions. An examination of the slopes of figures
3 and 4 would also suggest that the Horizontal Module at design
conditions achieved a higher degree of collection efficiency for
particles larger than 1.5 microns (aerodynamic) than the Vertical
Module at design operating conditions.
The particle size data presented in this paper refers to
aerodynamic particle diameter as measured by a calibrated cascade
impactors. The aerodynamic particle diameter takes into account
particle density and shape factor, and should not be confused with
the actual particle diameter. A curve showing the relationship
between aerodynamic and actual particle diameter for Mohave Genera-
ting Station flyash having a density of 2.2 g/cmj and consisting
primarily of spherical particles is presented in Figure 9. The
term geometric standard deviation is used in the text to identify
the distribution of particle sizes about an average size. For a
log-normal particle size distribution, a geometric standard
445
-------
FIGURE 9 AERODYNAMIC vs, ACTUAL PARTICLE DIAMETER
MOHAVE GENERATING STATION FLYASH
PARTICLE DENSITY: 2,2 GRAMS/CM^
l°1
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446
-------
deviation of two means that the size below which 84 weight percent
of the particles are found is twice the average size. Thus, for
a geometric standard deviation of one, all the particles are the
same size. A brief description of particle size distribution
analysis is included in Appendix 1 for reference.
As another prerequisite to obtaining high quality,
accurate data, detailed grain loading and size distribution studies
were conducted to determine the effects of stratification in the
ductwork. It was found that below a grain loading of about 0.05
gr/SCF, a variation of only + 15% in geometric mean particle size
could be expected due to stratification at various points in the
duct.
5.2 Characteristics of Mohave Flyash
The geometric mean particle size and standard deviation of
the flyash at the scrubber inlet changed as a function of grain load-
ing, as shown in Table 2.
An examination of Table 2 indicates that an increase in
total grain loading by a factor of twenty will cause an increase
in the grain loading of small particles less than 2 microns
(aerodynamic) diameter by a factor of only four. This means that
changes in grain loading are almost entirely determined by changes
in the weight percentage of very large particles. It would thus
be expected that the outlet grain loadings from the Test Modules
would change over a narrow range for a wide range of changes in
inlet grain loading. This has already been shown in Figures 3,
4 and 5.
5.3 Performance Predicted by Inertial Impaction
If interest is restricted to particle diameters more than
about 1.0 micron (aerodynamic), then Ranz and Wong6 have shown that
collection by inertial impaction is by far the most important
collection mechanism. Particle penetration for many types of
control equipment can be expressed as
* (1)
P = exp (-Kdp")* where K is a constant
Calvert 7'^ has indicated that the exponent B equals 2 for most
types of packed towers, sieve plate columns and venturi scrubbers.
If the particle penetration is integrated over the entire size
distribution, the average penetration can be determined
w .w
, „, 2N (dw) (2)
exp (-KdD ) —
- j P ffi - /
o
where w is the particle weight
447
-------
TABLE 2
ESTIMATED FLYASH CHARACTERISTICS
MOHAVE GENERATING STATION
SCRUBBER INLET AVERAGE GEOMETRIC
GRAIN LOADING PARTICLE DIAMETER STANDARD DEVIATION
(GR/SCF) (AERODYNAMIC MICRONS) <3~G
0,02
0,04
0,10
0,40
3,0
4,0
6,0
16,0
2,7
2,8
3,0
4,5
448
-------
Following Calvert, equation (2) can be solved for a log-
normal particle size distribution as a function of the inlet size
distribution parameters and the so-called "cut diameter". The "cut
diameter," dp50, is the particle diameter for which the collection
efficiency of the control hardware is 50%. Therefore, more than
half of the particles larger than the "cut diameter" will be
collected, and conversely, less than half of the particles smaller
than the "cut diameter" will be collected. The "cut diameter"
characterizes the overall particulate removal efficiency of a
control device for known inlet particulate characteristics. The
solution to equation (2) is plotted in Figure 10, and gives overall
penetration as a function of the inlet particle size distribution
characteristics and the "cut diameter" of the particulate control
device.
A method for determining the "cut diameter" of the
Horizontal and yertical TCA Modules is given in Figure 11, where
"cut diameter" is plotted as a function of the percent of net
electrical power required to operate the scrubber. Figure 11 was
prepared from the data in Figures 7 and 8 as well as control room
measurements of electric power requirements over a wide range of
operating conditions. Since the "cut diameter" of a given scrubber
depends on the energy supplied for droplet shattering and liquid/gas
contacting, it is reasonable to expect that the "cut diameter" would
correlate with percent of net electrical power required to operate
the scrubber. A control device using less electrical power to
achieve the same "cut diameter" would be a more efficient device
in terms of particulate removal. The data in Figure 11 includes
penalties for inefficient power consumption, such as pressure drop
losses due to ductwork restrictions, pumping losses due to elevation
differences and less than 100% mechanical efficiency of rotating
equipment.
As shown in Figure 11, a "cut diameter" of a 1.0 microns
(aerodynamic) could be achieved at 450,000 SCFM for an electric
power consumption factor of 1.7% with the Horizontal Module and
2.3% with the Vertical TCA Module. Both scrubbers could achieve
this result, but one is 35% less efficient than the other. When
turndown with the Vertical TCA Module is taken into account, the
electrical power consumption resulting from an increased liquid
flowrate required to achieve the same "cut diameter" increases
significantly.
For the engineer, this method of predicting overall
particulate removal knowing the inlet particulate characteristics,
the type of scrubber and its electric power consumption character-
istics can be a useful design tool.
449
-------
450
-------
451
-------
Example 1
Horizontal Module at 2.6 MW and 450,000 SCFM
Estimated MW capacity at 450,000 SCFM: 170 MW
Power consumption factor = 1.53%
"Cut diameter" from Figure 11: dp50 =1.20 microns
Select inlet grain loading of 0.04 gr/SCF
Inlet particulate characteristics from Table 2
dpg = 4.0 microns and
-------
6.0 CONCLUSIONS
Several conclusions can be drawn from the data presented
in this paper:
1. For the 170 MW scrubbers tested with Mohave Generating
Station flyash, more than 90% total particulate removal
can be achieved with mobile bed, fixed bed and spray
chamber scrubbers operating under design conditions at
450,000 SCFM and an inlet grain loading of 0.1 gr/SCF.
2. The effect of turning down the flue gas flowrate without
changing the liquid flowrate increases the particulate
removal of the Horizontal Module and significantly decreases
the particulate removal of the Vertical TCA and PPA Modules.
This could limit the particulate removal effectiveness of
the Vertical Module configurations during normal load
following operation.
3. The effect of changes in liquid flowrate on particulate
removal is more pronounced with the Horizontal Module than
the Vertical TCA Module.
4. Of all the miscellaneous effects tested with the Vertical
Module, including stages of packing in the PPA configura-
tion and static ball depth in the TCA configuration, only
minor changes in particulate removal were observed.
5. Increases in the number of stages and the nozzle pressure
with the Horizontal Module cause a significant increase in
particulate removal. However, incremental improvements in
particulate removal performance with more stages or higher
nozzle pressures exhibit practical limits according to the
law of diminishing returns.
6. Large increases in grain loading at the scrubber inlet
causes relatively small increases in grain loading at the
scrubber outlet. This is due to high collection efficiency
for large particles. For example, at design operating
conditions the Horizontal and Vertical TCA Modules both
achieve more than 95% removal for particles larger than
6.0 microns (aerodynamic).
7. At design operating conditions, the Vertical TCA Module
requires about 35% more electrical power than the Horizontal
Module to achieve the same degree of particulate removal.
453
-------
7.0 ACKNOWLEDGMENT
The Test Modules Program was a joint venture of the Navajo
and Mohave Power Project participants who are listed below:
Salt River Project Agriculture Improvement and Power District
Arizona Public Service Company
Department of Water and Power of the City of Los Angeles
Nevada Power Company
Tucson Gas and Electric Company
Bureau of Reclamation of the U.S. Department of the Interior
Southern California Edison Company
Funding for this program was provided by the participants in
accordance with their respective megawatt entitlements in the
Novajo and Mohave Power Projects. Southern California Edison
Company was the project manager of the Test Modules Program.
The conclusions presented in this paper represent the personal
opinions of the authors and are not intended to represent the
opinions or position of any of the project participants.
454
-------
8.0 REFERENCES
1. Shapiro, J.L. and Kuo, w.L. "The Mohave/Navajo Pilot
Facility for Sulfur Dioxide Removal," Second (2nd) EPA
Flue Gas Desulfurization Symposium, November 8, 1971,
New Orleans, Louisiana.
2. Weir, A., and Papay, L.T. "Scrubbing Experiments at the
Mohave Generating Station," Third (3rd) EPA Flue Gas
Desulfurization Symposium, May 14, 1973, New Orleans,
Louisiana.
3. Weir, A., Johnson, J.M., Jones, D.G., and Carlisle, S.T.,
"The Horizontal Crossflow Scrubber," Fourth (4th) EPA
Flue Gas Desulfurization Symposium, November 4, 1974,
Atlanta, Georgia.
4. Weir, A., Papay, L.T., Jones D.G., Johnson, J.M., and
Martin, W.C., "Results of the 170 MW Test Modules Program,"
Fifth (5th) EPA Flue Gas Desulfurization Symposium,
March 8, 1976, New Orleans, Louisiana.
5. Calvert, S., Goldschmid, J., Leith, D., and Jhaveri, N.C.,
"Feasibility of Flux Force/Condensation Scrubbing for
Fine Particulate Collection," EPA - 650/2-73-036, EPA
Control Systems Laboratory, October, 1973.
6. Ranz, W.E., and Wong, J.B., "Impaction of Dust and Smoke
Particles," Ind. and Eng. Chem., Vol 44, No. 6, 1952,
p. 1371-1381.
7. Calvert, S., "Engineering Design of Fine Particle Scrubbers,"
EPA/APT Fine Particle Scrubber Symposium, May 28, 1974,
San Diego, California.
8. Calvert, S., Goldshmid, J., and Leith, D., "Scrubber
Performance for Particle Collection," A.I.Ch.E. Symposium
Series 70 (137): 357, 1974.
455
-------
9.0 APPENDIX 1 - PARTICLE SIZE DISTRIBUTION ANALYSIS
Description of Cascade Impactors
Reviewing the theory which underlies how a cascade itn-
pactor works, a typical device is a series of plates con-
taining holes of a given size on each plate. The holes act
as gas jets when a sample is drawn through the cascade, and
the gas jet from the holes in an upstream plate, or stage,
impinge on areas which do not contain holes on the downstream
stage.
The efficiency of particle separation increases as the
particle diameter and gas velocity increase and as jet dia-
meter decreases. Thus one can set up a series (or cascade)
of impinging jets with increasing collection efficiency so
that smaller particles would be collected in succeeding
stages of the cascade. While the separating mechanism is
different, the effect is just the same as using a series of
screens with decreasing opening sizes. One arrangement of a
series of perforated plate and impacting surface pairs as
used in a typical impactor has the collection efficiency
characteristics shown in Figure 9.1.
Jet Orifice
Streamlines
Impact:
Plate
^7" 7 X 1 7 ////// / SI / / / S 7 I?
/• x / >/ / / / / / / / // / / / s / /~\
Trajectory of / Trajectory of Particle
Impacted Particle too Small to Impact
SCHEMATIC OF AN IMPACTOR STAGE
456
-------
Ul
0.1
.2
.3
.4 .5 1.0 2345
AERODYNAMIC PARTICLE DIAMETER, ymA
20
Figure 9.1 Cascade Impactor Stage Efficiencies (for 1 CFM and standard conditions)
-------
Note that the particle diameter used is called
"aerodynamic diameter" and it has the units of "aero-
dynamic microns", AimA. This is the effective diameter for
particle separation by inertial impaction and it takes into
account the effects of particle density and particle "slip"
between gas molecules. The collection efficiency of an
impinging jet depends on the inertial impaction parameter,
Kp, and the distance the jet travels from the orifice to
the impaction plate. For ratios of jet length to diameter
between 1 and 4 there is little difference in collection
efficiency. In that range of jet lengths the experiment-
ally determined cut diameter is given by:
x 104 (1)
and (Kpsq^ = 0.45 (2)
Where:
= diameter of particle collected with 50% efficiency
yum
= inertial impaction parameter for 5070 efficiency
collection of
p = particle density, g/cnH
C' = Cunningham "slip" correction factor
Uft = jet velocity, cm/sec
u = gas viscosity, Poise
dj^ = jet diameter, cm
All of the particle properties can be lumped together
in the aerodynamic particle diameter, which is defined as:
dpa =
The aerodynamic cut diameter, dpc, is therefore given
by:
Sc = M%/!!!h\% x 10* - 0.45 (!^> x 10* W
' VW \% 1
458
-------
If collection efficiency for round jet impaction is
plotted against the ratio of particle aerodynamic dia-
meter to aerodynamic cut diameter", one curve describes
all cases within the usual range of jet length to dia-
meter ratios. The cut diameter is the particle size which
is collected at 5070 efficiency. Figure 9.2 is such a plot
and it describes the shape of each of the individual
curves in Figure 10 The sharpness of fractionation pos-
sible with an impact or is clearly shown in Figure 9,2. Par-
ticle collection efficiencies for diameters other than
the cut size can be found by the use of Figure 9.2. This
relationship as well as the value of the impaction para-
meter for the cut size (equation (2)) has been determined
experimentally.
Effects of Temperature and Pressure
While the effect of gas flow rate is accounted for by
equation (1) , temperature and pressure have effects whose
prediction requires additional information. Temperature
affects both the gas viscosity and the Cunningham factor,
C1 . Pressure has a significant effect on C' but not much
on viscosity. Gas composition has an effect on both
viscosity and C1 but these variations are not large in the
usual air and flue gases encountered.
The situations one may encounter are as follows:
1. Aerodynamic cut diameter as defined at impactor
conditions is affected through effects on gas
viscosity and velocity.
2. Actual particle cut diameter, dp, is computed
from aerodynamic cut diameter with compensation
for particle density, ftp, and C1 at impactor con-
ditions.
3. Aerodynamic cut diameter at conditions other than
those in the impactor can be computed from the
actual diameter with compensation for particle
density at C1 at the new conditions,,
Thus, in order to take temperature, pressure and gas
composition effects into account, it is necessary to have
additional information relating these parameters tongas
viscosity and to the Cunningham factor. Gas viscosity
data can be found in standard chemical engineering and
chemistry handbooks. The Cunningham correction factor is
a complex function of the mean free path of the gas and
the particle diameter,, It increases as temperature increases
and as pressure decreases. For air at standard temperature
and pressure it is given approximately by:
C' = 1 + 0.165 (5)
459
-------
o
ss
w
i— i
CJ
o
h— I
E-
•-a
o
100
90
80
70
60
50
40
30
20
10 —
0
0.8 0.9 1.0 1.
RATIO OF PARTICLE DIA/CUT
1
DIA.
1.2
/a
pa pc
Figure 9.2. Generalized Collection Efficiency -
Size Relationship
460
-------
Illustration
To illustrate the performance of a cascade impactor,
assume that 100 mg of particles are carried into the
impactor by a gas stream at 1 CFM flow rate. Also assume
that the_size distribution is log normal, that the mass
median diameter (dpg) is 1.0/imA and that the geometric
standard deviation ((Tg) is 3.0. The line in Figure 9.3
represents this size distribution on a log-probability
graph.
The performance of each stage of the impactor can be
predicted from Figure 9.1 and the results presented in
Figure 9.4, a plot of mg of particles suspended in the gas
versus particle aerodynamic diameter. The solid line
represents the original size distribution (corresponding
to Figure 9.3)and the dashed lines represent the particle
suspension passing through, or penetrating, the various
impactor stages. The first stage is neglected because its
cut diameter is so high that a negligible amount of parti-
cles would collect there. Stage No. 2 has an 8.1/unA cut
point and it would collect 3% of the particles (i.e., 3 mg
out of 100 mg). Stage No. 3 has a 3.8/nnA cut point and
it would collect about 8.0 mg while allowing 89 mg to
penetrate as the "undersize" fraction,,
The mass of particles collected on a stage is very
nearly the same as would be collected if the stage had a
perfectly sharp size selection. For example, Stage No, 5
has a cut diameter of 0.9/amA and since 46% of the particle
mass is smaller than 0.9^umA, a stage which collects every-
thing larger than the cut size and nothing smaller would
allow 46 mg to penetrate.
