United States
Environmental Protection
Agency
Research and Development
Industrial Environmental Research
Laboratory
Research Triangle Pari NC 27711
EPA-600/9-80 039c
September 1 980
Second
Symposium on the
Transfer and
Utilization of
Particulate Control
Technology
Volume III.
Particulate Control
Devices
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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 nine series. These nine broad cate-
gories were established to facilitate further development and application of en-
vironmental technology Elimination of traditional grouping was consciously
planned to foster technology transfer and a maximum interface in related fields.
The nine 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
8. "Special" Reports
9. Miscellaneous Reports
This document is available to the public through the National Technical Informa-
tion Service, Springfield, Virginia 22161.
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EPA-600/9-80-039c
September 1980
SECOND SYMPOSIUM ON THE
TRANSFER AND UTILIZATION OF
PARTICULATE CONTROL TECHNOLOGY
VOLUME III. PARTICULATE CONTROL DEVICES
by
F.P. Veriditti, J.A. Armstrong, and Michael Durham
Denver Research Institute
P.O. Box 10127
Denver, Colorado 80210
Grant Number: R805725
Project Officer
Dennis C. Drehmel
Office of Energy, Minerals, and Industry
Industrial Environmental Research Laboratory
Research Triangle Park, NC 27711
INDUSTRIAL ENVIRONMENTAL RESEARCH LABORATORY
OFFICE OF RESEARCH AND DEVELOPMENT
U.S. ENVIRONMENTAL PROTECTION AGENCY
RESEARCH TRIANGLE PARK, NC 27711
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DISCLAIMER
This report has been reviewed by the Industrial Environmental Research
Laboratory-Research Triangle Park, North Carolina, Office of Research and
Development, U.S. Environmental Protection Agency, and approved for publi-
cation. Approval does not signify that the contents necessarily reflect
the views and policies of the U.S. Environmental Protection Agency, nor
does mention of trade names or commercial products constitute endorsement
or recommendation for use.
ii
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ABSTRACT
The papers in these four volumes of Proceedings were pre-
sented at the Second Symposium on the Transfer and Utilization of
Particulate Control Technology held in Denver, Colorado during 23
July through 27 July 1979, sponsored by the Particulate Technology
Branch of the Industrial Environmental Research Laboratory of the
Environmental Protection Agency and hosted by the Denver Research
Institute of the University of Denver.
The purpose of the symposium was to bring together research-
ers, manufacturers, users, government agencies, educators and
students to discuss new technology and to provide an effective
means for the transfer of this technology out of the laboratories and
into the hands of the users.
The three major categories of control technologies - electrostatic
precipitators, scrubbers, and fabric filters - were the major concern
of the symposium. These technologies were discussed from the
perspectives of economics; new technical advancements in science and
engineering; and applications. Several papers dealt with combina-
tions of devices and technologies, leading to a concept of using a
systems approach to particulate control rather than device control.
Additional topic areas included novel control devices, high
temperature/high pressure applications, fugitive emissions, and
measurement techniques.
These proceedings are divided into four volumes, each volume
containing a set of related session topics to provide easy access to a
unified technology area.
iii
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CONTENTS
Page
VOLUME I CONTENTS
VOLUME II CONTENTS
VOLUME IV CONTENTS
vi
ix
xii
Section A - Scrubbers
Page
FLUX FORCE/CONDENSATION SCRUBBER DEMONSTRATION
PLANT IN THE IRON AND STEEL INDUSTRY 1
R. Chmielewski, S. Bhutra, S. Calvert, D.L. Harmon, J.H. Abbott
COLLECTION CHARACTERISTICS OF A DOUBLE STAGE SCRUBBER
TO ELIMINATE THE PAINT MIST FROM A SPRAY BOOTH .... 16
T. Isoda and T. Azuma
APPLICATION OF SLIPSTREAMED AIR POLLUTION CONTROL
DEVICES ON WASTE-AS-FUEL PROCESSES 25
F.D. Hall, J.M. Bruck, D.N. Albrinck and R.A. Olexsey
EVALUATION OF THE CEILCOTE IONIZING WET SCRUBBER ... 39
D.S. Ensor and D.L. Harmon
DEMONSTRATION OF A HIGH FIELD ELECTROSTATICALLY
ENHANCED VENTURI SCRUBBER ON A MAGNESIUM FURNACE
FUME EMISSION 61
M.T. Kearns and D.L. Harmon
DROPLET REMOVAL EFFICIENCY AND SPECIFIC CARRYOVER
FOR LIQUID ENTRAINMENT SEPARATORS 81
J.H. Gavin and F.W. Hoffman
AN EVALUATION OF GRID ROD FAILURE IN A MOBILE
BED SCRUBBER 95
J.S. Kinsey and S. Rohde
OPERATION AND MAINTENANCE OF A PARTICULATE SCRUBBER
SYSTEM'S ANCILLARY COMPONENTS 104
P.A. Czuchra
LOWERING OPERATING COSTS WHILE INCREASING THROUGHPUT
AND EFFICIENCY OF REACTORS AND SCRUBBERS 117
R.P. Tennyson, S.F. Roe, Jr. and R.H. Lace, Sr.
OPTIMIZING VENTURI SCRUBBER PERFORMANCE THROUGH
MODELING 127
D.W. Cooper
THE IMPACT OF HUMIDIFICATION CHAMBER PHYSICS ON
WET GAS CLEANUP SYSTEMS 145
D.P. Bloomfield, M.L. Finson, G.A. Simons and K.L. Wray
V
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Page
IMPROVING THE EFFICIENCY OF FREE-JET SCRUBBERS 162
D.A. Mitchell
Section B - Fabric Filters
HIGH VELOCITY FIBROUS FILTRATION 171
M.J. Ellenbecker, J.M. Price, D. Leith and M.W. First
THE EFFECT OF DUST RETENTION ON PRESSURE DROP IN
A HIGH VELOCITY PULSE-JET FABRIC FILTER 190
M.J. Ellenbecker and D. Leith
ROLE OF FILTER STRUCTURE AND ELECTROSTATICS
IN DUST CAKE FORMATION 209
G.E.R. Lamb and P.A. Costanza
PRESSURE DROP IN ELECTROSTATIC FABRIC FILTRATION .... 222
T. Ariman and D.J. Helfritch
EXPERIMENTAL ADVANCES ON FABRIC FILTRATION TECHNOLOGY
IN JAPAN - EFFECTS OF CORONA PRECHARGER AND RELATIVE
HUMIDITY ON FILTER PERFORMANCE 237
K. Iinoya and Y. Mori
BAGHOUSE OPERATING EXPERIENCE ON A NO. 6
OIL-FIRED BOILER 251
D.W. Rolschau
NEW FABRIC FILTER CONCEPT PROVEN MORE FLEXIBLE
IN DESIGN, EASIER TO MAINTAIN, AND UNSURPASSED
FILTRATION 260
B. Carlsson and R.J. Labbe
EPRI'S FABRIC FILTER TEST MODULE PROGRAM: A REVIEW
AND PROGRESS REPORT 270
R.C. Carr and J. Ebrey
Section C - Granular Beds
ELECTROSTATIC ENHANCEMENT OF MOVING-BED
GRANULAR FILTRATION 289
D.S. Grace, J.L. Guillory and F.M. Placer
ELECTRICAL AUGMENTATION OF GRANULAR BED FILTERS .... 309
S.A. Self, R.H. Cross and R.H. Eustis
vi
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Page
THEORETICAL AND EXPERIMENTAL FILTRATION EFFICIENCIES
IN ELECTROSTATICALLY AUGMENTED GRANULAR BEDS 344
G.A. Kallio, P.W. Dietz and C. Gutfinger
AEROSOL FILTRATION BY A CONCURRENT MOVING
GRANULAR BED: DESIGN AND PERFORMANCE 363
T.W. Kalinowski and D. Leith
DEEP BED PARTICULATE FILTRATION USING THE
PURITREAT (TM) PROCESS 382
L.C. Hardison
Section D - Novel Devices
PILOT-SCALE FIELD TESTS OF HIGH GRADIENT
MAGNETIC FILTRATION 404
C.H. Gooding and C.A. Pareja
EXPERIENCES WITH CONTROL SYSTEMS USING A UNIQUE
PATENTED STRUCTURE 416
G.C. Pedersen
ELECTROSTATIC EFFECTS IN VORTICAL FLOWS 429
P.W. Dietz
CONDENSATIONAL ENLARGEMENT AS A SUPPLEMENT TO
PARTICLE CONTROL TECHNOLOGIES 439
J.T. Brown, Jr.
Section E - Specific Applications
WELDING FUME AND HEAT RECOVERY - THE PROBLEM,
THE SOLUTION, THE BENEFITS 448
R.C. Larson
PARTICULATE REMOVAL CONSIDERATIONS IN SOLVENT
EMISSION CONTROL INSTALLATIONS 472
E.A. Brackbill and P.W. Kalika
ARSENIC EMISSIONS AND CONTROL TECHNOLOGY - GOLD
ROASTING OPERATIONS 484
J.O. Burckle, G.H. Marchant and R.L. Meek
CONTROL OF SALT LADEN PARTICULATE EMISSIONS FROM
HOGGED FUEL BOILERS 508
M.F. Szabo, R.W. Gerstle and L. Sims
AUTHOR INDEX 526
vii
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VOLUME I
CONTROL OF EMISSIONS FROM COAL FIRED BOILERS
Section A - Electrostatic Precipitators
COST AND PERFORMANCE OF PARTICULATE CONTROL
DEVICES FOR LOW-SULFUR WESTERN COALS 1
R.A. Chapman, D.P. Clements, L.E. Sparks and J.H. Abbott
CRITERIA FOR DESIGNING ELECTROSTATIC PRECIPITATORS ... 15
K. Darby
EVALUATION OF THE GEORGE NEAL ELECTROSTATIC
PRECIPITATOR 35
R.C. Carr
EPA MOBILE ESP HOT-SIDE PERFORMANCE EVALUATION .... 56
S.P. Schliesser, S. Malani, C.L. Stanley and L. E. Sparks
PRECIPITATOR UPGRADING AND FUEL CONTROL PROGRAM
FOR PARTICULATE COMPLIANCE AT PENNSYLVANIA
POWER & LIGHT COMPANY 80
J.T. Guiffre
MODIFICATION OF EXISTING PRECIPITATORS TO RESPOND TO
FUEL CHANGES AND CURRENT EMISSION REGULATIONS 100
D.S. Kelly and R.D. Frame
PERFORMANCE OF ELECTROSTATIC PRECIPITATORS WITH
LOAD VARIATION 117
W.T. Langan, G. Gogola and E.A. Samuel
FLY ASH CONDITIONING BY CO-PRECIPITATION WITH
SODIUM CARBONATE 132
J.P. Gooch, R.E. Bickelhaupt and L.E. Sparks
PREDICTING FLY ASH RESISTIVITY - AN EVALUATION 154
R.E. Bickelhaupt and L.E. Sparks
S03 CONDITIONING FOR IMPROVED ELECTROSTATIC PRECIPITATOR
PERFORMANCE OPERATING ON LOW SULFUR COAL 170
J.J. Ferrigan, III and J. Roehr
DOES SULPHUR IN COAL DOMINATE FLYASH COLLECTION IN
ELECTROSTATIC PRECIPITATORS? 184
E.C. Potter and C.A.J. Paulson
ANALYSIS OF THERMAL DECOMPOSITION PRODUCTS OF FLUE
GAS CONDITIONING AGENTS 202
R.B. Spafford, H.K. Dillon, E.B. Dismukes and L.E. Sparks
viii
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VOLUME I CONTENTS (Cont.)
Page
BIOTOXICITY OF FLY ASH PARTICULATE 224
A.R. Kolber, T.J. Wolff, J. Abbott and L. E. Sparks
Section B - Fabric Filters
FABRIC FILTERS VERSUS ELECTROSTATIC PRECIPITATORS ... 243
E.W. Stenby, R.W. Scheck, S.D. Severson, F.A. Horney
and D.P. Teixeira
DESIGN AND CONSTRUCTION OF BAGHOUSES FOR
SHAWNEE STEAM PLANT 263
J.A. Hudson, L.A. Thaxton, H.D. Ferguson, Jr., and N. Clay
OPERATING CHARACTERISTICS OF A FABRIC FILTER ON A
PEAKING/CYCLING BOILER WITHOUT AUXILIARY PREHEAT
OR REHEAT 297
W. Smit and K. Spitzer
OBJECTIVES AND STATUS OF FABRIC FILTER
PERFORMANCE STUDY 317
K.L. Ladd, R. Chambers, S. Kunka and D. Harmon
START-UP AND INITIAL OPERATIONAL EXPERIENCE ON A 400,000
ACFM BAGHOUSE ON CITY OF COLORADO SPRINGS' MARTIN DRAKE
UNIT NO. 6 342
R.L. Ostop and J.M. Urich, Jr.
DESIGN, OPERATION, AND PERFORMANCE TESTING
OF THE CAMEO NO. 1 UNIT FABRIC FILTER 351
H.G. Brines
EXPERIENCE AT COORS WITH FABRIC FILTERS - FIRING
PULVERIZED WESTERN COAL 359
G.L. Pearson
FABRIC FILTER EXPERIENCE AT WAYNESBORO 372
W.R. Marcotte
A NEW TECHNIQUE FOR DRY REMOVAL OF S02 390
C.C. Shale and G.W. Stewart
SPRAY DRYER/BAGHOUSE SYSTEM FOR PARTICULATE AND
SULFUR DIOXIDE CONTROL, EFFECTS OF DEW POINT, COAL
AND PLANT OPERATING CONDITIONS 410
W.R. Lane
SELECTION, PREPARATION AND DISPOSAL OF SODIUM COMPOUNDS
FOR DRY SOx SCRUBBERS 425
D.A. Furlong, R.L. Ostop and D.C. Drehmel
ix
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VOLUME I CONTENTS (Cont.)
Page
HIGH VELOCITY FABRIC FILTRATION FOR CONTROL OF
COAL-FIRED BOILERS 432
J.C. My cock, R.A. Gibson and J.M. Foster
EPA MOBILE FABRIC FILTER - PILOT INVESTIGATION OF
HARRINGTON STATION PRESSURE DROP DIFFICULTIES 453
W.O. Lipscomb, S.P. Schliesser and V.S. Malani
PASSIVE ELECTROSTATIC EFFECTS IN FABRIC FILTRATION ... 476
R.P. Donovan, J.H. Turner and J.H. Abbott
A WORKING MODEL FOR COAL FLY ASH FILTRATION 494
R. Dennis and H.A. Klemm
Section C - Scrubbers
PARTICULATE REMOVAL AND OPACITY USING A WET VENTURI
SCRUBBER - THE MINNESOTA POWER AND LIGHT EXPERIENCE . . 513
D. Nixon and C. Johnson
PERFORMANCE OF ENVIRONMENTALLY APPROVED NLA
SCRUBBER FOR SOz 529
J.A. Bacchetti
DESIGN GUIDELINES FOR AN OPTIMUM SCRUBBER SYSTEM. ... 538
M.B. Ranade, E.R. Kashdan and D.L. Harmon
TESTS ON UW ELECTROSTATIC SCRUBBER FOR PARTICULATE AND
SULFUR DIOXIDE COLLECTION 561
M.J. Pilat
EPA MOBILE VENTURI SCRUBBER PERFORMANCE 570
S. Malani, S.P. Schliesser and W.O. Lipscomb
THE RESULTS OF A TWO-STAGE SCRUBBER/CHARGED
PARTICULATE SEPARATOR PILOT PROGRAM 591
J.R. Martin, K.W. Malki and N. Graves
AUTHOR INDEX 616
X
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VOLUME II
ELECTROSTATIC PRECIPITATORS
Section A - Fundamentals
Page
COLLECTION EFFICIENCY OF ELECTROSTATIC PRECIPITATORS
BY NUMERICAL SIMULATION 1
E.A. Samuel
THE EFFECTS OF CORONA ELECTRODE GEOMETRY ON THE
OPERATIONAL CHARACTERISTICS OF AN ESP 31
G. Rinard, D. Rugg, W. Patten and L.E. Sparks
THEORETICAL METHODS FOR PREDICTING ELECTRICAL CONDITIONS
IN WIRE-PLATE ELECTROSTATIC PRECIPITATORS 45
R.B. Mosley, J.R. McDonald and L.E. Sparks
LATERAL PROPAGATION OF BACK DISCHARGE 65
S. Masuda and S. Obata
THEORETICAL MODELS OF BACK CORONA AND
LABORATORY OBSERVATIONS 74
D.W. VanOsdell, P.A. Lawless and L.E. Sparks
CHARGE MEASUREMENTS ON INDIVIDUAL PARTICLES
EXITING LABORATORY PRECIPITATORS 93
J.R. McDonald, M.H. Anderson, R.B. Mosley and L.E. Sparks
OPTIMIZATION OF COLLECTION EFFICIENCY BY VARYING PLATE
SPACING WITHIN AN ELECTROSTATIC PRECIPITATOR 114
E.J. Eschbach and D.E. Stock
INTERACTION BETWEEN ELECTROSTATICS AND FLUID DYNAMICS
IN ELECTROSTATIC PRECIPITATORS 125
S. Bernstein and C.T. Crowe
PARTICLE TRANSPORT IN ELECTROSTATIC PRECIPITATORS ... 146
G. Leonard, M. Mitchner and S.A. Self
Section B - Operation and Maintenance
THE "HUMAN ELEMENT" - A PROBLEM IN OPERATING
PRECIPITATORS 168
W.J. Buchanan
ELECTROSTATIC PRECIPITATORS - ELECTRICAL PROBLEMS
AND SOLUTIONS 173
R.K. Raymond
xi
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VOLUME II CONTENTS (Cont.)
Page
ELECTRODE CLEANING SYSTEMS: OPTIMIZING RAPPING
ENERGY AND RAPPING CONTROL 189
M. Neundorfer
COMPOSITION OF PARTICULATES—SOME EFFECTS ON
PRECIPITATOR OPERATION 208
J.D. Roehr
INCREASING PRECIPITATOR RELIABILITY BY PROPER LOGGING
AND INTERPRETATION OF OPERATIONAL PARAMETERS - AN
OPERATORS GUIDE 219
P.P. Bibbo and P. Aa
ELECTROSTATIC PRECIPITATORS - START-UP, LOW LOAD,
CYCLING, AND MAINTENANCE CONSIDERATIONS 242
F.A. Wybenga and R.J. Batyko
ELECTROSTATIC PRECIPITATOR EMISSION AND OPACITY
PERFORMANCE CONTROL THRU RAPPER STRATEGY 256
W.T. Langan, J.H. Oscarson and S. Hassett
RAPPING SYSTEMS FOR COLLECTING SURFACES IN AN
ELECTROSTATIC PRECIPITATOR 279
H.L. Engelbrecht
LOW POWER ELECTROSTATIC PRECIPITATION - A LOGICAL
SOLUTION TO COLLECTION PROBLEMS EXPERIENCED WITH
HIGH RESISTIVITY PARTICULATE 296
J.H. Umberger
Section B - Advanced Design
HIGH INTENSITY IONIZER TECHNOLOGY APPLIED TO
RETROFIT ELECTROSTATIC PRECIPITATORS 314
C.M. Chang and A.I. Rimensberger
BOXER-CHARGER - A NOVEL CHARGING DEVICE FOR HIGH
RESISTIVITY DUSTS 334
S. Masuda and H. Nakatani
PRECIPITATOR ENERGIZATION UTILIZING AN ENERGY
CONSERVING PULSE GENERATOR 352
H.H. Petersen and P. Lausen
PRECHARGER COLLECTION SYSTEM - DESIGN FROM THE
LABORATORY THROUGH FIELD DEMONSTRATION 369
M. Nunn, D. Pontius, J.H. Abbott and L.E. Sparks
xii
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VOLUME II CONTENTS (Cont.)
Page
TOWARDS A MICROSCOPIC THEORY OF ELECTROSTATIC
PRECIPITATION 374
C.G. Noll and T. Yamamoto
ION CURRENT DENSITIES PRODUCED BY ENERGETIC ELECTRONS
IN ELECTROSTATIC PRECIPITATOR GEOMETRIES 391
W.C. Finney, L.C. Thanh and R.H. Davis
EXPERIMENTAL STUDIES IN THE ELECTROSTATIC
PRECIPITATION OF HIGH-RESISTIVITY PARTICULATE 399
J.C. Modla, R.H. Leiby, T.W. Lugar, and K.E. Wolpert
PILOT PLANT TESTS OF AN ESP PRECEDED BY THE
EPA-SoRI PRECHARGER 417
L.E. Sparks, G.H. Ramsey, B.E. Daniel and J.H. Abbott
Section C - Industrial Applications
PILOT PLANT/FULL SCALE EP SYSTEM DESIGN AND
PERFORMANCE ON BOF APPLICATION 427
D. Ruth and D. Shilton
THE SELECTION AND OPERATION OF A NEW PRECIPITATOR
SYSTEM ON AN EXISTING BASIC OXYGEN FURNACE 441
D. Ruth and D. Shilton
CONTROL OF FINE PARTICLE EMISSIONS WITH WET
ELECTROSTATIC PRECIPITATION 452
S.A. Jaasund
TUBULAR ELECTROSTATIC PRECIPITATORS OF TWO
STAGE DESIGN 469
H. Surati, M.R. Beltran and I. Raigorodsky
PRESENT STATUS OF WIDE-SPACING TYPE PRECIPITATOR .
IN JAPAN 483
S. Masuda
LOW FREQUENCY SONIC CLEANING APPLIED TO
ELECTROSTATIC PRECIPITATORS 502
S.B. Smith and J.A. Schwartz
AUTHOR INDEX 514
xiii
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VOLUME IV
SPECIAL APPLICATIONS FOR AIR POLLUTION
MEASUREMENT AND CONTROL
Section A - High Temperature High Pressure Applications
Page
FUNDAMENTAL PARTICLE COLLECTION AT HIGH
TEMPERATURE AND PRESSURE 1
R. Parker, S. Calvert, D.C. Drehmel and J.H. Abbott
PARTICULATE COLLECTION IN A HIGH TEMPERATURE CYCLONE . . 14
K.C. Tsao, C.O. Jen and K.T. Yung
EVALUATION OF A CYCLONIC TYPE DUST COLLECTOR FOR HIGH
TEMPERATURE HIGH PRESSURE PARTICULATE CONTROL .... 30
M. Ernst, R.C. Hoke, V.J. Siminski, J.D. McCain,
R. Parker and D.C. Drehmel
CERAMIC FILTER TESTS AT THE EPA/EXXON PFBC MINI PLANT . . 42
M. Ernst and M.A. Shackleton
HOT GAS CLEAN-UP BY GLASS ENTRAINMENT OF
COMBUSTION BY-PRODUCTS 64
W. Fedarko, A. Gatti and L.R. McCreight
THE A.P.T. PxP DRY SCRUBBER FOR HIGH TEMPERATURE AND
PRESSURE PARTICULATE CONTROL 84
R.G. Patterson, S. Calvert and M. Taheri
GAS CLEANING UNDER EXTREME CONDITIONS OF
TEMPERATURE AND PRESSURE 98
E. Weber, K. Hu'bner, H.G. Pape and R. Schulz
PROGRESS ON ELECTROSTATIC PRECIPITATORS FOR USE
AT HIGH TEMPERATURE AND HIGH PRESSURE 126
G. Rinard, D. Rugg, R. Gyepes and J. Armstrong
REDUCTION OF PARTICULATE CARRYOVER FROM A
PRESSURIZED FLUIDIZED BED 136
R.W. Patch
COMPARATIVE ECONOMIC ANALYSIS OF SELECTED PARTICULATE
CONTROL SYSTEMS FOR ADVANCED COMBINED CYCLE POWER
PLANTS 154
J.R. Bush, F.L. Blum and P.L. Feldman
CONCLUSIONS FROM EPA'S HIGH TEMPERATURE/HIGH
PRESSURE CONTROL PROGRAM 170
D.C. Drehmel and J.H. Abbott
xiv
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VOLUME IV CONTENTS (Cont.)
Section B - Fugitive Emissions
Page
WATER SPRAY CONTROL OF FUGITIVE PARTICULATES: ENERGY
AND UTILITY REQUIREMENTS 182
D.P. Daugherty, D.W. Coy and D.C. Drehmel
THE CONTROL OF DUST USING CHARGED WATER FOGS 201
S.A. Hoenig
SPRAY CHARGING AND TRAPPING SCRUBBER FOR FUGITIVE
PARTICLE EMISSION CONTROL 217
S. Yung, S. Calvert, and D.C. Drehmel
CONTROL OF WINDBLOWN DUST FROM STORAGE PILES 240
C. Cowherd, Jr.
THE CONTRIBUTION OF OPEN SOURCES TO AMBIENT
TSP LEVELS 252
J.S. Evans and D.W. Cooper
FUTURE AREAS OF INVESTIGATION REGARDING THE
PROBLEM OF URBAN ROAD DUST 274
E.T. Brookman and D.C. Drehmel
STATUS OF CONNECTICUT'S CONTROL PROGRAM FOR
TRANSPORTATION-RELATED PARTICULATE EMISSIONS 291
J.H. Gastler and H.L. Chamberlain
NEW CONCEPTS FOR CONTROL OF FUGITIVE PARTICLE
EMISSIONS FROM UNPAVED ROADS 312
T.R. Blackwood and D.C. Drehmel
DEVELOPMENT OF A SAMPLING TRAIN FOR THE ASSESSMENT
OF PARTICULATE FUGITIVE EMISSIONS 321
R.L. Severance and H.J. Kolnsberg
SECONDARY NEGATIVE ELECTRON BOMBARDMENT FOR
PARTICULATE CONTROL 333
W.E. Stock
Section C - Measurement and Analysis
HIGH TEMPERATURE AND HIGH PRESSURE SAMPLING DEVICE
USED FOR PARTICULATE CHARACTERIZATION OF A FLUIDIZED
BED COAL GASIFICATION PROCESS 338
S.P. Tendulkar, J. Pavel and P. Cherish
ON-STREAM MEASUREMENT OF PARTICULATE SIZE
AND LOADING 351
E. S. VanValkenburg xv
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VOLUME IV CONTENTS (Cont. )
Page
ANALYSIS OF SAMPLING REQUIREMENTS FOR CYCLONE
OUTLETS 368
M.D. Durham and D.A. Lundgren
ELECTROSTATIC EFFECTS ON SAMPLING THROUGH
UNGROUNDED PROBES 387
W.B. Giles and P.W. Dietz
OPTICAL PARTICULATE SIZE MEASUREMENTS USING A
SMALL-ANGLE NEAR-FORWARD SCATTERING TECHNIQUE .... 396
J.C.F. Wang
IN-STACK PLUME OPACITY FROM ELECTROSTATIC
PRECIPITATOR SCRUBBER SYSTEMS 411
L.E. Sparks, G.H. Ramsey and B.E. Daniel
TI-59 PROGRAMMABLE CALCULATOR PROGRAMS FOR
IN-STACK OPACITY 424
S.J. Cowen, D.S. Ensor and L.E. Sparks
UTILIZATION OF THE OMEGA-1 LIDAR IN EPA
ENFORCEMENT MONITORING 443
A.W. Dybdahl and F.S. Mills
EFFECTS OF PARTICLE-CONTROL DEVICES ON ATMOSPHERIC
EMISSIONS OF MINOR AND TRACE ELEMENTS FROM COAL
COMBUSTION 454
J.M. Ondov and A.H. Biermann
A SOURCE IDENTIFICATION TECHNIQUE FOR AMBIENT
AIR PARTICULATE 486
E.J. Fasiska, P.B. Janocko and D.A. Crawford
PARTICLE SIZE MEASUREMENTS OF AUTOMOTIVE
DIESEL EMISSIONS 496
J.D. McCain, and D. Drehmel
CONTROL STRATEGIES FOR PARTICULATE EMISSIONS FROM
VEHICULAR DIESEL EXHAUST 508
M.G. Faulkner, J.P. Gooch, J.R. McDonald,
J.H. Abbott and D.C. Drehmel
AN EVALUATION OF THE CYTOTOXICITY AND MUTAGENICITY OF
ENVIRONMENTAL PARTICULATES IN THE CHO/HGPRT SYSTEM . . 524
N.E. Garrett, G.M. Chescheir, III, N.A. Custer, J.D. Shelburne,
Catherine R. De Vries, J.L. Huisingh and M.D. Waters
AUTHOR INDEX 536
xvi
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FLUX FORCE/CONDENSATION SCRUBBER
DEMONSTRATION PLANT IN THE
IRON AND STEEL INDUSTRY
by
Richard Chmielewski, Sudarshan Bhutra, and Seymour Calvert
Air Pollution Technology, Inc.
San Diego, California 92117
and
Dale L. Harmon and James A. Abbott
Industrial Environmental Research Laboratory
U.S. Environmental Protection Agency
Research Triangle Park, North Carolina 27711
ABSTRACT
A flux force/condensation scrubber was designed, built, and operated
by Air Pollution Technology, Inc. to control the particulate emissions from
an iron and steel melting cupola. The cupola had a production rate of
approximately 14,000 kq/hr (15 tons/hr). The entire exhaust gas flow of
8.5 kg/s (14,000 dscfm) was cleaned in the scrubber system. The major
components included an afterburner, spray saturator, condenser, scrubber,
and fan. A cooling tower was used to reject heat from the system and a
parallel plate settler was used to separate solids from the circulation
water.
Flux force/condensation scrubbing technology has been extensively
developed by A.P.T. under EPA sponsorship. The flux forces of thermo-
phoresis and diffusiophoresis and the condensation of water vapor on the
particles are used to improve the collection efficiency for fine particles
and reduce the energy requirements.
Performance data were obtained by sampling for particulates at
several locations. Several operating conditions were tested and experi-
mental results were compared with theoretical predictions of particle
collection. The performance of the flux force/condensation scrubbing
system was evaluated and compared to high energy gas atomized scrubbing.
l
-------
FLUX FORCE/CONDENSATION SCRUBBER DEMONSTRATION
PLANT IN THE IRON AND STEEL INDUSTRY
INTRODUCTION
This paper describes a flux force/condensation (F/C) scrubber for con-
trolling emissions from an iron melting cupola. The air pollution control
system has been designed, built, and operated by Air Pollution Technology,
Inc. under contract by the EPA. The objectives of this project are to
verify the design, performance, and economics of F/C scrubbing.
F/C scrubbing uses the flux forces of diffusiophoresis and thermophoresis
to improve fine particle collection. Condensation of water vapor on the
particles and the subsequent increase in mass is the third mechanism which
improves performance.
This demonstration program is the most recent in a number of EPA contracts
to develop F/C scrubbing as a means of improving the fine particle collection
in wet scrubbers. Earlier work involved laboratory and pilot-scale projects
to verify the technical feasibility and develop design data and methods for
F/C scrubbing. This paper presents the results to date of a second demon-
stration plant which is operating on the total emissions from an iron melting
cupola. An earlier demonstration F/C system was designed and operated to con-
trol the emissions from a secondary metal recovery furnace.
DESCRIPTION OF SOURCE
The melting furnace is a No. 9, water-cooled cupola whose inside
diameter at the melting zone is approximately 180 cm. The melting capacity
of the cupola is 14,000 kg/hr (15 tons/hr). Figure 1 shows the process
flow diagram for the system. The raw materials for the cupola include
approximately 16,000 kg/hr scrap, 2,750 kg/hr coke and 500 kg/hr limestone.
The combustion air is supplied at a rate of 3.4 DNm /s (7,700 dscfm).
The uncontrolled emission factor for an iron melting cupola is 8.5 g
particulate per kg of metal charged. For the cupola at rated capacity
this corresponds to 136 kg particulate/hr. The emission limit is 14.4
kg/hr for this application where the process weight is 19 x 103 kg/hr.
The maximum allowable system penetration is 0.106.
2
-------
Figure 2 shows the measured particle size distribution for the emissions
from the cupola. These size data were obtained using University of Washing-
ton cascade impactors. The mass median particle diameter is about 0.9 umA.
The measurements were made at the inlet to the condenser vessel.
PERFORMANCE REQUIREMENTS
The pressure drop required to achieve a penetration of 0.10 for a high
energy gas atomized scrubber is predicted to be approximately 100 cm W.C.
This calculation was performed using the particle size distribution shown
in Figure 2. The existing air pollution control system at the foundry was
capable of providing a scrubber pressure drop in excess of 100 cm W.C.
The F/C scrubber system is designed to minimize the total scrubber sys-
tem pressure drop and thereby reduce power requirements. Collection of
fine particulate using flux forces and growth due to condensation reduce
the pressure drop required by the scrubber system.
Figure 3 shows the predicted "grown" particle size distribution
leaving the condenser. The grown particle size distribution will depend
upon the condensation ratio achieved and the particle number concentration.
The condensation ratio is defined as the mass of water condensed per unit
mass of dry gas. The particle number concentration is not measured directly
by the cascade impactor, however it can be calculated using the particle
mass size distribution and the mass concentration. Calculated number con-
centrations range from 10B to 109 particles/cm3.
The predicted scrubber pressure drop required to obtain the desired pene-
tration of 0.10 based on the grown size distribution shown in Figure 3 is
40 cm. W.C. An additional 13 cm W.C. is required for the condenser pres-
sure drop. The total pressure drop required of the F/C scrubber system for
this case is 53 cm W.C. versus 100 cm W.C. for a high energy gas atomized
scrubber. This pressure drop saving would translate to a 40 percent reduction
in fan power requirement.
PROCESS DESCRIPTION
The exhaust gases from the cupola bed are mixed with additional combus-
tion air which enters through the charging door. The gases then enter an
afterburner tank which ensures complete combustion of any carbon monoxide
produced. The combustion gases leave the afterburner at a maximum tempera-
ture of about 1,100°C. The exhaust gas is then cooled in the saturator by
water sprays and passes into the condenser at the saturation temperature of
about 75°C. Approximately 3.6 g/s of water is evaporated in the satura-
tion process.
3
-------
The saturated exhaust gases are cooled in the condenser by contacting
with cold water. It is in the condenser that the F/C effects are applied.
Water vapor is condensed and temperature and concentration gradients
established driving the flux forces.
The cooled gases pass into a gas atomized scrubber where the grown par-
ticles are collected before passing through the exhaust fan and out the stack.
The heat rejection loop of the system involves a pump to circulate the
hot water from the condenser through a cooling tower where the heat is
rejected to the atmosphere.
EQUIPMENT DESIGN
Table 1 summarizes the specifications for the major pieces of equipment.
Saturator
The function of the saturator is to cool the hot exhaust gases from the
afterburner to the adiabatic saturation temperature. The design chosen was
a horizontal tank with high pressure water sprays. Mass transfer calcula-
tions were performed on the evaporation process to ensure sufficient resi-
dence time for evaporation. The physical layout of the plant dictated that a
horizontal vessel be used. The material of construction is 316 stainless
steel. City water is supplied to the internal spray nozzles using a high
pressure centrifugal pump. The head developed by the saturator pump is
15 atm. The high pressure results in small spray drops which evaporate
more effectively than large drops. The co-current spray pattern minimizes
gas pressure drop across the saturator.
Condenser
The condenser is a key component of the F/C system. Water vapor is
condensed and temperature and concentration gradients established in this
part of the system. The design chosen was a counter-current packed bed
using 1.25 m of Intaiox packing having a nominal dimension of 5 cm. The
cooling water is distributed through a trough system across the top of
the packing. The gases pass from the saturator into the condenser. Flow
straighteners are used to direct the gas flow upward into the condenser.
Liquid drains from the bottom of the condenser into a sump tank adjacent
to the condenser. The material used for the condenser is 316 stainless steel.
Scrubber
The scrubber used to collect the grown particles is a gas atomized
scrubber having a variable area throat. The throat area can be adjusted
using a hydraulic cylinder from the instrument trailer or by manual
adjustment of the throat blades. The scrubber design incorporates an
integral liquid sump and entrainment separator.
4
-------
The scrubber entrainment separator is a tubebank separator made of 304
stainless steel tubes. The rest of the scrubber is made of 316 stainless
steel.
Cooling Tower
The function of the cooling tower is to reject the heat which is given
off by the exhaust gases as they are cooled in the condenser. Hot water
from the condenser is drained into a sump. The water is pumped from the
sump to a commercial-packaged cooling tower. The cooling tower is a stan-
dard design incorporating countercurrent air and water flows. The air
flow is provided by a set of three centrifugal fans in a forced draft
configuration.
Water Treatment
The water treatment for the F/C system involves two processes. Control
of water pH is required because of formation of sulfurous acid by absorp-
tion of sulfur dioxide from sulfur present in the coke. The pH of the water
is maintained above 6 by addition of soda ash.
The solids collected in the system amount to approximately 1,000 kg/day.
Particulate matter collects in the bottom of the afterburner tank and in the
bottom of the condenser. This material is removed manually from the after-
burner and pumped from the condenser using an air operated diaphragm pump.
Suspended solids are removed from the water system by use of a parallel
plate gravity settler. A side stream of 1.5 L/S is passed through the
settler where solids are separated from the water. A flocculent aid is
used to improve solids separation in the settler.
EQUIPMENT LAYOUT
The equipment layout is shown in Figure 4. The space available at the
site limited the equipment layout, requiring an innovative design to ensure
adequate access for testing and maintenance. The saturator runs north-
south with the center line approximately 2 m above grade.
The condenser vessel and scrubber are located on a concrete pad just
north of the existing water sump which is used in the new system. The
existing exhaust fan is being used and remains in its original location.
The ducting system has been designed with sampling requirements in mind.
The cooling tower is located on a platform above the condenser. This
design feature allows gravity flow of cooling water from the cooling tower
to the condenser eliminating one pump.
Figure 5 shows an alternate design with the scrubber in the forced draft
mode. This configuration will be tested in a later portion of the test
phase.
5
-------
TEST METHODS
The performance tests are being made using standard stack sampling
techniques. EPA Method 5 tests are made at the inlet to the condenser and
at the outlet of the scrubber to determine total mass efficiency of the F/C
system. Cascade impactor tests are also made at various locations to obtain
size penetration data.
A detailed test plan has been established which will yield performance
data at various operating conditions. The primary variable which will be
investigated is the condensation ratio achieved by varying the cooling
liquid flow rate and the scrubber pressure drop.
Performance Data
The tests have recently started and to date we have obtained performance
data on the collection efficiency in the condenser. The main mechanisms for
particle collection in the condenser are by inertia! impaction and by the
combined effects of thermophoresis and diffusiophoresis. Figures 6, 7, and
8 show the experimental penetrations and the predicted penetrations for
three different condensation ratios. Log-probability coordinates have
been used to facilitate the presentation of the data.
In each case the experimental penetration is lower than the predicted
penetrations. This is reasonable in view of the fact that the collection
prediction did not include that occurring in other parts of the system.
Testing is now under way to obtain data on the total system performance.
SUMMARY
A demonstration F/C scrubber system has been designed, built, and is
in operation on the emissions from an iron melting cupola. The system has
the potential of reducing the power requirements for control of particulate
emissions compared to high energy gas atomized scrubbers which are in use on
this type of source.
Results of tests on the performance of the condenser show that the experi-
mental penetration is better than predicted. Testing is now in progress to
determine the optimum performance of the F/C scrubber system.
6
-------
TABLE 1. SUMMARY OF EQUIPMENT DESIGN DETAILS
Component
Description
Size
Construction Materials
Saturator
Condenser
Scrubber
Exhaust Fan
Cooling Tower
Saturator Pump
Cooling Tower Pump
Scrubber Pump
Settler
Ducting
Piping
Horizontal Spray Tower -
APT Design
Counter Current Packed Tower -
APT Design
Gas Atomized Scrubber -
APT Design
Centrifugal Radial Blade -
Westinghouse
1.5 m x 10.5 m 316 Stainless Steel
2.75 ml x 6 m
2 m x 2 m
316 Stainless Steel with
Stainless Packing
316 Stainless Steel
Wheel Dia.1.6 m Carbon Steel
Baltimore Air Coil VST 400 3.7 m x 2.7 m x 5.5 m Epoxy Coated Galvanized
54 Z/s @ 15 atm Carbon Steel
All is Chalmers Two-Stage
Centrifugal
Peabody - Vertical Turbine
Carver Single-Stage
Centrifugal
Parkson-Lamella
754 i/s @ 2 atm Cast Iron
254 Si/s @ 1 atm Cast Iron
Settling Area 12 m2 Carbon Steel Shell
PVC Plates
0.8 m4> & 0.9 n*j> FRP
5 cm, 10 cm, 20 cm FRP
-------
COOLING TOWER
/ \
CONDENSER
FAN
AFTER-
BURNER
CUPOLA
SCRUBBER
WATER
SATURATOR
WATER
Figure 1. F/C demonstration system flow diagram.
-------
10.0
to
<
E
a.
DC
111
H-
111
2
<
O
O
>-
Q
O
CC
Ul
<
5.0
4.0
3.0
2.0
1.0
0.5
0.4
0.3
5 10 20 30 40 50 60 70 80 90 95 98
CUMULATIVE MASS PERCENT
Figure 2. Particle size distribution at condenser inlet.
-------
<
E
a.
«
DC
U1
h
UJ
2
<
Q
O
>-
Q
O
DC
LU
<
5.0
4.0
3.0
2.0
1.0
T 1 1 I
CONDENSATION RATIO = 0.22 g/g
PARTICLE CONCENTRATION = 108 particles/cm3
GBOANN
0.5
0.4
0.3
x
x
j.
X
_L
X
10 20 30 40 50 60 70 80
CUMULATIVE MASS PERCENT
90 95 98 99
Figure 3. Predicted grown particle size distribution at condenser outlet.
-------
O' -C
Spbw Sgmecrog.-
ft ^ »J"T
&FM. OOROE.R.
«^.cLl0S-d
cur ^rrttt >wvrfc.u-vtioo
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-------
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-------
t/4
98
95
90
80
* 70
%
O 60
< 50
C ,
l- 40
U1
m 30
flL
20
10
5
11 1 1 II 1 1 1 1 • 1 1 1
PREDICTED PENETRATION FOR SUPERFICIAL GAS VELOCITY
= 200 cm/s
-
CONDENSATION RATIO =0.10 g/g
\ ^ **
-Vv \
-
\\
\\
\
¥
_1 1 1 _L _ _J 1 1
2
5
10
20
30 *
40 o
z
50 UJ
O
60 C
u.
70 w
80
90
98
0.5 1.0 2.0 3.0 4.0 5.0 10.0
AERODYNAMIC DIAMETER, pmA
Figure 6. Predicted arid experimental penetration for condenser.
-------
98
95
90
80
2 70
P 60
C 50
H
Ui 40
z
Ml
S 30
20
10
5
1—i i m 1 1 1 1—i—r-
PREDICTED PENETRATION FOR SUPERFICIAL GAS VELOCITY
= 200 cm/s
CONDENSATION RATIO = 0.15 g/g
\
2
5
10
20
30
40 ^
50 ^
HI
60 o
70 £
111
90
95
J L
j i_
98
0.5 1 0 2.0 3.0 4.0 5.0
AERODYNAMIC DIAMETER. jjmA
10.0
Figure 7. Predicted and experimental penetration for condenser.
-------
i i r I i 1 p 1 1—i i i i
PREDICTED PENETRATION FOR SUPERFICIAL GAS VELOCITY
= 200 cm/s
CONDENSATION RATIO = 0.22 g/g
i i
'C>
\
\
\
\
\
\
i i i
>-
o
5
10
20
30
40
50 __
UJ
60 5
70 i
UJ
80
90
95
98
1.0 2.0 3.0 4.0 5.0 10.0
AERODYNAMIC DIAMETER,
Figure 8. Predicted and experimental penetration for condenser.
-------
COLLECTION CHARACTERISTICS OF A DOUBLE STAGE SCRUBBER
TO ELIMINATE THE PAINT MIST FROM A SPRAY BOOTH
By:
Toru Isoda and Tadahiro Azuma
Osaka Prefectural Industrial Research Institute
Enokojima, Nishi-ku, Osaka City 550, Japan
ABSTRACT
A 4.5 m3/min laboratory paint spray booth was built adopting a double-
stage scrubber with heavy oil as the scrubbing liquid. The relationship
between the collection efficiency E and the pressure drop AP was studied using
a cylindrical paper tilter and a Digital Dust Counter. Experimental results
indicated that the collection efficiency in this system is in accordance with
the collection efficiency equation based on the theory of inertial impaction.
With water used as the scrubbing liquid, the factor K is about 0.5 for single-
stage scrubbing, while it is 0.6 for double-stage scrubbing.
At a pressure drop of 60 mm of H2O, heavy oil used as the scrubbing
liquid gained about 2.0 Z more in collection efficiency than water. The
zigzag baffler equipped for removing heavy oil mist was also effective for
the paint mist collector, and 99.1 % of E was gained at a pressure drop of
135 mm H2O.
16
-------
INTRODUCTION
A high performance scrubber is urgently required in order to eliminate
paint mist particles that result from treatment of solvent gas which Is
emitted from a spray booth by a catalytic combustor. An experimental study
of the collection characteristics of paint mist particles is carried out using
a new type double-stage scrubber which is equipped with a zigzag baffler. The
test is carried out using both water and heavy oil as the scrubbing liquids.
This paper presents the experimental results of the relationship between the
collection efficiency, pressure drop and fractional efficiency.
EXPERIMENTAL APPARATUS
Figure 1 is a schematic diagram of the experimental scrubber system,
which includes a zigzag baffler, a paint spray system, a recorder for the gas
flow rate and the pressure drop, an apparatus with a cylindrical paper filter
for the measurement of the particle concentration, and a cascade impactor
for the measurement of the particle size distribution.
The zigzag baffler has the function of separating the oil mist generated
from the scrubber section. Table 1 shows the operating condition in this
experiment, and Table 2 presents the property of the paint that was used in
this study.
The paint particle concentration in the both the outlet and the inlet of
the scrubber was measured on the basis of the paint in its solid form because
of its volatility.
EXPERIMENTAL RESULTS
Given in Figure 2 are the relations hips between the counting rate of
the Digital Dust Counter and the weight concentration of the solid paint
particles. Figure 3 shows the particle size distribution of the solid in
spray paint particles.
Collection Characteristics of the Double-Stage Scrubbing Apparatus and the
Value of K in the Equation: Pt ¦ exp (- K Vty L)
Shown in Figure 4 is a comparison of the collection characteristics
against the solid paint particles on both single and the double-stage
scrubbing. The collection efficiency for a venturi scrubber is given
approximately by,
Pt = exp (- K /~ij) L) (1)
Where, Pt ¦ penetration
¦ dimensionless parameter for collection by inertial impaction
L - liquid to gas ratio
K ¦ factor
The dimensionless inertial parameter Is expressed as,
17
-------
p d 2u
* = P P ° (2)
18yg6
where, pp = particle density
dp = particle diameter
uQ = velocity of the aerosol stream
Ug = gas viscosity
<5 = dro-let diameter
In a previous report, the factor K was reported as being from about 0.7 to
1.5w (2) in a venturi scrubber. The collection performance of the scrubber,
a kind of "self-induced spray wet collector" as is the one in this study, is
also approximately in agreement with the value given by eq.(l). In this
case the liquid to gas ratio,vL and the droplet diameter 6 in eq. (1), (2)
is given by the equation'-*'* as follows.
Ps = 0.1023 ug2L (3)
4980
6 = + 28.8 L (4)
where, Vr = gas velocity relative to the liqud. Futhermore, in each case of
single and double-stage scrubbing, the factor K is about 0.5 and 0.6
respectively under the operating condition which ranges between 65 to 90
mmH20 with a pressure drop AP.
Effects of the Heavy Oil as a Scrubbing Liquid on Collection Characteristics
Figure 5 gives the collection characteristics of the double-stage
scrubbing unit which utilized water and heavy oil as scrubbing liquids. The
heavy oil exerts an influence on the ability to collect paint particles to
some extent under the operating condition of a low pressure drop. It appears
that an increase in collection efficiency is attributed to the property of the
paint particle which is hydrophobic.
Zigzag Baffler for Collection of the Paint and Oil Mist Particles
Shown in Figure 6 is the comparison of the collection characteristics of
the paint particles between the double-stage scrubber equipped with a zigzag
baffler and the scrubber alone. It can be seen that the zigzag baffler
effectively acts as the collector of the paint particles which are
agglomerated by the oil mist particles generated from the scrubbing section.
Figure 7 gives the relationships between a pressure drop AP and oil mist
generation in the scrubber. Given in Figure 8 is the collection efficiency
of the zigzag baffler for the paint particles on various oil mist generations.
The collection efficiency of the paint particles increases as the oil mist
18
-------
generation is increased. Therefore, it appears that the oil film formed on
the surface of the battle plates and the agglomarations of the fine paint
particles affects the collection efficiency of the paint particles. Given
in Figure 9 is the collection efficiency of the oil mist particles due to the
zigzag baffler.
Fractional Efficiency
Figure 10 gives the fractional efficiency of the paint particles. The
measurement of the fractional efficiency could not be carried out under a
range of particle size which is below l-2ym because of the agglomeration of
the particles which occurred in the area of the scrubber, zigzag baffler,
and the ducts.
REFERENCE
1. Johnstone, H.F., R.B. Feild, and M.C. Tassler "Gas absorption and aerosol
collection in a venturi atomizer" Ind. Eng. Chem. 46, 160, August 1954
2 Morishima, N., "Presumption of Scrubber Performance" Performance of Dust
Collector, ed. by Iinoya, K., Ind. Tech. Center Press (1976)p.l46
3. Calvert, S. "Source Control by Liquid Scrubbing", Air Pollution III, 2nd,
ed. by Stern, A. C., Acad. Press (1968) p.475
4. Nukiyama, S., and Tanasawa, Y., "Study of Liquid Atomization" Trans.
Soc. Mech. Engrs. (Japan), 5, No.18, 68-136, February (1939)
19
-------
Table 1. Operating condition
Gas flow rate(m3/min)
4.5
Scrubbing liquid—711
Water. Heavy oil
Scrubber, pressure dropCmmHaO)
60,75,90
Zigzag baffler,pressure dropCmmHjO)
45
Paint spray pressure(kg/crTf)
1.5, 3.5
Scrubbing liquid ci rculation( I/mi n)
20
Table 2. Property of paint
Specific gravity
Solid i n pai nt
1.05
472.3 g/1
baffle plate
zigzag baffler
baffle plate
scrubbing 2.
crubbing 1.
id
spray gun"- :
blower
(i)
recorder: pressure loss
recorder: gas flow rate
3k
Digital Dust Counter
0)
sampling apparatus: part iclecocer*
t rati on
sampling apparatus: particle size
distribution
loo e ooool
• IO »«•««.
!• • •• _•# •!
pump
Figure 1. Schematic of the experimental apparatus
20
-------
l/>
§10000
c
E
8.5000
¦E
D
8
a>
o
i_
CP
c
H 1000
3
O
o
500
Cspray pressure:1.5kg/cm*
A spray pressure: 3.5 kg/cmj
aoi 0.05 qio a20
Weight concentration,CCg/Nm1)
Figure 2. Counting rate vs. weight
concentration of solid paint particles
100
90
80
3 70
^ 60
cc 50
*40
N
£ 30
d 20
10
0 1 23456789 10
Particle diameter, Dp( jum)
Figure 3. Particle size distribution
of solids in paint
spray pressure:!.5kg/crrf\
spray pressure:3.5kg/cm2 y
21
-------
c
o
5
A
3 r
2 1
Q)
C
K=0.6
-A-singl stage scrubbing
-A-double stage scrubbing
calculated using eq.(1)
scrubbing liquid:water
gas volume:4.4Nm3/min
spray gun:l.5kg/cma
particle cocentration:
(basis on the solid in paint)
Cj: 0.760-1,510g/Nma
Cc-'OX^-O^Og/Nm1
X K=0.5
50 60 70 80 90
Pressure drop, AP(mmHtO)
Figure 4. Collection characteristics of
single and double-stage scrubbing
100
5
4
c
o
Oi
c
-------
o
o
CL
cf
o
a>
c
*1
0.9
08,
Oscrubber equipped with a
zigzag baffler
• scrubber only
— spray gun :1.5kg/cm1
—spray gun:3.5kg/cm2
n zigzag bafflerAP=45mmHa0
scrub.liq.:heavy oil
50 60 70 80 90 100 110 120 130 140 60
Pressure drop, AP(mmHa0)
Figure 6. Collection characteristics of a double-
stage scrubber with a zigzag baffler
Z
o>
c 0.4
o
-------
100
c
o
*10
o
o
0.3
zigzag baffler AP=A5mmHaO
spray gun :1.5kg/cma
paint perticle,
Ci-0.0U~0.031g/Nm3
Co:0.0l1-0.026g/Nnr?
j.
0.4
e
o
iIT
(Li
'y
•4—
95
"o
c
o
8
i
92
Li.
.Oscrubber equipped with a
zigzag baffler
Iscrubber only
scrubber AP= 75mm H^O
zigzag battlerAP=45mmH20
0 1 2 3 A 5 6 7 8 9 10
Particle diameter,Dp(jjm)
Figure 10, Fractional efficiency of
paint particles
Oil mist generation,G(g/Nm'
Figure 8. Effect of the zigzag baffler
in removing paint particles on various
oil mist generations
scru bber AP= 75mmHaO
gas volume: 2,7-9.0 Nm3/min
120
Pressure droD AP(mmHaO!
Figure 9. Collection efficiency of oil mist
particles by a zigzag baffler
24
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APPLICATION OF SLIPSTREAMED AIR POLLUTION
CONTROL DEVICES ON WASTE-AS-FUEL PROCESSES
By:
Fred D. Hall
John M. Bruck
Diane N. Albrinck
PEDCo Environmental, Inc.
Cincinnati, Ohio
Robert A. Olexsey
U.S. Environmental Protection Agency
Cincinnati, Ohio
The recovery of energy from the combustion of municipal solid wastes is
becoming an attractive alternative as landfill space becomes scarce and the
availability of fossil fuels decreases. Particulate emissions from "waste-as-
fuel" processes, however, may differ significantly in chemical and physical
properties from particulate emissions produced by firing only coal. Such
differences can affect the design and operation of air pollution control
equipment. Presented in this paper are the results of a 2-month test program
at Ames, Iowa, with a mobile electrostatic precipitator (ESP) and a mobile
scrubber supplied by the U.S. Environmental Protection Agency (EPA), Industrial
Environmental Research Laboratory (IERL), Research Triangle Park. PEDCo
Environmental, Inc., and Acurex Corporation jointly conducted the test program
to examine the effect of burning refuse-derived fuel (RDF) on particulate and
heavy metal control efficiencies. The mobile ESP was used only as a primary
control device, whereas the mobile scrubber was tested both upstream and
downstream of the existing full-scale ESP. This paper also presents a status
report on a PEDCo test program with a pilot fabric filter at Ames.
25
-------
APPLICATION OF SLIPSTREAMED AIR POLLUTION
CONTROL DEVICES ON WASTE-AS-FUEL PROCESSES
INTRODUCTION
PEDCo is currently engaged under an EPA contract in testing and evaluat-
ing prototype air pollution control devices on various waste-as-fuel processes
for the control of potentially harmful air emissions. As part of the contract,
a sampling and analysis program was conducted at the RDF and coal cofiring
facility in Ames, Iowa, This program involved evaluating the pollutant removal
efficiency of a mobile ESP and a mobile scrubber from IERL by slipstreaming
flue gas from the suspension fired boiler. The objective of the program was
to determine the pollutant removal efficiencies of each device for variable
boiler loads, RDF inputs, and device operating parameters. Installation and
testing of the mobile devices took place in October and November 1978.
TEST PROGRAM
The PEDCo study depended on procuring small-scale air pollution control
devices and on slipstreaming and testing appropriate waste-as-fuel facilities
to generate necessary data.
Early in the project, several waste-as-fuel facilities were visited as
candidate sites for this program. The Ames, Iowa, site was chosen because of
the cooperation of the management, availability of data from previous studies
and evaluations, high reliability of the preprocessing and utility operations,
and limited facility modifications required.
The execution of a program of this type relies heavily on the cooperation
of plant management. The administrative, operating, and maintenance staff of
the Ames Municipal Electric System provided EPA, PEDCo, and Acurex with invalu-
able assistance throughout the program.
This test program was jointly conducted by Acurex, under contract to the
IERL Fine Particulate Control Branch, and PEDCo, under contract to the IERL
Fuels Technology Branch. Acurex was responsible for operating the mobile
devices and measuring particulate loading and sizing at the inlet and outlet
streams of each device. PEDCo was responsible for interacting with the Ames
Municipal Electric System and performing more complex sampling and analytical
tasks, Involving characterization of fuels and of the inlet and outlet streams
of each device.
The project objective was to characterize the specific pollutant removal
efficiencies of each control device for different boiler loads, RDF inputs,
26
-------
and operating parameters. It has been assumed that any full—scale application
of these control devices to a coal- and RDF-fired boiler must operate with
varying flue gas temperatures and compositions. A plan was developed to
monitor and measure these two factors, but not to modify or control them.
Ames Boiler 7 was used for the test program. The unit, which was in-
stalled in 1968, burns pulverized coal with a tangential firing mechanism. The
specifications of Boiler 7 are given by Hall et al. (1978) as follows:
Manufacturer Combustion Engineering
Heat input 460 GJ/h (436 x 106 Btu/h)
Steam output 160 Mg/h at 491^0 and 5860 kPa
(360,000 lb/h at 900^ and 850 psi)
Furnace pressure Balanced draft
Coal firing equipment 2 pulverizers;
8 nontilting tangential burners
Dust collection equipment American ESP
PEDCo characterized emissions by simultaneously measuring total particu-
lates, halogens, selected trace metals, and elemental composition of particu-
lates at the inlet and outlet of each device. The operating parameters that
were varied during testing of the mobile scrubber included pressure drop,
actual gas flow rate, and scrubber liquor flow rate. The gas flow rate was
increased as the scrubber liquor flow rate was decreased in such a manner that
the liquid-to-gas ratio remained relatively constant. The operating parameters
that were varied during testing of the mobile ESP included the number of
energized fields, specific collection area (SCA), and actual gas flow rate.
Figure 1 shows the slipstream and sampling locations for the test program.
Both mobile control devices were slipstreamed into the emission effluent stream
upstream of the existing ESP. The mobile scrubber was also operated as a
secondary device and tested downstream of the plant ESP. Samples were taken at
the inlet and outlet of each mobile control device while boiler conditions,
fuel composition, and control device operating conditions were monitored.
The PEDCo sampling and analysis program is outlined in Table 1. A Method
5 source sampling train (SST) was used simultaneously at the inlet and outlet
of each device. Total particulates were determined from the front half of the
SST. Halogens and trace elements were determined from the front and back
halves of the SST and from specified grab samples of coal and RDF. Elemental
analyses were conducted on the particulate catch of Acurex samples that cor-
related to PEDCo runs.
Discrete grab samples of coal, RDF, ESP ash, and scrubber liquor were
collected at the time of emission testing. Coal samples were taken before the
pulverizer, processed refuse samples were collected at the Atlas storage bin
27
-------
COLORADO
COAL
IOWA
COAL
PULVERIZER
to
00
BOILER
FIRING
RDF AND
PULVERIZED
COAL
FULL SCALE ESP
PULVERIZED
STACK
SECONDARY
INLET
STACK
PRIMARY INLET
MOBILE
SCRUBBER
SLIPSTREAM
DUCTWORK
STACK
MOBILE
ESP
SAMPLING LOCATIONS:
ASH - ESP ASH COLLECTION HOPPER
C - COAL MIXTURE (UNPULVERIZED)
- ESP INLET
- ESP OUTLET
- SCRUBBER INLET
UNFILTERED SCRUBBER LIQUOR
SCRUBBER OUTLET
PREPROCESSED RDF
EI
EO
SI
SL
SO
R
Figure 1. Schematic of the waste-as-fuel test site, at Ames, Iowa.
-------
Table 1. SAMPLING AND ANALYSIS PROGRAM FOR THE COFIRING FACILITY AT AMES, IOWA
Sample streaa
Collection technique
Chemical species and other
items determined
Analysis method
Sampling and
analysis responsibility
Air
Inlet/outlet
Crab
°2
Fryrite
PEDCo/Acurex
C°2
Fryrite
PEDCo/Acurex
SST
Total particulates
Pb, Sb, Cd, Fe, Cu
CI
Br
F
Inorganic elements
SO2
Particle size distribution
EPA Method 5
AAS
a
a
EPA Method 13B
SSMS
EPA Method 6
Cascade impactor
PEDCo/Acurex
PEDCo
PEDCo
PEDCo
PEDCo
PEDCo/Acurex/Brehm
Acurex
Acurex
Solid
Coal/RDF
Crab/composite
Density''
Pb, Sb, Cd, Fe, Cu
CI
Br
F
Inorganic elements
Fuel characterization
Gravimetric
AAS
a
a
EPA Method 13B
SSMS
Ultimate and
proximate
PEDCo
PEDCo
PEDCo
PEDCo
PEDCo
PEDCo/Brehm
PEDCo
ESP ash
Grabc
PEDCo/Acurex
Liquid
Scrubber liquor
Grab''
PEDCo
a Specific ion probe/titratlon with AgNOj or Hg(N03>2.
^ Only RDF was neaaured.
Stored for future reference.
d Sent to Montgomery Engineering.
-------
before the fuel entered the pneumatic feed to the boiler. The ESP particulate
catch was collected from the mobile ESP daily, and samples of unfiltered
scrubber liquor were collected at the time of emission testing. The tempera-
ture and pH of scrubber liquor were measured immediately after sampling and
recorded. Ultimate and proximate analyses were performed on ground and homog-
enized coal and RDF.
TEST RESULTS
Test results are based mainly on analyses of PEDCo samples. Acurex,
however, supplied the Method 5 samples analyzed by spark source mass spectrom-
etry (SSMS) for particulate elemental composition and provided size distri-
bution data. The Acurex samples were taken at the same time as PEDCo samples
to allow comparison of analytical data.
Table 2 presents average fuel analyses for coal and RDF at Ames. The
samples analyzed were composites of coal and RDF taken daily during the test
program. On a dry basis, the ash content of RDF (14.83 percent) is higher
than the ash content of coal (11.06 percent), whereas the sulfur content of
RDF (0.38 percent sulfur) is significantly lower than the sulfur content of
coal (3.13 percent). Heating values for RDF and coal as received are 15,450
kJ/kg (6,644 Btu/lb) and 23,680 kJ/kg (10,180 Btu/lb).
Figure 2 shows the particulate size distribution at the control device
inlet for each test condition. Each distribution value is an average for
several runs. Given the uncertainty of particulate sizing techniques, the
distribution when RDF and coal are fired does not appear appreciably different
from the distribution when only coal is fired. As expected, significantly
fewer particulates less than 1 vim in diameter were found at the inlet of the
scrubber when used as a secondary device; submicron particulates accounted for
less than 5 percent by weight of the total particulate matter at the scrubber
inlet in the former case, but more than 10 percent in the latter case.
Figures 3, 4, and 5 show the particulate removal efficiencies of the con-
trol devices as measured during the test program. The scrubber operated at
pressure drops of 25.4 and 76.2 cm (10 and 30 in.) H2O. The pressure drop was
increased by increasing the gas flow rate through the scrubber and adjusting
the scrubber liquor flow rate to maintain a constant liquid-to-gas ratio of
about 2.1 liters/standard (15 gal/1000 scf). The efficiency of the scrubber
was consistently above 99 percent when used as a primary device, but consider-
ably less (72 to 97 percent) when used as a secondary device. The efficiency
of the ESP ranged from 94 to 98 percent. Neither scrubber nor ESP particulate
removal efficiencies appreciably changed as the portion of heat input supplied
by RDF increased.
Table 3 shows the differences in elemental composition between coal and
RDF fuel samples, and Table 4 shows the differences in elemental composition
between uncontrolled emissions from burning coal plus RDF and uncontrolled
emissions from burning only coal. The fuels were analyzed from composite
samples from each test. The factors for uncontrolled emissions were determined
from measured data and engineering judgement. The elements not shown in Table
30
-------
Table 2. AVERAGE FUEL ANALYSES
(values in percent except as shown)
Coal
RDF3
Fuel as received
Proximate analysis
Water
16.13
13.25
Ash
9.27
12.78
Volatile matter
36.54
61.36
Fixed carbon*5
38.06
12.62
Heating value, kJ/kg
23,680
15,450
(Btu/lb)
(10,180)
(6,644)
Dry fuel
Ultimate analysis
Ash
11.06
14.83
Carbon
73.53
50.31
Hydrogen
1.38
3.97
Oxygen
9.72
30.10
Sulfur
3.13
.38
Nitrogen
1.17
.14
Heating value, kJ/kg
28,210
17,800
(Btu/lb)
(12,130)
(7,654)
Average of grab samples from Atlas storage bin.
k By difference.
31
-------
20.0
SCRUBBER AS PRIMARY DEVICE WHEN ONLY
COAL IS FIRED
SCRUBBER AS PRIMARY DEVICE WHEN RDF
AND COAL ARE FIRED
ESP WHEN ONLY COAL IS FIRED
ESP WHEN RDF AND COAL ARE FIRED
SCRUBBER AS SECONDARY DEVICE WHEN —
ONLY COAL IS FIRED
SCRUBBER AS SECONDARY DEVICE WHEN
RDF AND COAL ARE FIRED
20 30 40 50 60 70
98 99
AMOUNT OF PARTICULATE MATTER SMALLER THAN INDICATED SIZE,
% by weight
Figure 2. Particulate size distribution at control device inlets.
32
-------
20,000 1 1 1 1— 1 1
lft nnn - A INLET LOADING *
# • OUTLET LOADING
16,000 - ''
14,000 -
12,000 t -
"e 10,000 - ~ >3
5 8000 - | S §
>- ~ 5 S
>> 6000 z t 5 r
¦O £ UJ ^ t
cr 4000 £ ^
E £ O ^ Z
z 2000 ^ £ S ^ E
2 1000 * G ¦ % Z.
t— „ . <-< ul /
5 cn £ w ^
| f ® t ^ f
LlJ tO
O '
S CTl
O 0>
0
£ 120 -
= 100-
—I •
£ eo - * •
5 - .. I •
40 -
20 -
01 I 1 I I I I
0 5 10 15 20 25 30
PORTION OF HEAT INPUT SUPPLIED BY RDF, %
T
A INLET LOADING
• OUTLET LOADING
>-
o
>-
o
>-
M
*
o
oj
ct>
CTl
>-
O
ID
a>
CJ1
to
cn
CTl
ir>
cn
r-~
cn
cn
>-
o
/*
ir>
cri
-------
5000
4000
3000
2000
1000
INLET LOADING
OUTLET LOADING,
CORRECTED FOR LEAKAGE
>-
o
QO
cn
to
«c
oc
=3
o
u.
o
o
UJ
to
<£
00
5 10 15 20
PORTION OF HEAT INPUT SUPPLIED BY RDF, %
25
2 3
Figure 4. Efficiency of mobile ESP with SCA of 59 m /m per second
34
-------
400 -
300 -
200 -
100 -
2 40 -
30
20 -
10 -
>-
o
LT>
vo
>-
o
00
o
CT>
>-
o
fx*
(/)
o>
—J
~ INLET LOADING
• OUTLET LOADING
o
r^.
•
r^.
CTl
>-
O
V8
o
ir>
N
LjJ
in
CVJ
CTl
I
1
15 20 25
PORTION OF HEAT INPUT SUPPLIED BY RDF, %
Figure 5. Efficiency of mobile scrubber as a secondary device.
35
-------
Table 3. DIFFERENCES IN ELEMENTAL COMPOSITION BETWEEN
COAL AND RDF FUEL SAMPLES FROM ALL TESTS8
Factor by which elemental
concentration in RDF differs from
elemental concentration in coal
Elements more concentrated in
RDF than in coal
Copper^
120
Leadb
40
Bromine
25
Fluorine
25
Zinc
10
Manganese
10
Antimony
5
Chromium
5
Cobalt
5
Rubidium
4
Tin
3
Barium
3
Silver
3
Cadmium^
2
Titanium
2
Strontium
2
Phosphorus
2
Elements less concentrated in
RDF than in coal
Vanadium
50
Beryllium
25
Scandium
7
Yttrium
6
Sulfurc
5
Molybdenum
5
Iron^
4
Thallium
3
Nickel
2
Cerium
2
Lanthanum
2
Zirconium
2
£
Based on SSMS analysis unless otherwise stated.
^ Based on AAS.
c The initial determination of sulfur concentration might have been low by a
factor of 3-10 because of high organic content, which necessitated low-
temperature ashing of the samples.
36
-------
Table 4. DIFFERENCES IN ELEMENTAL COMPOSITION BETWEEN
UNCONTROLLED EMISSIONS FROM BURNING COAL PLUS RDF
AND UNCONTROLLED EMISSIONS FROM BURNING ONLY COALa
Factor by which elemental
concentration in emissions from
burning coal plus RDF differs from
elemental concentration in emissions
from burning only coal^
Elements more concentrated in
emissions from burning coal
plus RDF than in emissions from
burning only coal
Gaseous chlorides
10
Lead
6
Zinc
5
Copper
3
Iron
2
Elements less concentrated in
emissions from burning coal
plus RDF than in emissions
from burning only coal
Zirconium
2
a Only elements that show definite differences are listed. Other elements, as
determined by SSMS and/or AAS, show concentrations too variable to determine
emission differences.
k Estimated from actual data and engineering judgement.
37
-------
A were too variable to determine emission differences between burning coal plus
RDF and burning only coal.
Elemental concentrations in fuel do not appear directly related to ele-
mental concentrations in uncontrolled emissions. For example, copper was
measured to be 120 times more concentrated in the RDF fuel sample than in the
coal fuel sample, and lead was estimated to be 40 times more concentrated.
Uncontrolled emissions from burning a mixture of 75 percent coal and 25 percent
RDF contain three times more copper and six times more lead than when burning
only coal. The data also indicate that gaseous chloride emissions are about 10
times greater when coal and RDF are burned than when only coal is burned.
The efficiency of trace element removal was typically less with the
scrubber than with the ESP. For example, lead removal efficiency was about 90
percent with the scrubber, but greater than 95 percent with the ESP.
ADDITIONAL TESTING
The use of a fabric filter as a primary control device for the removal of
specific pollutants from waste-as-fuel processes is being studied. A pilot
fabric filter designed and constructed for this project is currently installed
at the cofiring utility in Ames. A test program of about the same scope as the
program with the mobile ESP and scrubber is scheduled for completion by the end
of July 1979.
Following the Ames testing program, the fabric filter is scheduled for
testing at the Braintree Municipal Incinerator in Braintree, Massachusetts.
The Braintree program will also involve a pilot scrubber. The scope of this
test plan will be much the same #s the scope of the Ames program.
REFERENCES
1. Hall, J.L. et al. Compliance Tests of Steam Generator Unit 7 for Parti-
culate Effluent. Ames, Iowa, July 17-25, 1978.
38
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EVALUATION OF THE CEILCOTE IONIZING WET SCRUBBER
By:
David S. Ensor
Meteorology Research, Inc.
Altadena, California 91001
and
Dale L. Harmon
U.S. Environmental Protection Agency
Office of Research and Development
Industrial Environmental Research Laboratory
Research Triangle Park, North Carolina 22711
ABSTRACT
The Ceilcote ionizing wet scrubber installed on a refractory brick
kiln was evaluated with tests involving particulate mass emission, par-
ticle size distribution, and opacity. The overall efficiency was 93 per-
cent with an average outlet opacity of 8 percent on a 1.68 m (5.5 ft) path
length. The average particle cut diameter of the scrubber is 0.5 microns
with a theoretical power input of 67 watts/anr (2.5 hp/1000 acfm). The
theoretical power requirement for the ionizing wet scrubber was 41 watts/
am3 (1.54 hp/1000 acfm). A cooling tower supplying chilled water to
the prescrubber required an additional 26 watts/anr (0.96 hp/1000 acfm)
for a total system input of 67 watts/anr (2.5 hp/1000 acfm). It is
recommended that the scrubber be considered where practical for the
removal of fine particulate matter.
39
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SECTION 1
SUMMARY AND CONCLUSIONS
The Ceilcote ionizing wet scrubber (IWS) was evaluated with field
measurements of particle collection efficiency and an analysis of power
consumption.
This evaluation was one of a series of such evaluations being con-
ducted by the Industrial Environmental Research Laboratory of the U.S.
Environmental Protection Agency (EPA) to identify and test novel de-
vices which are capable of high efficiency collection of fine par-
ticles. The test methods used were not the usual compliance-type
methods but were, rather, state-of-the-art techniques for measuring
efficiency as a function of particle size using cascade impactors and
electrical aerosol size analyzers (EASA).
The following conclusions were made from the study:
1. The emission from the refractory kiln was a submicron
fume formed by condensation of volatile material baked
from the raw clay. The aerosol had a mean diameter of
0.6 microns and a geometric standard deviation of 5.0
as determined with a cascade impactor assuming a par-
ticle specific gravity of 1.8 g/cnr. The cooling of
the flue gas from 150° to 50°C (300° to 120°F) doubled
the concentration of particulate matter entering the
ionizers.
2. The average overall mass collection efficiency was
93 percent for 3 days of testing. The average inlet
concentration was 0.25 g/dsm^ (0.0076 gr/dsft3). The
average outlet opacity was 8 percent over a 1.68 m
(5.5 ft) path length as measured with a plant process
visiometer at about 99°C (210°F).
3. The particle cut diameter (the particle diameter col-
lected with a 50 percent efficiency) was 0.4 to 0.6
microns. The total theoretical power into the scrubber
(the power required to operate the scrubber would in-
clude the efficiency of power supplies, pumps, and fans)
was 67 watts/am^/mln (2.5 hp/1000 acfm). The ionizers
required the greatest percentage of the power Input at
48.4 percent and the cooling tower required 38.4 per-
cent. The use of energy to remove particulate matter
1n the scrubber is better than a theoretical high-
pressure drop venturl scrubber.
It 1s recommended that the Ceilcote ionizing wet scrubber be con-
sidered in applications where scrubbers are practical for the removal
of fine or submicron particulate matter.
40
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SECTION 2
SITE DESCRIPTION
PROCESS
Globe Refractories, Inc., located in Newell, West Virginia, makes
bloating-type refractories. These refractory products are used in the
steel industry to line ladles. The term "bloating" means the refractory
brick expands permanently when reheated, sealing the lining of the ladle.
The raw material is a local clay called Lower Kittanning clay com-
posed mainly of kaolinite, quartz, illite, and pyrite. Minor constit-
uents are organic matter, micas, and ammonium chlorides or fluorides.
The clay is formed into the required shapes, including bricks, sleeves,
nozzles, and pocket blocks and fired to about 1000°C (2000°F) in a
tunnel kiln under a controlled temperature profile over 4 to 6 days
cycle. The formed clay is loaded on tunnel kiln cars which are slowly
pushed through the kiln.
The chemical reaction during firing includes oxidation of organic
matter to carbon dioxide and water, oxidation of pyrite to iron oxide
and sulfur oxides, decomposition of kaolinite and illite to release
chemically combined water, decomposition of amnonium chloride or
fluoride to form ammonia, gaseous chlorides, and fluorides. The flux-
ing of alkalis forms a glass bonding the brick.
The major emissions problem from the kilns is believed to be a
NH4 HSO4 smoke from reaction of ammonia gas and sulfur oxides.
CONTROL DEVICE
General
The Ceil cote Ionizing wet scrubber consists of five sections:
• Quench unit
• Prescrubber (with cooling tower)
• First Ionizing wet scrubber
• Second ionizing wet scrubber
• Induced draft fans and stack
About 28.3 am^/sec (60,000 acfm) of flue gas at 150°C (300°F) is
piped to the scrubber through a 1.5 m (5 ft) diameter 30 m (100 ft)
long duct constructed of fiberglass reinforced plastic. A diagram of
the scrubber is shown in Figure 2-1.
Quench Section
In the quench section, the flue gas is reduced from 150°C (300°F)
to 60°C (140°F) by evaporative cooling. The quench water is supplied
at 7.6 £ /sec (120 gpm) at a 3.6 x 105 N/m2 (52 psig) pressure. The
unit is 2.1 m (7 ft) in diameter and 2.7 m (9 ft) long.
41
-------
STACK
IWS UNIT #1
CROSS FLOW SCRUBBER
COOLING TOWER
IWS UNIT 12
TUNNEL FILM STACK
IONIZER
FAN
EMERGENCY
VENT
DAMPER
OUTLET TEST
LOCATION
PRE SCRUBBER
(CONDENSER/COOLER)
INLET TEST
LOCATION
sQUENCH UNIT
Figure 2-1. Overview of Ceilcote ionizing wet scrubber.
-------
Prescrubber and Cooler
The gas temperature is reduced to 46°C (115°F) in the prescrubber.
The prescrubber has a gas flow area 3 m x 3 m (10 ft x 10 ft) with an
overall length of 5.5 m (18 ft). It is of cross-flow design with a
series of inlet baffles and a i.8 m (6 ft) deep bed of TelTerette
packing. There are continuous water sprays on the inlet baffles in
front of the packing and above the packed bed. Outlet baffles are
sprayed on a periodic basis. Most of the water for the sprays is
cooled in the cooling tower. The cooling tower reduces the temper-
ature of the prescrubber water from 49°C (120°F) to 29°C (85°F) with
a flow of 54 j^/sec (860 gpm) and a heat transfer rate of 3.78 x 10^
cal/hr (15 x 10® Btu/hr).
The cooling tower is 3.8 m (12.5 ft) in diameter and about 6.6 m
(21.5 ft) high. The tower has a stack extension for a total height of
11 m (36 ft). It has 5 cm (2 in.) Tellerette Type-R packing and the
entrainment separator is 0.3_m of 5 cm (2 in.) Tellerette Type-R pack-
ing. A fan rated at 27.3 anr/sec (58,000 acfm) at 10 cm HoO supplies
cooling air at the bottom of the tower. The cooled water is pumped
to the prescrubber at 52.4 //sec (832 gpm) at a 1.8 x 10^ N/m? (26
psig) pressure.
Ionizing Wet Scrubber (IWS)
The ionizing wet scrubber consists of two sections: an ionizer
or charger and a cross-flow scrubber. The ionizer consists of charging
wires suspended between irrigated grounding plates. The firstAionizer
operates at 30 kV with a current of 100 to 110 mA and the second oper-
ates at 31.5 kV with a current of 210 to 225 mA The scrubber con-
tains 1.2 m (4 ft) of Irrigated 5 cm (2 in.) Tellerette Type-R packing
and 0.3 m (1 ft) of unirrigated packing for entrainment separation.
The irrigated packing is sprayed from the front and the top. The en-
trainment separator is flushed periodically. The recirculation Dump
is rated at 36.3 ^/sec (575 gpm) at a pressure of 1.8 x 10$ N./mz
(26 psig).
Fan and Stack
The fiberglass reinforced polyester fan is rated at 20 anP/sec
(42,000 acfm) at 46°C (115°F) water saturated and 28 cm WC (11.in. WC)
static pressure. The fiberglass reinforced plastic stack 1s 1.67 m
(66 in.) in diameter and 46 m (150 ft) in height.
Test Locations
The test locations are Indicated on Figure 2-1. Both the Inlet
and the outlet were located on circular ducts with sufficient length
before and after the test port for smooth flow.
43
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SECTION 3
TEST METHODS
TECHNICAL APPROACH
Measurements were conducted in three areas: aerosol character-
ization, gas composition, and process streams.
SIZE DISTRIBUTION MEASUREMENT
Cascade Impactor
The MRI cascade impactor is an annular jet collector type sim-
ilar to that reported by Cohen and Montan. The body of the device
consists of quick connect rings supporting jet plates, collection
discs, and a built-in filter holder. The design permits flexibility
in various sampling situations.
The tests were conducted with procedures described by Harris.
The particles were collected on Apiezon grease-coated, 304 stainless
steel foil collection discs. The discs were weighted to 0.01 mg on
a Cahn electrobalance.
The inlet impactors were operated at duct temperature. The
outlet impactors were heated to 121°C (250°F) with electrical heating
jackets. The nozzles on the outlet impactors were also extended by
46 cm (18 in-) tubes which were also heated to 121 °C (250°F) to dry
the aerosol before entering the impactor.
The impactor data were reduced with a procedure described by
Markowski and Ensor which is similar to the method described by
McCain et al.
Fine Particles
The measurement of the size distribution of submicron particles
was a two-stage process. First, the aerosol sample was diluted with
clean, dry air; and secondly, the particulate matter in the diluted
aerosol was measured with a Thermosystems Model 3030 electrical aero-
sol size analyzer.
Two separate dilution systems were used. The inlet dilution
system consisted of a sampling probe with a cascade impactor precutter
to remove particles greater than 2.5 microns followed by an out-of-
stack, three-stage quantitative dilution system. The sample can be
diluted with filtered dry air from 3:1 to 1000:1 by adjustment of
control valves. The outlet dilution system was a single stage 1n-
stack mixing tee. In both systems, the clean air flows are measured
with orifices and the sample flows with Venturis.
44
-------
Thhe TSI Model 3030 electrical aerosol size analyzer (EASA) was
used at both the inlet and outlet. The EASA consists of a charger,
where a known charge is placed on the particles, and a mobility analyzer,
where the charged particles are attached to a central collecting rod.
The size of particles collected in the mobility analyzer depends on
the applied voltage on the collecting rod. The aerosol passing through
the mobility analyzer is detected by measuring the current transferred
by the particles. The aerosol distribution in 11 logarithmic steps
from 0.003 to 1 micron is measured.
OPACITY MEASUREMENT
The opacity at the inlet and outlet of the scrubber was measured
with an MRI plant process visiometer (PPV). The instrument was in-
stalled on a 3-inch port and the sample was heated to reduce relative
humidity. A diagram of the instrument is shown in Figure 3-1. The
aerosol particles in the chamber were illuminated by a flash lamp with
an opal glass filter. The scattered light was detected by a photo-
multiplier 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:
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 Bouger Law.
IT
0
where
bscat = the scattering coefficient due to
scattered light
3(5)= volume scattering function
e - scattering angle
Opacity (percent) ¦ [1 - exp (-bext L) 100
where
bext " ext"incti°n coefficient, m"*
L ¦ stack diameter, m
45
-------
A
CALIBRATOR
LIGHT RAYS
ELECTRONICS
FLASH
LAMP -—
DIFFUSER
-PHOTO MULTIPLIER
COLLIMATOR
SAMPLE
VOLUME
ASPIRATOR
76-394/?
Figure 3-1. Diagram of the plant process visiometer.
46
-------
The instrument is spanned with an internal calibrator consisting
of 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
calibration opal glass is calibrated with oil smoke with reference in-
struments using both an integrating nephelometer and a transmissometer.
The PPV is described in detail by Ensor et al.
The PPV at the inlet was mounted with the 1.9 cm (3/4 inch) probe
positioned in the center of the duct. The probe was insulated and the
chamber electrically heated. The PPV at the outlet was placed on the
ground and a 3 m (9 ft) probe was extended into the duct from the bottom.
The probe and chambers were electrically heated to about 93°C (200°F)
to ensure that the gas was above the water dew point.
GAS COMPOSITION
The concentration of O2, CO, and CO2 was measured with an Orsat
analysis following EPA Method 3. The water content of the flue gas
was obtained with the impinger catch during the cascade impactor tests.
PROCESS VARIABLES
The process variables were obtained as follows:
• The velocity was determined with an S type pitot probe
following EPA Methods 1 and 2.
• Pressure drop across the scrubber and pressure at the
draft fan were measured using pressure transducers and
recorded on a strip chart.
• Samples of water were obtained for determinations of
dissolved solids.
• The water flow rates, pressure, and power required were
estimated from the design specifications.
47
-------
SECTION 4
FIELD TEST RESULTS
MASS COLLECTION EFFICIENCY
The overall average efficiency of the scrubber as a unit shown
in Table 4-1 was 93.5 percent. The "ionizer off" test conducted
on November 10, 1978. maybe used as an indication of the concentration
of particulate matter entering the IWS units. If it is assumed no
particulate matter was removed by the IWS packing, the IWS efficiency
was 98.2 percent.
TOTAL POWER INPUT
The total theoretical power input to the scrubber is summarized
in Table 4-2. The electrical power to the cooling tower for air cir-
culation and water circulation referenced to the volumetric flow of
TABLE 4-2. ESTIMATION OF THEORETICAL POWER REQUIREMENTS
Scrubber Stage
Gas
watts/am^/mln
(hp/1000 acfm)
Mater
watts/amJ/mln
(hp/1000 acfm)
Corona
watts/amfymln
(hp/1000 acfm)
X
of
Total
Quench
1.97
(0.074)
3.0
Prescrubber
6.82
(0.256)
10.3
IMS l(«)
8.47
(0.318)
4.80
(0.180)
1.82
(0.068)
22.7
IMS ?(»)
8.47
(0.318)
4.80
(0.180)
3.89
(0.146)
25.7
Cooling Tower^)
16.7
(0.628)
8.79
(0.330)
38.4
Total
33.6
(1.264)
27.2
(1.02)
5.71
(0.214)
66.6
(2.50)(O
Percent of Total
50.6
40.8
8.5
T7J The gas pressure drop was divided between the two units
(b) Includes an estimate of the fan and the punp power requirements
(c) Note the total power on 11/21/78 measured with a Volt-amp meter was 149 hp or
3*02 hp/1000 acfm (80.5 watts/anP/mln)
48
-------
TABLE 4-1. SUMMARY OF OVERALL MASS COLLECTION EFFICIENCY
Date
(1973)
Inlet
Cone.
mg/m3(a)
Run (gr/ft^)
Run
Outlet
Cone.
mg/m3(a)
(gr/ft3) (a)
Penetration
(%)
11/10
(0
25
26
4a
259
256
271
16
17
46
756
471
360
262
(0.114)
(0.231)
201.9
11/18
27
28
349
318
13
19
16.9
34.8
333
(0.145)
?r\i 1
(0.0123)
7.77
11/20
(c)
29
30
5a
186
185
226
36
37
5c
6.84
15.10
7.41
199
(0.0331)
9.78
(0.00426)
4.91
efficiency Opacity^5)
(%) (%)
-101.9
81
92.2
11
95.1
3.7
Comments
IWS Power shut off
Phase II tests
after completion
of repairs
11/21
(c)
32
33
66
198
218
311
242
(0.106)
38 19.7
39 21.1
6a 9.38
16.7
(0.00728)
6.90
93.1
6.8
Note: Both impactors and in-stack filters were heated to 121° C (250° F) at the outlet.