A computation of the performance of Stage No, 5 inte-
grated over the example size distribution resulted in the
prediction that about 46.3 mg would penetrate. It has
been shown that the use of cut diameter to characterize
the collection efficiency of an impaction stage will give
results within a few percent of the correct answer so long
as the geometric standard deviation is larger than about 2.0,
The stage cut diameters masses collected, and other
data for this example are shown in Table 9.1, below. Parti-
cles which penetrate the last stage (No. 7) are separated
from the gas by means of a filter with an efficiency of
substantially 100%.
461
-------
MASS PERCENT OVERSIZE
10
10
5
4
1.0
.5
.4
.3
.2
0.1
"^^^^^^^'i!^^^^j^^^^^^^^^§.
:!if±j Solid line is the assumed^
FR size distribution. jrj
H}-|-
Points were computed as <_p;
described in text. ipf
3-LL
!«
ffF
4-4-
12 5 10 20 30 40 50 60 70 80
MASS PERCENT UNDERSIZE
Figure 9.3 Log Probability Plot
90 95 93 99
462
-------
100
bfl
CO
w
f-1
u
PL,
PH
o
CO
CO
2 3 4 5 6 7 8 9 10
PARTICLE DIA^ffiTER, pmA
Figure 9.4 Particle Mass Penetrating Cascade Stages
463
-------
TABLE 9.1
Stage
No.
2
3
4
5
6
7
Filter
TOTAL
Mg on
Stage
3.0
8.0
22.5
20.2
19.3
13.0
14.0
100.0
Stage
Cut Dia.
8.1
3.8
1.6
0.9
0.5
0,3
Cum. %
Oversize
3.0
11.0
33.5
53.7
73.0
86.0
100.0
Cum. %
Undersize
100.0
97.0
89.0
46.3
66.5
27.0
14.0
Data Analysis
Once the stage "catches" have been measured, usually
by weighing particle collection foils or papers, the data
analysis is relatively simple. Generally the objective is
to make a plot of particle diameter versus mass percent
oversize or undersize and to represent the size distribution
in terms of log-normal distribution parameters if possible.
Thus it is necessary to do the following:
1. Add all of the stage and filter collection weights
to get the total particle mass collected.
2. Compute either:
a0 Cumulative percent collected as the gas flows
through succeeding stages. This is "percent
oversize".
b. Cumulative percent penetrating as the gas
flows through succeeding stages. This is
"percent undersize".
3. Compute the cut diameters for the impactor stages,
taking into account gas viscosity (or temperature)
and gas sampling flow rate,,
464
-------
At this point one can plot the data on log-
probability paper or in any other format desired. To
avoid confusion about what cut diameter is associated
with what cumulative mass percentage, bear in mind the
following points:
1. The cumulative mass collected by a stage and all
those preceding it (i.e., through which the gas
flowed before reaching this stage) represents
the particles the same size and larger than the
stage cut diameter.
2. The cumulative mass penetrating a stage and all
those preceding it represents the particles
smaller than the stage cut diameter.
To illustrate, consider the Stage No, 5 given in
Table 9.1. The cumulative mass percent oversize is 53.7%,
which represents the particle size class which is 0.9 AamA
diameter or larger. One can plot this point on Figure 9.3,
which has a percent oversize scale on top of the plot and
a corresponding percent undersize scale on the bottom,,
The point is, as noted before, very close to the assumed
size distribution that was the basis for this illustration.
Data points for the example are plotted on Figure 9.3.
It is obvious that the process of fractionating the
particles through a hypothetical cascade impactor and
using the cut diameter method of data analysis has not
introduced much error. In practice, the weighing errors
and other experimental errors will overshadow any errors
introduced by the cut diameter approximation to define stage
efficiency.
465
-------
EVALUATION OF A PARTICULATE SCRUBBER ON
A FULL SCALE UTILITY BOILER
L. E. Sparks
Industrial Environmental Research Laboratory
Environmental Protection Agency
Research Triangle Park, North Carolina
D. S. Ensor
Meteorology Research, Inc.
Altadena, California
S. Calvert
Air Pollution Technology, Inc.
San Diego, California
D. V. Wallon
K. S. Campbell
Steams-Roger, Inc.
Denver, Colorado
467
-------
EVALUATION OF A PARTICULATE SCRUBBER ON A
FULL SCALE UTILITY BOILER
L. E. Sparks
Industrial Environmental Research Laboratory
Environmental Protection Agency
Research Triangle Park, North Carolina 27711
D. S. Ensor
Meteorology Research, Inc.
Altadena, California 91001
S. Calvert
Air Pollution Technology, Inc.
San Diego, California 92117
D. V. Wallon
K. S. Campbell
Steams-Roger, Inc.
Denver, Colorado 80217
Abstract
The paper presents results of a performance test and engineer-
ing analysis of a mobile-bed scrubber on a full-scale coal-fired
utility boiler. The scrubber nominally operated at the design par-
ticulate removal efficiency of 95 percent, but the concentration of
submicron particles was greatly influenced by mist entrainment.
The entrainment resulted in a. difference of aerosol penetration
through the scrubber as a function of elemental composition and
outlet submicron particle concentration independent of pressure
drop through the scrubber. The engineering analysis showed that
the 1972-installed cost was $29/kw and the annual operating cost
is 0. 5 mills/kwh (75 percent availability). An initial decline in
scrubber availability after start-up resulted from now-corrected
minor design problems. Steadily improving reliability is attrib-
uted to the utility's providing maintenance and solving operating
problems.
Presented at the Symposium. On
Particulate Control in Energy Processes
Sponsored by
Environmental Protection Agency
and
Electric Power Research Institute
San Francisco, California, May 11-13, 1976
468
-------
ACKNOWLEDGEMENT
The efforts of Public Service Company of Colorado personnel
in providing assistance during the source test and the engineer-
ing evaluation were vital to the success of the project. This
work was conducted under EPA Contract 68-02-1802.
469
-------
I. INTRODUCTION
Although scrubbers are commonly used for particulate control in
many industries, they are not commonly used in the utility industry. Be-
cause scrubbers are not sensitive to dust resistivity, they may provide an
acceptable means of controlling high-resistivity fly ash from combustion of
low sulfur coal. In order to assess the potential of scrubbers for particu-
late control in the electric power industry, Industrial Environmental Re-
search Laboratory-Research Triangle Park (IERL-RTP), sponsored an ex-
tensive evaluation of the TCA'1' scrubber at Public Service Company of
Colorado's Cherokee Power Plant. A summary of the performance tests
and engineering analysis of the evaluation is presented in this paper.
Details of the tests and test results are reported by Ensor et al. (1975).
II. DESCRIPTION OF SYSTEM
Power Plant
The Cherokee Power Plant is located in north Denver, Colorado,
near the South Platte River. The #3 unit has a nameplate rating of 150 mw.
The gas cleaning equipment consists of mechanical collectors, an electro-
static precipitator, and a wet scrubber arranged in series.
The unit normally burns coal or natural gas. The coal is mined in
western Colorado and has a sulfur content of about 0. 5 percent, ash con-
tent of 9 percent, and a heating value of 2.492 X 107 J/kg (10,712 BTU
per pound).
Scrubber
The Model 6700 Turbulent Contact Absorber Scrubber, designed by
UOP, Air Correction Division, was installed in 1972. A diagram of the
system, is shown in Figure 1. The flue gas from the precipitator passes
into two parallel induced draft fans. A bypass damper is used to direct
the flue ga-s either into the stack or into the scrubber. The flue gas (under
design conditions), 17,000 AM3/min, at 137 °C (610,000 ACFM at 280 °F),
enters the booster fans to offset the pressure drop through the scrubber.
In the presaturator, 1440 I/rain (380 GPM) of makeup water is sprayed
into the gas to reduce the temperature to approximately 52 °C (125 °F).
- Mention of manufacturers or brand names does not imply endorsement
or recommendation for use by EPA.
470
-------
Mullllelf
Control
ra /
IX*WR*trcam
Sample Point
3 Parallel Uucli
\
Flue G»«
Bollir
Collector Up»tnam
Sample Point
4 Parallel DucU
_L
Ili
-OtclroiUtic
L_
p^^ 1_ l-=b=^
Mill
Blinin.
W»«h
Isolation
/ Damper*
^/ Damper //
S~t=ih \
^ / r
A-r=fc!j ~~L_
i -/^jf-1 L—
alor
q
1
-nvwv-
,-•__• ~_J
Contactlfif
SpKrrea
1
S~ Reaeal 51
/~ MUl Kllm
..
*^
Draft Fan*
ScrubWr Ooo»l*r F*J%*
— Scrubber
Mcclrculallon
Starry DrawoH
Fig. 1. SCRUBBER SYSTEM
-------
From the presaturator, the gas enters the scrubber. The scrubber con-
sists of three stages of fluidized beds packed with 3. 8 cm (1. 5 in. ) diam-
eter plastic balls arranged into three separated parallel scrubber sections.
The two outer sections each handle 20 percent of the flow, while the center
section handles the remaining 60 percent. All three sections can operate
independently to provide flexibility of operation.
The scrubbed gas passes through chevron-type mist eliminators
made of fiberglass reinforced plastic where entrained droplets are re-
moved. The mist eliminators are sprayed once a work-shift from the
top to prevent accumulation of solids.
The gas is then heated by steam coils to 85 °C (185 °F) before enter-
ing the stack to prevent corrosion of the stack and duct work and to provide
plume buoyancy after discharge into the atmosphere. The steam coils are
equipped with two sets of soot blowers to remove fly ash from the heat
transfer surfaces.
The scrubber was designed to the following specifications (Raban,
1974):
Gas Flow 1, 036,.000 AM3/hr, 610,000
ACFM 138°C (280°F)
Liquid-to-Gas Ratio 7.4 //m3 (55 gal/1000 ft3)
AP 30. 5 cm H2O (12 in. H2
-------
techniques using Meteorology Research, Inc. (MRI), cascade impactors
for determining concentrations and size distributions on a mass basis for
particles having diameters between approximately 0. 3 \J.m and 20 urn, and
(3) standard EPA Method 5 mass train measurements for determining out-
let mass loadings. Details of the various instruments and test procedures
are reported by Ensor et al.(1975).
Two Meteorology Research, Inc. (MRI), JPlant Process Visiometers
(PPV) were installed--one at the scrubber inlet and the other at the outlet.
These instruments were used as real time monitors to detect upsets and
to determine in-stack opacity.
Impactor runs were conducted to obtain samples for elemental analy-
sis by Ion-Excited X-Ray Analysis.
All impactor runs were conducted using Apiezon L high vacuum
grease as a particle collection substrate. The Apiezon L was applied to
light-weight collection discs. All weighings were conducted using a Cahn
400 analytical balance with a precision of 0. 01 mg. Use of the Cahn bal-
ance and light weight substrates was essential to reduce errors due to
weighing.
*
The test program was divided into three ph'ases in order to system-
atically evaluate the scrubber. The test phases were:
• Phase 1 - Evaluation of the distribution of aerosol mass
concentration at the inlet. With the limited resources in
manpower and equipment, the evaluation of the source
was attacked one aspect at a time. All four inlet ducts
were sampled at the same time. The cascade impactor
trains were supplemented with in-stack filter probes to
establish the consistency of the concentration.
• Phase 2 - Evaluation of outlet distribution of aerosol mass
concentration. Similar parallel tests were used on the out-
let of the scrubber to indicate the variation in the sections.
• Phase 3 - Simultaneous inlet-outlet tests for scrubber
efficiency. The final phase of the field program was
the simultaneous inlet-outlet tests with both primary
and secondary testa.
The tests disclosed several problems with the scrubber which are
discussed by Ensor et al.(1975).
473
-------
IV. SCRUBBER EFFICIENCY RESULTS
The Phase 3 tests were used to compute the efficiency of the scrubber.
Sufficient data were taken to allow selection of the runs taken during steady-
state plant conditions. The data used in the efficiency determinations are
summarized in Table 1. The diffusion battery tests were restricted to sec-
tion B at the outlet because of the relatively nominal behavior of that section
of the scrubber.
Total Mass Collection Efficiency
The mass collection efficiency was computed using only the cascade
impactor data. The results were summarized along with some control
room data in Table 1. The pressure drop across the system was reported
to indicate the magnitude of the effect of the existence of the air reheaters.
The outlet gas flow was the total for the system as determined from the
velocity traverses. The lack of correlation between efficiency and the
pressure drop was an unexpected result.
The average efficiency of the December tests was 92 percent,
slightly lower than the design efficiency of 95 percent.
Mass Penetration as a Function of Particle Diameter
The particle mass penetration as a function of particle diameter
was determined for three days. These data are shown in Figures 2 to 4.
Generally, the size distributions and penetrations were consistent,
though loadings fluctuated considerably. The maximum penetration (min-
imum efficiency) was achieved in the region of 0. 2 |Jm actual diameter,
similar to results for another scrubber as reported by Sparks et al.(1974).
Elemental Chemical Analysis
The measurement of the concentrations of a. number of elements
was intended mainly to provide data to aid in the interpretation of the
scrubber performance. It was not planned to perform elemental bal-
ances or determine emission factors. The samples were taken during
December 10-11, 1974. The overall penetration of various elements is
shown in Table 2. The penetration as a function of particle diameter for
various elements is shown in Figure 5.
474
-------
Table 1. SCRUBBER COLLECTION EFFICIENCY
DATE
1974
II /ZO
11/2!
IZ/IO1
12/11
12/12
LOAD
mw
166
IU
157
i to
160
°2
Percen!
S.6
1.4
3.4
3.0
2.6
OUTLET
GAS FLOW
ACTUAL.
m /hr
at
a
9.47 x 10*
I0,£ x 10*
8.78 x 10*
SECTION A
AP
"YSTEM
eml! O
41
39
36
38
36
AP
BED
cmH O
9.9
9.6
IS. 2
14.7
14.7
AP
MIST
ELIMIN.
cm!f?O
0.76
0.76
1.7
1.5
1.8
EFF.
%
ND
ND
96. J
96.4
79.6
SECTION B
AP
SYSTEM
cmH2O
45
43
41
4Z
44
AP
BED
cmH <>
25
18
20. 6
22.1
22.9
AP
MIST
RLIMIN.
cmH O
2.5
1.8
2.5
J. 2
2.5
EFF.
%
84.7
89.9
92.6
93.2
93.1
SECTION O
AP
SYSTEM
cmH O
46
44
41
44
46
AP
BED
<;mH20
z4
20
18.5
22.4
24.1
AP
MIST
EUMIN.
cmH O
8. J
5. 1
5.8
2.S
3.8
KFF,
%
ND
NO
86.9
96.7
92, 1
LFull velocity traverses were not taken.
JThe control room data were incomplete. Interviews, data from other days and the log
book were used to supplement available information.
-------
1.0
2
O
J—t
H
U
2
O
H
H
2
H
P.
H
1
>— i
O
i— »
JH
eJ
0.1
Q.01
Q
"" "::;'";" ' '"
V7 DIFFUSION BATTERY RUNS #8 FT #9 <£
V I f-r
jj O IMPACTOR RUNS #110 $ #113
H A IMPACTOR RUNS #109 § #114
••"t
§ D IMPACTOR RUNS #116 § #119
ti
IMPACTOR RUNS #120 5 #118
•!-•'• i • I -t"1 till' ; • i
04 0.07 0. I 0.2 0.5 1.0 2.0
PARTICLE DIAMETER, ^
5.0
Fig. 2. COMBINED PENETRATIONS FOR DIFFUSION BATTERY
AND CASCADE IMPACTOR (December 10, 1974)
476
-------
1.0 \—
2
o
J—1
E-1
U
z
0
C4
E"
W
2
W
u
H-t
H
A DIFFUSION BATTERY RUNS
US
IMP ACTOR RUNS #122 § 1123 $
O IMPACTOR RUNS #124 § #125 »^
IMPACTOR RUNS #127 S, #126
1 • i nr if
0.01
0.02
0.05 0.10 0.3 0.5
PARTICLE DIAMETER,
1.0
3.0 5.0
Fig. 3. COMBINED PENETRATIONS OF DIFFUSION BATTERY AND
CASCADE IMPACTOR (December 11, 1974)
477
-------
1.0
z
o
I—*
E->
U
Z
9. o.i
E-*
t,1
0,
f_T
»—«
i—l
U
t—(
Pi
0. 01
^i^ga ;.-.L
O DIFFUSION BATTERY RUNS #10 § #12
A IMPACTOR RUNS #129 § #133 B 1140
D IMPACTOR RUNS #132 § #135 B1300
O IMPACTOR RUNS #130 5 ^134 A 1113
V IMPACTOR RUNS #131 5 #136 C 1300
o.03 o.os
0.10 0.50.5 1.0
PARTICLE DIAMETER, /irn
3.0 S.O
Fig. 4. COMBINED PENETRATIONS FOR DIFFUSION BATTERY
AND CASCADE IMPACTOR (December 12, 1974)
478
-------
Table 2. PENETRATION OF THE ELEMENTS THROUGH THE
SCRUBBER FOR DECEMBER 10, 1974
Element
Penetration
Average Outlet
Coacentrations
Micrograms/DSm
3 b
Al
Si
S
K
Ca
Ti
V
Cr
Fe
Ni
Cu
Zn
Br
Pb
For All Elements
Total Mass
0.029
0.033
3.4a
0.043
0.059
0.073
0.14
1.10a
0.18
0.95
2.9a
1.5a
0.28
0.64
0.108
0.074
326
658
1030
50
508
96
27
57
1500
33
668
501
5.7
120
Penetrations greater than 1 indicate generation of particles in flue gas
byvevaporation of the scrubber liquor.
5Dry Standard, 21. 1 °C, 760 mm Hg.
479
-------
Z
O
>—I
H
U
O
I—I
H
2
W
10"
10
PARTICLE DIAMETER.
Fig. 5. SCRUBBER PENETRATIONS FOR SELECTED ELEMENTS
480
-------
V.
DISCUSSION
The scrubber performance results were analyzed in depth for trends
and to allow comparison to data in the literature.
Examination of the particle penetration data for each simultaneous
pair of inlet and outlet runs shows a wide range of results, as summarized
in the following table.
TABLE 3. PARTICLE PENETRATION SUMMARY
RUN NO.
IN
56
55
60
114
119
121
125
133
135
124
113
127
134
115
120
123
136
OUT
53
57
61
109
116
123
127
129
132
125
110
126
130
119
118
122
131
SECTION
B
B
B
B
B
B
B
B
B
B
A
A
A
C
C
C
C
d a
P<=,
Jim
1.5
1.1
0.8
0.8
1.0
0.8
< 0.5
0.8
0.8
0.6
1.0
0.8
2.0
< 0.5
0.6
0.7
1.1
AP
cm H3 O
23.0
23.0
23.0
23.0
23.0
23.0
23.0
23.0
23.0
23.0
--
15.0
9.0
25.0
--
25.0
24.0
F. outb
F. in
0.31
0.81
1.2
2.9
0.39
0.55
1.2
2.0
0.47
3.1
0.81
0.41
6.7
0.76
1.0
1.1
3.5
Scrubber cut (50 percent penetration) diameter.
Ratio of outlet filter to inlet filter particle concentrations
(mg/DSm3)/(mg/DSm3).
481
-------
The scrubber performance cut diameter (i. e. , particle diameter
at 50 percent penetration), used to characterize efficiency, varied from
less than 0. 6 jam to 2. 0 [am. No pattern of correlation between cut diam-
eter and other parameters such as pressure drop was found. Thus, the
variation was due to a combination of system fluctuations and measure-
ment errors.
Penetrations found in this study are much higher than those re-
ported in a previous study of the same scrubber (Calvert et al. 1974)
and of another mobile bed scrubber on a coal-fired power plant (Statnick
and Drehmel, 1974). For example, Calvert et al. reported a cut diam-
eter of about 0. 35 |am and a penetration at 1. 0 (am of about 0.09. Statnick
and Drehmel reported that, for 25 cm of water scrubber pressure drop,
penetration did not exceed 0. 05 for any particle size and was about 0. 02
for 1. 0 U m particle diameter. These and some other points are com-
pared in the following table with representative results from the present
study.
TABLE 4. COMPARISON OF MOBILE BED STUDIES
INVESTIGATORS
Present
Calvert et al
(1974)
Statnick and
Drehmel (1974)
V um
0. 8
0. 35
PENETRATION AT PARTICLE
DIAMETER SHOWN
0.5|am 0.8 (am 1.0 jam 2.0|am
0.8
0. 3
0.07
0.5
0. 15
0.03
0.4
0. 08
0.02
0. 15
0.02
0.002
Approximate mean values for all runs, exclusive of 56/53 and 134/130.
There is evidence that the high outlet particle concentration measured
is due to entrainment from the scrubber. Facts which indicate the presence
of outlet particles introduced by entrainment are as follows:
482
-------
1. Outlet filter loadings are higher than inlets, based on
the same gas sample volume, for many of the tests.
2. Penetrations in the present study are higher than those
found by others.
3. Variations in penetration are not related to gas flow
rate, pressure drop, or other known parameters, but
can be attributed to variable entrainment.
4. Penetrations for scrubber section "A" are the same
as for section "B" despite the gas pressure drop for
"A" being about half that for "B".
5. Reheater and entrainment separator operating problems
occurred during the test.
6. Elemental analyses of inlet and outlet particles showed
an apparent "generation" of particles containing soluble
elements as described in the previous section.
The overall conclusion is that the scrubber performance data ob-
tained in this study are specific for the operating factors and scrubber
condition which existed during the test period. It is not possible to
establish a general mobile bed scrubber performance model from these
data because of the overshadowing and undefined influence of liquid en-
trainment. It appears that both scrubber performance and reliability
could be improved by improved entrainment separation.
VI. ENGINEERING ANALYSIS
The objective of the engineering analysis was to assemble the
following information:
• Capital costs
• Operating costs
• Major maintenance problems
• Scrubber reliability
• Estimated cost required to minimize
operating problems.
483
-------
Capital Costs
The total installed cost in 1972 for the scrubber was $4,400, 000.
Based on the boiler nameplate rating of 150 mw, the cost is $29/kw or,
based on the rated gas volume (see Section I), is $4. 18/1000 m3/hr
($7. 10/1000 ACFM). The detailed cost itemization is given by Ensor
et al. (1975). In 1975 dollars, the scrubber would cost $5,800,000.
Operating Costs
The total operating costs are approximately $495, 000/year (fourth
quarter 1973 and the first three quarters of 1974) based on 75 percent
availability of the scrubber or 0.50 mills/kwh.
Maintenance Problems
The scrubber had a number of maintenance problems, many of
which were solved during start-up and operation. The problems men-
tioned here are the more persistent ones which have defied solution.
The maintenance problems are discussed in detail by Ensor et al. (1975).
These include:
• 1.. Breakage of Mobile Bed Contactors - The plastic mobile bed
contactors have been a chronic problem from breakage from
wear. The desired lifetime of the spheres is 8000 hours, how-
ever, normally a lifetime of only 6000 hours was experienced.
However, damage to the pump liners and plugging of the noz-
zles can result from the fragments entering the liquor re-
cycle piping. Screens in the scrubber hopper have eliminated
this problem with the penalty of adding screen cleaning to the
required maintenance. Public Service Company of Colorado
has tested a number of different packings and is currently us-
ing polyethylene spheres in the scrubber.
2. Migration of Mobile Bed Contactors - The mobile bed contrac-
tors will also migrate from one section of the scrubber to an-
other if an opening the width of a contactor exists in the parti-
tions. Poor distribution of the contactors causes channeling
of flue gas and a reduction in particle collection efficiency.
3. Guillotine Dampers - The guillotine isolation dampers have
caused problems due to breakage of the damper when closing
against a buildup of fly ash and leakage.
484
-------
4- Recirculation Pumps - The recirculation pumps have been a
source of problems in the past due to mechanical failure. A
new pump has been tested by Public Service Company of Colo-
rado and has been providing good service.
5« Reheater Section - The scrubbed gases are heated by direct
contact with three banks of steam coils. These coils are sus-
ceptible to pluggage and corrosion in the wet flue gas scrubber
discharge. Addition of a second set of soot blowers and drying
ash during periods of the scrubber shutdown with heavy soot
blowing has had minimal effect. Corrosion problems have
rendered the heaters inoperable. Both upper and lower layer
reheat coils have been removed.
6. Weather-Related Problems - The freezing of lines has always
been a problem during cold weather. All lines must be heat-
traced and drained when the scrubber is shut down.
Another solution is to enclose the scrubber in a weatherproof
building. The major problem with this is the possibility of
leakage of flue gas into the structure.
Scrubber Reliability
For this study, availability of the scrubber was defined as:
hours of scrubber operation*- hours boiler was burning 100% gas
Avai ability - hourg Q£ boiler operation - hours boiler was burning 100% gas
The reason that the hours the boiler wag burning 100 percent gas was sub-
tracted from the total hours of scrubber and boiler operation is that the
scrubber is normally either left running with only water systems operating
or completely shut down during periods of 100 percent gas burning opera-
tion
It is interesting to note that the scrubber was out of service a majority
of the time during August and September 1974, when the unit returned to
mixed fuel burning. The reason for this is that after long periods of idle-
ness, due to either maintenance or 100 percent gas burning, the scrubber
*
- Hours of scrubber operation is related directly to the time the
boiler was operating.
485
-------
usually experienced minor start-up problems. When start-up troubles
with the scrubber were encountered, every effort was made to use any
available gas on the unit. Consequently the scrubber had a poor per-
centage availability for the months involved, but the overall availability
was not affected very much.
Values of availability a.re given for two capacity levels: 100 per-
cent of capacity and 80 percent of capacity or greater. There is a sig-
nificant amount of time that the scrubber operated at 80 percent of ca-
pacity. For example, in January of 1974, although the scrubber only
operated at 100 percent of capacity for 175 hours, it did operate at 80
percent of capacity for 547 hours. Although the scrubber is not treat-
ing all of the boiler flue gas, it is felt that some credit should be given
the scrubber for handling most of the flue gas. In most cases, running
at 80 percent of capacity will probably allow PSCC to meet particulate
emission standards.
The accumulative percent availability to date reflects the total
hours of scrubber operation since start-up. This value is plotted by
month in Figure 6. This figure shows that after an initial drop, the
availability of the scrubber has been increasing steadily. This is
typical of other PSCC scrubber operations. The initial drop is attrib-
uted to the many minor design problems which arose soon after start-
up. The following increase in availability is attributed to the solving
of these problems and the increased operating knowledge which was
gained through operating experience.
As of November 30, 1974, the scrubber had operated at 100 per-
cent capacity 59. 9 percent of the time and at 80 percent capacity or
greater 70. 9 percent of the time.
In order to more clearly see the increase in availability with time,
after the first year when each new month was added to the log, only the
previous 12 months availability was calculated. This- value is plotted in
Figure 7. Looking at the last month tabulated (November 1974), reveals
that the scrubber operated at 100 percent of capacity 78.6 percent of the
time and at 80 percent of capacity or greater 96. 7 percent of the time
during the previous year. The figure also shows that the scrubber had
been operating at 80 percent of capacity or greater with over 90 percent
availability since May of 1973.
486
-------
>t>
-p-
00
2
I
10
I X
* a
wn
5 1 S S I 4 1 I 5 I 8 I t 1 I I 5 * 3 i 5 i
1«74
MONTH
Fig. 6. CHEROKEE //3 SCRUBBER ACCUMULATIVE AVAILABILITY FROM START-UP TO
NOVEMBER 1974.
-------
00
00
>0
3 •»
i
= I
—r~
st
I 8
I
I
8
UOMTM
Fig. 7. CHEROKEE tf3 SCRUBBER ACCUMULATIVE AVAII-ABIUTY FROM START-UP TO
NOVEMBER 1974 AVERAGED OVER THE PREVIOUS YEAR ON A RUNNING BASIS.
-------
The feasibility of using the design and operating experience to im-
prove availability of the scrubber was investigated. The goal was to in-
crease the availability of the scrubber, realizing that the maintenance
required may slightly increase due to the extra equipment. The suggested
modifications are summarized below:
Identical Scrubber
Extra 33-1/3 Percent
Capacity Section
Indirect Reheat, Incre-
mental
Scrubber Enclosure
Miscellaneous Charges
$5,800,000'
1,200,000
200,000
90, 000
80,000
$7,370,000
Estimated Availability
(Percent)
60-70
10-15
5-10
75-95
The capital cost of $7, 370, 000 is $49/kw for a 150 mw unit.
The important aspect of maintaining and improving the availability
of the scrubber system is the attitude of the plant operating and mainten-
ance people in keeping the unit on line.
The steadily increasing availability of the Cherokee scrubber is an
indication that the Public Service of Colorado is committed to solving the
operating and maintenance problems.
VII. CONCLUSIONS
The performance data and engineering analysis indicate that scrub-
bers are capable of collecting fly ash from combustion of low sulfur coals.
Capital and operating costs for a scrubber are likely to be higher than
similar costs for a moderately sized electrostatic precipitator. However,
in applications where the dust resisitivity requires a large-sized electro-
static precipitator, a scrubber system is likely to be cost competitive.
* - 1975 dollars.
489
-------
REFERENCES
1. Calvert, S., N. C. Jhaveri, C. Yung: "Fine Particle Scrubber Perform-
ance Tests, " EPA-650/2-74-093. (NTS PB 240-325/AS), October 1975.
2. Ensor, D. S. , B. S. Jackson, S. Calvert, C. Lake, D. V. Wallon, R. E.
Nilan, K. S. Campbell, T. A. Cahill and R. G. Flocchini: "Evaluation of
a Particulate Scrubber on a Coal-Fired Utility Boiler, " EPA-600/2-75-074.
(NTB PB 249-562/AS), November 1975.
3. Raben, I. A. : "Use of Scrubbers for Control of Emissions from Power
Boilers "U.S. Paper No. 13. Proceedings Symposium on Control of Fine
Particulate Emissions from Industrial Sources. San Francisco, California.
January 1974.
4. Sparks, L. E. , J. D. McCain and W. B. Smith (1974): "Performance of a
Steam-Ejector Scrubber", J. Air Poll. Control Assoc. , 24, 958.
5. Statnick, R. M. and D. C. Drehmel: "Fine Particle Control Using Sulfur
Oxide Scrubbers". 67th Meeting of the Air Pollution Control Assoc., Den-
ver, Colorado, Paper No. 74-231. June 1974.
490
-------
COLLECTION MECHANISMS AT HIGH TEMPERATURE AND PRESSURE
Seymour Calvert
and
Richard Parker
Air Pollution Technology, Inc.
4901 Morena Blvd., Suite 402
San Diego, California
491
-------
COLLECTION MECHANISMS AT HIGH TEMPERATURE AND PRESSURE
By Seymour Calvert and Richard Parker
Air Pollution Technology, Inc.
4901 Morena Blvd., Suite 402
San Diego, California 92117
ABSTRACT
The paper reports the results of a theoretical study
of particle collection mechanisms at high temperature (1,100°C)
and high pressure (15 a tin) . The effects of high temperature
and pressure are presented for various collection methanisms ,
including: inertial impaction, Brownian diffusion, electrical
migration, thermophoresis, diffusiophoresis, centrifugal forces,
gravitational forces, and particle agglomeration. Some examples
of the effects of temperature and pressure on the performance of
particulate control devices are also presented.
In general, high pressure reduces the effectiveness of par-
ticle collection mechanisms. For a given pressure, high tempera-
ture further reduces the effectiveness of collecting large parti-
cles, but improves the collection of particles smaller than about
0.1 to 0.5 ym in diameter. The beneficial effects of temperature
are much more pronounced at lower pressures (near atmospheric).
492
-------
COLLECTION MECHANISMS AT HIGH TEMPERATURE AND PRESSURE
INTRODUCTION
The removal of particles from high temperature and pressure
gas streams is a problem which is encountered in many existing
and proposed energy processes. One important example is the re-
moval of particles from low-BTU coal gas before it is combusted
and used to drive a combined-cycle gas turbine/steam turbine power
generation system. The particles must be removed to protect the
turbine blades as well as to satsify the particulate emissions
standards, and must be removed with a minimum loss of pressure
and temperature. High temperature particulate removal is also
of interest in other processes, such as some metallurgical pro-
cesses, where heat recovery is not economical and expensive cool-
ing equipment is required to reduce the effluent gas temperature
to a temperature suitable for conventional particulate control.
The effects of high temperature and pressure on particle collec-
tion have been studied by many authors including Thring and
Strauss , Strauss and Lancaster^, Calvert, et al.3, and Rao, et
al.4. The present study represents a review and extension of the
material presented by previous authors.
All types of devices used for particle collection can be
characterized by a deposition velocity, "v^", which can be re-
lated to the particle collection efficiency, "E", through the
penetration, "Pt", as follows:
(Vj A
- -*—) (1)
where
A, = deposition area, cm2
Qr = volumetric flow rate, cm3/s
493
-------
The particle penetration is defined as:
p _ Outlet concentration ' , _-.