(a) 21.1° C, 76 cm, dry
(b) 1.676 meter (5-1/2 ft) path length
(c) In-stack filter samples - the other mass concentration results were computed from the
cascade impactor tests.
-------
flue gas. processed was considered as the energy input to the scrubber.
Also, the gas side pressure drop of 10 cm WC was assumed to result only
from the IWS units. The greatest single power input is from the gas
side pressure drops (it should be noted that this input includes the
cooling tower air flow). The major power input from these estimations
is in the cooling tower; thus, the need to condense components of the
flue gas stream in order to control the emission from the kiln results
in an additional energy requirement when compared to a process which
emits only solid particles. Also, the tests were conducted under atmo-
spheric temperatures of about 4°C (40°F) which leads to efficient cooling
of the water streams.
PARTICULATE MATTER FORMATION
Using the water analysis and the gas stream properties, the in-
crease in particulate matter concentration by evaporation is shown in
Table 4-3. It was assumed that the gas stream was initially cooled
by evaporation under conditions of constant enthalpy to saturation,
then cooled by sensible heat transfer to the observed quench temper-
ature. It was assumed that cooling of the gas from the saturation
temperature did not cause the removal of particulate matter.
The fraction of particulate matter formed by evaporation of the
scrubber water is estimated from the test day of November 10 when the
ionizers were shut off. Particulate matter formation from water evap-
oration accounts for about 30 percent, and about 70 percent results
from condensation directly from gas phase for the particulate matter
formed in the scrubber. The influence of the particulate matter
formed by the evaporation of water is shown in Figure 4-1 with an ana-
lysis of the tests. The outlet concentration and scrubber penetration
were directly related to the quantity of water evaporated in the quench
section and estimated particulate matter. Thus, additional improvement
in efficiency may be realized by reducing the dissolved solids in the
quench water. However, the emissions were below any applicable regu-
lation and the costs of providing fresh water for cooling the flue gas
may not be warranted at the present time.
SIZE DISTRIBUTION
The impactor and EASA results were combined to obtain particle
size distributions and fraction penetration curves. The last day of
testing, November 21, 1978, was selected for analysis in detail. All
measurements exhibited very good repeatability.
The differential size distribution is shown in Figure 4-2. The
figure illustrates many of the problems experienced with obtaining
fine particle data. It was noticed that the size distribution measured
at the inlet was related to the life of the silica gel used to dry the
dilution air. The size distribution obtained under conditions of fresh
silica gel was bimodal. However, as the silica gel was depleted, the
distribution grew to a single mode at 0.4 microns and appeared to match
the impactor results rather than the bimodal case. It is believed that
50
-------
TABLE 4-3. ESTIMATION OF AEROSOL FORMATION IN QUENCH SECTION
Oate
(1978)
11/10
11/18
Il/?0
11/21
Voluaetrlc
6as Flow
Inlet tend,
¦'/see
(acfa)
29.32
(62132)
22.72
(48148)
25.48
(53980)
21.(3
(45828)
Tenp.
Inlet
•C
m
in
(MS)
154
(313)
13S
(275)
138
(281)
Hater Cone,
by Yoluae
Inlet
Percent
7.69
3.63
3.53
S.33
Specific
Hwldlty
Inlet
9/9
(lb/lb)
0.1342
(0.1342)
0.067
(0.067)
0.0589S
(0.0589S)
0.0907
(0.0907)
So)Ids
In Water
-t/4
2600
2830
T After
Quench
•C
if)
48
(110)
43
(110)
41
(106)
41
(106)
Specific
Huaidity
After Quen:h
9/9
(lb/lb)
Specific
Hunidity
at Saturation
g/9(«)
(lb/lb)
T at
Saturation
•c(«)
m
Increase
In Specific
HuMdlty
9/9
(lb/lb)
Increase in
Particulate
Particulate Natter
fron evaporation
Coopared to
Inlet, Percent(d)
C.0755
0.19D
63
0.0558
163.5
(C.0755)
(0.190)
(145)
(0.0558)
(0.0713)
63
0.06
0.119
55
0.052
147.0
(0.06)
(0.119)
(131)
(0.052)
(0.0641)
44
0.053
0.095
52
0.0361
115
(0.US3)
(0.095)
025)
(0.361)
(0.0501)
60
0.053
0.130
56
0.0393
131
(0.053)
(0.130)
(132)
(0.0393)
(0.0571)
54
Particulate Natter fror
tvaporation Co«Mred tc
total Generated
(a) It Is assiacd that the Inlet gas Is cooled fey evaporation as a constant enthalpy process and then the
gas is cooled by sensible heat transfer
(b) Obtained by the product of the Mater solids concentration and increase in specific hmldtty, it is
itMNd after tarnation, the particulate sitter Is not rtaovedd fey condensation In the quench section
(c) Standard conditions 21.1* C, 76 cm Mg. Dry
(d) Ratio of the increase In particulate Batter formed fron evaporation to that Measured during the Inlet
tests
(e) Computed by dividing the increase In particulate Matter froo evaporation by the outlet of 11/10 froa
the average outlet concentrations froa 11/18, 20, and 21. Assuaes negligible reaoval of particulate
aatter in the prescrwbber «nd deenerglzed IMS.
5-13
-------
11
10
9
8
7
26
24
22
20
18
16
6
5
4
3
2
1'
5 14
»
I 12
10
8
6
4
2
O OUTLET CONCENTRATION,
• PENETRATION,
rc » CORRELATION COEFFICIENT
y « A + Mx
J UL
A
96.6
0.975 0.0894 -5.18
_L
-L
±
20 40 60 80
100 120 140 160 180 200
ng/ds nr
-t-
-4~
¦+-
i 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09
gr/dscf ,9.,61
CONCENTRATION OF AEROSOL FORMED IN QUENCH SECTION OF SCRUBBER
Figure 4-1. Scrubber performance as a function of
scrubber-generated aerosol.
52
-------
O Inlet Cascade Impactor
• Outlet Cascade Impactor
O Inlet EASA with Hat Dilution Mr
~ Inltt EASA with Dry Dilution Air
Corrected for Cross Sensitivity
& Inlet EASA with Dry Dilution Air
Uncorrected for Cross Sensitivity
+ Outlet EASA - Corrected for
Cross Sensitivity
A Outlet EASA - Uncorrected for
Cross Sensitivity
rz
J
1.01
0.1 1.0
PARTICLE DIAMETER, microns
79-263
Figure 4-2.
Impactor and EASA differential
size distributions for 11/21/78.
53
-------
the emission was very reactive with water. The dilution system was
operated with dilutions of up to 1000:1; therefore, with fresh silica
gel the aerosol would be measured with insignificant relative humidities
(absolute humidity of 10"^g Hftfq air). The cascade impactor measured
the particle size distribution under flue gas conditions of about 3 per-
cent relative humidity. The single mode distribution is the one most
likely to be present at the inlet of the scrubber. The size distri-
butions measured by the EASA were repeatable and the humidity phenomenon
was observed on other test days.
A second uncertainty is related to the data reduction procedures.
The EASA has cross sensitivity errors in the last four size channels (a
particle will be sensed in more than one channel). A computational pro-
cedure reported by Twomey was used to reduce the data to correct for
cross sensitivity; the instruction manual method which does not correct
for cross sensitivity was also used. The size distribution reported for
the EASA depends on the data reduction technique.
In these tests, the instruction manual method agrees better with
cascade impactor size distribution than the results corrected for cross
sensitivity. However, the cascade impactors also have cross sensivity
and the reduction technique used does not correct for nonideal behavior
in the impactor. Therefore, agreement of EASA and impactor size distri-
butions does not mean any particular size distribution is correct. For
this reason, results from different approaches are reported. The am-
biguity in reported size distribution indicates the difficulty of ob-
taining this information for a reactive condensible aerosol even when
state-of-the-art experimental techniques are used.
PARTICLE SIZE DEPENDENT PENETRATION
In Figure 4-3, the combined impactor and EASA penetration curves
are shown for November 21, 1978. The interpretation of the EASA re-
sults is an important consideration in computation of the penetration.
The data obtained with partially spent silica gel at the inlet (diamond)
are the most believable in the 0.1 to 0.3 micron range.
SCRUBBER PERFORMANCE
The aerodynamic diameter as defined by Calvert et al. is given by:
dactual
aero
where
C
Cunningham correction factor
Particle density, g/cm3
P
dactual = Actual cut diameter, microns
54
-------
100 I—
10
o
fc
2
fc
SB
£
1.0
*
r
0.1
J I I I I I 11
I
I
I
I
0 Clictdt lapictor
O EASA • Met Dilution
Air Corrected for
Cross Sensitivity
~ EASA - Dry Dilution
Air Corrected for
Cross Sensitivity
£ EASA - Uncorrected for
Cross Sensitivity
On* Stenderd Oevlitlon
Knits Indicated
I
I
I
J ¦ I I i I I I
' ' ' ' ' ' ¦ '
0.01
0.1 1.0
PARTICLE DIAMETER, microns
10
79-265
Figure 4-3. Particle size dependent penetration for
cascade Impactor and EASA results.
55
-------
The aerodynamic cut diameters were computed from the actual dia-
meter. The EASA size distribution was used because the cut diameter
was below the resolution of the impactor.
The square root of the Cunningham correction factor and density
yJC p was computed for the size range of interest in Figure 4-4. The
actual size cut diameters were taken from Figure 4-5 and reported 1n
Table 4-3. Depending on the measurement and data inversion technique,
the aerodynamic cut diameter was from 0.4 to 0.6 microns.
TABLE 4-4. SUMMARY OF CUT DIAMETERS
Technique
EASA uncorrected for channel
cross sensitivity
EASA Inlet size distribution
dry dilution air
EASA Inlet size distribution
water In dilution air
Actual Diameter
at SOX Penetration
(micron)
0.28
0.24
0.14
(g/cm*)1/2
2.16
2.28
2.75
Aerodynamic
Cut Diameter
(micron)
0.60
0.55
0.38
The aerodynamic cut diameter from Table 4-3 and theoretical power
from Table 4-2 are shown 1n Figure 4-4. This figure has theoretical
performance curves for a number of different scrubber types for com-
parison. These results suggest that the IWS is more efficient than a
theoretical venturi scrubber.
The aerodynamic cut diameter obtained for the whole scrubber on
November 21 Is believed to be valid for the IWS units Including the
aerosol generated in the quench and prescrubber section. The generated
particles are captured 1n the IWS in the 1 to 10 micron diameter range.
The penetration In the fine particle range, less than 1 micron, appears
to be unaffected by particle generation. Thus, the aerodynamic cut dia-
meter 1s unaffected by this phenomenon.
OPACITY
The wide ranges of concentration and opacity between Inlet and
outlet conditions allowed the correlation of mass concentration and
opacity as shown 1n Figure 4-5. The correlation coefficient of 0.96
suggests that the size distribution was fairly consistent.
56
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PRESSURE DROP. Inches H20
1 1.5 2 3 4 S 6 7 8 9 10 15 20 30 40 50 60 80 100
4.01
la, lb SIEVE PLATE SCRUBBERS
3.0-
2a. 2b VENTURI SCRUBBERS
IMPINGEMENT PLATE
2.0
PACKED COLUMNS
I
•
ae
ui
i
mm
O
l.o-
0.8-
5
0.5-
- A All
OIWS
Into Systen
Only
0.4
0.3
300
100
0.2
POWER, watts (an /m1n)
0.25
0.5
0.8 1.0
POWER, hp/1000 acfm
5.0
3.0
2.0
0.1
2
4 5 6 7 8 9 10
70 90100
200 300
30 40 50
3
20
PRESSURE DROP, cm H,0
2 77-413/2
Figure 4-4. Aerodynamic cut diameters of the Cellcote ionizing
wet scrubber compared to the theoretical performance
of other scrubber types (after Cooper and Anderson,
adapted from Calvert).
-------
10
20
-• 30
0.25
-- 40
-¦ 50
fc
1
60 -go
« *
0.50
fc . ¦ 0.0142 + 0.210 (MASS CONCENTRATION, gr/ft )
ICIt
CORRELATION COEFFICIENT 0.956
t=8
» If
70 o —
-¦ 80
0.75
1.0
0.1
0.2
gr/ft3
0.3
¦ H ¦ ¦ -4 ¦¦¦
400 500
100
ZOO
100
600
700
MASS CONCENTRATION, mg/nf
79-268
Figure 4-5. Correlation of opacity with mass
concentration.
58
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SECTION 5
REFERENCES
Calvert, S. "Engineering Design of Fine Particle Scrubbers." J.
Air Poll. Cont. Assoc. 24:929-934. 1974
Calvert, S.J., J. Goldschmid, D. Leith, and D. Mehta. Scrubber Hand-
book. EPA-R2-72-118a, NTIS No. PB-213-016, August 1972.
Cohen, J.J., and D.M. Montan. "Theoretical Considerations, Designs,
and Evaluation of a Cascade Impactor." Am. Ind. Hyg. Assn. J.
28:95-104. 1967
Cooper, D.W., and D.P. Anderson. Dynactor Scrubber Evaluation.
EPA-650/2-74-083a, NTIS No. PB 243-365 U.S. Environmental Pro-
tection Agency, RTP, NC, June 1975.
Ensor, D.S., L.D. Bevan, and G. Markowski. "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
Harris, D.B. Procedures for Cascade Impactor Calibration and Operation
in Process Streams. EPA-600/2-77-004, NTIS No. PB 263-623, USEPA,
RTP, NC, January 1977.
Markowski, G.R., and D.S. Ensor. "A Procedure for Computing Particle
Size Dependent Efficiency for Control Devices from Cascade Im-
pactor Data." 70th Annual Meeting of the Air Pollution Control
Assoc., Toronto, Canada, June 1977.
McCain, J.D., G. Clinord, L.G. Felix, and J.Johnson. A Data Reduction
System for Cascade Impactors. Proceedings: Advances in Pgj^2f']e
Sampling and Measurement. fPA-600/7-79-065, NTIS No. PB 293-363,
USEPA, RTP, NC, February 1979.
Perry, J.H. Chemical Engineer's Handbook. McGraw-Hill Book Company,
Inc., New York. 1963
Strauss, W. Industrial Gas Cleaning. Pergamon Press, New York,
pp. 333-334. 1974
Twomey, S. "Comparison of Constrained Linear Inversion and an Iter-
ative Nonlinear Algorithm Applied to the Indirect Estimation of
Particle Size Distributions." J. of Computational Physics. 18:
188-200. 1975
59
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DEMONSTRATION OF A HIGH FIELD ELECTROSTATICALLY
ENHANCED VENTURI SCRUBBER ON A MAGNESIUM FURNACE
FUME EMISSION
by
M. T. Kearns
Air Pollution Systems, Inc.
Kent, Washington 98 031
and
D. L. Harmon
Industrial Environmental Research Laboratory
U.S. Environmental Protection Agency
Research Triangle Park,
North Carolina 27711
ABSTRACT
•i
A 566m /m (20,000 acfm) permanent installation demonstration
system, consisting of the Air Pollution Systems' High Intensity
Ionizer and a variable throat venturi scrubber (called the Scrub-E)
has been installed on a magnesium recovery furnace. The furnace
produces submicron fume particles of MgO, MgCl2 and ZrC14. The
system is designed to demonstrate the effectiveness of the High
Intensity Ionizer versus high venturi pressure drop on the furnace
emissions. The High Intensity Ionizer array operates stably at
field strengths of 10-15 kV/cm and at velocities in excess of 18m/s
(60 fps) while maintaining high charging efficiencies. The report
covers the system design, technology, applications, and project
developments. An Environmental Protection Agency proposed charged
droplet Scrub-E is also discussed covering the design, technology,
and proposed demonstration program.
60
-------
DEMONSTRATION OF A HIGH FIELD ELECTROSTATICALLY
ENHANCED VENTURI SCRUBBER ON A MAGNESIUM FURNACE
FUME EMISSION
INTRODUCTION
The Particulate Technology Branch of the U.S. Environmental
Protection Agency (IERL-RTP) has had for the last 5 years the
responsibility to evaluate and bring to commercial feasibility,
devices based on new collection principles or concepts, or new
combinations of existing concepts. An emphasis has been placed
- on those devices that are applicable to the metallurgical
industry, and that can be easily retrofitted to existing equip-
ment. In 1977, the EPA issued a competetive procurement to
"...demonstrate at pilot or small full scale the technical and
economic feasibility for the most promising existing novel
particulate collection system for control of fine particulate
emissions from industrial sources." This competetive procure-
ment was won by Air Pollution Systems, Inc. (APS) to demonstrate
the Scrub-E which was tested in the APS laboratory under the
novel device evaluation program. A contract was funded in
September 1977 with APS to demonstrate the Scrub-E on a fine
particulate source (£ 3ym in diameter). A magnesium recovery
furnace which emits submicron particles of MgO, MgCl2, and ZrCl^
was selected as the demonstration site. The Scrub-E is currently
erected (July 1979) with startup to occur within the next few
months. (See Figure 1.)
Figure 1. EPA Scrub-E demonstration system.
61
-------
BACKGROUND
Substantiated research has concluded that fine particulate
(< 3ym in diameter) can result in serious environmental and health
problems. The impact of fine particulates on visibility and cloud
nucleation is compounded by their extremely low settling rate
(mobility) and resultant persistence in the atmosphere. The
above factors are compounded if the fine particulate is composed
of toxic substances. This has caused increased control require-
ments to be set by the air pollution control agencies.
The collection of fine particulate is presently limited to
equipment which is either very large and/or has high operating
costs. The choices include equipment such as precipitators and
baghouses which are large and have a high capital cost, or venturi
scrubbers which are relatively cheap, but require very high energy
consumption to collect fine particulate. The users of fine parti-
cle collection equipment are, therefore, faced with capital and/or
operating costs which are becoming prohibitive, especially in the
case of the venturi scrubber.
APS has taken the venturi scrubber system which represents
the smallest and consequently the cheapest capital cost and
attempted to reduce the operating costs to a reasonable level.
The resultant system, the Scrub-E, incorporates the APS High
Intensity Ionizer with the venturi scrubber system. (See Figure 2).
Ionizer T.R.
To Existing Scrubber
Outlet Test
Station
__n_
rl
Inlet Tent
Station .
High
Intent iLy
Ionizer
Variable Throat
Venturi i
Scrubber
Separator
Smoke llouse
^ Hg Recovery
Furnace
^ Mg CL Furnace
FAN
FAN
Recycle
Water
Tank
Recycle Pump
To Drain
Figure 2. EPA demonstration system schematic flew diagram.
62
-------
The Scrub-E utilizes the combination of electrostatic forces
and inertial forces for the removal of particulate. This combina-
tion maintains an improved efficiency on fine particulate while
requiring one-half to two-thirds the energy consumption of a con-
ventional scrubber (see Figure 3). The design of the ionizing
unit provides for a relatively small add-on device to precharge
the particulate.
-APS SCRUB - E
100
BO
60
8 «0
0
H
1
20
VENTURI SCRUBBER
APs 12 cm WC
PARTICULATE Ti<>2
I 0.472 m3/« LAB DATA |
WATER VOLUME REQ. = 1.06 I /m3
1 2
PARTICLE SIZE-Mm
Figure 3. Particle removal efficiency.
FUNDAMENTAL THEORY
Particle Charging
From classical theory, the amount of charge collected by a
particle in an electric field depends on the strength of the field,
the density of charge present as ions (or in some cases electrons),
the particle radius and dielectric qualities of its constituent
material, and the amount of time available for charging. This is
defined by Oglesby et al.
¦L as follows:
the saturation charge on particles due to field charging
<3S ~ (e/e+2 )TT£0a^E0 £2)
where e = relative dielectric constant
Eo= permittivity of free space
= 8.85 x 10~12 farads/m
a = particle radius, meters
E0 = electric field v/m
63
-------
In practice, the actual charge is time dependent. The
"charging efficiency" Np can be written as:
Np = t/t + T (2)
where T is the charging time constant and depends on the ion
density, N0, and the ion mobility.
The time, t, a particle resides in the charging field depends
on its velocity and the geometry of the field.
The ion density is a function of the supply current, field,
and ionizer geometry so that:
No = it e to y (R + r) EQ (3)
where I = supply current to ionizer - amps
to = width of active region - meters
(weighted average)
R = anode radius - meters
r = cathode radius - meters
Eq = average field V/m
y = ion mobility m /V-sec.
e = electronic charge - coulombs
Further charging of a particle beyond its field saturation
value can be accomplished by diffusion charging resulting from
the thermal motion of ions. This process is most effective for
particles in the submicron range, specifically with r < 0.25 ym.
Diffusion charging (White^) is a time function so that the
charging rate is:
where
dn
dt
=
tt a2 C NQ exp -(ne2/akT) (4)
n
=
number of ions
e
=
electronic charge -esu
k
=
Boltzmann's constant = erg/°K
T
=
° Kelvin
C
=
rms value of ion thermal velocity
t
=
time-seconds
64
-------
Integration yields the following formula for the diffusion
component of particle charge:
akT . , rraCN0e^tx /n..
n = ——— In (1+ — ) (5)
e2
The total charge on a particle is assumed to be the sum of
the field charging component and the diffusion charging com-
ponent. Although this is probably an over simplification, it
serves as a basis for the verification of experiments.
A further component in charging should be noted if there is
a free electron component in the ionizer gap. Because of the
high mobility of electrons, a small percentage of free electrons
can contribute measurably to particle charging with an impercep-
tible change in space charge. This effect was noted by Crynack
and Penney3 for fields near sparkover.
It is reasonable to assume the same effect will be found in
the APS High Intensity Ionizer because of the high fields in-
volved, but the magnitude is not known. Indications of excess
current is shown in Figure 4 where excess charging of particulate
was noted in charge-to-mass measurements of titanium dioxide dust.
INDICATED SATURATION CHARGE
3
8
r
e
Ui
I'
1 1 1
THEORETICAL VELOCITY
> r REDUCTION
CALCUU
JED SAT
JRATION I
:harge"~
TK>2 Dl
iST (mmd
• 11.6 kV/
cm
• •
1
EXPERIM
1
¦NTAL points
1
20
40
60
FP&
SO
100
120
—I—
12
IS
M/S
24
31
37
GAS VELOCITY
Figure 4. Effect of gas velocity on charging efficiency.
65
-------
Venturi Scrubber
A venturi scrubber utilizes inertial or impaction forces
to collect particulate on drops of water or scrubbing liquor.
When liquor is injected into a moving air stream (as high as
122 m/sec), it is atomized or broken up into small droplets. The
relative velocity between the droplets and the entrained particu-
lates supplies the energy for collection. Aerodynamic (drag)
forces act on the fine particles tending to draw them around
obstructions (such as water droplets) in the air stream. Opposing
the drag forces are inertial forces which resist any sudden change
of direction. Since the aerodynamic drag force is proportional to
the plane projected area of a particle and the inertial forces are
proportional to the mass, there is a strong size dependence over
which force will predominate. For a small particle the mass de-
creases faster than the plane area, hence for a given gas velocity
there comes a natural cutoff point where impaction forces become
ineffective in comparison to the drag force (see Figure 5).
too
•s
o BO
(X) MPT DATA.IOHI1ER OH
40 ca A p, J.J to 1.7 !/¦', O.Hm/i
@ APT DATA, lOMIZtK Off
40 cm Ap, 1.3 to 1.7 \/» * 0.30m /t
(7) MM FROMi
SCRUBBCR MHDtaoK. CaJwt ot *1.
PuMJstod J 97 J op. 5-177
40 cm Ap, l.S l/m1
© 100 cm Hp, l.S 1/m* J
@ ISO cm Hp, J.S 1/"J I
•0
70
Aerodynamic Particle Size, mxi
Figure 5. Particulate collection efficiency vs. size.
The above relationship can be expressed as a function of the
impaction parameter / (Ranz and Wong^) .
/IT = fC ^g]** dp
\18nDc J
66
-------
where
V9 "
Pp ,
c =
n =
dp =
DC =
gas velocity - gih/sec
3
particle density - gm/cm
Cunningham correction
air viscosity - gm/sec x cm
particle diameter - cm
collector diameter - cm
Equation (6) illustrates the dependence of collection
efficiency on particle size since the particle diameter dp is
proportional to / Y . Decreasing the collector size, d , or in-
creasing the gas velocity, Vg, will therefore improve cSllection
efficiency. The collection efficiency per droplet can be con-
sidered in terms of target area or percent of the projected droplet
area within which the approaching particulate will be collected
by impaction or interception and not swept around the droplet.
(See Figure 6.)
MJB-MICflON PAimCUS SUPtTMAM AAOUNO OftOPLfT
jeitusMft
INCJttAM) OROPUT TARQST ARIA
u6—
HIQHLY CMAM80 PAftTtClfS qOJ
70 nNCINT RCOUCTIOM IN HNITIMTtON
SCHUtMH WITH PftCCHAAQKR
Figure 6. Mechanisms of collection in scrubbers.
The target area of a droplet is limited to the projected
area of the droplet, hence the collection efficiency per droplet
due to impaction alone can never be greater than 1.
Two means which can be used to achieve greater particle
removal are increasing the relative velocity (resulting in greater
pressure drop) or increasing the water/air ratio. Both of these
means, however, have practical limitations and neither can com-
pletely overcome the reluctance of smaller particulate to be
collected.
67
-------
Application of Electrostatic Forces to Collection
Electrostatic forces have long been used in precipitation
for dry process control and have been studied for application to
scrubber collection and fiber bed collection. Of particular
interest in this case is the work of Kraemer and Johnstone^ and
Zebeio.
The total force for particle collection is the vector sum
of the inertial and electrostatic forces acting on a particle. The
total electrical force is the sum of all individual electrical
forces acting on a particle. Strauss? considers the total force
Fe^ as:
Coulombic force between a charged particle and charged
collector
F = £3-2 (7)
EC 47reQr^
Image force of induced charge on uncharged particle by
charged collector
F = - (£=k\ d3Q2 c (8)
fEI 'e+2 16TT£or5 l8)
Image force of induced charge on uncharged collector by
charged particulate.
Fem = q P - 29 ?Pr? 2 (9)
8ire r-J TT£_(4r^-D^)
o o
Space charge force of repulsion from surrounding like
charged particulate toward a relative charge void (the collector).
Fes = ~ 2402 {10)
q = charge on particle - coulombs
where Q = charge on collecting body - coulombs
r = distance between particle and collector -
cm
N = particle concentration per unit volume
#/cm3
e = dielectric constant of aerosol particle
eQ= specific inductive capacity of space
= 8.85 x 10~21 coulomb2 dyne""* cm-2
68
-------
Because the electrostatic forces are additive, the more
positive forces that can be utilized the higher the force of
collection. Furthermore, each force that can be increased in
magnitude should add to the maximum. As will be explained,
unusually high levels of charging are attained in the APS Ionizer
by field charging (up to 15 kV/cm) with diffusion charging and
electron charging components additive.
APS SCRUB-E DESIGN
Air Pollution Systems, Inc. has developed and patent-
ed a new electrode geometry (symmetry) which is successful in
establishing a highly stable, intense corona discharge. The
cathode is a solid metal disc supported by a structurally reliable
tube section centered in a cylinderical anode arrangement. The
electrode configuration produces a substantially uniform three-
dimensional field which is a principal factor in the greater
electrical stability. This geometry provides operating field
strengths of 10-15 kV/cm compared to 3-6 kV/cm for wire electrode
designs at atmospheric pressures. Figure 7 shows a typical E-I
curve for the ionizer.
r-
a
M/S (M N
i '
0
f
M
m
U
41 AIM (ST
•»
/ NO*
QPMi
RAJ
MAL
VTIIM
Ml
1 IJ 12J 117 20J
APPLIED ELECTRIC FIELD. kV /cm (AVQ.)
Figure 7. Typical E-I characteristics.
In addition to the very high fields, the electrode geometry
produces a concentrated field with ion densities of 109-101" ions/cc,
many times that obtained conventionally. As noted above, the
charge obtained on small particles is dependent on the ion density
and is therefore increased with the higher ion density in the APS
Ionizer. This increases the diffusional charging of the fine par-
ticulate. As a result of the higher fields and ion densities, both
69
-------
the level of charge acquired by the particulate and the rate of
charging are significantly higher. Particle charges approaching
half an order of magnitude higher than conventional is currently
being achieved. Particle charging efficiency was studied as a
function of velocity (residence time). Figure 4 illustrates the
effective charging of Ti02 dust (mmd = 0.5ym) in charge-to-mass
measurements Q/m versus velocity for normal design ranges. The
plotted theoretical and experimental data points show a good
correlation in expected efficiency decrease with the increase in
velocity. The saturation charge (per unit mass) was calculated
to be 87y coul/gram for an 11.6 kV/cm field. This degree of
charging at higher fields was greater than predicted based on
cascade impactor size distribution analysis. The charge levels
above the calculated saturation charge have been theorized to be
the result of direct electron charging due to the very high field
strengths.
The APS High Intensity Ionizer effectively charges the
particulate at velocities of 7-10 times higher than more convention-
al electrostatic electrode configurations. Velocities in excess
of 3 0 m/s can be maintained past the ionizer with charging times
of a few milliseconds. Figure 8 shows the velocity regime of
prime concern for the ionizer versus percent saturation charge.
Charging efficiency was found to be between 80% and 90% for normal
operating velocities. The high current densities are also princi-
pally responsible for the capability to efficiently charge the
particulate at velocities in excess of 30 m/s. This means that
the particles can be charged to a higher level in a smaller volume,
and as can be seen in Figure 2, the ionizer section is small com-
pared to the overall size of the equipment.
100 —
LAB AIR (STP)
90
DESIGN
RANG5
160
90 100
FPS
0 9 15 21 27 30 46
M/S
IONIZER VELOCITY
Figure 8. Calculated particulate charging efficiency.
70
-------
EPA SCRUB-E DEMONSTRATION SYSTEM
Air Pollution Systems, Inc. is currently involved in the
field erection phase (Task 4) on EPA Contract No. 68-02-2666.
The Scrub-E is installed on a magnesium recovery furnace at
Teledyne Wah Chang Albany, Albany, Oregon. (See Figure 1.) As
can be seen, the venturi scrubber-separator, structural steel, and
APS Ionizer module are erected. Currently (July 197 9), all equip-
ment is installed for operation with the exception of electricals,
instrumentation, and ductwork. The scheduled date for startup is
tentatively planned for late September or early October. (Delays
have been caused at the jobsite due to the availability of materi-
als. The ionizer emitting system, anode tubesheet, and venturi
variable damper are fabricated from zirconium 702 to resist HC1
acid attack. All other components are fabricated from PVC, FRP,
or stainless steel.)
Figure 2 displays the demonstration system schematic flow
diagram. The particulate gas stream to the Scrub-E originates
at four sources. Essentially, the magnesium recovery furnaces
(two each) recover nonspent magnesium metal after it has reacted
to reduce ZrCl4 to pure zirconium. The spent magnesium, as HgCl2(
is shipped to another facility for reduction to pure magnesium
metal. The MgCl2 furnace is a concentrator operation to recover
whatever remaining magnesium that is left in pure form. The smoke-
house emission is where the MgCl2 crucibles are heated in a burnout
operation. The magnesium furnaces produce emissions of MgO and
traces of ZrCl^ and MgCl2 as the pure magnesium is ladled from
the furnaces. All of the furnaces, although, emit primarily MgO
fume. Scanning electron microscope photographs (SEM) performed on
the magnesium furnace emissions have indicated the greatest pro-
portion of particulate to be less than 1.0 ym in diameter. Cascade
impactor size distribution analysis of the particulate verified the
SEM data with a D5Q by weight of 1.48 ym. This Dgg had a standard
geometric deviation of 3.0 with a linear regression correlation
coefficient of 0.995.
Objectives for the demonstration program are to remove 90%
of all particulate from 0.01 to 3 ym in diameter from the source
gas stream. An independent testing organization will be brought
on site to test the Scrub-E and to prepare a final report. Accu-
rate particle size fractional measurements require specialty
equipment. Diffusion battery techniques with condensation nuclei
counters (CNC) or mobility analyzers will be used to document the
concentration and size distribution of particle diameters from 0.2
to 0.01 ym. Inertial techniques (cascade impactors) will be
utilized for the particle distribution measurements between 0.3 and
25 ym. The standard mass train measurement techniques (Method 5)
will be used for the determination of total inlet and outlet mass
loading for overall efficiencies and performance.
71
-------
All of the data from the test program will be analyzed
and with supportive information, a final report will be prepared
documenting the system performance and economic factors pertaining
to the design.
RESULTS OF PRIOR SCRUB-E TESTING
Results of APS laboratory and field testing have shown
that the ionizer typically reduces the penetration through the
venturi scrubber by approximately 70% or more, without changing
any other parameter. These results have been verified by Air
Pollution Technology (APT) of San Diego, California. They are
an independent testing organization which evaluated the Scrub-E
under EPA contract number 68-02-1496- These results are reported
in Publication No. EPA 600/2-76-154* . Typical results of the APT
tests are shown in Figure 9 which compare favorably with APS lab
results in Figure 3. The two curves show fractional efficiency
results with and without the ionizer energized. The fractional
efficiencies of particulate were determined using both cascade
impactors and diffusion batteries. As can be seen in Figure 9,
fractional efficiencies are very high on submicron particulate
with the ionizer on. The test aerosol used was titanium dioxide
with an aerodynamic mass mean diameter of 1 ym. As can be
seen, the fractional efficiencies for particles < 0.8ym and > 2.0
92% and > 99%# respectively.
1.0
IONIZER OFF
Z
o
C.I.
i
5
ec
IONIZER ON
t-
lli
z
0.08
0.014-
0.08
1.0
5.0
0.1
PARTICLE DIAMETER -Jim
Figure 9. Penetration vs. particle diameter using both
diffusion battery and cascade impactar, Run 28,
A.P.T. test data.
72
-------
Another significant demonstration of the Scrub-E was on
a urea (carbonyl diamide) prilling tower. The particulate size
distribution was 80 to 90% by weight in the 0.4 to 1.0 ym in-
terval range with an inlet grain loading of 0.09 gm/m^. The
objective of the test program was to reduce the visible emissions
to a zero level which was estimated to be about 0.0034 gm/m3.
Collection efficiencies ranged from 90.75 to 96.25% performed
by the plant personnel, with a final outlet grain loading of
0.0034 gm/m . The pressure drop required to obtain this level
was 19 cm w.g. at a flow volume of 21 m^/min. There were no tests
conducted with the ionizer off, but visual observations revealed
a dense opaque plume without the electrostatics and a zero opacity
or 0.0034 gm/m3 with the ionizer on.
EPA CHARGED DROPLET SCRUB-E DEMONSTRATION PROGRAM
Initially in the performance of EPA Contract No. 68-02-2666,
Air Pollution Systems, Inc. was to include in the Scrub-E demonstra
tion system a positive polarity charged droplet vessel. However,
due to the limitations of space at the job site, this was not
feasible. APS proposed to the EPA to fabricate a smaller system,
85 m^/min (300° cfm) which would be mobile and, therefore, easily
adaptable to plant space availability. The system is comprised of
a single throat High Intensity Ionizer, followed by a venturi
scrubber, and horizontal separator (see Figure 10). The separator
vessel houses the charged droplet system which will allow liquid
to gas contact residence times of 3-5 seconds, dependent on total
gas flow volumes.
m
1IACK
CHAMfD
MOflll
HOC DING
IAHK
»-«
Kit
KVM
MIT HIMMAXM
KM BIHCH
Figure 10. Charged droplet venturi scrubber trailer mounted.
73
-------
The purpose of the charged droplet scrubber will be to in-
vestigate the feasibility of further reducing the pressure drop
requirements of the venturi scrubber, while maintaining the high
efficiency levels on submicron particulate by the introduction of
charged droplets. The reduction of the liquid to gas ratio and/or
the throat velocity will control the amount of charged particulate
passing through the venturi into the charged droplet section for
comparative efficiency measurements. Testing performed in the APS
laboratory has resulted in a method of applying a substantial
charging level to the water drops through the use of higher fields.
The relative charge measured by APS has shown levels approximately
2-3 times higher than obtainable by methods used conventionally.
Theoretically, the objective to achieving the highest percent
removal of particulate or collection efficiency with the charged
droplets is to:
a) maintain the liquid-to-gas ratio most economically feasible
for water consumption rate and induced pressure drop;
b) produce the smallest droplets possible that are economically
feasible to provide the highest liquid surface area per
quantity of liquor and per cubic meter of gas;
c) provide the most efficient charging mechanism that produces
the highest charge-to-mass ratio to the droplets, and
d) provide the highest residence contact time between particles
and droplets that is feasible for equipment size and, there-
fore, inplant space.
To achieve objective c), APS laboratory experiments have
found that the use of higher fields and direct charging of the
water results in a higher charge/mass ratio. It is theorized
that the configuration of direct charging is more efficient since
conduction of electric charges at the liquid source appears to be
along the surface rather than through the bulk of liquid. This
mechanism of direct charging also allows the liquid source to raise
in potential until it is ejected as droplets, thereby becoming
highly charged. Typical values of charging of 200 ym droplets
achieved by APS are 1.43 x 10~° coulombs/gram. This amounts to
approximately one-fourth that of the Rayleigh stability limit,
(see Figure 11).
Figure 11 demonstrates the Rayleigh stability curve as a
function of droplet size. The field strengths were calculated know-
ing the charge/mass for the water droplets. The curve proves that
the Rayleigh limit is equivalent to approximately 4 0 kV/cm, and
therefore higher fields effectively produce a higher charge/mass
to the water droplets.
74
-------
H
12 .
10 .
§.
u 8 .
3
Ul
10
£
g 8
U
s
4 -
RAYLEIGH'S STABILITY
LIMIT-
SATURATION
CHARGE 20 kV/cn
•APS DATA
"IE ToE HE 2^0 fto
WATER D*OPLET SIZE - wm
Figure 11. Water droplet charge.
Droplet surface perturbation stabilities have been predicted
by Rayleigh for the maxiumum charge that a conducting water or
liquid droplet with a given radius can obtain. Hendricks9 examines
Rayleigh's equation for the instability of charged liquid masses
against surface displacements. Those liquid masses, in order to
be stable against surface displacements, must have a charge/mass
ratio of less than:
= 10"6yli
Limit pAV* (11>
75
-------
where
Q/m = coulombs/kilogram
Y = liquid surface tension in dynes/cm
p = liquid density in kilograms/m^
r = droplet radius in meters
_ £
This quantity is equivalent to approximately 4.83 x 10
coulombs/gram:
where
y = 73.35 dynes/cm
p = 1000 kg/m3
r = 10~^m (100 microns)
To achieve objective d), the collection of fine particulate is
usually limited by the droplet interaction lifetime unless slow
velocities and hence high residence times are employed. Melcher,
et al.10 has identified the critical time constants to the collect-
ion of charged fine particles by large charged collectors (droo-
lets). ^
The characteristic time required for the removal of charged
particles by charged drops is given by:
Tc = e0/NQb (12)
The time constant given for the unipolar droplets to be
electrostatically self precipitated, or the effective droplet life-
time is:
Tr = eQ/NQB (13)
where
(MKS units)
e = dielectric constant of free space
o
= 8.85 x 10"12 farads/m
N = drop number density
Q = drop charge coulombs
b = particle mobility which from Stokes
« q/6irnca
76
-------
B = drop mobility which from Stokes
= _Q
6irrir
nc = gas viscosity (Cunningham slip by sub-
script)
g = particle charge - coulombs
a = particle radius - m
r = droplet radius - m
The limitations to overcome are the mobilities of the water
droplets which are greater than the fine particulate (B>b). This
results in a shorter characteristic time Tr than Tc. An increase
in the average charge on the water droplets (Q) decreases the
characteristic time Tc, which decreases the necessary residence
time for higher collection efficiency. This satisfies objective
d) above. An overall reduction in the liguid-to-gas ratio, or
drop number density (N) by reducing the actual L/G and droplet
radius will maintain Tr constant. This satisfies objectives a)
and b) above.
The removal of submicron particulate has proven to be diffi-
cult due to their low mobility, diffusive nature, and unfavorable
inertial properties. Because electrical forces and interactions
are effective in this size range, the addition of electrical
energy to charge collector sites (water droplets) is of great
interest. The addition of a high positive charge to the scrubbing
droplets should not only decrease the characteristic time Tc for
interaction, but increase coulombic forces of attraction (eguation
7) and image and space forces as well (eguations 8-10). It is
believed that these forces will be especially effective on the
smaller particulate fraction (£lym).
HII RETROFITTABILITY - CONCLUSION
The specialized APS electrode geometry applies an intense
electrostatic field. In addition, high ion densities promote near
saturation charging of particulate within a few milliseconds, thus
allowing use of high gas flow rates. It is because of the short
charging times that the physical size requirements of the charging
system can be greatly reduced. This means that the particles can be
charged to a higher level in a smaller volume and easily retrofitted
to existing collection equipment.
The present uses of the Ionizer are: to precharge the particu-
late entering a venturi scrubber, significantly reducing the
pressure drop required to obtain a specific collection efficiency
(the Scrub-E); to precharge particulate entering a dry ESP, thus
significantly reducing the overall size; and to be integrally in-
corporated into a wet ESP device called the Electro-Tube . The
77
-------
Electro-Tube is a pipe type ESP (acid mist ESP) which is
effective where the resistivity or stickiness of the dust prohibits
the use of a conventional unit. Emissions with these character-
istics are often found in the iron and steel, metallurgical, chemi-
cal, pulp and paper, and mining industries.
The utilization of electrostatics with venturi scrubbers as
incorporated in the Scrub-E is expected to expand the scrubbers use
into many application areas which have fine particulate emissions.
As the mass mean diameter of the particulate to be controlled de-
creases, the Scrub-E becomes a more attractive alternative. (See
Figure 9.) This is due to its small size and the high efficiencies
which can be obtained on fine particulate (<3ym in diameter) at low
energy levels. Traditional control equipment which can obtain high
efficiencies on submicron particulate is very large and, therefore,
has a high capital cost as well as utilizing a large portion of
valuable plant space. The small size and relatively low capital
costs, combined with the low energy requirements, favor the use of
the Scrub-E for fine particle control. The EPA demonstration
system will provide the necessary design criteria from data collec-
tion for an HII retrofit to any conventional venturi scrubber
system.
In addition, the EPA charged droplet Scrub-E is being studied
for its eventual use as an industrial air pollution control device.
Such a device should have broad applicability both as a charged
droplet retrofit to existing scrubbers and especially as a device
where scrubbers would be suitable if the collection efficiencies
were sufficiently high.
Among the manufacturers expected to be attracted to the
Scrub-E are the mining, metallurgical, chemical, and paper in-
dustries, salt soaked hog fuel boilers, phosphate rock dryers,
coal dryers, urea prilling towers, and electric arc furnaces.
ACKNOWLEDGEMENTS
The preparation of this paper was supported by the EPA (IERL-
RTP) Contract No. 68-02-2666. Special appreciation to Gerald L.
Sing and Thomas E. Nelson of Teledyne Wah Chang Albany for their
support and cooperation on this project.
REFERENCES
1. Oglesby, Sabert, et al, , "A Manual of Electrostatic Precipitator
Technology, Part I, Fundamentals," EPA report No. APTD 0610/
NTIS No. PB 196 JbO, August 1970.
2, White, H. J., Industrial Electrostatic Precipitation, Pergamon
Press, New York, 1963.
78
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3. Crynack, R. R. and Penney, G. W.,"Charging of Fine Particles
in Negative Corona Near Sparkover", Institute of Electrical
and Electronics Engineers (IEEE) Transactions on Industry
Applications, Vol. 1 A-1C, No. 4, July - August 1974.
pp. 524-531.
4. Ranz, W. E. and Wong, J. B., "Impaction of Dust and Smoke
Particles on Surface and Body Collectors," Industrial Engineer-
ing and Chemistry, Vol. 44, No. 6, pp. 1371-1381, 1952.
5. Kraemer, H. F. and Johnstone, H. F., "Collection of Aerosol
Particles in Presence of Electrostatic Fields," Industrial
Engineering and Chemistry, Vol. 47, p. 2426, 1955.
6. Zebel, G., "Deposition of Aerosol Flowing Past a Cylindrical
Fibre in a Uniform Electric Field," Journal of Colloid Science,
Vol. 20, p. 522, 1965.
7. Strauss, W., Industrial Gas Cleaning, Pergamon Press, Sydney,
Australia, 1974.
8. Calvert, S., et al., "APS Electrostatic Scrubber Evaluation,"
EPA-600/2-76-154a, NTIS No. PB 256-335, June 1976.
9. Hendricks, Charles D., Jr., "Charged Droplet Experiments,"
Journal of Colloid Science, Vol. 17, pp. 249-259, 1962.
10. Melcher, J. R., et al., "Charged Droplet Scrubbing of Submicron
Particulate," EPA-650/2-74-075, NTIS No. PB 241-262,
August 19 74.
79
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DROPLET REMOVAL EFFICIENCY AND
SPECIFIC CARRYOVER FOR LIQUID ENTRAINMENT SEPARATORS
By:
Joseph H. Gavin
Dart Environment and Services Company
Avon, Ohio 44011
and
Frank W, Hoffman
Calfran Industries, Inc.
Springfield, Massachusetts 01101
ABSTRACT
In this paper we will describe a new and highly reliable technique for
determining liquid carryover in a gas stream. In addition, we are presenting
a blade rating factor for ease of comparing various style liquid entrainment
separator blades.
80
-------
For many years various forms of blade-type liquid entrainment separators
have been in use. Construction has varied from an array of simple wooden
slats to complex airfoil shapes formed from plastics and metals. Blade-type
separators are used in numerous applications including pollution control
scrubbers and cooling towers.
Until recently, carryover of these separators was judged by visual in-
spection and true performance characteristics were never really obtained. As
environmental considerations became more demanding, attempts have been made
to monitor carryover by means of sensitized paper, materials balance and
isokinetic sampling. All are prone to error and none address themselves to
actual blade performance characteristics.
Prior to the discussion and data presentations, some comments are in
order. Calfran Industries and Fiberglass Equipment Division of Dart Environ-
ment and Services Company have been jointly studying liquid entrainment
separators for over five years. In that time both experimental and currently
marketed liquid entrainment separator modules from numerous manufacturers have
been studied. Also we have established a reliable technique for determining
carryover and obtaining liquid entrainment separator performance character-
istics.
Before proceeding, a definition of terms to be used in subsequent dis-
cussions is necessary.
SEPARATOR MODULE - A device (blades, vanes, mesh, etc.) through which the gas
containing entrained liquid droplets passes and which as a device separates
the liquid from the gas stream.
SEPARATOR GEOMETRY - This takes into account blade shape, size and spacing.
Throughout this discussion whenever the term "separator module" is used it
fixes the separator geometry. A new geometry, such as a variation in spacing
or different blade, would constitute a new liquid entrainment separator module.
INPUT DROPLET DISTRIBUTION - A percentage occurrence versus droplet diameter
graph giving the entrained liquid droplet size distribution contained in the
gas stream prior to passing through the liquid entrainment separator.
OUTPUT DROPLET DISTRIBUTION - A percentage occurence versus droplet diameter
graph giving the entrained liquid droplet size distribution contained in the
gas stream after passing through the liquid entrainment separator.
REMOVAL EFFICIENCY CURVE - This is a plot of percent of droplets passed
(or removed) according to micron size. The percentage is defined as the ratio
of the number of droplets passing (or removed) through the separator module
over the number entering the separator module for a given size class by
diameter. This graph can be referred to as either the percent passage (or
removed) curve. Percent removal equals 100 minus percent passage.
81
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In order to specify the performance of any liquid entrainment separator
module, the ultimate goal is to obtain a removal efficiency curve for a given
set of conditions.
FACE VELOCITY - This is the average speed which the gas and entrained liquid
droplets enter the separator module. Face velocity is defined as the flow of
air in volume per unit time divided by the cross sectional area of the opening
in which the liquid entrainment separator module is placed.
LOADING - Is the measure of the amount of liquid entrained in the gas stream
before encountering the separator module.
CARRYOVER - A measure of the amount of liquid entrained in the gas stream
after passing through the separator module. The unit used in this discussion
is grains per cubic foot.
RE-ENTRAINMENT - This term refers to liquid removed from the gas stream,
accumulated on the liquid entrainment separator surfaces, and then re-entrained
in the gas stream. It is considered part of the carryover.
The objectives of this discussion are two-fold. First, a detailed de-
scription of a technique to obtain a removal efficiency curve for a liquid
entrainment separator is given. This is followed by the derivation and in-
terpretation of a liquid entrainment separator module rating factor termed
"specific carryover."
It is not a goal of this paper to present all of the results. The intent
here is to describe and illustrate the fundamental principles and point out
the areas of concern that experience and testing has shown to be of major
importance. In keeping with this line of thinking, all reference to details
of a particular liquid entrainment separator module have been deliberately
left out.
The method for obtaining both the input and output droplet distribution
is droplet photography. Actual techniques will be discussed later. Diameters
of the droplets are measured directly from the projected negatives of these
photographs (Figure 1). What is produced eventually is a chart of the number
of droplets into the liquid entrainment separator module and the number of
droplets passing through the separator module for a given statistical size
class. The ratio of droplets out to droplets in is the passage ratio or
percent passage. .These numbers represent the removal efficiency for that
particular size class. Using information gained in a similar manner for other
size classes one can generate a removal efficiency curve. This curve is a
plot of removal efficiency as a function of diameter. This may be done by
plotting Percent Passage versus Droplet Diameter or Percent Removal versus
Droplet Diamenter. While industry classically has spoken in terms of removal
percent, it will be seen later in this discussion, percent passage is math-
matically convenient and more useful in analysis and computations.
82
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In order to generate the data we constructed a test facility. A schematic
representation of the laboratory test apparatus is shown in Figure 2. Figures
3 and 4 are photographs of the horizontal and vertical test units.
Air is drawn into the intake duct by use of a variable speed centrifugal
blower. Total flow of air is monitored by an annubar and inclined manometer.
This system allows adjustment of the face velocity across a given liquid en-
trainment separator module and compensation for variation in pressure drops
from module to module. Pressure drop is monitored by means of piezometric
rings built into the duct.
After the blower we have a droplet source. One or more nozzles introduce
a controlled water spray, co-current, counter current or both. Fine droplet
nozzles are used to produce carryover and coarse droplet nozzles add sufficient
spray to achieve the desired loading.
A pair of viewing ports, allowing a clear optical path across the duct,
are placed before and after the separator module compartment. The camera
unit and source optics may be transferred from one pair of windows to the
other.
The source optics, and ultra high speed strobe unit, are located opposite
the camera. An automated 35 mm camera is used and is tied into a drive unit
so photographs can be taken at predetermined intervals and variable distances
across the duct. All the optics are located outside the duct so nothing
protrudes inside to obstruct the airflow patterns. The camera system is
capable of photographing droplets of one micron diameter and larger.
For most removal efficiency testing the procedures are as follows. A
liquid entrainment separator module is installed and a pressure drop versus
air velocity curve is obtained. A face velocity is then selected and monitored
by the inclined manometer. The desired loading is attained by adjusting the
water flow rate to the nozzles and monitoring the water drain quantity from
the module. Loading and face velocity can be held constant to within 5%.
Once operating conditions have been set and allowed to stabilize, photo-
graphic scans through the input and output windows are made. The set of
photographs before the blades are analyzed to yield the input distribution
and the set of photographs after the mist elminator module are analyzed
to yield the output distribution. A computer performs the necessary cal-
culations to generate the removal efficiency curve. Data exists on effects
of loading, face velocity, spacing, spray distribution, just to name a few.
Data has been taken both in the laboratory and in the field under actual
operating conditions. Results have been verified and the measurement has
been proven accurate and reliable.
TYPICAL DATA - Por a given liquid entrainment separator module the four basic
sets of data for test purposes are illustrated in Figures 5, 6, 7, and 8.
Figure 5 is an input distribution curve. Not that the specific nature of
the input distribution is not critical, However the droplet size and volume
must allow carryover in a quantity that can be measured and provide sufficient
liquid at the separator module to allow for proper loading.
83
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The output distribution is given in Figure 6, Discrete data points from
the input-output data sets are entered into a computer. This program com>-
pares the corresponding size classes and then modifies these ratios by den-
sity values. Density values are obtained from the negative drop counts.
Since the same camera is used for both input and output photographs, the
sampling volumes are identical. The average droplet count per frame then
becomes relative density.
A plot of the ratio of percent passage versus droplet diameter appears
in Figure 7. Note the excellent agreement with a semilog plot. This agree-
ment is evident in all work done on blade type liquid entrainment separators.
Deviations at the small end of the droplet spectrum can occur if sufficient
"fines" are not present to insure good statistical analysis. Variations at
the large diameter end may be due to re-entrainment which, when it occurs,
tends to produce coarse droplets.
To obtain Figure 8, the percent removal curve, one just graphs 100 minus
percent passage. Of all the data then, Figure 7, the percent passage is the
most significant. The reason is two-fold. One, it contains the basic removal
data while, two, being in a convenient mathmatical form, i.e. a straight line.
Accuracy and interpretation of a given removal curve are of next concern.
Some basic ground rules for comparison testing are in order. The first is,
the efficiency curve is very sensitive to the geometry of the liquid entrain-
ment separator. Small variations in spacing can significantly affect the
removal efficiency. Velocity will also have an effect. It will be assumed
blade geometry, velocity, etc. are held constant.
What must be considered, assuming a given operating condition, are speci-
fic, statistical, and sampling accuracy for the input and output distributions.
Specific accuracy is the accuracy of measuring the diameter of any droplet.
With proper magnification this can be held to better than 5%. Statistical
accuracy is controlled by simply measuring sufficient numbers of drops in a
given sample to achieve the desired accuracy. Remember there is no magic
droplet count that is sufficient but the nature of the distribution determines
the statistical accuracy.
Sampling accuracy refers to sample points in the duct itself used to
obtain droplet diameters. This is by far the greatest source of error. Con-
siderable time was spent to arrive at a sampling scheme. Attempting to locate
a monitor point is difficult. Unless total uniform mixing can be achieved no
single monitor point is acceptable. As gas speed changes the "droplet pop-
ulation" moves around. The nozzle fixes the population, but the varying air
speeds redistribute the droplets. On the output side different liquid
entrainment separators produce different spatial variations. By experimen-
tation, we found the most accurate sampling plan is to scan continuously
across the middle of the duct in a line perpendicular to the edges of a blade
system. With appropriate sampling and maintenance of operating conditions,
reproducibility can be held within 10%.
84
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Interpretation of removal efficiency curves must be done after consid-
eration of two facts. First, the total amount of liquid removed depends
on the slope of the curve as well as the y-intercept. For example, one
liquid entrainment separator may have a low percentage removal with a cut-off
that is sharp. Another system may have a high percentage removal, but the
diameter cut-off curve is gradual. Both may remove the same amount of liquid.
Equally critical is the input distribution when trying to assess total
carryover. The liquid entrainment separator is a simple filter. What
actually comes out in each case is dependent on the input distribution.
SPECIFIC CARRYOVER - Specific Carryover is an attempt to compare removal
efficiency curves on a common basis. This is done by using an input droplet
distribution which is uniform. Simply stated, that means there is one drop
of every size present from zero diameter to an infinite diameter. Using a
given removal efficiency curve the normalized volume of liquid passing through
is computed. This value is termed "specific carryover". Now each separator
has attached to it a rating number representing its efficiency seeing this
uniform input droplet loading.
Before starting the derivation it is important to point out that there
are two cases. Case I represents "total cut-off" which is when the pro-
jection of the Percent Passage curve for zero micron droplet diameter falls
below 100%. The "partial cut-off" or Case II is when the extension of the
Percent Passage line reaches 100% at some value for droplet diameter greater
than zero microns. Both situations are illustrated in Figure 9. Consider
first the case of total cut-off (Figure 10).
To begin the derivation one must assume moisture removal is exponential
in nature. This is certainly reasonable, based on past experience, for
blade systems.
We have found that the graph of Percent Passage versus Droplet Diameter
is exponentially decaying and of the form:
Y-Ae-B* (1)
Where Y is the percent passage and x is the diameter in microns. A and B
are shape constants and are determined by actually testing the blades under
a given set of conditions.
In Figure 11, the shaded portion represents those drops which "get
through". The sum of their volumes represents the amount of entrained liquid
that gets through the separator. The key here is that this quantity of liquid
defined as specific carryover, is fixed in amount. Two quantities determine
the carryover. They are the number of drops and the volume of liquid in
each drop.
85
-------
SPECIFIC CARRYOVER - SUM OF (NUMBER X VOLUME)
-------
Integrating, / 2
a/tt ~Bxc I xc 2xc 2 \ -r 3
sc = AM e [ —+ -J-+ -p- ] + ^nixc (10)
Note that if x goes to zero this value for SC reduces to that found in Case
I. °
GENERAL APPLICABILITY - Removal Efficiency curves and Specific Carryover are
very useful in practical evaluation of liquid entrainment separators provided
they are applied properly and interpreted correctly.
The prime use of specific carryover is to compare various blade systems.
The independent variable may be spacing, gas velocity, or even basic blade
design. Specific carryover becomes the dependent variable. It is much
easier to graph specific carryover as a function of blade spacing than it is
to plot a family of curves whose y-intercept as well as slope vary greatly.
Actual carryover in grains per cubic foot can be calculated from the re-
moval efficiency curve. What is essential is the nature and amount of the
input entrainment. For the same liquid entrainment separator, the actual
amount of carryover is predominately determined by the input distribution.
Given this distribution and a removal efficiency curve for the operating
conditions carryover in the system can easily and accurately be calculated.
87
-------
FIGURE 1
Source
Optics
~
Eliminator Input
Spray
Input ,nJ*e,#r
Droplet I
«;?»
Camera
Unit
Vlawlm
Modula Window
u
Removal Efficiency
Test Apparatus
88
FIGURE 2
-------
FIGURE 3
-------
Ill
o
z
III
K
K
3
O
o
o
z
111
o
K
Ul
a
20
16
12
Input Droplet Distribution
DIAMETER IN MICRONS
FIGURE 5
u
z
Ul
cc
AC
3
o
o
o
111
o
BE
111
a.
20
16
12
Output Droplet Distribution
20 4 O 60 SO lOO 120 140
DIAMETER IN MICRONS
FIGURE 6
90
-------
too
REMOVAL EFFICIENCY
DIAMETER IN MICRONS
lO 30 40 SO
60
FIGURE 7
REMOVAL EFFICIENCY
Percent Removal
DIAMETER IN MICRONS
30 4 O SO 60
FIGURE 8
-------
PERCENT PASSAGE
S
IOO
Comparison Curves
REMOVAL EFFICIENCY
2500
SC
«
Total Cut-off
Partial Cut-off
DROPLET DIAMETER ImlcroiMl
O JO 30
40
SO
60
FIGURE 9
-------
e
SPECIFIC CARRYOVER
PARTIAL CUT-OFF
x
Droplet Diam«t«r
TOTAL CUT-OFF
Droplat Di«m«t«r
FIGURE 11
-------
AN EVALUATION OF GRID ROD FAILURE IN
A MOBILE BED SCRUBBER
By
3ohn S. Kinsey
AeroVironment Inc.
Pasadena, California 91107
Steve Rohde
Public Service Company of Colorado
Brush, Colorado 80723
Presented at the
Second Symposium on the Transfer & Utilization of
Particulate Control Technology
July 23-27, 1979, Denver, Colorado
ABSTRACT
The body of information presented in this paper is directed to those engineers
and maintenance personnel involved in the operation of mobile bed scrubbers.
Materials tests were performed on three specimens of austenitic stainless steel in
order to determine the possible causes of grid rod failure in a mobile bed scrubber
controlling particulate emissions from a pulverized coal fired steam electric
generating station. These three specimens were successively placed in an Instron
universal testing machine and loaded in tension until rupture occurred. The results
of these tests, in addition to other supporting data, were used in an attempt to
determine the cause of the grid rod failure during actual scrubber operation. It
was determined that the grade of stainless steel used, the scrubber water quality,
and certain operational problems could be causes of premature failure.
94
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AN EVALUATION OF GRID ROD FAILURE IN A MOBILE BED SCRUBBER
Standard practice for removing flyash from the fluegas of large coal-fired power
plants has traditionally been by mechanical separation and/or electrostatic precipitation
with fabric filtration becoming increasingly more popular. Wet scrubbers have also been
used however, depending on the required application. Public Service Company of
Colorado, at its Arapahoe Generating Station's Unit No. 4, has employed such a scrubber
downstream of the original electrostatic precipitator for the purpose of meeting the
applicable particulate emission standards promulgated by the State of Colorado. This
particular scrubber, manufactured by the Universal Oil Products Company, utilizes the
mobile bed concept to remove flyash from the fluegas stream prior to discharge. Figure 1
is a schematic drawing of the TCA (Turbulent Contact Absorber) scrubber installed on
Unit No. 4.
Since its installation in 1973, this unit has had a history of reoccurring maintenance
problems, some of which are as follows:
o Partial pluggage of presaturator sections
o Ball migration and grid failure
o Droplet carryover from the demister section
o Reheater pluggage
o Corrosion of outlet ductwork
o Inoperative guillotine isolation dampers
One of these problems, that of grid failure, will be discussed in this paper.
An experiment was conducted by the authors to compare certain physical and
mechanical properties of three specimens of austenitic stainless steel rods both with each
other and published values to determine possible causes of premature grid rod failure.
Specimens of both exposed and unexposed stainless steel were tested, using the procedure
outlined below.
DESCRIPTION OF SCRUBBER OPERATION
Fluegas from the electrostatic precipitator is pulled through the scrubber booster
fan and discharged into the presaturator section. Here, moisture is introduced into the
fluegas to cool the gas stream and allow for a smoother gas/liquid interface. From the
presaturator, the fluegas flows into three scrubbing sections in series containing poly-
ethylene balls. Each of these sections are separated by a grid of 3/8" diameter stainless
steel rods. By the turbulent action generated by the fluegas flow, the random motion of
the polyethylene balls, and a countercurrent flow of H-O from the spray nozzles, flyash is
scrubbed from the gas stream. The gas then flows fo a chevron-type demister, where
entrained water droplets are removed. Thereafter reheat potential is added to the fluegas
by steam fed heat exchangers and directed toward the stack.
The scrubbing action described above is quite dependent on a number of factors.
One of the most important of which is keeping the balls in their respective section and not
"migrating" to other sections. This migration is caused by the grid rods failing and, thus,
95
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allowing sufficient space for the balls to work through from one scrubbing section to
another. If enough of this migration takes place, the following operational difficulties can
occur:
o Channeling of gas flow, hence affecting contact area versus flow.
o Increased or decreased pressure differential across affected scrubber beds.
o Increased ball breakage and subsequent plugging of the reheater sections or
recirculation pump screens.
o Combinations of the above.
EXPERIMENTAL
Three specimens of austenitic stainless steel were marked, using the following
designation:
Specimen No. 1 - 316 stainless steel never exposed to environmental conditions
inside a scrubber
Specimen No. 2 - 316 stainless steel in service inside the scrubber for approximately
a six week period prior to replacement
Specimen No. 3 - 317 stainless steel never exposed to environmental conditions
inside a scrubber
These three specimens were successively placed in an Instron universal mechanical
testing machine, and an extensometer mounted on each specimen. The specimens were
then loaded in tension until rupture occurred. A plot of load versus strain was made, using
a strip chart recorder. In addition, certain other information on the chemical composition
of the specimens was obtained by the authors for subsequent analysis.
The data extracted from the strip chart recordings for load versus strain for each
specimen is presented in Table 1. The axial stress was calculated for each load recorded
and results of these calculations are also presented in Table 1. A plot of stress versus
strain, up to and including rupture, is presented in Figure 2 for Specimens 1 and 3,
respectively. A similar graph for Specimen 2 was not prepared, due to the fact that
rupture was premature.
The Modulus of Elasticity was determined for each specimen using a linear
regression analysis of the data presented in Table 1. Results of this analysis is presented
in Table 2. In addition, the tensile strength was also determined and presented in Table 2.
A chemical analysis of the 316 stainless steel used in this experiment was conducted
by an independent metallurgical laboratory.4 The results of that analysis is presented in
Table 3. In addition, it was also determined that the scrubber water contains a fairly
substantial concentration of chlorides and relatively low pH. Chemical analysis shows a
level of ^60 mg/1 of chloride (as NaCl) in the water and a pH « 2.0. The significance of
this information will be discussed later.
96
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RESULTS
The results of the calculations described above were analyzed for their significance
related to the problem at hand. The published values for Modulus of Eleasticity and
tensile strength for AISI (American Iron and Steel Institute) type 316 and 317 stainless
steel were included in Table 2 for comparison purposes.2 As can be seen, the Modulus of
Elasticity determined from experimental data for each specimen is considerably lower
than the applicable published value. In addition, it was also noted that both the tensile
strength and the yield strength for all these specimens far exceeded the standard value.
It should also be noted that the tensile strength of Specimen 2, which was exposed to the
environment inside the scrubber, is quite a bit less than that of Specimen 1, which has
never been exposed to those conditions.
The chemical analysis data obtained, along with the AISI Standard Classification
Ranges, are presented in Table 3.4 As can be seen, the chemical composition of all the
specimens does conform to AISI Classification. It should be noted, however, that
Specimens 1 and 2 are at the low end of the range for chromium, nickel, and molybdenum.
Upon additional research, it was discovered that cold working significantly increases
both the tensile strength and yield strength of stainless steel.2'3 Using published curves of
percent cold work versus yield and tensile strength, it was determined that Specimen //I
has been cold worked approximately 15%, and Specimen #3 has been cold worked 12-18%.
This fact is also supported by other reference data, which gives values for certain
mechanical properties of cold drawn Chromium-Nickel stainless steel wire.2
A careful inspection of each specimen was made prior to testing. Specimen #2
was found to have a wear pattern with alternate high and low areas. This type of wear
could be caused by:
o Crevice-type pitting due to flyash buildup.3,5
o Simple abrasion due to friction caused by the polyethylene balls and associated
flyash being forced against the bars and grinding down the metal.
o Combinations of the above.
EXPERIMENTAL ERRORS
The following experimental errors must be taken into account when evaluating the
data mentioned above:
1. Only one sample of each specimen of stainless steel was tested. Duplicate
samples, or even tripicate samples, are necessary to increase both accuracy
and reproducibility of the data.
2. Specimen //2 ruptured prematurely at a section in the center of the specimen
which appeared to have been severely affected by crevice-type corrosion.
This section of rod was located at a structural member inside the scrubber.
The data for this specimen is, therefore, somewhat questionable even in the
elastic range.
DISCUSSION
According to one author,' austenitic stainless steels are not, as a general rule, highly
resistant to hot sulphurous gases as would be found in a scrubber. It was also learned,
however, that the addition of sufficient amounts of molybdenum gives the required
97
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resistance to reducing environments of sulfurous acid. In addition, when hot, wet SO-
gases and quantities of h^SO^ are present, as found in a scrubber application, the addition
of copper to the stainless steel may also be necessary.3,6 It is also recommended that cold
and hot working, including welding, should be limited to minor operations to keep the steel
below a Rockwell Hardness B 96 3 in order to limit "sensitizing" the metal and causing
intergranular corrosion.6
Also reported in the literature is that stress-corrosion cracking can occur in both
316 and 317 stainless steel when exposed to a combination of sulfurous acid and >100 ppm
of metal chlorides.3,5 As was stated above, the concentration of NaCl exceeds the 100 ppm
criteria by a factor of ^.6.
CONCLUSIONS
From the information presented above, it was concluded that the following factors
contribute to premature grid rod failure at Arapahoe Unit *f:
Grade of Stainless Steel. All of the specimens tested have been cold-worked or cold
drawn in the range of 15%. This coupled with the low molybdenum content can
significantly reduce resistance to the conditions such as found in a scrubber. The
addition of a small percentage of copper to the 317 stainless steel alloy is likely to
increase the service life of the rods.
Scrubber Water Quality. As the chloride (as NaCl) content of the scrubber water
exceeds 100 ppm, stress-corrosion cracking of the grid rods can occur. This creates
a more severe corrosion problem than would be normally present with the low pH
compounding the problem.
Operational Problems. All specimens exhibit the same basic wear pattern due to a
number of possible causes. If the polyethylene balls are being forced up against the
grid bars on a continual basis, a problem may exist with the fluid mechanics of the
fluegas flow through the scrubber.
98
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ACKNOWLEDGEMENTS
We would like to acknowledge Mr. A.C. Bishard of the Colorado Department of
Health and Dr. Robert Pearson of the Public Service Company of Colorado for their
cooperation in the formulation of this paper.
REFERENCES
1 Baumeister, Theodore, Mechanical Engineers Handbook, Sixth Edition, 1958.
2 Mechanical and Physical Properties of Austenitic Chromium - Nickel Stainless Steel
at Ambient Temperatures, International Nickle Company, 1973. : ~~
3 Metals Handbook, American Society for Metals, Eighth Edition, Volume 1, 3uly 1967.
4Test Data, Public Service Company of Colorado, December 1977.
5Javetski, John, "Solving Corrosion Problems in Air Pollution Control Equipment -
Part I," Power, 72-77, May, 1978.
6 Corrosion Resistance of Austenitic Chromium-Nickel Stainless Steel in Chemical
Environments, International Nickel Company, 1963. ~~~
99
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GAS FLOW TO STACK
GENERAL DIAGRAM
OF
WET SCRUBBER
1
REHEATER SECTION
STEAM INLET
REHEATER PLAIN
TUBE SECTION (PRO
SOOT BLOWER
(ADDITION)
REHEATER PLAIN
TUBE SECTION (SEC)
STEAM RETURN
SOOT BLOWER
MAKEUP WATER
PUMPS
DEMISTER WASH NOZZLES
DEMISTER
SECTION
RECIRC. NOZZLES
T7IJTTT* t 7
i« w w >/» i\ i \ r\
EXHAUST GASES
FROM BOILER
PRECIPITATOR
SCRUBBER BALL
SECTION
ISOLATION DAMPER
J
PRESATURATOR
SECTION
1
RECIRC. WATER
FILL LINE
SCRUBBER
BOOSTER
FAN
SLURRY PUMPS
SLURRY
SURGE
TANK
RECIRCULATION PUMP
SLURRY TO ASH PONDS
Figure 1: SCHEMATIC OF TCA SCRUBBER INSTALLED AT ARAPAHOE UNIT NO.
100
-------
20
00
60
40
20
0.2
0.4
0.3
120
100
80
60
40
201
E=24.4(I0)S Iwi
auu=104.1 kii
<7^= 102.5 W
0.05 aio 0.15 0.20
Figure 2: STRESS VS. STRAIN CURVES FOR SPECIMENS 1 AND 3.
101
-------
Table No. 1: RAW DATA - SPECIMENS 1, 2, AND 3
Specimen No. 1 - 316 Stainless
Specimen No. 2-316 Stainless
Specimen No. 3 - 317 Stainless
Steel -
Low Grade -
Unexposed
Steel
- Low Grade
- Exposed
Steel -
3.75% Moly-Unexposed
Load
Stress
Strain
Load
Stress
Strain
Load
Stress
Strain
(kips)
(ksi)
(inch/inch)
(kips)
(ksi)
(inch/inch)
(kips)
(ksi)
(inch/inch)
1.00
9.05
0.000165
1.00
9.05
0.000395
1.00
9.05
0.000132
2.00
18.11
0.00059
2.00
18.11
0.00079
2.00
18.11
0.00003
3.00
27.16
0.00099
3.00
27.16
0.00112
3.00
27.16
0.00086
0. 00
36.22
0.00102
0.00
36.22
0.00105
0.00
36.22
0.00122
5.00
05.27
0.00182
5.00
05.27
0.00182
5.00
05.27
0.00155
6.00
54.32
0.00220
6.00
50.32
0.00220
6.00
50.32
0.00198
7.00
63.38
0.00287
7.00
63.38
0.00270
7.00
63.38
0.00251
7.58
68.63
0.00330
8.00
72.03
0.00369
8.00
76.05
0.00330
9.20
83.30
0.0066
9.00
81.09
0.00703
10.00
90.50
0.00508
10.00
90.50
0.01085
9.00
85.11
0.0132
10.36
93.80
0.00660
10.06
91.08
0.01650
9.60
86.92
0.0198
10.96
99.23
0.01320
10.22
92.53
0.0231
9.68
87.60
0.0260
11.20
101.01
0.0260
10.36
93.80
0.0297
9.70
87.83
0.0290
11.28
102.13
0.0396
10.06
90.71
0.0396
11.30
102.67
0.0528
10.60
95.97
0.0528
11.00
103.22
0.0660
10.78
97.60
0.0792
11.00
103.58
0.0792
10.80
97.78
0.0825
11.06
103.76
0.0990
10.90
99.05
0.1155
11.50
100.12
0.1650
11.08
100.32
0.1085
11.06
103.76
0.1980
11.20
101.01
0.1980
11.32
102.09
0.2086
11.32
102.09
0.2600
11.32
102.09
0.3300
11.22
101.59
0.396
11.10
100.50
0.0158
-------
Table No. 2: COMPARISON OF EXPERIMENTAL RESULTS WITH PUBLISHED VALUES
Mechanical
Property
Specimen
No. 1
Specimen
No. 2
Published
Value for 316
Stainless Steel
Specimen
No. 3
Published
Value for 317
Stainless Steel
Modulus of
Elasticity
Tensile
Strength
21.9(10) 3ksi
102.5 ksi
25.0(103)ksi
S7.8 ksi
28(10)3ksi
SO ksi
24.4(10)3ksi
104.1 ksi
28(10)3ksi
85 ksi
Table No. 3: ACTUAL VS. PUBLISHED CHEMICAL COMPOSITIONS
Chemical
Element
Specimen
No. 1 & 2
**AISI Standard
Classification Range
Specimen
No. 3
**AISI Standard
Classification Range
Chromium
16.18%
16-18%
*
18-20%
Nickel
10.10%
10-14%
*
11-15%
Molybdenum
2.08%
2-3%
3.75%
3-4%
*No Data Available
**Am. Iron and Steel Institute
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OPERATION AND MAINTENANCE OF A PARTICULATE
SCRUBBER SYSTEM'S ANCILLARY COMPONENTS
By:
P. A. Czuchra
FMC Corporation
Itasca, Illinois 60143
ABSTRACT
The proper functioning of a scrubber system is largely dependent on the
operating condition of its ancillary equipment. In this paper the author
discusses the common problems incurred with ancillary equipment and steps
which can be taken during the design and operation of this equipment to mini-
mize maintenance problems. The equipment discussed includes such items as
fans, pumps, clarifiers, drag chain tanks, rotary vacuum filters, motors,
agitators, ductwork, dampers, piping, valves, recirculation tanks, instruments
and motor starters. Maintenance schedules for the various pieces of equip-
ment are reviewed and possible solutions to such problems as corrosion,
abrasion and pluggage are discussed.
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OPERATION AND MAINTENANCE OF A PARTICULATE
SCRUBBER SYSTEM'S ANCILLARY COMPONENTS
SECTION I. INTRODUCTION
A great deal of the operating problems associated with a particulate
scrubbing system are due to the scrubber's auxiliary equipment. The auxiliary
equipment includes everything it takes to make a scrubber system operate such
as pumps, fans, clarifiers, vacuum filters, motors, agitators, ducts, dampers,
piping, valves, recirculation tank, instruments, and motor starters. This
paper will attempt to provide some basic guidelines for the maintenance of
these items, and some suggestions for the correction of the more common failings
of these pieces of equipment. It should be noted that this paper will serve
only as a general guideline and is not intended to replace the maintenance
manual of a specific vendor's piece of equipment.
As in any air pollution control system, a scrubber system is subjected to
a harsh environment. The abrasiveness of some types of particulate causes
equipment such as fan wheels and valves to wear away. Stack gases may also
contain a variety of components that, when wetted, produce compounds which
corrode equipment. This is true of any gas stream from a boiler which fires
sulfur bearing fuel and of gases from many chemical processes. All scrubber
components must be designed and maintained with these considerations in mind.
An important part of any maintenance program is developing a workable
maintenance file. This should consist of all the individual maintenance manuals
of every piece of equipment used in the system, the associated parts identifi-
cation drawings, and a complete and concise maintenance schedule for all
components. A record of all maintenance done on the system should also be
included in the file. A common problem in setting up such a file is that most
vendors supply only one type of literature with their equipment. This is a
combined installation, operation and maintenance manual. The section on
operation and maintenance is usually quite small in comparison to the instal-
lation section. After the installation of the system is completed this manual
is often misplaced or forgotten about. If this occurs, a new manual should be
ordered from the vendor as soon as possible. In order to keep the maintenance
file simple and workable, it is suggested that only those pages relating to the
operation and maintenance of the specific components be included in the file.
All pages relating to the installation of the equipment should be kept in a
separate file.
SECTION II. ROTATING MACHINERY
A. General
Rotating machinery consists of all items such as fans, pumps, clarifiers,
rotary vacuum filters, motors and agitators. Because these items are in con-
stant motion, a great deal of wear can take place. These components require
some amount of regular attention to assure the proper operation of the system.
Key areas of maintenance on rotating machinery are the bearings and any com-
ponents rotating in the fluid stream.
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There are also many other areas of maintenance unique to each type of
equipment. Wherever practical, spares should be included. This is especially
true for high maintenance items such as pumps. Wear resultiny from any abra-
sive solids in the liquid stream often dictates frequent inspection and repair.
When mounted before the scrubber, fan wheels may also exhibit a great deal of
abrasion from the particulate in the gas stream. Unfortunately, however, it
is rather impractical to utilize spare fans in a scrubber system due to the
difficulties in duct arrangements and the high capital cost of fans.
A stock of the more failure-prone components should be maintained. These
include such items as bearings, seals, bushings, pump housing liners (if pumps
with replaceable liners are used), purnp impellers, gaskets, rnake-up water con-
trol valves, pH control valves, pH probes, recirculation tank level control
sensors, and vacuum filter cloths.
B. Fans
As in all rotating machinery, various components of the fan must be peri-
odically lubricated. These components include all bearings on the wheel shaft,
couplings and dampers. Lubrication schedules and types of lubricants should
follow the equipment manufacturer's guidelines. Generally the lubrication
schedule for these items ranges from quarterly to annually. The frequency of
inspection of wear-prone components of the fan will vary with the severity of
the application. On a fan handling relatively clean gas, the wheel and housing
may go for up to one year without any need of inspection. Fans on dirty appli-
cations may require monthly or even weekly inspection. The fan wheel and
housing should be inspected for signs of corrosion, abrasion and particulate
buildup. If the fan is located upstream of the scrubber and the gas tempera-
ture is above 250°F, corrosion is generally not a problem. However, if the fan
is downstream of the scrubber, and the gas contains any corrosive compounds,
or if the fan is before the scrubber and the gas temperature is below the dew
point of any corrosive components in the gas, corrosion will be a maintenance
consideration. If the fan is downstream of the scrubber, moving the fan
upstream of the scrubber may eliminate the problem of corrosion if the gas
temperature is high enough. Should rearranging the fan not be feasible, the
fan will have to incorporate more corrosion-resistant materials of construction.
This requires going to a higher alloy on the fan wheel and possibly installing
a corrosion-resistant liner on the housing. Attempting to reduce corrosion by
collecting the corrosive gas in the scrubber system will generally not work.
This is due to the fact that any condensation on the fan wheel will tend to
pick up any residual corrosive gas which gets through the scrubber system.
This corrosive component when collected by the condensation will eventually
react with the fan materials no matter how small its concentration is.
Abrasion on the fan may be lessened by the installation of replaceable
wear liners on the wheel. Replacing the existing fan with a larger fan capable
of the same CFM and static pressure but at a lower RPM may help to reduce the
abrasion problem. If the fan is located before the scrubber, the installation
of a high-efficiency mechanical collector before the fan may help to reduce any
abrasion problems. Should a mechanical collector already be installed prior
to the fan, the mechanical collector's tubes should be checked for pluggage or
excessive wear. Should the existing mechanical collector be found to be
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working properly, consideration may be given to replacing it with a higher
efficiency mechanical collector or going to a two-stage mechanical collector.
Material buildup on the fan wheel may require a spray wash system to be
installed in the fan. If this fails to improve the situation enough, moving
the fan upstream of the scrubber may alleviate the problem.
Fan bearings should be inspected every eight hours for oil level, oil
color, oil temperature and vibration. Bearings found to be inadequately lubri-
cated may require a greater frequency of lubrication or the installation of a
forced lubrication system. The bearing housing should be inspected weekly for
leaks, cracks and loose fittings. Excessive bearing wear may be caused by a
higher fan operating temperature than was originally specified, or by exces-
sive vibration due to material buildup on the fan wheel. Misalignment of the
bearing mountings could also account for this problem. A major annual inspec-
tion should be made to check all bearing clearances and to detect any sign of
wear, pitting or scoring on the bearings.
Dampers should be inspected semiannually for wear, corrosion, cracking
or looseness. All stuffing boxes should also be inspected. Annual damper
inspection should include the tightening of all fastenings. The travel of
actuator and linkage versus damper leaf and linkage travel should also be
checked at this time. Shaft seals on the fan should be inspected visually
for wear monthly.
C. Pumps
All bearings on the pumps and pump couplings should be lubricated peri-
odically in compliance with the manufacturer's instructions. Many pumps include
oilers which should be maintained at a specified oil level. Some pumps are
equipped with grease lubricated bearings. These should also be lubricated
according to the manufacturer's specifications.
As with fans, the frequency of inspection will vary with the severity of
the application. Normally pump bearings should be inspected at least as fre-
quently as fan bearings with minor inspections every eight hours for bearing
oil levels and for seal leaks and vibration. Semiannual inspection should be
made for leaks, cracks and loose fittings. A major inspection for wear,
pitting, scoring and clearance should be made at least once per year.
The pump housing and impeller along with the seals are components which
are subject to abrasion and corrosion. A properly operating mechanical collec-
tor preceeding the scrubber will aid in reducing the amount of particulate to
be collected by the scrubber. This in turn will reduce the amount of abrasive
solids required to be pumped through the scrubber recirculation system and
should reduce the abrasion on the pump.
Increasing the bleed stream from the scrubber system will also reduce the
concentration of suspended solids handled by the pump. Should neither alterna-
tive be available, a pump with a replaceable rubber liner may be of benefit.
This should aid in minimizing pump replacement costs. The installation of a
water flush in the seals may help to reduce the wear on the seals.
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Corrosion may be combatted by installing a pump constructed of a high
alloy, or a rubber-lined pump, depending on the chemistry of the recirculation
system. Increasing the amount of chemicals added to a scrubber system can
help neutralize the recirculation stream and help preserve the equipment.
However, this may require a significant expenditure for purchasing chemicals.
For example, the addition of caustic will increase the scrubber's efficiency
to collect a corrosive gas such as sulfur dioxide. As more sulfur dioxide is
collected, the pH of the solution drops and increased amounts of caustic are
needed to neutralize it until an equilibrium is reached between the scrubbing
efficiency and the desired pH level. At a recirculating solution pH of about
seven, the collector's efficiency of sulfur dioxide using caustic is about 90%
as compared to a 60% efficiency when scrubbing with water at a pH of four.
Increasing the scrubber's bleed rate may help to lessen corrosion problems by
decreasing the concentration of the corrosive compounds in the recirculation
system.
D. Clarifiers and Drag Chain Tanks
Clarifiers and drag chain tanks fall into a general category of gravity
separation devices. A clarifier or drag chain tank may be used to remove the
particulate matter from the scrubber bleed stream when the particulate is
sufficiently large and has a short settling time.
Clarifiers rake the settled particulate to the center of the unit, where
it is removed from the bottom in a concentrated slurry. The slurry from the
clarifier is pumped to a final discharge or to an additional dewatering device
if a drier final discharge is required. The pipeline from the clarifier must
be periodically inspected to make sure that it is not plugged.
The rake mechanism in a drag chain tank moves the solids up an incline.
As the rake breaks the surface of the liquid, water falls off the rake and a
relatively dry solid is moved up the incline. The solids continue to be pushed
up the incline and over the edge of the tank where they fall onto a pile along
the side of the tank. The pile of solids must be regularly removed from the
side of a drag chain tank at a rate dependent on the application. A front end
loader may be used for this chore; however it is usually better housekeeping
to have the solids discharged into a dumpster and then have the dumpster re-
placed and hauled away when full. The rakes of this equipment require little
maintenance since the speed of rotation is very slow. However, sludge buildup
may put excessive torque on the rake mechanism. If this should occur, the
clarifier or drag chain tank must be shut down, drained and hosed out.
Bearings and gear reducers must be lubricated and inspected according to
the vendor's schedule. This is generally done semiannually. In addition to
the inspection of the bearings, the drive mechanism must be inspected for
temperature rise, sprocket alignment, sprocket wear, chain tension and oil
level. These inspections should generally be done after the first 100 hours
of operation and every 500 hours thereafter.
E. Rotary Vacuum Filters
Rotary vacuum filters are commonly a part of concurrent sulfur dioxide
and particulate control systems but may also be used on the bottom discharge
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from a clarifier in a particulate control system should a more concentrated
discharge be required. In a particulate control system, the fine particulate
matter may pass through an ordinary rotary vacuum filter. A precoat on the
rotary vacuum filter may be required for such an application. In a precoat
filter a layer of filter aid is formed on the filter drum. Once the filter
has been coated with the filter aid, the liquid to be clarified is fed into
the unit. The knife-edge then shaves off the particulate from the filtered
mass along with a small portion of the precoat. After the precoat has been
removed down to a depth where it can no longer effectively filter, the filter
must be shut down and recoated. Precoating of the filter may take as long as
one hour and the filtering cycle may last from 16 hours to one week depending
upon how fast the precoat is removed from the filter. The depth of the precoat
must be visually checked periodically. A spare filter or multiple filters
should be installed to insure continuous operation of the particulate removal
system.
Periodic lubrication of all knife-edge bearings, filter bearings, motor
bearings, gear speed reducers, and vacuum pump bearings must be done in accor-
dance with the vendor's schedule. The filter cloth, knife-edge alignment and
drive belt must be inspected occasionally and adjusted or replaced as required.
The frequency of inspection will again vary with the severity of the applica-
tion ranging from once per week to once per year.
F. Motors - Mechanical
Motors are basically quite simple and reliable devices from a mechanical
standpoint. However, as with all rotating machinery, the bearings require
some degree of care and attention. Motor bearings should be lubricated in
accordance with the manufacturer's specifications. The schedule for this may
vary from 750 to 8000 hours of operation depending on the motor RPM and the
severity of the conditions that the motor must operate under. Extreme care
must be exercised in order to avoid excessive greasing of the motor bearings.
Excessive greasing can cause the bearings to overheat and can contaminate the
motor's windings. When greasing motor bearings, the bearing drain or flush
holes must be unplugged to help guard against over greasing. Some manufac-
turers supply bearings which are prelubricated and sealed for life. This is
advantageous in that overlubrication is perhaps the greatest cause of motor
bearing failure. Bearings should be inspected in the same manner as described
for the fan and pump bearings.
G. Agitators
Agitators may be required on systems where sedimentation of the particu-
late in the recirculation tank is a problem, or if additional mixing is required
to mix any additives into the scrubbing liquor. The bearings and gear reducers
should be inspected and lubricated according to the manufacturer's specifica-
tions. The agitator's shaft and propeller should be inspected occasionally
for wear and corrosion. Should the shaft or propeller exhibit excessive wear
or corrosion, it should be replaced with a higher alloy. On side-entering-type
agitators, the packing and seals should be inspected for signs of leakage and
wear.
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SECTION III. STATIC EQUIPMENT
A. Duct Work
Duct work should be visually checked occasionally for leakage and excessive
flexing. The frequency of these checks will vary with the severity of the ap-
plication, but should be done no less than semiannually. Holes in the duct
work can be the result of a number of causes, but are most often the result of
corrosion or abrasion. For duct work located after the scrubber, corrosion
will be a primary cause of failure. The deterioration of the duct due to cor-
rosion may be alleviated by lining the duct with a corrosion resistant liner.
This, however, may not be possible if the duct has corroded too greatly or if
the gas temperature is too high. Should this be the case, the duct will have
to be replaced with an appropriate alloy or with fiber glass if the temperature
is low enough. Fiber glass has proven to be an appropriate material of con-
struction when located downstream of the scrubber. The duct can be protected
from temperature excursions due to water outages in the scrubber by installing
an emergency spray quench system in the scrubber and supplying it with water
independent of the normal scrubbing recirculation system.
Abrasion is a problem common to duct work located upstream of the scrubber.
The severity of this problem can be lessened by increasing the duct diameter,
thus slowing the velocity through the duct. The addition of a mechanical col-
lector prior to the scrubber system will lessen the particulate loading through
the duct work and therefore aid in reducing duct work abrasion. Should a
mechanical collector already be in the system and abrasion is a problem, the
collector should be inspected to insure that it is in proper operating condi-
tion. If it is found to be working properly, consideration should be given to
installing a higher efficiency mechanical collector. If excessive flexing or
cracking should develop, an expansion joint should be added to the duct work at
that point.
B. Dampers
Dampers, due to their infrequent operation, present special maintenance
problems. Guillotine zero-leakage dampers are perhaps the most effective for
air pollution control systems. Louver and butterfly dampers, however, may be
appropriate for less severe or demanding service. Since the maintenance on all
types of dampers is similar, concentration will be given to guillotine type
dampers.
Lubrication is an important part of any damper maintenance program. The
scheduling and types of lubricants should follow the manufacturer's specifica-
tions. Lubrication is normally done annually on all components of the drive
mechanism and bearings. The blower motor bearings should also be lubricated at
that time if one is included with the damper.
A program of regular inspection must be instituted to help insure that the
damper will operate when required. The damper blades should be moved in and
out of the duct six to twelve inches several times, at a frequency of no less
than once per month. More severe services may require a greater frequency of
this checking. Checks for leakage around the flanges should be made at least
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once per month. All portions of the damper's sealing elements should be in-
spected for wear semiannually. Alignment and operation of the drive mechanism
should also be checked semiannually. On a yearly basis, all drive components,
bearings, damper blades and blowers should be inspected for wear or leakage.
Please note that all these inspections should be carried out with strict adher-
ence to the manufacturer's instructions.
Guillotine dampers should be equipped with an air purge to prevent buildup
of debris in the blade seating surface trough when the blade is in an open
position. It is important, to insure proper damper operation, that the manu-
facturer's schedule for activating these purges be followed. The frequency of
these purges will depend again on the severity of the application; however, it
is recommended that it be done at least daily with six air bursts of five
seconds long with one to one and one half minutes between each burst.
C. Piping
Piping should be inspected at regular intervals depending on the applica-
tion. Piping is prone to three common problems: abrasion, corrosion and
pluggage. Abrasion is common to any scrubber system collecting a hard, insol-
uble particulate, such as coal-fired boiler flyash. Abrasion problems in the
recirculation piping are best alleviated by reducing the concentration of the
particulate that is being run through the recirculation system. This can be
done by increasing the bleed rate from the system, by installing a mechanical
collector ahead of the scrubber if none already exists, or by insuring the
proper operation of an existing mechanical collector. Reducing the pipe velo-
city by installing a larger diameter pipe may also help relieve an abrasion
problem. It must be noted, however, that a velocity which is too slow may pro-
mote settling and pluggage of the pipes. A pipe velocity of four to seven feet
per second is a reasonable compromise between preventing abrasion and pre-
venting pluggage.
Corrosion will be a problem in recirculation systems where the scrubber
is operating on a gas stream containing corrosive gases. The methods of
alleviating a corrosion problem in the recirculation system are similar to
those discussed for pumps. Addition of caustic will ultimately raise the pH
and neutralize the scrubber solution. Replacing the existing piping with
piping constructed of a more corrosion-resistant material may be a more eco-
nomical alternative in the long run. FRP pipe or plastic-lined cast iron pipe
has been successfully used where corrosion is a problem and the particulate
being collected is water soluble, such as sodium sulfate. Plastic-lined cast
iron has an advantage over FRP in that is it is more resistant to accidents on
the outside such as forklifts running into it. On a recirculation system
where both corrosion and abrasion are problems, stainless steel and rubber-
lined cast iron pipes have been shown to be effective.
Pluggage can be a problem on recirculation systems handling a slightly
soluble particulate or if the pipe velocities are too low, thus promoting
settling. On systems where the former is the problem procedures such as
increasing the bleed rate or raising or lowering the pH may help to alleviate
the pluggage. Should the latter be the case, an increase in the pipe velocity
by decreasing the pipe diameter may help to eliminate this problem. Piping on
ill
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a system which is prone to pluggage should be flushed out with clean water on
a yearly basis as a minimum. Consideration should also be given to rodding
out the piping if this is possible.
D. Valves
Valves will be subject to most of the same problems which have been dis-
cussed for piping: that is, abrasion, corrosion and pluggage. The solution
to these problems in valves will be similar to those solutions for the equiva-
lent problems in piping. Installing larger diameter valves, increasing the
bleed rate and installation of a mechanical collector before the scrubber will
all aid in reducing abrasion. Increasing system pH and installing valves of
corrosion-resistant material will help to eliminate corrosion problems. Plug-
gage can be alleviated by increasing the system bleed rate, adjusting pH or
increasing the valve velocity in cases of settling. On a system which exhibits
both corrosion and abrasion, rubber pinch valves have been found to be effec-
tive. The rubber provides resistance to corrosion, and the design of the valve
provides a smooth passage for the liquid stream containing the particulate,
thus minimizing abrasion.
Air-operated control valves may exhibit sticking problems. The problem
may stem from one of two sources. The valve may exhibit some pluggage due to
the particulate in the recirculation system, or it may have some water build
up at the air connection. The latter problem may be reduced by insuring a
supply of dry air to the valve or by regular inspection and drying of the air
connection. The frequency of this cleaning will depend on the amount of mois-
ture in the air supply. Since a control valve can be a rather vital component
in the operation of a scrubber system, it is recommended that the piping be
arranged so that any system containing a control valve can be temporarily
placed on manual operation while the control valve is repaired or replaced.
This is particularly important in the caustic addition system where loss of
caustic may severely damage the scrubber system's components. It is also
important in the scrubbing liquor make-up system where a malfunction of the
control valve may cause the scrubber to run dry, thus damaging the equipment
from excessive temperature. A malfunctioning control valve in the make-up line
may also cause the recirculation tank to overflow. A failure of either of
these control valves and the inability to resort to manual operation may
require shutdown of the entire scrubbing system.
E. Recirculation Tank
The scrubbing liquor recirculation tank in most systems will require very
little maintenance outside of an occasional inspection of the general condition
of the tank and liquid inside. Corrosion may be exhibited on a tank handling
an acidic scrubbing medium. Adding caustic, lining the tank with a corrosion-
resistant liner or replacing the existing tank with one constructed of FRP or
an appropriate alloy can alleviate the problem. If the scrubber is handling
a large, easily settleable particulate and if the tank has a relatively long
retention time, particulate may build up on the bottom of the tank. This
buildup may plug the recirculation pump nozzle and foul the instrument probes.
The addition of an agitator may help to solve this problem. Lowering the reten-
tion time by lowering the liquid level may also help, provided this can be done
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without causing cavitation in the pump. Certain contaminants collected in the
scrubber may cause a foaming problem. This problem may be reduced through
lowering the scrubbing liquor pH by reducing the amount of caustic added to
the system or by adding a slight amount of acid if the scrubbing solution is
originally basic. Should the scrubber system not be able to tolerate a lower
pH, foam-breaking sprays can be installed above the tank or anti-foaming agents
may be added to the tank. Adding more freeboard to the recirculation tank may
be of some benefit in relieving a foaming problem.
Pump cavitation may be a problem if the liquid level in the tank is not
high enough for the pump flow rate and the diameter of the recirculation pump
nozzle. A baffle or bellmouth may be used to prevent such cavitation.
SECTION IV. ELECTRICAL
A. Instruments
Instruments usually require very little maintenance. The philosophy is
generally to replace them as they fail. The accuracy and calibration of analog
instruments should be checked annually. All probes should be cleaned at a
frequency dependent on the application. The level controller in the recircu-
lation tank handling a highly soluble particulate may never require cleaning,
whereas the pH probes on a lime slaker, should lime be used as a neutralizing
agent, may require cleaning daily. It is therefore important to select types
of probes which will be appropriate for the frequency of maintenance. A flange-
mounted differential pressure transmitter, used for level control and located
at the bottom of a recirculation tank, may be appropriate for a scrubber system
operating on a highly soluble particulate. However, if the particulate tends
to settle easily, this type of level controller might require that the tank be
frequently drained in order to clean the probe when a sufficient amount of
solids have settled around it. A top-entering float-type level controller may
be more appropriate for this type of application. All pneumatic lines should
be cleaned and dried of moisture regularly at intervals depending on the qual-
ity of the air supplied to them. Orifice plates should be inspected at least
annually and may require more frequent inspection should the fluid contain any
abrasive solids. Excessive wear on orifice plates may necessitate replacement
with a harder material of construction. Ink and paper in the recorders must
be checked frequently.
B. Motor Starters
Motor starter switches should be lubricated occasionally according to
manufacturer's specifications. Cleaning of all electrical equipment is impor-
tant to insure reliable operation. Dirt in electrical equipment may decrease
heat dissipation and cause poor ventilation. Overheating and insulation break-
down may result from this. Insulation breakdown may also occur if the dust is
electrically conductive. All contacts, transformer insulation and cooling fans
should be inspected, cleaned and dried annually. Dirty environments may require
more frequent inspection and cleaning.
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C. Motors - Electrical
Motors require little attention from an electrical standpoint outside of
an occasional cleaning and check of temperature rise. Dust may be removed by
using dry low-pressure compressed air and a dry cloth. Grease when mixed with
dirt can form a gum which is difficult to remove. Oil can cause a deteriora-
tion of the insulating varnish which in turn can lead to a burnout. Any grease
or oil can be removed with a solvent such as naphtha. A fresh coat of insu-
lating varnish should be reapplied after such a cleaning.
Measurement of the temperature rise can provide an indication of problems
such as overloading, shorts or burned insulation. The three methods available
to measure the temperature rise are thermometer, resistance and imbedded ther-
mocouple. The thermometer method simply consists of placing a thermometer as
close as possible to the windings and core of the motor. This may be done by
removing a head bolt or eye bolt and inserting a common mercury or alcohol
thermometer. The resistance method requires that a Wheatstone or Kelvin bridge
be used to measure the winding resistance at room temperature and then once
again after the motor has operated for four to six hours at full load. The
temperature can then be calculated from the following equation:
T = RH-RC (234.5 + T, ) - (T--T,)
RC
Where T = temperature rise (°C)
Ru = hot winding resistance (ohms)
Rq = cold winding resistance (ohms)
T^ = ambient temperature, motor cold (°C)
T2 = ambient temperature, motor hot (°C)
The embedded thermocouple method involves connecting a bridge-type instrument
to a thermocouple which has been embedded in the windings. This will usually
give a reading directly in degrees centigrade. The actual temperature rise
should then be compared against the manufacturer's recommended temperature
rise for that type of motor. A motor can be too hot to touch and still be
within the proper operating temperature. Therefore it is important not to
attempt to measure the motor temperature by hand. On motors 300 H.P. and
above the thermocouple method should be used with the thermocouple connected
to a recorder and the temperature recorded continuously.
SECTION V. SUMMARY
The ease of maintenance is often a function of proper system design. A
properly designed system will often require little maintenance outside of an
occasional inspection and lubrication. The effects of corrosion and abrasion
must receive a great deal of consideration when designing, redesigning or
maintaining the ancillary equipment of a particulate scrubber system.
An ideal scrubber system from a maintenance and operation standpoint will
have a high-efficiency mechanical collector preceeding a low-RPM forced-draft
fan with replaceable blade liners. Duct velocities to and from the scrubber
will be between 50 and 70 feet per second. Isolation dampers are guillotine
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zero leakage types. Duct work from the scrubber will be FRP or lined carbon
steel with an emergency quench system incorporated into the scrubber.
The recirculation system will consist of a rubber-lined recirculation
pump and an installed full spare. The lining will be of a replaceable type.
The piping will be rubber lined cast iron sized for a velocity of four to
seven feet per second. All valves are rubber pinch valves wherever possible.
The recirculation tank will be constructed of FRP or an appropriate alloy.
The bleed rate will approach that of a once-through system.
A system of this design will not be a zero maintenance system. However,
proper system design and adherence to a routine maintenance program, based on
the individual equipment manufacturer's specifications, will go a long way
toward eliminating many of the problems which have plagued the operation of
particulate scrubber systems.
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SOURCES CONSULTED
Andco/Metrof1 ex Installation, Operation and Maintenance Manual. Buffalo,
Chemineer Agitators Instruction Manual. Dayton, 1978
FMC Corporation Service Manual. Itasca, Illinois 1978
Galigher Parts List and Operating Instructions. Salt Lake City, 1978
Goulds Pumps Installation, Operation and Maintenance Instructions. Lubbock
Texas 1978
Johnson, Karl E. Private Interview. Itasca, December 1978
Komline-Sanderson Installation, Operation and Maintenance Manual. Peapoc.
New Jersey~T978
Leiner, Roger E. Private Interview. Itasca, December 1978
Louis All is Instruction Manual. Milwaukee, 1971
Marinello, R. L. "The Plant Engineer's Bible: Maintenance Operating Manual
Plant Engineering, (October 26, 1978), 114-9
Moersfelder, Allan E. Private Interview, Itasca, December 1978
Susca, Vito Private Interview, Itasca, December 1978
Square D Company Instruction Manual. Middletown, Ohio, 1975
Westinghouse Corporation Installation, Operation and Maintenance Manual.
Boston, September 1977
116
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LOWERING OPERATING COSTS WHILE INCREASING
THROUGHPUT AND EFFICIENCY OF REACTORS AND SCRUBBERS
By:
Richard P. Tennyson
Sheldon F. Roe, Jr., P.E.
Robert H. Lace, Sr.
The Munters Corporation
Fort Myers, Florida 33901
ABSTRACT
This paper is directed to those engineers responsible for the design,
operation and/or maintenance of scrubber equipment for either quality control or
chemical process improvement.
This paper reviews the unique combination of high capacity stacked wetted
film contact packing with high velocity entrainment separation. The unique
geometric designs of both contactor and separator permit as much as 50% or
greater face velocity without the flooding, reentrainment, or excessive pressure
differential to which conventional devices would be subjected.
Detailed examples of combined particulate collection and gas absorption are
discussed, along with suggestions for potential applications in process and
emission control - including both cross flow and counter flow systems.
Upgrading of existing systems to achieve additional capacity or improved
performance efficiency are reviewed. Photos, geometric illustrations and
performance characteristics are presented.
117
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LOWERING OPERATING COSTS WHILE INCREASING
THROUGHPUT AND EFFICIENCY OF REACTORS AND SCRUBBERS
1.0 INTRODUCTION
Emphasis in this paper concerns wetted film contactors, rather than
auxiliary equipment such as mist eliminators and other liquid or gas handling
equipment. The purpose is to provide a basic description of the wetted film
contactor, and to provide some examples of its successful use.
2.0 DISCUSSION
2.1 What is a wetted film contactor?
Historically, spray towers and packed towers have been utilized for gas
and particulate removal from effluent or process gas streams. It could be
said, that the wetted film contactor combines the best features of spray and
packed towers. It overcomes the clogging problems with packed towers, while
maintaining the low pressure drop of spray towers.
However, at least two functions are added to the spray or packed towers.
First, the liquid film moves sideways to the gas flow direction, as shown in
Figure 1. This pays dividends in pumping head and increased residence time of
the liquid. The residence time caused by this lateral movement may be twice
what it is for a packing where the liquid flows straight. Second, the gas
phase all must take identical flow paths in contact with the turbulent mixing
action of the liquid film.
Construction is of corrugated sheets as illustrated in Figures 2 and 3.
Several surface areas are available as shown in Figure 2.
2.2 How is a wetted film contactor used?
Crossflow and counterflow are the two common configurations as illustrated
in Figure 4. Operating parameters must be optimized with the general selection
process illustrated in Figure 5. In addition, each of the above may be used in
staged separations. The Japanese have had considerable success with two stage
mist elimination in crossflow. This has also proven successful in counterflow
at Arizona Public Service and Texas Utilities. The first stage is coarse to
prevent clogging, while the second is finer, utilizing cleaner water.1
Actually, some of the mist eliminators may contain approximately 25 ft^/ft^
of surface area in the wetted film contactor context. However, this area is
arranged for higher internal gas velocities and correspondingly higher pressure
drop. Some molecular transfer (S02, etc.) might occur, as a result of this
surface area.
As applied to wetted film contactors, the staged separation technique is
proposed in a similar manner to the mist elimination.
118
-------
FIGURE 1
FLOW CONFIGURATION OF A WETTED FILM CONTACTOR
FRONT ELEVATION SIDE ELEVATION
ISOMETRIC VIEW
PHOTOGRAPH OF CONTACTOR
MODULES ORIENTED AS IN
DIAGRAM ABOVE
^ = LIQUID PATH
119
-------
FIGURE 2
REPRESENTATIVE SIZES OF WETTED FILM CONTACTORS
PLAN VIEW
SECTION A-A
ELEVATION VIEW
DESIGNATION
FLUTE HEIGHT
H IN
MM
FLUTE WIDTH
W IN
MM
SURFACE AREA
SQ FT/CU FT
SQ M /CU M
APPROX. VOID VOLUME
APPROX. MIN. FREE PASSAGE DIA.
IN
MM
MAX. LATERAL LIQUID SPREAD
IN/FT OF HEIGHT
mm/m OF HEIGKT
PRESSURE DROP
IN WG/FT OF HEIGHT
M BAR/M OF HEIGHT
AT 600 FPM VELOCITY & 8 GPM/SQ FT
12060 19060 27060
0.46
12
1.1
28
68
223
0.4
10
13.9
1158
0.18
1.5
0.75
19
1.9
48
42
138
0.7
18
13.9
1158
0.12
1
1.07
27
3.1
79
30
98
95,4 97.2 98.1
1.1
28
13.9
1158
0.07
0.6
120
-------
FIGURE 3
COUNTERFLOW POLYPROPYLENE MEDIA
USED IN THE S02 SCRUBBERS AT
CHOLLA AND MARTIN LAKE POWER PLANTS
STAINLESS STEEL CONSTRUCTION
121
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FIGURE 4
SURFACE MEDIA GENERAL
FLOW CONFIGURATION
LIQUID IN
GAS IN
LIQUID OUT
CROSSFLOW
GAS OUT
ft
GAS OUT
LIQUID IN
—or
LIQUID OUT
-------
FIGURE 5
OPTIMIZATION OF SURFACE CONTACT MEDIA
EXHAUST GAS
SO2)
GAS PHASE POLLUTANTS (ie
SOLID PHASE PARTIUCLATES
LIQUID PHASE POLLUTANTS (ie.
ACID MISTS, SCRUBBER ENTRAINMENT)
LIQUID
MAKE UP
¦d
RECYCLED
LIQUID
fTOO SMALL—^
I PLUGGING
[LOW CAPACITY)
TOO SMALL—
¦WATER EVAPORATES AND
I SOLIDS REMAIN IN MEDIA—|
|HIGH PRESSURE DROP
fTOolMAET ~~
lLARGE LIQUID HANDLING I
|FACILITIES |
I TOO SMALL— L
|REDUCED EFFICIENCY|
JIARGER TOWER j
Select crossflow vs counterflowI
I SURFACE CONTACT MEDIA i
("select media flute sizFI
1 9 5% VOID AREA, I
I 30 to 120 FT/FT^ |
-
r~* — — — — —«
I SELECT LIQUID FLOW RATE'
i .L- .22 gpm/*[t i
fsEI.ECT LIQUID-SOLI5I~CONCENTRATION}
'SELECT GAS VELOCITY
I 200 - 800 FPM I
{too LARGE— 1
-^¦REDUCED EFFICIENCY— I
iT.&Dnc vrccPT.c —nppprR RPn '
SELECT MATERIAL OF CONSTRUCTION I
STAINLESS ALLOYS OR PLASTICS I
LARGE VESSELS —DEEPER BED'
I LOW PRESSURE DROP— I
I ENERGY CONSERVED__ |
JtOO~LARGE--
"MEDIA FLOODING I
I LARGE WATER TREATMENT FACILITY--'
|HIGH PRESSURE DROP I
[TooTarTe-- ~~]
•SOLIDS BUILD UP!
TOO LARGE" ^
•HIGH PRESSURE DROP— I
,DESTROY TURBULENT !
"MIXING OF FILM AND LIQUID|
|DRAINAGE I
±
CLEANED
EXHAUST GAS
I
SPENT LIQUID
CLEANING PROCESS
123
-------
As a first stage, the media has a surface area of 30 ft^/ft^, and the second
stage 42 ft2/ft3. if very clean liquid is available, the surface area could be
68 ft2/ft3. This fine size is utilized for acid mists as from sulfuric acid,
for instance.
Since the surface contact media operates so well for gas with liquid, it
is logical to expect that the surface contact media has a beneficial effect on
gas flow distribution. The current practice of designing large duct diameters
accentuates the problem of flow distribution.
2.3 How does a wetted film contactor affect particulate removal?
Particulate scrubbing incorporates several wellknown principles in the
mechanics of particle collection, such as condensation, impaction and agglome-
ration. Each of these are effectively applied in the wet film contactor.
Turbulent mixing of the scrubbing liquid and the particulate laden gas will
occur as frequently as every quarter inch of gas, or liquid travel through the
pack. Unlike conventional packed scrubbers the gas and liquid flow through the
ordered contact surfaces is uniform without localized flooded or liquid void
areas, which reduce collection efficiency and increase pressure drop. In Table 1
below some published data of a prominent manufacturer of environmental control
equipment using this media, as well as results of independent tests conducted on
a scrubber for boiler flue gas particulate and SO2 on oil well steam generators
is presented.
Table 1
Illustration of Particulate Removal
Gas Velocity
Through Packing,
fpm
Feet of Pack
or Depth of
Bed, ft
Pressure
Drop,
in w.g.
Efficiency,
%
Manufacturers Data,
Case 1 :
750
3
.5
99% for 6 micron
particles
Manufacturers Data,
Case 2 :
750
3
.5
92.5% for 4 micron
particles
Flue Gas from Oil
Well Steam Generator
Test:
650
4
.75
83% *
*Ash content of crude oil represents 4% of total particulate collected in
sampling train. Sample was subjected to combustion in which it indicated a total
of 93% carbonaceous material, primarily carbon black. This would indicate that
the majority of the particulate collected was in a 5 micron or less particle size
range.
124
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2.4 Field applications and operating performance.
2.4.1 Utility S02 application.
The basic description of surface contact media has been given before —
now some actual case histories of applications will be discussed.
One early application of wetted film contactors now has over 5 years
experience, as was recently reported.^ In this case, the polypropylene wetted
film contactor pictured in Figure 3 was utilized. This system contained 4 feet
height of the wetted film contactor, followed by a two stage mist eliminator.
Comparisons were made with a spray tower, where the spray tower removed
14.4% SC>2, compared to 92.4% for the packed tower. Particulate removal effi-
ciency was 99.8, and 99.7% respectively. The packed tower had an L/G ratio of
6.5 l/m3 (48.9 gal./lOOO acf) and a gas velocity of 2.1 m/s (6.9 ft/sec.).
2
A second example was discussed, where the towers consisted of five stages.
Stages 1 and 2 are sprays, while stage 3 is the wetted film contactor, pictured
in Figure 3. Stages 4 and 5 are coarse and fine mist eliminators. Tower
removal efficiency was 99% with 4 feet of wetted film contactor. However,
throughput was increased 10% by using only 2 feet of packing with greater than
98% SO2 removal.
It should be noted, that both of the preceding sys'-.ems have been constructed
of polypropylene, which has resulted in good operating records.
Part of the dependability of these units is attributed to designs allowing
quick replacement from an ample supply of inexpensive spare parts. Although all
plastics and stainless steels are available, polypropylene is recommended,
because of its low cost and chemical resistance.
2.4.2 Pilot test.
The foregoing steam generator test also illustrates what is considered to
be a major advance in the relationship of pressure drop to removal efficiency.
This example concerns a crude oil (4-5% S, 15.5° A.P.I.) fired boiler with
no special modifications to the boiler and no special removal equipment, other
than the scrubber.
The scrubber contained 4 feet of height of the wetted film contactor
(19060 stainless steel, as described in Figures 2 and 3). Gas velocity through
the wetted film contactor was 600 fpm with an inlet gas temperature of 500° F
and outlet of 120° F. Liquid rates were approximately 8 gpm/ft2 of caustic or
sodium aluminate. Inlet particulate varied from 0.35 to 0.56 g/scf with
analytical results indicating a fine particle size. Inlet SO2 approximately
1500 ppm.
Results showed that, at a pressure drop of 1.25 to 1.50 in. of water, 91 to
95% removal of SO2, and 80 to 85% removal of particulate were accomplished.
125
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3. CONCLUSIONS
Surface contact media of the type described herein should be investigated
more thoroughly with the primary goal of high mass transfer with low pressure
drop.
Experience ranges from full scale utility plants with over 5 years of
outstanding operating records to recent pilot tests demonstrating 80 to 85%
particulate removal and 90 to 95% S02 removal at 1% to 1^ in. w.g. pressure drop.
For the future, staging, as has been demonstrated with two stage mist
eliminators, is recommended. With staging it is possible to prevent clogging
while providing a bed of fine structure near the final effluent point for
minimum discharge. It is possible to provide large beds (1000 ft2) of uniform
"pore" size (25 + 2 mm) where all of the gas must take identical flow paths.
In this regard, surface contact media should also be of value as a flow modifier
enabling economical handling of the large gas volumes being planned in today's
technology.
REFERENCES
1. Heacock, F.A., R.T. Travis. Cholla Station Unit 1 FGD System 5 Years of
Operating Experience. EPA Symposium on Flue Gas Desulfurization; Las
Vegas, Nevada; March 1979.
2. Richman, M. Limestone FGD Operation at Martin Lake Electric Station.
EPA Symposium on Flue Gas Desulfurization; Las Vegas, Nevada; March 1979.
3. Tennyson, R.P., S.F. Roe, R.H. Lace. Improved Mist Elimination Through
Advanced Design Concepts. EPA Symposium on the Transfer and Utilization
of Particulate Control Technology; Vol. 3, EPA-600/7-79-044c;
February 1979.
126
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OPTIMIZING VENTURI SCRUBBER PERFORMANCE THROUGH MODELING
By:
Douglas W. Cooper, Ph.D.
Department of Environmental Health Sciences
Harvard School of Public Health
Boston, Massachusetts 02115
ABSTRACT
EPA's programmable calculator version of a venturi scrubber model was
used to study scrubber optimization. It showed that in the equation relat-
ing penetration (Pt) to pressure drop (Ap), In Pt = -bApa the exponent, a,
becomes <1 as Ap increases (at a constant liquid-to-gas ratio), which we
have shown favors multiple-stage scrubbing over single-stage, for the same
total Ap. An example is given for which two Venturis in series would give
the same collection efficiency as a single stage but require 25 percent less
power. Penetration of various particle sizes for a single stage at a fair-
ly typical liquid-to-gas ratio (Ql/Qg = 1 x 10 ) is correlated with a new
dimensionless group for pressure drop, Apt/yg, formed from the gas viscosity
(Vq) and the particle aerodynamic relaxation time (r). For values of Qi/Qg
within the range of conventional practice, an example is given for which
achieving the same penetration is predicted to require quite different values
of Ap, in apparent contradiction of the "contacting power" hypothesis.
127
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OPTIMIZING VENTURI SCRUBBER PERFORMANCE THROUGH MODELING
INTRODUCTION AND GOALS
Venturi scrubbers are widely used for particulate control; they are more
efficient than such scrubbers as packed columns, sieve plates, and impingement
plates at pressure drops greater than about 10 cm H2O (9.8 x 10^ dyn/cm^ =
980 Pa), according to relationships given by Calvert (1977).* We used a pro-
grammable calculator venturi model to investigate the relationship between
penetration and pressure drop, an implicit variable in the model. This pene-
tration versus pressure drop relationship was used to study economic optimi-
zation of scrubber performance for a specific example and generalized by the
introduction of a new dimensionless group for pressure drop. The relation-
ship between penetration and pressure drop was also studied to investigate
further the hypothesis (Cooper, 1976)^ that multiple venturi stages in series
would be more efficient for collecting particles than would a single stage
having a pressure drop equal to the total across the multiple stages.
METHODS AND RESULTS
The Model
3
A programmable calculator program was prepared by Sparks (1978); it uses
a model for venturi scrubbers developed by Calvert et al. (1972)9 and de-
scribed by Calvert (1977).Penetration (Pt) is predicted by
Pt = exp
"2 «LVLV«'f'
L55Vg
(1)
where
and
o
= liquid volume flow, cm /s
Qq » gas volume flow, cm /s
uG = gas velocity in the throat, cm/s
= liquid density, g/cm
dj = droplet diameter, cm
Uq = gas viscosity, poise = g/cm-s
f = empirical factor, 0.5
K = uGCppdp /9yGdd
a Stokes number, having
128
-------
dp = particle diameter, cm
C = Cunningham slip correction (approximately 1 + 0.16 x 10"^ cm/dn at
STP)
PP
¦ particle density, g/cm^
and
F(K,f) = K_1 £(Kf + 0.7) + 1.4 ln(K^-7) + (2)
Calvert (1977)^ noted that the f factor includes the influence of other ef-
fects besides impaction, including perhaps particle growth due to water va-
por condensation, presence of drop sizes different from the predicted size,
maldistribution of liquid, etc.
Scrubber drop diameter is calculated from the correlation of Nukiyama
and Tanasawa (Calvert, 1977):
0.45 , „ \ 1.5
d - (—V* + °.°597 — (lOOO^A (3a)
d G Vl) ^ \ %}
where
3
PL = liquid density, g/cm
HL ¦ liquid viscosity, g/cm-s
= liquid surface tension, dyn/cm.
3
For water this becomes (Sparks, 1978):
— + 91.8
G \^G/
1.5
dd = 7T + 91.8 Itt1 I (3b)
The programmable calculator uses model equations (1), (2), (3b) to pre-
dict particle penetration by integrating numerically (trapezoidal rule) over
the particle size distribution, assumed to be log normal, having a mass med-
ian aerodynamic diameter of d„ and a geometric standard deviation of a .
© o
The Program
3
Sparks (1978) implemented the venturi scrubber model above by preparing
a program, designed to be written on two magnetic cards, for the Texas Instru-
ments SR-52 programmable calculator. After the first card is read by the
calculator, the user supplies all the input information (T, f, dg, ag, d^.Ad.df,
PL» Qi/^G' uG' vG» PP^ w^ich includes specifying the temperature (T) and the
integration Instructions, the initial particle diameter (dj), the final par-
ticle diameter (df), and the size of the steps in the integration (Ad).
Sparks (1978)^ stated that the following choices are usually adequate: Ad ¦
0.1 for integration from 0.1 to 1.0 pm; Ad « 0.5 for integration from 0.1 to
10 iim; Ad ¦ 2 for integration from 10 to 20 ym. (We used Integration step
widths this small or smaller.) The first card gives the droplet diameter as
well as the pressure drop across the scrubber due to the liquid gas interac-
129
-------
tlon:
Ap = 8.24 x 10"4(Ql/Qg)ug2 (4)
where
Ap = pressure drop, cm H^O.
After the second card is read, the user gets penetration from the integration
interval initially specified; one can also get penetration for adjacent size
intervals for particles which are larger than those in the initial interval
without having to reprogram. (The documentation3 should indicate the "re-
set" instruction must be used immediately before card 2 is read.)
The Model and Some Measurements. For the limited set of data shown in
Table 1 the scrubber model and SR-52 program did very well in predicting
pressure drop and penetration.
Table 1. COMPARISONS BETWEEN SCRUBBER MODEL (SR-52 PROGRAM)
PREDICTION AND MEASUREMENTS (SPARKS, 1978)3
Plant Pressure Drop, cm HqO Penetration
Predicted Measured Predicted Measured
1 46 45.7 0.016 0.015
2 25 25 0.028 0.025
3 105 100 0.004 0.005
The model for scrubber penetration was derived from physical reasoning,
measurements for single collector efficiency, and mathematical analysis. The
f factor is an adjustable empirical constant, generally 0.5. ^ Thus, the model
is semi-empirical. Here we compare results from the model with the experi-
mental results reported by Semrau et al. (1977)-*for their work with ori-
fice scrubbers, which are approximately venturi scrubbers of infinitely short
throat length. These results were reported in terms of the number of trans-
fer units
N = -In Pt (5)
versus the "effective friction loss," equivalent to the pressure drop through
the scrubber due to the droplets, Ap. The experiments were run at flow rates
up to 250 m3/hr (150 ft /min) and at pressure drops up to 320 cm H2O. The
orifice was located in a downward flow. The test aerosol was ammonium fluor-
escein atomized from solution. Semrau et al. (1977) stated it is non-
hygroscopic, only slightly soluble in water. The two aerosols for which com-
plete particle size data (obtained from cascade lmpactor measurements) were
given had dg = 1.05 ym with ag = 1.58 and dg = 0.68 ym with ag = 1.69; these
are only moderately polydisperse.
Sparks (1978)^ compared the penetration data of Semrau et al. (1977)^
with the predictions from this model, using the values of liquid-to-gas ratio
130
-------
and gas velocity of the experiments. The predicted penetration values were
within the spread of the data when plotted versus pressure drop. However,
we have found that the predicted pressure drops at QjVQq » 1 x 10-^ were
usually about twice as high as the measured pressure drops.
It has been argued (Semaru, 1977,^ for instance) that the penetration of
a particular aerosol through a scrubber will be almost wholly a function of
pressure drop alone. Calvert (1977) indicates that liquid-to-gas ratio is
also quite important, with Q^/Qq of about 1 x 10~3 generally optimal. We
used Qi/Qq ~ 1 x 10"-* and calculated the penetration versus pressure drop
for the aerosols used by Semrau et al. (1977)-* over the pressure drop range
10 cm H2O to 330 cm H20, corresponding to gas velocities u^ = 0.35 x 10^ cm/s
to 2.0 x 10^ cm/s. (Other values were T = 25°C, f = 0.5, pL = 1 g/cm^ = p ;
d^, dj, and Ad were chosen to give three significant digits for Ap and Pt.;
Figure 1 shows the two-branched lines used by Semrau et al. (1977)^ to summar-
ize their data and the values of penetration versus pressure drop calculated
by us from the scrubber model. The excellent agreement is puzzling: The
calculations were made assuming a 10"^ liquid-to-gas ratio. In their report,
Semrau et al. (1977)^ give Qj/Qg values that often were much greater than
1 x lO""^. That the penetrations agree fairly well with the predictions at
Ql/Qg = 1 x 10" supports the view that changing the Ql/Qg parameter is not
very important at a given pressure drop, but that seems to contradict other
results of the model which show Ql/Qg to ke quite important.(See Appendix.)
The Two-Branched Curve
One way of correlating the data for penetration versus pressure drop is
to use the transfer unit approach and the form:
N ¦ -In Pt = b(Ap)a. (6)
When this relationship is plotted on log-log paper, a straight line of slope
a results. Cooper (1976)2 noted that data from earlier work by Semrau and
Witham (1974)^ could be approximated not by one line but two, as in Figure
1. Further, he noted that if a relationship of the form of equation (6) held
for each particle size (or for a monodisperse aerosol), it would be more effi-
cient to use two or more scrubbers in series with total pressure drop Ap than
one scrubber with the same Ap, if and only if the value of the exponent a was
less than one. The extension of this principle was also presented for the
two-branched curve; again, whether or not to use multiple stages depended upon
whether or not the value of a was less than unity. Our calculated results for
penetration versus pressure drop at Q^/Qq ¦ 1 x 10" ^ for dg = 3.0, 1.05, 0.68
ym with CTg ¦ 1.05 (essentially monodisperse aerosols) are presented in Figure
2, with the lines characterizing the experimental results of Semrau et al.
(1977). The two-branched nature is still evident, and for the monodisperse
aerosols there exists a pressure drop range for which the value of a is less
than unity, the condition for which series scrubbing is attractive. The model
includes only impaction as a modeled collection mechanism which changes with
pressure drop, so that is is not necessary to invoke other collection mech-
anisms? to explain the existence of two branches, although the curved nature
of the penetration versus pressure drop relation (approximated by the two-
branched lines) may be partially due to the change in droplet diameter with
131
-------
10
T 1 1 1 1 1 I I
Q_
_c
i
ii
Z
tn
• MB
c
3
t-
(/)
C
a
• dfl
= 1.05
V
°"9
= 1.58
*d9
= 0.68
= 1.69
0.1
10
¦ ' ' » I i i i i I I 1 1—L
100
500
Pressure drop, Ap, cmH20
Figure 1. Penetration versus pressure drop: lines represent results
of experiments with orifice scrubber by Semrau et al. (1977);
circles and triangles are results of calculations with scrub-
ber model for Ql/Qq = 1 x 10"*3 and the same aerosol size
distributions used in the experiments.
132
-------
dfl
=
3.0
-g
=
1.05
dg
=
1.05
°g
=
1.05
a
0.68
=
1.05
I I I I I I 11
100 500
Pressure drop, Ap, cm l-LO
Figure 2. Penetration versus pressure drop: model calculations
for monodisperse aerosol at Qj^/Qg ¦ 1 x 10"3 and the
lines characterizing the experimental results of Sem-
rau et al. (1977) for moderately polydisperse aerosols.
133
-------
change in gas velocity (thus pressure drop).
Dimensional Analysis
It is generally useful to correlate data with dimensionless groups ob-
tained from dimensional analysis. For example, pipe flow pressure drop is
usually correlated with a friction factor f(Re), a function of Reynolds num-
ber. Scrubber penetration is dimensionless, so it must depend upon dimen-
sionless groups. Using a standard method of dimensional analysis, Bucking-
ham's Pi,° we found that penetration could be written as a function of some
common dimensionless groups:
Pn = Pn(K, Ap'/PgUg2, QL/QG> Re, dp/dd, dd/D) (7)
where
Re = PguGD/pG, the throat Reynolds number;
D = throat diameter (orifice diameter), cm.
(Other collection mechanisms, such as sedimentation, diffusion, electrostatic
migration, would have enlarged this list of dimensionless groups.) Incorpo-
rating most of these terms make intuitive sense. The impaction parameter (K)
characterizes the collection by a single droplet; the Reynolds number char-
acterizes the flow; the ratio of the particle diameter to the droplet diam-
eter characterizes interception; Q^/Qg indicates the concentration of col-
lectors. It did not seem sensible to characterize pressure drop by Ap'/pGuG^,
however, and this led to consideration of alternate dimensionless groups.
Equation (4) indicates we could characterize pressure drop by using a
combination of two dimensionless groups,&p/(pqUq^)(Q^/Qq), which would then
be constant for those conditions under which the equation (4) was applicable.
Clearly, plotting Pt versus Ap/(pgUQ )(Ql/Qg) was not likely to be informa-
tive, especially where equation (4) holds and the latter group is a constant.
To obtain a more relevant dimensionless group, we turned to the equations
for pressure drop and penetration in an infinitessimal collection volume of
cross-section A and length dL. For N
-------
where
ri(K) = impaction collection efficiency for a droplet.
The ratio of the change in penetration to the change in pressure drop is:
n(K) nn.
(JspGuGZ)f
which suggests that the dimensionless pressure drop should be
r\ (K) Ap' / QiP GUg2 ) f (Red).
There are several approximate forms for f(Re) and n(K); if one chooses Inge-
bo's correlation for accelerating droplets, f(Re) = 55/Re^ (Calvert et al.,
1972)," and n = 0.5K as a rough approximation for impactive efficiency, then
a particularly simple form of dimensionless pressure drop results: (2/55) x
Ap'x/yG or just Ap't/vig where t is the characteristic (or relaxation) time
of the particles, the terminal settling velocity divided by the gravitational
constant g. For spherical particles:
x = Cpd 2/18yr. (11)
P P °
In Figure 3, we have re-plotted the lines characterizing the data from
Semrau et al. (1977)6 and the results of our calculations using the scrubber
model. Here Pt is plotted against the dimensionless pressure drop Ap't/yg.
Note the degree to which this helps correlate experimental results for the
two aerosols, which have diameters such that the ratio of their characteris-
tic times is x(dg = 1.05 ynj)/T(dg = 0.68 pm) = 2.2. (Presumably, if the cor-
relation variable for pressure drop were n(K)Ap/f(Re^)(%>qUq ) the correlation
would be even better.) The correlation shows less agreement at the higher
pressure drops for both experimental and theoretical results, perhaps because
the finest particles in the experiment were measured least well or diverge
from the log-normal approximation, perhaps because the change in droplet
diameter is not included in the correlation Ap'-r/yg perhaps because the ap-
proximation n= 0.5K » TyG/dd is particularly poor here.
An alternative dimensionless group would be the ratio of the particle
diameter to the scrubber cut diameter, the particle size for which the pene-
tration is 0.50. This approach has been outlined by Calvert (1974), and
can sometimes be used conveniently, even though the cut diameter must be de-
termined implicitly from equation (1). For standard conditions, graphs have
been presented to simplify this (Calvert, 1974);^ for other conditions, such
techniques would have to be employed anew. It seems very convenient to know
that to obtain Pt « 0.1 (at Ql?Qq ¦ 1 x 10""*) for a given particle size re-
quires Ap't/uq of approximately 750 in any dimensionally consistent set of
units and that the exponent a becomes less than 1 at Ap't/Pq ¦ 5 x 10 . New
values of t and can readily be calculated for temperature and pressure
conditions different from 20°C and one atmosphere pressure. (Note, for ex-
ample, that Ap'x/yQ has a yg~ dependence.) This correlation group is not
expected to work well when collection mechanisms other than impaction are im-
portant or when f(Re) is not proportional to Re~*.
135
-------
X 0.68
_g
1.69 EX.
1.58 EX.
1.05 TH.
1.05 TH.
1.05 TH.
100
1000
¦ i I i i i )
Dimensionless pressure drop, A p't
Figure 3. Penetration versus dimensionless pressure drop, Ap't/yg:
theoretical results (TH.) for monodisperse aerosols
(df ¦ 0.68, 1.05, 3.0 ym) and the lines characterizing the
experimental results (EX.) for polydisperse aerosols
(dg " 0.68, 1.05 nm) from Semrau et al. (1977)
136
-------
It has been shown by Leith (1979)^® that at a constant pressure drop,
the theoretical minimum penetration is achieved when
Kf ¦ 2xuGf/dd « 1.10 (12)
according to the model represented by equations (1) and (2). With equation
(A), the model gives
Pt = exp[-0.0155 f^Ap'x/yG] (13)
at the optimum combination of Uq and Ql/Qq (thus optimum droplet diameter),
and the number of transfer units is linearly related to the pressure drop.
For f = 0.5 this equation predicts an optimal value of N = 2.9 or Pt = 0.055
in comparison with the non-optimal values shown in Figure 3 (N = 2.3, Pt =
0.1).
A simple example of the possible energy savings for a two-stage scrubber
versus a single-stage device is found by comparing the dimensionless pressure
drop needed for a single stage penetration of 0.01, Ap't/yg ¦ 2000, with the
dimensionless pressure drop for two stages in series (each having 0.1 penetra-
tion), Api'-r/yg + Ap2't/uq ¦ 750 + 750 - 1500. Thus the two-stage device con-
sumes 1500/2000 - 0.75 times the single-stage scrubber, a saving of about 25
percent. As is well-known, energy costs can predominate for high-energy
scrubbers, so this saving could be quite significant.
Scrubber Optimization
One of the reasons for seeking penetration as a function of pressure
drop is that the pressure drop times the volume flow rate divided by the fan
efficiency is the power consumption of the scrubber, usually the dominant
annualized cost item for high-energy (Ap > 50 cm H2O) venturi scrubbers.
Scrubbers can be designed to meet a given efficiency by specifying the throat
velocity (viq) and the liquid-to-gas ratio (Qi/Qq) » thus the pressure drop.
These three factors, with total gas flow rate, are needed to predict capital
and operating costs.
We have used the programmable calculator model to investigate the eco-
nomically optimal design of a scrubber to achieve a penetration of 0.05 on
the effluent from a coal-drying operation. In so doing, we have found a
wida variety of pressure drops which are predicted to give the same collec-
tion efficiency, in apparent contradiction to the "contacting power theory."
See, for example, Semrau (1977).? Table 2 from Dirgo (1979)^ gives the re-
sult of a literature search for venturi scrubber operating conditions con-
sidered typical. The velocity range is 0.18 x 10^ cm/s to 1.8 x 10 cm/s and
the liquid-to-gas volume flow ratios ranged from 0.25 x 10"^ to 5.5 x 10" .
The optimization described next was carried out within these ranges.
The conditions for the coal drying process were adapted from Calvert et
al. (1972)® and the Particulate Pollutant System Studya gas flow of
150,000 m /hr at 70°C containing particles (after precleaning with a cyclone)
of pp - 1.35 g/cnr*, dg ¦ 3.0 ym and ¦ 4.4, a very polydisperse aerosol.
Table 3 shows the results of more than two score trial-and-error calculations
137
-------
by Dirgo (1979) seeking penetrations of 0.050 (to two significant figures).
It is evident that penetration is not predicted to be a single-valued func-
tion of pressure drop within the range of throat velocities (yG) and liquid-
to-gas ratios (Qj/Qq) encountered in practice. From these results, the most
efficient Ql/Qg ratio in terms of pressure drop minimization was 1.2 x 10" 3
in agreement with the recommendation of Calvert (1977).^ The sensitivity of
the penetration to changes in the pressure drop for Ql/Q/-. = 0.9-1.25 x 10" 3
was APt = 0.0003 per A(Ap) = 1 cm HoO, which indicates tne predicted penetra-
tion penalty incurred by lowering the pressure drop, under these conditions.
Table 2. TYPICAL VENTURI SCRUBBER OPERATING PARAMETERS
Gas Velocity (ur:)3
Liquid/Gas Ratio(Qj_/Q^)
Source
150-160 ft/s
5-40 gallons/1000 ft3
Hesketh (1974)
4572-18288 cm/s
0.68-5.47 x 10~3 m3/m3
5000-15000 cm/s
0.25-4.0 x 10"3 m3/m3
Calvert et al.(1972)^
3500 ft/min
1778 cm/s
3-20 gallons/1000 ft3
0.41-2.74 x 10"3 m3/m3
(Kinley and Neveril,
1976)13
200-600 ft/s
5-7 gallons/1000 ft3
(NAPCA, 1969)14
6096-18288 cm/s
0.68-0.96 x 10" 3 m3/m3
200-400 ft/s
5-20 gallons/1000 ft3
(Calvert, 1977)4
6096-12192 cm/s
0.68-2.74 x 10"3 m3/m3
Parameters are reported first as they appear in the literature. They have
been converted to units of cm/s for Uq and m /m for Q-^/Qq for purposes of
comparison and because these are the units required by the SR-52 venturi
scrubber model.
Rather than minimizing pressure drop at a given efficiency, one would
prefer to minimize annualized costs (these include capital and operating ex-
penses). Dirgo (1979)*^ used cost equations from Calvert et al. (1972)9 and
from Kinley and Neveril (1976), 3 with costs inflated to May 1977; he noted
that his final capital cost estimates were in the high end of the range of
the reported values. The annual costs were dominated by the operating costs
and the operating costs were dominated by the power costs. Figure 4 has a
series of budget lines for operating costs (OC) from his analysis; these
show the operating costs for having the indicated Ap and Qj^/Qp (exclusive of
water pollution costs, which should depend upon liquid recycling practices).
By finding the poipt of tangency between the budget lines and the constant
penetration curve one can select the operating conditions which minimize
operating costs; here the predicted minimum in OC occurs for Q^/Qg between
1.1 and 1.2 x 10" 3 and Ap = 114 cm H2O. For high-energy Venturis, the op-
timal conditions for minimizing operating costs will be near those for
138
-------
180
100
#0.6M/yr
#0.55M/yr
#0.5M/yr
$0.45M/yr
#0.4M/yr
Ql/Q6 .10
-3
Figure 4. Locus of predicted pressure drop and liquid-to-gas
ratio values giving penetration = 0.050 for coal
drying process control example, along with budget
lines for predicted operating costs.
139
-------
minimizing pressure drop, because of the dominant cost of energy. For low-
energy scrubbers other operating costs could make the budget lines less
nearly horizontal in plotting Ap vs. Qj/Q(;> an(* the minimum cost might be at
a substantially higher pressure drop than the minimurr pressure drop.
Table 3. QL/QG AND Ap FOR Pt=0.050 FOR HYPOTHETICAL COAL DRYING APPLICATION
VQg
UG
Ap
(lO"3)
(cm/sec)
(cm H2O)
0.645
18000
172.2
0.75
15000
139.1
0.875
13075
123.3
1.00
11900
116.7
1.125
11150
115.2
1.25
10550
114.6
1.44
10000
118.7
1.82
9250
128.3
2.25
8750
141.9
2.70
8500
160.7
19
For the method of steepest ascent, one makes a small change from the
"base" case in one variable and measures the rate of change of the penetra-
tion with respect to pressure drop, A(Pt)/A(Ap). Next, the other variable
is changed from the base case and the rate of change determined. A new
base case is chosen, by changing either or both variables, guided by which
variable gave the greatest rate of change of penetration per change in pres-
sure drop. In Table 4, an example of a similar, simple optimization proce-
dure is given. The first column lists the operating conditions, their changes
(if any), and the results of such changes. The next columns give the values of
the variables in the first column for: the first base case; the first trial;
the difference between the base and trial 1; the second trial; the difference
between the base and trial 2; and, finally, the new base, determined by rais-
ing the liquid-to-gas ratio by 10 percent and lowering the velocity to keep
the pressure drop constant (guided by equation (4)). The example shows a de-
crease in penetration from 0.014 to 0.010 for this monodisperse aerosol, while
the pressure drop has been returned to its original value.
DISCUSSION
This discussion will focus on the implications of these results for mul-
ti-stage versus single-stage scrubbing. As has been shown elsewhere (Cooper,
1976),^ if the exponent in the expression -In Pt - N ¦ bApa is less than one,
it is optimal (for energy consumption) to use multiple venturi stages in ser-
ies rather than a single stage. The same reference gives the mathematical
criteria for single-stage versus multiple-stage scrubbing for a two-branched
140
-------
curve; even without the criteria, it would be easy to determine through trial
calculations whether or not multiple stages should be used, if Ap is high
enough that the exponent a is less than one. (The penetration through the
series is just the product of the penetrations through the individual stages.
One can also just sum their transfer numbers.) If the change in the expon-
ent a as pressure drop is changed were due to aerosol polydispersity, then
this analysis would have been incorrect, as noted previously (Cooper, 1976),
but that has been shown not to be the case (see Figure 2).
Table 4. EXAMPLE OF SCRUBBER OPTIMIZATION ITERATION
(dg - 1.05 pm, Og - 1.05, T - 25°C, f = 0.5)
Conditions
base
trial 1
(change)
trial 2
(change)
new base3
ug, 10^ cm/s
1.0
1.05
1.0
0.953
A(ug), 10^ cm/s
0.05
0.0
Ql^g » 10"3
1.0
1.0
1.1
1.1
A(Qj^/Qg), 10 ^
0.0
0.1
Pt
0.0142
0.0113
0.0080
0.0104
-A(Pt)
0.0029
0.0064
Ap, cm H20
82.4
90.85
90.64
82.4
A(Ap), cm H2O
8.45
8.24
-A(Pt)/A(Ap)
0.0034a
0.000753
a-A(Pt)/A(Ap) is greater for increasing Q^/Qg than for increasing ug, so
(Ql/Qg) should be increased to form new base, keeping Ap constant.
Superior collection by multiple-stage venturi scrubbers over single-stage
Venturis would follow from diminishing returns, in terms of collection effi-
ciency, at high pressure drops (high throat velocities). A convenient way to
picture this is to imagine that one raises the pressure drop in a test scrub-
ber by increasing gas velocity only (constant Q^/Qq)• If droplets were of
optimal size, then equation (.is) shows that the number of transfer units will
be linearly related to pressure drop. As one lowers the velocity, the num-
ber of transfer units will decrease more strongly than linearly because not
only is the pressure drop decreasing, but the droplets are becoming non-opti-
mal in size. As the velocity is increased, the number of transfer units will
increase less than linearly with pressure drop because the droplets again are
becoming non-optimal in size.
Calvert (1977)* pointed out that for large particles for which the single
droplet collection efficiency is almost unity, the penetration is limited by
Ql/Qg. A multiple-stage scrubber effectively multiplies the effectiveness of
Qt/Qg by th© number of stages employed, if all the water is added upstream of
tne first stage. Although a previous analysis (Cooper, 1976) showed that it
was optimal in terms of pressure drop to divide the pressure drop equally
141
-------
among the stages of a multiple-stage scrubber, it may not be best to have
Ql/Qg be the same for every stage...perhaps the first stage(s) should be at
low Ql/Qq and more water should be added immediately upstream of successive
stages, to optimize the product of droplet area, collection efficiency, and
path length for the particles reaching the latter stages, while keeping the
same pressure drop per stage.
Another advantage of multiple-stage scrubbing not evident in this analy-
sis is that particle growth by condensation may take place in the first
stage(s), thus facilitating capture; condensation may be enhanced by expan-
sion and cooling at the diverging portion of the first stage(s). Also, the
multiple-stage device would tend to have longer contact time for gaseous and
particulate diffusion.
Certainly, capital costs must be taken into account, but for submicrom-
eter aerosols that have to be scrubbed at high energy, the power savings from
multiple-stage scrubbing may outweigh increased construction costs.
CONCLUSIONS
q
The scrubber model developed by Calvert et al. (1972) and programmed by
Sparks (1978)^ was quite useful. The two-branched nature of a log-log plot
of transfer units (N = -In Pt) versus pressure drop was shown not to be an
artifact due to the use of polydisperse aerosols nor to require for explana-
tion of its shape any other collection mechanism than impaction. (Actually,
the fitting of two lines to the curve of log N versus log Ap is only one of
many possible approximations by any number of straight lines.) A dimension-
less pressure drop, Ap'T/yg, was found to be useful for correlating experimen-
tal data and theoretical predictions of penetration. The model predicts that
penetration is sometimes not uniquely related to pressure drop, even within
the ranges of gas throat velocity and liquid-to-gas ratio found in practice.
It was shown that Q^/Qq and Ap can be adjusted to minimize operating costs
with the help of the model. Results of the simulations support the hypothe-
sis (Cooper, 1976)^that multiple-stage venturi scrubbers would be more effi-
cient than single-stage venturi scrubbers for high-energy scrubbing (at the
same total pressure drop).
ACKNOWLEDGMENTS
The author gratefully acknowledges the aid of Richard C. Antonelli and
John A. Dirgo, both graduate students at the Harvard School of Public Health,
in carrying out many of the SR-52 programmable calculator runs, and the fi-
nanical support of the U.S. Public Health Service through its program grant
for environmental management,5 D04 AH 01475.
142
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REFERENCES
1. Calvert, S. Scrubbing. In: Air Pollution. Stern, A.C.(ed.). New York,
Academic Press, 1977.
2. Cooper, D.W. Theoretical Comparison of Efficiency and Power for Single-
stage and Multiple-Stage Particulate Scrubbing. Atmos Environ 10:1001-
1004, 1976.
3. Sparks, L.E. SR-52 Programmable Calculator Programs for Venturi Scrub-
bers and Electrostatic Precipitators. EPA-600/7-78-026. U.S. Environ-
mental Protection Agency. Office of Research and Development. Wash-
ington, D.C. 1978.
4. Calvert, S. How to Choose a Wet Scrubber. Chem Eng, August 29, 1977.
5. Semrau, K.T., C.L. Witham, and W.W. Kerlin. Relationships of Collection
Efficiency and Energy Dissipation in Particulate Scrubbers. (Presented
at the Second Fine Particle Scrubber Symposium, New Orleans, May, 1977.
Sponsored by the U.S. Environmental Protection Agency.)
6. Semrau, K.T., C.L. Witham, and W.W. Kerlin. Energy Utilization by Wet
Scrubbers. EPA-600/2-77-234. U.S. Environmental Protection Agency.
Office of Research and Development. Washington, D.C. 1977.
7. Semrau, K.T. Practical Process Design of Particulate Scrubbers. Chem
Eng, September 26, 1977.
8. Baumeister, T., E.A. Avallone, and T. Baumeister, III, eds. Marks' Stan-
dard Handbook for Mechanical Engineers. New York, McGraw Hill, 1978.
9. Calvert, S., J. Goldshmid, D. Leith, and D. Mehta. Scrubber Handbook.
EPA-R2-72-118, NTIS No. PB-213-016, NTIS, U.S. Department of Commerce.
1972.
10. Calvert, S. Engineering Design of Wet Scrubbers. J Air Pollution Control
Association. 24:929-934, 1974.
11. Dirgo, J.A. Personal communication. Harvard School of Public Health.
Boston, MA. 1979.
12. Hesketh, H. Understanding and Controlling Air Pollution. Ann Arbor,
Ann Arbor Science Publ., 1974.
13. Kinley, M.L. and R.B. Neveril. Capital and Operating Costs of Selected
Air Pollution Control Systems. EPA-450/3-76-014. U.S. Environmental Pro-
tection Agency. Washington, D.C. 1976.
14. National Air Pollution Control Administration. Control Techniques for
Particulate Air Pollutants. AP-51, United States Department of Health,
Energy, and Welfare. Washington, D.C. 1969.
143
-------
15. Midwest Research Institute. Particulate Pollutant Systems Study. Volume
III, Handbook of Emission Properties. Contract CPA 22-69-104. U.S. En-
vironmental Protection Agency. Washington, D.C. 1971.
16. de Neufville, R. and J.H. Stafford. Systems Analysis for Engineers and
Managers. New York, McGraw-Hill, 1971.
17. Sparks, L.E. Importance of Particle Size Distribution. (Presented at
the Symposium on the Transfer and Utilization of Particulate Control
Technology, Denver, CO., July 1978. Sponsored by the U.S. Environmental
Protection Agency.)
18. Leith, D.H. Personal communication. Harvard School of Public Health.
Boston, MA. 1979.
19. Cochran, W.G. and G.M. Cox. Experimental Designs. New York, John Wiley
and Sons, 1957.
APPENDIX
The similar performance of some scrubbers at a given pressure drop but
different liquid-to-gas ratios (equivalently, different velocities) seemed to
contradict the results in Figure 4, which showed quite different performance,
for different Ql/Qq at the same pressure drop, on an aerosol (dg ¦ 3.0 ym) .
Using the Calvert model as given in the body of the paper, and the pressure
drop group, one can write the number of transfer units as (Leith, 1979):^-®
N = (4/55)(TAp'/yG)K"1F(K,f). (A.l)
Therefore, the change in the number of transfer units at constant pressure
drop becomes:
AN - (4/55)(TAp'/yG)A[K"1F(K,f)] (A.2)
The change in number of transfer units is most sensitive to changes in the
bracketed group (K~*F(K,f)) for large values of tAp'/Pc (large particles,
high pressure drops or, simply, high collection efficiencies). For a par-
ticle of 1 ym aerodynamic diameter at Qt/Qq ™ 1 x 10"^ and lu ¦ 0.5 x 10 cm/s,
increasing the velocity by 10 percent (decreasing Ql/Qg ^y , Percent to keep
pressure drop constant)gave only a 4 percent decrease in K""*-F(K,f), changing
the number of transfer units by 0.055 and producing a penetration of 1.06
times the original, a negligible change. For the same conditions, a particle
of 5 ym aerodynamic diameter would have xAp'/yg he 22 times larger, AN « 1.01,
and the predicted penetration would increase three-fold, from Pt ¦ 0.0041 to
Pt = 0.011, a quite appreciable change.
144
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THE IMPACT OF HUMIDIFICATION CHAMBER PHYSICS
ON WET GAS CLEANUP SYSTEMS
By:
D.
P.
Bloomfield
M.
L.
Fins on
G.
A.
Simons
K.
L.
Wray
Physical Sciences Incorporated
Woburn, Massachusetts 01810
ABSTRACT
The performance of the humidification chamber in wet gas cleanup
systems is central to the effectiveness of the entire system. We have
presented the theoretical analysis and the results obtained from a pre-
liminary investigation of humidification chamber processes namely water
droplet evaporation and tar condensation. The major findings of the analy-
sis are that sharp temperature gradients and the absence of dust (seed
nuclei) in the chamber lead to the formation of fine tar droplets which may
be impossible to separate from the gas phase, A moderate temperature
gradient will favor heterogeneous tar condensation and yields tar droplets
of a more readily collectable size.
145
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THE IMPACT OF HUMIDIFICATION CHAMBER PHYSICS
ON WET GAS CLEANUP SYSTEMS
The removal of tars and particulate materials from the effluent
stream of a coal gasifier is of major importance in both a power genera-
tion sense and in an environmental sense. Gas cleanup systems represent
a major cost in powerplant design and their effective operation is crucial to
the successful operation of coal gasifier powerplants. The wet gas cleanup
system presently under study at the Morgantown Energy Technology Center
is an attractive approach to tar, particulate and sulfur removal.
Figure 1 shows the relative location of components in the wet gas clean-
up system. The humidification chamber in Figure 1 is a direct contact heat
exchanger where the gasifier effluent is cooled by a recycle liquor spray. As
the gas cools, high boiling tar condenses. Tar condensation may occur by
homogeneous nucleation or heterogeneous condensation. In the former case,
under conditions of intense subcooling, tar droplets form from the gas phase
The nature of the process is such that a very large number of extremely fine
droplets form. The formation rate can be extremely fast. Once formed,
the droplets coalesce, but this process is relatively slow. As a consequence,
the tar particle droplets at the exit of the humidification chamber may be
too small to capture in downstream scrubbers. If the temperature profile
is more gradual, and particularly if a large quantity of ash, is present, hetero-
geneous condensation can occur. In this case, vapor phase tar condenses on
dust nuclei. The resulting tar droplets can be relatively large and easier
to capture in the downstream separators.
.FIRST LAW ANALYSIS
Using the operating conditions specified in Ref. 1, one may determine
the composition of the gas stream entering the humidification chamber. For
an inlet gas flow rate of 180,000 SCFH, the total molar flow rate into the
humidification chamber is about 535 lb mol/hr. In order to cool the inlet
gas stream to an exit temperature of 320°F, about 167 lb mol/hr of water
must be added. This effectively setB the mol fraction of water at the exit
of the humidification chamber at about 38%, the total exit molar flow rate
being about 702 lb mol/hr.
By setting the water vapor exit mole fraction, the heat balance also
sets a limit on the system pressure. Increasing the pressure to 234 PSIA will
saturate the exit stream. Above this pressure decreasing amounts of water
are evaporated and exit temperature increases. In order to operate above 234
PSIA, either the gasifier steam to carbon ratio must be reduced, or the
allowable exit temperature from the humidifier must be increased above 320°F.
146
-------
42"
"Wellman
Galusha
Gasifier
clone
Coal Air
Dust
Humid
Q
Liquor
Pump
Decanter
Turb
Exhaust Air
Dis-
engage
Electro-
static
Precip.
1
COS-*U2S
I v
60% TOTAL TAR
Tar
Polish
ft
Final
St ret ford
Process
Filter
1
Sulfur
Fig. 1 Low Temperature Gas Cleanup
-------
DROPLET EVAPORATION
A fine liquid spray is injected into the humidification chamber. The
sole function of this spray is to cool the gas stream from its inlet tempera-
ture of 1100 F to a temperature of 320 F. The temperature gradient in the
humidifier is created by the transfer of the gas stream's sensible heat to
the water droplets of the spray.
The major difference between the humidification chamber and conven-
tional spray towers is that the difference between the gas and droplet tempera-
ture in conventional equipment is generally small, on the order of 100°F. In
the case of the humidification chamber, the difference between droplet and gas
temperature is on the order of ~1000 F.
An expression for droplet evaporation at high mass flux where heat con-
duction to the droplet is considered rate limiting is readily derived from first
law and continuity balances between the droplet phase and the gas phase.
Equation (1) shows the time rate of change of droplet diameter.
"('•% (¦- (t)*)» -
w"here k is the gas conductivity, X is the latent heat of vaporization,
p^is the liquid density, Tq is the inlet gas temperature, T^ is the inlet liquid
temperature, is the gas specific heat and Mj_, and Mgare the inlet liquid and
gas mass flow rates. The results of numerically integrating the above expres-
sion are shown in Figure 2. It is apparent from Figure 2 that if the injected
droplets are small.sharp temperature gradients will be formed. As we will show,
this has an important bearing on tar condensation.
If we assume that the diffusion of steam away from the droplet is rate limiting,
then droplet evaporation is described by Penetration Theory . Using Penetra-
tion Theory, the heat and mass transfer co-efficients are calculated using the
Chilton-Colburn analogy. These may be combined with a first law and con-
tinuity balance along with the properties of water in an algorithm which gives
droplet radius as a function of time. The penetration theory results we show
in Figures 3 and 4. Comparing the penetration theory results with the heat
transfer limited case leads to the conclusion that penetration theory is appli-
cable in the lower region of the humidification chamber and then only when
saturation conditions are approached.
TAR CONDENSATION PHENOMENA
While high boiling coal tar is obviously not a pure substance, a pre-
liminary treatment of tar droplet formation demands this type of simplifica-
tion. The physical tar properties used in the analysis are shown in Table 1.
148
-------
I I 111 III
GAS INLET 1100 °F
END OF EVAP. TIME 3
rrrn
I I I 1 Mil
000
20 30 40 60 80 100
DROPLET DIAMETER am
Fig. 2 Effect of Droplet Injection Diameter on Evaporation
Time - Heat Conduction Limiting
149
-------
1200
1100
1000
900
800
700
600
500
400
300
200
100
0
PA*TI<: LE : XLA.1V ETER = 0
5
3 Effect
10 15 20 25 30 35 40 45
ELAPSED TIME SEC (-FT FROM TOP OF TANK)
of Initial Particle Diameter on Temperature Profile - Penetration Theory
-------
Cn
g
zi.
i
a
«
H
w
2
5
Q
E-«
W
-1
6
O
P=J
P
Pi
W
H
<
£
6.0
5. 0
4.0
3. 0
2. 0
1. 0
10 15 20 25 30 35
ELAPSED TIME - SEC (~FT FROM TOP OF TANK)
40
45
50
Fig. 4 Effect of Evaporation on Drop Diameter -
Penetration Theory
-------
The most significant of these properties is surface tension.
TABLE 1
(7)
PROPERTIES OF HIGH BOILING TAR
Boiling Point Range
Temp °C
250
250-360
360-450
450
Average Molecular Weight
Average Specific Gravity-
Latent Heat of Vaporization
Surface Tension
Tar Ffaction %
20
15
15
50
250
1.1
75 cal/gm
36 dyne/cm (Temp Unknown)
In the absence of data on the surface tension at higher temperatures, this
property has been treated parametrically. Tar condensation is described by
two processes: nucleation and coalescence. In both of these processes, dust
particles may play a significant role. It is recognized that tar emulsified
in the recycle liquor may play an important role in the condensation process.
In this analysis, tar injected with the recycle liquor has been ignored.
TAR NUCLEATION
While the mass of dust particles in the inlet of the humidification cham-
ber may be significant, the number density of these particles may be quite
low. When the number density of dust particles is large, heterogeneous
nucleation may be dominant. If the number density of particles is small,
then homogeneous nucleation is the more important phenomenon. In our pre-
liminary treatment of homogeneous nucleation, we have examined initial rates.
A more accurate analysis requires the treatment of homogeneous nucleation as
a coupled phenomenon since the nucleation of liquid tar droplets will obviously
influence the concentration of tar vapor. The homogeneous nucleation process
requires the formation of critical radius (r*) particles. The critical radius,
concentration of particles of critical radius and rate of formation of critical
radius particles is determined by the gas stream temperature. In our analysis
152
-------
we have used the Clausius-Clapeyron form of the Thomson equation for
critical radius:
ZtfVp.
r* = (2)
X (1 - T/T )
o
where a is the surface tension, v» is the volume per molecule of tar, \
is the latent heat of vaporizationt T is the local temperature and T0 is the
tar saturation temperature. The concentration of particles of critical radius
(particles/cm3) is given by:
Ng* = N exp ^ j (3)
where N is the total number of tar molecules per cubic centimeter in th^ inlet
gas. The formation rate of critical radius particles (particles/sec-cm ) is:
Q * Ng» (4,
If KT/>/2m5cT
where P is the partial pressure of tar and m is the mass of-a tar molecule.
While the errors associated with calculated homogeneous nucleation rates
are sometimes large, they may be taken as the lower bound on the actual
nucleation rate. Classical theory predicts onset of nucleation accurately.
An examination of Eqs (2) - (4) shows why the surface tension of the
tar particle is so important. Fig. 5 shows the critical radius of tar parti-
cles as a function of gas stream temperature. It is apparent that a sharp
decrease in temperature in the humidification chamber represents a high
degree of subcooling of the tar vapor. Under these conditions, the criti-
cal bize becomes extremely small. Further, the critical particle radius
is proportional to the surface tension. Fig. 6 shows that the concentration
of particles is an exponential function of the gas temperature and Fig. 7
shows the impact of gas temperature on the rate of formation of critical
size tar droplets.
In the condition where the gas temperature in the humidification
chamber is dropped sharply (fine water droplets), a very large number of
very fine tar particles is formed very rapidly.
153
-------
10
w
£
O
&
u
W
D
i—i
P
C
ri
H
W
J
ft
O
pet
Q
J
<
U
hH
13
rt
u
1. o
o. 1
01
. 001
—
1
—
—
a = 36
DYNE/CM-a
CT= 18 D"J
'NE/CM -
a =
320°F
I
36 DYNE/C
A
T
SAT
y
400
500
600
700
TEMPERATURE - K
Fig. 5' Homogeneous Nucleation Effect of Temperature on
Critical Drop Radius
154
-------
10
10
10
10
cn
2
O
W
H
9
8
a
Q
¦!(¦
60
z
10
10
10
10
0= 3*
450 500
550
a =
0=3,
z
600 650
700
GAS TEMPERATURE - K
Fig. 6 Effect of Surface Tension and Temperature on Concentration of
Critical Size Droplets
155
-------
AGGLOMERATION
After nucleation, the tar particles grow by coalescence. This process
is described in general by the theory of Smoluchowski , For particle
coalescence in the Knuds^en regime (Knudsen No. >10) a simplified treatment
has been given by Ulrich . Coalescence in the Knudsen regime is somewhat
faster than coalescence in the continuum regime (Knudsen No. < 5). Tar
particles in the humidification chamber may be in either the Knudsen or the
continuum region, depending on the assumed value for surface tension. As
the particles coalesce, the process of coalescence changes from Knudsen dif-
fusion to continuum diffusion. The Knudsen number (Kj^) is defined as the ratio
of the mean free path (L) in the gas phase to the diameter of the particle.
Kn = L/2r (5)
L = (KT/^nd2) (6)
wTiere K is the Boltzmann constant, d is the diameter of gas molecules, the
gas P is the total pressure in atm and T is the absolute gas temperature.
For a total pressure of 100 PSIA (6.9 atm), a temperature of 320 F (433°K)
and assuming the diameter of gas molecules is 4 x 10"® cm^, the mean free
path is 0.05 jj,. In this case, the particle radius can grow to .014jj. Further
growth is via continuum diffusion. The coalescence of tar particles in the
region of Knudsen numbers greater than 10 takes the form:
N = N (1 + 3.4 x 10"12T 1/2C N 5/6t)"6/5 (7)
o o o
where N is the number of tar droplets per cm^ at time t, N© is the initial
droplet concentration, T is temperature, and CQ is the initial tar molecule
concentration in the gas phase. The coefficient 3.4 x 10"12 is obtained from
substituting tar properties into the coalescence equations. Note that at longer
times the number density of tar droplets is independent of the initial droplet
density.
With an expression for the number of tar particles, a simple mass
balance shows that the radius of the exit tar particles is a function of the
-2/5 power of the coalescence time.
In the continuum diffusion regime, the expression for the rate of particle
coalescence is derived from Smoluchowski theory using the Stokes-Einstein
form of the diffusivity . In this case, the number of particles is related to the
coalescence time by the equation.
N
N - o
_ (8)
1 + N
o
(i f')"
156
-------
—
— \
—
—
-t>" cr
* 36:
3YNE
'CM
a s 1J
DYN
/
2/CIU
<7= 3. 1
i DYN
CM
Z!~
—
—
\
.
450 500 550 600 650 700
GAS TEMPERATURE - °K
Fig. 7 Effect of Surface Tension and Temperature on Formation Rate
of Critical Size Droplets
157
-------
1. 0
a
n
I
CO
E>
HH
9
&
w
u
t-H
E-«
rt
2
18 DYNE/C
K <5
K> 10
Transit
Regiox
< DROPLET SIZE @ t = 0
001
• 4
-3
-1
Log (t - sec) COALESCENCE TIME
Fig. 8 Tar Droplet Growth Rate
158
-------
where ja is the viscosity of the gas.
The particle radius (rt) at some time after the onset of coalescence is
determined by the number density (N) at that time, and the particle radius
and number density at the initiation of coalescence by a simple mass balance.
r. = r (N/N) 1/3 .
to o
Fig. 8 shows the rate of particle growth assuming the nucleation of
tar droplets for a surface tension of ~18 dyne/cm. The tar droplet grows from
an initial radius of .0023^ to a radius of .0143(1 in about 2 msec. At this point,
the transition from the Knudsen diffusion to continuum diffusion occurs. At
about .02 sec., the particle has grown to a radius of .023^. Growth in the con-
tinuum diffusion region progresses slowly. At 100 sec., the particle diameter
has only grown to —1^.
HETEROGENEOUS CONDENSATION
Because of tar/dust interactions, heterogeneous condensation is a some-
what more complex phenomenon than homogeneous nucleation. We have
made a preliminary estimate of the condensation rate to form tar particles ot
a readily collectable size (~t0p). To do this, we assume the dust loading in the
gas is ~10 gm dust/gm air. We further assume that the dust has the com-
position of coal ash Mgm/cm ). For dust particles much larger than the mean
free path, the rate controlling process is the diffusion of tar molecules to the
surface of the dust particles. An estimate of the tar diffusion coefficient was
made using the Chapman-Enskog equation (D = 2.7 x 10-2 cm^/sec). Finally,
the process was analyzed at the tar saturation temperature («450°C). The con-
centration of tar particles in the gas stream is then:
CTAR = °TAR, O exP<-2r'D dDUST°DUST '
where CTAR,0 is «3.6 x 10*® molecules / cm^, D is the diffusion coefficient,
dpUST is dust particle diameter, and t is in sec. Substituting coefficient
values into the above expression we obtain (setting d^ST *10y):
C. s 3.6 x 10 ^ exp (- .32t).
TAR
The results of this equation are shown in Fig. 9. To reduce the tar vapor
concentration to 4% of the inlet value, a condensation time of *10 sec is required.
Note that increasing the dust loading will reduce the time scale exponentially.
While 10 sec is not an especially long time, the gas must not be appreciably
subcooled during this period. If the gas is subcooled during the process, then
159
-------
100
90
80
70
60
50
40
30
20
10
0
LOADING
T
72
o*
fO
dD
UST = 10
•ADING = 10"3 gm DUST/gm GAS—i
= 2. 7 x 1 -" cm /sec
IjL
dt
ne
UST
19
20 Pi
«LRTI
CLE,
S/cm
J
\
\
1 2 3 4 5 6 7 8 9 10 11 12 13 14 1
ELAPSED TIME AFTER INITIATION OF CONDENSATION
(SEC)
Fig. 9 Heterogeneous Condensation
160
-------
homogeneous nucleation will proceed concurrently with heterogeneous condensa-
tion. This will, in turn, reduce the amount of tar collected on the dust parti-
cles .
ACKNOWLEDGEMENT
We are grateful to the U.S. Department of Energy-, Morgantown Energy
Technology Center, for their support of this work. Funding was through DOE
Contract No. DE-AT21-78MCO8450.
REFERENCES
1. Cleaning Producer Gas from MERC Gasifier, A. Moore, IR No. 96
May, 1977.
2. Wet Scrubber System Study, Volume 1, ATP Inc. to Environmental
Protection Agency, July 1972, NTIS PB-213 016.
3. Transport Phenomena, R. Bird, W. Steward, E. Lightfoot, John Wiley
& Sons, I960.
4. Kinetic Theory of Liquids, J. Frankel, Doren Publications Inc. 1974.
5. Stochastic Problems in Physics and Astronomy, S. Chandrasekhar,
Reviews of Modern Physics, Vol. 15, No. 1, Jan 1943.
6. Particle Formation in High Temperature Systems, G. Ulrich, WSCI
71-37.
7. High Temperature Tar, J. Weiler, Chemistry of Coal Utilization Sup-
plementary Volume, H. H. Lowry Ed., John Wiley & Sons, 1963.
161
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IMPROVING THE EFFICIENCY OF FREE-JET SCRUBBERS
By:
Donald A. Mitchell
Hydro-Sonic Systems
P.O. Box 4138, Lone Star, Texas 75668
ABSTRACT
A free-jet scrubber Is an innovative gas cleaner in which the primary liq-
uid/gas contacting takes place in the turbulent mixing zone emanating from a jet
nozzle. The first such devices, powered by supersonic steam or air ejectors,
were exceptionally effective cleaners but were criticized as being high-energy
users. A great reduction in the energy consumption of free-jet scrubbers has
been accomplished in the last four years, a feat which may seem in contradiction
to the "contacting power rule". An explanation of what is happening in these
devices must take into account the effectiveness with which energy and water are
utilized, thereby leading one to a somewhat circular definition of "effective
contacting power". The theoretical implication of this analysis is primarily to
remind us that there are many paths between two states and some are more energy
efficient than others. The practical value is that such analysis generates in-
formation which aids in choosing between several design options, a direct bene-
fit of the various free-jet configurations.
162
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IMPROVING THE EFFICIENCY OF FREE-JET SCRUBBERS
Several recognized methods for capturing particulate and noxious vapors
involve the generation of fine water droplets and frictional forces which aid
in vigorous liquid/gas contacting. One class of cleaners utilizing this con-
cept, the gas-atomized systems, provide the means for this "scrubbing" action
by pumping the polluted gases through a restriction where the increased flow
velocity atomizes water as it is sprayed across the path of the high-velocity
gas. The relative energy of motion between the water spray and the polluted
gas is responsible for the frictional forces which atomize the water drops and
induce pollutant capture. This energy is dissipated as heat due to the irrevers-
ibility of frictional processes. The presently accepted theory of particulate
capture says that the performance of all such cleaners is basically a function
of energy dissipated by turbulence during liquid/gas contacting, a relationship
called the "contacting power rule", and which may be written as follows:
Nt = f(Pt) ^ a PtY (1)
or
H = 1 - exp[-f(Pt)] (2)
Where Nt = Number of transfer units = -ln(l-n)
r) = Fractional efficiency
Pt = Total contacting power
a, 7 » Constants peculiar to an aerosol
The implication of this theory, which is strongly supported by experiment,
is that an equal energy input to two differently configured scrubbers should
give equal cleaning performance. Of course some devices may fall below optimum
performance due to irreversible energy losses whenever water atomization and
droplet distribution are less than ideal or when the relative momentum is par-
tially reduced due to a poor flow pattern. This paper describes an innovative
class of gas-atomized scrubbers in which the basic operating principle permits
one to choose between designs in which there is a wide range in energy and water
options, the best choice depending on the particulars of a given application.
The essential distinguishing feature of free-jet scrubbers is that liquid/
gas contacting takes place in the turbulent mixing zone created by compressible
fluid flow through a jet nozzle. There are several ways in which this may be
accomplished. Several arrangements will be described, followed by a comparison
of their performance characteristics.
163
-------
The first free-jet cleaner (Figure 1) was called the "Steam-Hydro". In
this device, a steam ejector pulls the process gas around the nozzle where wa-
ter is sprayed around the steam jet, resulting in violent shattering of the wa-
ter and ejection of high-velocity droplets through the polluted gas. Excep-
tionally effective capture is achieved with a minimum of water usage (about 0.35
kg/kg of gas). Additionally, the system can be operated with a compressed air
ejector with no loss in performance, a fact which demonstrates that steam con-
densation is not the primary cleaning mechanism as has been suggested by some.
However, it can be seen that only a portion of the kinetic energy of the steam
or air can be effectively utilized in the contacting process. This is because
the water does not penetrate into the supersonic jet core, but remains in the
mixing zone where the flow is subsonic and thus of lesser kinetic energy. The
result of this energy loss is apparent from the performance data to be described
later.
STEAM /AIR EJECTOR
WATER INJECTOR
QUENCH (OPTIONAL)
GAS
CYCLONE SEPARATION
¥
INLET
Figure I Free-jet scrubber with ejector drive.
The "Fan Coalescer" (Figure 2) is an entirely subsonic free-jet design in
which maximum benefit can be made of kinetic energy. In this arrangement, a
fan is used to pump the gas through a nozzle where it flows out in a diverging
cone as shown. The inner cone retains its velocity and serves as the energy
source for atomization and pollutant capture. As with all free-jet configura-
tions, however, the contacting takes place in the turbulent mixing zone, although
some water may penetrate into the core. However, our experience has been that
any attempt to inject water into the core or into the nozzle throat will result
in somewhat poorer performance.
164
-------
SUBSONIC NOZZLE
COALESCER
: • • • • Tr-*' " " i
i - m ' ' i* 1 '{ftiTr i" '
WATER INJECTOR
r
Figure 2 Single stage fan-driven "Coalescer" cleaner.
For control of fine particulate and noxious vapors, the "Tandem Nozzle"
arrangement (Figure 3) is normally significantly better than the single nozzle
design. Dividing the contacting process into two parts allows better utiliza-
tion of energy and water, primarily due to enhanced saturation, adiabatic ex-
pansion, condensation, and droplet growth. The two-stage free-jet shows its
greatest advantage for processes in which emissions consist of very fine, low
density particles, especially where oily hydrocarbons are known to be present.
SCRUBBING AND INITIATION
"OF GROWTH
GAS
INLET
SEPARATOR
SUBSONIC NOZZLE
SUBSONIC NOZZLE
Figure 3 Two stage subsonic free-jet arrangement.
165
-------
The "SuperSub" (Figure 4) is a hybrid which incorporates a single subsonic
nozzle and a small steam or air ejector. A combination of energy sources allows
one to achieve efficient performance under normal conditions and to enhance fine
particulate capture during upset conditions simply by activating the small ejec-
tor. It requires only a 10% increase in water flow beyond that required by a
single subsonic nozzle design.