Inlet concentration
The deposition velocity for any collection mechanism depends
on the force balance between the driving force (deposition force)
and the resistance force of the gas. As will be shown later, the
fluid resistance force is generally affected by high temperature
and pressure much more than the driving force.
Table 1 lists the most important particle collection mechan-
isms for high temperature and pressure collection. The effects of
temperature and pressure on these mechanisms are presented below.
Some examples of the effects of temperature and pressure on the
performance of specific devices are also presented.
Table 1. IMPORTANT COLLECTION MECHANISMS AT
HIGH TEMPERATURE AND PRESSURE
1. Inertial impaction
2. Brownian diffusion
3. Electrical migration
4. Gravitational settling
5. Centrifugal separation
6. Thermophoresis
7. Diffusiophoresis
8. Particle size alteration
a. thermal coagulation
b. turbulent coagulation
c. sonic agglomeration
FLUID RESISTANCE FORCE
The major difference between the collection of particles at
normal conditions and at high temperature and pressure is in the
494
-------
fluid resistance force. The fluid resistance force is generally
approximated by Stokes' law modified to allow for non-continuum
slip flow effects.
uo
where
Fr ~ fluid resistance force, dynes
Pg = fluid dynamic viscosity, g/cm-s
d = particle diameter, cm
UQ = relative velocity between the particle and the gas, cm/s
C1 = Cunningham slip correction factor, dimensionless
The temperature and pressure dependence of equation (1) is
contained in the terms "Pr" and "C"'. The viscosity of a gas
increases iv'ith increasing temperature and at extreme pres-
sures, viscosity increases slightly with pressure. This effect is
not significant at pressures below about 20 atm.
The Cunningham slip correction factor may be calculated as:
c. = i + 2A. 1>257 + 0.40 exp (-1.1 d /2A)j (4)
dp [ P J
where
X = mean free path of gas molecules, cm
The Cunningham slip correction factor is a function
of temperature, pressure, and particle diameter, and is only
important for small particles, high temperatures, and low
pressures.
The particle deposition velocity for most collection mechan-
isms of interest is inversely proportional to the fluid resistance
force, and therefore proportional to the ratio "C'/yG". The effects
495
-------
of high temperature and pressure on this ratio, plotted as a func-
tion of particle diameter, are illustrated in Figure 1. At atmos-
pheric pressure, the ratio decreases with increasing temperature
for particles larger than about 0.6 ym. At 15 atm pressure, the
ratio decreases with temperature for all particle sizes considered.
Therefore, the particle deposition velocity will generally be
smaller at high temperature and pressure than at normal conditions.
COLLECTION MECHANISMS
Inertial Impaction
One of the most important mechanisms for collecting particles
greater than a few tenths of a micrometer in diameter is inertial
impaction. Inertial impaction is the collection of moving particles
by impinging them on some target. The relative effect of inertial
impaction for different particles and flow conditions may be charac-
terized by the inertial impaction parameter, "K ", defined as:
C' p d 2 u
K = 3L_J ° (5)
where
p = particle density, g/cm3
d = characteristic length for collector, cm
The inertial impaction parameter is equivalent to the ratio
of the particle stopping distance, "x ", to "d ". The particle
stopping distance is that distance the particle would travel before
coming to rest if injected into a still gas at a velocity "u ",
when only the fluid resistance force acts on the particle. By
considering the particle stopping distance divided by "UQ"> the
particle's inertia can be characterized by a relaxation time,
"T", defined as:
496
-------
NO. CONDITIONS
20°C, 1 atm
500°G, 1 atm
1,100°C, 1 atm
20°C, 15 atm
500bC, 15 atm
1,100°C, 15 atm
0.1
0.5 1.0 5.0
PARTICLE DIAMETER x, 10", cm
10.0
Figure 1. The effect of temperature and pressure on the ratio
of the Cunningham slip correction factor to the dy-
namic viscosity of air.
-------
x K d
T = _JL = _J1_
uo Uo 9 ^G
Figure 2 shows the effects of high temperature and pressure
on the particle relaxation time plotted as a function of the com-
pensated particle diameter (particle diameter times the square
root of the particle density). For large particles, inertial
impaction decreases with increasing temperature and pressure.
For small particles, high temperature and atmospheric pressure
curve 2 indicates that inertial impaction begins to improve at very
high temperature. High pressure effectively nullifies any benefi-
cial effect of high temperature" (curve 3) .
Brownian Diffusion
Small particles can undergo significant Brownian motion resul-
ting from the random bombardment of the particle by gas molecules.
The rate of diffusion is characterized by the particle diffusivity,
defined as :
C! k T
D = - (7)
Sir pG dp
where
D = particle diffusivity, cm2/s
k = Boltzman's constant, erg/°K
T = absolute temperature, °K
Figure 3 shows the effects of temperature and pressure on
particle diffusivity as a function of compensated particle diameter.
Smaller particles undergo higher rates of diffusion. High tempera-
ture increases the diffusivity for all particle sizes, however,
this effect is greater for smaller particles. High pressure
decreases the beneficial effect of temperature. The combined
effect of high pressure and high temperature is a net increase in
diffusivity.
498
-------
0°C, 1 atm
1,100°C, 1 atm
1,100°C, 15 atm
PARTICLE RELAXATION TIME, s
Figure 2. The effect of high temperature and pressure on particle inertia.
-------
7.0
Ul
o
o
0°C, I atm
1,100°C, 1 atm
1,100°C, 15 atm
0.3
10
- 7
10
- 6
PARTICLE DIFFUSIVITY, cm2/s
Figure 3. Particle diffusivity at high temperature and pressure.
-------
The diffusional deposition velocity can be shown to be pro-
portional to the square root of the diffusivity. That is,
UBD
Using equation (1) , the penetration can be related to diffusivity
as:
In Pt~
- !
£n PtQ
where
Pt-p p = penetration at temperature, T, and Pressure, P
Pt = penetration at standard conditions
DT p = diffusivity at "T" and "P", cm2/s
D^ = diffusivity at standard conditions, cm2/s
Figure 4 is a plot of the diffusivity ratio, "D- p/D ",
as a function of penetration for three assumed penetrations at
standard conditions. For a given temperature and pressure, the
diffusivity ratio can be obtained from Figure 3 and the effect of
this diffusivity ratio on the penetration can be obtained from
Figure 4 .
Electrical Migration
The migration of a particle in an electric potential field
is proportional to the field strength, the particle charge, and
the fluid resistance force. Electrical migration is generally
characterized by a deposition velocity (migration velocity) de-
fined as :
u =
6 3* * d
(io7) (10)
p
501
-------
0.1
z:
o
i—i
? .01
0.001
0 1 2 3 4 5 (
DIFFUSIVITY RATIO
Figure 4. Penetration as a function of particle diffusivity
relative to standard conditions.
502
-------
where
u = migration velocity, cm/s
\5
q = particle charge, coulomb
E = electric field strength, volts/cm
For a given field strength and particle charge, the effects
of temperature and pressure are contained in the ratio "C'/y"
and are illustrated in Figure 1. The particle charge and electric
field strength are complicated functions of temperature and pres-
sure, and are beyond the scope of this study.
Gravitational Settling and Centrifugal Separation
Using equation (3) to describe the fluid resistance force,
the gravitational settling velocity and the deposition velocity
of a particle in a centrifugal force field may be approximated as:
_ __
18 M
1 C' d 2 (Pr, - pr) u.2
-*- Tj 13 U L /" "1 O "S
u = — * * (12)
18 Uo R
where
u = gravitational settling velocity, cm/s
u = centrifugal force deposition velocity, cm/s
g = acceleration of gravity, cm/s2
u = tangential particle velocity at radius "R", cm/s
R = radial position of particle, cm
p~ = density of the gas, g/cm3
In general, even at relatively high pressures (<~50 atm),
the gas density is much smaller than the particle density and may
be neglected in equations (11) and (12). Therefore, the tempera-
ture and pressure dependence of equations (11) and (12) is con-
tained in the ratio "C'/yG" and is illustrated in Figure 1.
503
-------
Thermophoresis and Diffusiophoresis
Temperature and concentration gradients can give rise to
deposition forces which can improve the collection efficiency of
particulate control devices.
Thermophoresis is the result of gas molecules impinging on
the particle surface from opposite sides with different mean
velocities. The particle receives a net impulse opposite to the
temperature gradient in the gas. The magnitude of the thermo-
phoretic force was first devised by Epstein , and may be used with
equation (1) for the resistance force, to obtain the thermophore-
tic deposition velocity as:
3 C'yG
2 PG T
kG
2kG + kp
ud = ^ 2 VT (13)
where
kp = gas thermal conductivity, J/s-m-°K
k = particle thermal conductivity, J/s-m-°K
VT = temperature gradient, °K/cm
Uj = thermophoretic deposition velocity, cm/s
Figure 5 shows the effects of temperature and pressure on
the thermophoretic deposition velocity per unit temperature gra-
dient, for a 5 ym diameter silica particle, as calculated from equa-
tion (13). The deposition velocity increases slightly with increasing
temperature but decreases significantly with pressure.
More elaborate equations for predicting the thermophoretic
deposition velocity have been presented by Hidy and Brock and
7
Derjaquin and Yalamov . These equations are generally a function
of the thermal-accomodation coefficient between the molecule and
the particle surface. The accomodation coefficient has been shown
o
to be a strong function of temperature (Byers and Calvert ), at
relatively low temperatures. No information has been found con-
504
-------
5-1
z
w
H
Q
<
&
C3
§
W
CU
W
H
M 6
o, 6
u
CJ
o
K-3
W
2
O
o
cu
w
n
w
i
H
Particle diameter = 5x10
Particle thermal conductivity
= 1.0 J/s-m-°K
(0.6 BTU/min-ft-°F
3 15 atm =s=-:~:H£EE:
io-
200
400 600 800
AIR TEMPERATURE,. °C
1,000
1,200
Figure 5. Thermal deposition velocity as a function of
temperature and pressure
505
-------
cerning the effects of very high temperatures and pressures on
the accomodation coefficient. This lack of information repre- •
sents a major uncertainty in the extension of the theory of ther-
mophoresis to high temperature and pressure conditions.
Diffusiophoresis is generally important in situations where
there are large concentration gradients, or where vapor condensa-
tion is occurring. It is unlikely that such conditions will exist
in high temperature and pressure particulate removal systems. How-
ever, diffusiophoresis may be important in situations where high
pressure and low or moderate temperature particle collection is
required.
Calvert, et al. looked at the problem of diffusiophoresis
at high pressure and relatively low temperature. Figure 6 (from
Calvert el al. ) shows the diffusiophoretic velocity as a function
of vapor pressure gradient for various temperature and pressure
conditions. The diffusiophoretic velocity is greatly reduced at
high pressures.
Particle Size Alteration
One way to improve the collection efficiency for fine par-
ticles is to cause the fine particles to agglomerate into large
aggregates which can be collected more easily.
Thermal coagulation - Thermal coagulation is the agglomera-
tion of particles undergoing random Brownian motion. The rate of
agglomeration (or coagulation) is generally considered to be pro-
portional to the square of the particle number concentration.
That is,
= - K N2 (14)
dt
where
- 2
N = particle number concentration, cm
t = time, s
K = coagulation constant, cm3/s
506
-------
s
o
CJ
o
_J
w
U
Pi
o
(X
o
I— I
C/D
1=3
10
-1
10
- 2
10
- 3
0°C, 1 atm
25°C, 1 atm
50°.C, 1 atm
25°C, 10 atm ^
i 102 103 5
VAPOR PRESSURE GRADIENT, millibar/cm
Figure 6. The effect of high temperature and moderate
pressures on the diffusiophoresis of particles
507
-------
Q
Using equation (3) for the fluid resistance, Fuchs pre-
sents the following equation for the thermal coagulation constant.
A v T r '
KQ = 4iT D d = 1 K i L (15)
The thermal coagulation constant is shown as a function of
temperature, pressure, and particle diameter in Figure 7. The
agglomeration of particles increases with temperature and decreases
with pressure. It appears that at high pressure and high tempera-
ture, there is a small increase in the rate of agglomeration for
particles larger than 0.1 pm. At high temperature and atmospheric
pressure, the rate of agglomeration of fine particles^ would be
increased significantly, however, high pressures greatly
decrease the benefit of high temperature.
Turbulent Coagulation - Particles can also agglomerate as a
result of turbulence in the fluid. Turbulent coagulation has been
discussed by Beal and is proportional to the turbulent diffusion
coefficient. For the general case where the particle diameter
is much smaller than the turbulent microscale, Beal presents
the following equation for the turbulent diffusion coefficient.
(16)
where
D = turbulent diffusion coefficient, cm2/s
E = energy dissipation rate for unit mass of fluid, cm2/s3
v = kinematic viscosity, cm2/s
r = particle radius, cm
Therefore, "Dt" is proportional to "v"1/2". Figure 8 is a
plot of "v v 2 " against temperature for a range of pressure. Tur-
bulent agglomeration increases greatly with an increase in pressure
508
-------
Particle
Diameter
200
400 600 800
AIR TEMPERATURE, °C
1,000
,200
Figure 7. Thermal coagulation of particles at high temperature
and pressure.
509
-------
s
u
t/1
o
u
t/1
•H
U
•H
CtJ
200 400 600 800
AIR TEMPERATURE, °C
1,000 1,200
Figure 8. Turbulent agglomeration of particles
at high temperature and pressure.
510
-------
at low temperatures. Unfortunately this beneficial effect of
pressure is almost completely nullified at high temperatures.
Sonic agglomeration - Another way to cause particles to
agglomerate is by the application of sonic vibrations. Sonic
agglomeration has been studied by Mednikov and has been shown
to be inversely proportional to the square root of the product
of gas density and the speed of sound. That is,
(17)
where
K - sonic agglomeration coefficient
H.
C,-, = speed of sound in the gas
The speed of sound of an ideal gas is given by:
CG = (YRT)1/2 (18)
where
Y = the ratio of specific heats, dimensionless
R = gas constant, erg/g-°K
T = absolute temperature, °K
Figure 9 shows the relative sonic agglomeration coefficient for
a variety of temperatures and pressures. Once again, high tempera-
tures can slightly improve agglomeration at atmospheric pressure,
but high temperature and high pressure together significantly
reduce sonic agglomeration relative to standard conditions.
511
-------
C/}
l/>
CD
o
•H
t/1
£
-------
EXAMPLES
Single Stage Impactor
Using equation (3) to represent the fluid resistance force,
it is possible to predict the collection efficiency and work re-
quirements for particle collection by inertial impaction. Figure
10 shows the particle cut diameter as a function of flow work (or
specific power) for a single stage impactor at various temperature
and pressure conditions. For an impactor, it may be assumed that
all particles larger than the cut diameter will be collected.
The flow work is equivalent to the energy requirement per mass
of gas. The power requirement tvould be equal to the flow work
multiplied by the mass flow rate.
Figure 10 illustrates that the work required to collect
submicron particles by inertial impaction increases rapidly with
decreasing particle diameter. The effects of temperature and
pressure may be seen by comparing the four curves. Curve 1 re-
presents standard temperature and .pressure conditions. Curve 2
shows the effect of high temperature at atmospheric pressure. The
work required to collect submicron particles is not increasing as
rapidly as that required to collect larger particles. The reason
for this is that the fluid resistance force decreases as the mean
free path of the gas molecules increases relative to the particle
diameter, and the mean free path increases with temperature, for
a constant pressure.
Curves 3 and 4 are for simultaneous high pressure and high
temperature conditions. The beneficial effect of high temperature
on the mean free path is completely nullified by a decrease in
mean free path with increasing pressure. The work required to
remove submicron particles is greatly increased. For example,
the collection of all particles greater than 1 ym by inertial im-
paction would require about 450 J/kg at standard conditions (Curve
1) . It would require about 4,000 J/kg at 1,100°C and 15 a tin,
(Curve 4). This is equivalent to approximately a 9:1 increase
513
-------
0.01
SPECIFIC POWER, HP/MSCFM
0.1 1.0
P!
u
X
10 H
5 t
i.o
0.5 ..
0.1
density =1.0 g/cm3
NO. CONDITIONS
1 20°C, 1 atm
2 800°C, 1 atm
4tt 3 800°C, 10 atm
4 1,100°C, 15 atm
10
103
FLOW WORK, J/kg
10'
10
Figure 10
Flow work (specific power) for impaction from a round jet as
a function of temperature and pressure.
-------
in the power requirement to maintain a similar degree of particle
removal.
Cyclone Separator Efficiency
A further illustration of the effect of high temperature and
pressure on particle collection may be obtained by predicting the
collection efficiency of a cyclone separator operating at various
temperatures and pressure, for the same inlet velocity.