In comparing the performance of free-jet scrubbers, it is convenient to
plot efficiency or emissions versus equivalent pressure loss. The pressure drop
therefore indicates the power input per unit volumetric flow rate. The problem
arises as to how to compute an equivalent pressure loss for the ejector drive.
Actually it is a simple procedure, although the values assigned to gas densities
and efficiencies are somewhat arbitrary. However, the general features are not
very sensitive to these estimates. The equivalent pressure loss for an ejector
unit is obtained by finding the available energy loss for the ejector flow and
computing the pressure drop across a fan system necessary to drive the same pro-
cess gas while expending the same total available energy. Therefore, for the
steam or air ejector cleaners, the performance curve is not really a plot of
performance as a function of effective contacting power since one can only accu-
rately compute the kinetic energy of the ejector media, not the energy effective-
ly utilized in scrubbing. This is because of the unknown amount of energy lost
due to irreversibilities in the transition from supersonic to subsonic flow.
However, if one simply estimates the effective energy at sonic conditions, a
value comparable to that for the fan-driven cleaners will be obtained.
WATER
WATER INJECTED
AGGLOMERATION
INLET
SMALL EJECTOR NOZZLE
(SUPERSONIC)
SUBSONIC NOZZLE
FREE-JET MIXING PARTICULATE WETTED
STEAM OR COMPRESSED AIR
Figure 4 Hybrid supersonic/subsonic free-jet arrangement
166
-------
The open hearth steel refining process is one in which emissions are fairly
repeatable (Figure 5). While the absolute value of controlled emissions is
noteworthy, the more instructive feature is the great difference in cleaning
between the units at typical energy levels. As expected, the Steam-Hydro con-
sumes much more energy than the fan-driven units. However, it uses the least
quantity of water (about 3 gal/1000 SCF). The other cleaners also achieve high
cleaning, but do so with considerably less available energy expenditure, although
they do require three or more times the water rate. Note that the stack sample
measurement for the single nozzle system suggests that its performance is the
same as that for the SuperSub, a somewhat misleading result. Although no stack
samples were obtained for the single stage system at higher pressure losses, it
was obvious from the stack opacity that the higher performance could not be
achieved even at much greater energy levels and water rates. This is partly due
to the fact that the tandem unit induces condensation growth of submicronic fumes
and partly due to the fact that a single stage cannot effectively utilize an un-
limited quantity of water—that is, excessive water will tend to destroy the
free-jet action.
7
6
J*
~
~
~
~
5
B
~
4
A
~
A SINGLE NOZZLE
~ TANDEM NOZZLE
~ SUPERSUB
# STEAM HYDRO
3
2
100
200
300
50
400 500
Pressure toss or equivalent (cm W.G.)
Figure 5 Performance of free-jet scrubbers on an open hearth furnace.
167
-------
A second series of curves (Figure 6) illustrates the dramatic difference
between performance on a process with more limited particle emission types and
sizes, such as steel refining, and a process with a wide range of emissions,
such as for a gray iron cupola. The most obvious feature for this process is
that the devices requiring less energy can give remarkably better performance
than the devices needing higher energy, with the exception of the single nozzle
scrubber. (Data for the single stage unit appear as discrete points to illus-
trate an apparent non-conformity with simple theory.) What this data set sug-
gests is that water utilization and condensation growth are dominating factors
when one is scrubbing an off-gas with a large fraction of submicronic, low den-
sity emissions, especially where condensible hydrocarbons are present.
50
~
~
~ A
~
~
~
~
~
~
A
~
A
A
100
1 T
A SINGLE NOZZLE
~ TANDEM NOZZLE
~ SUPERSUB
# AIR HYDRO
200
300
400 500
Pressure loss or equivalent (cm W. G.)
Figure 6 Performance of free-jet scrubbers on a gray iron cupola.
168
-------
The selection of a particular free-jet configuration is greatly aided by
utilization of performance data as given in Figure 7. These curves indicate
something about the nature of the particulate as well as the relative advan-
tage of one particular type cleaner over another. However, such curves do not
indicate the influence of certain process variables and do not alleviate the
need to predict absolute emission levels. Furthermore, it is quite hazardous
to extrapolate efficiency from very low to high energy levels necessary for
opacity control unless a given device has been shown to have the capability to
achieve high performance. It is the ability of a cleaner to get the last bit
of fine particulate which really counts in opacity-limited processes, and this
is where the free-jet gas cleaners are unsurpassed.
7
(TANDEM
(SINGLE STAGE)
6
COAL DRYER
(10.6)
(Inlet ^"*
10.6 gm/m*N)
5
OPEN HEARTH
4
ELECT. ARC
FURNACE
(.74)
IRON CUPOLA
3
MUNICIPAL INCINERATOR
IRON CUPOLA
2
80
70
90
60
100
110
130 140 150
120
Pressure loss (cmW.G.)
Figure 7 Performance curves for selected processes.
169
-------
To summarize the characteristics of free-jet scrubbers which account for
their operation at sustained optimum performance? the following features are
fundamental:
1. The nozzles are designed to achieve maximum benefit of relative momentum.
2. The free-jet produces a turbulent mixing cone which extends the length
of the liquid/gas contacting region and increases the time of inter-
action.
3. The water is sprayed into the turbulent mixing region where pressure
gradients enable it to be used to maximum benefit, this effect being
most pronounced for the ejector-driven units.
4. Beyond the high-velocity core, water droplets are agglomerated, thus
aiding in liquid separation and minimizing mist carry-over.
In conclusion, the application of accepted theory to free-jet scrubbers
gives some insight into the performance characteristics of the free-jet scrub-
bing process. However, it adds little to the theory itself, except to remind
us that there are many energy paths and physical mechanisms possible between the
inlet and outlet of a gas cleaner. Also, it is humbling to remind ourselves of
how little available energy could be used to do the same job by the most effi-
cient process, a value which can be easily calculated using the methods of sta-
tistical mechanics. In principle, on the order of about 10 joule could "clean"
a cubic meter of typical off-gas. While such a low value may never be approached
in practice, it should serve to point out the biting truth of just how far we all
have to go to achieve the ideal.
170
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HIGH VELOCITY FIBROUS FILTRATION
By:
Michael J. Ellenbecker, John M. Price, David Leith, and Melvin W. First
Harvard School of Public Health
Department of Environmental Health Sciences
665 Huntington Avenue
Boston, Massachusetts 02115
ABSTRACT
The particle collection characteristics of fibrous beds at low filtration
velocities are well understood; many experiments have confirmed the collection
efficiencies predicted by theory. In this study, the particle collection
characteristics of a stainless steel fibrous material collecting resuspended
fly ash were investigated at very high superficial filtration velocities
(i.e., 1-10 m/s). Under these conditions the theoretical collection effi-
ciency due to impaction should have approached 100% for all but the smallest
particles present. Previous investigators have found, however, that penetra-
tion in this velocity range can be higher than predicted by theory, presumably
due to particle bounce and/or resuspension.
Collection efficiency was measured against particle size and filtration
velocity. Test results indicate that, under certain conditions, penetration
can increase both with velocity and with particle size. The possible use of
these anomalous collection characteristics in a unique fly ash collection
device is discussed.
171
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HIGH VELOCITY FIBROUS FILTRATION
INTRODUCTION
Mats of fibrous materials have been used for many years to collect par-
ticles from gas streams. Several different collection mechanisms, including
diffusion, impaction, and interception, can contribute to the overall collec-
tion efficiency of fibrous mats. These collection mechanisms are well under-
stood, and the theories describing them can be used to accurately predict col-
lection efficiency over a wide range of filter and operating characteristics.
The theory and experiments described here are concerned with a set of
filtration conditions which are outside the range of "normal" filtration.
The predominant characteristic of this filtration regime is the very high
filtration velocities used - in the range of 1-10 m/s (200-2,000 fpm). These
filtration velocities, when combined with the tested filter materials (ran-
domly oriented 8 urn stainless steel fibers) and dust (resuspended fly ash),
result in particle collection characteristics which cannot be described satis-
factorily by classical filtration theory.
This combination of particle sizes, fiber diameter, and filtration velo-
cities results in particle collection predominantly by impaction, even for
very small particles. Just as importantly, the kinetic energy of large par-
ticles and the drag forces on collected agglomerates are sufficiently large
so that particle bounce and reentrainment can be expected.
This research was undertaken to investigate fly ash collection charac-
teristics in the high inertia regime. In particular, the research is aimed
at determining combinations of fiber beds and filtration conditions where
small particles are effectively collected by impaction while large particles
bounce through or are reentrained. Since the large particles which would
penetrate such a filter can be collected by other simple devices (e.g., cy-
clones), this penetration characteristic could possibly form the basis for a
two-stage, high efficiency collection device.
THEORY
Particle Collection Mechanisms
Introduction. Filtration theory attempts to predict overall particle re-
moval in a filter based on an understanding of the interaction of particles
with a single filter element. This element can be an isolated fiber or pre-
viously collected dust particles.
The single element efficiency, ns» defined as the number of particles
collected by a filter element divided by the number of particles whose centers
172
-------
pass through the element's projected area as the particles approach it.2 The
major mechanisms by which particles collect in a filter element are inertial
impaction, diffusion, interception, gravity settling, and electrostatic de-
position.
A single element collection efficiency due to any one mechanism, rii»
can be estimated for each of the fundamental collection mechanisms; however,
this paper is concerned with sets of operating conditions where the impaction
mechanism dominates all the others. Under such conditions, the contribution
of all other collection mechanisms is insignificant and can be disregarded.
Inertial Impaction. As the gas flows past a collector element, the
streamlines curve around the collector. Near the collector surface, inertia
will cause particles to deviate from the streamlines. Particles with suffi-
cient inertia can cross streamlines and hit the collector. Inertial impac-
tion has been found to be proportional to the dimensionless impaction param-
eter,2 ip:
0 d v C
~ " -?^-C (1>
where: Pp » particle denisty (kg/m3)
dp - particle diameter (m)
v = superficial filtration velocity (m/s)
Cc ¦ Cunningham slip correction factor
U = gas viscosity (Pa's)
d^ = fiber diameter (m)
Much experimental and theoretical work has been devoted to determining
the relationship between the impaction parameter, ip, and the single fiber col-
lection efficiency from impaction, rij. One widely accepted empirical formula
was derived from experiment and theory by Landhal and Herrman.^ It was de-
veloped for a fiber Reynold's Number of 10 and can be expressed as:2
n - -3 \ (2)
ij» + 0.774 + 0.22
Typical Calculations. Theoretical impaction collection efficiencies can
be calculated for values of the variables used in this study, which are sum-
marized in Table 1. All experiments were conducted using 8 ym diameter fi-
bers. Figure 1 plots single fiber collection efficiency nj as a function of
particle size for the lowest velocity (1 m/s) and the highest velocity (10 m/s)
used here. Note that a significant fraction of even very small particles will
impact a fiber at these velocities.
Other Collection Mechanisms. Since the fibers used in this study were
very small (8 ym diameter), interception might be thought to be important.
However, velocities were sufficiently high so that impaction always completely
overshadowed interception. The theoretical single fiber collection efficiency
from impaction and interception combined was calculated assuming that the
mechanisms are independent but that a particle caught by one mechanism cannot
173
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be caught by the other; the combined collection efficiency never exceeded
the single fiber collection efficiency from impaction alone by more than 1%.
Likewise, particle collection from diffusion was negligible under the condi-
tions considered.
Overall Collection Efficiency
5
Dorman gives t
L and solidity a as:
5
Dorman gives the overall collection efficiency for a filter of thickness
-4Lans
n ¦ 1 " ^ndjU-a)1 (3)
where ns is the combined single fiber collection efficiency from all effects.
In this study, impaction dominates all other collection mechanisms so that
n y It (4)
s i
Equation 3 is difficult to apply in this case, because at the high velocities
used here the filter mats compress and L changes in a manner which is diffi-
cult to measure and/or predict. However, Equation 3 can be modified by making
use of the definition of solidity, a, defined as the ratio of fiber volume to
total filter volume:
li
a ~ VT ~ LA
(5)
m
<6>
where pf is the filter density and m, and A are the filter mass and cross-sec-
tional area, respectively. Solving Equation 6 for filter thickness L and sub-
stituting into Equation 3:
-4m n
° 1 - E^[,pfdfA(l-a)] <7>
The overall theoretical filter collection efficiency can thus be calcu-
lated from easily measured filter bed parameters (mf, Pf, dj, A), the theore-
tical single fiber collection efficiency from impaction (nj), and filter poro-
sity (1-a). Porosity changes with filtration velocity; however, the porosity
of a filter with no airflow can be measured using pycnometry, and changes in
porosity with filtration velocity can be estimated from the slope of the meas-
ured pressure drop versus velocity curve.
174
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EXPERIMENTAL APPARATUS
Apparatus
Figure 2 is a diagram of the experimental high velocity filtration sys-
tem. It contains a 30 cm square filter holder by which fibrous materials can
be packed to depths of up to 15 cm. A movable 4 mesh screen upstream and a
stationary 4 mesh screen downstream provide rigid support at high filtration
velocities. A downstream centrifugal blower draws the test aerosol through
the 20 cm round entry duct, the filter holder, and the Venturi meter which
measures the airflow rate. A Dwyer photohelic gauge is used with a motor
driven blastgate to maintain a constant airflow rate.
A National Bureau of Standards (NBS) dust feeder** generates the fly ash
aerosol at mass rates from 3 to 125 g/min. The test dust is an electrosta-
tically precipitated fly ash with a count median diameter (CMD) of 0.14 ym
and a geometric standard deviation (GSD) of 1.9. A polydisperse (CMD =0.7
um, GSD = 2.05) dioctyl phthalate (DOP) aerosol can also be generated using a
DeVilbiss nebulizer (The DeVilbiss Co., Somerset, PA). The test aerosol and
inlet air are thoroughly mixed before entering the filled section by a Stair-
mand disk7 located at the inlet duct.
Simultaneous up and downstream isokinetic samples are taken to determine
aerosol penetration of the filter medium. Microscopic and gravimetric analy-
ses of the sample filters determine particle size distribution of the fly ash
and downstream of the filtering section as well as overall collection effi-
ciency. Eight stage Anderson impactors (Anderson-2000 Inc., Atlanta, Georgia)
are also used up and downstream for particle sizing. DOP penetration is meas-
ured using a Particle Measuring Systems, Inc., laser scattering aerosol spec-
trometer (PMS Model CSASP-100).
Procedures
Upon packing a filter material into the filter housing and setting the
airflow rate, test aerosol is generated and directed to the filter. Inlet
dust concentration is held constant and after the filter reaches maximum pres-
sure drop (about 7.5 kPa) the filter is removed for inspection and cleaning.
During the dust loading, pressure drop across the filter is recorded and aero-
sol samples are taken at regular intervals. Cleaning of the filter material
can be done by blowing compressed air through the upstream side of the filter,
striking the filter and filter frame against a rigid surface, or shaking the
filter manually. After cleaning, the filter is placed in the filter housing
and retested at original flow rate and dust loading.
EXPERIMENTS
The fractional penetration of fly ash through filter mats of 8 vim stain-
less steel fibers was measured at five filtration velocities: 1.5, 2.9, 3.2,
6.9 and 8.0 m/s. The characteristics of the fly ash and filter material are
given in Table 1. At the lower velocities the filtration time between clean-
ings was long enough so that adequate Anderson lmpactor samples could be col-
lected in one cleaning cycle, while at the higher velocities multiple cycles
175
-------
were needed to obtain sufficient impactor deposits.
The fractional penetration of the polydisperse DOP aerosol was also
measured over the same velocity range. The DOP concentration was extremely
low, so that the filter exhibited no loading effects during the DOP tests.
RESULTS
Fractional penetration curves for fly ash collected at five different
filtration velocities are presented in Figure 3. Increased penetration with
both velocity and particle size is evident.
Similar measurements were made with the liquid DOP aerosol at six velo-
cities; the fractional penetration curves for these experiments are shown in
Figure 4. Shown also for comparison in Figure 4 are the 1.5 and 8.0 m/s fly
ash penetration curves from Figure 3. In contrast to the fly ash, the DOP
aerosol penetration follows the trends predicted by classical impaction
theory, decreasing with both velocity and particle size.
Dust concentration downstream of filters loaded with fly ash using clean
inlet air were also measured and found to be Insignificant at these velocities.
DISCUSSION
Particle Bounce
Equation 7 is a theoretical formula which predicts overall filter collec-
tion efficiency (rj) from the single fiber collection efficiency (rij) and gross
filter parameters (m^, p*, df, A, 1-a). However, Equation 7 can also be used
to calculate experimental single fiber collection efficiencies (n'j) from ex-
perimental filter collection efficiency data (nf). Such numbers are more use-
ful than overall collection efficiencies in comparing results to theory be-
cause they are independent of the particular filter configuration tested and
should only be a function of the particle-fiber impaction parameter (Equation
2).
Solving Equation 7 for n'j,
¦nrpfdf A(l-a) ln(l-n')
and substituting the filter parameters from Table 1 gives:
n'j - -0.145(l-a)ln(l-n') (9)
where the solidity, a, is a function of filtration velocity. As discussed
earlier, changes in a with velocity can be estimated from the clean filter
pressure drop curve; in any case, the quantity l-o is always close to unity
in these tests so that changes in a have little effect on the calculation of
176
-------
The empirical formula of Landall and Herrmann relating impaction param-
eter (4) and rij (Equation 2) is plotted in Figure 5. Also plotted are the
experimentally determined single fiber collection efficiencies for the lower
test velocity (1.5 m/s, a = 0.0085) and the highest test velocity (8.0 m/s,
a - 0.0285). The values for intermediate velocities are similar and have been
omitted for simplicity. As shown in Figure 5, the experimental and theoreti-
cal single fiber collection efficiencies are in agreement only for the smal-
lest Anderson impactor fly ash size category, when the fiber impaction param-
eter is approximately equal to 1. As impaction parameter increases the ex-
perimental n'x decreases; when rij equals 0.99 (i|i i 100), n'j equals only 0.2
for the lowest velocity and 0.05 for the highest velocity.
Experimental single fiber collection efficiency data for the liquid DOP
aerosol are also shown in Figure 5. In contrast to the fly ash data, the DOP
values closely fit the Landell and Herrmann curve. These data strongly
suggest that fly ash particles do impact the fibers but bounce through the
filter at high velocities. Indeed, for particles larger than 4 ym the data
indicate that approximately 80% of the particle-fiber collisions result in
particle bounce at 1.5 m/s and that 95% result in particle bounce at 8 m/s.
Adhesion Probability
The amount of particle bounce for different velocities and particle sizes
is indicated by the particle adhesion parameter, h, which is defined as the
probability of a particle adhering to a fiber after colliding with it.
Theoretically, the probability of particle adhesion can be calculated from a
consideration of the energies involved. A particle will not adhere to a
fiber upon collision if the kinetic energy of the particle on rebound is
greater than the energy of attraction (primarily due to London van der Waals
forces).^ Loffler® and and Dahneke^ present the energy balance equation and
the procedures for calculating critical conditions necessary for bounce to
begin. As pointed out by Dahneke® and Esmen, et al.,*® the equations are
difficult to solve for a polydisperse, irregular dust such as fly ash and a
multilayer fiber filter, so that adhesion probability cannot be predicted as
a function of particle size and velocity. However, experimentally measured
adhesion probability should show a correlation with particle kinetic energy.
The adhesion probability can be obtained from experimental data by taking
the ratio of the experimentally measured single fiber collection efficiency
to the theoretical single fiber collection efficiency. Since impaction pre-
dominates in this study,
The adhesion probabilities calculated for all of the data points in Fig-
ure 3 are plotted against particle kinetic energy in Figure 6. The expected
decrease in adhesion probability with Increasing kinetic energy is evident.
The data fall on a straight line (log-log plot) at low energies (high adhesion
probability) but exhibit increasing scatter at higher energies (low adhesion
probability).
177
-------
The spread in data at high energies may be caused by the complex par-
ticle impaction and rebound pattern which would exist in this regime. High
energy particles, after bouncing, will accelerate until reaching gas velocity;
during acceleration they may impact another fiber and rebound again or they
may adhere, depending on the velocity at impact. A very complex collection
pattern can thus be expected for a polydisperse dust in a multilayer filter
for particles with large kinetic energy. However, this mechanism cannot
account for the very low~adhesion probabilities measured on two experiments
(v » 5.2 m/s and 8.0 m/s) when the particlle kinetic energy exceeded 10-13 j.
Additional data will be collected in an attempt to clarify these results.
Relative humidity can have a great effect on particle adhesion,1^" and
thus could have contributed to the differences in adhesion probabilities meas-
ured in different experiments. Relative humidity ranged from 20-50% during
the experiments, but the adhesion data could not be correlated with the rela-
tive humidity data.
A regression line was fit to the data of Figure 6 and is plotted in Fig-
ure 7. Also plotted are data of Loffler® and Esmen, et al.*® Loffler's data
were collected for 1.8-10.0 pm diameter NaCl particles collected on 20 vim
diameter polyamide fiber filters while Esmen, et al., impacted 4.4 and 8.8 ym
diameter uranine spheres on a flat brass substrate. Considering the different
particles and substrates used, the agreement among the various experiments is
good. Figure 7 indicates a region of particle inertial energy where particle
adhesion can be expected to be very good (E^ < 10 J) and a region where par-
ticle adhesion can be expected to predominate (Ej^ > 10~^3j). In the inter-
mediate energy regime, (10~^J < Ejj < 10~^J), particle adhesion probability
will be a strong function of the particular particle and filter used.
Practical Design Implications
The penetration characteristics reported here suggest that a novel, com-
pact, high efficiency particle collector may be feasible. The device would
operate at high gas velocities, and would collect particles over a wide size
range by impaction. It would consist of two stages: 1) a fibrous media fil-
ter, which would efficiently collect small particles but allow some large par-
ticles to bounce through or reentrain; and 2) a cyclone, which can efficiently
collect large particles. When operated in series, the two devices should col-
lect all particle sizes with high efficiency.
Such a device would have many desirable properties. Our experiments in-
dicate that the pressure drop across a clean fibrous bed, when operated at
high velocities, (e.g., 1.5 m/s), is about the same as across a cyclone: 0.5-
1 kPa. The total pressure drop across the device would thus be fairly low -
on the order of 2 kPa, The device would be extremely simple in design and
easy to operate; the only area where moving parts may be needed is in the fil-
ter cleaning system, and that would depend on the cleaning method chosen. The
device, by operating at high gas velocities, would be very compact. It would
collect all types of aerosols (solid and liquid) and would use relatively small
amounts of energy.
178
-------
The design and testing of such a two-stage device is now in the planning
stage. An important design consideration now being addressed is the selection
of an optimum fibrous material. The ideal material would have a fractional
penetration similar to the stainless steel fiber beds but would be cheaper and
easier to clean in situ. Materials such as felts and sintered fiber mats are
being tested in preparation for the construction of a pilot two-stage device,
SUMMARY AND CONCLUSIONS
Particle bounce has been shown to play an important role in the collec-
tion of fly ash by mats of stainless steel fibers. The penetration of fly ash
particles through 8 ym fiber filters in the 1-10 m/s velocity range increased
with both particle size and velocity; since impaction was the controlling col-
lection mechanism, these results run contrary to classical filtration theory.
Particle adhesion probability was found to be inversely proportional to
particle kinetic energy. This result is in general agreement with previous
adhesion studies, which were performed with monodisperse aerosols under con-
trolled laboratory conditions. This study demonstrates that the particle ad-
hesion properties of a polydisperse dust (fly ash), under conditions similar
to those encountered by practical filters, can also be related to the particle
kinetic energy.
These results have important implications for the operation of fibrous
filters at high velocity, and may help explain cases where penetration is
higher than anticipated. These experiments also suggest the possible use of
a novel two-stage inertial collection device to take advantage of the high
velocity particle penetration characteristics noted here.
ACKNOWLEDGEMENT
This work was funded by the Environmental Protection Agency under grant
804-700, James H. Turner, project officer.
179
-------
NOMENCLATURE
A
2
filter cross-sectional area, m
Cc
Cunningham slip correction factor
df
fiber diameter, m
dp
particle diameter, m
h
adhesion probability
L
filter thickness, m
mf
filter mass, kg
V
superficial filtration velocity, m/s
Vf
3
total fiber volume, m
VT
3
filter volume, m
a
filter solidity
T1
theoretical filter collection efficiency
n'
experimental filter collection efficiency
ni
theoretical single fiber impaction efficiency
*'l
experimental single fiber impaction efficiency
combined single fiber collection efficiency
y
gas viscosity, Pa*s
pf
fiber density (kg/tP)
pP
particle density (kg/m^)
*
inertial impaction parameter
180
-------
REFERENCES
1. Davies, C.N, Air Filtration. New York, Academic Press, Inc., 1973.
2. Pich, J. Theory of Aerosol Filtration by Fibrous and Membrane Filters.
In: Aerosol Science, Davies, C.N. (ed.). New York, Academic Press,
Inc., 1966.
3. Landahl, H.D., and R.G. Herman. Sampling of Liquid Aerosols by Wires,
Cylinders, and Slides, and the Efficiency of Impaction of the Droplets.
J. Colloid Sci. 4(2):103-136, 1949.
4. Strauss, W. Industrial Gas Cleaning. 2nd ed. Oxford, Pergamon Press,
1975. p. 287.
5. Dorman, R.G. Filtration. In: Aerosol Science, Davies, C.N. (ed.). New
York, Academic Press, Inc., 1966.
6. Dill, R.S. A Test Method for Air Filters. Trans. Amer. Society of Heat-
ing and Ventilating Engin. 44:379-386, 1938.
7. Stalrmand, C.J. Sampling Gas-Borne Particles. Engineering. 41:141-143,
181-183, 1941.
8. Loffler, F. Adhesion Probability in Fibre Filters. Clean Air. 8(4),
1949.
9. Dahneke, B. The Capture of Aerosol Particles by Surfaces. J, Colloid
Interface Sci. 37(2):342-352, 1971.
10. Esmen, N.A., P. Ziegler, and R. Whitfield. The Adhesion of Particles Upon
Impaction. J. Aerosol Sci. 9:547-556, 1978.
11. Corn, M. The Adhesion of Solid Particles to Solid Surfaces, I. A Review.
J. Air Poll. Control Assoc. 11(11):523-528, 1961.
181
-------
Table 1. VALUES OF TEST VARIABLES USED IN THIS STUDY.
Dust
Type
Count median diameter
Geometric standard deviation
Particle density (Pp)
Filter
Fiber type
Fiber diameter (df)
Fiber density (p^)
Filter cross-sectional area(A)
Filter mass, mf
Filtration velocity (V)
resuspended fly ash
0.14 ym
1.9
2.2 x 103 kg/m3
310 stainless steel
8 vim
7.92 x 103 kg/m3
0.093 m2
0.032 kg
1 - 10 m/s
182
-------
S f-o
s
5: 0.6
10.0
0.5
/.O
PA/tr/CLE D/ANETEft, dp(fJLm)
Figure 1. Theoretical single fiber impaction efficiency vs.
particle diameter at 1 and 10 m/s filtration velocities.
-------
EXHAUST
PHOTOHELIC
GAUGE
^ MANOMETER
NBS
DUST
FEEDER
MOTOR
00
DAMPER
VENTURI
METER
FIBROUS
FILTER
MEDIA
Figure 2. High velocity filtration apparatus.
-------
CO
tn
fOO.O
\20.0
S
—A5 m/s
Q—£.S "
0—S.£ "
A—6.9 "
—e.o»
^ K/.0
6/? 6.0 JO.O
I I I
/5.o £ao
I I
A£fiODrMM/G D/AMEr£ft (^)
Figure 3. Fly ash penetration vs. particle size at five filtration velocities.
-------
jooo
00.O
so.o
40.0
20,0
/0.0
S.0
0.0
+0
*0
S A0
x **
0.6
I
^ 0.4
0.2
0./
1 TT
x
\
IS6EN0
A
O
~
A
~
V
A0 rrys P0P
2.0 "
3.0 "
5.2 "
s.s "
0,0 "
/.5 "
0.0 "
44 44
4$ 99
09 44
09 H
09 94
FIK4S#
I I 1
I
J 1 I I
.4 .0 .0/.0 2.0 4.0 S.0 0.0/0.0
4£fi00r#4M/C 0MAt£r£#
20.0
Figure 4. Dioctyl phthalate (DOP) penetration vs. particle
size at six filtration velocities.
186
-------
$
OJ
LAA/DAHl -tfEPPMA/V
CURVE
\
X
1
IE6END
¦ Fir ASH. /.srrys
• FLY ASH, S.O m/s
A DO P. All rsioctr/ss
/O ~l
/0O
/o
/MPACr/OAf PAPAMETEP, \ft
/02
JO*
Figure 5.
Theoretical and experimental single fiber impaction efficiencies
vs. impaction parameter for DOP and fly ash.
-------
/S m/s
/0~]
/0~15 /0-lh /0~lz /0~12
PA*r/ci£ xwsrtc smsask 4L U)
-i i
>-10
Figure 6. Fly ash adhesion probability vs. particle kinetic energy.
-------
AO
O.0
o.s
NT
\*0.£
Is
.OS
\
,0/
to
LESEMD
V ESMEN, 4.4
o iom.£*t/o.o pm
• • " SM m
O " * 2JS "
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1
PO*Efi Ct/WE
Ftr ro DATA,
was sroor
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ao
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AO~" A0~13 /O-11
PAAtTAClE K/NEr/C EMEfiS/, £ (J)
/o-
11
/0-10
Figure 7. Comparison of adhesion probability data from
this study with those of other investigators.
-------
THE EFFECT OF DUST RETENTION ON PRESSURE DROP IN A HIGH VELOCITY
PULSE-JET FABRIC FILTER
By:
Michael J. Ellenbecker and David Leith
Harvard School of Public Health
Department of Environmental Health Sciences
665 Huntington Avenue
Boston, Massachusetts 02115
ABSTRACT
Pulse-jet fabric filters exhibit abnormally high pressure drops when
operated at higher than normal superficial filtration velocities. The causes
of high pressure drop are investigated for a pilot-scale pulse-jet filter
collecting fly ash at 50, 75, 100, 125, and 150 mm/s filtration velocities.
The measured equilibrium pressure drops ranged from 0.35 kPa (1.4 in
H2O) at low velocity to 10.2 kPa (40.8 in H2O) at high velocity. Retention
of a large mass of dust on the fabric was found to have a significant effect
on equilibrium pressure drop at all of the tested filtration velocities.
The dust deposit's distribution, total mass, and specific resistance
(K2) all remained essentially constant at lower filtration velocities, but
changed dramatically at the highest test velocity. At the 150 mm/s velocity
the total dust mass doubled, the dust distribution equalized over the surface
of each bag, and K2 increased by a factor of 4.7.
190
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THE EFFECT OF DUST RETENTION ON PRESSURE DROP IN A HIGH VELOCITY
PULSE-JET FABRIC FILTER
INTRODUCTION
Fabric filters have been widely used for particle collection because
they operate with high efficiency, and will find increased future use as par-
ticle emission regulations become more stringent. Pulse-jet filters may be
well suited for both new and retrofit applications because they operate at
higher superficial filtration velocities (air-to-cloth ratios) than do re-
verse-air or shaker-cleaned systems, and are correspondingly smaller. It is
tempting to increase filtration velocity through a pulse-jet system still
further, since the decreased installation cost may more than compensate for
the increased cost associated with operating at a higher pressure drop. Oc-
casionally, space limitations may make a high velocity pulse-jet system the
only option which will both fit and provide the necessary collection efficien-
cy.
However, Increases in filtration velocity above the values commonly used
are associated with large increases in system pressure drop,1*2 which tend to
make high velocity operation both uneconomical and impractical. Fabric fil-
tration theories do not relate this increased pressure drop to system opera-
ting variables.
Redeposition of the dust deposit following a cleaning pulse has been
identified2 as one cause of increased pressure drop at high velocity. Rede-
position occurs when dust agglomerates removed from a fabric surface by a
cleaning pulse are recaptured on the same bag or a neighboring bag before they
can fall into the hopper. Another cause of increased pressure drop is the
failure to remove part of the dust deposit from the fabric during cleaning.
Unremoved and redeposited dust have the same effect on the system—excessive
dust on the fabric. In this paper they will not be distinguished, and their
total effect will be termed "mass retention." The mass retained on a fabric
after cleaning will be called "residual" mass.
Mass retention causes the dust deposit to bfcild up from cleaning cycle
to cleaning cycle until the system reaches an equilibrium condition in which
the total mass of dust on the bag is much greater than the mass of dust col-
lected during one cleaning cycle. The pressure drop across this large dust
deposit is much greater than the pressure drop across a deposit which consists
solely of "the dust collected during one cleaning cycle. Mass retention be-
comes very significant at high filtration velocities, presumably because of
the increased tendency to drag dust agglomerates back to the bag surface as
filtration resumes after cleaning.2 Therefore, it appears that redeposition-
caused dust deposit retention may be a key factor in preventing the economi-
cal use of high velocity pulse-jet filter systems.
191
-------
The research described here was undertaken to quantify the factors which
contribute to increased pressure drop when a pulse-jet fabric filter is op-
erated at high superficial filtration velocities. The relative contributions
of changes in filtration velocity, magnitude and distribution of the dust
mass, and dust deposit specific resistance were determined.
THEORY
Pressure Drop
Darcy's law, which is often applied to fabric filter systems,states that
the rate of flow through a porous medium is directly proportional to the pres-
sure gradient causing flow.^ In fabric filtration, the total pressure drop
Ap can be divided into the pressure drop across the clean fabric Apf and that
across the dust deposit Ap^:
in which Kj is the clean fabric resistance, is the dust deposit specific
resistance, v is superficial filtration velocity, and w is the areal density
of the dust deposit.
Most filtration models define Kj for a "conditioned" fabric rather than
a clean or unused fabric. is then the resistance of a fabric which has
been conditioned by dust which lodges within the fabric itself and is not re-
moved by cleaning, while Ko characterizes the resistance of the actual dust
cake which forms on the fabric. In high velocity pulse-jet systems, however,
large amounts of dust are always present within the structure of the felted
material, and a true dust cake is not always formed on the surface.7'9»10 It
is thus not possible to distinguish the dust which is conditioning the fabric
and the dust which is forming the cake; under these circumstances, It seems
more appropriate to consider the clean fabric separately and account for all
of the dust in the dust deposit specific resistance,
Theoretical calculations of pressure drop in a fabric filter system are
usually based on the assumption that the cleaning mechanism employed is 100%
effective In removing the dust deposit added since the last cleaning cy-
cle.^11"13 Under this assumption, the pressure drop immediately after clean-
ing is the pressure drop across the (conditioned) fabric only, Ap^:
In this case the pressure drop just before a cleaning pulse is equal to
Apf plus the pressure drop across the dust deposit added since last cleaning:
Ap - Apf + Apd
Darcy's Law then takes the following form:5"8
Ap = KjV + 1^2 vw
(1)
(2)
(3)
APb - Krv + K2vwq
(4)
where Wq is the incremental dust deposit added during one cleaning cycle.
192
-------
If dust mass is retained on the fabric, however, the model must be modi-
fied to include the pressure drop across this deposit. The residual dust
mass in a fabric filter can be characterized by the mass retention coeffi-
cient, y, defined as the fraction of the dust mass which remains on the bag
after pulse cleaning.3 If a fabric filter system is operated to equilibrium
with a dust deposit retention of y, the pressure drop before a cleaning pulse
is:3
Apb = Kxv + K2vwq
U-Y
(5)
The pressure drop after a cleaning pulse is:3
AP<
V + V0ll-Y
(6)
Dust Deposit Profiles
Mass retention phenomena can lead to an uneven distribution of the resid-
ual dust mass along and among the bags in a fabric filter system,3*14 which
will lead to a corresponding uneven velocity profile along each bag (assuming
a constant specific resistance). Since the pressure drop is the same at all
points on each bag, a bag area with a small dust deposit will have a higher
local filtration velocity than an area with a large dust deposit.
The effect of dust deposit profiles can be investigated by dividing the
bag into n segments, each with a constant dust deposit areal density. Assume
that and K2 are constant from segment to segment, and that the segments are
numbered consecutively from one end of the bag.
The total pressure drop across all segments is equal:
Ap ¦ vi(K1 + K2wi)
(7)
The total airflow through the bag must equal the sum of the flows through the
Individual segments:
n
Q = vA - X vA. <8)
i-1 1
Substituting Equation 7 into Equation 8 and rearranging,
¦1
A,
Ap ~ vA
n
£
i-1 K1 + K2wi
(9)
Equation 9 states that pressure drop can be calculated from the harmonic aver-
age of the local dust deposit areal densities.15
193
-------
Factors Affecting Pressure Drop
Equation 9 gives the pressure drop across the dust and fabric for an
unevenly distributed dust deposit; an examination of this equation reveals
the factors which can cause a change in the pressure drop. These factors
include:
1) changes in v;
2) changes in the magnitude of the wj values (i.e., changes in the aver-
age dust deposit areal density, wf;
3) changes in the distribution of the w^ values;
4) changes in the fabric resistance, Kj; and
5) changes in the dust deposit specific resistance, K2.
The possible contributions of changes in each factor listed above to
changes in pressure drop will be discussed.
Changes in Velocity. Equation 2 predicts a direct proportionality be-
tween pressure drop and superficial filtration velocity. If Ki and are not
functions of velocity and the dust deposit is constant, a doubling of the fil-
tration velocity will cause a doubling in the pressure drop. Conversely, any
change in Ap other than a doubling when v is doubled will reflect a change
in the magnitude of the dust deposit, the distribution of the dust deposit,
Kj_, and/or K2.
Changes with Dust Deposit Profile. Changes in the profile of the dust
deposit can influence pressure drop. Solutions in Equation 9 reveal that, for
a given mass of dust, the highest pressure drop occurs when the dust is even-
ly distributed over the fabric surface.16 The wider the variation in w^, the
lower the pressure drop.
Changes with Dust Deposit Magnitude. Changes in the magnitude of the
average dust deposit have a direct effect on pressure drop. Equation 2 states
that a doubling of the average dust deposit areal density will cause a doub-
ling of Ap, assuming constant v, K^, and 1^.
Changes with Ki and K?. may also vary with filtration velocity, per-
haps due to a compression of the fibers comprising the felt. In many cases,
however, the pressure drop across the fabric is much smaller than the pres-
sure drop across the dust deposit; when this is so, changes in will have
little effect on the total pressure drop.
Previous work has shown the dust deposit specific resistance, K^, to be
a function of the superficial filtration velocity. Billings7 summarizes data
which seem to indicate a linear relationship between filtration velocity and
specific resistance; i.e., "increases of velocity by factors of 2 or 3 tend
to produce similar increases in K2." Dennis17 found K2 to be proportional to
v0'^ for fly ash collected on woven fabrics. Rudnick1® found K2 to vary with
v0,2 in his laboratory-scale tests of woven fabrics, while Borgwardt15 meas-
ured values of K2- in laboratory and field tests of fly ash on woven fiber
glass which varied with v1,5.
K2 is usually determined by solving Equation 2, using a calculated or
194
-------
measured average value for the dust deposit areal density. If the dust de-
posit is not evenly distributed, however, this procedure will give an appar-
ent value for K2, K2a, which is the value K2 would take if the dust deposit
were evenly distributed and the pressure drop were the same as the value
measured. The actual value of K2 can only be determined if the dust deposit
profile on each bag is known. K2 can then be calculated by solving Equation
9 for K2.
EXPERIMENTS
Pressure drop and dust deposit profiles were measured on a pilot-scale
pulse-jet fabric filter.19 The fabric filter contained three commercial-
scale bags manufactured from polyester needled felt. The test dust was re-
suspended fly ash, fed at a constant mass rate to the filter inlet. Three
experiments were performed at each of five superficial filtration velocities:
50, 75, 100, 125, and 150 mm/s.
After each experiment, the dust deposit profiles along each bag were
measured using a beta gauge areal density measurement system.19 The bags
were 2.4 m long, and areal density measurements were taken every 0.1 m. The
pressure drops were measured with time, and each experiment was run until
equilibrium was reached.
Manometer fluid was accidentally deposited on one bag during the third
experiment, making it necessary to wash the bags. The dust deposit profiles
and equilibrium pressure drops were found to be significantly different be-
fore and after washing.19 Since this difference confounds the data, the re-
sults of the first three experiments are not included in the following dis-
cussion.
RESULTS
Dust Deposit Profiles
The dust deposit profiles from the fifteen experiments were described and
analyzed previously,19 and are summarized here. Figure 1 shows the dust de-
posit areal density profiles for each superficial filtration velocity, aver-
aged over the last twelve experiments. The average dust deposit areal density
is plotted against filtration velocity in Figure 2.
Pressure Drop
The equilibrium total pressure drop was measured at the conclusion of
each experiment. The total pressure drop (Apt) includes the pressure drop
across the dust deposit (Ap^), the pressure drop across the clean fabric
(Apf), and the pressure drop across the venturi nozzle (Apn). The pressure
drop appearing in the theoretical relationship of Equation 9 is the pressure
drop across the fabric and dust deposit (Apf+(j), and can be determined by sub-
tracting Apn from Apt.
Apn was measured at each filtration velocity by operating the baghouse
195
-------
0.8
H
55
0.6
-------
25 50 75 100 125 150
FILTRATION VELOCITY, MM/S
Figure 2. Average dust deposit areal density for
each of the last twelve experiments*
197
-------
without bags installed. The pressure drop across the fabric and dust deposit,
obtained by subtracting Apn from Apt at each filtration velocity, is plotted
against superficial filtration velocity in Figure 3.
According to Darcy's law, the pressure drop across the clean fabric is
a linear function of velocity (Equation 2). The fabric resistance (K^) of
the test fabric was 850 Pa-s/m (0.02 in H20/fpm), giving values of Apf which
range from 40 Pa (0.2 in H2O) for the lowest test velocity to 130 Pa (0.5 in
H2O) for the highest velocity. These values are small compared to the pres-
sure drops across the fabric and dust combined, as shown in Figure 3.
DISCUSSION
Summary of Factors Affecting Pressure Drop
The equilibrium system pressure drops, Apt, measured in these experiments
ranged from 350 Pa (1.4 in H2O) for one experiment at a filtration velocity
of 50 mm/s to 10.2 kPa (40.8 in H2)) for one experiment at a filtration velo-
city of 150 mm/s. The first number represents an acceptable pressure drop
at a fairly high air-to-cloth ratio (10 cfm/ft^), while the second number
represents a very high pressure drop at a very high air-to-cloth ratio (30
cfm/ft2). The dust deposit measurements taken during this series of tests
make it possible to determine the relative contributions of changes in fil-
tration velocity, dust deposit magnitude and distribution, and specific re-
sistance to this increase in pressure drop at higher velocities.
The factors contributing to a change in Ap when the superficial filtra-
tion velocity is changed are summarized in Tables 1 and 2. Table 1 quanti-
fies those factors contributing to a change in Ap^ when the filtration velo-
city is increased from one lower velocity (75 mm/s) to another lower velocity
(100 mm/s). Table 2 quantifies the factors contributing to the dramatic
change in Ap^ when going from the 125 mm/s filtration velocity to 150 mm/s
The specific resistances listed in these tables are the actual K2 values cal-
culated from Equation 9, not K2a values from Equation 2.
When going from the 175 mm/s filtration velocity to 100 mm/s, the pres-
sure drop across the dust deposit increases by a factor of 1.64. Most of
this increase is accounted for by the increase in velocity itself, since Ap^
should Increase by a factor of 1.33 according to Darcy's law. Changes in the
dust deposit distribution, dust deposit magnitude, and K2 all account for a
small fraction of the increase.
The causes of the nearly elevenfold increase in Ap
-------
25 50 75 100 125 150
FILTRATION VELOCITY, MM/S
Figure 3. Measured pressure drop across the dust
and fabric vs. filtration velocity.
199
-------
Table 1. FACTORS CONTRIBUTING TO THE CHANGE IN Apd AT LOWER FACILITIES
Factor
Contributing To
The Change
In Ap^
Superficial Filtration
Velocity (mm/s)
Ratio Of
Values At
The Two
Velocities
75
100
Ap^Pa)
330
540
1.64
v(ntm/s)
75
100
1.33
Distribution Of
w^ values*
0.951
0.969
1.02
Magnitude Of
w(g/m2)
394
420
1.07
% (S"1)
11,740
12,460
1.06
Ratio of the measured Apd to the value Apd would take if the dust
deposit was evenly distributed.
Table 2. FACTORS CONTRIBUTING TO THE CHANGE IN
Apd AT THE HIGHEST VELOCITIES
Factor
Contributing To
The Change
In Apd
Superficial Filtration
Velocity (mo/s)
Ratio Of
Values At
The Two
Velocities
125
150
Apd(Pa)
680
7,455
10.96
v(mm/s)
125
150
1.20
Distribution Of
Values*
0.946
0.998
1.06
Magnitude Of
ff(g/mz)
439
807
1.84
K2(s_1)
13,100
61,440
4.69
it
Ratio of the measured Apd to the value Apd would take if the dust
deposit was evenly distributed.
200
-------
Total Dust Mass
The previous discussion centered on the change in pressure drop at dif-
ferent superficial filtration velocities. It should be noted, however, that
a very large mass of dust remained on the bags after cleaning for all filtra-
tion velocities. This large dust mass contributed a great deal to the high
pressure drops measured in some of these experiments. As shown in Figure 2,
the average dust deposit areal density ranged from 0.39-0.44 kg/m^ at the
lower filtration velocities, and doubled to an average value of 0.81 kg/m at
the highest velocities. These values, when combined with the incremental
dust deposit used in these tests (w0»3.1 g/m ), resulted in mass retention
coefficients (y) of 0.993 at lower velocities and 0.996 at the highest velo-
city.
The increase in pressure drop due to mass retention can be determined
by solving Equation 4 to give Ap with no mass retention and comparing the
values obtained with the pressure drops measured with retention. The results
of such calculations at each filtration velocity are presented in Figure 4.
The curve labelled "no mass retention" represents the lowest possible pressure
drop at each velocity which could be obtained on the pilot system with the ex-
perimental conditions used. The "measured mass retention" curve represents
the pressure drops measured during the experiments; the difference between the
two curves represents the increased pressure drop caused by retention of the
dust deposit.
The curves show that mass retention at velocities through 125 mm/s, while
significant, did not cause the system pressure drop to rise to levels higher
than 2,000 Pa, or 8 in H2O . Mass retention would have to be reduced signifi-
cantly at 150 mm/s, however, to give a pressure drop in this range.
Dust Deposit Specific Resistance
Magnitude. At velocities below 150 mm/s, pressure drops measured in
these experiments were relatively low, even though a large mass of dust was
deposited on the fabric. This reflects the relatively low specific resistance
(K2) values measured here, compared to those measured on woven fabrics7»l0,17,18
The K2 values measured at 150 mm/s, while much higher than the others measured
here, are still lower than typical values for woven fabrics. Lower specific
resistance is due to the nature of the dust deposit on non-woven materials,
where some particles are deposited within the material on more-or-less iso-
lated fibers.^® The porosity of this type of deposit is much greater than
when the same particles form a particle-to-particle dust layer on the surface
of a woven fabric.
Pust Profile Effects. The actual and apparent dust specific resistance
values, K2 and K2a, are compared at each filtration velocity in Table 3. The
values listed are the averages from the experiments at each velocity. The
actual values were calculated by substituting the 72 areal densities measured
on the three bags during each experiment into Equation 9 and solving for K2
using Newton's method.
201
-------
10
7
. ~ MEASURED MASS
RETENTION
NO MASS
RETENTION
or 2
« 0.7
0.4
$2 0.2
0 25 50 75 100 125 150
FILTRATION VELOCITY, MM/S
Figure 4. Pressure drop with measured mass retention
and pressure drop with no mass retention vs.
filtration velocity.
202
-------
Tables III. COMPARISON OF APPARENT AND ACTUAL VALUES
OF DUST DEPOSIT SPECIFIC RESISTANCE.
Superf icial
Filtration
Velocity
(mm/s)
Dust Deposit Specific Resistance, s~*
9
Apparent
Actual
m
Error
50
10,400
10,800
4.1
75
11,200
11,700
4.6
100
12,800
13,400
4.9
125
12,500
13,100
4.9
150
61,200
61,400
0.5
203
-------
The errors in the apparent values of Kg at lower velocities ranged from
4 to 5%, while the error at 150 mm/s was only 0.5%. The low error at the
highest velocity reflects the essentially flat profiles measured at that ve-
locity, while the higher errors reflect the characteristic profiles measured
at lower velocities. Greater variations in dust deposit along the bags would
cause greater errors in K2a
Changes in Filtration Velocity. The changes in K2 with filtration velo-
city do not follow the pattern noted by others,7,15,17,18 who found K2 to be
proportional to va, where 0.2 < a <; 1.5. Here, K2 increased only slightly
at the lower test velocities, showing a 20% increase when the filtration velo-
city was raised from 50 to 125 mm/s. The nature of the dust deposit was evi-
dently only slightly affected by the deposition velocity over this range.
The dust-fabric structure changed dramatically at the highest test velo-
city, however, as indicated by the nearly five-fold increase in K2 accompan-
ied by an almost doubling of w. The sudden change in the dust deposit and
specific resistance seem to indicate that a fundamental change in the dust de-
posit occurs when the filtration velocity is raised above some critical value.
It seems probable that the same factor which was responsible for the doubling
of the amount of dust retained on the fabric was also responsible for the in-
creased flow resistance of the deposited dust. The amount of dust deposited
at the 150 mm/s velocity may have been sufficient to form a true dust cake
on the surface of the felt; such a cake could account for the higher K2 val-
ues measured at that velocity. Alternatively, some "collapse" mechanism may
have occurred, causing the dust-fabric structure to become more firmly packed,
making the dust more difficult to remove, and increasing K2.
SUMMARY AND CONCLUSIONS
Dust deposit retention had a significant effect on equilibrium pressure
drop at all of the tested filtration velocities.. The measured values of the
mass retention coefficient (y) ranged from 0.993 to 0.996, indicating that
over 99% of the dust deposit was not removed by each cleaning pulse. It
would be desirable to identify basic changes in the cleaning process which
will increase the cleaning efficiency.
The dust deposit's distribution, total mass, and specific resistance all
followed the same pattern with changes in superficial filtration velocity.
They remained essentially constant at all lower velocities, but changed dra-
matically at the highest test velocity. At the 150 mm/s velocity the total
mass doubled, the distribution equalized over the surface of each bag, and
the specific resistance increased by a factor of 4.7, A fundamental differ-
ence in the dust deposit formation and removal process must exist at the
highest velocity to account for these changes .
The measured equilibrium pressure drops ranged from very low (0.4 kPa,
1.4 in H2O) to very high (10.2 kPa, 40.8 in 1^0). All of these values could
have been reduced if mass retention would have been reduced; the highest pres-
sure drops were due more to the increase in specific resistance, however.
204
-------
A test program is planned to measure pressure drop and areal densitltes
at many velocities between 125 and 150 mm/s, in an attempt to determine the
precise manner in which the dust profile and K£ change over this range. Tests
are also planned to determine the relative contribution of cleaning efficiency
and redeposition to mass retention, and to measure local variations in K2.
ACKNOWLEDGMENTS
This work was supported in part by the Bird Companies Charitable Founda-
tion, Inc., and in part by EPA grant 804-700.
205
-------
NOMENCLATURE
Symbols
A area of filter bag
clean fabric resistance
K.2 dust deposit specific resistance
K2a apparent dust deposit specific resistance
Ap pressure drop
n number of bag segments
Q air flow rate
v superficial filtration velocity
w dust deposit areal density
y mass retention coefficient
Subscripts
a after cleaning pulse
b before cleaning pulse
f fabric
n venturi nozzle
t total
0 incremental quantity (one cleaning cycle)
l,2,...,n bag segments
206
y
-------
REFERENCES
1. Leith, D., S. N. Rudnick, and M. W. First. High Velocity, High Efficiency
Aerosol Filtration. EPA Report EPA 600/2-76-020. National Technical In-
formation Service, Springfield, Va., 1976.
2. Leith, D., M. W. First, and H. Feldman. Performance of a Pulse-Jet Filter
at High Filtration Velocity - II. Filter Cake Redeposition. J. Air Poll.
Control Assoc. 27(7):636, 1977.
3. Ellenbecker, M. J. and D. Leith. Effect of Dust Cake Redeposition on
Pressure Drop in Pulse-Jet Fabric Filters. Proceedings of 3rd Interna-
tional Powder & Bulk Solids Handling & Processing Conference. Rosemont,
111., May 1978.
4. Carman, P. C. Flow of Gases Through Porous Media. New York, Academic
Press, 1956.
5. Williams, C. E., T. Hatch and L. Greenburg. Determination of Cloth Area
for Industrial Air Filters. Heat. Piping Air Cond. 12(4):259, 1940.
6. Silverman, L. Filtration Through Porous Materials. Heat. Vent. 47(7):
67, 1950.
7. Billings, C. E. and J. Wilder. Handbook of Fabric Filter Technology.
National Technical Information Service Publication No. PB-200 648.
Springfield, Va., 1970.
8. Davis, W. T., P. J. LaRosa and K. E. Noll. The Generation and Evaluation
of Fabric Filter Performance Curves From Pilot Plant Data. Filtr. Sep.
13(6):555, 1976.
9. Davies, C. N. The Separation of Airborne Dust and Particles. Proc. Inst.
Mech. Eng. London. 1B(1):185, 1952.
10. Holland, C. R. and E. Rothwell. Model Studies of Fabric Dust Filtration:
1. Flow Characteristics of Dust Cakes Uniformly Distributed on Filter
Fabrics. Filtr. Sep. 14(1) :30, 1977.
11. Walsh, G. W. and P. W. Spaite. An Analysis of Mechanical Shaking in Air
Filtration. J. Air Poll. Control Assoc. 12(2):57„ 1962.
12. Robinson, J. W., R. E. Harrington, and P. W. Spaite. A New Method for
Analysis of Multicompartmented Fabric Filtration. Atmos. Environ. 1(4):
499, 1967.
13. Theodore, L., J. Reynolds, A. Corvlni and A. Buonicore. Particulate Con-
trol by Pulsed-Air Baghouse Filtration: Describing Equations and Solution.
Proceedings of 2nd Specialty Conf. on: The User and Fabric Filtration
Equipment. Buffalo, N.Y.,1975. p. 90-103.
207
-------
14. Stephan, D. J. and G. W. Walsh. Residual Dust Profiles in Air Filtra-
tion. Ind. Eng. Chem. 52(12):999, 1960.
15. Borgwardt, R. H., R. E. Harrington, and P. W. Spaite. Filtration Char-
acteristics of Fly Ash from a Pulverized Coal-Fired Power Plant. J.
Air Poll. Control Assoc. 18(6):387, 1968.
16. Ellenbecker, M. J. Pressure Drop in a Pulse-Jet Fabric Filter. Sc.D.
Thesis. Harvard School of Public Health, 1979, Appendix C.
17. Dennis, R., R. W. Cass and R. R. Hall. Dust Dislodgement from Woven
Fabrics Versus Filter Performance. J. Air Poll. Control Assoc. 28(1):
47, 1978.
18. Rudnick, S. N. and M. W. First. Dust Cake Compaction in Fabric Filtra-
tion. Paper 78-62.7 presented at the 71st Annual Meeting of the Air
Pollution Control Association. Houston, Texas, June 1978.
19. Ellenbecker, M. J. 0£. cit., Manuscript II.
20. Rothwell, E. Fabric Filter Failures—Relating Laboratory Observations
to Practice. Paper presented at the Filtration Society's Conference on
Filtration, Productivity and Profits at the Dust Control, Air Filtration
& Gas Cleaning Exhibition. London, England, Sept. 1977,
208
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ROLE OF FTLTER STRUCTURE AND ELECTROSTATICS
IN DUST CAKE FORMATION
By
George E. R. Lamb and Peter A. Costanza
Textile Research. Institute
Princeton, New Jersey 08540
ABSTRACT
This study has examined the relationship between filter
performance and dust cake formation. The latter is found to be
dependent on filter structure and on electrostatic effects. An
important feature of the dust cake is th.e distribution of mass
with distance from th.e fabric surface. When the cake forms
close to the upstream surface, lower dust retention and lower
pressure drop result, as well as better capture efficiency. This
upstream concentration of the dust cake can be achieved by proper
choice of the fibers of which the filter is made, by certain
layered constructions of the filter, and by application of an
electric field at the filter surface.
209
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ROLE OF FILTER STRUCTURE AND ELECTROSTATICS
IN DUST CAKE FORMATION
INTRODUCTION
At TRI a program of research aimed at studying basic mechanisms
of gas-particulate filtration has been active fox several years.
Various approaches to the optimization of filter performance have
been taken. These have included modification of the component
fibers and of the fabric structure, as well as the application of
D.C. electric fields between suitably located electrodes. The
changes in performance due to these modifications are assumed to
be accompanied by changes in the structure of the dust cake. This
structure, however, is difficult to observe and quantify. Micro-
photographic analysis has, in general, not yielded useful data
[Miller et al. (.1978) Leith et al. (1976) ]. The approach
taken in the present studies has been to use filters formed of a
number of thin layers. By weighing these before and after
filtration, the distribution of dust mass with depth into the
filter can be characterized. This macroscopic technique has given
useful indications of the changes in cake structure accompanying
filter modification, which are invisible to an observer looking
at the filter surface. Such information should aid the design of
better filter fabrics.
EXPERIMENTAL METHODS
All measurements were made with "patch" filters 12.5 cm x 12.5
cm. These were clamped in holders that exposed a 10-cm diameter
circle to the aerosol stream. The aerosol was generated by feeding
pulverized coal flyash into the path of a compressed air jet. The
aerosol stream moved upwards in a 10-cm duct towards the filter
at the top, so that larger particles and agglomerates with a set-
tling velocity greater than the upward flow fell to the bottom and
were not part of the measured dust cake. Dust penetrating the
test filter was collected on a full-flow membrane filter. Effi-
ciencies were determined from the mass of dust collected on both
filters. Pressure drop across the test filter was measured using
a pressure transducer and recorder, as well as a liquid manometer.
FIBER GEOMETRY
The effects of the geometry of the component fibers on the
performance of a filter fabric has been studied extensively [Miller
et al. (1977)3]. It has been shown that variations in the diameter
of the fiber (at constant filter weight and density) bring about
210
-------
considerable response in both capture efficiency and pressure
drop. However, these variations offer only a trade-off, since a
gain in one response is accompanied by a loss in the other.
Changes in fiber cross-sectional shape, however, have been shown
to result in both increased efficiency and lower pressure drop.
This is illustrated in the "performance curves" in Figure 1, which
compare filters made of fibers having different lotxed cross
sections. A performance curve is obtained by plotting pressure
drop OP) vs. penetration (mass fraction of dust coming through
the filter) for filters operating under a standard set of
conditions. Each curve contains points for filters made of fibers
with equivalent cross-sectional shapes but different linear
densities; that is, the curve is characteristic of the fiber
shape. Ideal performance is represented by a point at the origin,
and the closer a curve is to the origin, the better its relative
performance. Figure 1 compares filters made of fihers with round
and trilobal cross sections. The variations in lobe depth are
characterized by the "aspect ratio" CAR), which is the ratio of
the diameters of the circumscribed to the inscribed circles
(Figure 2). It can be seen that as the aspect ratio increases,
the performance improves. The greater efficiency would be expected
because the projected area of a lobed fiber increases Cat constant
linear density) with increasing aspect ratio. Electrostatic
effects, however, also appear to play a significant part. It has
been shown theoretically [Mi Her et al. C1978}'] that in an electric
field, lobed fibers have greater capture efficiency than round
ones, and that the efficiency increases with the number and depth
of the lobes. In any filter, local electric fields must be created
by charges on fibers and captured particles. It is possible, there-
fore, that differences in efficiency caused by changes in fiber
shape are due to the different responses of the fihers to these
local fields.
The following experiments in which local differences in electric
potential were eliminated by coating the fihers with metal appear
to support tho above hypothesis. Felts were made with nylon fibers
of round and of shallow CAR = 1.4) and deep CAR = 2) trilobal cross
sections. The surface conductivity of one set was increased by
vacuum deposition of gold. In Figure 3 penetration is plotted as
a function of number of conditioning cycles. Without a gold coating,
as expected, penetration is lower through the deep-lohed samples.
With the gold coating, this improvement disappears, and in addition,
the penetration for all three samples increases sharply at about
the 20th cycle. It appears, therefore, that electrical effects play
a part both in capture and in the stability of the dust cake,
It might not be apparent why fiber cross-sectional shape should
influence pressure drop. If anything, increased lobe depth should,
by offering greater resistance to air flow, bring about pressure
drop increases. However, the observed lower pressure drops
associated with lobed fibers are attributable to differences
211
-------
AP
(mmHeO)
100
80
-
N. o ROUND (AR"1)
Y
X TRILOBAL lAR-t.T)
60
. I
" Y "
? TRILOBAL I AR-4)
40
-
20
'
' '
0 .001 .002 .003
PENETRATION
Figure 1. "Performance curves" comparing filters made of
fibers with different aspect ratios (AR).
Figure 2. Cross section of fiber with aspect ratio 2.
212
-------
PENETRATION (%)
100 r
AS IS GOLD COATED
D DEEP <3>
S SHALLOW ®
R ROUND ®
Figure 3. Penetration through fabrics made of fibers with
different aspect ratios, with and without gold coating.
in the way dust distributes with respect to depth into the filter.
Table 1 shows amounts of dust captured in various layers of clean
4-layer filters. It is seen that the trilobal filter collects a
greater fraction in the upstream layers. This results in more
thorough cleaning, as indicated in Figure 4 which shows the amounts
left in the various layers as the layered filters are progressively
conditioned. It is seen that the residual cake mass is lower in
the trilobal filters.
The improved cleanability of trilobal filters also appears
to be partly electrostatic in origin. Table 2 shows the effect of
gold coating on the pressure drop after conditioning obtained with
round fiber and trilobal fiber filters. Whereas with the control
filter the trilobal one gave a lower pressure drop, gold coating
makes the difference vanish. The difference in the mass of resi-
dual cake likewise disappears.
213
-------
Table 1. DISTRIBUTION (%) OF DUST CAKE MASS IN VARIOUS
LAYERS OF FOUR-LAYER FILTERS MADE OF TRILOBAL AND OF ROUND FIBERS.
Layer number
Trilobal
Round
1
93.0
88.3
2
3.4
6.3
3
1.5
3.0
4
0.8
1.0
Penetration 1.2 1,5
FLYASH RETAINED
(g)
0.9
O Round
A Triloba!
0.8
All Layars
0.7
0.6
0.5
/ A
0.4
//
0.3
Remaining
Layers
0.2
40
30
CYCLES
Figure 4. Retention of dust cake in layers of multilayer
filters made of trilobal and round fibers in first forty
cycles of conditioning.
214
-------
Table 2. FILTRATION PERFORMANCE OF GOLD-COATED
POLYESTER FILTER FABRICS AFTER 50 CYCLES OF CONDITIONING.
Pressure
drop
(mm H20)
Residual cake
after cleaning
C% of filter mass]
Round (control)
Round (gold-coated)
Trilobal (control)
Trilobal (gold-coated)
145
145
142
115
122
102
75
104
COMPOSITE FILTERS
The above findings, namely that collection in the upstream
regions of a filter promotes improved performance, suggested that
other means (besides use of lobed fibers) of causing a similar
collection pattern might result in similarly enhanced performance.
A simple modification of fabric structure which would be expected
to achieve this end consists of forming the fabric in layers, and
arranging that the upstream layers are made of finer fibers.
Card webs containing minor proportions of fibers having low
softening points were prepared. Composite felts were then made
by combining layers of fine and thicker fibers and compacting
them by pressing at 120°C. The effects on filtration performance
of varying the thickness of the upstream layer, and also of
reversing the order of the layers, were examined using the eight
samples described in Table 3.
Table 3. COMPOSITION OF TWO-LAYER FILTERS
MADE FROM 3 AND 6-DENIER FIBERS.
Upstream / Downstream
100% 6d
83% 6d
50% 6d
17% 6d
100% 3d
83% 3d
50% 3d
17% 3d
0% 3d
17% 3d
50% 3d
83% 3d
0% 6d
17% 6d
50% 6d
83% 6d
215
-------
Experimental
conditions were
as
follows:
Fabric weight
Fabric density
Fiber type
Face Velocity
Inlet concentration
Cycle time
0.25 kg/m2 (7.3 oz/yd2)
0.130 g/cm3
Polyester
60 mm/s (12 ft/min)
3.5 g/m3
5 min
Measurements were made not only of the penetration and the pres-
sure drop of these filters, but also of the residual dust after
cleaning (expressed as percentage of the base filter mass).
Figures 5,6, and 7, respectively, show these measurements as
a function of %6d/%3d in the two-layer composite filters; that is,
percentage of 6-denier fibers upstream increases to the right,
whereas percentage of 3-denier fibers upstream increases to the
left. It can be seen now that the composites with 3-denier fibers
upstream allow less penetration and also exhibit a lower pressure
drop than those with 6-denier fibers upstream or with 100%
3-denier fibers. The 17% 3d upstream/83% 6d downstream composite
appears to be the best combination. These effects may be explain-
able by the correlation between the curves of Figures 5 and 6 with
those of Figure 7. It appears generally that penetration and
pressure drop are directly related to the amount of dust retained
after cleaning. Finer 3d fibers upstream prevent dust from lodging
deeper into the filter whence it cannot be blown back. This in
turn keeps the penetration lower and, since unremovable dust
increases resistance to air flow,also keeps the pressure drop lower.
ELECTRIC STIMULATION
The mechanisms whereby a fiber collects particles from an
aerosol stream more efficiently when an electric field is present
are well known [Zebel (1965)5]. Less clear are the reasons why in
some circumstances the presence of an electric field also causes
a reduction in pressure drop. It appears that, as in the cases
examined above, the pressure drop response is associated with
changes in the cake structure.
Figure 8 shows the distribution of cake mass in the different
layers of a filter made up of thin separable nonwoven fabric.
These were weighed after five minutes of filtering a 3-g/m3 flyash
aerosol. One set of results was obtained with no applied field.
The second (marked 4 kV/cm) was obtained with 6 kV applied to
alternate parallel wire electrodes resting on the upstream surface
of the upstream layer (no. 1). With the applied field, there is a
measurable shift in the distribution in an upstream direction. At
the same time, there is a large reduction in the penetration (from
216
-------
120
-
no
-
100
-
90 1
<
> /
PENETRATION (X)
0.81-
upatraam
0/100 17/83
90/80 63/17 100/0
%8d/%3d
0/100 17/83
90/80
%6d/%3d
83/17 100/0
Figure 5. Pressure drop across composite Figure 6. Penetration for the composite
polyester filters having different propor- filters of Figure 5.
tions of 3 and 6-denier fibers in two layers.
OUST RETAINED
AFTER CLEANING
(%of«Uobrle)
300-
upnraam
(TOO 17/83 90/80
%8d/%3d
83/17 100/0
Figure 7. Residual dust after cleaning
the filters in Figure 5.
217
-------
100
90
>
4 kV/cm
80
-OKV G
70
-
60
-
50
II
40
II
30
w
20
VI
10
\
0
i i X-
layer number
O OkV
* 4kV/cm
Pentlrotion
Figure 8. Distribution of dust cake mass and overall pene
tration for four-layer filter with and without applied
electric field.
7% to 0.2%) and also a decrease in pressure drop. The above
result appears paradoxical, since if in the presence of an electric
accumulation in a fabric is shifted in an upstream
greater mass concentrated in the first layer would
lead to greater pressure drop. In fact the opposite
occurs, and we must conclude that cake formation in the uppermost
field the dust
direction, the
be expected to
layer is modified by the presence of the field.
Figure 9 shows photographs of the upstream surface of a needled
polyester commercial filter fabric after 6 minutes of filtering a
5-g/m3 flyash aerosol without an applied field (Figure 9a) and with
a 6-k.V/cm field (Figure 9b) . In contrast with the compact cake
seen in Figure 9a where surface fibers are free of dust, in
Figure 9b there is heavy deposition on these surface fibers. Final
pressure drop measured with the cake in Figure 9a was 18 mm of
water; with the cake in Figure 9b it was 3 mm of water. The
explanation of the paradox is thus that the dust cake, on being
218
-------
Figure 9. Photographs of dust on the surface of a nonwoven
fabric.
a) Dust collected with no field.
b) Dust collected with 6-kV/cm field.
219
-------
shifted upstream by the field, is also made to form in a region of
the fabric where the packing density Cor volume fraction of fibers)
is lower. It follows that pressure drop reduction by an applied
field depends on the presence on the upstream surface of the filter
of a region of low packing density.
Accordingly, the dependence of pressure drop reduction on the
packing density a of the surface layer was studied using a series
of two-layer composite fabrics in which the upstream layers (bonded
by a thermoplastic binder) were made with different packing densi-
ties by being formed at different pressures. Figure 10 shows
plots of pressure drop ratios (PDR) versus applied voltage for the
different composites. PDR is the ratio of the pressure drop
obtained with the applied field to that without a field. It can
he seen that the response is indeed strongly dependent on the
packing density of the upstream layer. These results have a
significant influence on the power consumption accompanying a
certain level of pressure drop reduction, as seen in Figure 11 for
two of the composites in Figure 10. The lower surface density is
seen to be associated with much lower power levels.
a *0.136
a »0.18
2 4 6 6 10 12 14
APPLIED POTENTIAL DIFFERENCE (kV)
Figure 10. Dependence of PDR on applied voltage for filters
with surface layers of indicated packing density ex. Face
velocity 6 cm/s.
220
-------
09
OB
or
OjS
05
0.4
03
0.2
0.1
V«6 cm/«
a* 0.09
V*6 cm/1
a *0.015
0.001 0.01 0.1 1
POWER CONSUMPTION (W/m2)
10
100
Figure 11. Pressure drop ratio vs. power needed to obtain
it for two surface layer packing densities.
ACKNOWLEDGMENTS
This work was carried out under Grant Number 8*80^926
from the Environmental Protection Agency (Project Officer: James
H. Turner), and tfte TRI research project "New Technology for
Filtration of Gases by Fibrous Media,w supported by a group of
Corporate TRI Participants.
REFERENCES
1. Miller, B., G. Lamb, P. Costanza, D. O'Meara, and J. Dunbar.
Studies of Dust Cake Formation and Structure in Fabric
Filtration. EPA-600/7-78-095, June 1978.
2. Leith, D. , S. Rudnick, and M. First. High Velocity, High-
Efficiency Aerosol Filtration. EPA-600/2-76-020, January 1976.
3. Miller, B., G. Lamb, P. Costanza, and J. Craig. Nonwoven
Fabric Filters for Particulate Removal in the Respirable Dust
Range. EPA-600/7-77-115, October 1977.
4. Lamb, G., P. Costanza, and D. O'Meara. Electrical Stimulation
of Fabric Filtration. Part II: Mechanism of Particle Capture
and Trials with a Laboratory Baghouse. Textile Res. J. 48:
566-573, October 1978,
5. Zebel, G. Deposition of Aerosol Flowing Past a Cylindrical
Fiber in a Uniform Electric Field. J. Colloid. Sci. 20: 522-
543, 1965.
221
-------
PRESSURE DROP IN ELECTROSTATIC FABRIC FILTRATION
By:
Teoman Ariman
Department of Aerospace and Mechanical Engineering
University of Notre Dame
Notre Dame, Indiana 46556
and
Dennis J. Helfritch
APITRON Division
American Precision Industries
Charlotte, North Carolina
ABSTRACT
Fabric filtration systems have been employed in industry for over a cen-
tury with relatively few technological modifications. However, with the re-
cent substantial increase in energy costs, conservation in energy consumption
has become vitally important. As a result, the filtration systemsof yester-
year may not be the best approach for future applications. Recently, an ex-
ternal electrical field was considered in fabric filtration of industrial dust
with very promising initial results. An increase in the collection efficiency,
particularly for fine particulates, and a decrease in the pressure drop was
observed. In this paper the further results of an experimental program in the
investigation of pressure drop in the electrostatic fabric filtration in in-
dustrial dust control are presented. The basic apparatus, a bench-side electro-
static fabric filtration system, creates a representative dust cake under
specific conditions of operating parameters and charge levels. The results
clearly indicate that filter and dust cake resistance,or pressure dropjde-
creases substantially with the increased electrostatic field strength for all
industrial dust samples tested,regardless of fabric type and other relevant
parameters.
222
-------
1. INTRODUCTION
The energy consumption of a fabric filter air pollution control system is
a function of the pressure drop across the fabric-dust cake combination and of
the work required to remove collected dust from the fabric. Billings and
Wilder (1970)1 in a survey of operational systems in the field noted that the
average pressure drop across a fabric filter is approximately 6 inches of water,
a value which creates an electrical cost of 35f cfm-year (assuming a charge
rate of 3.5
-------
cc
UJ
I
IL.
o
£2
X
o
z
u
u
z
llJ
o:
uj
u.
u.
o
UJ
tr
3
U)
V)
UJ
a.
a.
70
60
50
40
30
A ¦ No charge
B - Positive charge
C = Negative charge
0 12 3
WEIGHT OF DUST ( Oz. per Sq. Ft.)
Flfuro 1
From both the theoretical and experimental standpoints, the application
of electrostatics to fabric filtration systems tends to augment the already
comparatively high value of collection efficiency (with the obvious exception
of a case in which particles and fabric-dust cake are unipolarly charged).
However, work is needed to advance the state-of-the-art in formulating and
predicting pressure losses across the fabric-dust cake system. In particular,
as indicated in the work of Buttersworth (1964)2, Ariman and Lane (1972)^,
Ariman and Helfritch (1977)^ and Ariman (1978)", (1979)^ there, have been few if any
initial attempts to show the energy savings associated with the application of
electrostatics. It appears that there is tremendous potential for significant
advances in both modelling and in experimental work in this area.
When dealing analytically with the collection of dust from an air stream
past a cylindrical fiber, inertia, interception, sedimentation and diffusion
of particles are generally taken into consideration. Electrostatic effects
are mostly neglected although they often occur and may play an important role
as indicated by Ariman et al (1973)®, Fuchs (1966)^, Davies (1973)^® and
Frederick (1961)H. The difficulty of including electrostatic considerations
in modelling of ordinary fibrous filters is that it is usually impossible to
make any statements about the electrical charge on fibers. Theoretically,
this difficulty is eliminated by placing a filter in an electrical field. In
this spirit, Kraemer and Johnstone (1955)12 made a comprehensive analytical
and experimental study of aerosol collection on a spherical collector. They
obtained computer solutions assuming potential flow around the collector and
they also derived several approximate equations for the collection efficiency
of isolated cylindrical fibers. Dawkins (1958)13 extended Kraemer's computer
solutions to cylinders in potential flow, but was unable to solve the viscous
224
-------
14
flow case. Natanson (1957) dealt with the deposition of charged and un-
charged aerosol particles on a single uncharged fiber; coulombic and
polarization forces, as well as image forces, were taken into account. For
essentially isolated spheres and cylinders comparisons of the collection ef-
ficiency for electrical forces with that due to other mechanism shows that
for low velocities, the electrical collection may be clearly dominant. An
experimental confirmation of this prediction was given by Lundgren and Whitby
(1965)15, Natanson's problem has also been dealt with by Yoshioka jet^ al
(1968)16 w^0 computed numerical solutions. They carried out experiments on
the filtration efficiency of homogeneous aerosols with particle radius of
0.5 pm charges from 0 to 70 electrons per particle.
17 18
Rossano and Silverman (1954) and Goyer jit al (1954) experimentally
measured the effect of fiber charge and particle charge on filtration. Their
results show an increase in efficiency due to particle charge, fiber charge,
and filter packing density. Silverman et al (1956)19 investigated the effect
of electrostatic charge on the filtration of atmospheric dust. The continual
motion of a fabric over a fibrous filter was used to produce a continually
charged filter. Their results show filter efficiency increases inversely
with velocity and directly with fiber charge. Rivers (1962)20 compared some
experimental data with theoretical calculations on a commercial electrified
filter. His results show an increase in filter efficiency can be obtained by
placing a strong electrical field across the filter. An extensive review of
the subject of the electrical behavior of aerosols is given by Whitby and Liu
(1966)21. Gillespie (1955)22 obtained rough estimates of deposition on a
cylinder in potential flow under the action of electrostatic forces. Neglecting
inertia and making assumptions for the net effective charge on fibers, he was
able to obtain a reasonably good agreement between theory and experiment. For
different degrees of electrification of filter and particles, he presented
penetration versus particle size curves. The collection on a single fiber
model by electrostatic mechanisms was also investigated by Thomas and Woodfin
(1959)23, Sweitzer (1961)24 Dennis et al (1958)25, Anderson and Silverman
(1958)26, and Lapple (1970)^7.
Havlicek (1961)28 and Silverman et al (1952)29 have markedly improved the
efficiency of conventional fiber filters through application of an external
electrical field. Walkenhorst and Zebel (1964)30 and Zebel (1966)3! adopted
a different course. They did not place the conventional fiber filter between
sievelike electrodes but employed an open arrangement of some 30 to 100 layers
of regularly arrayed fibers. The strong electrical field was thus homogeneous
compared with other experiments, so that all fibers and aerosol particles were
uniformly polarized. The experimental results showed that a satisfactory de-
position was obtained. Because of the large distance between fibers it was not
necessary to have recourse to terms such as porosity, packing and density.
Zebel (1968)32 theoretically investigated the deposition of uncharged and
charged aerosol particles from an air stream upon an isolated circular cylind-
rical fiber perpendicular to the stream in a homogeneous electrical field.
The particle trajectories were calculated for both potential and viscous flows
by Lamb's formulation. It should be noted that if higher collection efficien-
cies are present, the single fiber model generally overestimates deposition
as indicated by Ariman et^ al (1978)3 and Ariman (1979)33.
225
-------
g
Recently Ariman and his co-workers have (1973) developed an improved
model for fabric filters. This model assumes several layers of evenly ar-
ranged fibers of cylindrical cross section of radius b. The axes of the
electrical field E is assumed. All cylindrical fibers and spherical dust
particles are assumed to be uniformly polarized by the electrical field. The
flow and electrical fields in the vicinity of an individual fiber influence
the motion of a particle. These fields are in turn subject to the interference
of the two immediately neighboring fibers. Here two cases were considered,
the first for Fu << 1, where Fu = 2b/d(0 < Fu < 1), and the second for Fu
-------
3. EXPERIMENTAL PROGRAM
The basic apparatus which was employed in the experimental program is
shown schematically in Figure 3. Primarily, it consists of a means for
creating a representative dust cake on a fabric sample under varying conditions
of relative humidity and charge levels. At constant face velocity, suspended
dust particles were filtered by an initially clean fabric filter. The particles
were charged prior to deposition by means of passage through a corona discharge
(Figure 4). Degree of charge was controlled by the magnitude of the field
strength, and pressure drop was recorded after a specified weight of dust had
been deposited. Then the similar testing is performed for different dust/
fabric type combinations.
Air flow rate through the system is measured with the aid of a laminar
flow element and a pressure gage. Relative humidity is sensed with an electric
hygrometer and the temperature in the sample holder is monitored with the aid
of a thermocouple system. Pressure drop across the filter is measured by an
inclined manometer. The system permits a variation in the flow rate over a
range such that a filtering velocity range of 0-15 ft/min can be obtained and
that relative humidity can be controlled over the range of 15-80 percent.
The electric field strength is varied between 0-12 kv/in.
Nine dust samples provided by Apitron division of the API were utilized in
the bench-scale tests. The resulting extensive data were used in the formu-
lation of the effects of electric field strength, flow rate, dust type, fabric
type, etc. or pressure drop. The dust samples were:
a.
Electric Furnace Fume
b.
Limestone
c.
Asphalt Plant Rock Dust
d.
Phosphate
e.
Gypsum
f.
Calcium Carbide
g-
Asbestos
h.
Wedron Silica
i.
Fly Ash
Some selected results are presented in the following section.
3. BENCH SCALE TEST RESULTS
In Figure 5, pressure drop values in inches of water are presented as a
function of the filtering velocity (ft/min) for charged and non-charged gypsum.
Fabric type was wool felt with a weight of 16 ounce per square yard and re-
lative humidity and temperature were kept at 36% and 81°F respectively during
the test runs. Three tests were performed at 5, 10 and 15 ft/min filtering
velocity with approximately the same amount of gypsum. Thus in the graph CW
represents the cake weight and is determined by weighing the initially clean
flat fabric filter sample (9" x 11.5") before and after the test. The fabric
filter sample was cleaned for each test by a brush. It is seen that in all
tests electric charge on particles causes particularly for U - 15 ft/min a
227
-------
AIR
INLET
DUST HOPPER
MOTOR -v
AC SUPPLY
FLEXIBLE TUBING
TEST TUBE
BRUSH INSIDE'
-VIQRATING engraving
CONDENSING DOX
OC HIGH VOLTAGE
SUPPLY
80ILEH
CORONA OEVICE
SLIDING
VALVE
DISPERSION
DUCT
\y FILTER MOUNT
|B3| ©
ELECTRIC HYGROMETER
INCLINED MANOMETER
SENSOR
LAMINAR FLOW
ELEMENT
PRESSURE CAGE
(O-SIN. H20)
EXHAUST
FAN
RHEOSTAT
~~ I I
AIR OUTLET
Figure 3
) ELECTRODE
INLET
CORONA
WIRE
+ ) ALUMINUM
TUBE
(2in. I.D.)
WEIGHT AND
STABILIZER
PUTTY
FUNNEL
Figure 4
228
-------
considerable decrease in pressure drop or filter resistance. It appears that
the majority of the decrease takes place for V = 7 kv/in with little additional
decrease with V * 10 kv/in. This situation more or less was valid for all dust
samples tested.
GYPSUM
T * 80* F V ¦ 7 KV/in.
GYPSUM
6 I
5
o _
c\j 4
X ^
3 -
a
<
2 -
WOOL FELT (I6 0J.)
T ¦ 8I*F
RH ¦ 36*
CW.24.5gr.
No Chargt
7 KV/in
10 KV/in.
u ft. min.
8 "
o
N
X
¦C, 6
a.
«
FIBERGLASS TEFLON B ( 26 oz )
RH » 34%,
POLYESTER FELT (22 oz.)
RH = 57 % -
WOOL FELT 116 oz.)
RH » 36 7.
\ CW = 24 ar
ORLON FELT (l6oz.)
RH * 38%
U ft. mm
25 gr.
Figure 5 Figure 6
The effect of fabric type on pressure drop for the charged case of V ¦ 7
kv/in is demonstrated in Figure 6. Orion felt, wool felt, fiberglass teflon B
and polyester felt fabric samples were utilized in the tests. It appears that
under the similar operating conditions orlon felt (16 oz) and wool felt (16 oz)
were the best performers and fiberglass teflon (26 oz) was the worst one from
the pressure drop point of view.
Figure 7 represents the pressure drop variations for limestone and for
non charged and charged cases with Polypropylene (15 oz) and Orion (16 oz) at
the filtering velocities of 5, 10 and 15 ft/min. For all tests it is observed
that electric charge on limestone particles causes a substantial decrease in
pressure drop particularly for higher filtering velocity values of 10 and 15
ft/min.
229
-------
LIMESTONE PHOSPHATE
No Charge
POLY PROP. (I5oz. )
T = 84° F
RH = 52 %
CW= 30 gr
O
c\
X
c\j
c
7 KV/in.
a.
<
(0 and 12
KV/in.
VELOCITY ( ft./min.)
CW - 52 gr.
POLYESTER FELT <22oz.)
T - 74* F
No Charge
52 gr
47gr
7 KV /in
48 gr
VELOCITY (ft/min)
10
NOMEX
& 6
No Charge
CO
53 gr
7 KV/in.
VELOCITY f ft/mini
NoCharg«i
ORLON (16 02.)
CW • 30gr.
O
c,
X
c
Q_
<1
7 KV/in
¦12 KV/in.
lOKV/in.
VELOCITY (ft./min.)
Figure 7 Figure 8
As stated earlier the degree of charge was controlled by the magnitude of the
field strength, namely through the voltage values of V = 7, 10 and 12 kv/in
cases the tests were performed for a dust cake weight of 30 grams. It is re-
markable to notice that for the Orion 16 oz filter sample and noncharged case
pressure drop is 8.2 inches of water while for the charged case of 7 kv/in the
corresponding pressure drop value is substantially lowertnamely 2.1 inches of
water. It is therefore estimated that by charging limestone particles, the
filtration rate per unit fabric area of a given fabric can be increased by a
factor of up to four. This implies that with the aid of electrostatics it
might reasonably be expected fabric filters can be reduced in size by a factor
of four, while achieving better collection efficiency. It is further seen from
the figure that there was no noticeable difference for the cases of 10 kv/in
and 12 kn/in. Again the majority of the decrease in the pressure drop takes
place for v = 7 kv/in case.
Figure 8 presents pressure drop results for phosphate and for Nomex (9 oz)
and Spun Nomex (9 oz) fabrics again for two difference face velocities of 10
and 15 ft/min for charged (7 kv/in) and non-charged cases. In accordance with
the test results of the previously stated dust samples, a considerable decrease
in pressure drop is again observed due to the effect of electric charge on
particles.
Pressure drop - face velocity relation is presented in Figure 9 for electric
230
-------
ELECTRIC FURNACE FUME
SPUN NOMEX (9 02.)
NoCharge
7 KV/in.
o
10
5
VELOCITY (ft./min.)
6
SPUN GLASS
No Charge
4
2
7 KV/in.
0,
VELOCITY (ft./min.)