Figure 11 shows a typical efficiency curve for a high effi-
ciency cyclone (curve 1). The cyclone cut point (50% efficiency)
occurs at a particle diameter of about 1 ym, and the cyclone is
better than 99% efficient for particles larger than 15 ym. This
curve was obtained using the theory presented by Calvert, et al3
and attributed to Leith and Licht . The cyclone was assumed to
be about 15 cm (6 inches) in diameter with a volumetric flow rate
of about 1,400 Jl/min (50 f t3/min) .
Curve 2 shows the estimated cyclone efficiency for a gas at
1,100°C and atmospheric pressure, and for the same inlet gas velo-
city as in curve 1. The cyclone efficiency has dropped signifi-
cantly and now has a cut point occurring at 2.0 ym and is only
96% efficient for 15 ym particles.
Curve 3 shows the estimated cyclone efficiency for a gas
at 1,100°C and 15 atm, for the same inlet velocity. The cyclone
efficiency has decreased again slightly for small particles, but
is relatively unaffected for larger particles. The cut point now
occurs at 2.5 ym and the efficiency for 15 ym particles is 95%.
Specific Power for a Cyclone Separator
Figure 12 presents another example of high temperature and
pressure effects on the performance of a cyclone separator. The
specific power ratio is the ratio of the specific power (HP/MSCF)
at high temperature and pressure to that at standard conditions.
The curves in Figure 12 show the power requirement relative to
standard conditions to collect various particle sizes while main-
taining a constant collection efficiency for each particle size.
515
-------
100
20°C, 1 atm
1,100°C, 1 atm
1,100°C, 15 atm
2 4 6 8 10 12 14 16 18 20 22 24 26 28 30
PARTICLE DIAMETER x lO1*, cm
Figure 11.
The effects of high temperature and pressure on
the collection efficiency of a high efficiency
cyclone.
-------
500
1,000
2,000
AIR TEMPERATURE, K
Figure 12, The specific power requirements
for a cyclone as a function of
temperature and pressure.
517
-------
Efficiency of a Fiber Bed
Figure 13 shows the effects of temperature and pressure on
the collection efficiency of a fiber bed. The curves were cal-
culated from the theory presented by Calvert, et al. , and attri-
buted to Jorgeson. The collection efficiency is the combined effi-
ciency resulting from inertial impaction, interception, and Brown-
ian diffusion.
From Figure 13 it is apparent that the filter collection effi-
ciency for particles larger than about 0.5 ym in diameter is re-
duced significantly at high temperature and pressure. This is
a result of the larger inertial impaction parameter (equation 2)
at high temperature and pressure. For particles smaller than 0.5
pm, the collection efficiency is somewhat increased because of the
increased Brownian motion at high temperatures.
At high temperature and atmospheric pressure the collection
efficiency is greatly increased for particle diameters up to about
0.9 ym. For particles larger than 1 pm the collection efficiency
is reduced at high temperature and atmospheric pressure, although
not as severely as at high temperature and high pressure.
CONCLUSIONS
In general the particle collection efficiency is significantly
reduced at high temperature and pressure for particles with dia-
meters larger than a few tenths of a micrometer. Particle collec-
tion at high temperature and pressure is likely to be very diffi-
cult, especially for fine particles.
Some mechanisms seem to increase significantly at high tempera-
ture and low pressure (for example, Brownian diffusion). In this
case, it is likely that collection of submicron particles would be
easier at higher temperature. However, the collection of larger
particles would still be more difficult.
It is unlikely that a new device will be devised which can
remove particles from high temperature and pressure gases more
efficiently than at standard conditions. It is more probable that
518
-------
CJ
S;
w
I— I
u
I— I
Hi
PH
w
z
o
h- 1
H
U
H-3
O
U
100
90
80
70
60
50
40
30
20
10 -
Constant face velocity
It
1
2
3
20°C, 1 atm
1,100°C, 1 atm f
1,100°C, 15 atm \.
0.1
Figure 13
0.5 1.0
PARTICLE DIAMETER' x 101*, cm
5.0
The effects of high temperature and pressure on
the collection efficiency of a fiber bed.
-------
high temperature and pressure particle cleanup will have to be
achieved by modified conventional equipment (cyclones, metal or
ceramic filters, granular beds) operating at higher costs (larger
power consumption) than would be needed for operation at standard
conditions.
520
-------
REFERENCES
1. Thring, M. W. , and W. Strauss, "The Effect of High Temperatures
on Particle Collection Mechanisms," Trans. Instn. Chem.
Engrs. . Vol. 41, p. 248, 1963.
2. Strauss, W., and B. W. Lancaster, "Prediction of Effectiveness
of Gas Cleaning Methods at High Temperatures and Pres-
sures," Atmospheric Environment. Vol. 2, p. 135, 1968.
3. Calvert, S., J. Goldshmid, D. Leith, and D. Mehta, "Scrubber
Handbook," A.P.T., Inc. report to the E.P.A., July, 1972.
4. Rao, A. K., M. P. Schrag, and L. J. Shannon, "Particulate
Removal from Gas Streams at High Temperature/High Pres-
sure," M.R.I, report to the E.P.A., EPA #600/2-75-020,
August, 1975.
5. Epstein, P. S., "Zur Theorie des Radiometers," Z.~ Phys.,
Vol. 54, p. 537, 1929.
6. Hidy, G. M., and J. R. Brock, "The Dynamics of Aerocolloidal
Systems," in Int'l Rev. Aerosol Phys. Chem., Vol. 1,
Pergamon Press, N.Y., 1970.
7. Derjaguin, B. V. and Yu. I. Yalamov, "The Theory of Thermo-
phoresis and Diffusiophoresis of Aerosol Particles,"
in Int'l Rev. Aerosol Phys. Chem., Vol. 5, ed. G. Hidy
and J. Brock, Pergamon Press, N.Y., 1972.
8. Byers, R. L., and S. Calvert, "Particle Deposition from Tur-
bulent Streams by Means of Thermal Force," I. and E.G.
Fundamentals, Vol. 8, No. 4, p. 647, 1969.
9. Fuchs, N. A., The Mechanics of Aerosols, Pergamon Press, N.Y.,
1964.
10. Beal, S. K. , "Turbulent Agglomeration of Suspensions," Aerosol
Science, Vol. 3, p. 113, 1972.
11. Mednikov, E. P., Acoustic Coagulation and Precipitation of
Aerosols, Consultant Bureau, N. YY, 1965.
12. Leith, D., and W. Licht, "The Collection Efficiency of Cyclone
Type Particle Collectors - A New Theoretical Approach,"
paper presented at A.I.Ch.E. meeting in S.F., December,
1971.
521
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MEASUREMENT METHODS AT HIGH TEMPERATURE AND PRESSURE
D. R. Blann
Acurex Corporation, Aerotherm Division
485 Clyde Avenue
Mountain View, California
523
-------
MEASUREMENT METHODS AT HIGH TEMPERATURE
AND PRESSURE
D. R. Blann
Acurex Corporation, Aerotherm Division
485 Clyde Avenue, Mountain View, California
Symposium on
"Particulate Control in Energy Processes"
May 11-13, 1976
San Francisco, California
ABSTRACT
The current evolution of coal conversion processes has generated a
need for a new class of particulate and gaseous sampling instrumen-
tation. Many new processes, such as fluidized bed combustion, coal
gasification, and combined cycle power plants will operate at condi-
tions of high temperature (some greater than 1000°C) and high pres-
sure (up to 100 atmospheres). The combined risks of temperature,
pressure, corrosion, and toxic and flammable chemicals will present
a host of difficult problems in mechanical design, sample handling
procedures, and safety when designing sampling systems for these
processes.
This paper presents a brief review of some approaches which have
been, or are being, used to sample HTHP processes, and a discussion
of some of the major problem areas which must be considered in HTHP
sampling.
524
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MEASUREMENT METHODS AT HIGH TEMPERATURE
AND PRESSURE
INTRODUCTION
A very significant obstacle facing the developers of coal conversion
technology today is a lack of information on the service conditions of the
conversion processes themselves. For many processes the actual gas composi-
tion, particle loading and physical state are not known with any significant
degree of certainty. What data are known are based on bench-scale experi-
ments or computer simulation. Getting this information, from pilot- or full-
scale facilities, is required for the proper design, process control, and
pollution abatement of these processes.
However, many of the new coal conversion processes will operate at
conditions of high pressure (up to 100 atmospheres) and high temperatures
(some greater than 1000°C). Sampling at these conditions will require the
development of a whole new class of sampling technology. The combined risks
of temperature, pressure, corrosion, and toxic and flammable chemicals pre-
sent a host of difficult problems in mechanical and thermal design, sample
handling procedures, and safety when designing sampling hardware for these
conditions.
The hardware, procedures, and methodologies to accomplish these HTHP
sampling tasks do not yet exist in any standard form or approach. It is not
the purpose of this paper to solve these problems or even suggest any stan-
dard approaches. Only research and development will do that. We will,
however, indicate some of the approaches that have been used and are being
used (a cursory state-of-the-art survey) and identify the major problems
which HTHP sampling will pose.
STATE-OF-THE-ART IN HTHP SAMPLING
State of the art in HTHP sampling technology is something of a mixed
bag of tricks. In general, HTHP sampling environments have not lent them-
selves to standard sampling approaches; each specific process tends to have
its own set of difficulties. Examples of HTHP hardware one finds in the
literature are generally custom designed to solve specific access or technical
525
-------
problems, sometimes with strict economic constraints as well. An exception
to this is in the area of flame composition sampling where there are com-
mercial suppliers who offer off-the-shelf and specialty items as standard
D-orhict lines (see Figure 1).
Space does not permit a comprehensive state-of-the-art review of HTHP
sampling techniques, but the following paragraphs present some notable ap-
proaches that have been or are being used,
Sampling in the Bureau of Mines Coal-Fired Gas Turbine
In the early 1960's the Bureau of Mines conducted extensive develop-
•-•D-;t work on a coal burning gas turbine at its Morgantown (W. Va.) Coal
/esoarch Center. Coal ash erosion of the turbine blades was a critical fac-
tor, the evaluation of which required accurate and reliable knowledge of the
duct characteristics at the turbine inlet and exhaust. A. S. Moore of the
USBM developed a rather sophisticated particulate sampling probe for sampling
(2)
the 19-inch I.D. ducts leading to the blade sections.
Gas temperature and pressure for this application were 1300°F to 1500°F
and 70 to 80 psig respectively. The probe was 1/2-inch I.D., 316 S.S., with
nozzles of either titanium carbide or Hasteloy HS-31. Both the probe and the
t',:be were water cooled. The probe had a powered traverse mechanism for pro-
-'ing the duct (Figure 2(a)).
Australian Coal Burning Gas Turbine Research
The Australians have also conducted extensive direct-fired coal com-
bustion research between the years 1948 to 1970. Much of their work was
directed at determining the effects of blade erosion due to particulates and
to characterize such erosion it was necessary to sample both mass and size
of the particulate entering and exiting their turbine.
They obtained their data by direct sampling and counting particles.
particulates were collected either by small cyclones or absolute type fil-
(3)
ters. ' More detailed data is not available at this time.
Gas Sampling at the Bureau of Mines Stirred-Bed Reactor (Gasifier)
Figure 2(b) shows a schematic of the gas sampling and analysis system
also used at the Morgantown Coal Research Center for sampling their stirred-
bed gasification process currently under development.
526
-------
SELECT FROM THESE STANDARD GAS SAMPLING TYPES AND MODELS
Description
TYPE GA
Uncooled, of welded construction, this
probe is suitable tor use up to 2000°F and
Is the smallest standardly offered.
TYPE GB (Inconel Tip)
This probe can be used either as a heated or
cooled probe. The cooling passages are
especially designed for unrestricted flow of
cooling or heating fluids. For some applica-
tions. it is important to keep the sampled
gas above 500° F to the analyzer to prevent
condensation of evaporated hydrocarbons.
In this case, the probe is heated by blowing
hot air through the cooling passages. This
is more satisfactory than trying to heat it
with an electric coil because the variation of
heat loss conditions from the surface of the
probe makes it almost impossible to main-
tain an even temperature. For higher tem-
peratures. the probe can be kept hot enough
to prevent condensation but cool enough
not to rnelt by blowing unheated shop air
through the coolant passages. Type GC is
recommended for temperatures exceeding
3000- F.
TYPE GC (Copper Tip)
This probe has a copper tip with high
enough conductivity to prevent burning out
at the higher temperatures and velocities.
Since the tip is easily damaged, it is not
recommended for lower temperature
applications.
TYPE GA-F (Internal Pressure & Water Taps)
All Stainless Steel construction. Reinforce-
ment at pressure take off section offers
female tap for ease of installation. Performs
exactly like Type GA.
Ordering
Part No.
GA-12-.050
GA-24-.050
Probe'
Length
12"
24"
Dimensions
GB-12-,050
GB-24-.050
24"
GC-12-.050 12"
GC-24-.050 '. 24"
GA-F-12-.050J 12"
GA-F-24-.050! 24"
GB-F-12-.050i 12"
i
JGB-F-24-.050.I 24"
TYPE GB-F (Internal Pressure & Water
Taps, S/S Tip)
Same as Type GB except internal pressure
taps. Barrel is 3/V O.D. and sensing head is
W O.D. Specially designed for ease of
installation.
TYPE GC-F (Internal Pressure & Water
Taps, Copper Tip)
Same as Type GC except internal pressure
taps. Barrel is %" O.D. and sensing head is
>/2" O.D.
'Longer or shorter lengths available Consult United Sensor.
tlnlet holes may be enlarged upon request.
GC-F-12-.050
12"
i/l NM
1/16 NPt TAP
GC-F-24-.050I 24"
J_
Figure 1 Standard gas sampling probes.
527
-------
I Position indicator
2. Limit switch
3. Probe and pttoMube carriage
4. Fourrpm speed-reductor shaft
with change gears
5. Slop switch
6. Lead screw
7 Waler-jackeied pitoi tube
8. Waler-jackefed probe
9. Pressure seal
SO. Probe to pilot-tube coupling
Nineleen-inch diameter inlet duct
12. Probe and pilot-tube cantilever
support
Ln
M
OO
Figure 2(a) Sampler for USBM coal burning gas turbine.
(2)
-------
100CTF - IKHTF
100 - 300 psig
(Filter)
Pressure
let-down
Heat
exchanger
350°F
Side stream
gas
Filter
(tar)
N
To analyzers
(HCN, NH3, organic sulfur)
Heat
exchanger
Ice bath
To analyzers and stack
(fixed gases)
Figure 2(b)
Gas sampling and analyzing system for
minor constituents.
Light oils,
water
529
-------
Stream conditions are 100 to 300 psig, 1000°F to 1100°F. Particu-
lates are caught at high temperature in a bulk filter. After pressure and
temperature reduction, the condensing tars are caught on an alundum or sili-
cate filter; gases are analyzed at several stages in the sampling process,
as shown.
Sampling at the BCURA Fluidized Bed Installation, Leatherhead, England
Recent telecons with personnel at the BCURA pressurized fluidized
bed combustor in England have revealed how they approached sampling their
process. Figure 3 is a drawing of their pilot-scale MKII combustor; it
(4)
substantially represents the real installation/ '
BCURA sampled for particulate size at 800°F to 925°F and 4 atmospheres
following their primary and secondary cyclones on the horizontal run exiting
the second cyclone, which is a 10-inch I.D., 4-foot long duct followed by a
6-inch I.D., 4-foot long water quench section.
Access to the duct was through a gland installed on an elbow at the
end of the water quench section. They inserted a 5/8-inch S.S. probe axi-
al ly some 4.5 feet up the horizontal duct (which gets the sampling nozzle
ahead of the quench section) and withdrew a sample isokinetically (120 ft/
sec), collecting samples in a cyclone. Sizing was done by coulter counter.
Particles have been found to be primarily less than 5 microns.
Westinghouse Developed Sampler for Pressurized Fluidized Bed Combustor
Sampling
Also in the area of pressurized FBC process sampling, Westinghouse
engineers designed and constructed a small single point sampling system
which consisted of a sampling probe (~l/4-inch tubing), static pressure
probe, impactor (Brinks or Anderson) scalping cyclone, heated oven, and
miscellaneous valves and hardware. Figure 4 presents the system somewhat
schematically.
The device was designed to be installed and removed only when the
process was not in operation. The impactor was placed within a pressure
chamber inside the oven (350° capability) so that neither it nor the oven
had to withstand system total pressure (10 atmospheres).
530
-------
14
15
13-
16
17
12
21
18
1 WATER INLETS AND OUTLETS
2 FIRST STAGE CYCLONE
3 RECIRCULATION CYCLONE
4 BALANCING AIR SUPPLY
5 STARTUP GAS BURNERS
6 BED REMOVAL PIPE
7 PRESSURE SHELL
8 WATER-COOLED LINER
9 COMBUSTOR CASING
10SECOND STAGE CYCLONE
11 AIR INTAKE
12 CASCADE
13 ALKALI SAMPLING PROBES
14 MIXING BAFFLE
15 NOX SAMPLING POINT
16 WATER SPRAYS
17 DEPOSITION PROBE
18 DUST AND GAS SAMPLING PROBE
19 TO PRESSURE LETDOWN VALVE
20 RECIRCULATION CYCLONE
21 COAL INLET
22 AIR DISTRIBUTOR
ASH OUTL
DETAIL SHOWING ARRANGEMENT
OF TUBES
IN FLUIDIZED BED
Figure 3 BCURA pressurized fluidized bed
combustor (Reference 4).
531
-------
U1
OJ
1-0
Refractory
"plug"
3" Nipple
n-
Inconel/ / /£">•>•»-sa^
lining / // |
I— 2" hole
• Standard
3" flanges
Figure 4 Westinghouse 10 atm participate sampler.
Oven
-------
High Temperature. High Pressure Sampling at Aerotherm
Aerotherm Division of Acurex Corporation has developed several sam-
pling probes in the past 10 years which have been applied to high tempera-
ture and/or high pressure environments.
Figure 5 shows a miniature water-cooled gas sampling probe utilized
in pinpoint sampling of high enthalpy, high velocity gas flows typical of
those issuing from an arc-plasma generator. Environmental conditions in-
cluded temperatures to 7000°F, 20 atmospheres total pressure, 15,000 ft/sec
velocities and corrosive gases. Survival in this extreme environment was
attained by a high performance liquid cooling system which employed nucleate
boiling phenomena to affect adequate heat transfer rates. Coolant flow pres-
sure drops were on the order of 1000 psi. A similar probe has been supplied
to the Combustion Research Section of the EPA for flame sampling.
A somewhat more conventional example is shown in Figure 6 which shows
one of a line of water-cooled stack sampling probes designed for use at tem-
peratures up to 2400°F. The nozzles, which were uncooled, were fabricated
P
of Inconel for high temperature resistance. In some applications, special
P
high temperature Cermet coatings such as Sermatel were applied to provide
corrosion resistance.
During the past year, Aerotherm has nearly completed development of a
sophisticated particulate train for the Navy designed for operation in the
exit plane of military jet engines. The system, shown in Figures 7 and 8,
will automatically isokinetically sample the jet stream over a velocity range
of Mach 0.1 to Mach 1.3. Gas stream conditions include temperatures to 1400°F
and pressures to 50 psi. The system gives real-time particle size distribu-
tion data for submicron particles by means of a charge mobility analysis de-
vice by Thermo-Systems, Inc.
A rather interesting feature of this system is a microprocessor with
programming capability which permits the rapid automatic calculation and flow
control necessary to respond to engine operating point changes, and which will
ultimately permit automatic positioning of the probe anywhere on a two-
dimensional x-y plane.
533
-------
o
2
INCHES
Figure 5 Aerotherm Miniature Water-Cooled
Gas Sampling Probe
534
-------
Figure 6 Water
probe.
-------
Jet engine sampler.
Figure 7 Control console with microprocessor. Figure 8 Sampling train and support stand.
-------
Current HTHP Sampling Research and Development at Aerotherm
Aerotherm currently has under development (for the EPA), a sampling
system designed to acquire particulate mass loading and size distribution
data in process gas streams such as pressurized fluidized combustion pro-
cesses. The system was designed to be applicable to a rather broad range
of process conditions, shown in Table 1. As a result it contains certain
features not normally associated with process-specific systems such as
those we have been reviewing in the foregoing paragraphs.
The probe is hydraulically activated, due to the high blow-off loads
experienced at process pressures up to 20 atmospheres. Probe cooling and
sample conditioning is accomplished by a Dowtherm circuit. Dowtherm per-
mits sample conditioning to 600°F regardless of process conditions and
avoids the "cold wall" condensation problem. The specified temperature
and pressure range represents a turn-down of approximately 10:1 in sample
mass flowrate. Figure 9 presents a pictorial overview of the system; sev-
eral detail views of the probe assembly are presented in Figure 10.
Development of this system is being carried out as one subtask of an
extensive level-of-effort program sponsored by the Industrial Environmental
Research Laboratory (W, Kuykendahl, Project Officer) entitled "Measurements
for High Temperature, High Pressure Processes," The total program extends
3 years and is basically structured to extend the state of the art in par-
ticulates and gaseous sampling technology to 1500 psi, 2000°F environments.
Emphasis will be on coal conversion processes (primarily gasification) and
pressurized fluidized bed combustion. Program structure is shown in Fig-
ure 11.
MAJOR PROBLEM AREAS ENCOUNTERED IN HTHP SAMPLING
High temperature, high pressure sampling presents several formidable
problem areas to the system designer:
• Pressure and temperature
t Corrosion and erosion
• Condensation of process constituents
• Toxicity of process constituents
t Flammability and explosion hazard of process constituents
537
-------
TABLE 1 HIGH TEMPERATURE, HIGH PRESSURE
SAMPLING SYSTEM SPECIFICATIONS
Sample Environment
• Temperature
• Pressure
• Stream Velocity
• Participate Grain
Loading
* Participate Size
Range (for classi-
fication)
• Duct Size
Traverse Capability or
Penetration of Nozzle
into Duct or Vessel
Access Process Port
Requirements
1200°F - 1800°F
650°C - 1000°C
3-20 atmospheres
15-100 fps
0-15 gr/ft3
0.2 - 7.0 microns
Variable depending on probe;
std. is 8 inches I.D. minimum
Approximately 26 inches either
in situ or extractive configu-
ration (some dependence on in-
ternal configuration of duct
vessel)
Standard: 4 inches IPS minimum,
300 Ib flange acess through
4 inches IPS alloy gate valve
538
-------
n-
2
VO
HYWWUUC SUPPLT SYSTEM
COWTWH. CONSOLE
High temperature, high pressure sampling system.
Figure 9
-------
GATE VALVE
-PROBE SHOWN IN FULLY
RETRACTED POSITION
REF
TRUSS ROD TIE
BAR 4 MOUNT
-BEARING
PEAR ACCESS
=>LATE
GATE VALVE
SEE VIEW C
TRUSS ROD
(4 REQD)
-HYDRAULIC
CYLINDER
PROBE ASSEMBLY
PROBE OUTER TUBE
HOUSING
PROBE HOUSING
HYDRAULIC CYLINDER
MICRO-SWITCH
(2 REQD)
— MICRO-SWITCH
/ ACTUATOR
ACTUATOR ROD
ACTUATOR ROD
GUIDE TUBE
PROBE HOUSING
HYDRAULIC
CYLINDER
TRUSS ROD TIE
BAR i MOUNT
ACTUATOR ROD,
MICRO-SWITCH
REAR ACCESS
PLATE
0-RING
PROBE OUTER
TUBE HOUSING
VIEW C
-PROBE HOUSING
BEARING
Figure 10(a) High temperature pressure probe assembly.
-------
ACTUATOR ROD
GUIDE TUBE
PROBE HOUSING
TUBE MOUNT
PROBE OUTER TUBE HOUSING
HYDRAULIC--.
CYLINDER ^
HYDRAULIC
CYLINDER
TRUSS ROD
(4 REQD)
SECTION A-A
•TRUSS ROD TIE BAR
AND MOUNT (3 REQD)
ACTUATOR ROD
GUIDE TUBE
TUBE MOUNT
MICRO-SWITCH
MICRO-SWITCH MOUNT
SHAFT
HYDRAULIC
CYLINDER
PROBE
HOUSING
BEARING
PROBE HOUSING
MICRO-SWITCH
PROBE OUTER
TUBE HOUSING
SHAFT, HYDRAULIC
CYLINDER
SECTION -
TRUSS ROD
(4 REQD)
Figure 10(b) High temperature pressure probe assembly.
541
-------
MEASUREMENTS FOR HIGH TEMPERATURE, HIGH PRESSURE PROCESSES
PROJECT ORGANIZATION
STATE-OF-THE-ART REVIEW
• Evaluation of existing
HTHP instrumentation
• Process Review
PROGRAM MANAGER
Fred Moreno
PROJECT ENGINEER
Creighton Hartman
DESIGN & DEVELOPMENT
0 Near-term techniques
— Design
- Lab test
— Field demonstration
• Advanced techniques
SAMPLING SUPPORT SERVICES
t Review of EA/TD programs
• Sampling support services
• Preparation of guidelines
and procedures
Figure 11 Program structure for EPA-sponsored HTHP process measurements program.
-------
Pressure and Temperature
From a design point of view, the fundamental problems are those which
arise as a result of the extremes of temperature and pressure. If the sam-
pling system fails catastrophically when subjected to the duct environment
due to inadequate structural integrity, other experimental problems (sticky
particulate, probe deposition, cold wall condensation) or scientific nice-
ties become totally irrelevent.
Two factors form keystone elements in any design related to HTHP
sampling. These are:
• Safety — personnel operating the equipment should not be subjected
to any substantial risk of injury to life or limb due to inadequate
design
• Reliability - the device should operate over many cycles with no
significant degradation of performance or increase of the proba-
bility of failure
The purely technical considerations for designing for extreme condi-
tions have mostly to do with the structural behavior of materials. Critical
material properties are severely degraded at temperatures above about 1000°F.
At high pressures, this property degradation becomes critical because of the
tremendous structural loads and stresses that can be imposed on the machine
elements.
Seals are also critical. Insertion of a sampling device requires
seals somewhere at the point of access. Safety and reliability both require
there should be little possibility of a process leak. Even if thermally pro-
tected, a small leak of high temperature gas due to improper design, con-
tamination of the sealing surface, or inadequate or careless seal installa-
tion, can lead to catastrophic consequences.
Attaining access to the high pressure environments for sampling is
a challenging problem in itself. The larger the port required, the larger
the blow-off loads on the system restraints. For a seal diameter of 4 inches,
a representative port size for conventional stack sampling, the blow-off load
is almost 20,000 pounds at 100 atmospheres!
543
-------
To maximize safety and reliability in HTHP environments, sampling
techniques, methodologies, and equipment should be well proven. The uti-
lization of good pressure vessel design practices and thorough application
and interpretation of analytical procedures cannot be overemphasized. Non-
destructive tests, of which there are many types, should be utilized where
possible to prove a sampling system or demonstrate basic integrity of cri-
tical component parts. Proof tests and simulation tests (where possible)
in the laboratory create even greater confidence in a design.
In summary, safety and reliability can be maximized by:
• Thorough attention to detail
• Rigorous design and analysis
• Rigorous quality control of fabricated parts and certification
of critical materials
• Thorough laboratory tests before field operations
t Detailed test planning, preparation, assembly procedures (leave
nothing to chance)
• Thorough post-test inspections of equipment
Corrosion and Erosion
The environments associated with coal conversion technology can be
extremely hostile. The most severe conditions exist, of course, in the main
gasifier vessel and the transfer-lines which see high temperature, high
pressure, corrosive gases, and high velocity char and ash particles which
can cause severe erosion.
Fortunately, the sampling system designer generally does not have to
deal with these problems on a long term basis as does the process designer.
A material loss rate of 500 mils/1000 hrs may be acceptable to a sampling
system used intermittently for short time exposures, whereas such a loss rate
for a process component would be absolutely intolerable if reasonable life-
time is to be realized. However, some sampling components are sensitive to
dimensional stability for proper operation (such as nozzle inlets, cyclones,
inertia! cascade impator plates); thus, materials performance becomes critical
544
-------
even for intermittent operation, and it is certainly true for sampling equip-
ment used for In situ process monitoring, such as particulate monitoring up-
stream of a combined cycle turbine.
Both gaseous and liquid phase corrosion must be considered. Gaseous
corrosion may consist of oxidation, carbonization, sulfadation and hydrogen
embrittlement. Liquid corrosion may result from condensing liquid deposition
from the gas onto surfaces below the dewpoint, by slag deposits from coal ash
components, and by reactions of sulfur with nickel to form relatively low
melting constituents in the metal.
There is currently little information available on erosion in high
temperature, high pressure environments. Potentially, the most severe damage
resulting from impact erosion is the removal of the protective oxide scale
which protects from corrosion.
Space does not permit more in-depth review of these problems and even
less so a discussion of their potential solutions. However, in general, it
can be stated that a significant degree of protection from corrosive effects
can be obtained by additions of nickel (>22 percent), chromium (>22 percent)
and molybdenum (>3 percent) to the alloys. It has been found in some well
planned experiments being performed by the Materials Properties Council of
the Anerican Gas Association that additional protection in gasification-type
environments can be obtained by cladding with 671 alloy (50 percent nickel —
50 percent chromium) which is too brittle for most structural applications,
by aluminizing, coating, and/or cooling.
Sometimes the erosion problem can be reduced by hard facing the metal
with ceramic coatings. While this does not necessarily reduce corrosion, it
will protect the oxide scale from erosion and thus the corrosion rate will
not be accelerated by erosion.
Applications expertise in materials is critical to success. For ex-
ample, many ceramics (which show great promise in corrosive/erosive environ-
ments) lack thermal shock resistance, which can lead to total failure by
disintegration if not properly restrained or configured. Expense and lack
of fabricability are factors which often mitigate against the use of exotic
materials no matter how great their potential. Ultimately, the best design
advice is (where possible):
545
-------
• Cool components to reduce operating temperature
0 Reduce flow velocities, especially particle-laden streams
• Design for low stresses
Condensation of Process Constituents
Many HTHP processes under consideration for particulate and gas sam-
pling contain many chemical species which may be liquids or solids at room
temperature. Therefore, any proposed sampling approach must consider the
possibility that under certain conditions some species may change physical
states, even several times, during the sampling process.
For example, many of the "coal tar" derivations (cresol, phenol,
toluene) have critical pressures in the region of 40 to 50 atmospheres.
Their critical temperatures are above 300°C. If a sampling probe operating
at 500°C and 70 atmospheres in a coal processing stream (gasification, liqui-
fication, etc.) samples an appreciable volume of those components and cools
them below 300°C (which would be reasonable) before reducing the pressure,
these species are quite likely to condense and plug the flow channels with
spurts of liquids. Even worse, if the pressure is then reduced after cool-
ing, these components can later flash from liquid to gas before finally being
collected.
Condensation can be a significant problem in the "simple" combustion
environment. Streams with very high percentage levels by volume (20 percent)
of water vapor can have condensation temperatures as high as 300°F at pres-
sures of 300 psia. However, gasification probably represents the most severe
potential problems of high temperature condensation. A LURGI gasifier, for
example, produces a broad variety of tars (boiling point 500°F+), tar oils
(B.P. 285°F - 680°F), and naptha oils (B.P. 120°F - 320°F). If sample integ-
rity is to be preserved and a representative sample obtained in such systems,
the sample must be held at very high temperatures indeed; quite possibly as
high as 750°F to 1000°F in some cases.
Compounding this problem are those constituents such as phenol, re-
sourcinal, napthalene, etc., which have relatively high vapor pressures at
low temperatures. It would be very easy to "lose" these compounds during
sample handling.
546
-------
To illustrate this problem more explicitly, Figure 12 shows how high
temperature condensibles could affect the sampling process, The process
stream conditions are taken from the Clean Fuels from Coal Symposium, Sep-
tember 1973, sponsored by IGT, Chicago. For illustration, worse case con-
ditions were assumed to prevail, and a poorly designed, but not totally
unrepresentative, sampling system was assumed.
Clearly, the chemical composition of the process stream must be con-
sidered in the selection of sampler operating conditions, not only in the
sampling probe itself, and in the particulate removal stages, but in all
the subsequent gas sample steps which follow.
Toxicity of Process Constituents
Sampling personnel should be aware of the potential health hazards
that exist with respect to many of the process constituents in HTHP pro-
cesses. Many of the "coal tar" compounds found in gasification processes
are known carcinogens or have other toxic properties, which makes operating,
cleaning, and maintaining the sampling equipment a safety consideration.
Flammability and Explosion Hazards
In gasification processes, the gas flow through the sampler will be
flammable and/or explosive. Significantly, sampling rates are such that
after expansion, considerable volume is represented which may preclude sim-
ple venting if there is any possibility of sparks, static charge build-up,
or open flame.