Figure 9
furnace fume. In accordance with the procedure of the previous tests no charge
and charged (7 kv/in) cases were investigated. Nomex (14 oz) and Spun Nomex
(10 oz) were utilized for the tests of Figure 9. As expected a considerable
decrease took place in pressure drop for the charged case. It is further noted
that for the similar operating conditions Spun Nomex fabric filter causes some-
what higher pressure drop as compared to that for Nomex fabric filter. Namely
3.6 inches of water for Spun Nomex versus 3.2 inches for non charged limestone
particles and 1.6 inches versus 1.3 inches for the charged case. Note that
this interesting observation was made despite the fact that Nomex (14 oz)
filter is heavier than the Spun Nomex (10 oz) one. Therefore it appears that
in addition to the fabric weight, fabric characteristics need also to be con-
sidered in the pressure drop investigations.
In order to investigate the effect of the increase in the charge on
particles in a continuous form, a series of tests were performed with Wedron
Silica by using Orion (16 oz) fabric filter sample at the face velocities of
10 and 15 ft/min and with no charge and V « 5, 6 and 7 kv/in cases (Figure 10).
For the higher face velocity between V=0 and V"5 kv/in no noticeable change
took place in the pressure drop. However, following the formation of the
corona discharge around 5.3 kv/in, a sudden decrease in the pressure drop was
observed. Namely from 3.8 to 2.1 inches of water for V-6 kv/in. A further de-
crease with less magnitude also took place when V was increased to 7 kv/in.
A continuous decrease in pressure drop with an increased rate was observed
wity higher V values for the face velocity of IMO ft/min.
231
-------
WEDRON SILICA
ORLON (16 oz. >
T =¦ 89°F
RH * 58%
C W = 23 gr.
15 ft./min.
10 ft./min.
0
2
4
6
8
VOLTAGE (KV/in, )
Figure 10
ASPHALT PLANT ROCK DUST
ORLON (16 oz.)
T = 81°
RH¦70 %
CW« 25gr.
15 ft./min.
4 ¦ •
10 ft/min.
7
-7
VOLTAGE (KV/in.)
Figure 11
232
-------
Finally the effect ot the change of sign of corona discharge field was
also investigated for an Asphalt Plant Rock Dust and Orion (16 oz) filter.
It is quite interesting to note that for both filtration velocities of 10
and 15 ft/min there is no noticeable difference on pressure drop with the
change of the sign of the electric field (Figure 11).
5. CONCLUDING REMARKS
An experimental program was carried out for the investigation of pressure
drop of electrostatic fabric filtration in industrial dust control. The re-
sults of the experimental program clearly indicate that filter and dust cake
resistance or pressure drop decreases substantially with the electrostatic
charge on particles. Although there are quantitative differences in the magni-
tude of the effect of charge on dust particles prior to filtration, the effect
is qualitatively identical for all industrial dust samples tested.
For a possible explanation of the substantial reduction in the pressure
drop, the microstrueture of each dust cake sample was studied under a micro-
scope. As a characteristic example two pictures of an asbestos cement dust
cake for the uncharged (Figure 12) and charged (Figure 13) are presented
here. It appears from the comparison of these pictures that the microstructure
of the dust cake of the charged asbestos cement dust differs a great deal from
the noncharged cake. The dust cake formed on an industrial fabric sample with
the deposition of charged asbestos cement particles looks much more porous
and fluffy than the one formed with the deposition of uncharged particles on
the same type of fabric sample. In turn it is quite clear that more porous
dust cake would give rise to a small resistance against the dust laden gas
flow. Note that a substantial increase in the porosity of the charged dust
cake was observed for all the dust and fabric samples used in the test programs.
Furthermore it also appears that again due to the quite loose and fluffy
microstructure of the charged dust, it would be easier to clean an industrial
bag when an external electric field is present. Thus, it would appear that
considerably less frequent fabric cleaning could be employed in electrostatic
fabric filter systems with an additional energy savings and also an extension
of bag life which typically lasts for a year. This latter factor is an im-
portant consideration since a typical expenditure of 20<:/cfm - year is required
for labor and materials to replace fabrics.
It is shown in this research program that the electrostatics appears to have
a profound effect on the flow resistance of fabric filters. It is seen that
by charging industrial dust a decrease in the pressure drop up to 80% is obtained.
Thus, it appears that substantial savings could be achieved in the energy costs
from incorporation of electrostatics into bag house filtration in industrial
dust control. Furthermore a considerable increase in the collection efficiency
especially for fine particulates, mass mean diamter less than 3 micron, is
another important outcome of the electrostatic fabric filtration.
233
-------
UNCHARGED ASBESTOS CEMENT DUST
DEPOSITED ON FABRIC - 5X MAGNIFICATION
Figure 12
CHARGED ASBESTOS CEMENT DUST
DEPOSITED ON FABRIC - 5X MAGNIFICATION
Figure 13
234
-------
REFERENCES
1. Billings, C.E. and Wilder, J. Handbook of Fabric Filter Technology, Volumes
I-IV, GCA-TR-70-17-G, prepared for NAPCA (EPA)(1970).
2. Butterworth, E., Manufacturing Chemist, 65, Feb. (1964).
3. Air Polution Control, Part 1, Edited by W. Strauss, 337, Pergamon Press,
New York (1966).
4. Ariman, T. and Lane, M.J., "On Experimental Determination of Collection
Efficiency and Pressure Drop in Fabric Filters in an Electrostatic Field,"
unpublished report, (1972).
5. Helfritch, D.J. and Ariman, T., "Electrostatic Filtration and the Apitron
Design and Field Performance,"Novel Concepts, Methods and Advanced Tech-
nology in Particulate/Gas Separation, Teoman Ariman, Editor, University
of Notre Dame.
6. Ariman, T., "Pressure Drop in Electrostatic Fabric Filtration," UND,
Final Report, December (1978).
7. Ariman, T., "Electrostatic Fabric Filtration in Industrial Dust Control -
A Review," UND-DOW-EFF TR. No. 1, April (1979).
8. Ariman, T., Rao, K.S., Yang, K.T. and Hosbein, R.L., "Collection of Dust
by Fabric Filtration in an Electrostatic Field," Proceedings of the Second
Annual Environmental Engineering and Science Conference, p. 555 (1973).
9. Fuchs, N.A., The Mechanics of Aerosols, Pergamon Press, Oxford (1966).
10. Davies, C.N., Air Filtration, Academic Press, New York (1973).
11. Frederick, E.R., Chem. Eng., 68, 107 (1961).
12. Kraemer, H.F., and Johnstone, H.F., Ind. Engng Chem., 47, 2426, (1955).
13. Dawkins, G.S., Technical Report No. 15, Engineering Experiment Station,
University of Illinois (1958).
14. Natanson, G.L., Proc. of the Academy of Science, U.S.S.R., Physical Chem-
istry Section, 112, 95 (1957).
15. Lundgren, D.A.. and K.T. Whitby, I and E.C. Fund (1965).
16. Yoshioka, N., Emi, H., Hattori, M., and Tamori, I., Kagaka Kugaku, Chem.
Eng. Japan, 32, 815 (1968).
17. Rossano, A.J., Jr., and Silverman, I.., AEC NYO-1594. Air Cleaning Lab-
oratory, Harvard School of Public Health (1954).
18. Goyer, G.G., Gruen, R., and LaMer, V.K., J. Phys, Chem.t 58, 137 (1954).
235
-------
19. Silverman, L., Conners, E.W., Jr., and D.M. Anderson, AEC NY0-4610, Air
Cleaning Laboratory, Harvard School of Public Health (1956).
20. Rivers, R.D., ASHRAEJI, 4_» 37 (1962).
21. Whitby, K.T. and B.Y.H. Liu, Aerosol Science, Academic Press (1966).
22. Gillespie, T., J. Colloid, Sci., 10, 299 (1955).
23. Thomas, J.W. and Woodfin, E.J., Applic. and Industry, (AIEE)(1959).
24. Sweitzer, D.,"Electrets, Literature SearcH'No. 308, Jet Propulsion Lab-
oratory, Cal. Inst, of Tech., (1961).
25. Dennis, R., Kristal, E., and Silverman, L., AEC NY0-4614, Air Cleaning
Laboratory, Harvard School of Public Health (1958).
26. Anderson, D.M., and Silverman, L., AEC NYO-4615, Air Cleaning Laboratory,
Harvard School of Public Health (1958).
27. Lapple, C.E., Advances in Chemical Engineering, JS, Academic Press, New
York (1970).
28. Havlicek, V., Int. J. Air Water Pollution, j4, 225 (1961).
29. Silverman, L., Billings, C.E., and Bennis, R., AEC NYO-1592, Air Cleaning
Laboratory, Harvard School of Public Health.
30. Walkenhorst, W., and Zebel, G., Staub, 24, 444 (1964).
31. Zebel, G., J. of Colloid Sci., 20, 522 (1965).
32. Zebel, G., Staub, (English Translation) 26^, 18 (1966).
33. Ariman, T., "Collection of Dust by Fiber Filters in an Electrostatic Field,"
Survey Article, to be published.
34. Scheidegger, A.E., The Physics of Flow Through Porous Media, The Macmillan
Company, New York (1960).
ACKNOWLEDGEMENT
This investigation has been supported by the APITRON division of American
Precision Industries and also by the Department of Energy, Industrial Energy
Conservation under the contract EC-77-S-2-4428 to the University of Notre Dame.
The senior author acknowledges the encouragement and interest of Mr. John R.
Rosmeissl, the project monitor during the course of this research program.
236
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EXPERIMENTAL ADVANCES IN FABRIC
FILTRATION TECHNOLOGY IN JAPAN
- Effects of a Corona Precharger and Relative
Humidity on Filter Performance -
By:
Koichi Iinoya and Yasushige Mori
Department of Chemical Engineering
Kyoto University
Kyoto, Japan 606
ABSTRACT
1. The effects of a corona precharger on the performance of a fabric filter
have been studied experimentally in air with controlled humidity. Test dust,
fine calcium carbonate powder, is precharged and then ducted to a filter
fabric. The charged dust for the most part reduces both the pressure loss
and the particle penetration through the filter, on which the same amount of
dust is collected as in the experiments without the precharger.
2. Fundamental experiments have been conducted to clarify the effects of
air humidity on the performance of a fabric filter. Three kinds of test
dust (calcium carbonate, Kanto-loam and flyash) are used. In general, the
humidity of air does not have much of an effect on the pressure loss and the
particle penetration, when the agglomerated particles size distributions of
the airborne dusts are controlled so as to be almost constant in spite of
the different humidity of air.
237
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1. EFFECT OF CORONA PRECHARGER ON FILTER PERFORMANCE
There are several experimental reports on the performance of a
combination dust collector of an electrostatic precharger and a fabric filter,
which reduces the pressure loss.l)~5) However, they are mostly concerned with
the qualitative results of fundamental experiments or the overall performance
of a pilot scale filter. A preliminary but quantitative experiment on the
hybrid filter is conducted herein by use of a bench scale test facility.
1.1 Experimental Apparatus and Procedure
Figure 1 shows a flow diagram of the experimental apparatus. Fine
calcium carbonate is uniformly fed to a mixer type disperser as a test
dust by a Micro Feeder (Sankyo Dengyo Co.), and is ducted to the test filter
through both a corona precharger and a settling (or mixing) chamber. The
test section is a series combination of a test fabric filter and a back-up
high performance Millipore filter, and is followed by an orifice flow meter
and a suction pump. The air flow rate is controlled so as to be constant
during a test period by use of an automatic regulator, and is recorded on a
chart with the pressure loss of the loaded test filter.
The corona precharger is a kind of a short electrostatic precipitator,
which is constructed with 8 discharge electrodes of tungsten wire (0.2 mm
x 100 mm) and 10 plate electrodes (11 mm x 100 mm). The distance between
the above two electrodes is 12 mm. The applied voltage is about 7 kV D.C.,
and the electric current is kept constant between 3 and 4.5 mA. during a
test period.
The dust collection area of a test fabric is about 50 cm2, and the
fabric is a polyester (Tetron) felt, of which the weight is 640 g/m^ and the
air permeability is 15 cm/sec for the pressure loss of a half inch of water
column. Electrostatic charge of dispersed test dust is measured just before
the test filter by use of a Faraday cage with a glass fiber filter.
Penetration of test dust is calculated from the ratio of dust weight
collected by the test filter and the weight by the back-up filter. Figure 2
shows particle size distributions of dispersed test dust sampled just before
the test filter by use of a slit type cascade impactor (Shimachsu CI50).
The size distributions of charged dust are almost the same as those of
uncharged dust, because a part of electrostatically agglomerated particles
may settle down in the settling chamber. The filtration velocity is 3.1 m/min
for all of the test runs.
6)
1.2 Experimental Results
Figure 3 shows a picture of the ragged surface of the charged dust bed
collected on a test fabric. Figure 4 is also a picture of a rather smooth
surface of the charged dust bed on the same fabric. Figure 5 is another
picture of the smooth surface of the uncharged dust bed on the same fabric.
In the case of Figure 3, pressure loss of the loaded test filter is much
lower than that of the case of Figure 5. However, the pressure loss of
Figure 4 is nearly the same as that of the uncharged dust bed shown in
Figure 5.
238
-------
Figure 6 shows a relation between drags (pressure loss divided by
filtration velocity) and dust loads deposited on a test filter for various
relative humidity of filtered air. The ragged (nodular deposit) surface often
takes place at a relative humidity of 50 to 75%. The reduced pressure loss is
apparently due to the uneven nodular deposits of test dust collected on the
fabric by the electrostatic charge effect.
Figure 7 shows the cumulative collection efficiencies by weight for
the same experimental conditions as those in the pressure loss. The
efficiencies of charged dust are usually higher than those of uncharged dust
and have no correlation with the air humidity nor the shape of the powder bed
surface.
Figure 8 shows the measured electrostatic charges of test dust just
before a test filter at various air humidities. The lower points mostly
represent the charges just before the test period, and accordingly the upper
points correspond to those just after the test period. The charges at an
air humidity of 60 to 70% are somewhat higher than those at lower or higher
humidities.
1.3 Conclusion
Pressure loss of a fabric filter may be reduced by use of electrostatic
corona precharger in front of the fabric filter at an appropriate condition.
The reduced amount is less than half of the original pressure loss for the
equal dust load deposited on a filter.
On the other hand, collection efficiency of a fabric filter is usually
improved with a corona precharger. Penetration of dust is reduced up to
10 to 50% of the original values. Therefore, the precharger seems to be a
promising device for improving the performance of a fabric filter, if the
economic cost is allowable.
References
1) Helfritch, D.J. Performance of an Electrostatically Aided Fabric Filter.
Chem Eng Prog. 73:54 - 57, August 1977.
2) Lamb, G. E. R., and P. A. Costanza. Improving Performance of Fabric
Filters. Chem Eng Prog. 73:51 - 53, January 1977.
3) Lamb, G. E. R., and P. A. Costanza. Electrical Stimulation of Fabric
Filtration. Textile Res J. 47:372 - 380, May 1977.
4) Penney, G. W. Using Electrostatic Forces to Reduce Pressure Drop in
Fabric Filters. Powder Tech (Netherlands). 18:111 -116, 1977.
5) Penney, G. W. Electrostatic Effects in Fabric Filtration: I. Fields,
Fabrics, and Particles. EPA - 600/7 - 78 - 142a., September 1978.
6) Costanza P.A., G.E.R. Lamb, and J.B.Dunbar. Electrical Stimulation of
Fabric Filtration (ESFF) for Cotton Dust Control. J Eng for Ind, Trans
ASME. 101: 65 - 68, February 1979.
239
-------
II EFFECT OF HUMIDITY ON FILTER PERFORMANCE
There are few experimental reports about the effect of relative humidity
on the pressure loss of a fabric filter. Higher relative humidity seems to
reduce the pressure loss because of particle agglomeration!)~3). However,
two previous papers present different experimental results^-*), in one of which
humidity does not have much of an effect on both the pressure loss and dust
dislodgement of a filter when calcium carbonate is used as a test dust.
Therefore, we have tried again to confirm the effect by use of our experimental
facility and three kinds of test dust.
2.1 Experimental Apparatus and Procedure
Figure 9 shows a flow diagram of the experimental apparatus. A test dust
is uniformly fed into a mixer type disperser by a Micro-Feeder (Sankyo Dengyo
Co.) and is ducted to a test filter section through a settling or mixing
chamber with a bipolar ion generator as an electrostatic charge neutralizer.
The test filter is followed by two back-up Millipore filters in series, of
which the latter serves to calibrate the humidity effect on the blank weight.
The air flow rate is measured by an orifice meter and controlled so as to be
constant during a test period by an automatic regulator. Finally, a suction
pump induces air flow with suspended dust through the whole system.
The collection surface of the test filter is round and has an area of
about 32 cm^, and the fabric is a polyester (Tetron) felt, of which the
weight is 640 g/m^ and the permeability is about 15 cm/sec at the pressure
loss of 12.7 HH11H2O.
Dust penetration is calculated from both the dust weight collected on a
test filter and that on a back-up filter. Local penetrations of oversized
particle fractions are measured by use of a light scattering particle counter
(Bausch-Lomb) and a flow sampling line. Local penetrations of the whole size
range of test dusts are also measured by use of a piezo-electric mass balance
(Thermo System Inc.). Electrostatic charges of test dusts are also measured
by use of a Faraday cage with a filter just after the settling chamber.
Three kinds of test dust are fine calcium carbonate, Kanto loam (JIS No.8)
and fly ash (JIS No.5). The filtration velocity is constant at 2.7 m/min for
all test runs.
2.2 Experimental Results and Discussions
2.2.1 Pressure Loss. Figures 10, 11 and 12 show the drags of a filter
against dust loads deposited on a test fabric for three kinds of test dust re-
spectively. Relative humidity of filtered air does not have much of an effect
on the pressure loss of a filter and these tendencies are different from the
results obtained by others. ^ This experimental discrepancy may be due to
the difference of the agglomerated size distributions of test dust just before
the test filter, as given in the next section.
2.2.2 Particle Size Distributions of Test Dusts. Particle size distri-
butions of three kinds of test dust, measured by a cascade impactor just
before the settling chamber, depend on the relative humidity of air as shown
in Figures 13, 14 and 15. The higher the relative humidity, the coarser the
240
-------
particle size. Higher humidity seems to promote particle agglomeration.
However, particle size distributions of the same test dusts, measured by
an impactor just after the settling chamber, do not show much of a change for
a wide variation in air humidity, as shown in Figures 16, 17 and 18. This
seems to be due to the sedimentation of agglomerated coarse particles in the
settling chamber.
2.2.3 Collection Efficiency or Particle Penetration. Figures 19, 20 and 21
show particle penetrations against dust loads on a fabric for three kinds of
test dust respectively. Relative humidity of filtered air does not have much
of an effect on the collection performance of the filter. The reason seems
to be the unchanged size distributions for a wide range of humidity just after
the settling chamber, as mentioned in the previous section.
Figures 22, 23 and 24 show local number penetrations of particles above
one micrometer against dust loads on the filter for three kinds of test dust
respectively. The effect of relative humidity on the number penetration has
no definite tendency, and it depends on the kind of test dusts.
2.2.4 Electrostatic Charge of Test Dust. Electric charges of dispersed test
dusts sampled just before the test fabric are measured by use of a Faraday
cage with a glass fiber filter and a high sensitivity electrometer. Figure 25
shows that these electrostatic charges of three kinds of test dust depend on
humidity in the air. However, these data show no definite tendency.
2.2.5 Moisture Content of Test Dust. Moisture contents of three test dusts
change with air humidity, as shown in Figure 26. The moisture content is
nearly proportional to relative humidity.
2.3 Conclusion
Both pressure loss and collection efficiency of a fabric filter are not
affected by the relative humidity of filtered air, when particle size dis-
tributions of any test dust in the air do not change with humidity just before
the test filter. However, the moisture contents of test dusts are nearly
proportional to humidity, and electrostatic charges of the test dust are in
the order of 10-^ to 10"? coulomb/g.
References
1) Ariman, T., and D.J. Helfritch. How Relative Humidity Cuts Pressure Drop
in Fabric Filters. Filtration & Separation (London). 14:127-130,
March / April 1977.
2) Ariman, T., and D.J. Helfritch. Effect of Humidity on Pressure Drop in a
Baghouse Simulator. A I Ch E, Symp Series. 74: No.175 10-15, 1978.
3) Durham, J.F., and R.E. Harrington. Influence of Relative Humidity on
Filtration Resistance and Efficiency of Fabric Dust Filters. Filtration
& Separation (London).8:389-398, July / August 1971.
4) Donovan, R.P., B.E. Daniel, and J.H. Turner. EPA Fabric Filtration
Studies: 3. Performance of Filter Bags Made from Expanded PTFE Laminate.
EPA-600/2-76-168c., December 1976.
241
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5) Iinoya, K., and Y. Mori. Fundamental Experiments of Fabric Filters.
EPA-600/7-79-044b. Symposium on the Transfer and Utilization of
Particulate Control Technology: Vol.2, Fabric Filters and Current Trends
in Control Equipment.
dust
feeder
recorder
dust
disperser
settling
chamber
paper filter test fabric
filter
pump
Faraday
cage
electrometer
high voltage
power supply
blower
Figure 1 Experimental apparatus for electrostatic
effects in fabric filtration
O with corona precharger
~ without corona
CaC03(D)
after precharger
at inlet Of filter
by cascade impactor
0.5 1 2 5
particle diameter, Dp (^m)
Figure 2 Particle size distributions
of test dust
242
-------
Figure 3 Photograph of
ragged surface of
deposited dust with
corona precharger,
m = 83,5 g/m2,
R.H. =73 %
Figure 4 Photograph of
smooth surface of
deposited dust with
corona precharger,
m = 86.7 g/m2,
R.H. =83 Z
C , '
Ap f"
>£L
a i
/
• :
/A
nw
Figure 5 Photograph of
smooth surface of
deposited dust without
corona precharger,
m = 138 g/m2,
R.H. = 55 %
243
-------
c\6
£
I
E
E
<
a>
a
60
SO
40
30
20
10
0
R.H.
(*/.)
with corona precharger
nodular deposit
smooth deposit
-49
—
¦
50-75
O
•
00-
A
~
* without corona precharger
" o O u-3.1 *yr
0 o O CaC03(D)
min
felt fablic( B9650S)
x
_L
30 40 50 60 70
dust load , m(fl/m2)
Figure 6 Relation between drags
and dust loads on a filter
00 90 100
O nodular deposit with corona precharger
• smooth deposit with corona precharger
x without corona precharger
u«3-' ^mln
fett fabric(B9650S)_
CaC03(D)
S
1
2.0 ~
s
1.0 -
fm
©
J
IP
S
0.5 -
%
!
111
1
a.
>r
0.2 -
_>
O
o
O
0.1 -
i
3
U
0.05 -
0.03 L
5
.
v t
<»> V
_l ' I L.
10 20 50 100 200
dust load . m ( 9^2)
Figure 7 Relation between
cumulative penetrations
and dust loads on a filter
CaC03(D)
30 40
50 60 70 80
relative humidity , R.H.( •/.)
Figure 8 Effect of relative
humidity on electrostatic
charge by corona precharger
244
-------
rdust disperser
settling chamber with ion neutralizer
_=yONj>ressure loss (J) mass monitor
pW ] (Piezobalance)
(Dparticle counter
(Bausch & Lomb)
recorder
dust
feeder
dust j .gas flow rate
conc.j i
& ^
(D indicater
inlet chamber
' *|air
test fabric
chamber
test section
filter *brifice
blower
manometer
Figure 9 Experimental apparatus for effect
of humidity on filter performance
I 1
CaC03 (D)
u= 2.7 m/t
felt fabric
(B9640SF
R.H. (*/•) -
~ 35 -40
A 55-60
O 80-85
50 100 150
dust load, m ( Vm2 )
Figure 10 Relation between
drags and dust loads for fine
calcium carbonate at various
relative humidity levels
c\E
E
o
~r
E
E
175
150
125
100
75
50
25
1—
i
T 1
Kanto loam (JIS No. 8) p
u s 2.7
'Vmin
/
felt fabric (B9640SF) O*
1
o o
/ 0
-
°7
R.H.(V.)
~ 25 - 40
A 55-63
O 83-92
- A
•
'
1 1
50 100
dust load
150 200
m (8/m2)
250
Figure 11 Relation between drags and
dust loads for Kanto loam (JIS No.8)
at various relative humidity levels
245
-------
50
fly ash (JI5 No.5)
felt fabric (B9640SF)
30
20
R.H. (•/.)
~ 30 - 40
A 55-70
O 81 - 96
10
0
0
50
100 150 200 250 300
dust load, m ( 3/m2)
Figure 12 Relation between
drags and dust loads for
fly ash (JIS No.5) at
various relative
humidity levels
*
N
'Si
S
o
R. H. (•/•)
38 ~
56 A
90 O
CaCOa(D)
before settling chamber
t»y cascade impactor
J I I—L
02 03 050.7 1 2 3 5 7 10 20
particle size, Dp (nm)
Figure 13 Particle size
distributions of fine calcium
carbonate before settling
chamber at various relative
humidity levels
R. H.(*/•)
D 39
A 54
O 84
02 03 050.7 1
Kanto loam (JIS No.8)
before settling chamber
by cascade impactor
J 1 l I I I
3 5 7 10 20
particle size, Dp (tim)
Figure 14 Particle size
distributions of Kanto loam
(JIS No.8) before settling
chamber at various relative
humidity levels
246
-------
ft
(A
I
99
R .H. (•/•)
37 ~
59 A
92 O
fly ash (JIS No. 5)
before settling chamber-
by cascade impactor
J—i—i i i i i i
05 050.7 1 2 3 5 7 10 20
particle size. Dp (fim)
I l—i—i
R.H. (•/•)
~ 35
A 66
O 86
a so
CaCOj (D)
after settling chamber
by cascade impactor
05 0507 1 2 3 5 7 X)
particle size, Dp (urn)
20
Figure 15 Particle size
distributions of fly ash
(JIS No.5) before settling
chamber at various relative
humidity levels
Figure 16 Particle size
distributions of fine calcium
carbonate after settling
chamber at various relative
humidity levels
-i 1—i—r
R.H. (*/•)
~ 28
A 63
O 92
Kanto loam-
(JIS No.8)-
after settling chamber _
by cascade impactor
» ¦ i I—i—i—
Q2a3 0.5 1 2 3 5 7 10 20
particle size, Dp (urn)
Figure 17 Particle size
distributions of Kanto loam
(JIS No.8) after settling
chamber at various relative
humidity levels
i—i—r
R.H. (•/•)
~ 32
A 67
O 93
fly ash (JIS No.5)
after settling chamber
by cascade impactor
i i ¦ ' '
0.3 0507 1 2 3 5 7 10 20
particle size, Dp(iim)
Figure 18 Particle size
distributions of fly ash
(JIS No.5) after settling
chamber at various relative
humidity levels
247
-------
St
o ^
^ UJ
Si
o J 0.5
if03
0.2
[He
R ,H.( */•)
35-40 ~
55-60 A
80-85 O
CaC03(D) Nvft
u=2.7 m/min «v*]~
felt fabric (B9640SF)
5 10 20 30 50 100 200
dust load, m ()
£
5.0
5s 2.0
? 1-0
8 0.5
¦S 0.3
R.H.C/.)
25-40 ~
55-63 A.
80"92 O
felt fabric
(B9640SF)
Kanto loam
-(JIS No. 8 )
u= 2.7 ^
3
u
10 20 30 50 100 200 300
dust load, m (^m2)
Figure 19 Relation between
cumulative penetrations
and dust loads for fine
calcium carbonate at
various relative humidity
levels
Figure 20 Relation between
cumulative penetrations
and dust loads for Kanto
loam (JIS No.8) at various
relative humidity levels
2x105i
o
BY
IS
i I
2
1
0.5
0.3
0.2
0.1
i r
felt fabric _
(B9640SF) ^
flyash(JIS No. 5)
u = 2
1—1 1 i
-I
RHC/.)
30 ~A0 a
50-70 A'
81-96 O
10 20 30 50 100 200
dust load , m (9*m2)
500
Figure 21 Relation between
cumulative penetrations
and dust loads for fly ash
(JIS No.5) at various
relative humidity levels
felt fabric
(B9640SF)
CaCO^D)
R.H.(•/•)"
23 ~
68 A
87 o
I | 1*10<-
20 40
dust load,m(9"m2)
Figure 22 Relation between
local penetrations and dust
loads for fine calcium
carbonate at various
relative humidity levels
248
-------
2*10®
_ felt
fabric
(B
Kanto
- loam
_(JIS Na0)
I i
20 40 GO
dust load, m( Vm3)
Figure 23 Relation between
local penetrations and dust
loads for Kanto loam (JIS
No.8) at various relative
humidity levels
R.H.C/.)
~ 39
& 61
O 83
felt fabric (B9640SF)
fly ash, u = 2.7%,n
100 125
dust load ,
150 175
Figure 24 Relation between local
penetrations and dust loads
for fly ash (JIS No.5) at various
relative humidity levels
249
-------
SI.
I * -20
at inlet chamber
203040 50 60 70 80 90
relative humidity of air, R.H.(*/»)
Figure 25 Relation between electrostatic
charge and relative humidity
0.8
o-_
0.6
CaCOa(D)
Kanto loam
(JI5 No. 8)
0.4
0.2
fly ash (JIS No. 5)
relative humidity, R.H. (*/•)
Figure 26 Relation between moisture
contents of test dusts and
relative humidity
250
-------
BAGHOUSE OPERATING EXPERIENCE ON
A NO. 6 OIL-FIRED BOILER.
BY:
David W. Rolschau, President
DaVair, Inc.
Minneapolis, Minn. 55405
ABSTRACT
In June 1977, the University of Illinois, Chicago, Illinois,
went operational on the nation's first high ratio, reverse gas,
glass bag baghouse operating on a No. 6 oil-fired HTHW boiler.
This paper discusses factors influencing baghouse design and
selection, capacity control, furnace pressure stability, decision
to eliminate pre-coating, cold start procedures, operating
experience, and maintenance costs.
251
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BAGHOUSE OPERATING EXPERIENCE ON
A NO. 6 OIL-FIRED BOILER.
PROBLEM
The University of Illinois, Circle Campus, downtown Chicago,
Illinois, operates four 75 MBTU/hr, HTHW gas-oil fired boilers.
In 1975, when gas was in short supply, the decision was made to
install a stack baghouse system so as to make possible the burning
of low sulphur No. 6 oil in a residential setting. The boilers
were currently operating within code limits except during soot
blowing, which occurred once every fourth hour at full load.
SOLUTION.
The decision was made to manifold a single baghouse system,
sized for 75 MBTU/hr gas capacity (35,000 ACFM @ 450° F), to all
four stacks and to valve the system as shown in Figure 1, such that
the baghouse system would operate continuously on one of the four
stacks and would be manually alternated between stacks in order to
clean full stack capacity during soot blowing. The baghouse, there-
fore, was to handle approximately four times normal particulate
loading.
Boiler
/
/
Diff.
Temp.
Sensors
/
/
/
B H
a
B H ID
Fan
Figure 1
BOILER - BAGHOUSE SCHEMATIC
252
-------
No other baghouse system had ever been put in operation in
the country on No. 6 oil-fired boiler service.
A baghouse capacity control system was needed to match the
baghouse to the stack being cleaned. A baghouse gas dilution
capacity control system was designed to couple the baghouse to
any one of four different capacity stacks. The system matched
boiler and baghouse gas capacity by mixing stack gas with 5 per
cent atmospheric stack downflow air for dilution cooling, and
sensed the resulting difference between stack and mixed gas
temperature. The baghouse inlet temperature sensor was inter-
locked with one of the four stack temperature sensors and with
one of the four inlet manifold valves to provide automatic bag-
house volume control by regulating the I D fan to maintain stack
temperature 25° F higher than mixed gas temperature.
Since each HTHW boiler had an operating turn-down ratio of
10:1, a DC motor driver was specified for the baghouse I D fan.
The system specification, prepared by Murphy Engineering of
Chicago, Illinois, addressed the unknowns of oil-fired operation
by requiring performance guarantees with substantial penalties
in the following areas:
1. Gas effluent quality better than 0.015 Gr/ACF.
2. Noise level 65 dbA or less at customer's property line.
3. Abrasive wear for one year.
4. Labor and material cost for bag failure and/or excessive
pressure loss at rated capacity, for one year.
5. Assigned responsibility for selection and sizing of the
baghouse I D fan and D C motor driver to the baghouse
manufacturer.
The bid by Independent Mechanical Industries, Chicago, 111.,
using a DaVair DCR 200-15 baghouse at 4.5:1 cloth ratio was
accepted based on lowest first cost, compliance with all perform-
ance guarantees, and technical confidence in the unique bag clean-
ing features of the patented reverse gas design. A diagram of a
DaVair DC Series baghouse is shown in Figure 2.
253
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Figure 2.
DAVAIR DC BAGHOUSE.
Each DaVair DC Series baghouse compartment is a totally self
contained unit, complete with its own head mounted reverse gas
fan and automatic bag cleaning system. The highly sectorized
design is particulary well suited for small boiler service because
the pressure drag variations, which occur as all baghouses cycle
into and out of cleaning mode, are so small that furnace pressure
normally varies less than 0.1 in. W C from set point.
254
-------
The DC Series design creates downflow of the dirty gas stream.
Particulate removed from the bag during cleaning cycle is trans-
ported to the hopper by this downflow. Baghouse performance is
not dependent upon the ability of the particle to agglomerate to
a larger size, nor is performance adversely affected by the
increased gas viscosity of high temperature operation. Refer to
Figure 3, Stokes Law, terminal velocity of ash particles in flue
gas.
600
500
400
£
1
> 300
To
c
0}
H
200
100
400
200
300
100
500
0
Particle Size, Microns
Figure 3.
TERMINAL VELOCITY OF PARTICLES IN FLUE GAS.
255
-------
The DC Series baghouse operates with a design bag cleaning
velocity of 8 FPM. This velocity, which is approximately four
times higher than found in conventional reverse gas designs,
allows operation with heavier weight fiberglass bags to improve
bag life, and assures adequate cleaning energy availability for
maintaining cake porosity with sticky particulate, such as is
typically found in oil-fired service.
OPERATING EXPERIENCE.
Pre-Coating.
Based on our experience with the DC Series high energy clean-
ing system, we recommended that no bag pre-coating system be
installed or used. (A pre-coat is normally recommended if the
baghouse manufacturer anticipates that particulate stickyness will
exceed the capability of the bag cleaning system). That decision
proved to be correct. All four boilers have been routinely
brought on line from cold start-up many times since the system
went into operation in June 1977. No bag blinding has occurred.
Operation Below Dew Point.
During the first five months of operation, the baghouse ran
continuously on one boiler, firing at ten per cent load. Baghouse
outlet temperature never exceeded 240° F, which is substantially
below acid dew point on 0.9 per cent sulphur oil. No adverse bag
effects were detectable, although considerable amounts of sulphur
salts condensed on the metal surfaces of the baghouse clean gas
plenum.
Bag Life.
The first set of 200 fiberglass bags was destroyed by fire
aftdr nine months of continuous operation. No bag failures had
occurred within that time; in fact the bag section had not been
opened for any reason at any time prior to the fire.
The baghouse was not used during the 1978-79 season. No
additional data on bag life is available from this installation.
The bags have performed as guaranteed in the specification.
Pressure Drag.
During initial operation, a partial failure occurred in the
DaVair designed air cylinders which operate the cleaning system.
Sufficient reserve energy existed in the cleaning system, such
that the boiler-baghouse system operated at full load for three
winter months within the capability of the I D fan.
256
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Normal pressure drag is 4 in. W C, flange to flange, during
cleaning cycle, at 4.5:1 cloth ratio (35,000 ACFM). Drag stabil-
ity is 0.4 in W C.
Baghouse Capacity Control.
Heat radiation losses from the insulated baghouse and
breeching system prevent the automatic baghouse capacity control
system from functioning at boiler capacities below 40 per cent
load. Manual speed control of the baghouse I D fan is necessary
at the lower load conditions. An explanation of the situation
is shown graphically in Figure 4.
400
350
F.
300
250
Boiler Dis
Eme &
Radiation
Dilution
Automatic
Mcinual
50 75
% Boiler Load
100
Figure 4.
BAGHOUSE CAPACITY CONTROL.
257
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The differential temperature between boiler and baghouse
depends upon both heat radiation losses from the ductwork and
dilution mixing of the stack gas with atmospheric air. If the
stack gas is cooled 10° F by radiation loss at full load, that
cooling will be 20° F at half load, 40° F at one quarter load,
and 80° F at one eighth load. When the radiation loss curve
crosses line B, the differential temperature sensor, sensing cold
baghouse inlet, shuts down the I D fan. Switching the I D fan
control to manual mode at low load allows continued operation of
the baghouse, but increases the possibility of high oxygen content
in the stack gas and the risk of a baghouse fire.
Part Load Radiation Losses.
The specific value of the radiation loss curve in Figure 4
is that it graphically depicts the problem of maintaining baghouse
temperatures above acid dew point on installations where the bag-
house is subject to high turn down ratios, such as occurs when
more than one boiler is breeched into a single baghouse system.
Baghouse Fire.
On March 31, 1978, a baghouse fire occurred which consumed
the bags only. The baghouse structure survived, intact, without
deformation of the tube sheet. The baghouse can be put into
operation at any time, and is budgeted to go into service this
year, but increased availability of natural gas in the Chicago
area may cause postponement of that decision.
The fire occurred during an aborted cold start on a boiler
that was being brought on line after extended maintenance. The
baghouse was hot at the time, and had a larger than normal
quantity of high carbon ash stored in its hopper. A higher than
normal oxygen content exists in the gas stream during start-up,
and this also increased the risk of fire.
Costs.
The 1976 installed cost of the particulate control system
described was approximately $12. per ACFM.
Operating costs during the first nine months of operation
consisted of power costs for operating the I D and reverse gas
fans, plus ash disposal costs. Bag and equipment maintenance
costs were zero.
Probable bag life, based on experience, should exceed two
years.
258
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Conclusion.
The DaVair DC Series baghouse has demonstrated the ability
to provide compliance effluent quality at 4.5:1 cloth ratio,
without pre-coating, withlow pressure drag, good drag stability,
and good bag life, on No. 6 oil-fired boiler service.
Figure 5.
OIL-FIRED BAGHOUSE.
259
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NEW FABRIC FILTER CONCEPT PROVEN MORE
FLEXIBLE IN DESIGN, EASIER TO MAINTAIN
AND UNSURPASSED FILTRATION
BY:
Bengt Carlsson
Bahco Systems, Inc.
Atlanta, Georgia 30340
Robert J. Labbe
Bahco Systems, Inc.
Atlanta, Georgia 30340
-------
NEW FABRIC FILTER CONCEPT PROVEN MORE
FLEXIBLE IN DESIGN, EASIER TO MAINTAIN,
AND UNSURPASSED FILTRATION
The United States is well recognized for its advanced technology
in many fields. There is an old American saying, "necessity is the
mother of invention". Both the United States and Europe share the
necessity of meeting government regulations in air pollution. However,
European governments' regulations seem to be more rigid. I think this
is largely due to the rigorous enforcement imposed by the pollution control
agencies.
When industry is forced to make an unprofitable capital investment
like a dust collector, (I base this statement on general calculations
for return on investments), the total investment, i.e. capital investment
plus maintenance expense must be compared for the available alternatives.
If, in addition to the economical consideration, we evaluate the
difficulty and willingness of the worker to maintain a baghouse from the
inside, features of any alternative system become highly significant.
I want to quote something that I read in the Fabric Filter Newsletter
- "If the dilute emission of flyash (which has been identified as a suspected
carcinogen) to the atmosphere is considered harmful to the general public,
then why would a corporation allow their men inside a confined space in a
concentrated atmosphere to change out bags?"
Outside maintenance and compactness have been achieved by almost all
manufacturers through the use of compressed air cleaning. This method of
cleaning is a United States designed system with a proven success record
established in the last several years. What is Europe able to contribute
to fabric filter technology? Is it possible to achieve or even increase
capacity of a fabric filter utilizing the compact design with outside
maintenance? Questions arise as to: the life of the fabric with high air
to cloth ratios? Bag blinding as a problem? The following should give some
answers to these problems:
A very important consideration in the design of the fabric filter is
the gas inlet. (Figure 1). A well designed gas inlet will prevent abrasion
caused by coarse dust impinging on the fabric media. It is established and
proven that using "off line" cleaning allows us to use higher air to cloth
ratios with a minimum of risk of blinding, and improved fabric life
expectancy. For several reasons, European manufacturers of fabric filters
have tried several configurations other than the conventional bag. For
example, the envelope, pocket and cassette. A description of the conven-
tionally used bag filter cleaning technique is required to demonstrate the
reason for the above referenced different configurations.
261
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Clean, dry and oil free compressed air is introduced through a header
located in the clean air plenum of the baghouse. (Figure 2). Because
this header is located in the clean gas plenum, secondary air is induced
into the venturi by the high pressure compressed air jet. This high
velocity air moves down the bag in a wave fashion causing the fabric to
expand extremely rapidly. (Figure 3). This very rapid expansion of the
fabric will cause the dust cake to penetrate into the fabric. The
major portion of the dust cake is shocked loose. (Figure 4).
When the fabric has reached the most expansive point, it then returns
striking against the cage like a "whip lash". The dust penetrating the
fabric is discharged into the clean air side of the bag during this "whip
lash" phenomenon. The fact that the fabric is flexing from a maximum to a
minimum with dust penetrating into the fibers means that the fibers are
mechanically abraded by this penetrated dust. We are all aware that
a dirty suit will not last as long as a clean suit.
This fabric dust penetration phenomenon was discovered after several
laboratory tests while flexing a conventional bag. The same tests were
performed on the rigidly designed cassette where minimum flexing was
experienced and dust penetration was substantially reduced. (Figure 5).
The emission after the bag was ten times greater than the emission measured
after the cassette.
Employing the fabric filter as the dust collector on a coal fired
boiler has become more and more common today, due to the fact that a very
high efficiency is required and capable of being achieved - 99.9 plus
per cent. This alternative, when considering efficiency performances on
western low sulphur coal, is usually more desirable than the electrostatic
precipitator option.
Recognizing that high efficiency and continual cleaning performance is
achievable on the fabric filter, the key question left to be answered is:
what type of fabric has to be used? Several test programs have been
performed with all kinds of experts involved. An annotated summary of the
state of the art findings might look something like this ...
WOVEN GLASS, WITH DIFFERENT COATINGS;
Plus
Minus
High temperature resistance
The fabric is rather fragile
and has low mechanical
abrasion resistance
Good acid resistance
Long time experience from different
installations available
Limited capacity per square
foot which makes it difficult
to be used in higher air to
cloth ratio units.
262
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FELTED GLASS:
Plus
High temperature resistance
Good acid resistance
Good capacity and able to be used
on rather high air to cloth ratio
units
Minus
Fragile material with low
mechanical abrasion resistance
Limited experience from
installations available
TEFLON FELTED:
Plus
High capacity
High temperature resistance
High acid resistance
Experience data available to
some extent
Minus
High initial cost
Limited mechanical abrasion
resistance
CHEMICALLY TREATED NOMEX:
Plus Minus
High capacity Limited experience data
available
Good mechanical abrasion resistance
Temperature limitation to
Acid resistance 400 degrees Fahrenheit
With the continued rapid development of fabrics suitable for filtering
fly ash generated by coal fired boilers, the resultant increased number
of fabric filter installations shall create a focal point on maintenance
and operation. The older technology low ratio shaker and reverse air
fabric collectors, which have been recently, and are presently being
installed on coal fired boilers, are very expensive to operate and maintain.
Even though these low ratio collectors have a high capital cost and
operating cost, the reliability and long term experience is the basis of
their present usage.
263
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The high ratio compact cell filter design, allowing for complete
external maintenance, will definitely have an impact on fabric filter systems
for coal fired boilers.
Because of the more stringent pollution control regulations, the low
maintenance mechanical dust collectors no longer suffice as a viable control
alternative. Also, electrostatic precipitators that must operate on low
sulphur coals at very high efficiencies become economically not feasible in
many cases. The fabric filter alternative becomes virtually the only viable
solution. With this thought in mind, it becomes an essential mission of both
the consulting engineer and the manufacturer to inform the user of the
importance and cost of maintenance and operating procedures for the various
fabric filter alternatives.
264
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Filtration
Figure
Cleaning
-------
y
V'L. - "«^T ^
Figure 2
-------
Figure 3
FILTRATION AIR PULSE WHIPLASH
¦'¦' ¦ ;¦$'•.' ¦•¦•'; - ¦; -i* >;'' y/.:;A :-y£^
-¦;>:¦ .v.;;.;.: -.'.Y' -v.;
-------
FILTRATION
AIR PULSE
ro
a-
oo
Asa-
•, i » v* •
Figure 4
WHIPLASH
clean
side
-------
Filtration
Figure 5
Cleaning
Compact Cell
-------
EPRl's FABRIC FILTER TEST MODULE PROGRAM:
A REVIEW AND PROGRESS REPORT
R.C. Carr
Electric Power Research Institute
Palo Alto, California 94301
John Ebrey
Lodge-Cottrell Operations/Dresser Industries, Inc.
Houston, Texas 77005
ABSTRACT
The use of fabric filters (baghouses) for control of particulate emissions
from pulverized-coal fired utility boilers is a young but rapidly evolving tech-
nology. To insure that the electric utility industry has accurate design and
operational information available, the Electric Power Research Institute (EPRI)
has initiated a major fabric filter pilot plant research program. The program
consists of the design, fabrication, installation and testing of a 850 m3/min
four compartment fabric filter pilot plant at the EPRI Advanced Particulate
Control Test Facility in Denver, Colorado. The pilot plant design reflects the
range of commercially available technology offered by the major equipment vend-
ors. An extensive series of tests will be performed to quantify and optimize
the effects of air-to-cloth ratio, cleaning techniques, fabric type, bag geo-
metry and gas flow distribution on operating pressure drop, reliability, stack
opacity and total/fine particulate control. This paper provides a progress re-
port on the project including objectives, schedule, fabric filter hardware spe-
cifics and proposed test plan.
270
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INTRODUCTION
A key environmental problem facing the electric utility industry is
the increased emphasis by regulatory agencies on application of high effi-
ciency particulate control devices to pulverized-coal fired boilers. This
increased emphasis is-manifested by the revised New Source Performance Stand-
ard of 12.89 grams/10 Joules (0.03 lb./lO BTU) recently promulgated by EPA.
In addition, considerable attention is being focused on control of fine par-
ticulates (less than 2 micron diameter), trace element emissions and plume
opacity.
In response to these constraints the utility industry, through the
Electric Power Research Institute (EPRI), is conducting a major research pro-
gram to improve existing technologies and develop cost effective alternatives
which promise high efficiency particulate control. One such research effort
defining the optimum design and operating parameters of fabric filters (bag-
houses) is described in this paper.
Although baghouses have been in commercial use for many years, their ap-
plication to the electric utility industry is a relatively new technology.
For example, the first units to come on line were the Sunbury and Nucla station
baghouses in 1973 and 1974, respectively. The Sunbury station, owned and oper-
ated by Pennsylvania Power and Light Company, consists of four baghouses in-
stalled on four nominally 44 Mw p-c boilers firing an anthracite/petroleum coke/
bituminous coal mix. The Nucla station, owned and operated by Colorado - Ute
Electrical Association, consists of three baghouses installed on three 13 Mw
spreader-stoker boilers firing a submituminous coal. Although both install- ^
ations have reported very successful baghouse operation, [Ensor Et. Al. (1976)
Wagner, (1978) ] they are not considered typical of utility boilers due to their
small size, firing method and/or type of coal fired. More recently, baghouses
have been applied to larger more typical pulverized - coal fired utility boilers
ranging in size from £5-550 Mw. Preliminary results reported [Beaton (1978) ,
Faulkner, Ladd (1978) ] for many of these installations have been mixed, i.e.
some units are experiencing excessively high pressure drops (up to 250 mm)
whereas other exhibit more acceptable pressure drop characteristics (50 -
100 mm).
The specific reasons behind the anomalous baghouse performance observed
to date are unclear, although there are several possible explanations which
immediately come to mind: much of the initial experience was gained on atyp-
ical units which may not be applicable to more conventional utility boilers;
several vendors have entered the market offering baghouses with wide ranges
in hardware design and recommended operating philosophies; an inability to
measure or control critical parameters; underestimation of the importance of
gas flow distribution; laboratory performance data obtained with resuspended
flyash which may have led to erroneous scale-up conclusions due to subtle
changes in particle size distribution and subsequent filter cake dynamics; and
a lack of comprehensive pilot scale research on utility boiler flue gas streams.
Consideration of the current situation leads to the conclusion that there
is an urgent need to provide the electric utility industry reliable design and
operating baghouse specifications. Accordingly, EPRI has initiated a baghouse
271
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pilot plant program which is governed by the following constraints: pilot plant
of sufficient scale to properly simulate a full scale baghouse; the use of fly
ash and flue gas from a pulverized-coal fired utility boiler; hardware which per-
mits a complete and systematic evaluation of all the major design and operating
parameters without restriction to a particular manufacturer's design; and well
controlled and characterized operating conditions. Following this line of rea-
soning, EPRI has contracted with Lodge-Cottrell/Dresser Industries, Inc. to
design fabricate and assist in testing a 850 m^/min., four compartment baghouse
pilot plant which is currently being installed at the EPRI Advanced Emissions
Control Test Facility. The pilot plant will form the basis for an extensive
series of tests with the ultimate objective to define the optimum design and
operating criteria as well as identify cost effective improvements for fabric
filters applied to pulverized - coal fired utility boilers.
SPECIFIC OBJECTIVES
The electric utility industry have currently installed or committed to
over 5000 Mw of baghouses on pulverized - coal fired utility boilers. Based
upon the range of design and operating criteria evident in the baghouses pur-
chased to date (e.g. air-to-cloth ratios from 0.5 - 1.0 m/min), it appears
that choices are being made by utilities with less than complete information
at their disposal. This situation is likely in many cases to lead to higher
capital and operating costs than necessary without any assurance of satisfying
regulatory constraints. Accordingly, it is the intent of this project to pro-
duce results which will remove the intuitive aspects from baghouse design and
quantify the design and operational factors necessary for optimum performance.
It is expected that this characterization program will enable the utility in-
dustry to prepare optimized baghouse specifications for both retrofit and new
coal-fired boilers. Specific areas of interest are:
Establish criteria for proper gas flow distribution in inlet/outlet
manifolds, hoppers and cell plates.
0 Define the most effective method of fabric cleaning.
Establish the dependence of pressure drop and particulate penetration
on air-to-cloth ratio, time between cleaning and inlet concentration.
° Determine the importance of bag aspect ratio (*7d)
° Evaluate different levels of compartmentalization
Examine the effect of filter cake particle size distribution on
pressure drop and particulate penetration.
° Determine optimum start-up and shut down techniques.
Identify improved fabrics or methods of cleaning which require less
off-line cleaning time.
272
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TEST FACILITY DESCRIPTION
The EPRI Advanced Emissions Control facility is a pilot plant test fa-
cility designed to evaluate advanced concepts in both particulate and gaseous
emission control for pulverized-coal fired utility boilers. The facility is
located on the 110MW unit 4 at the Arapahoe Station of Public Service Company
of Colorado and consists of five flue gas slip streams ranging in size from
142-991 m^/min. (1-10 MW) as shown in figure 1. The slip streams are directed
to five pilot plants which are involved in a wide variety of testing. Unit 1 is
a nominal 142 nr*/min. pilot plant electrostatic precipitator used primarily as a
technology screening tool. Unit 2 is a nominal 991 m-^/min. precipitator which is
used for prototype evaluation of promising design and operating improvements.
Unit 3 is a nominal 283 m^/min. pilot plant which is used to evaluate more novel
concepts such as granular bed filtration and electrically augmented fabric
filters. Unit 5, the unit of interest here, is a nominal 850 m-^/min. fabric
filter pilot plant described in more detail below. Unit 6 is a 283 m^/min. pilot
plant designed to assess the impact of integrating the various emission control
technologies currently available into a reliable, cost effective system.
UNIT 5 FABRIC FILTER PILOT PLANT
The design of the fabric filter pilot plant (FFPP) presented some inter-
esting and formidable problems since it was decided at the outset that the
most meaningful results would be obtained from a flexible research facility
which also retained the representative character of current commercial utility
baghouses. After evaluating several configurations, the general arrangement
shown in figures 2, 3 and 4 was selected. This particular design consists of a
four compartment, hopper inlet, inside bag collector similar to existing util-
ity baghouses. However, there the similarity ends as described below:
General Specifications
To maximize the flexibility of the FFPP, each of the four compartments
is capable of independent operation and control, which explains the intricate
system of ducts shown in figures 2, 3 and 4. The ducting arrangement shown for
each compartment includes a dirty gas inlet, compartment bypass, reverse gas
cleaning, clean air preheat and clean outlet gas ducts, respectively. Refer
to figure 2 showing the process flow control diagram.
An important feature of this arrangement is the bypass duct, which main-
tains constant volume flow conditions through the on line compartments when a
fourth compartment is removed from service. Accordingly, the FFPP can be operat-
ed as one four compartment baghouses or four-one compartment baghouses depending
upon the particular situation. Note that flow control and measurement of the
inlet (dirty) gas is accomplished by calibrated venturi flow meters rather
than orifice plates to minimize the potential for dust fouling and flow dis-
turbances. Although requiring a more elaborate duct arrangement, the Venturis
were placed on the inlet rather than the outlet to permit flow control during
compartment bypass situations, which by definition eliminates flow in the
outlet duct. All flow control dampers depicted in figures 2, 3 and 4 are cast
iron wafer type round dampers designed for a maximum leakage rate of below 1%.
273
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Table 1 summarizes the major design specifications for FFPP. Note that
sufficient lattitude is provided to evaluate the wide range in hardware design
and operating philosophy currently available in commercial baghouses. Before
proceeding further, a comment on the range in air-to-cloth ratio (A/C) reported
in Table 1 is appropriate. With all full size bags ( 304 mm diameter x 10.4 m
long) and compartments in service the maximum A/C ratio possible is 0.6 m/min
with the existing booster fan. However, higher A/C ratios will be achieved by
capping an appropriate number of bag thimbles in each compartment, thus driving
the gas through the remaining elements at A/C ratios as high as 1.2 m/min.
Similar levels of A/C ratio can also be achieved by operating with one or two
compartments out of service. A combination of techniques will permit testing
at even higher A/C ratios, if desired. In any event, the fan and cleaning
system are sized to adequately handle operational excursions of this nature.
It should also be noted in table 1 that inlet particulate concentrations
ranging from 0.5 - 16 grams/m^ are possible. The wide range in inlet concen-
tration is accomplished by use of either a multiclone or flyash re-injection
system incorporated into the FFPP. (The particulate concentration for the
Arapahoe Boiler averages roughly 2.3 - 4.6 gram/m^). In addition, provisions
have been made for the injection of water, S0X compounds, and dry alkali reagents
for possible future evaluation of dry SO2 removal technology, operation of bag-
houses downstream of wet SO2 scrubbers and/or high-sulfur coal baghouse appli-
cations.
TABLE 1
Minimum Maximum
Flow, m3/min 425 850
Temperature, °C 93 177
Dust Loading gram/nP 0.5 16
Flue gas moisture, % vol. 7.6 15
Static Pressure at inlet, mm H2O -380 Ambient
Air-To-Cloth Ratio, m3/m2/min. .3 1.2
No. compartments - 4
Bags compartment - 36
Bag Length, m. 6.7 10.4
Bag Diameter, mm. 150 304
Cleaning Shaker, reverse air, or
shaker/reverse air
274
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Filter Bag Types And Dimensions
Each module of the FFPP is approximately 3 meters square by 13.7 meters high
and will contain 36 bags in a 6x6 array, each on square 355 mm centers for
testing nominal 304 mm diameter bags. The entire bag support assembly is ad-
justable, thus permitting evaluation of 6.7, 7.9, 9.1 and 10.4 meter long bags.
In addition, extra cell plates are available having appropriately sized thimbles
to permit evaluation of 150 and 203 mm diameter bags. These additional cell
plates retain the 355 mm center spacing. It is believed this combination of
bag lengths and diameters will allow a quantitative determination of the effect
of bag aspect (*7D) ratio on particulate penetration and pressure drop.
For ease of substitution, the dust hopper under each module is supported
independent of the housing and fitted with an 0.5 meter high spool piece to the
underside of the cell plate. To test different diameter bags, the spool piece
will be unbolted, top and bottom, and slid out to one side. The existing cell
plate may then be lowered, thus providing clearance above the thimble tops, and
exchanged for a cell plate serving bags of the desired diameter.
The initial bags to be installed in the FFPP will be 10.4 meter long, 3x1
twill, silicone-graphite - teflon (Tri Coat) coated fiberglass bags manufactured
by Globe Albany, Inc. Follow-on testing will include other combinations of bag
size, coating and weights.
Bag Cleaning
Since effective bag cleaning is an essential ingredient to efficient bag-
house operation, unusual flexibility has been incorporated into the FFPP to per-
mit complete characterization of this important aspect of the technology. A
variable reverse flow system is provided at rates ranging from 0.18 to 1.8 m/min.
using filtered flue gas. Shut-off, modulating and by-pass dampers along with the
associated ductwork are included to insure that the operating conditions
in the three on-line compartments are not affected during cleaning (see figures
2, 3 and 4).
In addition, a novel shaking mechanism will be installed. This mechanism
can be adjusted in finite increments to provide predominantly vertical or hor-
izontal movements, the maximum for either movement being 476 mm with nu-
merous intermediate combinations. As delivered, this mechanism will be adjust-
ed to provide natural spring-loaded, dampened frequency oscillation of the bags
after snap release. The frequency of activation can be varied from 2 to 40
cycles per minute during this mode of shaker cleaning, as required. To evaluate
the more conventional method of shaking used in several of the currently oper-
ating utility baghouses, the shaker can be easily modified to reinforced oscil-
lation with a frequency range of 10-200 cycles per minute.
Finally, combinations of reverse-air and shaker are possible, since the
systems are fully independent.
275
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Instrumentation And Controls
The prerequisite of operating four independent baghouse compartments
has dictated the use of monitoring, control logic and supervisory equip-
ment far exceeding ordinary commercial requirements, Combined with the
necessity of obtaining data on numerous operating variables from multiple
stations simultaneously, these aspects of the program were seen as a limiting
constraint on both the accuracy and time required to collect data if the
FFPP were operated and tested in a conventional manner. Fortunately, re-
liable micro-processor hardware and technology is now available which makes
management of such complex situations both cost effective and efficient.
The micro-processor utilized has the capability of performing ladder log-
ic, mathematical functions and proportional-integral-differential loop con-
trol. This system provides for 224 discrete inputs/outputs, 12 analog in-
puts and 16 analog outputs. If desired, this system may be expanded at the
job site.
To examine, control and record the baghouse operating variables the fol-
lowing major instrumentation and control systems are provided:
Qty. (5) - Flow control systems consisting of
Venturi metering tube
Differential pressure transducers for input into micro-processor
Thermocouple input into micro-processor
(A) - Cell Plate - differential pressure display
<1) - System differential pressure display
:6) - Temperature displays
'1) - Reverse air control system
[4) - Recorder channels for compartment gas volume
(1) - Recorder channels for reverse air gas flow
[50)- Mode selection switches for shake and reverse air cycles
(8) - Event recorder channels for shake and reverse air system in each
compartment,
[4) - Manometers
[1) - Air pre-heater system
'1) - Graphics display panel
[4) - Variable speed shaker motor systems
276
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TEST PROGRAM
The FFPP will be operated and characterization tests performed during an
initial fifteen month test period. The major test variables include types of
bags and fabrics, cleaning mechanism and duty cycle, gas volumetric flow rate
(A/C ratio), inlet/outlet concentration and pressure drop as summarized in
table 2. The test program will be working under the hypothesis that for a
given fabric cleaning technique, an optimum combination of A/C ratio and clean-
ing cycle exists which yields a minimum pressure drop for the required level of
outlet concentration. In addition, a second premise is that with proper char-
acterization of the FFPP the results will be of direct application to full
scale utility units constructed by different manufacturers and handling various
types of flyash.
Although not yet finalized the FFPP test plan is currently divided into
four basic sequences: 1) Air-Load testing; 2) Start-up and conditioning
tests; 3) First year characterization tests; and 4) Long term testing of
more advanced concepts.
This comprehensive test effort will remain flexible to respond to results
as they become available.
1. Air Load Testing
Prior to start-up of the FFPP with flue gas, an air-load will be performed
to insure that all systems are functional. This will include a verification
that all fans, dampers, motors and controls operate properly; leak testing
of each module; calibration of all transducers and instrumentation; and de-
termination of clean bag weights. In addition, the pressure drop across
each cell plate and the entire FFPP will be measured as a function of vol-
umetric flow both with and without bags. And finally, the velocity profile
across the bottom of the cell plate (bag inlets) and in the inlet duct work
will be measured as a function of volumetric flow rate to verify proper
flow distribution. It is appropriate to mention here that 1/4 scale lab-
oratory gas flow distribution tests are being conducted at the Lodge-
Cottrell Model Test Facility, to ensure that the capability is available
to assess the impact of gas flow mal-distribution on baghouse performance.
2. Start Up And Conditioning Tests
One aspect of utility baghouse operation which is currently in a state of
uncertainty is the optimum method of start-up, particularly with respect
to the need and appropriate procedures for pre-coating of the bags. To
address this issue, the FFPP start-up testing will include a bag pre-coat
with resuspended flyash on at least two of the four compartments. The
FFPP will then be operated with flue gas on a continuous basis to condition
the bags. The inlet and outlet streams will be monitored continuously dur-
ing this period for pressure drop, volumetric flow, and opacity. The clean-
ing technique to be used during start—up has yet to be selected, although a
reasonable approach may be to use shaker/reverse air assist on two compart-
ments and reverse air only on the remaining compartments. A major object-
ive of this phase of the test is to determine the long range effects of the
277
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start up procedure, i.e. will the differences in method of start-up be de-
tectable when all compartments eventually stabilize, or will the effect of
a lack of pre-coat at the outset persist in the form of high pressure drop
well after start-up.
3. First Year Characterization Test
This series of tests will be directed toward quantifying the effects of
cleaning cycle parameters, bag conditioning, bag aspect ratio (L/D), A/C
ratio and inlet particulate concentration and size distribution on operating
pressure drop and outlet concentration. As outlined earlier the primary
objectives are to establish the following:
0 The effectiveness of shaker, reverse air or shaker/reverse air combin-
ation cleaning.
0 The optimum cleaning cycle for each method of cleaning (i.e. shaker
frequency, shaker direction of motion, reverse air gas volume, times
for dwell, null, overlap, rate of compartment re-pressurization, etc).
0 Relationship between particulate penetration and the fundamental con-
trolling parameter of bag ash loading (product of inlet concentration,
face velocity (A/C) and time between cleaning).
° Relationship between pressure drop across the bags and bag ash loading.
° Significance of bag aspect ratio (^/D).
° Importance of inlet velocity distribution on pressure drop, particulate
penetration, and cleaning effectiveness.
° Identify the most successful fabrics from the standpoint of minimum
pressure drop and filter cake release.
0 Effect of inlet concentration on pressure drop, bag cleaning and partic-
ulate penetration.
The importance of items 3 and 4 (above) cannot be over-emphasized. Although
the relationship of particulate penetration and pressure drop has in the
past been correlated with A/C ratio, recent work by [Drehsen and Ensor
(1979)6, Dennis and Surprenant (1978)^] has shown that the fundamental
controlling parameter is bag ash loading. This parameter includes the
effects of cleaning cycle time, inlet concentration and A/C ratio and
more properly characterizes baghouse operation. Accordingly, it is
believed that establishment of quantitative pressure drop and particulate
penetration versus bag ash loading performance curves will provide the
necessary information for utilities and manufacturers alike to utilize
in specifications of baghouse design for optimum performance.
278
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Long Term Testing
The specific nature of these tests will be dependent upon the results of
previous testing of the FFPP as well as new developments in baghouse technology.
However, some general items of interest include evaluation of new types of fab-
rics, dry SC>2 removal technology, evaluation of the role of electrostatics, and
advanced bag cleaning concepts which require less off-line cleaning time.
Test Schedule
Air load testing is due to start in November 1979, with start-up and con-
ditioning tests commencing during late 1979 early 1980. It is anticipated that
characterization tests will be conducted in 1980. However the ultimate test
time will be highly dependent upon results obtained. Schedules for the long
term testing program will be developed at a later date. In order to expedite
the transfer of information to the electric utilities intermediate reports
will be issued as meaningful information becomes available.
MEASUREMENT METHODS AND PHILOSOPHY
Fabric filters have several unique operational characteristics which re-
quire special consideration when conducting emission testing. First, the very
efficient collection capability places a heavy demand on the ability to measure
gas streams with a 10,000 fold difference in particulate concentration. Outlet
concentrations as low as 0.23 milligrams/cubic meter have been measured in the
field, thus requiring long sampling times with high flow rate test equipment
and very sensitive instrumentation. Also, complex data reduction schemes are
required to compute particle size dependent penetration under these conditions.
A second unique aspect of fabric filtration is that the behavior is simi-
lar to a "batch process" because of the need to periodically remove compartments
from service and clean the collected particulate matter from the fabric. Ac-
cordingly, the outlet emissions will vary significantly as compartments are re-
moved from service and cleaned, which dictates the need for careful testing and
characterization of the cleaning cycle.
A final important factor which must be considered is the long time con-
stants associated with baghouse operation. The response of a baghouse to op-
erational changes is not instantaneous and in some cases several days may
be required to attain steady state. The parameters requiring the longest time
to achieve steady state are those dependent upon the ash layer on the bags.
For example, it is not uncommon for new bags to require several weeks of con-
ditioning to reach a stabilized level of pressure drop and collection effi-
ciency. Other process changes such as variable volumetric flow or cleaning
cycles can also introduce significant time lags prior to achieving equilibrium.
The consequences of these generic characteristics of fabric filtration are
that the FFPP operating conditions must be carefully controlled, monitored and
characterized on a continuous basis. This will be accomplished by equipping
the FFPP with real time continuous aerosol instrumentation, gaseous species
monitors and pressure transducers to track the dynamic behavior of the FFPP. The
specific instrumentation and methodology used will be similar to that reported
for previous particulate control device testing supported by EPRI [Ensor et. al.
(1976), Drehsen and Ensor (1976)6, Ensor, et. al. (1979)8].
279
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The testing of the FFPP will include both routine day-to-day monitoring and
intensive test periods to examine in detail specific operating conditions which
require more extensive resources and concentrated effort. The routine monitoring
will be performed between the intensive test periods to screen various operating
conditions, determine when steady state conditions are achieved and evaluate the
operating trends of the FFPP over long time periods. Measurements will include
continuous recording of gaseous species and the important baghouse parameters of
cell plate and flange-to-flange pressure drop, flue gas and reverse air volumet-
ric flows, temperature, static pressure, fan motor currents, damper settings, and
cleaning cycle sequences. In addition continuous inlet and outlet opacity in-
strumentation will be used to detect operating problems and provide a qualitative
indication of aerosol penetration.
The intensive tests will focus primarily on the characterization test se-
quence described above. The measurements conducted will include inlet velocity
distributions, particulate mass penetration, size dependent penetration from
0.01-20 micron particle diameter, SO3, and size dependent trace element analysis.
Where possible information will be data logged to provide a common time base
which is useful in correlating parameters and detecting subleties in the data.
A summary of measurement techniques is provided in table 3.
SUMMARY
A 850 m^/min four compartment fabric filter pilot plant (FFPP) is currently being
installed at the EPRI Advanced Particulate Control Test Facility in Denver,
Colorado. The FFPP will be the subject of an extensive test series with object-
ives to quantify the design and operational factors necessary for reliable and
cost effective performance. Specific areas of interest include determination
of gas flow distribution effects; most effective fabric cleaning techniques and
cycle; relationship of pressure drop and particulate penetration on air-to-cloth
ratio, time between cleaning and inlet concentration; appropriate start-up and
shut-down techniques; effects of bag aspect ratio and fabric characteristics.
It is expected that the program will produce results to enable the utility
industry to prepare optimized baghouse specifications for both retrofit and
new coal-fired boilers.
280
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REFERENCES
1. Ensor, D.S., R.G. Hooper, and R.W. Scheck. Determination of the Fractional
Efficiency, Opacity Characteristics, Engineering and Economic Aspects of a
Fabric Filter Operating on a Utility Boiler. EPRI Report FP-297. November,
1976.
2. Wagner, N.H. Current Status of Bag Filters at Pennsylvania Power and Light
Company. Third International Fabric Alternatives Forum Proceedings. Sep-
tember, 1978. pp. 3-1 to 3-32.
3. Beaton, R.J. Continuing Operation and Maintenance Experience of the Four
Reverse Air Fabric Filters at the Kramer Station. Third International
Fabric Alternatives Forum Proceedings. September, 1978. pp. 1-1 to 1-20.
4. Faulkner, G.R. and K.L. Ladd, Jr. Start-up, Operation and Performance
Testing of Fabric Filter System at Harrington Station Unit Number 2. Sep-
temper, 1978. pp. 2-1 to 2-15.
5. Green, G. Personnel Communications. Public Service Company of Colorado.
Cameo Baghouse Start-up Performance. June, 1979.
6. Drehsen, M.E. and D.S. Ensor, Evaluation of the Kramer Baghouse - Phase
II B. MRI Report to EPRI. February, 1979.
7. Dennis, R., and N.F. Surprenant. Particulate Control Highlights. Research
on Fabric Filtration Technology. U.S. Environmental Protection Agency
Report No. EPA - 600/8-78-005d. June, 1978.
8. Ensor, D.S., et al. Evaluation of the George Neal Electrostatic Precip-
itator. EPRI Project RP 780-1, Final Report. Publication date July, 1979.
281
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Table 2
FFPP MAJOR TEST VARIABLES
Variable
Independent
Controllable
Fixed by
Uncontrolled Design Dependent
Type of baghouse
Bags
Area of bags
Cleaning equipment
Cleaning cycles
Fabric
Gas Volumetric
flow rate
Gas composition
Gas temperature
Concentration In
Size distribution In
Particle composition
Opacity In
Concentration Out
Size distribution out
Ash loading on bags
Pressure drop
Opacity Out
Collected Ash
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
282
-------
Table 3
MEASUREMENT SUMMARY
PARAMETER
TECHNIQUE
COMMENTS
1. 02
2. NOx
3. S02
4. CO. C02
5. N0/N02
6. C02• 02. CO
7. H20
8. Opacity
9. Total Particulate
Mass and Trace
Elements
C
C
Electrochemical
Electrochemical
Electrochemical
Nondispersive infrared detector
Chemiluminescence
Orsat
Important to monitor boiler operation
and dilution in duct
Important to monitor boiler operation
and dilution in duct
Important to interpret resistivity
measurements
Monitor boiler operation and dilution in
duct
Obtain average molecular weight of stack
gas
Impinger catch during Impactor Test
Plant Process Visiometer Light
Scattering
Inlet of baghouse to monitor boiler
operation
Outlet of baghouse to monitor real-time
variations in baghouse performance, i.e.
cleaning cycles, approach to equilibrium
leaks, etc.
Method 5 Train with volatile trace
element absorbing solutions
10. Velocity
S-Type Pitot Probe
Gas volumetric flow rate, flow distri-
bution
-------
Table 3 (Cont'd)
PARAMETER
TECHNIQUE
COMMENTS
11. Total Particulate
Mass
12. In situ Resistivity
13. S03
Thimble
Point to plane probe
Goksoyr/Ross controlled
condensation
Resistivity Measurements in laboratory
Inlet, Outlet
14. Particle Size
Distribution
0.5 to 20 um
Cascade impactor
In situ - Inlet-Outlet
15. Trace Element
0.5 to 20 um
16. Particle Size
Distribution
0.05 to 0.5 um
17. Particle Size
Distribution and
Trace Elements
0.05 to 20 um
18. Compartment
AP
Cascade Impactor with Kapton
Substrates, Ion-Induced X-Ray
Analysis, Neutron Activation
Analysis
Electrical Aerosol Size Analyzer
Low Pressure Cascade Impactor
Kapton Substrates, Ion-Induced
X-Ray and Neutron Activation
Analysis
Pressure Transducer
In situ - Inlet-Outlet
Aerosol removed from duct and diluted
in-stack (inlet) and out-of-stack at
outlet
Alternative size distribution and trace
element technique.
Inlet and outlet.
Real-Time continuous pressure drop
monitoring
NOTES: C - Continuous analyzer recorded on strip charts and a magnetic
tape logger.
Inlet/Outlet impactor measurements simultaneous
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Figure 3
Elevation.
287
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^ MODULES
Figure 4
Plan view inlet and outlet.
288
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ELECTROSTATIC ENHANCEMENT OF MOVING-BED GRANULAR FILTRATION
Dale S. Grace
John L. Guillory
Fernando M. Placer
Combustion Power Company, Inc.
Menlo Park, California
ABSTRACT
The electrostatic granular filter developed by Combustion Power
Company combines a moving-bed filter with an electrostatic grid for high-
efficiency collection of particulate. The coulomb charge occurring natur-
ally on particulate from combustion or other processes in which tribocharging
occurs is sufficient such that imposition of an electric field substantially
enhances mechanical collection in the packed filter, A mathematical model
is presented which accounts for the major parameters affecting electrostatic
performance. Experimental data developed in an industrial unit retrofitted
with electrostatic equipment and a pilot-scale filter demonstrates dramatic
improvement in opacity and outlet loading; the industrial unit achieved
99% collection efficiency with less that 4 Iwd pressure drop. Data from
the pilot-scale filter has shown that collection of micron-size particulate
is especially enhanced with the electric field and has confirmed the 99%+
overall collection capability demonstrated in the industrial unit. The
design of a 420,000 acfm system being installed at Weyerhaeuser Company's
Longview, Washington plant is also discussed.
INTRODUCTION
Combustion Power Company's first commerical moving-bed granular filter,
the Dry Scrubber^, was installed in 1974. Although granular filters have
been used for industrial processes since the early 1900's, they were
usually of the fixed bed type, requiring intermittent cleaning of the filter
media and often disrupting the process. The Dry Scrubber was first conceived
for use with CPC's process research on fluid bed combustors and later de-
veloped for industrial applications, in part as a response to new and more
stringent regulations on particulate emissions. Indeed, interest in the
289
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environment has become more intense in recent years, and the public has come
to realize that further degradation in air quality is unacceptable. Indus-
trial capacity continues to grow, albeit at a slower rate than the first
half of the decade, and more efficient, control of pollution is necessary.
Since society bears the cost of pollution in any case— whether damage to
public health in the community, workers' health in the plant, or damage to
property and aesthetic values—a high degree of control is mandated. It is
in this context that the moving-bed granular filter is used and finding
many applications. The granular filter concept was designed with these
objectives in mind: to meet specific particulate control performance while
minimizing total annual cost. Design objectives include: moderately low
pressure drop, high volume flow, low operating and maintenance costs, and
dry operation. A filter utilizing both mechanical and electrostatic
mechanisms, the ElectroscrubberTM, was developed as an extension of this
concept.
DESCRIPTION
Figure 1 shows the general schematic of the Electroscrubber granular
filter. The system consists of a cylindrical vessel containing two con-
centric, cylindrical louvered screens. The annulus between the screens is
filled with 1/8 x 1/4 (pea sized) gravel media. Particulate-laden gas
enters the filter and travels through the inner screen into the media at
velocities ranging from 100-150 feet per minute. The particulate is removed
from the gas stream by contact with the media. Clean gas exits through the
outer screen and through the appropriate breeching to an exhaust stack.
Note that the gas flow may be reversed for some applications.
The filter media moves continuously downward by gravity to prevent a
filter cake from forming on the face of the filter and to prevent a high
pressure drop. To provide complete cleaning of the louver face, the louvers
are designed so that some of the media falls through each louver opening,
thus preventing any bridging or buildup of particulate material. The
particulate-laden media is continuously removed at the bottom of the granular
filter where it is transported by a pneumatic conveying system to the air/
particulate de-entrainment section of the system. Media recirculation rate
is controlled through the use of an L-valve, shown in Figure 2. The injector
air forces the media into the main lift air flow, transporting the media up
the lift pipe. The lift air velocity is approximately 80 feet per second
and the media velocity is less than 15 feet per second. The media and air
then enter the de-entrainment vessel, shown in Figure 3. The ash or dust
is removed from the media by the mechanical interaction of the media as it
is being conveyed up the lift pipe and by the scrubbing action of the air
as it lifts the media. The clean media then drains by gravity to the top
of the filter unit for another cycle. The particulate is pneumatically
conveyed to a particulate separator and storage silo.
290
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PARTICULATE TO
FINAL DISPOSAL
iih r /
ELECTROSTATIC
GRID
ELECTROSCRUBBER
ELEMENT
MEDIA LIFT
AIR BLOWER
DE-ENTRAINMENT
VESSEL
PNEUMATIC MEDIA
RECIRCULATION
& PARTICULATE
REMOVAL SYSTEM
(LIFT PIPE)
MEDIA FILL HOPPER
Figure 1 Electroscrubber Granular Filter Schematic
291
-------
DIRTY MEDIA FROM
ELECTROSCRUBBER ELEMENT
DIRTY MEDIA TO
DEENTRAINMENT VESSEL \
/V-
INJECTOR CONTROLLING
MEDIA FLOW RATE
CONTROL AIR ON
LIFT AIR ON
Figure 2 L-valve.
DEENTRAINMENT
CHAMBER
t—r
— >
DIRTY MEDIA
CLEANED MEDIA
MEDIA
OVERFLOW PIPE
RATTLER
PURGE AIR
RATTLER
BARS
DUST/AIR FLOW
REMOTE DUST
REMOVAL &
STORAGE SILO
MEDIA LEVEL
CONTROL SECTION
MEDIA RETURN PIPE
\ CLEANED MEDIA RETURNED
TO ELECTROSCRUBBER
DUST LADEN MEDIA
FROM ELECTROSCRUBBER
Figure 3 De-entrainment vessel.
292
-------
The electrostatic grid, in the form of a cage, is positioned within the
media bed of the Electroscrubber granular filter. A high voltage is applied
to this conductor and the electric field generated between the conductor and
the inner and outer screens enhances the collection of particulate in the
gas stream.
Commercial Electroscrubber granular filters meet a variety of process
requirements. Depending upon the application, inlet loadings up to 5 gr/acf
and temperatures from 250 to 850 F are specified. The filter vessels are
from 9 to 13 feet in diameter and 35 to 60 feet in height, with each vessel
cleaning from 25,000 to 125,000 acfm. For very low outlet loadings the
current modular design features up to four elements per module, and some
functions can be integrated for multiple installations. Figure 4 shows the
modular design for a multiple installation, in which large volumes of gas
are cleaned.
MECHANICAL COLLECTION
Collection of particulate in a packed bed of granular material by
purely mechanical means has been studied by many investigators. Neglecting
electrostatic effects (other than those which hold the particle to the
collector after capture) for the purposes of this discussion, granular
media collects aerosol particles by at least four mechanisms: inertial
impaction, interception, diffusion, and sedimentation. Of these effects,
impaction and diffusion are the mechanisms of principal importance in
particle size distributions associated with most pollution control
applications.
Collection by impaction occurs when the inertia of a particle causes it
to depart from the gas streamline and collide with the collector. Theore-
tical models indicate that the Stokes number is a satisfactory correlating
dimensionless group for these analyses; The Paretsky modeU for impaction
collection efficiency associated with a packed bed is given by the relation:
= St
nimp e
where St is the Stokes number and e is the void fraction.
Brownian diffusion caused by random motion of small particles being
bombarded by gas molecules enhances the possibility of a particle being col-
lected. Even though the particle may be generally following a gas stream-
line, random motion occasionally allows the particle to approach a collector
surface and causes it to contact that surface and be captured. Friedlander^
recommends that collection efficiency for a single isolated collector by
diffusion can be expressed as:
293
-------
GROUND ROD 5/8" 0
BURIED MIN. 8' 0"
IN GROUND
IBONOEO COPPER)
HIGH VOLTAGE
POWER SUPPLY
GAS OUTLET
-PNEUMATIC
CLEANUP SYSTEM
CLEANUP SYSTEM-
- DEENTRAINMENT
CHAMBER
~ CLEAN MEDIA
INVENTORY CHAMBER
£ INLET
fc~
MEDIA OVERFLOW PIPE
COUNTERFLOW AIR
MEDIA ADD HOPPER
- SUPPORT LEG
- SLIDE VALVES
MEDIA LIFT PIPE
SEAL LEGS
— AIR BLOWER
Figure 4 Electroscrubber Granular Filter Modular Design
294
-------
diff =
4.36Pe
-2/3
where Pe, the Peclet number, is related to the ratio of bulk to diffusive
mass transport and , like the Stokes number, is a simple function to
dimensions and properties.
From the physical description of these phenomena, it can be seen that
inertial impaction most effectively operates on larger particles (with a
significant mass/area ratio) whereas diffusion is more effective on small
particles whose trajectory can be influenced by molecular motion. The
other effects, interception (a limiting case of impaction wherein a particle
grazes the collector surface) and sedimentation (which describes the effect
on gravitation on particle trajectory) are small and will not be discussed
here. Reference 3 treats these phenomena more thoroughly.
Considering the various collection mechanisms to represent n "stages"
(each of efficiency n^) which operate successively on the original dust
load, the overall collection efficiency n" can be written as:
(1-rf) = n (1 -n •)
i = l 1
or, considering the two mechanisms discussed above,
(1-n) -
, ..(a)-
1 . pe"2/3
where m is an empirically determined exponent which depends upon the
Stokes number range.
ELECTROSTATIC COLLECTION
In addition to the mechanical collection mechanisms, the electro-
statically enhanced granular filter utilizes an electric field in the media
for particulate collection. In an electrostatic precipitator (ESP) a
corona discharge is created between a high voltage wire and a ground elec-
trode. The field used to generate the corona collects particulate which
has been charged by the same corona. The Electroscrubber granular filter
differs from the ESP in that there is no corona. Collection due to
295
-------
electrostatic forces relies on the particulate charge which occurs naturally
from upstream processes. Frictional charging (or tribocharging) will occur
in flow through pipes and cyclones and in high mixing environments. Collec-
tion occurs when the direction of travel of charged particulate is charged
by virtue of the electrostatic force, in opposition to the fluid flow field
and viscous forces, and moves through the fluid toward a collection surface.
Although the force is directed perpendicular to equipotential surfaces, the
particulate undergoes many changes in direction through a packed media bed,
and has many opportunities to come in contact with the media. The electro-
static collection in the granular filter is similar to an electrostatic
precipitator with very narrow and winding walls and therefore increased
effective surface area. Of course, the mechanical collection mechanisms of
impaction and diffusion contribute significantly to the overall collection.
The electric field strength in the filter is a function primarily of
applied voltage and geometry of the electrodes. Secondary effects include
presence of the media and the gas as highly resistive dielectrics. Thus,
within the well-defined overall field there are local changes in field
strength and direction due to the media and, on a smaller scale, the parti-
culate. Local gradients in field strength provide another small force on
the particulate resulting in dielectrophoresis. The collection of parti-
culate is therefore related to a number of factors. The actual collection
is affected by one or more of the collection mechanisms: impaction, inter-
ception, diffusion, and electrostatic forces. The one or two mechanisms
that dominate in an electrostatically enhanced granular filter will depend
on the following physical parameters: particulate size, density, charge,
collector size, and fluid velocity. In addition, re-entrainment, the
rubbing off and recapture of particulate from the media, affects the net
collection efficiency.
MATHEMATICAL MODEL
Some investigators have attempted to model the electrostatic effects on
particulate4»5; however, there are few theoretical models describing the
behavior of charged particles in a complex flow field. Besides solving for
the fluid flow field, a detailed model must account for inertial effects
and small variations in the electric field within the granular filter. The
following simple model includes the major parameters affecting particulate
collection, while assumptions are made regarding geometry and particulate
1oad i ng.
It was shown previously that the total penetration is equal to the
product of the penetrations due individual mechanisms. For a system in
which both mechanical and electrical effects are important,
^_T1^tot ^-T^elec ^"n^mech
296
-------
In order to include the electrostatic effect within the system of
equations, an expression is needed for the collection efficiency as a func-
tion of particle charge, average electric field, etc. An approach similar
to the Deutsch equation^ first developed for standard ESP's is used: the
force on a charged particle due to an electric field is equated with the
viscous force when it reaches terminal velocity. Relating this migration
velocity to the geometry of the collecting surface and flow field yields:
nelec = 1 - exp
^ q CAE
3irp C
qTI
f P
E = Average Electric Field Strength
A = Collection Area
c
= Volumetric Gas Flow Rate
C = Cunningham Correction Factor
q^ = Particle Charge
d = Particle Diameter
P
y = Gas Viscosity
Considering the geometry of a granular bed, Ac/Q^ is approximated as:
V«f
6(1-e)t
V
U = Gas Superficial Velocity
e = Void Fraction
t = Bed Thickness
d = Media Granule Diameter
Substituting,
elec = 1 - exp
^ (l-e)t E q C
Try p
U d d
9 P
297
-------
From the above it is noted that increased efficiency can be obatined by
increasing: electric field strength, particulate charge and filter thick-
ness, or by decreasing: flow velocity and media granule diameter. An
increase in temperature is seen to reduce efficiency through the increase
in gas viscosity. According to the model, larger particles actually have
increased efficiencies, as the particle charge is assumed proportional to
the square of the diameter (surface area)./
Two factors of importance in utilizing the electrostatic concept are
the particulate charge and the media resistivity. Tribocharging of the
particulate to nominally 10% of saturation is more than sufficient to pro-
vide the desired electrostatic performance. Media resistivity is important
because it affects the power required by the grid.