Certainly, sampling equipment for such processes should be grounded,
and all motors, switches, and controls should be explosion-proofed. What-
ever alternatives are selected for handling these gases, protection of per-
sonnel and process must be of primary consideration.
CONCLUSION
It is anticipated that over the next several years important develop-
ments will be forthcoming in coal conversion technology, as the effort con-
tinues to utilize this nation's abundant coal reserves in an environmentally
acceptable manner. A necessary component of this development will be the
547
-------
IM PACTO(Z
Ul
.p-
00
PRESSOKS
l?epuc
VALV/6
Cj4AM'SE(2
SOME INJ
PROCESS
FROM BIG? AS
VARIOUS COAL
•V PARTICULATES
FLOATING) ON
O £l F ICES
VALVE
TO «£.TEAM
- 7 %
PAeTicu-
-48%
LATE5
NtCS
-i-
Figure 12 Condensation problems arising from hypothetical
inadequate sampling approach.
-------
evolution of new particulate and gaseous sampling instrumentation, methodolo-
gies, and techniques. This paper has presented only briefly some of the
challenges that face this evolutionary development.
549
-------
FBC PARTICULATE CONTROL PRACTICE
AND FUTURE NEEDS; EXXON MINIPLANT
R. C. Hoke
EXXON Research and Engineering Company
Linden, New Jersey
551
-------
FBC PARTICULATE CONTROL PRACTICE
AND FUTURE NEEDS: EXXON MINIPLANT
R. C. Hoke
Exxon Research and Engineering Company
Linden, NJ
ABSTRACT
The successful development of the pressurized fluidized bed
coal combustion system is dependent on the removal of particulates from
the high temperature, high pressure flue gas before it is allowed to
expand through a gas turbine. Current estimates of the allowable par-
ticulate loading at the turbine inlet, based on turbine blade erosion
limits, are in the range of 0.001 to 0.05 gm/m3 (0.0004 to 0.02 gr/SCF).
These limits are lower than the current EPA particulate emission standard
of 0.115 gm/m3 (0.05 gr/SCF). Corrosion of the turbine by alkali sulfates
is also a potential problem and may impose even tighter allowable inlet
particulate loadings.
Reducing particulate loading down to these low levels requires
very efficient devices capable of operating at higher temperature (950°C)
and high pressure (10 atm). A number of devices were considered and the
system chosen consists in two stages of cyclones followed by a high
efficiency granular bed filter. The granular bed filter must be capable
of 90 to 99% removal efficiency to meet the estimated particulate loading
requirements. Although granular bed filters are still in the development
stage and have not as yet demonstrated these high removal efficiencies,
they offer the most promise compared to other available devices. A
granular bed filter has been purchased from the Ducon Company to be
installed on the Exxon FBC miniplant. A test program will be carried
out to determine particulate removal efficiency, maintenance of removal
efficiency and mechanical performance.
552
-------
FBC PARTICULATE CONTROL PRACTICE
AND FUTURE NEEDS: EXXON MINIPLANT
Fluidized bed combustion (FBC) Is a new direct concept tech-
nique which permits the use of coal in an efficient manner with minimum
impact on the environment. Two versions are being developed. The first
operates at atmospheric pressure and is being developed for use in both
electrical power generation and production of steam for industrial use.
The second operates at higher pressure, up to 10 atm absolute, and is
being developed for use in a combined cycle power generation system. A
sketch of a pressurized FBC system is shown in Figure 1. Coal, along
with limestone or dolomite which act as S02 sorbents, is injected into
the bottom of the pressurized boiler. Coal is burned in the limestone
bed which is fluidized by the incoming combustion air. 862 formed in
the combustion process is removed by the limestone bed. Steam coils
immersed in the fluidized bed remove the heat of combustion and maintain
the bed temperature in the range of 1500 to 1700°F. Steam thus generated,
is used to operate a steam, turbine. The desulfurized flue gas passes
through a particulate removal system and is then expanded across a gas
turbine which generates additional power and also operates the combustion
air compressor. A portion of the particulates removed from the flue gas
is recycled to the boiler to increase the carbon combustion efficiency.
The particulate removal system must reduce the particulate loading down
to levels .sufficiently low to protect the gas turbine and meet current
emission standards.
Figure 1 also shows a regeneration section coupled to the boiler.
This is an adjunct to the boiler which regenerates the sulfated limestone
to lime which is then returned to the boiler. An off gas with S02 con-
centration sufficiently high to be fed to a sulfur recovery system is
also produced. The regeneration system is currently at an earlier stage
of development than the boiler system. Although an FBC system can operate
with limestone used in a once-through manner, the regeneration of sulfated
limestone would reduce the quantity of limestone fed to the boiler and the
disposal of used limestone.
An atmospheric pressure FBC system is similar to the pressurized
system with the exception of the flue gas turbine xvhich is not used.
The FBC system offers a number of advantages over a conventional
coal fired boiler. It effectively controls S02 emissions without external
scrubbing systems. NOX emissions are also very lox^ because of low com-
bustion temperatures and also because conditions in the boiler favor NO
destruction reactions. For example, under pressurized FBC conditions,
NOX emissions are typically in the range of 100 to 150 ppm. An FBC system
can also handle a variety of fuels, gaseous, liquid and solid. It also
provides higher generation efficiency at lower capital and operating
costs.
553
-------
GAS TURBINE
TO
STACK
DISCARD
STEAM TURBINE
CONDENSER
COAL AND
MAKEUP SORBENT*
AIR
COMPRESSOR
SOLIDS
TRANSFER
SYSTEM
TO SULFUR
RECOVERY
1 SEPARATOR
DISCARD
BOILER
FUEL
REGENERATOR
Figure 1
PRESSURIZED FLU1DIZED BED COAL COMBUSTION SYSTEM
554
-------
The work currently in progress at Exxon Research and Engineering
Company is sponsored by the U.S. Environmental Protection Agency (EPA).
A number of other FBC programs are now underway in the U.S. and the U.K.
under sponsorship of the EPA, the U.S. Energy Research and Development
Administration (ERDA) and the Electric Power Research Institute (EPRI).
These programs have progressed from the laboratory and bench scale to
process development units burning 5 to 10 T/D of coal. Larger pilot
plants, of atmospheric and pressurized design in the range of 10-30 MWe,
are now under design and construction. Environmental assessment and
economic evaluation studies are also in progress.
The program at Exxon Research is being carried out in two
pressurized FBC units, a bench scale unit and a process development unit.
The process development unit, also known as the "miniplant" is capable
of burning up to 6 T/D of coal at pressures up to 10 atm. The design
of the unit was described in a report submitted to .the EPA1. The combustor
section of the miniplant has been in operation since July 1975 and has to
this date logged over 1200 hrs of coal combustion time. This includes a
10 day continuous demonstration run which was completed in November 1975.
One of the objectives of the FBC program at Exxon Research is
the evaluation of particulate removal systems. These systems must satisfy
three performance criteria as shown in Table 1.
Table 1
Particulate Emission Control Requirements
Required Efficiency
Requirement Allowable Level of Third Stage Device
(gr/SCF) (%)
Environmental 0.05 67
Turbine Erosion
Range of Estimates 0.02-0.0004 87-99.7
Tentative Level 0.002 98.7
Turbine Corrosion ? ?
The first performance requirement is based on the environmental
emission standard as set by the EPA. The second is protection of the gas
turbine from excessive erosion and the third is protection of the turbine
from particulate induced corrosion. The environmental emission standard
is 0.1 Ib particulate/10^ BTU fired. This is equivalent to a particulate
concentration in the flue gas of about 0.05 gr/SCF. At the present time,
turbine erosion limits are not well defined. Westinghouse Research
Laboratory recently estimated allowable levels based on limited data and
model studies.2 The estimates covered a range of 0.02 to 0.0004 gr/SCF.
Based on these estimates, Westinghouse suggested a tentative allowable
555
-------
level of 0.002 gr/SCF. Insufficient data are available to estimate the
allowable level to prevent corrosion, but some qualitative estimates sug-
gest that these levels may be even lower than the erosion limits.-^ The
particulate concentrations determined by the environmental and erosion
requirements are lower than can be met with simple particulate removal
systems such as high efficiency cyclones. It is estimated that a two
stage cyclone system could reduce the particulate loading to a level of
about 0.15 gr/SCF. Therefore, a third stage particulate removal device
would be also required to reach the levels set by the environmental and
erosion requirement. The efficiency of such a third stage device is also
shown in Table 1. To meet the environmental requirements, an efficiency
of 67% would be required. Efficiencies of 87 to 99.7% would be required
to meet the range of particulate levels set by erosion limits. If the
tentative limit suggested by Westinghouse is used, the required removal
efficiency would be 98.7%. Obviously, a very efficient third stage system
will be required.
Some surveys have been made of the type of particulate removal
system which may be capable of satisfying these requirements. Stone and
Webster^ and Westinghouse^ have published such surveys. Other evaluation
programs sponsored by EPA, EPRI and ERDA, are also now underway. Three
systems have been mentioned in these studies which could provide the
required degree of particulate.:removal. These are granular bed filters,
high temperature metal or ceramic filters and a low temperature scrubber
combined with an efficient heat exchanger. The consensus appears to
favor granular bed filters as the type which offers the best chance of
meeting the required high removal efficiencies.
A number of groups are currently developing granular bed filter
systems and are described in the above referenced Westinghouse and Stone
and Webster reports. The Ducon Company has tested granular bed filters
on refinery and other waste gas streams. A type of filter, called the
panel bed filter is currently being studied at the City College of New
York under sponsorship of EPRI. The Rexnord Company has installed a
number of "gravel bed" filters on cement kiln and other industrial off
gases. Combustion Power Company has also installed a number of "dry
scrubber" filters on flue gas from wood waste boilers. Combustion Power
is also developing a similar system for application on FBC.
A sketch of the Ducon filter is shown in Figure 2. The filter
consists of a series of beds containing the filter medium, stacked vertically
to form a filter element. A number of elements are contained in a pressure
shell. Dirty gas passes through an inlet screen, through the filter
medium and out through an outlet screen into a central collecting tube.
Clean gas exits the vessel at the bottom. Each element is periodically
cleaned by a short pulse of high pressure clean air flowing in reverse
flow through each element. The dust is blown out through the inlet screens
and collects in the bottom cone of the pressure vessel. The filter medium
is retained in the beds.
556
-------
Figure 2
DUCON GRANULAR BED FILTER
REGENERATOR
FLUE GAS INLET
CLEAN GAS
OUTLET •*
FILTER
ELEMENT
COLLECTED
FINES
BLOWBACK
GAS PORTS
FILTER ELEMENT INTERNALS
OUTER
SCREEN
GRANULAR
SAND BED
INNER
SCREEN
557
-------
The CCNY panel bed filter is described in Figure 3. Thin sections
of the filter medium are formed between louvred panels as shown in cross
section in Figure 3. Four panels would be placed at right angles to form
a hollow square cross section. Dirty gas is filtered by passing through
the granular filter medium. The medium is periodically cleaned by a "puff-
back" of clean air in the reverse direction. This blows off the dust and
some of the filter medium which is collected, cleaned externally and reeir-
culated to the top of the bed. After each puff-back, cleaned medium drops
into the panel bed to replace material removed by the puff-back.
A sketch of the Rexnord gravel bed filter is shown in Figure 4.
Dirty gas enters into a cyclone pre-cleaner which is an integral part of
the system. The pre-cleaned gas passes out the vortex tube and passes
down through the gravel beds and out the clean gas exit. The beds are
periodically cleaned by a reverse flow of clean gas. During the reverse
flow step, the gravel beds are raked by a rotating raking mechanism to
aid in removing the dust from the filter medium. The dust is then blown
in reverse direction through the cyclone and collects at the bottom of
the pre-cleaner.
Figure 5 is a sketch of the Combustion Power dry scrubber.
Dirty gas enters the containing vessel and flows through a louvred anular
vessel containing the filter medium. Clean gas exits through the central
collector. The filter medium is constantly fed into the top of the vessel,
continually moves down through the louvred containing vessel, passes out
the bottom, is cleaned externally and recycled back to the inlet.
The Ducon filter was chosen for testing on the Exxon/EPA miniplant
after considering all the available systems. The selection was based on
previous experience with the Ducon system which indicated that it had the
potential of providing high removal efficiency. Another desirable feature
of the Ducon system is the retention of the granular filter medium in the
filter vessel. In all other systems, the medium is removed, cleaned
externally and recycled back to the filter vessel. The Ducon system to
be used on the miniplant will consist of four filter elements, each 1 ft
in diameter by 6 ft long. The elements will be enclosed in a refractory
lined pressure vessel designed for 10 atm operating pressure. The system
will be capable of filtering 1200 SCFM of hot flue gas. It is currently
being fabricated and will be installed late in 1976. It will be evaluated
in an experimental program in which the particulate removal efficiency and
efficiency maintenance will be measured as a function of operating para-
meters. Equipment life will also be studied. A cooperative program is
also being developed by EPA, ERDA and a gas turbine manufacturer to install
and test samples of candidate gas turbine materials in the filtered flue
gas. A second particulate removal system will also be selected, installed
and tested as part of the EPA program after the evaluation of the Ducon
filter is completed.
558
-------
Figure 3
CCNY PANEL BED FILTER
FRESH SOLID
DIRTY-GAS
INLET
CLEAN-GAS
OUTLET
COMPRESSED
GAS FOR
PUFF-BACK
DUST AMD
SOLIDS OUT
559
-------
Figure 4
REXNORD GRAVEL BED FILTER
CLEAN GAS
DIRTY GAS
DUST
560
-------
Figure 5
COMBUSTION POWER GRANULAR BED FILTER
DIRTY GAS
Y
DUST AND
GRANUlfS
CLEANED GRANULES
561
-------
REFERENCES
1. Skopp, A., et al, "Studies of the Fluidized Lime-Bed Coal Combustion
Desulfurization System," Exxon Research and Engineering Company,
Report to EPA Contract CPA 70-19, December 1971.
2. Keairns, D. L., et al, "Fluidized Bed Combustion Process Evaluation,"
Westinghouse Research Laboratory, EPA-650/2-75-027-C, September 1975.
3. Robson, F. L., et al, "Fuel Gas Environmental Impact: Phase Report,"
EPA-600/2-75-078, November 1975.
4. Zabolotny, E. R., et al, Stone and Webster Engineering Corp., Report
to EPRI, November 1974.
562
-------
PARTICULATE CONTROL REQUIREMENTS FOR GASIFIED
COAL-COMBINED-CYCLE POWER PLANTS
Fred L. Robson
and
William A. Blecher
United Technologies Research Center
East Hartford, Connecticut
563 '
-------
PARTICULATE CONTROL REQUIREMENTS FOR GASIFIED
COAL-COMBINED-CYCLE POWER PLANTS
Fred L. Robson and William A. Blecher
United Technologies Research Center
East Hartford, Connecticut 06108
ABSTRACT
The combination of low-Btu gasification processes with combined-cycle power
generation offers an attractive means of generating electricity from high—sulfur
coal while minimizing pollutants. Because the mass flow rate of low-Btu gas is a
significant fraction of the combustor air flow, the ability to use the large
amounts of sensible heat contained in the fuel gas at combined-cycle efficiency
is a key factor in system performance. Best performance would be obtained by
using the gas without cool down after leaving the gasifier. However, this means
that all cleanup systems, including those for particulates must operate at extremely
high temperature while producing a gas of a quality that is acceptable to the
gas turbine and that will meet effluent requirements. In order to identify the
particulate removal problem it is necessary to define the particulate content
of the gas entering the cleanup system, as well as limits imposed by the turbine
and the environment.
To realistically assess a candidate particulate removal system it is necessary
to consider the alternatives. The conventional approach to the problem would be
to reduce gas temperature to a level compatible with a water scrub. The low-
temperature gas could then be used directly in the burner or reheated by regen-
eration against the dirty gas to minimize the performance decrement. The perfor-
mance and cost associated with these alternatives are presented and their inter-
action with other cleanup functions, particularly those for sulfur compounds and
ammonia, are discussed in an attempt to define the benefits that can be expected
to offset the cost of high-temperature particulate removal.
The work described in the following paper was supported, in part, by EPA Contract
68-02-1099- Mr. W. J. Rhodes of the Industrial Environment Research Laboratory,
RTP, is the Project Monitor.
564
-------
PARTICULATE CONTROL REQUIREMENTS FOR GASIFIED COAL-COMBINED-CYCLE
POWER PLANT
INTRODUCTION
Our ever increasing reliance on foreign sources for energy can be relieved
to some extent by changing the manner in which electricity is generated. At present,
about hO% of the electric power in this country is based upon the use of oil and
natural gas, about 53$ on coal and 1% on nuclear energy. While the U.S. has vast
amounts of coal reserve, much of the coal near the major urban areas contains too
much sulfur to be burned directly. Of the several ways of generating nonpolluting
power using coal currently under investigation, gasification in conjunction with
an advanced power system such as the Combined Gas and Steam (COGAS) system appears
to be one of the more attractive possibilities.
The COGAS system (Fig. l) is based upon the use of advanced, high-temperature
gas turbines such as those being pursued under a recently announced ERDA program
to develop the technology for 2600 F turbine inlet temperature machines. Like
present day utility gas turbines which operate at under 2000 F, these advanced
turbines will require quite, clean fuels; fuels which have occasionally been refer-
red to as being of pharmacutical grade.
This cleanliness is required for several reasons, the major ones being erosion
and corrosion. Many of the turbine blades have very thin coatings of oxidation
resistant materials on their surface. Thus, not only the problem of catastrophic
erosion which changes the blades aerodynamics must be considered, but erosion by
small particles which could abrade the coatings would now be a problem. Once the
coatings have been penetrated, alkali metal salts, usually sodium or potassium
sulfates, attack the high-strength, high-temperature alloys which form the blade.
To keep blade metal temperatures to reasonable levels, below 1600 F, some type
of cooling is used (Fig. 2). One method uses small diameter holes (0.03 to 0.05
in.) in various parts of the blade to release cooler air from the blade interior
to form a film of cooling around the blade. It is possible that deposition of
solids could block these holes and cause hot spots thereby leading to localized
blade failure.
Thus, it is important that substances which are not only harmful to the
environment, but also harmful to the power system be removed or reduced to accept-
able levels. It still remains, however, to fully define the scope of the problem.
CLEANUP SYSTEM GOALS
The off gas from a low-Btu coal gasifier contains CO, H2, N2, C02, and E^O
as its major constituents. Unfortunately, the sulfur content in the coal is
565
-------
Ln
ON
AIR
COAL •
COMPRESSOR
GASIFIER
POWER TURBINE
BURNER
COMPRESSOR TURBINE
STEAM BOILER
TO STACK
t
PUMP
CONDENSER
ELECTRIC
GENERATOR
ELECTRIC
GENERATOR
Figure 1: Combined Gas—Steam Turbine System
-------
IMPINGEMENT
FILM
EARLY DESIGNS
PRESENT
CONFIGURATION
\
TEMPERATURE
CAPABILITIES
2200 F
2400 F
2600 F
Figure 2: Turbine Blade Cooling Configuration
-------
onverted mainly to H.2S and other sulfur bearing compounds, and in some types
of gasifiers, the organically bound nitrogen in the coal is converted to ammonia
which, upon combustion would form NOX. Of course, coal ash, carbon carryover,
and other solids will be carried in the off gas. Depending upon the gasifier
type, this gas may be between 1000 F and 3000 F and between 1 atm and UO atm.
For example, the system to be described in the following sections has a two-stage
gasifier with the off gas exiting at 1800 F and approximately 30 atm. Using
Illinois Ho. 6 coal, the off gas contains 0.51 mol percent sulfur compounds,
0.38 mol percent nitrogen compounds and approximately 0.8 gr/SCF of particulate.
(A solids carryover equivalent to 10$ of the ash content has been assumed. Actual
values of particulates from various gasifiers have yet to be verified.)
The size distriubtion of the particulates from the gasifier has not been
adequately documented. Since pulverized coal is the fuel to the gasifier, it has
been assumed that the size distribution in the ash carryover would resemble that
of the fly ash from a P.F. boiler as shown in Table 1.
The goals for cleanup given in Table 2 are based upon both environmental
and turbine requirements. The present EPA standards for large coal-fired stations
of 1.2 Ib S02/per 106 Btu, 0.7 Ib K0x/per 106 Btu and 0.1 Ib particulate/per 106 Btu
are easily met if these goals can be attained.
CLEANUP SYSTEM IDEITIFICATIOI
The gas from the gasifier can be cleaned to the levels indicated in the sug-
gested goals using commercially available processes. However, these processes
operate at low temperatures (below 250 F) and thus require considerable fuel gas
cooling. The various irreversibilities during this cooling process impose a thermo-
dynamic penalty. Also, higher process utility loads for the low-temperature
process results in an energy penalty. The attractiveness then of the high tempera-
ture cleanup system arises from the potentially better performance that the overall
integrated system (gasifier/cleanup/COGAS) might have. To demonstrate this, the
performance and cost of the low- and high-temperature system will be compared.
Integrated Power System with Low-Temperature Cleanup Process
An integrated power system using a low-temperature cleanup process is shown
schematically in Fig. 3. The cleanup process is the Allied Chemical Corporation
Selexol system and is capable of reducing the sulfur compounds to the 100 ppm
level; and with additional water wash, reducing the ammonia content to acceptable
levels. Since this process involves aequeous scrubbing in a number of stages,
the particulates are removed to very low levels at a cost and performance penalty
that is not easily differentiated from the sulfur/ammonia removal process.
A brief description of the salient operating characteristics of this system
are given in Table 3. The cost and performance are given in Table k and Table
5, respectively.
568
-------
TABLE 1
ESTIMATED PARTICLE SIZE DISTRIBUTION
(Flyash From P. F. Boiler)
Range Conmlative Percent (wt)
50-75 19
25-50 , 5^
10-25 82
5-10 9^
2-5 98
< 2 10°
569
-------
TABLE 2
FUEL GAS CLEANUP
Low-Btu Gas Typical Current Spec
Sulfur 0.05 mol/£ or less than <1.0 mol% or less than
amount to form 0.6 ppm amount to form 5 ppm
alkali metal sulfate alkali metal sulfate
o
Particulates h ppm wt or 0.0012 30 ppm or 0.01 gr/ftj
gr/ft3 > 2p
Metals
Vanadium < 0.003 ppm wt < 0.002 ppm wt
Sodium/Potassium See sulfur spec < 0.06 ppm
Nitrogen 500 ppm as NHg
570
-------
CLEAN GAS TO BURNER
ro
I
o
Figure 3: BCR/Selexol System
-------
Power System
Temperature - F
Pressure - psia
• Unit Size - MW
Gasifier
Exit Temperature - F
Pressure - psia
• Fuel Gas HHV-Btu/SCF
TABLE 3
SYSTEM CHARACTERISTICS
Gas Turbine
2600
350
180 (Horn)
1800
Hoo
135-150
Steam
1000
1250
300 (Horn)
572
-------
TABLE
Gas Turbine Out-MW
Steam Out-MW
Auxiliaries-MW
Net-MW
Efficiency
Emissions Ib/10 Btu
SO,
NO
Particulates
SYSTEM PERPORMMCE
Low-
Temperature
73^
321
82
973
38.6$
0.39
0.29
0.01
High- High-
Temperature Temper atur e/ww
858
297
77
1078
0.55
5-5
0.01
770
293
78
985
39.0$
0.55
0.8
0.01
573
-------
TABLE 5
SYSTEM COST
(Millions of Dollars)
Gasifier
Desulfurization
Particulate Removal
Other Process Cost
Power System
Cost
Cost - $/k¥
Power Cost-mills/kWhr
Low-
Temperature
92
27
82
222
High- High-
Temperature Temperature/ww
92
23
27
92
23
21.5
hk .
2kl
^33
iK}2
19 = 5
82
227
x*
1^30
21.2
574
-------
Integrated System With High Temperature Cleanup Process
Fig. k shows a schematic of an integrated system with a high-temperature
cleanup process. The sulfur content is reduced to about 500 ppm in a CONOCO
(CONOCO Coal Development Center) half-calcined dolomite fluid-bed reactor which
operates at 1600 F. This system is one of several types of sulfur cleanup processes
which operate at high temperature that are currently in the pilot stage. In this
process, as in most high-temperature processes there is no mechanism for removal
of fuel-bound nitrogen compounds.
The high-temperature particulate removal system, shown schematically in
Fig. 5, consists of three separate units. The first unit would be a high effi-
ciency cyclone of a more or less conventional design, but constructed of high-
temperature alloys. Approximately 85$ of the particles above 20y could be
removed in this device. Following the cyclone, a series of small diameter cyclones
or multielones, also of high-temperature alloys would remove 90% of the remaining
particles above lOp. The 'final filter' would be selected from the various granular
bed, wire or ceramic mesh, or other devices briefly described in Table 5. These
devices would remove 99% of the residual particulate above 2y.
The actual performance of such devices has not been well established at the
temperatures and pressure of interest. Work at the Argonne National Laboratories
on pressurized fluid beds operating at 10 atm and 1600 F has shown that cyclones
followed by two stages of metal mesh achieved 99%+ removal of micron sized particles
from a slipstream. Similarly, work at CONOCO Coal Development Center indicates
that metallic mesh removed not only particulates, but that it could potentially
remove alkali metal vapor aerosols. It must be cautioned that the foregoing
results are based upon small-scale tests. It remains to be seen if these devices
can be scaled to the required sizes (~ 105 ACFM).
It is difficult to project costs for systems which have yet to leave the
pilot stage, but it would appear that high temperature particulate cleanup is not
an inexpensive problem. It was mentioned before that low-temperature particulate
removal costs were lost in overall cleanup systems. A 'grass roots' venturi or
packed bed scrubber costs under $5/ACFM(I0, since a great amount of gas/liquid
contact is necessary as part of the sulfur/ammonia removal, the costs assigned
to the particulate removal would be appreciably less.
High-temperature particulate removal, however, requires the use of high-
strength, high-temperature alloys for cyclone and pressure vessel construction.
Also, since most of the collection mechanisms for the smallest particles require
periodic cleaning, two systems are required if continuous overall system operation
is desired. As part of the Energy Conversion Alternatives System (EGAS) study
carried out by the General Electric Company and Westinghouse Electric, high-temperature
575
-------
CLEAN GAS TO BURNER-
iT
DOLOMITE
ABSORBER
DOLOMITE
MAKE-UP—»•
H2S
DOLOMITE SLUDGE
STEAM
\7
DOLOMITE
REGEN
C02
SEPARATOR
PROCESS WATER
CONDENSATE
Figure 4: BCR/Conoco System
-------
CYCLONE
MULTICLONE
FINAL FILTER
FROM
GAS1FIER
85%
-20M
99% +
TOGAS
TURBINE
Figure 5: High—Temperature Particulate Removal System
-------
cleanup systems (cyclones, multielones and granular bed filters) vere investigated
and costs projected.(5,6) The equipment along with the necessary installation,
auxiliaries, piping, etc. was estimated to cost in the order of $200/ACFM, or,
for a nominal 10QO-MW plant over $20,000,000 some 5-7$ of total plant cost. As
part of an EPRI-sponsored study, Stone and Webster has indicated that a system of
cyclones and metallic mesh filter would be in the same cost neighborhood.^''
The effect of these costs are reflected in the cost of power from the high-
temperature systems as shown in Table U.
The Cost of NOX Control
The estimated HOX emission from the example integrated power plant using high-
temperature sulfur and particulate control does not meet the EPA regulations
(Table k). The problem is two-fold: (l) HOX due to thermal effects and (2) N0x
from fuel-bound nitrogen. The high fuel temperature (l600 F) precludes the use
of premix burners; i.e., burners in which the fuel and combustion air are inti-
mately mixed prior to entrance into the combustion chamber. Premixing could
reduce thermal NOX by 10% or more.'"' The fuel bound nitrogen is converted to
NOX during combustion on essentially a mol to mol basis. As can be seen in Table
k, over 5 lb/10^ Btu is projected.
To reduce the NOX and also to remove particulates and alkali metals, a
third system has been included (see Tables k and 5). In this system a water wash
is used subsequent to the high—temperature sulfur removal. The inclusion of the
water wash allows the attainment of the EPA goals, but only at a performance and
economic penalty.
RESEARCH AND DEVELOPMENT NEEDS
Without a better definition of both the particulate conte.nt of the fuel gas
leaving the gasifier and the ability of the turbine to withstand particulate
injestion, it is difficult to identify in any more depth the requirement for
particulate cleanup. Thus, research in these areas is of first priority.
In addition, work on high temperature cleanup systems currently in the pilot
stage should be reviewed with the better problem definition in mind, and those
devices which show continued promise should be scaled to larger sizes.
Although the fuel bound nitrogen problem is associated with only a few of the
gasifier types, it would be desirable to develop methods of removing nitrogen
compounds at high temperatures.
578
-------
CONCLUDING REMARKS
The potential advantages of high-temperature sulfur and particular cleanup;
e.g., an approximately 10% lower electricity cost than a comparable power system
with low-temperature cleanup, make such processes of great interest. However,
problems associated with meeting the NO^ emission standards due to the fuel-bound
nitrogen in the off gas from some gasifiers, may limit the application of high-
temperature cleanup systems.
Much of the work on particulate removal devices which would operate at the
1600 F level is directly applicable to pressurized-fluid bed systems. Thus, there
is a dual incentive to continue efforts in this area.
579
-------
REFERENCES
1. Robson, F. L., et. al.: Fuel Gas Environmental Impact: Phase Report. EPA-
600/2-75-078, November 1975-
2. Vogel, G. J., et. al.: Reduction of Atmospheric Pollution by the Application
of Fluidized-Bed Combustion and Regeneration of Sulfur Containing Additives.
EPA-650/2-7U-10H. September 197^.
3. Private Communication from M. Pell, CONOCO to F. Robson, UTRC. September 18,
1975.
U. Robson, F. L., et. al.: Analysis of Jet Engine Test Cell Pollution Abatement
Methods. AFWL-TR-73-18, May 1973.
5. Gorman, J., et. al.: Study of Advanced Energy Conversion Techniques for
Utility Applications Using Coal or Coal Derived Fuels. NASA Contract NAS-
3-19^06. GE Oral Briefing February 1976.
6. Hamm, R., et. al.: Energy Conversion Alternatives Study. Westinghouse Oral
Briefing February 1976.
7. Private Communication from C. Jones, Stone & Webster to F. Robson, UTRC.
April 20, 1976.
8. Robson, F. L., et. al.: Fuel Gas Environmental Impact: Phase k Report.
To be Published.
580
-------
TECHNICAL REPORT DATA
• (Please rfad Instructions on the reverse before completing/
IEPORT NO.
EPA-600/7-76-010
2.
4. TITLE AND SUBTITLE
IYMPOSIUM ON PARTICULATE CONTROL IN
ENERGY PROCESSES
3. RECIPIENT'S ACCESSION NO.
5. REPORT DATE
September 1976
6. PERFORMING ORGANIZATION CODE
. AUTHORIS)
8. PERFORMING ORGANIZATION REPORT NO
David E. Blake, Editor
9. PERFORMING ORGANIZATION NAME AND ADDRESS
Aerotherm Division/Acurex Corporation
485 Clyde Avenue
Mountain View, California 94042
10. PROGRAM ELEMENT NO.
EHE624
11. CONTRACT/GRANT NO.
68-02-1318, Task 22
12. SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
Industrial Environmental Research Laboratory
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Proceedings; 10/75-7/76
14. SPONSORING AGENCY CODE
EPA-ORD
15. SUPPLEMENTARY NOTES IERL_RTp
Drop 61, 919/549-8411 Ext 2925.
officer
report is D. C. Drehmel, Mail
The proceedings are a compilation of papers presented at a symposium,
cosponsored by EPA's Industrial Environmental Research Laboratory (Research
Triangle Park) and the Electric Power Research Institute, and featuring speakers
representing the utility industry, research organizations, and government agencies.
Purpose of the symposium was to examine the current state of particulate control
technology for energy processes, to discuss practical solutions to problems with
particulate from Western U.S. coals, and to consider recent progress in high-
temperature/high-pressure energy processes. Control of particulate emissions from
conventional power generation is well developed and reasonably effective. However,
some recent technological developments--increasing use of low-sulfur, high-ash
Western U.S. coals, and new high-temperature/high-pressure processes--have
increased the difficulty of achieving effective particulate control.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.IDENTIFIERS/OPEN ENDED TERMS
c. COSATl Field/Group
Air Pollution
Dust
Energy Conversion Techniques
oal
Air Pollution Control
Stationary Sources
Particulate
Western U.S. Coal
High-Temperature, High
Pressure Processes
13B
11G
10A
2 ID
3. DISTRIBUTION STATEMENT
Unlimited
19. SECURITY CLASS (This Report)
Unclassified
21. NO. OF PAGES
584
20. SECURITY CLASS (This page)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
581
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