EXPERIMENTAL DATA
Substantial testing has been performed on full-size and pilot-scale
moving-bed granular filters. Both gravimetric and impaction classifier data
have been obtained for a veriety of operating conditions. The Snoqualmie
Falls Electroscrubber, located at a Weyerhaeuser mill in Washington, is a
Dry Scrubber retrofitted with an electrostatic grid. The Low-Outlet-
Loading Electroscrubber is a pilot unit located at Combustion Power Company's
Menlo Park facilities.
Snoqualmie Falls Installation
The Snoqualmie Falls installation brought on line in 1974 consists of
a Dry Scrubber designed to handle 45,000 ACFM of gas flow from a hog-fuel
fired boiler. This unit has an 18-in thick media bed and the cleaning/
transport system consists of a feeder-screener and bucket elevator rather
than a pneumatic lift system.
CPC undertook a development program with a full-scale installation that
would demonstrate the dramatic increase on filtration performance due to
electrostatic effects within the filter bed of a moving-bed granular filter.
For this purpose the Dry Scrubber was upgraded and retrofitted with an
electrostatic grid placed in the filter media. The electrostatic grid is
energized by a high voltage power supply capable of delivering 60 ma at
25 KV DC. Figure 5 is a schematic of the installation after it was
retrofitted.
After completion of the high voltage system installation, a test pro-
gram was initiated with the objective of evaluating the concept of electro-
static cleanup on an operating industrial system, determining any functional
problems and collecting data on the enhancement of the filter collection
efficiency as a function of grid voltage.
298
-------
EXISTING
CYCLONE
SEPARATORS
FLY ASH
FEEDER-SCREENER
TO FILTER
INLET
OPACITY
METERQ
SAMPLING
PORT
EXISTING
STACK
BUCKET
ELEVATOR
lj—.OPACITY
METER
FROM
GRID -
BOILER
DAMPER
SAMPLING
PORT CK
DUST
FEEDER [~j"
RATTLER
300 ACFM
BAG TYPE
DUST COLLECTOR
WITH FAN
Figure 5 Snoqualmie Falls Electroscrubber Granular Filter
-------
Table 1 is a summary of the test results. Inlet loadings range from
.138 gr/sdcf to 1.128 gr/sdcf (12% C02). The high inlet loadings were
obtained by dust injection to supplement the loading generated by the boiler.
The outlet loadings ranged from .006 gr/sdcf with the electrostatic grid
energized (20 KV) to .050 gr/sdcf with the electrostatic grid completely
de-energized. Efficiencies of 99% are achievable.
Figures 6 and 7 show the effect of grid voltage on outlet loading and
opacity. It is shown that mass outlet loading and outlet opacity can be
reduced by a factor of 5 or more from baseline values by energizing the grid
to 20 KV DC.
From the particle size distribution samples the fractional efficiencies
were calculated. Figure 8 shows the fractional efficiency for the Snoqualmie
Falls unit for different grid voltages. From these curves two basic con-
clusions can be drawn: electro-enhancement increases (65% to 95%) the
collection efficiency for submicron particles, and electrostatics tends to
flatten the fractional efficiency curve, i.e., makes particle collection
efficiency less dependent on particle size.
Another way of representing the data that has been found useful is the
plot of particulate penetration through the filter as a function of a
dimensional grouping that characterizes the filter performance. Figure 9
is a plot of penetration versus aPL-j/M, where AP is the pressure drop (Iwd),
L-j is the particulate inlet loading (gr/acf) and M is the media circulation
rate (lb/min). The best straight line curve fit was obtained from a least-
squares analysis of the points. The substantial decrease in penetration
is most apparent from this figure.
Low-Outlet-Loading Electroscrubber
Experimental data was also obtained from a pilot-scale unit, the Low-
Outlet-Loading Electroscrubber granular filter (LOL), shown in Figure 10.
This unit was designed with a media circulation system similar to the system
on the elements being installed on Weyerhaeuser Company's boiler unit at
Longview, Washington. The inner screen has an inside diameter of 17-1/2"
and an active screen height of 10 feet thereby providing a flow area of
45.8 ft^. The media bed thickness is 24 inches.
The media circulation and cleaning system is pneumatic and has a cap-
ability of lifting 300 lb/min of 1/4 x 1/8 filter media to 60 feet and
returning the cleaned media to the test unit through two return pipes. The
particle-laden air is separated from the media and passed through a small
baghouse and exhausted to atmosphere. The principal air supply is a 50Hp
induced draft blower with manual inlet vanes control. Air flow capability
is 7000 acfm @ 14 Iwd and 200 F. The air is heated before passing through
the media bed by a packaged natural gas burner capable of increasing the
inlet air temperature to approximately 300 F.
300
-------
Table 1 SNOQUALMIE FALLS DATA
GRID GRID FILTER OUTLET SINGLE PASS COLLECTION
TEST VOLTAGE CURRENT D/S AP GAS FLOW OPACITY EFFICIENCY
NUMBER (K Volts) (ma) (IWd) (acfm) (%) (%)
1 0 0 2.8 24889 5-6 83.3
2 10 3 2.8 28890 2 89.6
3 10 13 3.0 26971 4 94.1
4 10 4 2.7 26467 2 89.1
5 15 9 3.0 25945 1 94.2
6 15 14 3.2 25009 1.2 94.2
7 15 6.5 3.3 26887 1 93.5
8 20 16-30 3.2 25910 1-5 95.2
9 20 14 3.0 24762 1-3 97.1
10 0 0 2.7 25105 12 80.8
11 10 6-9 2.8 25144 2.5 94.3
12 0 0 2.4 25685 5 73.6
13 20 (3-2) 5.4 37073 1-1.5 95.8
14 20 16-24 2.7 24156 1-1.5 95.8
15 20 7.5-13.5 2.3 25729 .5-1 95.8
16 20 14-16 2.8 23376 .5-1 96.8
17 20 20-29.5 2.5 24303 — 97.4
18 20 19-39 4.9 36736 1.5 93.5
19 (20 36-53 3.3 26851 1.5 99.0
20 20 20-56 3.5 28439 1.7 98.6
21 10 14-16 6.4 47740 5-6 08.6
22 20 14-28 4.6 38552 1.5 96.1
23 0 0 2.1 23872 7 79.3
24 10 11.5-13.5 4.0 27656 8 98.0
Note: Efficiency calculated using front half loadings by EPA
Method 5 at inlet and outlet of filter.
301
-------
05
04
03
¦M gr/sdcf
02
.15 gr/sdcf
01
0
10
20
GRID VOLTAGE, KVOLTS
Figure 6 Snoqualmie Falls test results -
outlet loading vs. grid voltage
and inlet loading.
>-
o
UJ
—J
.40 gr/idcf
UJ
z:
CO
.15 gr/idcf
GRID VOLTAGE, K VOLTS
Figure 7 Snoqualmie Falls test results -
single pass opacity vs. grid
voltage and inlet loading.
-------
100
20K VOLTS
90
15 K VOLTS
80
10K VOLTS
u
2
(J
Z
o
u
e 60
O
u
u
cc
<
50
40
30
7
3
4
B
2
.7
.3
.4
.6
10
.2
.1
1
PARTICLE AER00YNAMIC OIAMETER IMICRONSI
Figure 8 Fractional Efficiency, Snoqualmie Falls Electroscrubber
Granular Filter
V ~
GRID DEACTIVATED
-
r 20 K VOLTS
\ °
0
^VP N.
°
OV
1
1 l n\
1 2 5 10 20
A? Lj ^.3 r Iwd x gr/icf "1
M L Ib/min J
Figure 9 Penetration vs. Filter Parameter, Snoqualmie Falls
Electroscrubber Granular Filter
303
-------
MEDIA RETURN
INLET PIPE
INNER SCREEN
BURNER
FAN
ADD SYSTEM
n
i
DE ENTRAPMENT
VESSEL
TO DUST REMOVAL
SYSTEM
STACK
LIFT AIR PIPE
SEAL LEG
LIFT AIR FROM BLOWER
INJECTION AIR
Figure 10 Low-outlet-loading Electroscrubber Granular Filter
Test Facility
304
-------
The dust injection system uses the vacuum generated by the induced
draft blower to inject any type of dust into the system inlet. Dust is in-
troduced by a screw feeder at a mass rate dependent on the physical proper-
ties of the dust being injected and the feed speed. An electrostatic grid
consisting of a series of vertical pipes is installed in the media cavity.
A 75 kilovolt high voltage power supply (15 ma max) supplies the potential
for the gird. The system has been operated a total of over 800 hours in
a series of functional and performance tests. Particular attention was given
to tests of the pneumatic system to explore actual capabilities of the
system for media lifting and cleaning.
Filtration performance of the system was evaluated by collecting
simultaneous inlet and outlet total loading samples using standard EPA
Method 5 procedures. The inlet and outlet particle size distributions were
obtained by sampling with an Andersen 2000 Mark III impaction classifier.
The baseline conditions with the grid voltage at 0 kv were obtained first,
sampling simultaneously at the inlet and the outlet. The grid was then
energized to 20 kv and an outlet sample taken without changing upstream
conditions. This data indicates that, for those operating conditions, the
outlet loading can be reduced over 90% by applying a voltage of 20 kv to
the grid. Dramatic increase of collection efficiencies for particles smaller
than 1 micron was noted. Increasing the voltage beyond 20 kv does not
provide significant increase in efficiencies. The particulate used in these
tests was collected by an existing Dry Scrubber installed on a hogged-fuel
fired boiler. Further tests are planned using lime-kiln dust and coal-fired
boiler ash. Also, further tests are planned for a thinner (12") media bed.
Figure 11 is a plot of penetration versus the dimensional parameter
APLn-/M. The straight line is from a least-squares analysis. As found
previously with the Snoqualmie Falls data, this grouping satisfactorily
characterizes the filter operation performance. These curves again demon-
strate the substantial decrease in penetration when using electrostatic
enhancement.
L0NGVIEW ELECTROSCRUBBER
Figure 12 is a photograph of the new Electroscrubber installation at
Weyerhaeuser Company's Longview, Washington plant. Figure 13 is an isometric
drawing showing some of the details of the installation. The electro-
statically-enhanced granular filer system is configured as described earlier
in Figures 1, 2, and 3. The unit is designed to handle up to 420,000 acfm
of gas from a hog-fuel fired boiler at temperatures of 350 F and inlet par-
ticulate loadings up to .4 gr/sdcf corrected to 12% C0o. The installation
consists of 12 elements, each capable of handling 35,000 acfm arranged in
3 modules with each module having an exhaust stack. Each element has a
separate media pneumatic lifting/cleaning system. Each module has a
separate baghouse to which the particulate-laden air is pneumatically
305
-------
.1
.05
.02
I-
<
cc
}f .010
Q-
.005
.002
~
~
S^O
~
^-O K VOLTS
—¦
/— 20 K VOLTS
—
o
1
1
1
1
10
20
50
100
AP Li
L x 10J
M
Iwd x gr/acf "|
Ib/min J
Figure 11 Penetration vs. Filter Parameter, Electroscrubber
Granular Filter Test Facility
306
-------
Figure 12 Longview Electroscrubber.
EXHAUST STACK
BAG HOUSE
INLET DUCT
OLD BOILER STACK
DE ENTRAPMENT
VESSEL
BAGHOUSE
FD FAN
ELECTROSCRUBBER
ELEMENT
Figure 13 Longview Electroscrubber schematic.
307
-------
conveyed. The exhausts of the 3 baghouses are manifolded to a single exhaust
stack. Each element has its own electrostatic grid powered by a power supply
capable of delivering 120 ma at 30 kv dc. The elements are filled with 1/4
x 1/8 gravel and the pneumatic circulation systems will circulate the filter
media.
SUMMARY
The utilization of electrostatic forces in granular filters provide an
additional mechanism for collection of particulate and is especially effec-
tive on micron-size particles. Data obtained from an industrial unit
retrofitted with electrostatic equipment and a pilot-scale filter has demon-
strated a considerable increase in collection efficiency using electrostatic
enhancement. Besides the mechanisms of impaction and diffusion, the electro-
static mechanism provides an important means for improving filter performance
at minimal additional capital and operating expense. Unlike electrostatic
precipitators which depend on corona-charging of the particulate, the
Electroscrubber granular filter uses the interaction of the tribocharged
particles with the electric field and small-scale electric field effects
to enhance contact of the particulate with the media. The use of electro-
statically enhanced moving-bed granular filters offers a relatively new
and effective means of meeting the particulate control objectives of the
next decade.
REFERENCES
1. L. Paretski et al, J. Air Poll Control Assn., Vol 21, p. 204 (1971).
2. S. K. Friedlander, Communication to CPC (Unpublished, 1977).
3. K. Phillips et al, Granular Bed Filter Development Program Final Report.
DOE Report # FE-2579-19 (1978).
4. W. E. Ranz and J.B. Wong, Impaction of Dust and Smoke Particles on
Surface and Body Collectors, Industrial and Engineering Chemistry,
June 1951.
5. J. R. Melcher & K. Zahedi, Electrofluidized Bed in the Filtration of
a Submicron Aerosol, Journal of the Air Pollution Control Association,
April 1976.
6. M. Crawford, Air Pollution Control Theory, 1976.
7. H. J. White, Industrial Electrostatic Precipitation, 1963.
308
-------
ELECTRICAL AUGMENTATION OF GRANULAR BED FILTERS
S. A. Self, R. H. Cross and R. H. Eustis
High Temperature Gasdynamics Laboratory
Stanford University
Stanford, California 94305
A study of the enhancement of the collection efficiency of granular bed
filters by electrical means is reported. By applying electric fields of a
few kV/cm (dc or ac) across a bed of insulating granules, the efficiency for
submicron charged aerosol is greatly increased, to the point where the
efficiency minimum normally observed with such filters in the size range
0.1-1 ym is removed. The performance of such filters is explored as a function
of granule size, applied field, face velocity and charge state of the entering
aerosol with both a bench-scale flow system (to 40 ft^/min) and a wind tunnel
(to 600 ft3/min). Some theoretical estimates are made to identify the most
likely physical mechanism of electrical enhancement, and a number of potential
application possibilities are discussed.
309
-------
INTRODUCTION
A great variety of air filtration devices for particulate emission control
have been explored and developed over the years, varying from purely mechanical
devices such as cyclones, scrubbers and fabric, fiber and granular bed filters,
to purely electrical ones exemplified by electrostatic precipitators. Because
electrical forces can be so much stronger than mechanical ones, it appears that
incidental electrical effects often play an unrecognized role in nominally
mechanical devices. In recent years, various attempts have been made to capi-
talize on this fact, and enhance the collection efficiency of basic mechanical
devices by deliberately incorporating electrical features. Examples include
electrified scrubbers^->2, electrified fiber^ and fabric^ filters, and electri-
fied granular beds, both packed and fluidized^-?. An overview of electrostatic
devices for control of submicron particles has been made by Melcher et al.®.
In the present work, the enhancement of the collection efficiency of a
packed granular bed filter by electrical means has been explored. The work
has been primarily experimental, utilizing a bench scale flow apparatus
operating at throughputs of up to 40 ft^/min; some additional experiments were
made in a tunnel facility at throughputs up to 600 ft3/min. In addition some
preliminary theoretical work has been undertaken concerning the nature of the
electrical interaction.
To provide a background framework for interpreting this work, a summary
is given in the next section of the present state of understanding of filtra-
tion by granular beds in the absence of electrical effects. More extensive
reviews of this topic have been given by Spielman^, by Tardos et al.^0 and by
Schmidt et al.H.
FILTRATION BY GRANULAR BEDS
Filtration is characterized by the simultaneous action of forces of fluid
mechanical, inertial and gravitational origin, together with electrical forces
acting between the particles and the collectors. These combined forces govern
the particle trajectories, which determine whether particles intercept the
collectors and are retained. A rational approach to understanding filtration
in an electrified bed is to first consider the mechanisms operative in the
absence of electrical effects, and then consider the modifications due to their
inclusion.
The theory of filtration neglecting electrical effects is usually discussed
in two steps. First, collection by an isolated filter element (sphere or cylin-
der) is considered. This subject is well established and is summarized below.
Second, collection by fibrous or granular bed filters is developed from the
theory for isolated collectors, but modified to give the efficiency per repre-
sentative element taking account of the flow modifications produced by the
surrounding elements. This approach, outlined below, is less firmly established,
and is better justified for tenuous media (fibrous or fluidized bed filters)
than for dense media such as fabric filters or packed granular beds.
310
-------
Aerodynamic Capture by an Isolated Sphere
This is based on the solution for flow past a sphere, which is well docu-
mented for the cases of viscous creeping (Stokes) flow (Re « 1) and potential
flow (Re » 1). Solutions for intermediate values of Reynolds numbers are listed
by Tardos et al.-^.
12
As discussed in the classic text by Fuchs and the more recent one by
Friedlander^, the collection efficiency is then calculated as the sum of four
physical processes, namely, interception, impaction, diffusion and gravitational
settling, though the latter is often neglected. These processes determine
whether a particle reaches the collector surface; however, whether it is per-
manently captured depends on the balance of attaching forces (van der Waals,
capillary, electrical, etc.) against the fluid mechanical forces tending to
re-entrain it. The overall collection efficiency then depends on the net effects
of collection and re-entrainment.
The cross section for interception is taken as the area normal to the in-
coming flow, extending from the axial streamline out to the radius of the limit-
ing streamline which passes within one particle radius of the collecting sphere's
surface. The collection efficiency is then taken as this area divided by the
projected area of the sphere. For potential flow, Fuchs^ and Ranz and Wongl^
suggest the following efficiency formula, with coefficients 8 of 2 and 3/2,
respectively:
nlnt(p) "B(ap/ac) (1>
where ap and ac are the particle and collector radii, respectively. For
viscous flow, Fuchsl2 amd paretsky-^ suggest the formula
"int^ " I (ap/ac)2 • (2>
Since (a /ac) is « 1, eq. (2) yields very much lower efficiencies than eq. (1).
In either case t is independent of velocity, viscosity and particle density.
Impaction relates to the departure of the particle trajectories from the
fluid streamlines due to inertia. For a given flow field the trajectories can
be computed and the cross-section for impaction is then the area normal to the
upstream flow extending from the axis out to the limiting trajectory which just
intercepts the sphere surface. As for interception, the efficiency is then the
ratio of this area to the sphere's projected area. According to theory, for a
given Reynolds number, the efficiency should be a function only of the Stokes
parameter
2
2p a u
Stk = 9|m <»
c
311
-------
where p is the particle density, u is the free stream velocity and y is
the gas Viscosity. For example, for air at STP, y ~ 1.8 x 10~5 Kg/m*s, and for
the typical value p = 2 x 10^ Kg/ra^, a = 10 ym, a = 1 mm and u = 1 m/s,
Stk * 2.5. P P
For both potential flow and viscous flow, calculations show that the
efficiency for impaction ri^m is zero below some critical Stokes parameter
Stkcr, increases rapidly throBgh values of Stk of order unity, and asymptotes
to unity (100%) for Stk > 10. For potential flow, Langmuir^^ gives
StkCr = 0.083 and the approximate formula for Stk >_ 0.2
'"'imp
<"> "(stkfo.s)2- <4>
For viscous flow, he found Stkcr = 1.214 and the approximate formula
T, . 0.75 In (2Stk)~|
imp " L1 + Stk - 1.214 j ' (5)
Later workers have obtained similar results which indicate that impaction is
very efficient for large particles but is insignificant for small particles,
with a rapid dependence on particle size. The collection site for impaction
is concentrated around the upstream facing surface of the collector.
Diffusion due to Brownian motion of submicron particles promotes their
deposition during flow past a collector. The diffusion coefficient for Brownian
motion is given by the Stokes-Einstein formula
D = KTC/6 tt ya
(6)
-23
where k is the Boltzmann's constant, 1.38 x 10 J/K, T is the absolute
temperature, and C is the Cunningham correction factor to Stokes' law when
a is comparable to or less than the mean free path X:
C = 1 + — 1.26 + 0.40 exp(-1.10 a A)
a L P
(7)
The steady state convective diffusion equation
u
Vn = D V n
(8)
where n is the concentration of particles in a given narrow size range, has
been solved by several workers-^ with very similar results for the efficiency
for diffusive collection by an isolated sphere:
312
-------
/ d -2/3 T kTC "1^/3
n,._- = 4.0 [——) ~ 4.0 (Pe) z/j = 4.0 KTC (9)
diff \2acu/ Ll2TryapacuJ K J
where Pe is the Peclet number. For instance, for air at STP, ac = 1 mm,
ap = 0.1 ym and u = 1 m/s, we have C * 1.87, D = 1.22 x 10"10 m2/s,
Pe = 1.64 x 10 and = 6.8 x 10-^. It may be noted that the Brownian
diffusion coefficient Is typically very small, and that eddy diffusion due to
weak turbulence in flows of low Reynolds number may in practice be comparable
to or larger than that due to Brownian motion.
The calculated fractional efficiencies versus particle size for an isolated
sphere by interception, impaction and Brownian diffusion for typical conditions
of interest are shown in Figure 1. Also shown are the total fractional effi-
ciencies for the cases of potential flow and viscous flow. Both show a minimum
efficiency due to the fact that, with increasing size, diffusive capture de-
creases while capture due to interception and impaction increases. Such a
minimum, usually falling in the size range 0.1-1.0 ym is characteristic of
packed bed and other filters. For potential flow, the minimum is at ~0.08 ym,
while for viscous flow the minimum is at ~2.0 ym and about an order of magnitude
smaller in value. The case illustrated has a Reynolds number of ~70 so that
one would expect a rigorous treatment to yield a solution falling between these
limiting cases of viscous and potential flow. Langmuir suggested the inter-
plation formula for impaction
n = tn(v) + n(p)(Re/60)] ,10)
imp [1+ (Re/60)] ' UU'
Aerodynamic Filtration by a Packed Bed
Model computations of the collection efficiency of fibrous and packed bed
filters are based on the theory for collection by isolated cylinders and spheres,
respectively, but modified to give an average efficiency per representative
element (cylinder or sphere) which takes account of the modification of the flow
by the surrounding elements. Such models are called cell models, and are
inevitably of dubious validity because they endeavor to account for the inter-
ference effect of the neighboring collectors in only a very approximate way.
In the case of typical fibrous filters, where the packing fraction is very low
(0.01-0.1), this approach has greater validity than for a close packed array
of spherical collectors with a typical packing fraction of 60-70%.
The overall fractional efficiency of the bed as a function of particle
size is then obtained by integrating the following differential relation for
the decrease in particle concentration
iS = _ 3otn_n
dx 4a Ui;
c
313
-------
where n is the total fractional efficiency per cell, a is the packing frac-
tion, and x is the flow direction. The fractional penetration P for a
bed of thickness L is then given as a function of particle size by
(12)
and the fractional efficiency of the bed is then
nB - 1 - P-
(13)
Equation (12) allows effective fractional efficiencies per cell, or sphere, to
be calculated from measured values of the penetration.
The total mass collection efficiency for a given input aerosol size distri-
bution is given by integrating the mass-weighted fractional efficiency over the
size distribution.
20
The most widely used cell model is that of Happel and Brenner , which
treats a representative sphere occupying the center of an idealized cell. The
cell consists of two concentric spheres, an inner one representing the collector,
and an outer sphere representing a fluid envelope having a free surface, i.e.,
having no tangential stress. The ratio of the volumes of the inner and outer
spheres is taken equal to the packing fraction of the real bed. The viscous
flow equations are solved for the cell, which allows the pressure drop or
specific flow resistance to be evaluated. The flow field also serves as the
basis for calculating particle capture.
Numerical solutions for the efficiency of impaction per sphere using
Happel's cell model with viscous flow were obtained by Paretsky et al.-^.
In the limit a -* 0, for an isolated sphere, they reproduce the result of
Langmuir for viscous flow with Stkcr = 1.214. With increasing packing frac-
tion, the efficiency curve (see Figure 1) is translated to lower Stokes numbers;
for instance for a ~ 60%, Stkcr is reduced by about an order of magnitude to
~0.1. This is probably attributable to the effect of fluid jets in the flow
passages increasing the effective Stokes parameter above that calculated on
the basis of the superficial velocity.
The collection efficiency per sphere by diffusion was evaluated by
Pfeffer21 for viscous flow as
(14)
where
p(a) = [2 - 3a1/3 + 3a5/3 - 2a2]/(I - a5/3).
(15)
314
-------
In the limit of an isolated sphere, a -* 0, p -+¦ 2, and eq. (14) reduces to
eq. (9). For a close packed bed with a = 0.6, p(a) - 1.15, so that the
Pfeffer model gives a very modest increase in efficiency per sphere over that
for an isolated sphere.
9
As Spielman comments, there is considerable disagreement among workers
as to which, if any, of the different models is best suited to predict particle
capture by fibrous and granular filters." Cell models completely ignore the
tortous flow passages associated with the topology of a real packed bed.
Payatakes et al.^2 disapprove of representative sphere models for densely
packed beds. As an alternative they discuss a constricted tube model which
requires numerical solution to predict particle capture. It would appear that
some such model, which takes a more realistic view of the actual geometry of
a packed bed will be necessary to obtain more useful formulae for the capture
efficiency in close-packed beds. Such a model would also appear to be necessary
to deal realistically with electrified beds where, as discussed below, the
electric field is very inhomogenous.
Experimental
A number of experiments on filtration by packed beds have been reported,
as reviewed by Tardos et al.^®. However, the data often covers only limited
parameter ranges and is not particularly well correlated with theory. Measure-
ments of penetration are commonly made as a function of superficial velocity
for a fixed monodisperse aerosol. When plotted as fractional efficiency versus
aerosol diameter, the data typically shows a minimum in the size range 0.1-1 ym.
Recent work by Goren23 extends over a wider range and is empirically correlated
in dimensionless plots of single grain efficiency against Stokes number, Peclet
number and also a gravity number Grv = 2 Pp ajj Cg/9yu, where g is the gravi-
tational acceleration. Nevertheless, the correlation of theory and experiment
leaves much to be desired. To quote Tardos et al., "There is still considerable
work to be done in order to explain the operation of granular filters. This
includes theoretical models that will take account of electrical forces, the
entrainment of dust from the collection surface, unsteady characteristics of
the filter, and lastly comparison of these theoretical results with well-
designed experiments."
Despite being imperfectly understood, granular bed filters are, in practice,
effective filters for particulate emission control, and are being investigated^f
along with other devices, for high temperature applications.
For conventional stack gas temperatures, commercial-sized units (to 200,000
acfm or more), known as dry scrubbers, are manufactured by Combustion Power
Company, and have been installed on a number of hog fuel boilers and other
systems where they are competitive with alternative devices such as electrostatic
precipitators.
In this form the bed of gravel granules (1/8" - 1/4") occupies the annulus
between two concentric vertical louvered steel cylinders. The dusty gas flows
radially through the bed from an input plenum, and the cleaned gas exhausts
upwards through the outlet duct. The gravel medium moves slowly downwards in
315
-------
approximate plug flow and exits at the bottom, where it is cleaned of dust and
recycled to the top of the bed. Such filters are sometimes used with a pre-
cyclone inertial separator to remove the bulk of the large size fraction to
avoid loading the bed unduly. The optimum design of the cyclone-filter combina-
tion depends on the particle size distribution and involves trade-offs between
the pressure drops and collection efficiencies of the two components.
Recent reports^ of field tests of such a granular bed filter, fitted with
high voltage electrodes in the bed, indicate a dramatic decrease in plume
opacity on a hog fuel boiler, and are attributed to a significant electrical
enhancement of the efficiency for subtnicron particles.
ELECTRIFIED BEDS - THEORETICAL CONSIDERATIONS
A considerable variety of distinguishable electrical effects are possible,
depending on the charge state of the entering aerosol and the configuration of
the bed with regard to the electrical properties of the granules, the electrode
configuration, the magnitude and character (dc or ac) of the applied voltage, etc.
Furthermore, the electrical states of both the particles and the bed can change
as a result of their mutual interaction.
Here we consider first the field (and charge) distribution in the bed alone,
and then the interaction of particles with the bed in terms of several distinct
electrical mechanisms.
Field Distribution in an Electrified Packed Bed
We consider a packed bed of semi-insulating granules immersed between
electrodes between which an external voltage is applied. The electrodes are
assumed to have large radii of curvature relative to the granule size, so that
they do not support corona discharges, and the scale of the field variation is
determined primarily by that of the granules and not of the electrodes. Because,
even for a perfectly close-packed bed of uniform spheres, the geometry is quite
complex, the analysis is restricted to the simple case of a regular cubic array
of contacting spheres between external electrodes located along principal lattice
planes (Figure 2(a)). This model has been discussed by McLean^ in connection
with the resistance of precipitated ash layers in electrostatic precipitators.
Though idealized, it apperas that the model describes the essential features
of the field distribution in practical beds.
Several points should be made at the outset. First, it is important to
recognize that no insulator is perfect, so that in a steady state there will
be a (small) current flow, and the distribution of potential and electric field
is determined by the bed's resistive properties, i.e., it is a problem in resist-
ive current flow rather than one of electrostatics for perfect dielectrics.
Furthermore, if the bed is used to filter an aerosol carrying a net charge,
the bed resistance must be low enough to leak off the current due to the
collected particles without producing fields comparable to those applied.
Otherwise the bed will charge up, bias off the applied field, and be self-
limiting. Even for an aerosol which is grossly neutral, but is bipolarly
charged, the resistance must be low enough to allow current flow between the
collection sites of the oppositely charged particles without inducing fields
316
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comparable to those externally applied. Fortunately, it appears that these
conditions will be met unless highly insulating granules are used under very
dry conditions, so that in practice any modification of the bed field due to
the aerosol is negligible.
Second, there are two modes of current conduction through the bed:
(i) through the volume of the media and (ii) on the surface of the media.
In general both contribute to determining the potential distribution, but it
seems likely, especially in moist combustion gases, that surface conduction
predominates.
Third, as discussed below, the contacts between the individual bed
elements are critical, because the resistance is primarily determined by the
contact area, and the electric fields are concentrated around the regions of
contact. There is evidence that localized intermittent discharges occur in
the narrow gaps around the contact points. It also seems probable that these
regions of strong field play an important role in the filter action.
The basic element of the array of spehres shown in Figure 2(a) is a single
sphere with current entering at one pole, 9=0, and leaving at the opposite
one, 9 = tt, as shown in Figure 2(b). An analysis of the potential distribution
for the cases of volume conduction and surface conduction can be then given
as follows:
Volume Conduction. For a sphere of radius a, volume resistivivity p
and for a total current, I, the potential at any interior point is given2'' by
v=2l\±.±
2tt K \
2a
Jin
R (1 + cosa )
o o_
R (1 + cosa )
TT 7T
(16)
where R and R are the distances from the poles, and a and a are
O TT O 7T
the angles to the polar axis (Figure 3(a)).
For a point on the surface, r = a, we have (Figure 3(b)) R = 2a sin(0/2),
R^ = 2a cos(0/2), aQ = (tt/2)-(0/2) , = (6/2). Then °
V(a,0) =
_ _£L
Ana )sin(0/2) cos(6/2)
- Jin
sin(9/2)(l+sin(8/2)
cos(0/2)(l+cos(9/2)
(17)
Pi
4"/ra
F (6) (say).
(18)
Thus the diametral plane 0 = it/2 is the equipotential V = 0, and the
potential has a simple pole at each contact point.
> PI = PI
R -> 0 21TR 2lTa0
0 o
(19)
317
-------
Around the contact points the equipotentials are approximately spherical
centered on the contact point, and of value inversely proportional to their
radius. The function F(0) is graphed in Figure 4.
The case of a finite contact area between the two spheres can be modelled
by assuming a perfectly conducting sphere of radius R = 6 « a surrounds
the contact point. The potential across the half-sphere is then
V = —— » —— fprn
o 2tt6 2TTa0
o
where 0o is the polar angle of the contact circle.
The equipotentials and field lines appear approximately as shown in
Figure 3(c).
To calculate the field surrounding a cubical lattice of spheres, assuming
zero conduction in the surrounding space, we should solve Laplace's equation
V^V = 0, subject to the boundary conditions that V is given by eq. (17) on
the surface of the sphere and that the diametral planes are equipotentials.
However, this is rather complex. For our purposes, a satisfactory solution
is obtained by assuming that the field is given approximately by the potential
difference divided by the gap height.
Now the gap half-height (Figure 3(d)) is
h = a(l - cosG) « a(02/2), (21)
and the field in the gap at polar angle 0 is
V(0) - V(6 )
E(0'V = h— * <22)
For small angles this can be written
E(0,0 ) s
o
E 0
av o
a 2E
(0 - 0 )
o
av
(23)
This has a maximum when 0 = 30 /2 of value
o
Emax " 17 ' <2*>
o
Here Eav is the potential drop across the sphere divided by its diameter.
Thus the field enhancement around the contact region can be very large
and scales inversely as the square of the diameter of the contact region.
318
-------
For instance, for a = 3 mm (~l/4" diameter balls) and a contact radius of
6 = 0.05 mm (-0.002"), 0q ~ 1°, so that (E^x/E^) ~ 1000. For an average
applied field of 3 kV/cm, the maximum field would be 3 MV/cm.
Clearly, with such large fields, there would be a partial (intermittent)
breakdown of the gap, which would effectively increase the contact area and
lower the maximum field. For instance, if we assume the gap breaks down for
E > 30 kV/cm, then with Eav = 3 kV/cm, we should have an enhancement of
10 and 0 ~ 10°, 6 ~ 0.5 mm (-0.020").
o
Surface Conduction. Consider now a sphere, radius a, with surface
resistivity s (ohms/square) with current entering and leaving at the poles
0=0, 7T as before. The volume resistivity is assumed to be infinite.
28
The potential on the surface can be shown to be
V = tanh 1(cos9) = ^ G(0) (say). (25)
For 0 = tt/2, V = 0, on the diametral plane. For small 0
V-^£n(6/2). (26)
Thus unlike the case of volume conduction, where V has a simple pole at
0=0, the case of surface conduction gives a logarithmic pole at 0=0. The
function G(0) is plotted for comparison with F(0) in Figure 4.
If the contact is a circular area subtending a half-angle 0Q at the
center of the sphere, then the potential drop across the half-sphere is
V = tanh'
o 2tt
1 (c°seo) - - § to(r) • (27>
For two spheres in contact over the angle out to 0=0, the electric field
in the gap is given approximately by
_ [tanh "'"(secS ) - tanh ^"(secQ)]
E(0,e ) = ~ rr ^ ¦ (28)
o 2tt a(l - cos0)
For small angles this may be written
2E Me/e )
E
-------
av
^ " ae2 s,n<2/e ) ' (30)
O O
Comparing this with the case of volume resistance, it is seen that the numeri-
cal factors are similar (1/e compared with 8/27), but in the surface conduction
case the fields are smaller by the factor &n(2/0Q) in the denominator.
For instance, for 0 =1°, this factor is 4.7, so that ®max^av ~
instead of 1000 in the vo?ume conduction case. To make the comparisonVanother
way, if we require Emax^av = set incidence of discharge, then for
the volume conduction case 0 ~ 9.9° whereas for the surface conduction case
0 = 6.5°. °
o
Thus if the conduction mode is via the surface, the field enhancement is
somewhat less, and the breakdown strength of the gap will be exceeded out to
a rather smaller radius from the contact point than for volume conduction.
The foregoing analysis reveals that in an electrified packed bed, the field
is highly inhomogeneous and is locally enhanced around the contact points by a
large factor. In practice the enhancement is probably limited by the occurrence
of intermittent microdischarges which effectively increase the contact area out
to where the field is no greater than a factor -10 over the average field values
(~3 kV/cm) typically applied to the bed experimentally. Near the sphere's sur-
face, away from the contact points, the field is reduced somewhat below the
average value.
In the region of contact, the fields are directed to the surfaces, across
the gaps, whereas away from the contact points the fields are more generally
parallel to the surfaces. These facts suggest that the regions around the
contacts between the granules probably play a dominant role in the collection
of fine particles in an electrified packed bed.
The presence of intermittent localized discharges around the granule
contact points, which will give rise to a degree of gaseous ionization, might
play some role in particle collection, for instance by discharging particles.
However, at the small currents likely to be met in practice, such effects are
probably negligible.
Interaction of Particles with an Electrified Bed
The state of charge of the entering aerosol is clearly important in
determining the interaction of particles with the granular bed. This applies
not only when the bed is externally electrified, but even when it is uncharged
(grounded), because of the existence of image forces.
In general an aerosol consists of particles having a distribution of
charge magnitudes extending over both polarities, which depends on how it was
formed and its subsequent history, and is in general correlated with its size
distribution. The net aerosol charge may be characterized by the net charge/
mass ratio as determined by a sampling Faraday cage filter. However, it is
important to recognize that aerosols having little or no net charge may o ten
have significant levels of bipolar charge.
320
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To establish a rough quantitative scale for the charge magnitudes likely
to be encountered, we may use the fact that there is a limit to the charge that
a particle can carry, set by the condition that the electric field Eg at its
surface becomes large enough to cause breakdown of the carrier gas. For a
spherical particle of radius a and density p, the limiting charge is given by
qA = 47ra2eoEs, (31)
while the charge/mass ratio is
q0 3e E
& _ o s
m pa
(32)
6 3
For instance, taking E = 3 x 10 V/m, 2a = 1 ym and p = 2 kg/m , we find
~ 8 x 10~17 c, or ~s500 electron charges, while (q^/m) ~ 8 x 10~2 C/kg,
or ~ 80 yC/g. The limiting value of q a^, while the limiting value of
(q/m) ec (1/a) .
This limiting charge, set by breakdown, can be compared with the saturation
charge due to field (bombardment) charging in a monopolar corona discharge.
% - 12™2eo (m) v2 • (33)
Here E is the charging field and e is the relative permittivity of the
particle. It is clear that for Ec - 5 x 10^ V/m, typical of single stage
precipitators q^ < q^.
There is relatively little documentation of the state of charge of aerosols
produced by various processes, but it is probably true to say that, more often
than not, they are significantly charged relative to the limiting value, unless
special measures are taken to neutralize them.
For instance, when powders are dispersed in air by various means to form
an aerosol, it appears that the resulting aerosol commonly has a significant
level of bipolar charge due to deagglomeration of the powder. In addition, a
net aerosol charge may be induced by tribocharing if the particles make contact
with surfaces in the dispersal process^ >30 ^
31
Further, measurements by White using a charge analyzing probe on the fine
particle effluent from electrostatic precipitators fitted to coal-fired power
plants show monopolar charge/mass ratios ranging up to 100 yC/g. More impor-
tantly, measurements of the input aerosol, i.e., the exhaust from the combustor,
showed that while it was approximately grossly neutral, bipolar charge/mass
ratios of the order + 10-20 yC/g existed. The origin of this bipolar charge
condition, and whether it is a typical or universal condition for fly ash
emissions from coal coinbustors, is unknown, but it is clearly an important factor
to be reckoned with in designing particulate control devices.
321
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Turning now to the question of the electrical interaction of aerosol par-
ticles with a granular bed, it is clear that we should consider the possibilities
of the particles being uncharged or charged to either sign, and also the possi-
bilities of the aerosol being grossly neutral or having a net charge.
Thus there are several distince physical situations and corresponding
mechanisms that can play a role in producing electrical forces on the particles.
We discuss these in turn and then compare their relative magnitudes.
(i) Uncharged particles situated in an electric field E0 have an
induced dipole moment
M
47re fcoj E a3
o (e+2) o p
(34)
where e is the vacuum permittivity and e is the relative permittivity of
the particle (~2 for fly ash). In a uniform field there is no net force on
the particle, but in a nonuniform field there is a force MV||Eo where V||
is the field gradient along the field direction. For the purposes of estimation,
the value of V j|E^ due to the inhomogeneous field in an electrified bed may
be taken as
radius.
(E0/ac) where E0 is the average field and
Then the force is of order
a~ is the collector
F, « 4tre
(e-1)
(e+2)
f2 &3
E p
o ~x-
a
c
(35)
(ii) In the case of an aerosol having a net charge, there is a self field
E , due to the space-charge of the particle cloud with which each particle
interacts causing the cloud to disperse towards the surrounding collector
surfaces. Poisson's equation gives
-+
E
m.
£
(36)
where q is the average charge on the particles. Taking the characteristic
dimension of the interstices of the bed as of order ac we have for the self
force qE on the particles near the collector surfaces
f ~ 5-1-
2 ~ £
(37)
(iii) Charged particles also experience an image force when close to the
collector. If the bed is conducting, the image force is
4TCo (2d)2
(38)
322
-------
where d is its distance from the conductor. In the case of an uncharged
dielectric collector of relative permittivity ec, this force is modified by
the factor (e_-l) (e +D •
C V
(iv) Finally, in an external field a charged particle experiences a force
F4 - q Eq. (39)
To compare the relative magnitudes of these four forces, we must estimate
some typical magnitudes for the quantities involved. For the particle charge,
we assume it is given by eq, (33) for bombardment charging which is valid for
particles down to at least ~ lym diameter. Also, the particle concentration can
be expressed as
3L
„ (40)
4ttp a
P P
where L is the mass loading.
Using eqs. (33) and (40) in eqs. (35), (37), (38) and (39) neglecting
factors involving e which are of order unity, and putting Ec ss E , we have
for the four forces
ft)
F1 - eoEoapl^l
F„ ~
e E2a2
o o p
F„
~ £oEoa2 (
o o y
P'v P-
a v2
_E.|
d/
F, ~ e E2a2
4 o o p
(42)
(43)
(44)
It is clear that F^, the force on an uncharged particle in a nonuniform field
is smaller than F4, the force on a charged particle in an applied field by the
ratio (ap/ac) ~ 10" . Also, F3, the image force on a charged particle only
becomes comparable with F4 when d ~ a^, i.e. when the particle is within
distance of the order of the particle radius of the collector. Relative to F4,
the self force F2 has a factor (ac/ap) ~ 103 and a factor (L/pp) « 1 which
is proportional to the aerosol mass loading. For a typical value L ~ 2 gram/m3
(~ 1 grain/ft3) and pp - 2 kg/m3, the factor (L/p ) ~ 10"3 and we find F2
comparable to F^. P
323
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Thus we should expect F^ and F3 to be negligible compared with F4,
while F2 becomes comparable with F^ at high loadings for monopolarly
charged aerosols.
There are three practical situations of interest where these estimates
of electrical forces can be applied, as follows. For an electrified bed and
uncharged particles, the only force is F^; for an unelectrified bed with
charged particles, the relevant forces are Fo and also F2 if the aerosol
is monopolarly charged; for an electrified bed and charged particles the
relevant forces are F4 and also F2 if the aerosol is monopolarly charged.
From the foregoing estimates one concludes that the strongest electrical
augmentation should be observed with charged particles and an electrified bed,
and the effect should be maintained at vanishingly small dust loadings (due to
F4). Significant augmentation should be observed for monopoalrly charged
particles and an unelectrified bed, but the effect should fall off with dust
loading, i.e., the later layers of the filter become progressively less
effective. With uncharged particles and an electrified bed one should expect,
in comparison, negligible augmentation.
These general predictions are borne out by experimental comparisons of
the enhancement of collection efficiency with and without pre-charging and with
and without voltage applied to the bed, described in the next section.
ELECTRIFIED BED - EXPERIMENTAL WORK
Most of the experimental work was performed in a bench scale flow system
at throughputs of up to 40 ft^/min as described below. Some exploratory work
was also done in a test tunnel facility at throughputs up to 600 ft3/min, and
is described below.
Bench Scale Apparatus
The bench scale test rig employed for most of this work is shown in
Figure 5, and consisted of a vertical, square section (4" x 4") duct of plexi-
glass containing a fixed bed 8" high of alumina balls 1/4" in diameter. The
bed was contained between grounded metal screens at the top and bottom and
had a similar grid at the central section which served as a high voltage
electrode.
Ambient air entered via a diffuser section and was exhausted at the top
via a fan. The flow velocity was in the range 1-6 ft/sec corresponding to
pressure drops between 0.4" and 4" water across the bed. Fly ash was fed from
a screw feeder via an "Air-Vac" aspirator to disperse the ash, at grain loadings
in the range 1-20 gr/ft3.
The aspirator incorporates a coaxial sonic nozzle, fed by compressed air
at 60-80 psi, and very effectively deagglomerates the ash due to the rapid
acceleration of the ash-laden aspirated air and the presence of shocks and a
high level of small scale turbulence. In the process, it produces a significant
level of bipolar charge in the aerosol which is approximately grossly neutral.
324
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No quantitative study was made of the charge state of the aerosol, but measure-
ments with a charge analyzing probed developed subsequently, have indicated
bipolar charge levels in the range 10-20 yC/g with this type of dispersal system
used with fly ash.
Ahead of the bed was fitted a corona discharge particle-charging device,
consisting of two fine wires situated between three ground planes. By selecting
the polarity of the high voltages applied to the wires, the charger could be
operated to give positively or negatively charged particles. Also, by applying
alternating voltages, it could be used to neutralize the tribo-charged dust
delivered by the dust dispersal system.
Sampling probe stations were located immediately upstream and down
of the bed. Since the principal interest in this study concerned the collection
of submicron particles, extensive use was made of an electrical aerosol analyzer
(EAA) (TSI model 3054) to measure grain loading and efficiency. This was sup-
plemented by measurements with a cascade impactor for the size range 0.36-4.4 ym
and also a Coulter counter for the range 1-30 Vim.
The ash used in these studies was fly ash from the Black Dog power plant.
This was relatively coarse, having a mass mean diameter of - 15 ym and only
~2% by mass below 1 ym. Consequently the EAA was used with a pre-cyclone
having a Djq cut point at 2 ym, and a dilution system to prevent overloading
the instrument. The EAA indicated that there was a negligible component of ash
below ~0.06 ym after it was dispersed in the test rig. It is not known whether
the ash as delivered was deficient in such particles or whether deagglomeration
was ineffective for very fine particles. In any event, this set a lower limit to
efficiency measurements of -0.1 ym.
Experiments with Tribo-charged Ash
In an initial series of experiments, the particle charger was not used,
and measurements were made of the filter's total mass efficiency and fractional
efficiency as a function of particle size, as other parameters were varied,
notably the voltage applied to the bed, the flow velocity, and the dust loading.
Figure 6 compares the efficiency n as a function of particle size of
the electrified bed (E ¦ 4 kV/cm) with that of the unelectrified bed, at a
superficial velocity of 2 ft/sec and an input grain loading of ~4 gr/ft^. The
overall mass efficiency was 79% for the unelectrified bed, increasing to 85%
for a field of 4 kV/cm, produced by applying 40 kV dc to the central grid. It
will be noted that the unelectrified bed has an efficiency of ri ~ 80% for
particle diameters D > 3 ym, but it falls rapidly to r) -v 20% for D < 1 ym.
This is characteristic of most filters and is due to the rapid onset of col-
lection by impaction for Stokes parameters increasing through a value of order
unity. The expected increase in efficiency due to diffusive collection for
small particles is not seen here, but can be expected to occur for D < 0.1 ym.
The most significant feature displayed by the data of Figure 6 is the
dramatic enhancement of efficiency in the submicron range brought about by
application of the field. The large drop in efficiency for submicron particles
exhibited by the unelectrified bed is completely removed, and the efficiency
remains high (~80%) down to D ~ 0.1 ym.
325
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The foregoing filter efficiency measurements were made by measuring sequen-
tially with the sampling probe stationed first before the bed and then after the
bed. This procedure poses some uncertainties related to ensuring that the flow
and sampling conditions are constant and identical for the two measurements. A
simpler procedure, which avoids some of these uncertainties and isolates the
effect of electrifying the bed, is to sample sequentially with the probe down-
stream of the bed, first with no field and then with the field applied. This
simpler, quicker procedure was adopted for the measurements described below,
where various parameters were varied.
Two related measures of the enhancement of filter efficiency due to appli-
cation of the field can be defined in terms of the flux P of particles
penetrating the filter, as follows:
Electrical Effectiveness e = ^^p(E-O)^'^"
Electrical Factor of Improvement F « . (46)
In either case the penetrating flux P can be a number flux or a mass flux, and
can be for a specified size range or integrated over the prevailing distribution.
Clearly, the two measures are related by
e = 1 - (1/F) (47)
or
F - 1/(1 - e) • (48)
Thus, as the electrical enhancement increases from zero to the maximum possible,
e runs from zero to unity while F runs unity to infinity.
Figure 7 shows the effectiveness e (based on the number flux of particles
< 1 ym) as a function of applied field E (positive polarity of applied voltage)
for conditions otherwise the same as in Figure 6. It is seen that e increases
with E and asymptotes to a saturation value of e ~ 75% for E ~ 5 kV/cm.
Figure 8 shows the factor of improvement as a function of particle diameter
from measurements with the mobility analyzer, cascade impactor and Coulter counter.
There is some lack of numerical agreement among the three techniques, which is
hardly surprising in view of their widely differing principles of operation, but
the trends are in consonance. A factor of improvement in the range 3-5 is indi-
cated for submicron particles, whereas a value F ~ 1 (no significant electrical
enhancement) is indicated for particles above D - 3 ym, where the efficiency of
the unelectrified filter is already high (-80%).
Figures 9 and 10 show the effectiveness £, (based on the number flux of
particles < 1 ym) as a function of grain loading and superficial velocity,
respectively. It is seen that e is a weakly increasing function of grain
loading, while it decreases with velocity from e ~ 80% at u - 1 ft/sec to
e ~ 55% at u ~ 6 ft/sec.
326
-------
Figure 11 compares the effectiveness (based on submicron particle number
flux) as a function of field strength for steady applied voltages (of either
polarity) and a 60 Hz alternating voltage. It is clear that for the same rms
field, ac is as effective as dc, which has obvious practical consequences of
economy in large scale application.
Figure 12 compares the effectiveness, based on submicron particle number
flux, for fine Ti02 powder (0.2 ym mass mean diameter), with that for the Black
Dog fly ash, and shows that there is no significant dependence on the presence
of large particles.
Figure 13 compares the effectiveness, based on submicron particles, for
two different sizes of alumina spheres used for the bed.
Finally, in Figure 14, a comparison is made of the collection efficiency
with and without applied voltage for two different electrode geometries, namely
the original screen geometry and one in which the screens were replaced by rows
of rods (~0.1" dia). The latter configuration (with vertical rods) was expected
to be more suitable for use in a bed moving slowly down in plug flow. As
expected, for the same applied voltage, the rods yield an enhancement only
slightly less than the screens.
It should be emphasized that in all the aforementioned experiments, the pre-
charging corona discharge was not energized, though it was known that the parti-
cles were significantly tribo-charged in the ash dispersal system, and that this
was likely to be important to the observed electrical enhancement effect.
Experiments with Pre-charged and Neutralized Ash
To investigate the effect of the state of charge of the ash entering the
filter, a series of experiments were made utilizing the two-wire, three-plate,
corona pre-charger shown in Figure 5. Such a charger also acts as a single-
stage precipitator. Figure 15 shows its efficiency for submicron particles
as a function of negative corona current, measured with the EAA at the lower
sampling station by switching the current on and off. No problem was experienced
with back corona, even with the highest current densities, because the Black
Dog ash is of low resistivity and the relative humidity was high (-60%).
In Figures 16-19 the filter efficiency in terms of the number flux of
particles in the size range 0.24-1.0 ym is plotted as a function of electrical
conditions. Figure 16 shows the efficiency as a function of the current to
the pre-charger for an unelectrified bed (the electrodes grounded). It indicates
that although the efficiency increases with particle charge, it does not attain
the high values observed with tribo-charged particles and an electrified bed.
Figure 17 shows the efficiency for submicron particles with both the pre-
charger energized (to 50 \xk) and the bed electrified, as a function of field
applied to the bed. The four combinations of the precharger and bed polarities
show essentially identical results and indicate efficiencies in excess of 95%
for submicron particles at E ~ 4 kV/cm. This is higher than that observed for
the tribo-charged ash (-85%) and is also higher than the overall mass efficiency
(i.e., including the bulk of the ash above 1 pm) for these conditions.
327
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These results suggested that the nominally uncharged ash used in the
earlier experiments was, in fact, quite strongly tribo-charged bipolarly, al-
though it was grossly neutral, as indicated by measurements with a Faraday
cage total filter. To elucidate this matter, measurements were made with 60 Hz
high voltage applied to the corona charging wires. It is well known that such
an ac corona discharge contains roughly equal concentrations of positive and
negative ions, and is effective in neutralizing aerosol charges.
Figure 18 compares the efficiency for submicron particles as a function
of field applied to the bed, with the charger off and with it energized with
ac. It clearly demonstrates that the high efficiencies (~80% at E ~ 4 kV/cm)
observed with nominally uncharged particles is due to their actually being
bipolarly tribo-charged, since when the ash is charge-neutralized the efficiency
is much lower.
Finally, a comparison was made of the collection efficiencies, with pre-
charged ash, of an unelectrified insulating bed and a grounded bed of conducting
steel balls of the same size. Figure 19 shows that while the conducting bed is
somewhat better than the insulating bed, it falls far short of the electrified
insulating bed.
In summary, the foregoing results show conclusively that to obtain the
large enhancement of collection efficiency in the submicron range requires a
strong field applied across an insulating bed combined with a significant charge
on the incoming aerosol. This is in agreement with the theoretical considera-
tions discussed in the previous section.
Experiments in a Test Tunnel
To explore the concept of the electrified granular bed filter at a larger
scale, and under more realistic conditions, that might be extrapolated to a
pilot scale, a filter was constructed to fit into a wind tunnel facility built
for studies of electrostatic precipitation. This facility consists basically
of a tunnel of rectangular cross section 10" wide by 30" high, with a flow
capability of up to 2000 ft^/min, and has provisions for heating and humidifying
the air flow to simulate conditions of a power plant exhaust stack. Fly ash
can be injected and dispersed using a screw feeder and aspirator arrangement,
similar to that employed for the bench scale rig but of larger scale, at ash
loadings of up to ~10 gr/ft^.
A plan view of the filter is shown in Figure 20. The shell is constructed
of particle board and the filter media Is contained between grounded wire mesh
screens having a separation of 18" in the flow direction.
In a practical granular bed filter it is necessary to arrange for the gran-
ules to move slowly under gravity in plug flow, to allow their extraction for
cleaning and re-use. Accordingly, the electrode system was chosen in the form
of vertical rods to be compatible with bed motion. As shown, three rows of
1/4" diameter rods were used, with the center row grounded and the other two
rows connected to a bus-bar for electrical energization.
Because it was judged that the high quality alumina spheres employed in
the bench—scale rig -might be too costly for large-scale application, the use
328
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of a lower grade bed material was investigated. Alumina granules or nuggets of
a grade commonly used for grinding were employed. These were irregularly shaped,
of approximately 1/4" mean size, and of a dark grey color, in contrast to the
white alumina spheres. The color is attributed to impurities. Their suitability
as an insulating bed material was first tested in the bench-scale rig. It was
found that they were limited to lower average electric fields than the pure
alumina spheres by the onset of electrical breakdown. The maximum field was
found to be ~2.1 kV/cm, compared with ~5 kV/cm with the pure alumina spheres.
By chemical cleaning the nuggets the maximum field could be increased to ~3 kV/cm.
Figure 21 compares the efficiency based on the number flux of sub-micron
particles, as a function of average applied field, for the nuggets and the
spheres, measured in the bench-scale rig using tribo-charged fly ash. It is seen
that the efficiencies tend to the same asymptote (-85%), but that the efficiency
for the nuggets approaches saturation more rapidly with increasing field. Thus
the lower maximum field for the nuggets may not be a disadvantage in practice.
A series of experiments were then made with the filter, filled with alumina
nuggets, in the test tunnel. A single wire-plate corona charging unit was
fitted upstream of the filter, and sampling ports were provided at stations
immediately upstream and downstream of the filter.
Figure 22 shows, as a function of face velocity, the pressure drop and also
the total mass efficiency for tribo-charged Black Dog fly ash, as measured with
a total filter sampling probe. No significant difference in total mass effi-
ciency was detectable whether the bed was electrified or not. However this is
to be expected, because the enhancement with an electrified bed occurs for the
submicron fraction which comprises only 1-2% by mass of the fly ash. The rather
rapid decrease in total mass collection efficiency for face velocities greater
than 1-2 ft/sec must be attributable to re-entrainnment of impacted ash from
the granules.
The electrical enhancement effects observed for submicron particles are
shown in Figure 23, where the efficiency for submicron particles is plotted
against face velocity for three cases; tribo-charged ash with and without the
bed electrified, and with corona-charged ash with the bed electrified at
1.2 kV/cm, dc. In the latter case, the submicron efficiency was 90-95% for
face velocities of 1-2 ft/sec.
During these experiments the ambient relative humidity varied by about 10
percentage points and it was observed that the bed current was a sensitive
function of humidity. At 60% relative humidity the filter drew 4 mA at 15 kV
(1.2 kV/cm), falling to 1.2 mA at RH ~ 55%. This supports the view that the
bed resistance is controlled by the surface resistivity rather than the bulk
resistivity of the granules.
To investigate this effect, the gas temperature was varied up to 150°F
by operating the natural gas burner in the tunnel test facility, which reduced
the relative humidity considerably. At this temperature the bed current was
reduced to the very low value of -10 ]iA at 1.2 kV/cm field. The current tended
to be unstable under these conditions, and it appeared as though the filter were
charging up and then breaking down to leak off the accumulated charge.
329
-------
The electrical effectiveness for submicron particles is compared in
Figure 24 for ambient and hot-flow conditions, Tribo-charged dust was used
with a filter field of 1.2 kV/cm in both cases. It is seen that under hot
conditions, the effectiveness is reduced by a factor ~2. It is not clear
whether this is due entirely to the change in bed resistance, or whether it
Involves also the ash resistivity which is also a strong function of tempera-
ture and humidity. From resistivity cell measurements it is estimated that the
Black Dog ash resistivity increases from ~109 ohm-cm at room temperature to
-1011 ohm-cm at 105°F. The hopper of the screw-feeder was heated in all the
experiments reported, so that the state of charge of the dispersed ash should
be independent of the main tunnel gas temperature, although its resistivity
can be expected to rapidly assume that corresponding to the tunnel gas tempera-
ture and humidity. Thus the principal effect of ash resistivity can be expected
to be on the bed resistance, as the bed accumulated dust, especially around
the contacts between granules.
O
For a face velocity of ~2 ft/sec, a grain loading of ~5 gr/ft and a bi-
polar charge/mass ratio of + 20 yC/g, the bipolar current intercepted by the
bed is ~+20 yA. This Is comparable with the external current through the bed,
and it seems probable that the reduced effectiveness at the elevated temperature
is due to the high bed resistance. Under these conditions, the bed fields due
to the applied voltage are reduced or biased off due to charge deposited by the
captured ash, which cannot leak off fast enough.
This is a critical point for operation of a granular bed in stack gases
and needs further investigation.
DISCUSSION
The principal result of this work is that the collection efficiency of a
granular bed filter for submicron particles can be dramatically enhanced by
electrical augmentation. The bench-scale experiments and theoretical considera-
tions indicate that electrification of an insulating bed to average fields of
a few kV/cm, together with significant levels of particle charge are necessary
to achieve the maximum effect, where the usual rapid fall off in efficiency of
granular beds for submicron particles is completely removed.
The potential of electrified granular bed filters for large-scale applica-
tion to particulate control will depend on an engineering evaluation of a number
of factors not fully addressed in this preliminary study. These include the
trade-offs between the face velocity, pressure drop and efficiency loss due to
re-entrainment, the problems and cost of auxiliary equipment for recycling
the bed material, efficiency loss associated with relative motion of the granules
in a moving bed, and optimization of the bed resistance to avoid, on the one hand,
unnecessarily large power consumption, and on the other, electrical biasing of
the applied field with too high a bed resistance.
Its potential will also clearly depend on the particular application and
how it is integrated with other equipment. For conventional coal-fired power
plant applications, it might be used with an inertial pre-separator or electro-
static precipitator to remove the bulk of the coarse fly ash fraction, which
330
-------
would greatly simplify the bed cleaning problem. In this sense it might be
attractive as retrofit equipment to control submicron emissions on existing
devices which do not meet the impending stricter requirements for such emissions.
It may also have potential for hot gas clean up for coal-fired gas turbines, if
suitable bed materials are available with low enough resistivities at high
temperatures.
The electrified bed filter has a number of inherent advantages over various
competing devices. It is rugged and fireproof so that, compared with bag filters,
it is immune to boiler upset. Moreover, it works equally well with ac and dc
voltages, which implies a considerable saving in power supply costs relative to
precipitators.
An intriguing and potentially important possibility, is that the ash emitted
by combustion systems may be sufficiently highly bipolarly charged that one might
eliminate the need of a pre-charging stage, with a consequent saving in complexity
and cost. The evidence on this point^l is inadequate, and further work is neces-
sary to characterize the charge state of the particulate emissions from common
combustion systems and other industrial processes. However, the fact that dra-
matic reductions in plume opacity have been reported25 0n hog-fuel boilers with
electrification of the granular bed indicates that the emissions were most
likely significantly charged in that case.
ACKNOWLEDGEMENT
It is a pleasure to record the useful advice and contributions of
Prof. M. Mitchner, Dr. L. Collins and Mr. E. Kushner to this work. This work
was supported by the Electric Power Research Institute.
REFERENCES
1. Pilat, M. J., "Collection of Aerosol Particles by Electrostatic Droplet Spray
Scrubbers," J. APCA, 25, 176-178 (1975).
2. Nielson, K. A. and Hill, J. C., "Capture of Particles on Spheres by Inertial
and Electrical Forces ,"I&EC Fund., 15, 149-157, (1976).
3. Bergman, W., Taylor, R. D., Miller, H. H., Bierman, A. H., Hebard, H. D.,
daRosa, R, A., and Lum, B. Y., "Enhanced Filtration Program at Lawrence
Livermore Laboratory—A Progress Report," 15th DOE Air Cleaning Conference,
Harvard, August 1978 (Preprint UCRL-81512).
4. Lamb, G.E. and Custanza, P.A., "Electrical Stimulation of Fabric Filtration,"
Textile Research J., 47, 372-380 (1977).
5. Zahedi, K. and Melcher, J. R., "E.lectrofluidized Beds in the Filtration of
a Submicron Aerosol," J. APCA, 26, 345-352 (1976).
6. Zahedi, K. and Melcher, J. R., "Collection of Submicron Particulate in
Bubbling Electrofluidized Beds," I&EC Fund., 16, 248-254 (1977).
331
-------
7. Alexander J. C. and Melcherf J. R,, "Alternating Field Electrofluidized
Beds in the Collection of Submicron Aerosols," I&EG Func., 1£, 311-317
(1977).
8. Melcher, J. R., Sachar, K. S. and Warren, E. P., "Overview of Electrostatic
Devices for Control of Submicrometer Particles," Proc. IEEE, 6^5, 1659-1669
(1977).
9. Spielman, L. A., "Particle Capture from Low-Speed Laminar Flows," Ann. Rev.
Fluid Mech., 9, 297-319 (1977).
10. Tardos, G. I., Abuaf, N. and Gutfinger, C., "Dust Deposition in Granular
Bed Filters: Theories and Experiments," J. APCA, 28, 354-363 (1978).
11. Schmidt, E. W., Gieseke, J. A., Gelfand, P., Lugar T. W., and Furlong, D. A.,
"Filtration Theory for Granular Beds," J. APCA, 28, 143-146 (1978).
12. Fuchs, N. A., The Mechanics of Aerosols, Pergamon Press, New York, 1964.
13. Friedlander, S. K., Smoke, Dust and Haze, Wiley-Interscience, New York,
1977.
14. Ranz, W. E. and Wong, J. B., I&EC, 44, 1371 (1952).
15. Paretsky, L., Theodore, L., Pfeffer, R., and Squires, A. M., "Panel Beds
for Simultaneous Removal of Fly Ash and Sulphur Dioxide: II Filtration of
Dilute Aerosols by Sand Beds," J. APCA, 21, 204-209 (1971).
16. Langmuir, I., J. Meteor., J5, 175 (1948).
17. Levich, V. , Physico-Chemical Hydrodynamics, Acad. Sci., USSR, Moscow, 1952,
English Translation, Prentice-Hall (1962).
18. Friedlander, S. K., "Mass and Heat Transfer to Single Spheres and Cylinders
at Low Reynolds Numbers," J. AICHE, 3, 11 (1957).
19. Aksel'rud, G., Zh. Fiz. Khim., 22, 1445 (1953).
20. Happel, J. and Brenner, H., Low Reynolds Number Hydrodynamics, Prentice
Hall (1962).
21. Pfeffer, R., "Heat and Mass Transfer in Multiparticle Systems," I&EC Fund.,
2, 380 (1964).
22. Payatakes, A. C., Tien, C., and Turian, R. M., J.AICHE, 19, 58, 67, 1036,
(1972).
23. Goren, S. L., "Aerosol Filtration by Granular Beds," Symposium on Transfer
and Utilization of Particulate Control Technology, Denver, Colorado,
24-28 July, 1978, Session C3.
24. Geffken, J., Guillory, J. L. and Phillips, K. F., "Performance Character-
istics of Moving-Bed Granular Filters," Ibid., Session C3.
332
-------
25. Reese, R. G., "The Electroscrubber, An Electrostatic Granular Bed Filter,"
Combustion Power Company, Menlo Park, Calfiornia.
26. McLean, K. J., "Cohesion of Precipitated Dust Layer in Electrostatic Pre-
cipitators," J. APCA, 27, 1100-1103 (1977).
27. Smythe, W. R., Static and Dynamic Electricity, 2nd ed., McGraw-Hill (1950),
ch. 6, p. 239.
28. Ibid., p. 243.
29. Stay, P. R., "Role of Gas Breakdown in the Charging and Discharging of Macron
Clouds," Proc. IEE, 120, 523-526 (1973).
30. Cheng, L. and Soo, S. L., "Charging of Dust Particles by Impact," J. Appl.
Phys., 41, 585-591 (1970).
31. White, H. J., Industrial Electrostatic Precipitation, Addison-Wesley, (1963)
p. 152, also private communication 1979.
32. Self, S. A., Paul, P. and Kushner, E., "A Charge Analyzing Probe for Aerosol
Plows," IEE-IAS Conference, Cleveland, October 1979 (to be presented).
333
-------
Air ot STP
ac * I mm
- 2
-3
'inf(v)
-8
0.01
0.1
»*I06 to7 3 * I07 10® 0.1 0.» I 3 10
Pt Btk
Figure 1. Theoretical collection efficiency by isolated
sphere for potential(p) and viscous(v) flow.
(0)
(b)
Figure 2. (a) Cubic array of spheres between plane electrodes
(b) Sphere with current entering and leaving at opposite poles.
334
-------
(«)
Equl-
Potentials
(d)
(c)
Figure 3. Geometry for current flow through resistive spheres,
(a) volume conduction, (b) surface conduction,
(c) potential and field distribution, and
(d) geometry in contact region.
F(e)
G(»)
Figure 4. Functions FC9) and G(0) for potential distributions
for current flow across a sphere.
335
-------
&5 Th. copperplofes
r jOO^dio. wire®
j'l'.u
Wife ondPlote
Preeharger
8»Di1tu»er
¦To flow volve ,fon ond
vent system
Velocity Probe
•TSI Sample Probe
(otter filter)
4"* 4"* 8* high granular
bed filters
(shown with 3electrodes)
TSI Somple Probe
(before filter)
Screen for Turbulence
Control Mixing
Oust Injection-
(Sonic Nozile Screw Type Feeder)
Room Air
Figure 5. Bench scale test apparatus.
77%
100 -
80
60
40
20
4J0 6.0 8010.0
40
Figure 6. Filter efficiency as a function of particle size, combined results
Coulter Counter (> 1 ym) and Electrical Aerosol Analyzer (< 1 ym).
336
-------
Filter Effeetiveneee oi e (unction of
Applied Field; Poeftive Power
Supply Polority
Z 40
UJ 30
Figure 7.
2 3 4
E , kv/cm
Filter effectivenss, E, as a function of applied
field strength.
(F)
i i i i i m
• Couotie Impoefor
• Electrical Aeroiol Anolyter
• Coulter Counter
•I -2 .4 .6 .8 1.0 2.0 4j0 6.0 8.0 tO.O
OIA. fitn
Figure 8. Factor of improvement (F) in collection of fine particles with
applied field of A kV/cm as a function of particle size as
measured with the TSI Electrical Aerosol Analyzer, the Cascade
Impactor and the Coulter Counter,
337
-------
5 10 15 20 25 30
LOADING, gr/ft3
12 3 4 5
VELOCITY, ft/sec
Figure 9. Filter effectiveness (e) as a Figure 10. Filter effectiveness (e) as
function of dust loading using a function of superficial
an applied field of 4 kV/cm velocity,
with a free stream velocity of
2 ft/sec.
€%
100
80
40
• AC(RM3),(2t»$ft)
Q Oe (2f»$t»)
20
E KV/cm
100
€ %
80
60
o Titanium Oioxide
% "Block O09" Flya»h
E, KV/cm
Figure 11. Comparison between AC (BMS) Figure 12. Comparison between filter
and DC operation of similar effectiveness (e) as a
filters under similar condi- function of field strength
tions. Effectiveness (e) vs. observed with submicron ti-
applied field. tanium dioxide and Black Dog
Power Station fly ash.
338
-------
• 1/8 dio Alumino Sphere)
~ 1/4" dio Alumina Sphtras
E, kV/cm
Figure 13. Effectiveness (e) as a
function of applied field
with two different filter
media sizes.
60r—
2.0
• E » 0 (Rod*)
o E 1 0 Iscritna. average
o< ZfisW>
»E'4 (Rods)
o E» 4 Uoteri»,ev#roj8.
of 2 tesflj
.2 .4 .6 .81.0 2.0
PARTICLE DIAM. (microns)
Figure 14. Filter efficiency, a compari-
son between rod and screen
electrodes as a function of
particle size.
60
60
§
o
C
u.
iu
tt
IU
h.
40
20
_L
J.
0 50 100 150 200 250
NEGATIVE CORONA CURRENT (*iA)
Figure 15. Charger's collection effi-
ciency for particles
between 0.24 and 1.0 yim
diameter.
O 60 100 150 200
NEGATIVE CORONA CURRENT (/iA)
Figure 16. Filter collection efficiency
as a function of precharger
current. Electrodes in fil-
ter grounded. Size range of
measurement - 0.25->1.0 ym.
339
-------
£ 80
O Chorger ond filter Negotive
a Chorger and filter Positive
A Positive Chorger, Negative Filter
x Negotive Chorger, Positive Fitter
I
0 1 2 3 4 5
APPLIED FIELD ON FILTER (kV/cm)
Figure 17. Filter efficiency as a function of applied field
with precharged fly ash. Corona current = 50 yA.
Size range of measurement = 0.24 -* 1.0 lim.
100
O»l0 kV AC on Chorger
x»Chorger turned off
0 1 1 1 '
0 12 3 4
APPLIED FIELD ON FILTER (kV/cm)
Figure 18. Efficiency as a function of applied field—
comparison between neutralized and tribo-
charged fly ash. Size range of measurement ¦
0.24 ¦* 1.0 microns.
340
-------
100
£ 80
>-
o
z
lii
o
E
Ul
a:
Ui
Figure 19.
60
40
20
o Chrome Steel Medio (Grounded)
x Alumino Medio (Grounded Electrodes)
x
o
0 50 100 150 200 250
NEGATIVE CORONA CURRENT (fiA)
Filter collection efficiency as a function of
charger current. Comparison between steel and
alumina media (0.25"). Size range of measurement,
0.24 + 1.0 ym.
Media lltlalning Screen
owl «ml electrode
FLOW
/ H.V. Connector
Locoticn of vertical electrode* OA" dlo rodt)
Figure 20. Top view of granular bed filter portion of test section
for ESP test facility 0-8" long in flow direction,
29.5" high, 9.84" wide) shows location of vertical elec-
trodes used i provide electric field.
341
-------
1
0
2
3
4
E kV/cm
Figure 21. Comparison between alumina spheres and Aluminum Oxide
Nuggets in bench scale test apparatus—submicron
efficiency vs. applied electric field.
15
10
8
12 3 4
FREE STREAM VELOCITY ft/sec
St
IP
100
80
60
40
20
0
Figure 22a. Pressure drop as a function
of free stream velocity in
the granular bed filter in
the ESP test facility.
2 3
U ft/tec
Figure 22b. Mass efficiency of granular
bed filter in ESP test
facility as a function of
velocity.
342
-------
X Filter Only, No Fields
• Filter With 1.2 kV/cm Field
o Filter With Precharger +1.2 kV/cm Field
?%
J
0 12 3 4 5
U. ft/sec
Figure 23. Comparison between baseline filter performance and
electrified performance in ESP test facility for
submicron particles. Efficiency vs. velocity.
X Ambient Conditions
• Hot (150*F)
.8
FILTER FIELD
Figure 24. Filter performance in test facility with heated airstream.
Shows effect of dust resistivity, hence bed current upon
electrical enhancement effect. No precharger.
343
-------
THEORETICAL AND EXPERIMENTAL FILTRATION EFFICIENCIES IN
ELECTROSTATICALLY AUGMENTED GRANULAR BEDS
G.A. Kallio, P.W. Dietz, and C. Gutfinger*
General Electric Company
Corporate Research & Development
Schenectady, New York 12301
ABSTRACT
Filtration efficiencies for fine particles by electrically-enhanced gran-
ular beds are computed and compared with available experimental data. Collec-
tion efficiencies of a single bed sphere in a "unit cell" are calculated by
numerically solving the equations of particle motion which incorporate the
effects of inertial impaction, interception, and charged particulate-imposed
electric field interaction in both Stokes and potential flows. The "unit
cell" model enables one to include the effects of surrounding bed particles
on the collection process. Total efficiency is computed by integrating the
single-sphere collection efficiency over the bed volume. These calculated
total bed efficiencies compare satisfactorily with data obtained from an ex-
isting granular bed facility. Both theory and experimental data show that
the electrical force significantly enhances particulate collection in granular
beds.
•Faculty of Mechanical Engineering, Technion - Israel
Institute of Technology, Haifa, Israel
344
-------
THEORETICAL AND EXPERIMENTAL FILTRATION EFFICIENCIES IN
ELECTROSTATICALLY AUGMENTED GRANULAR BEDS
I.INTRODUCTION
Increasing attention is being directed toward the emission of fine par-
ticles because of their potential responsibility for adverse health effects.
Since fine particles (below 5 - 10 Mm in diameter) are predominantly formed
by condensation, their chemical composition is often distinctly different
than that of coarse particles which are usually formed by mechanical proces-
ses. For example, studies have shown that certain heav^ metals are more con-
centrated in fine particles (Natusch and Wallace (1974) ). Because these
fine particles pass through conventional pollution control devices uncollected
and are also the most likely to be deposited in the respiratory tract, the
potential for serious health problems is clearly established and air quality
standards may soon be promulgated which require removal of these particles.
One device which has the demonstrated ability to remove particles as
small as 0.2 ym in diameter is the electrostatically-enhanced granulac bed
filter or ele^trofluidized bed (EFB) (Zahedi and Melcher (1976, 1977) * 5
Zahedi (1976) )• The fundamental mechanism responsible for this high col-
lection efficiency can be easily understood. Before entering the EFB, the
fine particles are electrically charged. Upon entering the bed, the parti-
cles experience an electrical force due to the electric field between the
particles. Essentially, the bed particles act like the plates of miniature
electrostatic precipitators (ESPs). As can easily be seen from the classic
Deutsch equation for ESP performance, the consequence of many small ESPs is
high efficiency.
Numerous authors (Whipple and Chalmers (19M)"*; Kraemer and Johnstone
(1955) ; George and Poehlein (1974)'; Nielsen and Hill (1976) ) have studied
the electrostatically-enhanced collection of fine particles on isolated spher-
ical collectors. In 1976, Zahedi proposed a model for the collection of
charged fine particles in a granular bed with an imposed electric field.
This model assumed that each particle in the granular bed acted like an iso-
lated sphere and, thus, its collection could be modeled by the Whipple and
Chalmers result. Recently, Dietz (1979) developed an analytic model for
the collection of fine (inertialess) particles in a granular bed with an im-
posed electric field. This model is based on the identification of a "single
filter element" or "unit cell" which oonsists of a single bed particle and
the adjacent gas. The flow in the vicinity of this element is then calculated
and single element efficiencies are computed. The overall efficiency of the
granular bed can thus be obtained by integrating over all of the collecting
bed particles.
This general approach to ^odeling the flow through granular beds has
been employed by Happel (1958) and Kuwabara (1959) to predict the pressure
345
-------
drop througb^granular beds, and by Gutfinger, Tardos, and Abuaf (197^, 1978,
1979) ' to predict the mechanical collection of fine particles by gran-
ular beds.
In the present paper, the concept of a unit-cell is employed to predict
the collection of inertial particles in an electrified granular bed. The
differential equation describing the motion of fine particles is numerically
integrated based on assumed fluid flow models. The resultant particle tra-
jectories are used to compute single element efficiencies and, subsequently,
overall bed efficiencies.
The results of these calculations are compared with data obtained by
Alexander (1978) . In conclusion, model extentions are suggested future
work.
II. THEORY
A. Equation of Motion
The equation of motion for a solid, spherical, fine particle of radius
r and mass density p in a gas flow can be approximated as
r r
r 3P„ = 6irt* (U - V) + Fe (1)
3 p p dt p
where U = local gas velocity
V = fine particle velocity
M. = gas viscosity
F = arbitrary electrical force
The assumptions and considerations incorporated into Eq. (1) are:
1) The mass of the displaced gas is negligible;
2) Stokes drag law is valid (Re < 1);
3) The fine particles do not perturb the gas streamlines;
4) The fine particles are large enough that mass diffusion, or Brownian
movement, can be neglected;
5) Gravity is neglected.
B. Bed Model
The granular bed is modeled as a homogeneous assemblage of equal "cells";
each cell consisting of a bed particle of radius a surrounded by a concentric
shell of gas with radius & (Tardos, et al, (197*0 ). See Figure 1. The
effect of void fraction upon filtration is incorporated into the model by
prescribing the shell radius i to be such that the void fraction within the
346
-------
Uo, Eo O
BED
PARTICLE
Figure 1. Unit cell model for granular bed.
cell equals the average void fraction a of the bed:
I = a(l-a)"1/3 (2)
In this way, the effect of neighboring bed particles upon the collection rate
can be modeled. Also, this approach establishes a boundary at which initial
conditions can be set without having to specify the location of all the bed
particles.
C. Gas Flow Models
__ The gas flow throughout the bed is assumed laminar. Two velocity fields
(U) are used in Eq. (1): Stokes flow (flea << 1) and potential flow (zero
vorticity, Re » 1) for which the corresponding a3Ctsymmetricfistream func-
tions have been derived by Kuwabara (1959) and Lamb (1915) :
Stokes flow
s (A/r + Br + Cr^ + Dr**) sin^Q (3)
347
-------
where
A =
-U a3 j 3
0 (1 - ^r)
B =
C =
D =
4k
3V
4k
U
2k
(1
5&"
a*
2lJ
JV
20 kS,3
. . 1 ,av6
and k = - (-g)
Potential flow
5(|)6 _ g(A)5 + 5(|)3 _ 1
a a a
, "Uo , A3
= 2 A3-a3
) (r^- —) sin^0
r
(4a)
(4b)
(4c)
(4d)
(4e)
(5)
Figure 1 displays the coordinate system employed in the analyses. The
components of the gas velocity are given by
1
Ur = " . 2 .
u„ =
r sin0
1
36
14
0 rsin0 3r
(6a)
(6b)
D. Electrical Force Model
,Ml .Q
The electrical force (F ) is due to a uniform external electric field
of strength E interacting with electrically charged fine particles and neu-
tral bed particles. A voltage is applied across the granular bed between
two parallel grids to create the field which is defined to be collinear with
the superficial gas flow. The imposed field polarizes the bed particles (i.e.,
induces a charge separation) to produce an attractive force between the charged
fine particles and the leading hemispheres of the bed particles. The follow-
ing assumptions apply:
1) The bed particles are semi-insulating, and retain a zero net charge;
2) Space charge effects due to the charged fine particles are negligible;
3) The image force between charged particles and neutral collectors is
neglected;
4) Dipole interactions are neglected.
The particulate efflux is charged by a high voltage corona source prior
to entering the bed. For this analysis, the fine particles are assumed to
348
-------
17
be conducting and charged to saturation (Melcher and Sachar (1974) ), thus:
«p = 12TCorpEc m
where q is the charge acquired by one particle and E is the corona charging
electriB field strength. Therefore, the total electrical force experienced
by a fine particle is
F® = qp E(r,0 ) (8)
where E(r,0) is the local (intensified) electric field within the bed.
The derivation of E(r,0) for the cell model is analagous to that of the
gas flow field. Specifying the potential, <&, on the surface of the unit cell
to be the average potential at that position, the potential within the cell
can be computed:
E
* = 3+ot(K-1) (K+2)r - (K-1)r~2 cos6 (9)
where K = The corresponding electric field components are given by:
Er = " I? <10a>
% = - 11$ (10b)
r 36
E. Normalization
The mathematics can be more easily managed by normalizing all distances
with respect to a, all velocities with respect to U , and the electric field
with respect to EQ, for example °
V
U
—*
= UV
°_*
= UoU
(11a)
(11b)
E
: EE*
0
(11c)
r
= ar*
(lid)
therefore
t
= at*/U
o
(lie)
Eq. (1) can now be written concisely as two coupled second-order differential
equations to describe the trajectory of a fine particle:.
Radial
dv *
"dti s
-------
Tangential
dvQ*
= (u0* - vQ*)/St + GEe* (12b)
P r2U
where st = | p P ° (13)
9 ya
£ aE E
G = 9 ° °?° (14)
p r U
P P o
The Stokes number, St, is a dimensionless quantity which relates the
ratio of the fine particles inertia to the viscous drag it experiences. There-
fore, a large, dense particle would correspond to a high Stokes number. The
dimensionless electrical parameter, G, relates the ratio between the electrical
force and inertial force on the fine particle. These two parameters play
an important role in describing the particles' trajectory.
Ill METHOD OF SOLUTION
A. Initial Conditions
In order to compute the trajectory of a given fine particle approach-
ing a bed particle, initial velocity conditions must be imposed at the cell
boundary r = &. Here the velocity of a fine particle is assumed to be equal
to the sum of the local gas velocity and the velocity due to electrical at-
traction:
V = U + bE at r = A (15)
where b is the fine particle's electrical mobility, defined as
b = 6fTyrp
This statement is equivalent to neglecting acceleration in Eq. (1). Incorpo-
rating normalized quantities, Eq. (15) becomes:
V* = U* + (St)(G)E# at r = I (17)
B. Collection Efficiency
Collection, or capture, of a fine particle is assumed to occur when it
contacts the surface of a bed particle. Thus, since the fine particles are
of a finite size, any particle whose trajectory passes within one radius (r )
of the surface of the bed particle is collected. The "critical trajectoryl,p
350
-------
is defined to be that trajectory which just grazes the bed particle. This
trajectory is characterized by the initial height, y , at which it enters
the cell (see Figure 1). All fine particles enterin§ the cell within y are
collected (a "capture window"), while those entering beyond y are swep£ around
the bed particle. Since the particulate flux is assumed to bS uniform upon
entering the unit cell, the single-sphere collection rate, T, can be computed
by
r = "oV'o2 (18>
where n is the inlet concentration density of the fine particles. A single-
sphere collection efficiency, n», can be defined as the ratio of this col-
lection rate to the total rate at which particles would pass through the area
occupied by the bed particle in its absence. Thus
n' = r/nUna2 = y 2/a2 (19)
OO C
Then, the single-sphere efficiency can be directly computed once the critical
trajectory height, y , has been found from the governing differential equa-
tions (Eqs. (12a) ana (12b)). Since these equations are highly non-linear,
they have been numerically integrated on the computer using a Kutta-Merson
predictor-corrector routine.
Figures 2, 3 and 4 show computer-drawn trajectories (pathlines) of fine
particles approaching the leading hemisphere of the bed particle. Each figure
demonstrates a particular collection mechanism that has been accounted for
in the efficiency calculations. In Figure 2, no electrostatics are applied
and the fine particles are assumed inertialess, so their trajectories iden-
tically follow the gas flow streamlines. Collection can only occur in such
a case when the fine particles follow streamlines which pass within one radius
(r ) of the surface of the bed particle. This mode of collection is termed
"interception" and dominates at low Stokes numbers. Note that the subsequent
capture window is quite small. Figure 3 shows how the trajectories of fine
particles with finite inertia deviate from the gas flow. Such collection
is called inertial impaction, which generally leads to greater efficiencies
than interception. In Figure 4, an external electric field is applied.
Charged particles will follow these attractive trajectories in the absence
of gas flow. Note that the capture window is large, resulting in single-sphere
collection efficiencies that may exceed 1.0.
Once the single-sphere efficiency has been determined for a particular
situation, the collection efficiency of the entire bed, H, can be computed.
An expression for T1 is derived by equating the change in particulate flux
across a thin bed layer to the collection rate of the enclosed bed particles
as determined by n1 (Dietz, (1979) )• Integration of the resulting differ-
ential equation over the bed length L yields
n = 1 - exp ^ -3(fca) n'L j (20)
351
-------
OAS FLOW
STRKAMLINES
BED
PARTICLE
Figure 2. Trajectories of inertialess fine particles for Stokes flow.
INERTIAL IMPACTION
PATHLINES
BED
PARTICLE
Figure 3. Trajectories of inertial fine particles for Stokes flow.
352
-------
ELECTROSTATIC
PATHLINES
BED
PARTICLE
Figure 4. Trajectories of charged fine particles
with an applied electric field.
IV. RESULTS AND DISCUSSION
A. Theory
Figures 5 and 6 display single-sphere collection efficiencies as a func-
tion of Stokes number for both Stokes and potential flow models. The effec-
tiveness of the various collection mechanisms — interception, inertial im-
paction, and electrostatic attraction — can be seen. The ratio r /a indicates
the extent to which interception plays a role. For r /a = 0.001, correspon-
ding to the lower curve in both figures, interception is essentially non-
existent and thus inertial impaction is the only collection mode. When
r /a = 0.05, interception does indeed become significant — being dominant
a? low St as shown by the middle curves. One might expect lowest efficiency
when St -* 0, however, a minimum is observed between St = 0.01 and 0.1. For
these Stokes numbers, computer results show that a fine particle traveling
along the critical trajectory closely follows the gas streamlines as it starts
to accelerate around the bed particle. However, at some point beyond the
inflection of the trajectory, the fine particle's inertia becomes great enough
in the accelerating gas flow that it deviates from the gas flow and moves
away from the bed particle. Consequently, lower collection efficiencies are
predicted.
353
-------
G =
STOKES FLOW
MODEL
O 0.8
o = 0.5
I 0.6
55 0.4
/ a = 0.05
G =
0.2
001
1
10
100
STOKES NUMBER (St)
Figure 5. Single-sphere efficiency vs. Stokes number for Stokes flow.
g =
POTENTIAL FLOW
MODEL
O 0.8
a = 0.5
I 0.6
.001
0.2
10
100
STOKES NUMBER ( St)
Figure 6. Single-sphere efficiency vs. Stokes number for potential flow.
354
-------
The upper curves of Figures 5 and 6 show the significant enhancement
that an external electric field can incur upon the collection of charged par-
ticles. The degree of electrical attraction is quantified by G, where G = 1
represents a typical situation. Note that n' exceeds 1.0 for large values
of the Stokes number. Also note that the minimum discussed previously is
essentially masked out by the electrostatic effects.
The single-sphere efficiencies obtained for potential flow are consis-
tently higher than those using Stokes flow. This is due to the fact that
potential flow predicts higher local gas velocities in the vicinity of the
bed particle. Consequently, both inertial impaction and interception are
increased. Therefore, collection efficiencies are sensitive to the choice
of flow solution, with Stokes flow and potential flow setting the lower and
upper limits, respectively, to efficiency. Stokes /low should be applied
where Rea < 1 and potential flow applied where Rea > 50. Intermediate Reynolds'
numbers are better approximated by Oseen flow and other solutions (Gutfinger,
et al (1979) ).
Since St, G, and r /a contain common terms, Figures 5 and 6 are difficult
to use to predict the sinsitivity of collection efficiency to certain param-
eters. Figures 7 and 8 show total bed efficiency (Eq. (20)) plotted versus
fine particle radius with and without an applied electric field. Here, bed
parameters have been specified and trends can be better observed. Besides
fine particle radius and electric field strength, the dependence of collection
efficiency upon bed particle radius and superficial gas velocity has been
shown also. In both figures, the potential flow solution has been used to
model the local gas velocity. Reynolds number (Re ) ranges from 1 to 50 for
the bed values specified. Therefore, the efficiencies presented in Figures 7
and 8 can be considered "best-case" predictions based upon the flow model.
There are several important trends that these plots reveal:
1) Predicted efficiencies are 0.99+ for fine particles with radii greater
than 2 ym.
2) Below 2 ym, the unelectrified bed shows drastically reduced efficiencies.
This drop is characteristic of such beds due to the rapid onset of
collection by inertial impaction for St> 0.1.
3) Collection efficiency is dramatically enhanced in the submicron range
by the application of an electric field and the charging of fine par-
ticles. In fact, 0.99+ efficiency is predicted throughout the 0.1 ym
to 10 ym 3ize range for moderate-to-high field strengths.
*0 A bed consisting of smaller particles is more efficient.
5) A lower superficial gas velocity enhances collection in an electro-
statically-dominated bed due to an increased residence tim^ of the
charged fine particles (note that G is proportional to Uq~ ).
355
-------
~ 0.8
ft 0.6
t 0.4
a = 0.5 mm
Un = 1,0 m / sec
0.5
PARTICLE RADIUS ( rp J,
Figure 7a. Overall bed efficiency vs. fine particle radius.
En = 105VI m
It 0.6
_i
f 0.4
0.2
0.1 mm
« 1m/sec
0.5
PARTICLE RADIUS (r.), ftm
Figure 7b. Overall bed efficiency vs. fine particle radius.
356
-------
c-
0.8
u, 0.6
P 0.4
0.2
a = 0.5 mm
10
0.5
PARTICLE RADIUS (rp),
Figure 8a. Overall bed efficiency vs. fine particle radius.
ov/m
0.8
0.6
P 0.4
0.2
mm
1 m/sec
0.1
0.5
1.0
5
10
PARTICLE RADIUS (rp),
Figure 8b. Overall bed efficiency vs. fine particle radius.
357
-------
6) A lower gas velocity in a mechanically-dominated bed (inertia and
interception) usually decreases collection efficiency; however, if
the bed is operating near the minimum in Figure 6, an increase in
efficiency may be obtained.
Parameter values used in Figures 7 and 8 are:
B. Data Correlation
Figure 9 displays electrofluidized bed data taken independently at MIT
by Alexander (1978) . In these experiments, monodisperse aerosols of liquid
dioctylphthalate (DOP) were collected in granular beds of New Jersey sand
and efficiencies were measured with a TSI mass monitor. The use of a liquid
aerosol insured adhesion to the bed particles, thus, no reentrainment could
occur. Uniform DOP droplets were generated by nebulization and optically
sized. The droplets were charged by ion impact and electrical mobilities
measured by a small, laminar-flow ESP. An electric field was applied col-
linear to the fluid flow with a series of screen electrodes placed across
the bed.
L = 10 cm
a = 0.5
E = 5•10 V/m
Pp = 2-103 kg/m3
y = 2*10 kg/m"see
E0 = 4 • 1 05V / m
~
DATA :
o E0 = 0 V / m
• E0 = 2 • 104
~ E0 = 4 • 104
* E0 = 4 • 106
0.1
0.5
1.0
PARTICLE RADIUS (rp ),
5
10
Figure 9. Alexander's EFB data correlation.
3S8
-------
The efficiency data shown in Figure 9 is plotted versus fine
(aerosol) radius as in Figures 7 and 8 with applied electric field
eter. The experimental EFB was typical of the parameters expected
systems:
a =
0.5 mm
U =
1 m/sec
L° =
10 cm
a =
0.5
K =
5 „
II II
a0*
103 kg/m3
2-10 kg/m-sec
The curves in Figure 9 represent the corresponding theory as predicted by
Eq. (20) and computer results. Note that both theory and data predict 0.99+
collection efficiency regardless of electrostatics for particle radii greater
than approximately 2 ym. Below this value, however, there exists significant
variability between different field values. This size range corresponds to
those particles which tend to lodge in the lower respiratgy tract so their
collection is of greater importance (Miller, et al (1979) ). Data and theory
show that modest electric field strengths can greatly improve bed efficiency
in this region.
The theory presented here generally under-predicts the data gathered
at MIT for fine particles less than 2 Jim. However, the curve shapes and
relative magnitudes compare favorably. The discrepancy is attributed to two
causes:
1) The model does not include collection due to the image force between
a charged fine particle and neutral bed particle. This separate
electrostatic effect would increase predicted efficiencies.
2) Collection due to diffusion, or Brownian motion, is not included
in the model. This mechanism generally becomes significant below
particle radii of about 0.2 Mm.
It is believed that the incorporation of the above mechanisms into the model
will better predict EFB performance.
particle
as a param-
in practical
REFERENCES
1. D.F.S. Natusch and J.R. Wallace, "Urban Aerosol Toxicity: The Influence
of Particle Size," Science, 186, 695 (197*0.
2. K. Zahedi and J.R. Melcher, "Electrofluidized Beds in the Filtration
of a Submicron Aerosol," JAPCA, 26, 3^5 (1976).
3. K. Zahedi and J.R. Melcher, "Collection of Submicron Particles in Bub-
bling Electrofluidized Beds," I&EC Fund., 16, 2H8 (1977).
359
-------
4. K. Zahedi, PhD Thesis, Massachusetts Institute of Technology, Cambridge,
MA (1976).
5. F.J.W. Whipple and J.A. Chalmers, "On Wilson's Theory of the Collection
of Charge by Falling Raindrops," Quart. J. Roy. Met. Soc., 70, 103 (1974).
6. H.F. Kraemer and H.F. Johnstone, "Collection of Aerosol Particles in
Presence of Electrostatic Fields," I&EC EDE, 47, 2426 (1955).
7. H.F. George and G.W. Poehlein, "Capture of Aerosol Particles by Spherical
Collectors: Electrostatic, Inertial, Interception and Viscous Effects,"
Environ. Sci. and Tech., 8, 49 (1974).
8. K.A. Nielsen and J.C. Hill, "Capture of Particles on Spheres by Inertial
and Electrical Forces," I&EC Fund., 15, 157 (1976).
9. P.W. Dietz, "Electrostatic Filtration of Inertialess Particles by Gran-
ular Beds," in preparation, 1979.
10. J. Happel, "Viscous Flow in Multiparticle Systems: Slow Motion of Fluids
Relative to Beds of Spherical Particles," Am. Inst. Chem. Eng. J.,
4,197 (1958).
11. S. Kuwabara, "The Forces Experienced by Randomly Distributed Parallel
Circular Cylinders or Spheres in a Viscous Flow at Small Reynolds Num-
bers," J. Phys. Soc. Japan, 14, 527 (1959).
12. G. Tardos, C. Gutfinger and N. Abuaf, "Deposition of Dust Particles
in a Fluidized Bed Filter," Israel J. Tech., 12, 184 (1974).
13. G. Tardos, N. Abuaf and C. Gutfinger, "Dust Deposition in Granular Bed
Filters: Theories and Experiments," JAPCA, 28, 355 (1978).
14. C. Gutfinger, G. Tardos, and N. Abuaf, "Analytical and Experimental
Studies on Granular Bed Filtrations," Proceedings of the Symposium on
the Transfer and Utilization of Particulate Control Technology, 3, 243
(1979).
15. J.C. Alexander, PhD Thesis, Massachusetts Institute of Technology,
Cambridge, MA (1978).
16. H. Lamb, Hydrodynamics, Articles 93-95, 6th Edition, Dover Publications,
New York (1945).
17. J.R. Melcher and K.S. Sachar, "Charged Droplet Scrubbing of Submicron
Particulate," Environmental Protection Technology Series EPA-650/2-
74-075 (1974).
18. F.J. Miller, et al, "Size Considerations for Establishing a Standard
for Inhalable Particles," JAPCA, 29, 610 (1979).
360
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NOMENCLATURE
constants used in Eqs. (3) and (4)
external electric field
local electric field
radial and tangential components of local electric field
charging electric field
arbitrary electrical force
dimensionless electrical parameter
bed particle dielectric constant
granular bed length
bed particle Reynolds number
fine particle Reynolds number
Stokes number
superficial gas velocity
local gas velocity
radial and tangential components of local gas velocity
fine particle velocity
radial and tangential components of fine particle velocity
bed particle radius
electrical mobility
unit cell radius
inlet fine particle concentration
fine particle charge
polar coordinates
fine particle radius
time
critical trajectory height
normalized parameters
bed void fraction
fine particle collection rate
bed particle permittivity
single-sphere collection efficiency
total bed collection efficiency
gas viscosity
361
-------
fine particle mass density
electric potential function
gas stream function
362
-------
AEROSOL FILTRATION BY A CONCURRENT MOVING
GRANULAR BED: DESIGN AND PERFORMANCE
By:
Thomas W. Kalinowski and David Leith
Department of Environmental Health Sciences
Harvard School of Public Health
Boston, Massachusetts 02115
ABSTRACT
Aerosol filtration by a concurrent moving granular bed filter has been
investigated. The effect of an intergranular dust deposit on mass penetra-
tion and pressure drop is described. A 2^ factorial experiment was conducted
to identify significant control variables, including: superficial velocity
(100 and 250 m/s), bed depth (130 and 230 mm), granule size (2.1 and 2.7 mm),
and the intergranular dust deposit expressed as percent by weight of collected
dust in the bed (1% and 5%). The test aerosol was resuspended utility boiler
fly ash. All tests were conducted at ambient temperatures and pressure. A
static, intergranular dust deposit produced high filtration efficiency (>99%)
at the low superficial velocity; granular motion produced reentrainment.
363
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AEROSOL FILTRATION BY A CONCURRENT MOVING
GRANULAR BED: DESIGN AND PERFORMANCE
INTRODUCTION
In a concurrent moving granular bed filter (CMGBF), aerosol passes down-
ward through a descending bed of granules while clean granules are added con-
tinuously at the bed top. To the aerosol, the slowly moving bed functions as
a graded media filter with coarse collectors (clean granules) encountered
first followed by an increasingly dense deposit of fine collectors (previous-
ly captured dust). By adjusting the granule flow rate and the superficial
gas velocity, the intergranular dust accumulation can be controlled. A
unique feature of the CMGBF is the controlled formation of an intergranular
dust deposit to enhance filtration. Particle collection in the CMGBF more
closely resembles cake filtration than filtration by a bed of clean granules,
in that particle collection depends more upon the nature of the intergranular
deposit than on the clean granules.
The CMGBF offers advantages over alternative granular bed configurations.
In continuously moving cross-flow filters, formation of a uniform dust de-
posit is not possible so that efficiencies greater than 98% are found only
with thick beds (390 mm) or small granules (0.8 mm); this leads to pressure
drops in excess of 2.5 kPa (10" H2O) according to Wade et al. (1978).^ In
intermittently moving beds, cake formation at the bed surface provides high
efficiency (> 99%) but periodic cleaning by reverse air or high pressure
pulsed air is required. McCain (1978)2 found increases in outlet concentra-
tion associated with periodic backflushing in an intermittently cleaned
granular bed filter. Even if such periodic peaks in outlet concentration do
not degrade average performance, they may be unacceptable for applications
involving downstream gas turbines. The CMGBF offers the advantages of high
efficiency provided by the intergranular particle deposit and continuous oper-
ation without bypassing for cleaning.
This study was undertaken to determine the fundamental mechanisms by
which an intergranular dust deposit forms in a CMGBF and the manner in which
this deposit influences bed efficiency and pressure drop. This paper describes
the CMGBF experimental apparatus and reports the results of a screening experi-
ment to investigate the effects of superficial velocity, (V), bed depth, (B),
granule diameter, (D), and the amount of intergranular dust deposit (K) ex-
pressed as percent by weight of collected dust in the bed. The first three
factors are important in clean-bed filtration theory; the last reflects the
condition of the dust-laden bed.
EXPERIMENTAL APPARATUS
A schematic diagram of the CMGBF experimental apparatus is shown in
Figure 1. The test aerosol is electrostatically precipitated coal fly ash
364
-------
from a utility boiler Ihe a&h was sieved through a screen with 348 ym openings
to remove large particles, then stored at 150°C. The fly ash had a mass
median aerodynamic diameter of 6.2 ym and geometric standard deviation of 2.4.
The fly ash was resuspended by a National Bureau of Standards dust feeder
(Dill 1938).^ The aerosol passed through a charge neutralizer using a 50
mCi Sr source designed according to the criteria of Cooper and Reist (1973).^
The air flow rate through the CMGBF was always greater than that coming
from the aerosol generator, with the difference provided by room air. Flow
was measured by a calibrated Stairmand disk at the inlet, which also mixed
the aerosol to provide a uniform inlet concentration (Stairmand 1941).-'
Air flow was controlled by a globe valve in the outlet pipe and a bypass
valve as shown on Figure I. Downstream were a surge tank and type AF24 Roots
blower. Relative humidity was determined by dry bulb and wet bulb tempera-
tures downstream of the bed.
Upstream and downstream isokinetic sampling were identical and consisted
of fixed 4.7 mm diameter probes center-line mounted in the inlet and outlet
pipes. Each probe was connected to a pair of 47-mm filter holders. One fil-
ter was used at the start of sampling for purging the probe and setting flow.
As shown in Figure 1, a solenoid valve diverted flow from the purge filter to
the sample filter during sampling. Mass samples were collected on MSA 1106 BH
glass filters.
The CMGBF consisted of three gas-tight sections: a granule storage bin,
the bed column, and a granule control mechanism. Granules flowed by gravity
from a 39 L storage bin mounted on top of the column. The bin was flanged to
an aluminum distribution plate from which four funnel spouts, each 19 mm in
diameter and 460 mm long, extend downward into the column. The storage bin-
funnel assembly could move up or down on external support rods.
The moving bed was retained in a 203 mm diameter column made from acrylic
plastic. The inlet pipe entered the column wall above the bed surface. Inlet
screen plugging has been reported for other granular filter designs (Wade et_
al. (1978)1 and Hoke and Gregory (1977)^), but was not found in the CMGBF
which had no inlet screens as aerosol passed directly into the top of the bed.
The bed outlet consisted of a nearly continuous circumferential port lo-
cated 200 mm from the bottom of the column. This port was sufficient to pro-
vide an average exit velocity equal to the superficial gas velocity through
the bed, thereby preventing gas acceleration at the exits which could contri-
bute to reentrainment of collected dust. The granules were retained by a
stainless steel screen with 0.3 mm openings. The exit port area was contained
within an enclosed annulus from which the outlet pipe exited.
Two rows of pressure taps were located 90° from each other along the col-
umn wall at 20 mm intervals; three more taps were located in the outer ring of
the exit annulus. Using these taps, total pressure drop across the bed and
pressure drop at intermediate bed depths could be obtained.
365
-------
The bed depth, B, was defined as the distance between the end of the
funnel outlets and the center line of the gas exit ports. Bed height was
varied by adjusting the elevation of the storage bin-funnel assembly. Mound-
ing of granules at the funnel outlets caused the top of the bed to be uneven;
this led to slightly lower pressure drops than would have been obtained had
the bed surface been flat. Based on pressure drop, bed depth as defined
above was found equivalent to that of a full right cylindrical bed approxi-
mately 30 mm shorter than the depth defined by the funnel outlets. Pressure
drop data are reported as pressure drop per unit bed depth (AP/B) based on
averages of intermediate pressure drop measurements to remove end effects of
the uneven bed surface and the outlet screen.
Takahashi and Yanai (1973)^ found that a turntable located at the column
bottom provided plug flow in granular beds. Plug flow was highly desirable,
to minimize granule-to-granule motion. As shown in Figure 1, a 298 mm diam-
eter turntable was centered directly below the column and enclosed in a hop-
per with acrylic front and back plates. The 200 mm bed height below the gas
outlet screen was provided to ensure a depth to diameter ratio greater than
one for all bed depths, to enhance plug flow and to prevent the bed from ro-
tating as a solid body on the turntable. The turntable was calibrated for
mass flow rates between 60 and 1000 grams per minute against motor RPM. These
calibrations were used as guides, and actual granule flowrates were obtained
for each run by weighing collected granules. The dirty granules were col-
lected in collection bottles below the hopper.
Two sizes of high density (pg = 3.25 g/cc) alumina granules were ob-
tained from Electrorefractories and Abrasives Division of Ferro Corporation
in East Liverpool, Ohio. Ferro Corporation currently markets these granules
as "Arlcite" in 8-12 mesh and 6-8 mesh sizes which were found to correspond
to a mean diameter of 2.1 mm for the 8-12 mesh size and 2.7 mm for the 6-8
mesh size.
To simulate the loose packing porosity in a moving granular bed, the Hap-
pel (1949) slow re-inversion method was used. The porosity calculated using
this procedure was found to be 0.41 for both granule sizes. Happel (1949)®
found this method to be within one percent of the flowing bulk density of
moving catalyst beds. For uniform, smooth spheres in the absence of wall ef-
fects, Eastwood et ail. (1969)^ found the random loose porosity to be 0.43 ±
0.02. Generally, porosity was found to be less for non-uniform granules,
as was found for the Arlcite granules in these tests.
The pressure drop characteristics of the unused alumina granules are
shown in Figure 2, a plot of &P/B against superficial velocity, V. Pressure
drop was linear with bed depth, and each experimental data point in Figure 2
is the slope of the least squares linear regression of the AP versus B data
over a range of 50 to 300 mm depth. The Ergun (1949)10 pressure drop equa-
tion fitted these data well if bed porosities of 39% and 38% were used for
the 2.1 mm and 2.7 mm granules, respectively. These porosities may be within
the range of experimental error for the 41% porosity measurements made using
the Happel slow Inversion technique.
366
-------
After use, the collected dirty granules were screened and reused. The
cleaned granules were not entirely free of fly ash as indicated by an ob-
served average increase of 0.2 Pa/mm in the pressure drop per unit depth
over that for new granules; however, complete separation of collected dust
from the granules may not be necessary or even important for effective opera-
tion of the CMGBF. Lee (1975)^ found higher collection efficiency for
cleaned sand compared to new sand in a panel bed filter where cake formation
was also desired. Recycle of dusty granules may allow the intergranular de-
posit to form faster than would occur with completely clean granules.
INITIAL PERFORMANCE TESTS
A screening experiment identified significant control variables which
affect particulate mass removal efficiency and pressure drop in the CMGBF.
A randomized 2^ factorial experiment was performed with the variables listed
in Table 1. The experiment consisted of a total of 16 runs conducted at room
temperature and atmospheric pressure.
The low superficial velocity of 100 m/s was limited by the NBS dust
feeder air flowrate. The high velocity of 250 mm/s was selected to be suffi-
ciently below the 350 mm/s velocity reported by Goren (1978)^2 to cause bounce
and/or reentrainment of 3.91 um potassium bipthalate particles in a static bed
of 2 mm alumina granules. Granule size was limited by the two available sizes
of Arlcite granules. The bed depths were selected to be in a range that would
produce moderate efficiency (< 97%) according to clean-bed theory for these
test conditions. Furthermore,the time to reach steady state as measured by
bed volume replacement became excessive for depths greater than about 230 mm.
The fourth factor, K, was selected as a measure of the intergranular dust
deposit and defined as:
W
K% - ^ (100) (1)
g
where W
-------
portion of the bed between the bed surface and bottom of the gas exit ports
was replaced. The time for bed volume replacement was calculated from the
dimensions of the column, the bed depth, and granular mass flowraire. At the
end of two bed replacements, steady state operation was expected and was
found in some runs as indicated by leveling off of penetration and pressure
drop. Although it is not certain that a true steady state was achieved in
all cases, two bed replacements appeared adequate for the purposes of these
screening experiments.
Mass penetration and pressure drop measurements were made throughout
each run, at quarter bed volume increments. Granular mass flowrate, W ,
was determined three times per run by weighing collected granules. Relative
humidity was determined from dry bulb and wet bulb temperatures early in each
run before the wet bulb wick became dirty. At the end of each run special
penetration measurements weremade as described later. Upstream and downstream
samples were also taken at two bed volumes for particle size analysis by
scanning electron microscopy (SEM). Isokinetic samples were collected on 0.2
ym pore diameter Nuclepore filters. Particle size analyses will be reported
in a subsequent paper.
RESULTS
The experimental conditions and outcome variables for each of the 16
runs are listed in Table 2. Analysis of variance was used to identify sig-
nificant control variables and their interactions. The sums of squares for
3-way and 4-way interactions were pooled to estimate the error mean square,
or error variance. Where appropriate, significant control variables were used
in multiple regression equations to summarize results.
Overall mass penetration ranged from 2.6% to 24%. Analysis of variance
identified the intergranular dust deposit, K, superficial velocity, V, and
bed depth, B, to be significant (p < 0.001, 0.001, and 0.01, respectively).
The first order interactions KV and VB were highly significant (P < 0.001).
The KB and KVB interactions were also significant (P < 0,05). The effects
on penetration of K, V and B and their interactions were positive, i.e.,
penetration increased as each of these factors increased. The effect of gran-
ule diameter, D, was also positive, i.e., penetration increased as D increased,
but was not significant at the 5 percent level. The presence of the inter-
actions, particularly KVB, indicate a multiplicative, rather than a simple
additive effect.
2
Figure 3 shows the least squares regression fit (r * 81%) of the mass
penetration data at two bed volumes to the Interaction term KVB:
Pt(%) = 2.42 + 0.69 x 10~4 KVB (% - mm2/s). (2)
Slight improvement in the correlation coefficient is possible by adding the
separate significant terms in a multiple regression equation; however, consider-
ing the size of the experiment and the large r2 provided by the single inter-
action term, the simplicity of equation 2 appears appropriate. It should be
noted that the powers of K, V, and B in equation 2 may be other than one, but
368
-------
a two level factorial experiment cannot reveal them. Relative humidity
ranged from 25 to 48 percent and was not statistically significant.
The pressure drop per unit bed depth (AP/B) ranged from 1.7 to 18.7
Pa/mm. For AP/B at two bed volumes, the analysis of variance revealed K, V
and D to be highly significant (P < 0.001). Only the KV and KD interaction
terms were significant at P < 0.05. The K, V, and KV terms had positive ef-
fects, increasing pressure drop. The effect of D was negative, as was the
KD effect. Bed height, B, and its interactions showed no significant effect,
as would be expected for the normalized AP/B variable.
2
Figure 4 shows the correlation (R = 96%) between observed pressure drop
data and values predicted by the simple additive model for the effects of K,
V, and D for the dust laden CMGBF:
AP/B(Pa/mm) = 8.44 + 3.05(K%) + 0.0099(V mm/s) - 4.39(D mm). (3)
Only a slight improvement in the multiple correlation coefficient is possible
by inclusion of the two significant interaction terms.
The outlet screen, which was designed to retain granules much finer than
those used in this experiment, was subject to plugging in some runs as indi-
cated by pressure drop rise across the outlet screen. Equation 3, however, is
based on intermediate bed depths and as such does not include pressure drop
across the outlet screen.
DISCUSSION
The numerical value of the coefficients in regression equations 2 and 3
are of less importance than the trends indicated by signs on the coefficients
and by the terms found to be significant. This section compares these re-
sults with conventional static, clean granular bed theory and identifies some
of the unique aspects of the CMGBF.
The penetration relationship, equation 2, is surprising because it op-
poses what would be predicted by clean-bed theory and because it is opposite
what was expected from knowledge of cake filtration, Conventional clean-bed
theory predicts decreasing penetration with increasing bed depth, B, and with
decreasing granule diameter, D. Present data show an opposite effect for bed
depth and no significant effect of granule diameter. Although the difference
between the two granule sizes used was not great, it appears that granule size
is of less importance in the CMGBF than might be predicted by conventional
theory.
For the granule sizes the fly ash used in these experiments, inertial
impaction should be the dominant collection mechanism (in the absence of sub-
stantial electrostatic effects). Thus, penetration should decrease as gas
velocity increases. It was expected that particle removal would be greatly
enhanced beyond that expected from impaction on the granules alone, due to
presence of the intergranular dust deposit. Using a 50 mm diameter bed of
Na2C02 pellets at a superficial velocity of 100 tnm/s, Rudnick et al. (1976)^
observed fly ash penetrations less than 0.02 weight percent in a CMGBF of
369
-------
somewhat different design.
The time trend of outlet concentration, shown in Figure 5 for a typical
run, is useful in understanding the apparently anomalous penetration data.
Figure 5 shows outlet concentration as a function of the number of bed volume
replacements. Outlet concentration decreased at first, as expected due to
intergranular deposit formation. However, at approximately three-quarters
bed volume, outlet concentration rapidly increased and then leveled off for
the remainder of the run. Three-quarters bed volume corresponded to the
point at which the loaded bed reached the upper edge of the outlet screen.
This breakthrough suggests reentrainment of previously collected dust at the
outlet screens. Wade jet al. (1978)1 found reentrainment to occur in a contin-
uously moving cross-flow granular bed filter. Accordingly a series of end-
of-run penetration measurements were conducted to identify the significance
of possible penetration processes including:
(1) straight through
(2) air scouring
(3) scouring by granular motion.
After the normal penetration measurements were completed, bed motion was
stopped and penetration was immediately measured again. This dust-on/bed-
stopped situation measured the penetration straight through the loaded, static
bed. Next the dust feeder was stopped and the downstream concentration was de-
termined with the bed still stopped, to measure the effect of air scouring at
the superficial velocity of the run. Finally, the bed motion was restarted
but with no dust fed. This dust-off/bed-moving situation allowed sampling of
the collected dust reentrained from the loaded bed as the result of granular
motion.
The results of these experiments are presented in Table 3. Expressed as
a percentage of the outlet mass concentration, the straight through case (1)
for the static, loaded bed averaged 17%. Air scouring, case (2), at the su-
perficial velocity averaged 2.7% but was essentially zero for the low velocity
runs. Case (3), reentrainment due to granular motion, was found to represent
an average of 72% of the outlet concentration. Recognizing that air scouring
case (2) is probably a component of both case (1) and case (3), these averages
sum to roughly 90%, representing a closure error of 10 percent. Because of
the disruptive effect of stopping and restarting the bed to make the case (3)
measurement, errors in the summation are understandable. The significant con-
clusion here is that the fraction of outlet concentration attributable to gran-
ular motion was always greater than that due to straight through penetration
and generally was about four times greater.
Visually, the downstream samples for case (1) and case (3) were quite
different. Penetration of the static, loaded bed generally produced a uni-
form discoloration of the downstream sample filter. However, the downstream
samples from case (3) situations often showed distinct, large particles or
agglomerates scattered across the filter surface. These observations will be
verified by analysis of SEM samples which were also taken for each run; these
analyses are currently underway and results will be reported later.
370
-------
Based on the data in Table 3, penetration through the static, loaded
bed averaged less than 0.5% by weight at 100 mm/s and 3.0% at 250 mm/s, ig-
noring the levels of the other factors. Increased penetration at the higher
velocity may be largely due to air scouring. Efficiencies in excess of 99.5%
at the lower velocity for the relatively fine test aerosol indicate the po-
tential for high efficiency if reentrainment due to granular motion is reduced
or eliminated.
Reentrainment may be reduced by lowering the granular velocity. If gran-
ular motion at the gas outlet screens is the major cause of reentrainment in
the present design, an alternative outlet configuration in which the gas exits
through a layer of stationary or more slowly moving granules may improve per-
formance. However, the increased penetration with increased bed depth indi-
cates that reentrainment at the outlet screens is not the only adverse effect
of granular motion. The presence of the bed depth, B, in equation 2 suggests
that reentrainment increases even within the bed as the amount of dust avail-
able (KB interaction) for reentrainment increases. A cascading effect can be
envisioned by which collected dust agglomerates break off and are redeposited
or are reentrained throughout the bed.
Even without major modification, it may be possible to operate the CMGBF
as an agglomerating filter which.can capture greater than 95 percent of the en-
tering fly ash mass. In addition, a significant portion of the reentrained
agglomerates might be collected in a cyclone downstream.
Finally, it is recognized that the alumina granules used in this experi-
ment offer advantages with regard to mechanical and thermal properties but
that their relatively large size and smooth surface may preclude adequate
adhesion by the fly ash particles and agglomerates. ^£00^ pellets used by
Rudnick et al. (1976)13 were softer and subject to humidity problems but may
have superior fly ash adhesion properties so that reentrainment was reduced.
Equation 3 for pressure drop indicates expected trends. Pressure drop
across a granular bed is known to increase as superficial velocity increases
and as granule diameter decreases, as predicted by the Ergun equation for
clean beds. Similar trends are confirmed in this experiment for loaded granu-
lar beds. In addition, equation 3 shows that the amount of intergranular dust
deposit, K% by weight, has a significant positive effect on pressure drop.
In this experiment for a given combination of superficial velocity and
granule size, increasing K from 1% to 5% generally produced five-fold increase
in AP/B. However, based on experiments at two levels of K, it cannot be con-
cluded that the effect of K is linear. Additional experiments over a range of
values for the intergranular dust deposit will be required to determine this
relationship. Because pressure drop is very sensitive to bed porosity, pressure
drop should rise as increases in the intergranular dust deposit reduce bed
porosity.
Based on the pressure drop measurements made at intermediate bed depths,
a uniform intergranular dust deposit is established within a few centimeters of
the bed surface. It appears advantageous to operate the CMGBF with a shallow
bed to reduce overall pressure drop without significantly affecting efficiency.
371
-------
SUMMARY
An experimental apparatus has been constructed to investigate aerosol
filtration by a concurrent moving granular bed. Screening experiments have
been performed to identify significant control variables. Initial results
show that conventional collection theory for static, clean granular beds is
inadequate to describe the performance of loaded, concurrently moving beds.
This is indicated by trends of increasing mass penetration with increases in
superficial velocity and bed depth, opposite to what is predicted by clean-
bed theory. Granule diameter was found to have little effect on penetration
in a bed with an established intergranular dust deposit.
Formation of an intergranular dust deposit results in performance char-
acteristics that are significantly different from those of comparable clean
beds. It appears that the intergranular dust deposit functions as a highly
efficient filter for collection of dust in the inlet stream. The collection
efficiency of the static, loaded bed averaged 99.5% by weight at 100 mm/s su-
perficial velocity and 97% by weight at 250 mm/s superficial velocity. How-
ever, agglomeration and reentrainment of collected dust were significant when
the bed was moving. Over the range of conditions in this experiment, about
three quarters of the penetrating dust resulted from reentrainment of col-
lected dust caused by granular motion.
Experiments are under way to investigate these and other effects further.
Fractional efficiency data will also be analyzed to determine fine particle
collection efficiency, extent of particle agglomeration within the CMGBF, and
nature of reentrained dust.
ACKNOWLEDGEMENT
This work was supported by the National Science Foundation, grant number
ENG 77-26975.
372
-------
REFERENCES
1. Wade, G.L., et al. Granular Bed Filter Development Program: Final Re-
port. Department of Energy Report FE-2579-29, April 1978.
2. McCain, J.D. Evaluations of Novel Particulate Control Devices. Environ-
mental Protection Agency Report EPA-600/7-78-093, June 1978. p. 33-62.
3. Dill, R.S. A Test Method for Air Filters. Trans. Amer. Society of Heat-
ing and Ventilating Engin. 44:379-386, 1938.
4. Cooper, D.W. and P.C. Reist. Neutralizing Charged Aerosols with Radio-
active Sources. J. Colloid and Interface Science. 45:18-26, October 1973.
5. Stairmand, C.J. Sampling Gas-Borne Particles. Engineering. 41:141-143,
181-183, August 22, 1941.
6. Hoke, R.C. and M.W. Gregory. Evaluation of a Granular Bed Filter for Par-
ticulate Control in Fluidized Bed Combustion. In: Proceedings of the EPA/
DOE Symposium on High Temperature High Pressure Particulate Control, Wash-
ington, D.C., September 1977. Environmental Protection Agency Report EPA-
600/9-78-004. p. 112-131.
7. Takahashi, H. and H. Yanai. Flow Profile and Void Fraction of Granular
Solids in a Moving Bed. Powder Technology. 7:205-214, 1973.
8. Happel, J. Pressure Drop Due to Vapor Flow Through Moving Beds. Indus-
trial and Engineering Chemistry. 41:1161-1174, June 1949.
9. Eastwood, J., et aJL. Random Loose Porosity of Packed Beds. British Chem-
ical Engineering. 14:1542-1545, November 1969.
10. Ergun, S. Fluid Flow Through Packed Columns. Chemical Engineering Pro-
gress. 48:89-94, February 1952.
11. Lee, K. Filtration of Redispersed Power-Station Fly Ash by a Panel Bed
Filter with Puffback. Ph.D. Thesis. The City University of New York,
April 1975.
12. Goren, S.L. Aerosol Filtration by Granular Beds. In: Proceedings of the
First EPA Symposium on the Transfer and Utilization of Particulate Con-
trol Technology, Denver, May 1978. Environmental Protection Agency Re-
port EPA-600/7-79-044c. p. 459-469.
13. Rudnick, S.N. Particulate Collection in a Low Level Radioactive Waste
Incinerator. In: Proceedings of the 14th ERDA Air Cleaning Conference,
Sun Valley, Idaho, August 1976. ERDA Report C0NF-760822. p. 80-101.
373
-------
Table 1 FACTORS IN 24 FACTORIAL EXPERIMENT
Control Variables
K - Intergranular dust deposit
V - Superficial velocity
B - Bed depth
D - Mean granule diameter
Level
Low
Hifih
1
5% by wt.
100
250 mm/s
130
230 mm
2.1
2.7 mm
374
-------
No
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
*
Table 2 RESULTS OF 2 FACTORIAL EXPERIMENT
Conditions
K V B D
+ + + +
- - - +
- + + -
+ — + +
+ + + -
+ + - -
- + - +
- - + 4-
+ - - +
+ - + -
- + + +
- +
+ +
+ -
Penetration
@ 2 Bed Volumes
(% by weight)
4.63
24.0
4.67
7.28
4.00
22.0
12.3
6.94
2.60
3.06
5.03
5.50
4.73
12.1
15.4
6.79
AP/B
2 Bed Volumes
(Pa/mm)
2.5
13.3
2.2
3.6
10.0
18.7
17.4
2.8
1.8
1.7
11.2
3.7
15.4
2.7
13.7
16.2
shown in Table 1; high level as + and low level as -.
375
-------
Table 3 END-OF-RUN PENETRATION TESTS
o
Penetration
Run
Outlet
Concentration(g/m )
% of C2BV
*
Case (1)
No.
C2BV*
C(l)
C(2)
C(3)
(1)
(2)
(3)
(% by weight)
1
0.204
0.027
0.000
0.072
13.2
0.0
35.3
0.51
2
1.31
0.600
0.072
1.64
45.8
5.5
125.
7.66
3
0.166
0.030
0.004
0. 121
18.1
2.4
72.9
0.86
4
0.133
0.002
0.000
0. 153
1.5
0.0
115.
0.12
5
1.02
0.053
0.026
0.735
5.2
2.5
72.1
0.20
6
1.87
0.292
0.034
1.10
15.6
1.8
58.8
3.20
7
1.05
0.170
0.049
0.646
16.2
4.7
61.5
2.65
8
0.150
0.040
0.012
0.107
26.7
8.0
71.3
1.59
9
0.085
0.015
0.000
0.072
17.6
0.0
84.7
0.41
10
0.134
0.008
0.000
0.115
6.0
0.0
85.8
0.17
11
1.24
0.072
0.004
0.738
5.8
0.3
59.5
0.41
12
0.101
0.030
0.002
0.082
29.7
2.0
81.2
1.66
13
1.27
0.080
0.004
0.250
6.3
0.3
19.7
0.27
14
0.268
0.030
0.008
0.167
11.2
3.0
62.3
1.76
15
1.38
0.451
0.153
1.63
32.7
11.1
118.
5.29
16
1.43
0.251
0.015
0.481
17.6
1.0
33.6
0.69
Average
16.8%
2.7%
72.3%
SD
11.9
3.2
29.8
ic
Outlet concentration under normal operating conditions at the end of two
bed volume replacements.
(1) Straight through — dust-oi^bed-stopped
(2) Air scouring — dust-ofij/bed-stopped
(3) Scouring by granular motion — dust-off/bed-moving
676
-------
20.3 cm £?MM£r£fi £££>
\
COMPP£SS££>
/}/£
GMAX/L£ STO/MG£ 0/A/
PLOW COA/TPOL
KALf£S
A DtSST
k—' rem
£.5 cm £>//}.
/.J cm &/A.
0Br mr
CAL/0Mr£D/*
ST/1/PM4//D
£>/SK
&cm*A M
47-mm
l£/ir£/?
HOLD£f?
Sl/PG£
urn
S0L£A/0/D MLl/£
Is
tHxi—Q
poors
&LOW£P
S4MPL£ Pi/MP
GP4A/VL£S
COLL£Cr/OA/
sorri£
/4CL/ISM
G/4UG£
^por/iM£r£P
Figure 1 Concurrent moving granular bed filter
experimental apparatus .
-------
£.5
£0
£
E
CD
Q_
CQ
S
D_
<3
a
LU
CD
Z
<
LU
—I
O
/s
/0
O.S
1 '
1
I T
1
i
—-
ERGUN EQ.
D = 2.1
mm e = 0.39
/°
ERGUN EQ.
r-»
CM
II
cs
mm e = 0.38
o
EXPT. D =
2.1 mm
A
/
/
A
EXPT. D =
2.7 mm
—
y /
/ /
/ /
/
P
/
m
-
' /
/
/
—
1. L_
1
« •
1 ,
. 1
/00
200
J0O
V (mm/s)
Figure 2 Clean bed pressure drop versus
superficial velocity-
378
-------
&
20
/S
• 1
¦ i
' 1
III
1 O
g/ "
-
° y/
r R2 = 81%
o
/D
-
—
—
X
o
1
-
¦vX°
o
o
—
~
i
-
.i 1
1 1
• 1
1 1 1
/2 /*
KVB (%-mm^/s)
REGRESSION EQUATION: Pl% = 2.42 + 0.69xl0-ll(KVB %-mm2/s)
X 10
Figure 3 Penetration at two bed volumes
regression equation.
379
-------
i
LU
LU
CO
CO
o
OO
/e
AP/B (Pa/mm)
- PREDICTED BY REGRESSION -
AP/b (Pa/mm) = B.kk + 3-05(K %) + 0.0099(V mm/s) - *».39(D mm)
Figure A Loaded bed pressure drop
regression equation.
380
-------
CONDITIONS:
K = 1.0%
V = 250 mm/s
B = 230 mm
D = 2.1 mm
AVG. INLET CONC.
= 1 -75 g/nr_
±
JL
J-
±
J L
AO
AO
NUMBER OF BED VOLUMES
Figures Outlet concentration versus number
of bed volumes - run no. h.
381
-------
DEEP BED PARTICULATE FILTRATION
USING THE
PURITREAT(TM) PROCESS
By :
L. C. Hardison
Air Resources, Inc.
600 N. First Bank Drive
Palatine, IL 60067
for:
Presentation at the
Second Symposium on the
Transfer and Utilization of
Particulate Control Technology
University of Denver
July 23, 1979
382
-------
DEEP BED PARTICULATE FILTRATION
USING THE-
PURITREAT(TM) PROCESS
INTRODUCTION
Particulate control of tarry, semi-solid materials
generated by industrial processes has presented a special
problem because of the failure of conventional high efficiency
equipment such as electrostalic precipitators and fabric col-
lectors. Deep bed filtration with disposable media works well
but replaces the air pollution emission problem with a solid
waste disposal problem.
This paper describes a unique granular moving bed filter
in which raw material for the process serves as the medium,
and is "disposed" by introduction into the process. This
approach was developed by Union Carbide Corporation and is
offered for industrial use by Air Resources, Inc. for a variety
of processes such as carbon calcination and ring furnaces, in
which it is an energy efficient alternative to thermal inciner-
ation.
383
-------
PROBLEM DESCRIPTION
The collection of particulate emissions from industrial
processes is practiced to prevent air pollution and to recover
raw materials or products which would otherwise be discharged
into the atmosphere. A wide variety of techniques for emis-
sion control allows selection of equipment to fit the process
and the properties of the particulate matter, and for many
applications, the system designer has a choice of several al-
ternatives. For example, utility power stations utilize wet
scrubbers, electrostatic precipitators, and recently, fabric
collectors for fly ash removal.
This is not generally true for processes which emit tarry
organic fumes. The aerosol fumes produced by curing of carbon
electrodes, which are composed of petroleum coke particles
bound together with heavy petroleum tar or pitch, are diffi-
cult to collect because they are very small in particle size,
and, more importantly, because they are sticky. Thus, the
collection techniques most applicable to small particle size
solid particulate collection are not workable because the col-
lected material cannot be removed mechanically. This elimi-
nates such conventional techniques as
(1) Mechanical (cyclonic) collectors
(2) Electrostatic precipitators
(3) Fabric filters
Wet collectors are a possibility but the small particle
size limits the effectiveness and the wet tarry slurry tends
to foul the internals and prevent good mechanical operation.
The remaining devices which can be used for mechanical
collection of the fume materials are
(1) Disposable filters, such as the High Energy
Air Filter (HEAF) developed by Johns Manville
Corp .
(2) Disposable sand bed filters and, because the
fumes are combustible organic materials:
(3) Thermal incineration.
Of these three, only thermal incineration has found any wide
spread application for the abatement of pitch fumes, and there
have also been some applications of wet electrostatic precipi-
tation equipment.
384
-------
In the production of carbon electrodes for use in
metallurgical furnaces", prebake anodes for primary aluminum
smelting, and other processes for making graphite products,
petroleum coke and pitch are used as raw materials. The coke
is ground to a relatively small particle size, generally on
the order of 1/4" maximum, and blended with heated pitch to
produce a semi-solid material. This paste is then formed into
the desired shape in extrusion presses or molds. The formed
product is heated to a temperature in excess of 1000°F in the
presence of air to oxidize the pitch binder and solidify the
mass. Subsequent high temperature treatment of the electrodes
may be carried out at temperatures as high as 5500°F to bring
about crystallization or graphitization of the carbon.
At each step in the process described, organic fumes are
generated by vaporization of a fraction of the pitch binder.
In particular, the initial heat treatment of the electrodes
at temperatures over about 700°F gives rise to substantial
emissions of tarry pitch fumes.
A prebake oven for curing the pitch binder in preformed
electrodes might typically discharge 10,000 SCFM to 30,000
SCFM of flue gas, which is largely air containing a small
amount of organic aerosol - in the range of a few pounds per
hour - which generates a dense visible plume when discharged
into the air. The fume material is typically a small particle
size aerosol, with a maximum particle size of a few micrometers
in diameter. The particles may vary in composition but are
most frequently heavy, complex organics of high boiling point,
which may contain such objectionable compounds as polycyclic
organics (Benzo-a-Pyrene is a member of this group) and partial
oxygenates formed by oxidative decomposition of the pitch.
Usually the particles form sticky or tarry semi-solid
coatings on surfaces upon which the contaminated gas streams
imp inge.
385
-------
SYSTEM DESCRIPTION
The Puritreat(TM) method for capturing these aerosol
fumes at minimum cost involves the use of a bed of coke par-
ticles which forms, in effect, a deep bed particle filter.
The fume laden air stream is passed downward through a layer
of particles several inches deep, at velocities up to 1000
ft/min, which results in a pressure drop on the order of 30"W.C.
This results in collection of the aerosol particles at
high efficiency by impingement of the droplets at high velo-
city on the larger particles making up the bed. This type of
collection is well known, and "sand filters" or other granular
beds have been subjects of a great deal of experimentation if
not commercial application, for dry solids collection.
The novel approach utilized by UCCPD involved the use of
a part of the granular carbon raw material for the electrode
manufacturing process as the collection medium. Figure 1
illustrates the basic mechanism for collection and "disposal"
of the fouled carbon by feeding it into the electrode manufac-
turing process.
Figure 2 indicates the further improvement of the system
by use of an edge seal device which permits automatically ad-
vancing bed which is indexed forward as the pressure drop
increases, indicating a significant reduction in the void
volume available for retention of collected tars.
Figure 4 is a photograph of the feed hopper on a 30,000
cfm commercial unit operated on an electrode curing operation.
Figure 5 is a photograph taken from the opposite side of
the unit. The coke hopper receives coke particles through a
six-inch diameter screw conveyor. The preferred particle size
range is obtained by screening the raw granular coke to between
three and six mesh or about 1/4"
Figure 6 shows the blower and motor. Spent coke is
raised to a spent coke hopper by a screw conveyor on the far
side.
Figure 7 shows the coke bed and the typical yellow color
of the collected pitch fumes.
386
-------
LOADING HOPPER
INDEXING
MOTOR a
DRIVE
SUPPORT SCREEN
RISE IN FAN AMPERES
CAUSES AUTOMATIC INDEX
TO FRESH BED POSITION
COLLECTION BIN
FOR RECYCLE TO
PLANT PROCESS-
FIGURE /
SCHEMATIC DIAGRAM FOR PUR/TREAT (TM) COKE BED FILTER
387
-------
FIGURE
GATE FOR AUTOMATIC INDEXING OF COKE BED
388
-------
Valve
—•- Clean zs/i2.
Ps.t Coke Or Mem
ASed/vm Met Coke
Flat S teel Bar
/ns talleo A t Proper
Aa/gle To Reta/n Coke
Clean
W
00
ID
•Sect/on A-A.
FIGURE 3
ALTERNATIVE BED CONFIGURATION
-------
FIGURE 4
30,000 CFM Commercial PuritreatTM Unit
FIGURE 5
Coke Hopper and Screw Feeder
390
-------
FIGURE 6
Blower and Spent Coke Conveyor
FIGURE 7
Coke Bed
391
-------
ALTERNATIVE SOLUTIONS
Several alternatives are available for handling the
abatement of tarry fumes. These are
a) Combustion
b) Physical removal
Because the tarry materials are almost entirely combus-
tible organic materials, it is possible to remove them from
the gas stream by incineration. Both catalytic and thermal
incineration systems are capable of high removal efficiencies
provided that the tarry materials are handled as vapors rather
than condensed particulate matter.
Capital and operating costs for incineration are high,
however, and are likely to involve operating costs of 25 to
50 times as high as the cost of operation of the Puritreat(TM)
system. Heat recovery is difficult to practice on fume streams
containing tarry materials because of the tendency of these
materials to foul the heat exchange surface, and to cause
fires. The capital and operating costs for both thermal in-
cineration and Purltreat(TM) are compared in Figure 8. Cata-
lytic incineration is capable of operating at about half the
fuel cost, but has considerable uncertainty in many applica-
tions in carbon processing because of the high temperatures
which may be required to assure that the organics are totally
vaporized when they reach the catalyst. For this reason, the
economic comparison with catalytic incineration was omitted.
High Energy filtration (HEAF) utilizing fixed filter ele-
ments composed of fiberglass or fixed granular beds is unwork-
able because the filter elements are rapidly fouled with tar
and cease to function. HEAF filters with disposable, roll-type
elements can be utilized, but have very high initial capital
costs and high operating costs, as indicated in Figure 9.
For this comparison, two alternative processes with
different fume loadings were checked because the fume loading
greatly influences the cost of filter replacement. HEAF
filters have a further disadvantage in that they generate
large volumes of contaminated fiberglass filter elements which
must be disposed of in an environmentally sound manner.
Wet scrubbers are capable of collecting tarry fume mater-
ials of relatively large particle size. The capital cost is
relatively independent of the efficiency, and operating costs
are comparable to the Puritreat(TM) unit. However, the effi-
ciency of removal is low and the separation of the tarry slurry
from the circulating water is extremely difficult.
392
-------
OPERATIONAL DATA
Preliminary tests of the Puritreat(TM) system were
carried out using a pilot unit configuration, as shown in
Figure 10. The pilot unit data were then used to construct
the commercial unit now in operation at the Union Carbide
plant in Columbia, Tenn. The test data discussed in the fol-
lowing paragraphs was taken from a similar unit at Welland, Ontario.
Efficiency tests were carried out for total particulate
collection and also for the removal of the specific polycyclic
organics, benzo (a) pyrene (BaP) and benzo (k) fluoranthrene
(BkF). Concentrations of total gaseous hydrocarbons were also
measured at the bed inlet and outlet using a continuous total
hydrocarbon analyzer.
Particulate loadings were measured by isokinetic sampling
with a glass fiber filter, as shown in Table 1. The particu-
late collection efficiency ranged from 72.7% to 97.9% and
averaged 83% over all of the tests. Operating conditions
during the test period are given in Table 2.
The removal efficiencies of BaP and BkF are summarized
in Table 3 and average overall average loadings during the
test program are given in Table 4. Efficiencies for BaP and
BkF tended to be slightly higher than for total particulate
matter.
It should be emphasized that there are many variables
involved in establishing the efficiency in a given situation.
Many of these, such as particle size and density, are difficult
to characterize except by on-site pilot test work. A pilot
unit is available for on-site tests and should be used for
characterizing existing applications.
393
-------
600
400-
w
*
200
BASIS FOR OPERATING COST ANALYSIS.
NATURAL GAS, 1000 BTU/FT* fo&O/MCF.
ELECTRIC POWER, $ 0.014/KWH.
25/l2 PARTICLES, $40.00/M#NET.
LABOR, $ 9.00/HR.
FIBERGLASS MAT, f 0.16/FT
NOTE *
OTHER HOURLY OPERATING COSTS
ARE INSIGNIFICANT
PiJRITREAT(TM} CAPITAL COST
PUR/TREAT(TM) 4/hR. TOTAL COST
SCFM X 1000
FIGURE 8
CAPITAL & GAS OPERATING COSTS FOR THERMAL INCINERATION
WITHOUT HEAT RECOVERY AND PUR/TREAT (TM)
-------
ioooA
IOO"\
FIGURE 9
CAPITAL 3 TOTAL OPERATING COSTS FOR H.E.A.F. a
PUR IT RE AT (TM)
CAPITAL COST $M, H.E.A.F.
/
/
/
/
/
/
/
/
/
y
/
/
/
V
/
/
804
Ul
<£>
U1
$
60-\
40-
20 A
CAPITAL COST ?M, COKE FILTER-
TOTAL OPERATING C0ST4/HR, H.E.A.F. ON COOLERS -
TOTAL OPERATING COST#/hR. H.E.A.F. ON P.L
TOTAL OPERATING COST #/HR., COKE FILTER ON COOLERS-
TOTAL OPERATING COST#/hR, COKE FILTER ON P.I.
5
—r
io
15
—r~
20
—r~
25
~J-
30
SCFM X 1000
-------
ADDITIONAL ADVANTAGES
The obvious advantages of the Puritreat(TM) system are
mainly concerned with minimizing the capital and operating
cost. Both of these are lower than for comparable systems
such as thermal incineration, HEAF filtration with disposal
of the filter medium, or wet scrubbing with disposal of the
environmentally sound waste water and sludge.
Several less obvious advantages are also inherent in the
Puritreat(TM) system.
(1) The recovered material recycles to the fur-
naces and is incorporated into the product,
adding, if only a small amount, to the pro-
duct yield, rather than to air pollution.
(2) There is no "byproduct" to dispose, as
would be the case with either high energy
filtration using disposable media, or with
water scrubbing.
(3) Incineration - if it could be practiced
in view of the very high fuel costs involved-
would probably not effectively remove the
heaviest polycyclic aromatic compounds and
might, conceivably, add to the emission of
benzo-pyrene and similar objectionable sub-
stances. The carbon bed is effective in
removing such materials.
(4) Because the process is completely "dry"
(even though the collected material may be
liquid or semi-solid, it is handled as a
dry solid in combination with the granular
collection medium), the problems associated
with wet systems such as corrosion and ero-
sion are avoided. This contributes to in-
creased reliability.
396
-------
Ft GURE fO
SKETCH OF PILOT UNIT
r COUP
§ DIA. HOLE
—FOR PITOT
SILENCER -
4" 1.0. x 4-6"LONG
Tee FOR
PITOT
BUFFALO FORGE
BLOW BR WI82
S DIA-
¦FLUE PfPE
G.E. MOTOR 345O RPM
D/RECT DRIVE
SLOPE DISH
BOTTOM
-------
TABLE 1 TEST RESULTS FOR PARTICULATES
Test No.
Lo cat i on
Weight of
Parti culate
Og)
Loading
(grains/ft3)
Particulate Emission
Rate C g/sec)
1 *
Inlet
1067.8
0. 104
1. 03
1
Out 1et
52. 3
0.00041
0.0529
2
Inlet
1006.0
0. 373
4. 16
2
Outlet
20.1
0.0064
0.0875
3 *
Inlet
422. 7
0. 107
0. 94
3
Outlet
9.9
0.00319
0.0241
4
Inlet
00
00
0.0287
0. 302
4
Out let
18. 2
0.00646
0.0823
5
Inlet
00
VD
o
0.0256
0. 140
5
Out let
16. 8
0.00423
0.0301
*Glass fibre filter was contaminated by condensed oil running down the sampling
probe in the vertical position.
-------
TABLE 2 SUMMARY OF DUCT AND STACK CONDITIONS
Average Flow
Average Stack
Particulate
Locat ion
Rate (scfm)
Gas Temp.
Loading
(° F)
(grains/ft3) *
Inlet
8, 470
67. 2
0.1424 1 J
Out let
10,620
00
*fr
r 21
0.0047 1 J
* Dry gas at 70°
Moisture content was assumed to be approx. 2%
(1) Simple arithmetic average excluding contaminates samples 1 and 3
(2) Simple arithmetic average
-------
TABLE 3 PURITREATCOLLECTION EFFICIENCY
: — (TM)
Test
Particulate
%
Benzo (a) Pyrene
%
Benzo (k) Fluorathene
%
1
(contaminated inlet filter)
-
-
2
97.9
86.0
79. 3
3
(contaminated inlet filter)
-
-
4
72. 7
94. 9
89. 4
5
78.5
98.3
99.2
AVERAGE
83. 0
93. 1
89 . 3
-------
TABLE 4 AVERAGE TEST LOADING
Location
Average Particulate
Loading
(grains/ft3)
Average Benzo (a) Pyrene
Loading
(grains/ft3)
Average Benzo (k) Fluorathene
Loading
(grains/ft )
Inlet
Outlet
(1)
0.142
0.0047 C2')
7.2 x 10" 5 ^
3.6 x 10"6 ^
CD
1.2 x 10 L J
-6 f 21
7.6 x 10 L J
(1) Simple arithmetic average excluding contaminated samples 1 and 3
(2) Simple arithmetic average
-------
APPLICATION OF THE PURITREAT(TM) SYSTEM
The application of the Puritreat(TM) system to existing
processes is relatively straightforward when on-site pilot
studies can be used to characterize the variables. The fol-
lowing design parameters can be evaluated in the pilot unit
and an optimum combination selected for the particular job.
1. Filter particle size, which is generated by screen-
ing the charge material to obtain a "cut" with good
filter characteristics.
2. Bed depth is generally under 6" and can be varied
considerably.
3. Filter velocity must be high enough to provide good
collection efficiency without unnecessary fan power
usage.
4. Temperature should be low enough to assure conden-
sation of all of the heavy organics prior to the
filter bed.
A reasonably complete test program can be carried out at
an operating plant for a cost of several thousand dollars.
For new equipment, test data of this sort are unavailable,
and it is necessary to design conservatively and estimate effi-
ciencies on the basis of similar operations. The results given
in Tables 1 through 5 are typical of a carbon electrode pitch
impregnation process.
402
-------
SUMMARY AND CONCLUSIONS
The Puritreat(TM) system can be used for abatement of
tarry particulate emissions from industrial processes which
use carbonaceous solids as raw materials at acceptable effi-
ciency levels. In addition to the particulate reduction, the
emission of objectionable polycyclic aromatic materials from
pitch impregnation and similar processes is good.
Operating conditions should be selected or confirmed by
pilot test work if possible, and conservative design para-
meters based on other similar applications used for new in-
stallations .
The Puritreat(TM) system offers an environmentally sound
and economical alternative to incineration and fixed element
filtration.
403
-------
PILOT-SCALE FIELD TESTS OF HIGH GRADIENT MAGNETIC FILTRATION
by
C.H. Gooding and C.A. Pareja
Research Triangle Institute
P.O. Box 12194
Research Triangle Park, N.C. 27709
ABSTRACT
Bench-scale and lab pilot-scale experiments have indicated that high gradi-
ent magnetic filtration can be an efficient and economical method of controlling
particulate emissions from selected sources in the iron and steel industry. A
5100 m3/hr (3000 CFM) mobile pilot plant has been constructed, and tests are
underway at a Pennsylvania sintering plant. The objective is to characterize
the collection efficiency of the process with respect to gas throughput, applied
magnetic field strength, collection media construction, and particle size. When
the tests are completed, the data will be used to make economic projections of
the application of magnetic filtration to sintering plants and other iron and
steel processes. This paper describes the HGMF pilot plant and gives a status
report of the field program.
404
-------
PILOT-SCALE FIELD TESTS OF HIGH GRADIENT MAGNETIC FILTRATION
BACKGROUND
Since 1975 the Research Triangle Institute has been working under EPA
sponsorship to investigate and develop high gradient magnetic filtration
(HGMF) as a particulate control process. HGMF is a magnetically enhanced
filtration process that has been proven commercially to be an effective and
economical method of separating fine, weakly magnetic particles from non-
magnetic solids and liquids (Oder, 1976).
The work at RTI began with a theoretical analysis of the particle collec-
tion mechanisms active in HGMF and a study to identify potential applications
in particulate emission control. It quickly became evident that the most
likely candidates for application were the iron and steel processes listed in
Table 1. All of these processes produce large quantities of waste gas that
contain significant concentrations of magnetic particulate. The first experi-
mental work at RTI was conducted on a bench-scale HGMF unit with a flow
capacity of 130 m3/hr (75 CFM). Dust from a basic oxygen furnace (BOF) was
dispersed in air and passed through the HGMF device. Variations were made in
the superficial gas velocity, the applied magnetic field strength, and the
density and depth of the steel wool filter. These experiments demonstrated
that BOF dust could be collected with greater than 99 percent efficiency.
A 2550 m3/hr (1500 CFM) pilot plant was subsequently designed and con-
structed at EPA's Particulate Aerodynamic Test Facility located in Research
Triangle Park, N.C. Additional tests were conducted using BOF dust as well as
dusts from an electric arc furnace (EAF) and a sintering machine. Once again
the fractional collection efficiency was determined under a variety of operating
conditions. The ranges of conditions tested are summarized in Table 2.
Figure 1 shows an experimentally determined magnetization curve for each
of the dusts as a function of applied field. The BOF dust was particularly
easy to collect because of its high magnetization. Satisfactory collection
efficiencies were also obtained with the electric arc and sinter dusts although
somewhat higher fields and more dense or deeper filters were required.
Increasing the superficial gas velocity was observed to have a relatively weak
but detrimental effect on collection efficiency with all dusts. Collection
efficiency increased with particle size up to about 2 ym, above which the
efficiency tapered off apparently due to reentrainment effects. As expected,
temperature had no significant effect on particle collection within the range
tested.
A preliminary economic evaluation showed that a full-scale application of
HGMF to any of the three steel processes should require a capital investment
in the range of $1.8-3.6 per m3/hr of gas flow ($3-6/CFM) for uninstalled
primary equipment; i.e., magnets and power supplies. The total energy require-
ment should be in the range of 3.2-4.8 kJ/m3 (2-3 hp/1000 CFM) to power the
magnets and the fans that would be required to move the waste gas through the
filters.
405
-------
Table 1. CANDIDATE PROCESSES FOR APPLICATION OF HGMF
Process
Dust
Concentration
g/m3
Mass Median
Diameter
Vim
Iron
Composition
% Total Fe
Noteworthy Gas
Characteristics
Sinter Machine
Windbox
1-2
10
25-50
5-15% H2O, hydrocarbons,
fluorides, SO , 120-180°C
X
Discharge End
5-12
10
25-50
120-180°C
Blast Furnace
10-25
20
35-50
20-40% CO, 2-6% H2,
200-300°C
Basic Oxygen Furnace
Open System
10-25
1
55-70
250-300°C
Closed System
40-70
2
55-70
70% CO, 250-300°C
Electric Arc Furnace
0.2-7
1
15-40
40-120°C
Open Hearth Furnace
4-7
5
55-70
7-15% H20, 250-350°C
Scarfing Machine
0.5-1
0.5
50-70
H2O saturated, 50-60°C
-------
Table 2. CONDITIONS TESTED IN PREVIOUS HGMF PILOT PLANT
Dust source:
Basic oxygen furnace
Electric arc furnace
Sintering machine windbox
Filter material:
AISI Type 430 stainless steel wool,
medium grade (equivalent strand
diameter = 100 ym)
Filter depth:
0.075-0.30 m
Filter packing
density:
0.0050-0.0100 volume fraction
Superficial gas
velocity:
4.9-11.9 m/s
Applied magnetic
field:
0.06 tesla
Gas temperature:
25 and 130°C
Filter pressure
drop:
0.15-6 kPa (0.6-25 inches H20)
407
-------
220
210
200
BASIC OXYGEN FURNACE
190
180
170
160
150
140
o 130
E 120
£110
ELECTRIC ARC FURNACE
<100
2 90
70
40
SINTER PLANT
0 20 40 60 80 100 120 140 160 180 200 220 240 260 280
APPLIED FIELD, A/m * 10 3
Figure 1. Magnetization curves of three steel industry
dusts.
408
-------
Additional details on theoretical aspects of the HGMF process and more
complete results of the bench-scale and pilot-scale experiments are available
in a series of published papers and reports (Gooding, et al., 1977; Gooding,
et al., 1978; Gooding and Drehmel, 1979). The remainder of this paper describes
the third stage of experimental development that is now being conducted by
RTI.
OBJECTIVES AND SCOPE OF PILOT-SCALE FIELD TESTS
In August 1977 RTI received funding from EPA to design, construct, and
field test a 5100 m^/hr (3000 CFM) HGMF pilot plant. Backed by experimental
data and economic projections derived from the previous work, RTI approached
the American Iron and Steel Institute to cultivate their interest and support
for the intended field program. In December, 1977 a presentation was made to
the AISI Technical Committee on Environmental Quality Control. As a result of
that presentation a member company, of AISI expressed an interest in providing a
sinter plant site for the field tests. Although sinter dust was not the
obvious choice for a first evaluation of HGMF based on dust magnetization, it
represents one of the more difficult environmental control problems presently
facing the steel industry. Hence, selection of the field test site was influenced
by the need of an alternative particulate control technique as well as by the
potential for successful application of the HGMF process. Following a period
of planning and negotiation, a Pennsylvania sintering plant was selected as the
site for the field tests.
The overall objective of the field program is to develop sufficient infor-
mation to compare the economics of a full scale HGMF application to the economics
of competitive control technologies. To accomplish that objective, tests are
first being conducted to verify data on the effects of important operating param-
eters of the HGMF process. We will then seek to demonstrate reliable and
efficient operation of the HGMF pilot plant for up to 500 hours while the sinter
plant is operated in a routine manner.
DESCRIPTION OF THE PILOT PLANT
The mobile pilot plant is housed in a 12.8 m (42 ft) freight van. It has
a nominal flow capacity of 5100 m3/hr (3000 CFM) based on the previous experi-
mental work although the ultimate practical capacity will be determined from
the field data. Figure 2 is a flow schematic of the portion of the system
that is inside the trailer.
The dirty gas enters the trailer via a 0.317 m ID, 316 stainless pipe (12",
Schedule 5) and passes by test ports through which samples can be drawn to
determine the size distribution and concentration of the dust. The gas is then
directed to one of two functionally identical HGMF devices. Two filtration
paths are provided so that one can be cleaned while the other is in operation.
Magnet A was constructed by Magnetic Corporation of America (MCA), Waltham,
MA, and Magnet B was originally constructed by Sala Magnetics, Inc., Cambridge,
MA for use in the earlier pilot plant. Magnet B was later modified by the
addition of new pole pieces, stand, and DC power supply purchased from MCA.
Each of the magnets now consists of an iron-bound solenoid surrounding a
canister that measures 0.432 m (17") ID by 0.305 m (12") long. The filters
409
-------
PINCH
VALVE
CLEANING
AIR
STORAGE
V8A
V6A
CLEAN GAS
TEMPERATURE
V7A
CLEAN
GAS
CLEAN
GAS
DIRTY
GAS
DIRTY GAS
TEMPERATURE
V7B
V6B
PINCH
VALVE
MAGNET
CLEANING
AIR
STORAGE
CYCLONE
Figure 2. Flow schematic of HGMF pilot plant.
-------
contained in the canisters are made of medium grade AISI Type 430 stainless
steel wool. The equivalent cylindrical diameter of the irregular fibers is
approximately 100 ym. Each of the magnets is capable of producing an applied
field of up to 0.5 tesla throughout the canister volume.
After passing through the filter, the cleaned gas travels past another set
of test ports and exits the trailer. It passes through an orifice and an induced
draft blower and is then exhausted to the atmosphere through a 10.7 m (35 ft)
high stack. An additional set of test ports is provided in the stack to give
sufficient access for Method 5 testing.
The filters are cleaned by backflushing with a pulse of compressed air.
The volume of the air tank associated with each filter is approximately 0.28 m^
(10 f£^). The compressed air is released through an 8" hypalon-lined, Galigher
Delta valve (Galigher Company, Salt Lake City, UT), which is a pneumatically
activated pinch valve with a throat diameter of approximately 20 cm. The
agglomerated dust that is blown off the filter is removed from the pulse air by
a cyclone. Any carryover is routed back into the dirty gas stream.
The blower that moves the gas through the pilot plant is a Centrifan Model
RB50-2 (Centrifan Company, Greenville, SC). The blower will exhaust the
required 5100 m^/hr (3000 ACFM) at a suction pressure of -18.7 kPa (-75 inches
H2O) and a temperature of 150°C (300°F). With the exception of the Galigher
valves, pneumatically or manually actuated Norris butterfly valves (Dover
Corporation/Norris Division, Tulsa, OK) are used throughout the system. The
entire system is designed so as to allow continuous operation at up to 205°C
(400°F).
The front quarter of the trailer contains an enclosed, air-conditioned
laboratory and control room. Two automated devices are incorporated to simplify
operation of the pilot plant. A Xanadu Model UPT100-10-10 solid state program-
mable timer (Xanadu Controls, Springfield, NJ) sequences the operation of the
butterfly valves, the pinch valves, and the magnets as the system cycles from
one flow path to the other. The total cycle duration and the sequence of
events may be easily changed in a few seconds by inserting a pencil-coded
computer card and adjusting a thumbwheel switch. A Robertshaw DCM-1000 controller
(Robertshaw Controls Company, Anaheim, CA) maintains constant gas flow through
the pilot plant. The controller receives its signal from an orifice located
in the clean gas pipe via a differential pressure transmitter and adjusts a
butterfly valve located on the blower exhaust. The orifice pressure drop and
the pressure drop across the magnetic filters are displayed on two strip chart
recorders.
The laboratory/control room also contains bench space, a wet sink, a lab
oven, a solvent sink and a lab hood. MRI Model 1502 cascade impactors (Meteor-
ology Research, Inc., Altadena, CA) are used to determine the particle size
distribution and concentration. A Perkin-Elmer Model AD-2Z microbalance
(Perkin-Elmer Corp., Norwalk, CT) is ust.l to weigh the impactor substrates. A
Climet Model 208A particle size analyzer (Climet Instruments Co., Redlands, CA)
is also available to monitor transient, conditions in the particle size distribu-
tion and concentration on the clean side of the magnetic filters.
411
-------
Figures 3 through 6 show several views of the pilot plant as it is
presently installed at the sintering plant.
STARTUP AND PRELIMINARY RESULTS
The pilot plant arrived at the sintering plant on June 18, 1979. Over
the next several days, the exterior piping and insulation were installed, and
the water and electrical connections were made. Initial checkout and de-
bugging of the system then began. The major components of the system operate
off 480 VAC, which was not available at the fabrication site, so most of the
equipment had not been operated prior to the field startup. A number of minor
problems were identified and corrected. The 3600 rpm blower created a severe
noise problem in the work area, but this was solved by installing a silencer
on the stack and an insulated enclosure around the blower.
The first fully integrated operation of the pilot plant was accomplished
on July 23. A factorial test was begun to evaluate the effects of superficial
velocity, applied magnetic field, filter density, and filter depth on particle
collection. Preliminary results from the first few tests show the outlet dust
loading from the pilot plant to be approximately 0.11-0.18 g/Nm3 (0.05-0.08
gr/DSCF), which is not low enough to meet applicable emission regulations.
These runs were made with a relatively porous filter, however, so the results
should improve as the tests proceed. The filter cleaning system appears to be
working quite well with a pulse air pressure of 170 kPa (25 psig) and a total
cycle duration of 15 minutes. The filter pressure drop increases linearly
with time during the 7 1/2 minute filtration time and returns to its original
value when the filter is cleaned and returned to service. Additional tests
are planned to study the transient effects of loading on dust penetration
asing the Climet particle analyzer. The transient pressure drop and pene-
tration characteristics will provide valuable data to select the optimum cycle
iuration for long-term operation. The results of the factorial tests will
ilso be used to select appropriate levels of the operating parameters for the
Long-term, continuous operation that will be conducted later in the field
>rogram. The field work will be completed by November 1979, and a full report
giving the results and conclusions from the tests will be submitted to EPA in
Jecember.
ACKNOWLEDGEMENTS
The authors are grateful to Doug VanOsdell, John Sauerbier, and Daryl
mith of Research Triangle Institute for their contributions to the design,
onstruction, and initial operation of the pilot plant.
412
-------
Figure 3. HGMF mobile pilot plant setup at the sintering plant.
im/iTPumiu
njf fAcmrv
Figure 4. Side entrance to pilot plant; gas from plant
enters via pipe that passes under steps.
413
-------
Figure 5.
Rear view of pilot plant; gas enters
from plant at right and exits through
stack at left.
Figure 6. Clean gas piping, blower enclosure
and stack with silencer.
-------
REFERENCES
Oder, R. R. High Gradient Magnetic Separation Theory and Applications. IEEE
Trans. Magn. Mag-12: 428-435, September 1976.
Gooding, C. H., T. W. Sigmon, and L. K. Monteith. Application of High-
Gradient Magnetic Separation to Fine Particle Control. EPA-600/2-77-230.
NTIS No. PB 276663/AS, November 1977.
Gooding, C. H., T. W. Sigmon, L. K. Monteith, and D. C. Drehmel. Application
and Modeling of High Gradient Magnetic Filtration in a Particulate/Gas System.
IEEE Trans. Magn. Mag-14: 407-409, September 1978.
Gooding, C. H., and D. C. Drehmel. Application of High Gradient Magnetic
Separation to Fine Particle Control. JAPCA. 29: 534-538. May 1979.
41S
-------
EXPERIENCES WITH CONTROL SYSTEMS
USING A UNIQUE PATENTED STRUCTURE.
By:
George C. Pe&ersen, P.E.
KIMRE, INC.
This paper will present practical design
and operating experiences using Kimre's patented
structure for particulate removal. The vide
range of sizes allows unique capabilities:
effectiveness to one-half micron or below can be
achieved, as can unexcelled pluggage resistance
and cleanability.
416
-------
EXPERIENCES WITH CONTROL SYSTEMS
USING A UNIQUE PATENTED STRUCTURE.
I. SUMMARY
The unique structure used by Kimre, Inc. is described and its uses
explained. Kimre's history, position, and approach is described. Then a
summary listing our experience is provided with an indication of the of
uses. The major part of the paper is consideration of four case histories:
siliceous particulate,
fluoride,
urea, and
"smoke."
This structure as used in the patented B-GON™ mist eliminator and
KON-TANE™ Tower Packing products, has been proven effective in an extreme-
ly wide range of problems.
II. INTRODUCTION AND BACKGROUND
The patented structure is totally unique in the industry. It is based
on two sets of monofilament fibers interlacing perpendicularly in a ladder
like arrangement, with specific three dimensional orientation between the
fibers. This structure provides a number of novel features compared to
other materials:
(a) Any type of monofilament materials can be used, in-
cluding metals. However, only thermoplastics are
routinely offered commercially.
(b) The fibers lie perpendicular to the gas flow for
maximum effectiveness.
(c) The interlacings involve a very small fraction of the
fiber but yeild a high compressive strength and high
void fraction material.
(d) A very wide range of fiber diameters is easily
produced providing the great design flexibility to
handle a variety of circumstances.
(e) The media is made in large sections; each over
1830 mm (6*) wide by many feet long.
(f) Since the individual fibers are straight, or nearly
straight each in its own direction, the media has
extreme dimensional stability and is very strong.
This allows the large pieces to be installed with
out difficulty and minimal support.
Cg) Build up of foreign materials, or corrosion
products does not cause "Mechanical, shrinkage,"
417
-------
as has occurred with knitted mesh materials particular-
ly for metals.
(it) The structure is easily, and always, heat stabilized to
well over the recommended operating temperature provid-
ing a substantial safety margin.
(i) Although strong, and resilient, thermoplastic materials
are compressible, flexible and almost impossible to
damage mechanically. They can be rolled up, twisted,
compressed, jumped on, and in fact be driven over.
(j) Flexibility and compressibility is particularly interes-
ting since it seems to be this feature that makes the
material uniquely cleanable.
Since this paper deals with some specific details I must familiarize you
with the standard nomenclature.
1. Our standard style designation splits fiber diameters
and void fractions by a slash with fiber diameters
(in mils) on the left and the void fraction
(percentage) on the right. Thus Style k/S6 is made of
H mil material and has a nominal 96# void fraction.
2. The material of construction is shown separately.
Polypropylene, Kynar™, Teflon™ and Tefzel™ are
common materials of construction (Kynar is the trade-
mark of Pennwalt Corporation for their polyvinylidene
fluoride, Teflon and Tefzel are trademarks fcr the
DuPont Corporation for some of their fluoro-polymer
resins.)
Kimre, Inc. was formed in 1973 and became active in 197^. It is a small
privately held Florida Corporation with Headquarters and Manufacturing facili-
ties in Perrine, Florida (a Miami suburb).
Kimre's position is unusual, possibly even unique. Primarily manufac-
turers, suppliers of our media (and technical expertise in their use), this
company also supplies consulting services, conceptual design, and accessories
such as support grids. Detailed engineering designs have been supplied
although our preference is to work through, and in conjunction with, original
equipment accounts and engineering companies.
We do not make scrubbers, absorbers, or any type of vessels.
Since 197^ we have grown at a gratifying rate, this has been done by
boot strapping, without any goverment assistance or outside financing; in fact,
the only real support has come from the Pennwalt Corporation. As a consequence
Kimre operates on a very careful financial balance without any excess discre-
tionary funds being available. I mention this, not to impress you with our
418
-------
"Great American Initiative" "but rather to lead into the consequences of our
position which have a hearing on our market posture and the type of results
that we will report:
We cannot spend much money on testing, test facilities,
pilot plants, or any thing else not having a significant
dollar return in a reasonably short period.
We cannot commit resources on large high risk project
areas even if we have the best available technology.
When the Company originally started, as you might imagine, we got projects for
which there were severe problems: usually an emergency where our GLS priority
service allowed delivery before anyone else, or correcting a problem for which
there seemed no other viable solution. Our product range, expertise, and
acceptance have increased substantially. Many scrubber manufacturers now use
our products regularly for routine as well as for especially difficult
operations.
Scientists may be shocked at our approach and we anticipate criticism
at the lack of specific information. Most of our customers are not willing
(or at least not anxious) to authorize publication of company identification
much less specific test information. Of course Kimre is not staffed to carry
through, on details to obtain approvals.
Our approach is frequently crude, and is always results oriented over
99% of our thousands of installations have been successful. Specific case
histories are given below and are discussed in some detail. A partial
summarization of our experience in air pollution control and chemical
processing follows:
(a) Pulp and Paper for:
Smelt dissolving tank vents
scrubbers on recovery boilers
Chlorine absorbers
Chlorine dioxide absorbers
Chlorine scrubbers
Hydrogen recovery
starch or clay dust
(b) Sulfuric Acid for:
intake filters of sulfur burners
drying towers
inter-pass absorbers
final absorbers
Sulfuric acid concentrators
by product acid plant
Oleum vaporization
419
-------
(c) Phosphorous. Processing for:
wood process phosphoric acid fluoride scrubbers
Mono-ammonium phosphate
Di-ammonium phosphate
GTSP
storage building dust scrubbing
rock dryers
electric furnace
food grade phosphoric acid production
(d) Chloro-Alkali for:
Chlorine drying
Hydrogen purification
brine filtration
(e) Others including:
HC1 scrubbing from a variety of sources
Hydrazine
oil mist removal
odor and dust from foundries
odor from sewage treatment plants
odor from animal by-products
fumes such as from brazing lines
organic vapor absorption into liquid
cooling towers
pickling
plating
fiberglass insulation
secondary smelters
desalination evaporators
etc. etc.
(f) One other interesting application area is incinerators,
we have had particular success on chemical waste
incinerator application.
Our experienced areas include: high vacuum to high pressure, cryogenic
service to over 200°C, rates from one cubic nr/hr to 860,000 m3/hr (h CFM to
500,000 CFM), particulate and mist removal from submicron to rocks or rain
drops, efficiencies from minimal effects to sub ppb.
III. CASE HISTORIES
A. Phosphorous Furnance Application
B. Phosphoric Acid Fluoride Scrubbing "
C. Urea
D. "Smoke" Scrubber
420
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A. Phosphorous Furnace Application (Prefilter for a candle-type unit.)
In 1976 our engineering firm clients approached us to help with the
design of a pollution control system for an electric arc-phosphorous furnace.
It was necessary to control water soluble P2O5 and water insoluble siliceous
materials. The materials ranged from over 10 microns in diameters to much
less than half a micron. It was a nasty and fairly difficult job. Size
distribution of the components is shown in Table I.
A high-efficiency candle-type mist eliminator would remove the
emissions but would plug up from the insoluble materials. The approach taken
was to use the Kimre materials to remove the bulk of the solids down to
approximately one-half micron. Removal of the coarser particles would
protect the candles from pluggage.
A small, 1+00 mm (16") diam pilot plant was installed with a two-page
configuration. The details of the configuration and design parameters are
shown in Table II. The pilot plant operated as predicted, except that
pressure drop was slightly high. The reason for the high pressure drop was
never determined although no pluggage occurred. Possibly high irrigation
rate or higher than design velocities were responsible. Our pressure drop
estimations are usually quite reliable or even a little high.
This story has, for us, a sad ending. Our joint venture technology was
not accepted and the job went to another company. There may, however, be a
silver lining: the unit selected has never worked well and our same concept
is now being considered for a new installation by the same client. We have
also learned that the original pilot plant was considerably over-designed and
a much lower pressure drop unit would have been quite acceptable.
421
-------
Table I.
Droplet
Diameter
>l*tU
1U-9U
9-6y,
6—b • 2\l
1+.2-2.7U
2.7-l.^U
1.1*-.85u
.85-.56U
c 5611
% of P205
% Insolubles
10.7
1+0
.16
2
ro
ro 0
.23
2
• 59
17.1
20
18.1
10
26.1
10
28.5
10
Design Pene-
tration of
Insolubles^
in %
.1
2
98
iThe insolubles are primarily siliceous in nature.
Table IX.
PAD A
3 Layers Style 32/9^, 3 layers Style 16/96, 2 layers Style 8/96, 5 layers
Style U/96, 1 layer Style 8/96, 1 layer Style 32/9I+.
FAD B
18 layers Style 2/96, 1 layer Style 8/96, 1 layer Style 32M.
This design option for a l6" diam, 600 fpm, pressure drop of 8.5"
yields an estimated effectiveness of 96% overall on all insulubles with
98# of remaining particulates less than .56 microns (p.). Note that the
solubles were not a factor in the design.
422
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B. Phosphoric Acid Fluoride Scrubber
This is a central Florida phosphate processing complex using a conven-
tional crossflow scrubber design. This unit had never operated satisfactorily
and work has been in progress since 1912. A great many steps were taken to
improve the performance. Yet, despite this ongoing program with many changes
and tests made, the unit was never capable of operating consistently below
0.0U5 Kg fluoride/ton P2&5 (0.1 lb fluoride/ton P2O5). One of their consultants
had indicated that mist carryover was a problem, and a high-efficiency chevron-
type mist eliminator (imported from Germany) was being utilized.
As a result of other work in the same complex we became aware of this
problem in late 1976 and in early 1977* However, we were not called in to
help solve the problem until July, 1978. At that time a review of the test
data made it very clear that the conventional "approach-to-equilibrium"
situation was probably not the major problem. We recommended a test sequence
to the plant personnel. These tests were selected to define the "form" of
the fluoride being emitted. Without getting into the technical details of
the tests run we determined that emissions from the unit were a combustion
of gaseous fluorides as well as fluorides in non-gaseous form. An appreci-
able percentage of the emitted fluorides were in non-gaseous form. We
determined further that the non-gaseous fluorides were in a physical form that
was not effectively collected by the equipment being used. A significant
portion of the fluorides (present as non-gaseous fluoride) was in the size of
1-5 microns. This size is substantially below the size that is effectively
collected in either the chevron-type mist eliminator or in the dumped packing.
In order to achieve the desired results we suggested a change in the
scrubbing section of the unit to include the addition of a small amount of
our media. Specifically we suggested the addition of two layers of Style
37/97» followed by two layers of Style 37/9^ in polypropylene with a total,
thickness of 90 mm (3.6"). To install the media it was necessary to relocate
some spray headers. Kimre KON-TME®* media was installed directly ahead of the
existing retaining frame and attached directly to it. Thus, there were no
structural changes on the inside of the unit. In front of the existing
chevrgn-type mist eliminator we suggested a high-efficiency B-GON™ Mist Elimi-
nator be installed. We suggested that this be comprised of two layers of
Style 37/9^ and three layers of Style 16/96, plus three layers of Style 8/96 in
an irrigated mode of use. The mist eliminator was supplied as two pieces
2390 mm x 1830 mm (9V' x 72") and two pieces 1780 mm x 6l0 mm (70" x 2h") to
conform to the superficial area of the chevron-type mist eliminator. These
76 mm (3.0") thick units were installed directly in front of, and attached to,
the chevron-type mist eliminators.
The units were installed in October 1978, during one of the plants
regular down days. Several problems occurred. First of all we had bypassing.
It was necessary to put in a sealing piece around the edges, for which we have
standard procedures. After this they found that the test results (although 50%
better than the previous results) were not those anticipated. A check of the
unit showed that the Style 8/96 material had sealed up with siliceous deposits
as is a commonly problem with these operations. The Style 8/96 was removed.
423
-------
Subsequent to this change the unit has consistently achieved the
objective of 0.0091 Kg fluoride/ton P2O5 (0.02 lb fluoride/ton P2O5). It has
been necessary to clean the media periodically as is easily determined "by
pressure drop of only a few inches WG.
Beplacement of the original type of packing with K0N-TANE® Tower Packing
is planned.
424
-------
C. Urea
This was a retrofit on existing prilling towers, with details of the
installation accomplished on the basis of pilot plant testing performed "by
customer's personnel. The pilot plant test included electrostatic precipita-
tors, charged droplet scrubbers, and Venturis. The venturi required seven
times the pressure drop of the KON-TANE®1 Tower Packing and B-GON™ Mist
Eliminator materials to achieve the same collection. The plant has a
capacity of 1+50 tons/day and an initial emission rate of 35 kg/hr (77 lbs/hr).
Size distribution given as: 90# minus two microns, 85# minus one micron, 30#
minus 0.5 microns, and a median diameter of 0.75 microns.
The scrubbing section is comprised of a total of 17 layers of materials,
using our Styles 37/9^» l6/97> and 8/96. The mist eliminator section uses
the same materials. However, it has only 11 layers and is a different
composite. The scrubbing modules are installed on top of the stacks. They
are set up so that the flow comes up vertically in the internal of the unit,
turns and moves radially, and horizontsilly, out through the scrubbing sec-
tion, the mist eliminator section and then out to the discharge. Each of the
six units handles 85,000 m3/hr (50,000 CFM). The scrubbing sections operate
at 2.3-2.1+5 m/sec (j%-8 fps) while the mist eliminator section operates at
1.5 m/sec (5 fps).
It was found that the liquid rate control is critical, having a
requirement of approximately .1+3 l/m^ (1+ gal/1000 ft^). Emissions as low as
11 mg/m^ (0.001+8 g/ft3) were obtained). Pressure drops of only 90-100 mm WG
are experienced in operation. All the units have been in service
since Fall of 1977. There is no indication of anything to suggest any
difficulties with the units which will limit their service life.
These units were unique in that the scrubbers were designed to fit on
top of the existing stacks and caused a very low pressure drop. This
obviated the necessity for massive reinforcing, or replacement, of the
existing stacks; or the need to duct the entire gas flow down to ground level
for treatment and then issue through a new stack. Installation in this
manner obviously saved a lot of money and down time.
We think it interesting that the measured performance far exceeds
projected theoretical performance. Actually when we were first approached on
this there was a serious question in our mind whether or not we could achieve
suitable results. Similar things have been observed, prior to this
installation, and subsequent to it. The reasons for this delightful discre-
pancy are not known but are possibly due to the multitude of many Venturis
(up to 62 million/m3 in Style 2/96) or the interaction of von Karm&n eddies
from one set of filaments which intersect the next set of filaments at right
angles. Whatever the reason the effect is dramatic and always occurs to
extend the useful range of collection.
425
-------
D. "Smoke" Scrubber
In November of 1976 one of our original equipment manufacturer accounts
(who supplies a variety of wet scrubber types) came to us with a "smoke"
problem where they had installed a conventional crossflow packed-bed scrubber
with 50 mm (2") packing. This was a 6800 m3/hr (1+000 CFM) scrubber with 2515
mm (8*3") of irrigated packing and a single mist eliminator stage which we
had provided.
The "smoke" was originating from a plastic extrusion operation where
nylon was being extruded and chopped into pellets. When the hot plastic
contacted the cool room air significant quantities of "smoke" were given off;
and as a result, everything in the area was coated with an oily coating. This
"smoke" is a composite of many different things, a combination of caprolactam
and oligomers, forming what seemed to be a mixture of mist, solid and gelled
particles. The original installation was essentially of no effectiveness in
removal of this "smoke." No mass emission figures were provided. Some
preliminary tests indicated that our material, Style U/96 in polypropylene,
would definitely remove the "smoke." However, provision had to be made to
handle gels and solids. A three-stage system was eventually installed in the
shell left after removal of the conventional dumped packing.
The final stage, Pad C, is a simple mist eliminator operating
at around 2.5 m/sec (500 fpm).
Pad B, the intermediate stage, is irrigated with fresh water at
a very low rate. This section operates at around h.3 m/sec
(lU ft/sec) and utilizes 100 micron fiber, present in our Style
h/96, as the primary scrubbing material. This stage, the
actual functioning stage on the unit to remove the "smoke", is
less than 50 mm (2^") thick.
The initial stage, Pad A, is irrigated with recycled solution
from a holding tank. This pad is comprised of a variety of
materials designed to give a high degree of pluggage resist-
ance while removing coarse materials and dirt present in the
system from room air. It operates at 2.5 m/sec (500 fpm).
The intake to the actual scrubber is maintained at 50 mm (2") pressure
with a manifold arrangement. The unit is set up so that if any of the extru-
ders is operating it automatically switches into the manifold which causes
an adjustment in the supplementary air, to maintain a constant total flow
through the scrubbing system. This control system is necessary as Stage B
is sensitive to velocity variations. Since its installation over a year
ago the entire unit has been operating at the anticipated 180 mm (7")
pressure differential without difficulty or pluggage. The measured efficiency
has not been released to us. However, it operates at an emission level that
is judged to be satisfactory. Prior to the installation of our material,
the plume was essentially 100% opaque emitting from a 380 mm (15") stack
with even one extrusion line in operation. With our materials in service
it is barely possible to see that anything is being emitted. In addition
426
-------
to that, the high speed blower operated completely dry and erosion of the
blades, a previous operating problem, has ceased.
As a bonus they recover the monomer.
The only other device plant personnel considered proven effective was
a "candle" type mist eliminator. That type of unit:
1. is much larger,
2. is much more expensive,
3. has higher pressure drop and
U. gets plugged up.
427
-------
IV. CONCLUSION
The structure has a wide range of useful applications. The complete
application area has not "been defined. Current limitations are:
for sub micron particulates and mist,
temperature over 260°C, and
application using very high liquid and gas rates for
counter current absorption.
428
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ELECTROSTATIC EFFECTS IN VORTICAL FLOWS
P.W. Dietz
General Electric Company
Corporate Research and Development
Schenectady, New York 12301
ABSTRACT
Electrostatic forces have been shown to significantly enhance the col-
lection efficiency of cyclones.* In this paper, an analytic model will be
developed for the effect of an applied electric field on charged particle
trajectories within a laminar, two-dimensional vortical flow. Boundary con-
ditions for particle injection are developed and collection efficiencies are
predicted.
*W.B. Giles, "Electrostatic Separation in Cyclones"
Symposium on the Transfer and Utilization of Particulate Control Technology,
Denver (1978)
429
-------
ELECTROSTATIC EFFECTS IN VORTICAL FLOWS
INTRODUCTION
In a conventional reverse-flow cyclone (see Figure 1), particle-laden gas
is introduced tangentially to produce a vortical flow. Within this acceler-
ating flow, the particles experience a radial force which tends to propel them
to the cyclone wall. However, the radial motion of the particle is resisted
by the combined action of turbulent mixing and fluid drag resulting from the
inward gas flow. The separative performance of cyclones is thus determined by
the competition between these forces (Leith and Licht 1972 and Dietz 1979).
Figure 1. Conventional Cyclone Geometry
Since electrostatic forces are recognized as having a paramount effect in
determining the particle trajectories in gas-solids flows (Soo 1971), it
would seem obvious to attempt to augment cyclone performance through the
application of electrostatic techniques and, indeed, several early investiga-
tors pursued this approach (Petroll and Langhammer 1962 and Molyneux 1963).
However, in spite of demonstrated improvements in performance, this technique
has not been pursued.
430
-------
Soo (1973) has formulated equations for electrostatic effects in cyclones
and has suggested that they may be important - especially in larger cyclones.
However, his equations require numerical integration and have not been
solved.
Recent experimental work by Giles (1978) has confirmed that electrostatic
effects are significant in cyclones and, indeed, can dominate their perform-
ance. By varying the electrical charge carried by the dust entering the
cyclone, Giles was able to demonstrate large changes in cyclone performance.
Unfortunately, the level of charging was not characterized in these experi-
ments and, thus, information on the fundamental process resulting in the
electrostatically-enhanced performance is not available.
To address the fundamental issues of collection mechanisms and cyclone
scaling, a program was initiated at General Electric Company's Corporate
Research and Development Center. What is reported in this article is a model
for collection in a two-dimensional, electrostatically-enhanced cyclone
(termed "electrocyclone"). Based on this simple model, hydrodynamic and
electrostatic cut-sizes are computed and the implications of the results for
scaling to large cyclones are investigated.
THEORY
The study of charged-particle motions is inherently quite complex due to the
coupled nature of the equations. The motion of the particle is governed by
the flow pattern and the electric field. However, both the flow and field can
be modified by the presence of particles. Thus, to analytically examine the
effects of electrostatic forces on cyclone performance, certain simplifying
assumptions are introduced:
1) the flow is two-dimensional (see Figure 2);
2) the dust loadings are low enough that the particles do not affect the gas
flow;
3) the tangential velocity of the particle is the same as that of the gas;
4) the tangential velocity of the gas is given by the equation for a modi-
fied free vortex (Alexander 1949),
U /r \m
t i _S ) '
Uto ' V r/
where R is the cyclone radius, is the tangential inlet velocity and
0.5 < m <1.0 (note: for a free vortex in an inviscid flow, m = 1);
5) the drag on the particle is given by Stokes' Law;
6) the electric field is imposed (thus, space charge effects are neglig-
ible) ;
7) the cyclone is operating in the steady-state (no transients).
431
-------
Sintered
Metal
Electrode
OUTER
ELECTRODE
Figure 2. Two-dimensional Cyclone Geometry with Distributed
Tangential Inlets and Axial Outlet
With these assumptions, a set of simple, tractable equations can be written
describing the motion of the particles. First, conservation of particles
gives
1
r
d_
dr
r r
= o
(2)
where r is the net radial flux of particles
T = n Upr , (3)
where n is the number density of the particles and U is their radial
velocity at position r. A simple force balance on a particle yields
V - bEt + 0r + 6 i uj , (4)
where U is the radial gas velocity, b is the mobility of the particle in the
radial electric field E ,
r
b =
3 iry dp , (5)
432
-------
and
d2 P
P P
18y '
(6)
where d is the diameter of the particle of mass density p and y is the gas
viscosrty. ^
Since the radial velocity and electric field are not affected by the pres-
ence of the particles, they are given by
= -U
and
o\ r
(7)
= -E
o\ r;
(8)
where
V
E = o
o ~ R In (R /R )
c co
(9)
and VQ is the applied voltage.
Introduction of eqns. (1 and 3-9) into eqn. (2) and integration yields
2m BU
r r
o c
to
R n
L c o
rn
(2m+l)R
2m+l
_c
-1
(10)
where r and n are the particle flux and number density at the outer wall,
o o
Because of the non-linear nature of the centrifugal force, stable orbits are
possible. Particles which enter the cyclone can be "trapped" at a radial
position at which all of the forces on the particle balance. This possibil-
ity is responsible for the singular behavior of eqn. (10). These particles
are assumed to be collected (in practice, they would drop from the system
under the influence of gravity). Thus, all particles above a certain critical
diameter - termed the "cut-size" - are collected. The cut size is determined
by the solution to the equation
8u
bE
to
2m-1
U
(11)
433
-------
where it is important to recognize that both b and 3 are functions of the
particle size! This abrupt transition from no collection to complete collec-
tion is shown in Figure 3.
100%
CUT-SIZE
(STABLE ORBIT)
COLLECTION EFFICIENCY
OF ENTRANCELAYER
Figure 3. Efficiency of Cyclone Including Both Cut Size
and Entrance Layer Effects
To complete the solution to the problem, it is necessary to specify the
inlet conditions. At first glance, it would seem appropriate to set the inlet
particle flux equal to the volume flow rate through the cyclone times the
inlet number density. However, since particle flux conserved, this would
imply that all entering particles pass through the cyclone uncollected (ex-
cept for those trapped into stable orbits). In actual fact, the boundary
condition at the outer wall needs careful examination. Because the flow is
introduced over a finite thickness (see Figure 4), the radial velocity of the
gas at the outer wall goes to zero. Consequently, even extremely fine part-
icles can be thrown against the outer wall and collected.
AIR
Figure 4. Close-up View of Distributed Inlets
434
-------
Direct integration of the governing equations (eqns. 1-9) over this thin
layer results in
r - n
o o
8oto
0 - —— - bE
OR o
c
(12)
Since the total rate at which particles enter the cyclone is given by
r in ' Vo' (13)
the efficiency of the outer entrance layer of the cyclone becomes
r - r + bE
in o to o
ni = —r " u (14)
in o
This result is shown in Figure 3. Note that all particles above a critical
diameter, d , are collected in the entrance layer and do not participate in
the main vortical flow. Thus, only those particles in the shaded region of
Figure 3 are trapped into stable orbits.
DISCUSSION
The efficiency of a cyclone can be characterized by its cut-size. The
general equation for the cut size of an electrocyclone is given by eqn. (11) .
In the absence of electrical effects, the inertial cut size, d is
2m-1
18UU R.
o 1
dpi -/ ^ ^ I • (15)
This result is the classic hydrodynamic cut size for a cyclone (Bhatia and
Cheremisinoff 1977).
A similar cut size can be computed for electrostatic effects once the
mechanism responsible for the particle charge is identified. In many electro-
static particle removal devices, the charge is imparted by a corona charger.
For these systems, the saturation charge obtained by a conducting particle in
a corona charger is given by
q * 3Tre0dpEc ' (16)
where e is the permittivity of free space ( e ¦ 8.854 x 10~*^ F/m) and E is
the electric field in the charging zone.
435
-------
Thus, for a particulate suspension which has received its electrical
charge from a corona source, the electrostatic cut-size, d , can be computed
from eqn. (11) in the limit of zero inertial effects
d
pe
(17)
To compare the electrostatic cut size to the hydrodynamic cut size, it is
necessary to specify a particular cyclone configuration. However, it is
possible to make several general statements about the manner in which these
cut sizes scale.
For example, varying the flow through a given cyclone will alter both cut
sizes. With increasing flow rate, the electrostatic cut size increases while
the hydrodynamic cut size decreases (see Figure 5). This result agrees with
the increase in efficiency observed by Giles at low flow rates.
Figure 5. Qualitative Comparison of the Effect of Volume Flow Rate
on Both the Electrostatic and Hydrodynamic Cut-size of a
Specific Cyclone
As the size of conventional cyclones is increased, the efficiency de-
creases (cut-size increases). This decrease is a consequence of the erosion/
saltation phenomena which limit the inlet velocity to approximately 30 to 40
m/sec. Since the tangential inlet velocity remains constant as the size of
the device increases, the radial velocity also remains constant for geome-
trically similar cyclones. Thus, the cut size increases as the square-root of
the cyclone size.
h
LOG Idpc)
INERTIAL (dpi)
ELECTROSTATIC ldpe)
LOG (Qv)
436
-------
In contrast, the electrostatic cut size is unaffected by increases in
size (at fixed inlet tangential velocity).
A comparison is made in Figure 6 between the effect of size on the
electrostatic and hydrodynamic cut sizes.
h
LOG (dpC)
HYDRODYNAMIC (dpi)
ELECTROSTATIC (dpe)
LOG (Rc>
Figure 6. Qualitative Comparison of the Effect of Cyclone Size
on Both the Electrostatic and Hydrodynamic Cut Size
CONCLUSION
A simple model has been developed for the collection efficiency of an
electrostatically enhanced cyclone. The electrostatic forces are shown to
augment the inertial forces in these electrocyclones resulting in improved
performance.
Equations for the hydrodynamic and electrostatic cut sizes are presented
and these results are used to investigate the scaling of electrocyclones with
size and flow rate.
ACKNOWLEDGEMENT
This work was performed under DOE Contract No. EX-76-C-01-2357 to the Gen-
eral Electric Energy Systems Program Department by Corporate Research and
Development.
437
-------
REFERENCES
R. McK. Alexander, "Fundamentals of Cyclone Design and Operation", Proc.
Aust. Inst. Mining and Met, 152-153, 203 (1949).
M.V. Bhatia and P.N. Cheremisinoff, "Cyclones" Chapter 10 in Air Pollution Con-
trol and Design Handbook, P.N. Cheremisinoff and R.A. Young, editors, Marcel
Dekker, New York (1977).
P.W. Dietz, "Collection Efficiency of Cyclone Separators", submitted
J.A.P.C.A. (1979).
W.B. Giles, "Electrostatic Separation in Cyclones", Symposium on the Transfer
and Utilization of Particulate Control Technology, Vol. 3, EPA-600/7-79-044c,
291 (1979).
D. Leith and W. Licht, "The Collection Efficiency of Cyclone Type Particle
Collectors - A New Theoretical Approach" AIChE Symp. Series Vol. 126, 68, 196
(1972).
F. Molyneux, "Electrostatic Cyclone Separator", Chem and Process Eng, 517
(1963).
J. Petroll and K. Langhammer, "Vergleichsversuche an Zyklonabscheidern" Frei-
berger Forschunsheft A-220, 175 (1962).
S.L. Soo, "Dynamics of Charged Suspensions", International Reviews in Aerosol
Physics and Chemistry Vol. 2, G.M. Hidy and J. Brock, Ed., Pergamon Press
(1971).
S.L. Soo, "Some Basic Aspects of Cyclone Separators", Proceedings of the First
International Conference in Particle Technology, 9 (1973).
438
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CONDENSATIONAL ENLARGEMENT AS A SUPPLEMENT
TO PARTICLE CONTROL TECHNOLOGIES *
By:
Janes T. Brown, Jr.
Department of Physics
Colorado School of Mines
Golden, Colorado 80401
ABSTRACT
An experimental study has been undertaken at the Colorado School of
Mines to evaluate the feasibility of using condensational enlargement to
enhance the collection efficiency of existing dust control technologies for
sub-micron particulates. Although the study was begun to consider the problem
of respirable dust control in the underground coal mine environment, this
enhancement technique shows promise for a more general class of applications.
In the experiments to be reported on, particulates are injected into a
continuous flow thermal diffusion chamber and subjected to varying
environmental conditions; saturation ratio (rel. humidity), temperature,
and residence time were the parameters. The results show that condensation
on a dust nucleus (~0.1 micron) can bring it to a size of several microns
in a sufficiently short period of time as to indicate engineering
applicability of the concept.
* This project was supported by Grant No. ROI OH 00565-02 awarded by the
National Institute of Occupational Safety and Health.
439
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INTRODUCTION
An experimental study of the properties of respirable particulates was
undertaken at the Colorado School of Mines in 1975. The object of this
study was to consider methods by which the control of respirable coal dust
in underground mines might be enhanced. It was felt that particle enlarge-
ment via condensation of water vapor showed substantial promise for dust
control enhancement and therefore, knowledge of the interaction of water
vapor with respirable particulates became very important. The experimental
results of our studies are the substance of this report.
A number of industrial situations exist wherein conventional dust
scrubber technology cannot be applied because of constraints imposed by
space, noise, safety or other environmental considerations. Such a situation
exists in underground coal mines. Space in the working area near the face
is very limited and therefore dust collecting devices must operate in close
proximity to the miners. Hence noise is a factor. In addition, the mine
atmosphere often contains a mixture of methane with the coal dust which
could build up to explosive concentrations inside a collector. This, plus
the overall mine ventillation requirements place lower limits on the air
flow through the collector.
Inertial separators, besides losing efficiency for sizes below about
five microns, generate considerable noise when operating at the volumes
required in a typical coal mine operation. Nevertheless, inertial separators
such as cyclones and wet Venturis, some combined with filtering mechanisms,
are being tried as machine-mounted scrubbers. Baghouses, which can be quite
efficient even at smaller sizes, have not been employed to any great extent
because of the ventillation requirements and heavy dust loading. Electro-
static precipitators, also efficient at small sizes, require operating
voltages of 30 - 100 kV; the frequent arcing which occurs at these voltages
presents a serious explosion hazard in the presence of methane and coal dust.
The mainstay for dust control in almost all mining operations has been the
water spray.
Water sprays and most inertial separators, unless operated at very high
energy input levels, fall to near zero efficiency for particles smaller than
five microns. It is only those particles smaller than five microns that
have an appreciable probability of penetrating the upper respiratory track
and only those smaller than about one micron are likely to be trapped in the
alveoli. Those particles that are deposited in the alveoli give rise to coal
workers' pneumoconiosis (black lung), silicosis, and other debilitating and
sometimes fatal respiratory pathologies. Consequently, one concludes that
although sprays and inertial devices reduce the mass loading of the air,
little is done to reduce the respirable dust content. Since the efficiency of
such commonly used control methods is strongly size-dependent (falling
rapidly with decreasing particle size) particle enlargement if practical
can improve their efficiency and decrease the health hazard of dusty
environments.
Figure 1 hints at the difficulty in using water sprays for the control
of small particulates; in order to attain optimal efficiency, the spray
440
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droplet must be comparable in size to the dust particles, with the maximum
occurring for a size ratio of about 0.7 (taken from Hocking, (1950)*).
Figure 2 shows the same results but for larger droplets (taken from
Mason, (1971)2).
ol 0 4 ii'5 i>5
-------
vapor through air. In the work being presented, we have not included an
analysis of the initiation of water droplet growth, i.e., heterogeneous
nucleation, on the dust particles. However, a study of this phenomenon and
the associated time lab is currently underway in our laboratory.
The classical model for droplet growth by condensation is due to
Maxwell (of electricity and magnetism fame) and may be taken as a zeroth-
order description for use here and is presented in detail by Mason (1971)2.
A widely accepted generalization of the classical model has been presented
by Fukuta and Walter (1970)11 and should be used for a detailed description
of condensational growth, particularly for small droplets (several microns).
In the Maxwell model it is assumed that the diffusion and temperature fields
are in a steady state, that the water vapor may be described as an ideal
gas, and that the Clausius-Clapeyron equation gives the temperature depen-
dence of the vapor pressure during the phase change. The resulting non-
linear differential equation may be integrated to yield
(1) r2(t) = ro2 + 2 (S - l)t/(a + b)
a « diffusion related constant
b = conduction related constant
S = saturation ratio
In the Fukuta-Walter model the form of the equation remains the same but
the constants a and b assume some implicit size dependence. When important,
curvature and solution effects may be readily included into the model.
For the sake of the experiments under discussion, the second term in
equation (1) is the important one, in particular the factor of (S - 1). When
the ambient conditions are saturated, S = 1 and no growth occurs. If the
relative humidity were 102%, then S = 1.02 and the growth rate would be
twice as large as when the relative humidity is 101%. The growth rate
could be increased dramatically by simply increasing the relative humidity
from 101% to 110%. The conclusion is that the rate of growth is a very
sensitive function of the relative humidity or saturation ratio. The results
to be presented below correspond to values of S equal to 1.006, 1.011, 1.020,
and 1.032.
EXPERIMENT
The experimental measurements of the condensational enlargement of sub-
micron particulates were made in a continuous flow parallel plate thermal
diffusion chamber. This chamber is a modification of one used earlier by
Saxena and Fukuta (1974)12 and has recently been thoroughly analyzed (Brown
and Schowengerdt 197913). Figure 3 is a schematic drawing of the diffusion
cell of the chamber. Because of the continuous range of supersaturations
produced, it is often referred to as a spectrometer. It consists of three
functionally different parts; the inlet chamber, the thermal circuit or
diffusion cell, and the exhaust chamber.
442
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Injection
Edge
Figure J. Continuous flow parallel plate thermal diffusion cell.
The inlet chamber (not shown) is designed to provide for the uniform
distribution of pre-humidified carrier air over the cross section of the
diffusion cell and for the alignment of the injection probe through which
the dust nuclei and respirable particulates are injected into the air
stream. It also insulates the inlet end of the diffusion cell from heat
exchanges with the surroundings.
Similarly, the exhaust section is designed to maintain laminar flow
at the exit, align the sampling probe of the particle counter with the
inlet probe, and to insulate the outlet end of the cell from heat exchanges.
The supersaturated conditions (S > 1, see equation 1) necessary for
droplet growth are produced in the diffusion cell. The heated and cooled
junctions (Figure 3), which are separated by an insulating strip, establish
an approximately linear thermal gradient along the metal top and bottom
plates (the side opposite the heated and cooled junction is also conducting).
This "folded" thermal circuit results in a continuously variable thermal
gradient across the gap between the plates. The gradient has a maximum at
the heated and cooled side and a minimum near zero on the opposite side next to
the conducting junction. The inside surfaces of the chamber are covered with
canvas which is periodically wetted in order to maintain saturated boundary
conditions.
Figure 4 (taken from Wieland 195611*) shows qualitatively how the meta-
stable supersaturated state results. At any point within the chamber the
vapor density and temperature vary, to a high degree of accuracy, linearly
from top to bottom. The Clausius-Clapeyron equation gives the saturated
vapor pressure as a function of the temperature. The top line shows the
actual vapor pressure which results from the diffusion of the water vapor
from the saturated hot side to the saturated cold side where the vapor
density is lower. As can be seen, the vapor pressure that results from
the diffusion process is greater than the saturated pressure at the same
443
-------
temperature. The ratio of the two is the saturation ratio:
(2) S = P(T)/Psatm
The internal temperature of the chamber is monitored with an array of 60
thermocouples which extend through the metal plates and into the wetted
canvas.
>20
Temperature (°C)
Figure 4. The saturation vapor pressure (solid
line) and the vapor pressure resulting from the
diffusion of water vapor between the plates of
diffusion cell of Figure 3 (dashed line).
(from Wieland (ISSe)").
In making the droplet growth measurements the test aerosol is injected
into the carrier air stream through the movable inlet probe shown in Figure
3. Care must be taken to insure that all transients in the velocity,
temperature, and diffusion fields have disappeared before the injection of
the test particulates (see Drown and Schowengerdt 197913). The velocity of
the carrier air is determined through a series of time-of-flight measure-
ments. The residence time in the supersaturated environment is then inferred
from the position of the inlet probe and the carrier air velocity.
An optical particle counter (Model 208, Climet Instruments, Redlands,
California) is used to count and size the grown droplets. The output pulses
from the white-light detector are processed and stored in a multichannel
analyzer and droplet size data are obtained from it in the form of peak
locations (modal diameters). The input particulates are the naturally
occurring ones in the lab air and their size is monitored with a laser
particle counter (ASAS, Particle Measuring Systems, Boulder, Colorado).
Latex spheres are used for calibration.
RESULTS AND CONCLUSIONS
The results of the experiments are shown in Figures 5-8. The dashed
curve is obtained from the application of the Fukuta-Walter model to con-
ditions of the diffusion cell (see Brown and Schowengerdt 197915 for details
of the analysis). The agreement with theory is good. There is, however,
a clear indication of a small systematic disagreement in the slopes of the
curves.
444
-------
Residence Time (Sec.)
Figure 5. Experimental and theoretical growth curve for
water droplets. S • 1.006, T » 28.3°c.
Residence Time (Sec)
Figure 6. Experimental and theoretical growth curve for
water droplets. 5 « 1.011, T * 27.5°C.
Residence Tine (Sec.)
Figure 7. Experimental and theoretical growth curves
for water droplets. S • 1.020, T ¦ 27.5°C.
I 2
Residence Time (Sec.)
Figure 8. Experimental and theoretical growth curves
for water droplets. S • 1.032, T ¦ 28.2°C.
445
-------
The conclusions to be drawn from these growth curves are as follows:
1. even for modest saturation ratios as in Figure 8, particles
of initial size less than 0.2 microns can be enlarged to
approximately 7 microns in about 2 seconds.
2. the growth rate is indeed a sensitive function of the
saturation ratio and there is good reason to expect
much more rapid growth to even larger sizes for increased
saturation ratios.
The possibility of increasing the size of particles as small as 0.1
microns, which are far to small for most control devices to collect, to as
large as 10 microns, offers the potential for enhancing the efficiency of
several existing control methods. Since the particles in the 0.1 - 1.0
micron range constitute the major source of the degradation of meterologic
visibility (Willeke and Brockman, 197716) and are also largely responsible
for respiratory disorders due to inhalable particulates, their efficient
control is of profound importance.
REFERENCES
1. Hocking, L.M., Quart. J. Roy. Meteorol. Soc. 85, 44 (1950).
2. Mason, B.J., The Physics of Clouds, Clarendon Press, Oxford, 1971.
3. Fanhoe, R., Lindroos, A.E., and Ahelson, R.J., Ind. Eng. Chem. 43,
1336, (1951).
4. Lapple, C.E. and Kamack, H.J., Chem. Eng. Progr. 51_, 110, (1955).
5. Semrau, K.T., Marymowski, C.W., Lunde, K.E., and Lapple, C.E.,
Ind. Eng. Chem. 50, 1615, (1958).
6. Rozen, A.M. and Kostin, V.M., Intl. Chem. Eng. 7, 464, (1967).
7. Fuchs, N. and Kirsch, A., Chem. Eng. Sci. 20^, 181, (1965).
8. Truitt, J. and Davis, R.J., Am. Ind. Hyg. Assoc. J., p. 583, Sept. 1971.
9. Semrau, K.T. and Witham, C.L., APCA Proceedings, 68th Annual Meeting
Paper #75-30.1.
10. Yoshida, T., Kousada, Y. and Okuyama, K., Ind. and Eng. Chem. Fund.
15, 37, (1976).
11. Fukuta, N. and Walter, N.A., J. Atmos. Sci. 27, 1160, (1970).
12. Saxena, V.K. and Fukuta, N., Denver Research Institute Report No. 2657,
University of Denver (December 1974).
13. Brown, J.T. and Schowengerdt, F.D., J. Aerosol Sci. 10, 339, (1979a).
446
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14. Wieland, W., Z. Angew. Math. Phys. 7_. 428, (1956).
15. Brown, J.T. and Schowengerdt, F.D., J. Aerosol Sci., October 1979
(in press).
16. Willeke, K. and Brockman, J.E., Atmos. Environ. 11, 995, (1977).
447
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WELDING FUME AND HEAT RECOVERY
THE PROBLEM, THE SOLUTION, THE BENEFITS
By:
Richard C. Larson
Technical Sales Manager
Torit Division, Donaldson Company, Inc.
St. Paul, Minnesota 55164
ABSTRACT
Welding is an important and constantly evolving technology enjoying
a higher growth rate in the metalworking industries than the general indus-
trial growth. Welding as with most industrial processes does have its
inherent problems. One of the problems is the fumes and gases generated
during the welding process. This paper presents a complete technical
analysis of the problems associated with most of the more common types
of industrial welding. In addition, alternative engineered solutions are
investigated in detail which includes a descriptive analysis of several
types of air cleaning devices. Finally, the control of the welding fume
problem is analyzed based on actual test installations demonstrating the
ability to achieve acceptable particulate and gas concentrations levels
in the worker's environment, while achieving significant heat recovery
(or conditioned air) by a unique approach to air recirculation.
448
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WELDING - THE PROBLEM
The use of welding in joining metals has had a phenomenal use since
about 1930; this growth has been faster than the general industrial growth.
Most common everyday items are, in one way or another, dependent upon welding
for their economical construction. Automobiles, electronic equipment, house-
hold appliances, kitchen utensils, bridges, ships, aircraft, buildings, farm
machinery - all are made more economically in today's technology by using
one or more methods for joining by welding. Welding continues to be a grow-
ing industry all over the world. In the last 28 years, the usage of welding
has increased 900%.
Welding, as with most industrial processes, does have its inherent
problems. One of those problems is the fume as gases generated during the
welding process. The fumes and gases generated present real and distinct
hazards not only to operational equipment, but especially to the health
and well being of the personnel who undertake such tasks.
Historically, the approach to the fume and gas problem has been either
to ignore it, or a mediocre attempt to ventilate this contaminated air by
either general ventilating powered roof exhausters or localized fume
exhausters in which in either case the air is exhausted outside. It is
now apparent to most production companies, which utilize welding, that a good
concerted effort must be made to handle the welding fume problem. This has,
in large been brought about by the Occupational Safety and Health Act (OSHA)
of 1970 and perhaps even to a greater degree by the energy conservation
consciousness of most industrial plants. The impact of OSHA and rising fuel
costs behooves companies which use welding processes to tackle the problem
of welding fume more directly, to analyze the alternative solutions, and to
justify the selected approach by any realized tangible and/or intangible
benefits that may be gained.
ANALYZE THE PROBLEM
In the process of analyzing the welding fume problem, it is important
not to lose sight of the final objectives, i.e. to control the fume to the
point that it is not an industrial hazard while keeping constantly alerted to
energy conservation.
Obviously, this will require a given air volume and an air moving device
to remove this contaminated air from the working area and some sort of air
cleaning device to hopefully clean the air to an acceptable level such that
it would not have to be exhausted outside.
This problem analysis would include the type of welding, the metal being
welded, the dimensions and layout of working areas, the number of welding
stations or welders or both (gas and electric arc), the rate of welding, the
presence or absence of air curtains or cross-drafts, the OSHA requirements,
and the aspect of energy conservation.
449
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The most common types of welding would include:
1.
gas welding
9.
thermal spraying
2.
torch brazing and soldering
10.
ultrasonic welding
3.
shielded metal arc
11.
laser welding
4.
gas shielded arc
12.
diffusion welding
5.
plasma arc
13.
explosive welding
6.
submerged arc
14.
resistance welding
7.
flux cored arc
15.
friction welding
8.
oxygen and arc cutting
Analyzing the fume problem in light of the impact of OSHA, has only
recently become rather obvious. There are, as briefly discussed earlier,
many potential fume and gas hazards produced in most welding processes.
A more complete list is tabulated in the Appendix (See Figure 1). OSHA
has set Threshold Limit Values (TLV's) for all of these materials, and
no worker is allowed to be exposed to concentrations above these levels
(See Figure 2). These Threshold Limits are the maximum atmospheric con-
centrations, in milligrams per cubic meter of air, which is felt safe
for the normal individual to breathe for eight hours a day, five days
a week. Evaluation of a process involves all of the factors previously
discussed and may include an air sampling program.
If there are existing air moving devices; fans, fume exhauster, etc.,
the problem should further be analyzed as to the amount of air and corre-
spondingly the amount of heat which is presently exhausted outside.
Having analyzed all of the necessary aspects of the welding fume
problem, the alternative solutions are investigated.
WELDING - THE SOLUTION
Ventilation for welding operations are dictated by all of the factors
considered thus far; however, other factors such as weather, the amount of
heat generated and the presence of volatile solvents also contribute to the
need for a ventilation system. Welding screens, which shield nearby workers
and onlookers from the ultraviolet light from the welding arc should not
restrict ventilation. They should be mounted about 2? above the floor.
Depending on the welding process and the potential fumes and gases
generated, the ventilation system could be one of three types: general
ventilation; source ventilation; or a combination of source and general
ventilation.
Ventilation for general welding and cutting, meaning for metals not
specifically covered later is required when welding under any of the
following conditions.
. In A space of less than 10,000 cubic feet per welder.
. In a room with a ceiling height less than 16'.
. In confined space or where the area contains
partitions or other structures that significantly
obstruct cross-ventilation.
4S0
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SOURCE VENTILATION
Source ventilation, as we shall define it, consists of localized hoods
interconnected with duct work hopefully with a high efficiency air cleaning
device and a fan.
One of the more common type hoods used in welding fume control is the
freely suspended hood. This type of hood is very similar to the articulating
fume exhauster hood. The air volume required to control the fume is depen-
dent on the distance ' x', which is the distance from the face of the hood to
the point where the arc is struck. It is important to remember that it is
not the volumetric air flow that you have in your hood that controls the
fume. The control is achieved by the capture velocity at the point of
fume generation. The capture velocity for welding fume is 100 feet per
minute. If a variation of a source ventilation hood is used, remember
you must achieve 100 feet per minute capture velocity. The American Welding
Society has studied capture velocities for many types of welding under many
types of conditions and has found that the required capture velocity can
vary from 80 feet per minute to 200 feet per minute. (See Figure 3)
Another type of source collection hood that can be used in welding
fume control is the cross-draft table. This type of hood is very adaptable
to fixed position or fixture welding. The air volume is dependent on the
length of the bench with a maximum bench width of 24 inches. Variations of
this hood are frequently used. (See Figure 4)
Another type of hood that is sometimes used in fume control is the
downdraft table. Note that in my opinion this type of hood is not
recommended. Temperatures as high as 6,000 degrees centigrade exist at
the arc zone and the fume will rise away very quickly for the first one
to three feet. The thermal gradients that are created and the swift plume
velocities make if difficult to pull the fume back down into the hood. In
addition, if the weldment is a large piece and it covers most of the table,
you have lost all air flow or capture velocity at the arc zone. (See
Figure 5)
Once the proper source collection hoods have been designed, they are
usually connected to a central duct system which must be designed for proper
air balance. (See Figure 6)
An example of a source collection installation is at the Cornelius
Company, which manufactures Vending Machines. The hoods that were used are
slotted crossdraft hoods at about a 45 degree tilt. The slots are located
all the way around the work piece at the critical weld areas. The complete
absence of fume from the general plant air is easily viewed by looking down
the length of this plant and noting there is no light scattering underneath
the fluorescent light fixtures. If the exhaust system is turned off, the
plant will fill up with fumes in a matter of 10 minutes. This system
consists of the hoods and duct work mentioned earlier as well as a Torit
TD-6120 cartridge collector and a fan which pulls approximately 15,000 cubic
feet per minute of air. The cartridge collector has been tested to be 99.7%
efficient on this welding fume which has an average particle size of 0.6
451
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microns. All of the cleaned air from the TD is recirculated back into the
plant during the heating season. The returned air has a particulate level
of over 2,000 times below the OSHA Threshold Limit Value.
OSHA has established Ventilation Standards for General Welding and Cutting
in Confined Spaces.
FEDERAL OSHA VENTILATION STANDARDS FOR GENERAL WELDING AND CUTTING IN
CONFINED SPACES
. Adequate ventilation is required to prevent accumulation of
toxic materials or oxygen deficiency for welders, helpers
and other nearby personnel.
. All replacement air shall be clean and respirable.
. If it is impossible to provide ventilation, approved
airline masks or hose masks are required.
. In areas that are immediately hazardous to life, approved
hose masks with blower or self-contained breathing apparatus
shall be used.
. If welders and helpers are provided with hose masks, hose
masks with blower or self-contained breathing apparatus,
a worker shall be stationed outside of the confined space
to ensure the safety of those working within.
. Oxygen shall never be used for ventilation.
Operations involving specific toxic metals or other substances require
source ventilation or approved respiratory protection without regard to work-
room size and the other aforementioned conditions that govern whether venti-
lation is needed for general welding. The OSHA Standards that cover these
materials are as follows:
. Operations involving fluorine (as a chemical compound, not a gas)
in confined spaces require adequate ventilation to prevent the
accumulation of toxic material or oxygen deficiency. The same
standards as those for general welding in confined spaces apply.
Outside of confined spaces the need for local exhaust ventilation
or airline respirators depends on contaminant levels although
"experience has shown such protection to be desirable for fixed-
location production welding and for all production welding on
stainless steels."
. Welding or cutting involving zinc in confined spaces must meet
the same requirements as general welding in confined spaces.
Indoors these operations require local exhaust booths or hoods
meeting the standard of 100 feet per minute in the welding zone.
. Welding or cutting involving lead-base metals in confined spaces
requires local exhaust ventilation. When these operations are
conducted indoors, local ventilation or airline respirators are
required. Outdoors approved respirators are needed. In all
cases workers in the immediate vicinity of the operation shall
be protected as necessary by local exhaust ventilation or air-
line respirators.
452
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. Welding or cutting involving beryllium-containing base or filler
metals whether conducted indoors, outdoors or in confined spaces
requires local exhaust ventilation and airline respirators unless
air sampling under the most adverse conditions shows that beryllium
exposures are below acceptable OSHA limits. As with lead, nearby
workers require protection by ventilation or airline respirators
if needed.
. Welding or cutting involving cadmium-bearing or cadmium-coated base
metals indoors or in confined spaces requires local exhaust venti-
lation or airline respirators unless air sampling under the most
adverse conditions shows that worker exposures are within acceptable
limits. Outdoors such operations require approved respiratory pro-
tection such as fume respirators.
. Welding or cutting indoors or in confined spaces involving metals
coated with mercury-bearing materials including paint requires local
exhaust ventilation or airline respirators utiless air sampling under
the most adverse conditions show that workers' exposures are within
acceptable limits. Outdoors such operations require approved
respiratory protection.
. Cutting of stainless steels with oxygen using either a chemical flux
or iron powder, or using a gas-shielded arc, requires mechanical
ventilation adequate to remove the fumes generated.
. Degreasing or other cleaning operations involving chlorinated hydro-
carbons must be located so solvent vapors do not reach any welding
operation. In addition, trichloroethylene and perchloroethylene
should be kept out of atmospheres penetrated by the ultraviolet
radiation of gas-shielded welding operations. This requirement is
to avoid the decomposition of solvent vapors into toxic substances.
It is important to point out that air sampling of the welders breathing
zone in a given plant is necessary to determine if one is in compliance with
the OSHA Standards and/or the degree of ventilation that is required.
GENERAL VENTILATION
Natural ventilation is usually considered adequate to prevent the accumu-
lation of excessive amounts of fumes when welding on mild steel with electrodes
that do not contain fluorides in the covering, provided that room volume is
not less than 10,000 cubic feet per welder, the ceiling height is at least 16'
and cross-ventilation is not blocked by partitions, equipment, etc. In spaces
that do not meet these requirements, general mechanical ventilation is required.
The air volume that is required for general ventilation could be computed
by one or more of the following methods.
General Ventilation Air Volume Requirements
1. One air change every five (5) minutes.
2. 5 feet per minute updraft velocity per square foot of floor space.
3. 2000 cubic feet per minute per welder (OSHA recommendation).
A. American Conference of Governmental Industrial Hygienists
recommendations.
453
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General Ventilation, where local exhaust cannot be used:
Rod, diameter
5/32
3/16
1/4
3/8
CFM/welder* A. For open areas, where welding fume can
1000 rise away from the breathing zone: cubic
1500 OR feet per minute required = 800 X pounds
3500 per hour rod used.
4500 B. For enclosed areas or positions where
fume does not readily escape breathing
zone: cubic feet per minute required =
1600 X pounds per hour rod used.
*For toxic materials higher air flows are necessary and operator
may require respiratory protection equipment.
Each of the above methods must be thoroughly investigated since no one
method is ideal; however, methods 3 and 4 are the most consistent direct
methods since they deal with air volumes based on the amount of particulate
that could be generated, i.e. the number of welders, the size of welding rod
used and the amount of rod used per hour. Methods 1 and 2 are based on air
volumes taking into account only the physical dimensions of the welding area
without regard to the concentration of welders in this welding area.
In principle, general ventilation removes the fume-laden air usually at
the roof level, or at intermediate levels above the welding areas. Since our
objective is to not only control the welding fume, but also to conserve energy,
our general ventilation system will include an air cleaning device and will
return the cleaned respirable air back into the plant atmosphere. A typical
layout for a general ventilation system is depicted in Figure 7. This is the
general ventilation layout we have found to be the most successful. As is
schematically illustrated, the warm fume—laden air is removed at or near the
roof level. Since our objective is to not only control the welding fume but
also to conserve energy, our system includes a high efficiency TD cartridge
collector and we return the cleaned respirable air back into the plant. Note
that the system includes a fan, a by-pass damper and low velocity exhaust
grilles located at the floor level. Also note that the return air duct work
is located up near the roof trusses so as to further take advantage of the
elevated air temperatures by virtue of radiation convection heating. By re-
moving the heated fume-laden air at the ceiling and reintroducing the cleaned
recirculated air at the floor level combined with the thermal gradients that
exist in all plants from the floor to the ceiling, our general ventilation
system becomes a very efficient "push—pull" system. Note the inline centri-
fugal fan and the by-pass damper. This by-pass damper is normally used in
the summer months to discharge the "super-heated" air to atmosphere. This
causes induced natural ventilation creating a cooling effect within the plant.
Additionally, this by-pass damper can be used for emergency by-pass if there
should be an equipment failure.
The strategic locations of the exhaust hoods and TD units are based on
the criteria mentioned earlier as well as the fact that the exhaust hoods and
recirculation grilles must be located in a certain pattern to achieve an
effective push-pull system.
454
-------
One example of a General Ventilation system, based on this design, is
located at the Bowen-McLaughlin-York Company in York, Pennsylvania. They
manufacture military armored vehicles. The system consists of 27 Torit
TD-6120 cartridge collectors. The system filters fume-laden air from the
plant at the rate of 364,000 cubic feet of air per minute. At Bowen there
are over 250 welders and this system collects more than 4,000 pounds of weld-
ing fume per month. This one million dollar air pollution control system will
pay for itself in less than five years. An independent testing laboratory
was unable to measure any particulate in the collector's return air grilles
after 8 hours of sampling.
Occasionally, it may be necessary to design and use a combined general
ventilation and source ventilation system. This may be necessary for several
reasons, including a welder welding in an enclosed area or perhaps due to
localized toxic fume and gas generation. For example a combined system could
consist of our previously discussed general ventilation system and a separate
hood and fan to remove the welding fume from an enclosed area. This fan blows
the removed fume to the ceiling where the general ventilation system takes
over and filters the air through the TD unit (See Figure 8). Another version
of a combined system is to use a separate hood, fan and small TD collector to
control fume generated in perhaps an isolated area within a plant. Note that
this small TD unit also returns the clean air back into the plant working
environement. (See Figure 9).
SELECTION OF AIR CLEANING EQUIPMENT
There are basically three types of air cleaning equipment that could be
used to filter or clean the fume-laden air. This includes an electrostatic
precipitator, or mechanically shaken baghouse and the Torit TD Cartridge
Collector.
The principle of operation for a modular electrostatic precipitator is
that the dust-laden air passes first through an ionizer section which charges
the dust particles either positively or negatively depending on the design.
These charged particles then pass through the collector plate section and are
electrostatically attracted to these plates. Then the clean air (in theory)
passes out of the collector plate section. (See Figure 10).
The mechanical shake baghouse uses filter tubes or bags of either natural
or synthetic filter media. The contaminated air enters the unit through an
inlet and then is filtered through the cloth bag from the inside to the out-
side. The cleaned air then passes out the outlet. Periodically a compartment
is isolated or the unit is shut off and the material is removed from the bags
by mechanically shaking them. (See Figure 11).
The Torit TD Cartridge Collector combines the technology of a continuous
duty pulse jet baghouse with the technology of cartridge filtration. Each
cartridge is 12-3/4" in diameter by 26" long and contains 300 square feet of
filter media. The TD collectors we discussed earlier each contain 32 cart-
ridges or the equivalent filter area of 9,600 square feet per unit. The
contaminated air enters the inlet and the fume is removed from the air stream
by passing through the cartridge from the outside to the inside. This ultra-
455
-------
clean air is then recirculated back to the plant. As monitored on a pressure
drop basis, the cartridges are cleaned automatically, on a "as required"
frequency, by backflushing them with sonic velocity pulses of compressed air.
The collected fume then falls into the storage hopper. (See Figure 12).
Capital expenditures are important, however, it should not be the base
criteria for the selection of air cleaning equipment. A more meaningful
criteria in the long term is to examine all of the advantages and disad-
vantages of each of the individual pieces of equipment, based on the opera-
tional characteristics and on a Life Cycle Analysis.
Electrostatic Precipitator (Modular)
Advantages
Compact
No complex duct work
Low power cost
Continuous duty
Low pressure drop
Disadvantages
High Maintenance
Creates ozone and N0X
Creates space charge
Mechanical Rappers - moving parts
or liquid wash system
Decreased efficiency overtime
Able to handle only light loadings
.004 grains per cubic foot or less
Mechanical Shake Baghouse
Advantages
High efficiency
Low operating cost
Disadvantages
Intermittent duty
Moving mechanical parts
Large space required
Normally requires precoating
Requires external duct work
Torit TD Cartridge Collector
Advantages
High efficiency
Low maintenance
No moving parts
Compact - low headroom
Continuous duty
Low air-to-cloth ratios
Disadvantages
Requires compressed air
Normally require® external duct work
As a final step in the welding fume analysis, once the ventilation system
has been designed and the fume collector and associated peripheral equipment
has been selected, any and all benefits should be examined.
WELDING FUME - THE BENEFITS
The realized benefits resulting from a properly designed and installed
welding fume ventilation system fall into two major categories. That being
the benefits of a cleaner plant and the heat recovered from a properly applied
air cleaning device. The benefits of a cleaner plant would include:
456
-------
1. Fewer occupational hazards - respiratory, eyes, skin.
2. Improved worker efficiency and morale.
3. Longer machine life.
4. Lower plant maintenance.
5. Improved product quality.
Obviously, there will be some intangible benefits of a cleaner plant,
which haven't been listed.
There are several ways of examining the energy conservation due to being
able to recirculate this cleaned ventilation air. The most direct approach
is to assume that the welding area were properly ventilated and that this
contaminated air were simply exhausted to the outside atmosphere. If this
were the case, make-up air would have to be supplied to the plant either by
direct or indirect means. The dollar figures you see here represent the
additional fuel burden per heating season if you exhausted 100,000 cubic feet
per minute of air (See Figure 13). These dollar figures are based on the
average industrial fuel costs in April of 1978. Escalating fuel costs will
change these values dramatically.
The recovered heat from a general ventilation system, which recirculates
the cleaned air, would be considerably higher in dollars per year than these
figures. This is primarily due to the fact that in this unique general venti-
lation approach, the contaminated air is removed at the proximity of the roof
level. Warm air rises and it has been measured that the temperature gradient
from the floor level to the roof level can vary as much as AO degrees Fahr-
enheit to 50 degrees Fahrenheit. The source of this heat can be from many
sources such as the welding process itself, lights, operating equipment,
motors, etc. After this super-heated air has been cleaned, we reintroduce it
at the floor level primarily to facilitate the "push-pull" principle of the
general ventilation system. This returned warm air is further heated due to
the compression of air across the fan and to a greater degree by radiation-
convection since the return air supply passes through the super-heated roof
level air for the width of the welding area. The total heat recovery can be
calculated knowing the air volume, the temperature gradients that exist with-
in a building and the outside temperature. A good example of the considerable
amount of heat that can be recovered is from the installation at Bowen -
McLaughlin, which was discussed earlier, in which the entire plant heat load
is maintained by the ventilation system, and the two newly purchased boilers
were never installed.
Even though the welding process produces fumes and gases, it is possible
to comply with current hygienic standards and recover enormous amounts of
energy whether in the form of heated or cooled air. In order to do this one
has to thoroughly analyze the problem and carefully investigate the alterna-
tive solutions which will enahle compliance and energy conservation.
457
-------
in
00
WELDING FUMES AND GASES
Particulates
Pneumoconioses
Harmful
I
| Fibrotic~l Nonfibrotic]
[ Silica
Asbestos
Copper |
Relatively
Harmless
Carbon
Tin
(Aluminum
Pulmonary
Irritant or
Toxic Inhalants
Cadmium
X
Chromium
Fluorides
'
Lead
Manganese
I
Magnesium
X
Mercury
X
[ Iron 1 f Molybdenum
-j
Nickel
I
Titanium
X
Vanadium I
» —'
Zinc
Gases
Primary Pulmonary
1
Non-Pulmonary
Ozone
Oxides of Nitrogen
Phosgene
Phosphine
Carbon Monoxide
Carbon Dioxide
Figure 1 Welding fumes and gases.
-------
Common materials that
may be encountered
in welding
Threshold
Limit Values
mg/m3 (ACGIH)
Threshold
Limit Values
mg/m3 (OSHA)
Antimony
0.5
0.5
Arsenic
0.5
0.5
Beryllium
0.002
0.002
Carbon dioxide
9,000.0
9,000.0
Carbon monoxide
55.0
55.0
Cadmium oxide fumes
0.05
0.1
Chromates (as CrO^)
0.05
0.1
Chromium (metallic)
0.5
1.0
Cobalt
0. 1
0.1
Copper Fumes
0. 1
0.1
Fluoride
2.5
2.5
Hydrogen fluoride
2.0
2.0
Iron oxide fumes
5.0
10.0
Lead
0. 15
0.2
Magnesium oxide fumes
10.0
15.0
Manganese
5.0
5.0
Mercury
0.05
0.1
Molybdenum
10.0
15.0
Nickel carbonyl
0.35
0.007
Nickel (metallic)
1.0
1.0
Nitrogen dioxide
9.0
9.0
Ozone
0.2
0.2
Phosgene
0.4
0.4
Selenium
0.2
0.2
Silver
0.01
0.01
Tellurium
0. 1
0.1
Titanium dioxide
10.0
15.0
Uranium
0.20
0.25
Vandium fumes
0.05
0.1
Zinc oxide fumes
5.0
5.0
Zirconium
5.0
5.0
Figure 2 Threshold limit values for materials commonly
encountered in welding and cutting.
459
-------
FLEXIBLE DUCT-
S'' FLANGE
WORK
PORTABLE EXHAUST
X, INCHES
PLAIN DUCT
CFM
FLANGE OR CONE
CFM
UP TO 6
335
250
6-9
755
560
9- 12
1335
1000
FACE VELOCITY = 1500 FPM
DUCT VELOCITY = 3000 FPM MINIMUM
ENTRY LOSS = 0.25 DUCT VP
ALSO SEE "GRANITE CUTTING" VS-909
Figure 3 Source collection (freely-suspended hood).
-------
45° SLOPE MIN
SLOTS-SIZE FOR 1000FPM
•P*
lr
BAFFLES ARE DESIRABLE
Q = 350 CFM/LINEAL FT OF HOOD
HOOD LENGTH = REQUIRED WORKING SPACE
BENCH WIDTH = 24" MAXIMUM
DUCT VELOCITY = 1000 - 3000 FPM
ENTRY LOSS = 1.78 SLOT VP +0.25 DUCT VP
I MAXIMUM PLENUM VELOCITY
I 1/2 SLOT VELOCITY
Figure 4 Source collection crossdraft table.
-------
EXHAUST DUCTS
Q = AV
CLEANOUT DOORS
ENCLOSE BASE OF TABLE
SLOTTED BAFFLE - SIZE SLOTS FOR 1000 FPM
Figure 5 Source ventilation downdraft table (not recommended)
-------
TORIT TD-6120
-p*
ON
t
Source collection layout.
-------
Figure 7 General ventilation layout.
-------
Figure 8 Combined general ventilation and source removal layout.
-------
Figure 9 Combined general ventilation and source collector layout.
-------
AIR
DUST
LADEN
COLLECTOR
FLOW
CLEAN
AIR
10,000
VOLTS
POWER
SUPPLY
5,000 VOLTS
110
AC
Figure 10 Electrostatic precipitators (modular).
-------
NORMAL OPERATION SHAKE CYCLE
SHAKER MOTOR
SHAKER
CABINET
CLEAN AIR
—KOUTLET
FILTER
TUBES
CONTAMINATED
AIR INLET =
O O CrO O
;o O O'O O
ool
DUST
HOPPER
SLIDE GATE
Figure 11 Mechanical shake baghouses.
-------
NORMAL OPERATION
ELEMENT PURGE
.p*
a>
to
BLOW PIPE
BLOWER OUTLET
(CLEAN AIR)
PLENUM
VENTURI
7
CONTAMINATED
AIR INLET
INDUCED
AIR
AIR
PULSE
INLET DEFLECTOR
DUST
Figure 12 Torit TD cartridge collectors.
-------
NATURAL GAS FUEL OIL (//2) COAL
Atlanta, Georgia
$10,009
$12,201
$ 8,261
Boston, Massachusetts
18,179
22,161
15,005
Chicago, Illinois
21,012
25,615
17,343
Cleveland, Ohio
19,902
24,262
16,427
Dallas, Texas
7,406
9,028
6,113
Denver, Colorado
17,803
21,703
14,695
Detroit, Michigan
20,142
24,544
16,625
Honolulu, Hiwaii
0
0
0
Minneapolis, Minnesota
26,875
32,762
22,182
New York, New York
15,679
19,113
12,941
Nome, Alaska
46,328
56,476
38,239
Pittsburg, Pennsylvania
19,090
23,272
15,757
Richmond, Virginia
13,930
16,981
11,498
St. Louis, Missouri
14,508
17,686
11,795
Sacramento, California
9,194
11,208
7,589
Salt Lake City, Utah
19,350
23,589
15,972
Seattle, Washington
16,769
20,442
13,841
Wausau, Wisconsin
27,768
33,851
22,920
Figure 13 Benefits of energy conservation.
The cost of exhausting 100,000 cfm of air at current fuel prices (4/78) per heating season.
470
-------
Addendum:
Recirculation of Filtered Air
In many cases, the economic and energy conservation factors are so
important as to warrant the additional capital expense of safety when
providing for recirculation.
In designing and using a recirculation system, the following items
should be followed.
1. The air cleaning system should be capable of achieving
a high enough efficiency to provide an exit concentration
of at least one-half of the Threshold Limit Value.
2. Provisions should be made for a bypass of the recirculated
air to the outdoors or shutdown of the process if there is
an equipment failure. If a system is intended to conserve
heat in the winter months and if adequate window and door
openings permit sufficient "make-up" air when open, the
system can discharge outdoors in warm weather.
3. It may be necessary to supply fresh makeup air in addition
to the recirculated air to provide continuous dilution of
any gases which may tend to otherwise accumulate.
4. The operation should be tested by a certified industrial
hygienist after installed to insure compliance.
5. Routine testing, maintenance procedures and records should
be maintained.
471
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PARTICULATE REMOVAL CONSIDERATIONS IN SOLVENT
EMISSION CONTROL INSTALLATIONS
A Case Study
By:
Eugene A. Brackbill
Peter A. Kalika
TRC - THE RESEARCH CORPORATION of New England
125 Silas Deane Highway
Wethers field, Connecticut 06109
ABSTRACT
Organic emissions from surface coating processes are receiving
increased regulatory attention. One of the commonly considered control
techniques is activated carbon adsorption. This is very favorable when
the emitted solvents have a cost saving reuse potential. However, the
presence of condensed, submicron organic particulates and their impact
upon the carbon is often overlooked.
This paper examines the case of a paper coating process utilizing
a phenol-formaldehyde resin system. A pilot study revealed that the
condensed organic particulate would greatly reduce the carbon's operating
life. Two high efficiency particulate control devices were evaluated to
determine their economic impact on the total system cost. The costs
associated with both devices effectively eliminated the solvent recovery
cost savings.
472
-------
PARTICULATE REMOVAL CONSIDERATIONS IN SOLVENT EMISSION CONTROL
INSTALLATIONS - A Case Study
INTRODUCTION
Recently, EPA has focused attention on the volatile organic compound
(VOC) emissions from surface coating processes. Activated carbon adsorp-
tion has been defined as one of the reasonably available control technol-
ogies (RACT) for reduction of VOC emissions from many of these processes.
Carbon adsorption is particularly attractive since it allows recovery of
the collected solvent for reuse. The resulting process operation cost
savings can offset the control system cost over a period of a few years.
However, the solvent recovery potential often overshadows consideration of
operational problems which severely limit the applicability of carbon
adsorption.
The presence of condensible organic compounds in the process exhaust
presents one of the most difficult problems. The condensation aerosols
formed by these compounds are typically submicron and may form a polymeric
coating which occludes the carbon pores. Such a coating usually cannot be
removed by conventional on-line regeneration methods.
This paper examines the particulate problem as characteristic to
specialty paper processes employing phenol-formaldehyde resins as binders
and surface coatings.
PROCESS DESCRIPTION
The manufacture of some specialty papers includes the saturation or
coating of the web with a resin system formulated to provide working
strength to the paper. The basic unit operations comprising the process
include web formation, wet saturation, and resin thermosetting.
The thermosetting process is typically accomplished in a recirculated
flow, convection oven. The emissions from the oven are primarily alcohols
and/or aromatic solvents. Additionally, small amounts of the resin
components and partially polymerized materials may be emitted. It is
these materials which exist in a transitional state in the exhaust stream
and condense when mixed with ambient air, being evidenced by a visible
plume.
473
-------
IMPACT OF PARTICULATE ON ACTIVATED CARBON PERFORMANCE - PILOT STUDY
A 14 acmm (500 acfm) pilot scale carbon adsorption system was connected
to a split stream from the curing oven exhaust. The purpose of the pilot
study was to determine the activated carbon performance parameters relative
to the carbon operating life. It had been predetermined that the solvent
collection efficiency would be greater than 90 percent.
Although the initial measured performance was excellent, 99 percent
solvent removal, the phenolic materials caused the carbon to plug and
rapidly deteriorate in its ability to adsorb and retain the solvents.
Carbon samples were extracted from the bed and laboratory tests performed
to evaluate three important physical characteristics: activity, retentivity,
and volatile matter content. Activity refers to the carbon's ability to
adsorb organic vapors from a solvent laden airstream. After two weeks of
operation the activity had decreased 39 percent and after one month the
decrease was 52 percent. Retentivity, the ability to hold organic vapor
after it has been adsorbed, likewise decreased by 51 percent and 84 percent
following two weeks and one month on stream, respectively. The carbon
volatile matter content after steam desorption showed an increase from
the normal value of approximately 3 percent by weight to 11.5 percent after
one month of operation. The deterioration trends of these parameters are
depicted in Figure 1.
The rapid deterioration rate had not been anticipated even thcugh the
presence of some particulate material in the oven exhaust was expected.
Additionally, the pilot unit was preceded by a foam type filter which had
been thought adequate for the trace particulate removal. In an effort to
more accurately describe the problem, measurements for both filterable and
condensible particulate material were performed downstream of the foam type
prefilter using the EPA Method 5 sampling train. The sample train filter
was maintained at approximately 110°C (230°F). The condensible fraction
was collected in the impinger train with the outlet temperature from the
last impinger maintained at approximately 21°C (70°F). Additionally,
particle size determinations of the filter catch were performed by con-
ventional optical microscopy techniques.
The results revealed an average total particulate loading of 16 g/hr
(0.036 lb/hr). Although seemingly small, its impact becomes more apparent
when viewed with respect to particle size. The particle size data,
Table 1, indicated that the particulate were predominantly submicron.
Therefore, it was realized that particle deposition would occur throughout
the bed due to diffusional mechanisms. The volatile matter test had
indicated this possibility since high residual levels were found in the
carbon samples regardless of the location from which they had been extracted.
As a result of the pilot study, it was concluded that activated carbon
adsorption was not a viable control technique unless the submicron
particulate could be eliminated.
474
-------
Figure 1. Activated carbon performance deterioration.
LEGEND
- ACTIVITY
- RETENTIVITY
- CARBON VOLATILE CONTENT
z
UJ
o
CSL
UJ
a.
WEEKS
475
-------
TABLE 1
PARTICLE SIZE DISTRIBUTION
Size Range, ym
Percent by Number
>3
3.3
1-3
31.2
<1
65.5
PARTICULATE CONTROL TECHNOLOGIES
Several control devices were considered but the particulate's submicron
size and "tarry" characteristics eliminated the majority of obvious
filtration and wet scrubbing choices. Electrostatic precipitation, both
conventional and low voltage two-stage, was rejected because of the fire
hazard potential and anticipated collected material removal problems.
Two control devices were eventually selected for economic evaluation.
These were the Andersen HEAF (High Efficiency Air Filter), formerly marketed
by Johns-Manville, and the Ceilcote IWS (Ionizing Wet Scrubber).
Andersen HEAF
This equipment has successfully been applied for control of phenol-
formaldehyde emissions in several industries, most notably fiberglass
impregnation. The unit employs a fiberglass medium for particle collection.
The medium is a 25.4 mm (1 in.) thick mat which closely resembles home
insulation in appearance. The contaminated gas stream is drawn through
the mat at superficial velocities ranging from 400 to 550 meters/min
(1,300 - 1,800 ft/min). Under these conditions the mat compresses to about
3.2 mm (0.125 in.) thick, increasing the volumetric fiber density. The
particulates are removed by impaction and, to a lesser extent, diffusional
mechanisms. Low viscosity liquid droplets agglomerate on the mat to form
a nearly continuous liquid phase which migrates through the mat and is
sheared by the airflow exiting the mat. The resulting larger droplets
are removed by sedimentation, centrifugal fan separation, and mesh pad
mist elimination. The more viscous compounds collected eventually blind
the medium since they are retained rather than drained. When the airflow
resistance reaches a predetermined level, the filter automatically advances
to expose clean medium to the gas stream.
476
-------
Ceilcote IWS
The ionizing wet scrubber combines the principles of electrostatic
particle charging, image force attraction, inertial impaction, and gas
absorption to collect submicron solid and liquid particles and gaseous
contaminants simultaneously. A high voltage ionization section applies
an electrostatic charge to the particles prior to their entering a con-
ventional packed bed wet scrubber section where they are removed either
by inertial impaction or attraction of the charged particles to the
electrically neutral packing surface. The collected particles are flushed
from the bed by the scrubbing liquid which, with the addition of chemical
agents, also functions to absorb gaseous contaminants.
ECONOMIC EVALUATION
Cost estimates for the two control systems were prepared in terms of
capital cost to purchase and install, and annualized cost. An estimate
was also prepared for an activated carbon system. The process and economic
parameters are presented in Table 2.
The overall accuracy of the estimates is expected to be in the range
of ± 15 percent. In preparing the capital estimates, the equipment costs
were based on manufacturer quotations. The installation, engineering, and
contingency figures were calculated as a percentage of the equipment costs.
Available space in and around the plant was limited. This was a major
factor affecting the installation cost.
The annualized cost data were prepared employing standard procedures.
The operating expenses included were plant overhead, maintenance, and
utilities. Plant overhead is the cost associated with space utilization,
administrative charges, insurance, and taxes. The equipment maintenance
costs were difficult to assign since there are many design and operating
factors which influence them. The estimated costs represent a best
judgment but are likely generous. The electrical costs are essentially
related to the fan motors. Additional costs characteristic to the devices
are included and explained in the individual estimates which follow.
Activated Carbon
Table 3 describes the costs for the activated carbon control system.
The operating expense includes the cost for a total bed replacement once
each year. The carbon cost is based on 32 kg/acm (2 lb/acf) exhaust flow
and $2.20/kg ($l/lb) carbon.
The solvent recovery reuse credit was calculated assuming that the
adsorption efficiency would be 95 percent, desorption efficiency 90 percent,
and the distillation efficiency 85 percent. Therefore, the overall recovery
477
-------
TABLE 2
PROCESS AND ECONOMIC ESTIMATION BASES
Process
Nominal exhaust volume
Exhaust temperature
Moisture content
Contaminant
Concentration
Estimated solvent evaporation rate
1840 acmm (65,000 acfm)
93°C (200°F)
5%, volume
50% isopropanol, 50% methanol
1010 ppm
2721/hr (72 gal/hr)
Economic
Depreciation
Interest rate
Electrical
Steam
Water
Recovered solvent value
Operation
12 year, straight line
10 percent
$0.035/kw-hr
$3.30/1000 kg ($1.50/1000 lb)
$0,004/1000 1 ($0.15/1000 gal)
$0.31/1 ($1.18/gal)
6,240 hrs/yr
478
-------
TABLE 3
ACTIVATED CARBON SYSTEM COST SUMMARY
(1840 acmm - 65,000 acfm)
CAPITAL COSTS
Control device $350,000
Ductwork 54,000
Fan 8,000
Evaporative cooler 30,000
Installation 9 8,000
Engineering 52,000
Contingency 58,000
TOTAL $650,000
ANNUALIZED COSTS
Depreciation $54,000
Interest 33,000
Operating
Overhead 6,000
Maintenance 11,000
Activated carbon 32,000
Steam 74,000
Water 26,000
Electrical 23,000
TOTAL $259,000
SOLVENT RECOVERY $383,000
NET ANNUAL COST $124,000 credit
479
-------
would be 73 percent. Applying this to the estimated solvent evaporation
of 272 liters per hour (72 gal/hr), there would be a 198 liter per hour
(52 gal/hr) recovery for reuse. At $0.31/liter recovery value, this
equates to a savings of $383,000 per year.
The estimates reveal the carbon adsorption cost benefit. The
difference between the annualized cost to own and operate and the recovered
solvent value yields a cost savings of $123,950 per year. At this rate,
the system capital cost would be recovered within 5.25 years (simple
payback).
Particulate Control Equipment
The estimates for the HEAF and IWS, Table 4, were prepared in an
identical manner as the carbon system. The IWS capital cost was based
on two units in series. The filter cost associated with the HEAF operating
cost was calculated from Andersen data on operating life and the price of
replacement media.
Total System Estimates
The particulate control-activated carbon adsorption systems estimates
are presented in Table 5. Obviously, the combined systems' capital costs
became a significant concern. Of equal importance, the particulate control
equipment contribution to the total annualized costs effectively eliminated
the solvent recovery cost savings. Still, the annualized costs for both
combination systems remained very reasonable, especially when compared with
those had the solvent recovery credit not been available. In light of the
increasing solvent cost and fluctuating supply situation, either control
system remained economically attractive.
The combined systems approach presented another area of concern. The
complexity of the control devices is such that maintenance and operation
would become a labor intensive effort requiring specially trained personnel.
That is, it was not anticipated that a simple planned maintenance schedule
for the control system could be incorporated into the existing plant
program.
In this case, it was decided to convert the process to a water-based
resin system. Although process equipment changes would be required and
some products eliminated, the overall economic impact on the plant would be
less than that of controlling the solvent based process emissions.
480
-------
TABLE 4
HEAF AND IWS COST SUMMARIES
CAPITAL COSTS
Control device
Ductwork
Fan
Pumps, piping, tanks
Installation
Engineering
Contingency
TOTAL
Andersen HEAF
$110,000
10,000
included
1,000
100,000
25,000
24,000
$270,000
Ceilcote IWS
$640,000
10,000
6,000
7,000
225,000
25,000
25,000
$938,000
ANNUALIZED COSTS
Depreciation
Interest
Operating
Overhead
Maintenance
Filter cost
Electrical
TOTAL
$22,000
14,000
3,000
2,000
25,000
98,000
$164,000
$78,000
47,000
5,000
7,000
37,000
$174,000
481
-------
TABLE 5
TOTAL SYSTEMS COST SUMMARIES
CAPITAL COSTS
Activated carbon
Particulate control
TOTAL
Andersen HEAF
$650,000
270,000
$920,000
Ceilcote IWS
$650,000
938.000
$1,588,000
ANNUALIZED COSTS
Activated carbon
Particulate control
TOTAL
$259,000
164,000
$423,000
$259,000
174,000
$433,000
SOLVENT RECOVERY
CREDIT
NET ANNUALIZED
COST
$383,000
$40,000
$383,000
$50,000
482
-------
CONCLUSIONS
The particulate emissions associated with some surface coating
processes must be considered if activated carbon adsorption is selected to
control organic solvent emissions. The capital and annualized costs
attributable to particulate control equipment capable of high efficiency
submicron particle removal can be a significant portion of the total
system cost and offset the potential cost benefit of the solvent recovered
for reuse by the carbon system.
483
-------
ARSENIC EMISSIONS AND CONTROL TECHNOLOGY
GOLD ROASTING OPERATIONS
John O. Burckle
Industrial Environmental Research Laboratory
Office of Research and Development
U.S. Environmental Protection Agency
Cincinnati, Ohio 45268
G. H. Marchant
Southern Research Institute
2000 Ninth Avenue South
Birmingham, Alabama 35205
Richard L. Meek
Southern Research Institute
2000 Ninth Avenue South
Birmingham, Alabama 35205
ABSTRACT
Since arsenic is a significant compound of ores processed in the non-
ferrous metals industries, it is of interest to characterize and evaluate con-
trol strategies which have demonstrated the potential for lowering the emission
rates of arsenic and other hazardous effluents from smelter operations. The
Campbell Red Lake Mines Gold Smelter at Balmerton, Ontario, Canada, has devel-
oped and implemented a successful control strategy for arsenic emissions from
a nonferrous smelting operation. The control system was designed and installed
by Hatch Associates, Toronto, Canada. The Red Lake smelter uses cyclones and
a hot electrostatic precipitator to recover metal values from roaster dusts
with subsequent air quenching to condense (or desublime) arsenic trioxide
which is recovered in a low-temperature baghouse. This paper is a review of
a test program conducted at Red Lake designed to characterize the control
systems and to evaluate the potential for transferring the technology to
smelting operations in the United States.
484
-------
ARSENIC EMISSIONS AND CONTROL TECHNOLOGY
GOLD ROASTING OPERATIONS
INTRODUCTION
The Campbell Red Lake Mines operation at Balmerton, Ontario, Canada has
developed and implemented a successful control strategy for arsenic emissions
from a nonferrous smelter. In the process, a gold-bearing arsenopyrite ore is
roasted to eliminate arsenic and sulfur, the hot roaster calcines are recov-
ered in cyclones and an electrostatic precipitator, and the hot gases from
the precipitator are quenched with air to condense (or desublime) arsenic
trioxide. The finely divided arsenic trioxide is then collected in cold bag
filters. The overall collection efficiency of the system exceeds 99.9% for
arsenic trioxide.
Since arsenic is a significant component of many ores processed in this
country by the nonferrous metals industry, the characterization and evaluation
of control strategies which have demonstrated capabilities for reducing the
emission rates of arsenic and other hazardous emissions from smelters is of
interest. The test program at Red Lake was conducted to evaluate the per-
formance of their "hot ESP/cold bag" system and to assess the potential for
transferring the technology to nonferrous smelting operations in the United
States.
PROCESS DESCRIPTION
A simplified flowsheet of the overall operation of the Campbell Red Lake
Mines Limited installation at Balmerton, Ontario, Canada, is shown in Figure 1.
Ore from an underground mine is subjected to crushing, grinding, ball milling,
flotation and tabling to obtain three cuts for further processing. The first
cut (tabling overflow) is amalgamated directly for recovery of gold, the
second (flotation tails and underflow) is relatively low in arsenic and sulfur
and can be treated by cyanidation to liberate the gold, but the third cut
(flotation and tabling overflow), which is high in arsenic and sulfur,
requires roasting to eliminate arsenic, antimony, and sulfur before the gold
can be economically extracted. Two-stage fluid-bed roasting effectively
removes the arsenic, antimony and sulfur from the concentrate. The calcines
from the roasters, cyclones, and hot electrostatic precipitator can then be
treated by cyanidation to recover the gold. The hot gaseous effluent from
the electrostatic precipitator is quenched with cold air to condense AS2O3
which is recovered in a cold baghouse.
A more detailed flow diagram of the roasting, gas cleaning, and arsenic
system at Campbell Red Lake Mines is shown in Figure 2. The concentrate
(nominally 9% As, 18% S, 23% Fe) is fed to the first roaster as an 80% slurry
485
-------
FROM MINE
FLOTATION TAILS
TO CYANIDATION
UNDERFLOW TO
AMALGAMATION
ROASTER
AIR
COOLING
AIR
COOLING.
AIR
Ai203 TO UNDERGROUND
STORAGE
DILUTION
COOLER
COLD
BAGHOUSE
CYCLONES
FLOTATION
COARSE ORE
STORAGE
FINE ORE
STORAGE
CRUSHING
CIRCUIT
GRINDING
CIRCUIT
FLUID BED
ROASTERS
TABLING
HOT
ELECTROSTATIC
PRECIPITATOR
GASES
TO STACK
Figure 1. Simplified flow diagram.
486
-------
CYCLONES
400°C
(750°F)
AIR 7.4 m3/min
ELECTROSTATIC
PRECIPITATOR
(2 IN PARALLEL)
370°C (700°F)
00
-J
CONCENTRATES
2400 Kg/ht
(5400 Ib/hr)
SOLIDS AS
80% SLURRY
IE U„
°
ujO„„
200 Kg/hr (435 Ib/hr)
COOLING AIR
300 m3/m
(10,650 SCFM)
AIR
50 m^/min
(1730 SCFM)
AIR
26 m^/min
(940 SCFM)
CALCINES
TO CYANIDATION
2157 Kg/hr
(4750 Ib/hr)
STACK GAS
395 m^/min
(13.930 SCFM)
93°C
(200°F)
STACK
BURNER
FAN
BAGHOUSE
( 2 BANKS OF 2 )
>
MIXING CHAMBER
107°C(225°F)
AS2O3 TO
UNDERGROUND
STORAGE
253 Kg/hr
(560 Ib/hr)
STACK
Figure 2. Roasting and Gas Cleaning Process (Design Flows).
-------
at a design rate of 2500 Kg/hr (2.7tons/hr) (dry basis). The first roaster is
operated at about 540-565°C (1000-1050°F) where most of the As, Sb, and S
are oxidized and volatilized. The overflow and underflow from the first
roaster are further treated in the second roaster at about 500-525°C
(925-975°F). The gases from the second roaster are fed through 2 parallel
sets of 3-stage cyclones operated at about 400°C (750°F), and the calcine
catch from the cyclones is combined with the roaster bed underflow for return
to a cyanidation circuit for recovery of gold.
The gases from the two banks of cyclones are combined and diluted with
air to reduce the gas temperature to 370°C (700°F) before being fed to the
electrostatic precipitator which consists of 2 units in parallel with 2
chambers in each. As will be shown by the data, sublimed arsenic (primarily
AS2O3) is not condensed in the precipitator and only particulate arsenic
(perhaps as AS2S2 carry over, non-volatile oxides such as As20s, or a com-
bined complex) is removed with the other entrained particulate from the
cyclones. The hoppers from the electrostatic precipitator are emptied peri-
odically and the catch is combined with the roaster and cyclone calcines.
The hot precipitator off-gases are then mixed with ambient air to
reduce the temperature to about 107°C (225°F) to condense the sublimed
arsenic (AS2O3). This particulate is caught in a 4-chamber baghouse before
the gases are sent to a double-walled stack. The arsenic trioxide dust from
the baghouse is removed periodically and conveyed to underground storage.
The reheat stack burner shown in Figure 2 is not required in normal operation.
Descriptive parameters for the precipitator are given in Table 1. The
precipitator is a double-walled unit, and electrically-heated air is circu-
lated in the shell to maintain temperature and avoid wall condensation of
arsenic. There are two chambers in each of the two parallel banks; however,
one transformer rectifier powers all four chambers of the precipitator.
During the test period, the roaster system was operated at around 90% of
design capacity so the actual SCA during the test was about 60 m2/(m3/sec)
or about 300 ft2/1000 acfm.
Parameters for the baghouse are shown in Table 2. The baghouse is
insulated, and the stack is also a double-walled unit to avoid wall condensa-
tion.
A key feature of the Red Lake System is the mixing chamber used for
cooling the hot gases from the precipitator to condense or desublime the
AS2O3. A simplified diagram of this mixer-cooler is shown in Figure 3.
Ambient air is fed into the center of the mixer, and the hot gases from the
precipitator are fed tangentially to create a swirling mixing action.
Ideally, condensation (or desublimation) of AS2O3 occurs only at the juncture
of the hot and cold gas streams. Wall condensation is avoided by keeping the
hot gases on the outer periphery and by heat tracing and insulating the walls.
SAMPLING PROGRAM
During the period of September 18-28, 1978, personnel from Southern
Research Institute and Radian Corporation carried out a test program in the
488
-------
TABLE 1. ELECTROSTATIC PRECIPITATOR
DESCRIPTIVE PARAMETERS
Item
Collection electrode area
(4 chambers-2 sets parallel)
No. of fields per chamber
Collection electrode spacing
Collection electrode dimensions
Corona electrode dimensions
(round wire)
Wire spacing
Number of gas passages
(per chamber)
Gas passage length
Volume flow rate (design)
Design temperature
Design efficiency
Design specific collection
area (SCA)
Measured efficiency (particulate)
Arsenic capture (total)
Metric
187.3m2
1
25.4 cm
1.83 x 1.83m
0.268 cm
22.86 cm
7
1.83 m
3.44 m3/sec
371°C
98%
54.37 m2/(m3/sec
98.3%
15%
English
2016 ft2
1
10 in.
6 x 6 ft
0.1055 in.
9 in.
7
6 ft
7300 acfm
700°F
98%
276.2 ft2/
1000 acfm
98.3%
15%
TABLE 2. BAGHOUSE DESCRIPTIVE PARAMETERS
No. of compartments
No. of bags (per compartment)
Bag material
Bag diameter
Bag length
Air-to-cloth ratio
(actual)
Measured efficiency (particulate)
Arsenic capture (total)
230
(10 rows of 23)
DrayIon T (acrylic)
10.16 cm (4 in.)
2.44 m (8 ft)
9.36 x 10"3 (m3/sec)/m2
(1.84 ft3/min/ft2)
99.9+%
99.8%
489
-------
TO BAGHOUSE
AMBIENT AIR
HOT GAS
HOT GAS
Figure 3. Gas mixer-cooler*
* U.S. Patent 4,126,425 and Can. Pat. 993,368 to Hatch Associates, Toronto,
Canada.
490
-------
Campbell Red Lake Mines Ltd. gold smelter at Balmerton, Ontario, Canada.
Radian Corporation personnel operated an integral sampling system
designed to obtain samples for elemental analysis, and a gas sampling system
for determining concentrations of sulfur oxides.
Southern Research Institute personnel operated cascade impactors for
determination of the particle size distribution of the particulate in each
gas stream, and a total particulate sampling system for determination of
total particulate loading in each stream. SRI personnel also monitored the
pressure drop across the baghouse, the operation of the roaster, and measured
the in situ resistivity of the particulate entering the baghouse (for possi-
ble modeling of a precipitator which could replace the baghouse).
The primary purpose of the test program was to evaluate the performance
of the gas treatment system which consists of a hot electrostatic precipitator
for recovery of metal values, a mixer-cooler for condensation (desublimation)
of AS2O3 and Sb203, and a cold baghouse for removal of fine condensed particu-
late. A simplified sketch indicating sampling points is shown in Figure 4.
Sample Point 1 at the precipitator inlet was under positive pressure, and
some difficulties were experienced in isokinetic sampling. The duct was ver-
tical with an upward gas flow and an internal diameter of 0.43 m (17 in.).
Sample Point 2, which was on one leg of the precipitator outlet, was
highly inaccessible and was not regarded as a satisfactory sampling point for
several reasons: (1) Sampling of only one leg of the precipitator required
an assumption that both sides in the parallel configuration were operating
identically, (2) Possible dust accumulation in the bottom of the duct could
affect flow measurements, and (3) The location was too close to converging
and diverging ducts for uniform particulate sampling. In spite of these
limitations, samples were taken from this 0.25 m (10 in.) square duct to
provide an approximation of precipitator performance. Because the precipi-
tator is closely connected to the mixer-cooler, this was the only sampling
point possible for estimation of emissions at the precipitator outlet/
mixer-cooler inlet.
Sample Point 3 at the outlet of the mixer-cooler provided data for the bag-
house inlet. Gas flow was vertically downward through the 0.71 m (28-in.) duct.
Sample Point 4 in the stack provided data for the baghouse outlet. The
internal diameter of the stack was 0.76 m (30-in.).
Insofar as possible, all samples were taken isokinetically to assure that
representative measurements were made.
There were some variations in operation during the test period that
affected the sampling program; however, adequate data were obtained to evalu-
ate the overall performance of the electrostatic precipitator and the baghouse.
A summary of plant operating data for the test period is shown in Table
3. These data show that the plant was in reasonably stable operation during
491
-------
REHEAT
BURNER
ELECTROSTATIC (2
TEMPERING Q) PRECIPITATOR
AIR
N)
AIR
rA
rS rS
r6!
BAG FILTER
—©
STACK
(TO STORAGE)
CYCLONES
DUST
ROASTERS
Figure 4. Sampling locations.
-------
TABLE 3. PLANT OPERATING DATA
Avg. Range
Freeboard Temperatures
Reactor No. 1, °F 1021 990-1037
°C 549 532-558
Reactor No. 2, °F 951 938-946
°C 511 503-508
Air Flow-to-Wind Box
Reactor No. 1, cfm 1200 1190-1220
m /sec 34.0 33.7-34.6
Reactor No. 2, cfm 920 860-1000
m /sec 26.1 24.4-28.3
ESP and Baghouse Temperatures
ESP Inlet, °F 737 713-768
°C 392 378-409
ESP Outlet, °F 586 531-620
°C 308 277-327
Baghouse Inlet, °F 245 229-250
°C 118 109-121
Baghouse Outlet, °F 232 228-236
°C 111 109-113
Stack, °F 216 212-219
°C 102 100-104
Precipitator*
AC Amps 10.1 9.5-10.5
AC Volts 418 413-423
DC Milliamps 40 35-45
DC Volts x 1000 57.3 56.2-58.6
* Excluding data at end of test period.
493
-------
the tests even though there was one operational shut-down during the test
period, and the precipitator had an electrical malfunction near the end of
the test period.
PARTICULATE MEASUREMENTS
Modified Brink cascade impactors were operated at the ESP inlet and
baghouse inlet (Sample Points 1 and 3), while Andersen Mark III cascade
impactors were operated at the precipitator outlet (Sample Point 2), and
University of Washington Mark III (Pilat) cascade impactors were operated
at the baghouse outlet (Sample Point 4). The collection substrates (glass
fiber) used in the Brink and Andersen impactors were conditioned by acid
washing in the laboratory to decrease the amount of substrates weight gain
due to gas phase reactions. The collection substrates used in the University
of Washington (Pilat) impactors were greased stainless steel substrates
which were prepared in the laboratory prior to use in the field. All impac-
tor data obtained at Red Lake were reduced using a computer program described
in EPA Report Number 600/7-78-042, A Computer-based Cascade Impactor Data
Reduction System, by J.W. Johnson, G.I. Clinard, L.G. Felix, and J.D. McCain,
March 1978.
The impactor data collected at the ESP inlet sampling location are pre-
sented in Figure 5 as cumulative mass loading vs. particle size. Figure 6
presents the same data as cumulative percent vs. particle diameter. As can
be seen from Figure 6, 50% of the particulate was less than 16 ym. A max-
imum particle diameter of 25 ym was used in the impactor data reduction pro-
gram for each sampling location, since the size of the largest particle
found in the cyclone catches of the Brink impactors operated at the ESP and
baghouse inlets was 25 ym.
A limited number of impactor runs were made at the precipitator outlet
to approximate the cumulative mass loading of particulate less than two
microns in diameter which would be present at the outlet of the condenser
in addition to particulate resulting from condensation (or desublimation).
Figures 7 and 8 present data on the cumulative mass vs. particle size from
the ESP outlet and baghouse inlet. Figure 7 shows that the mass loading of
particulate below 2 microns diameter at the ESP outlet was less than 7 X 103
mg/ACM at a gas flow rate of 3888 dscm/hr (from Table 5). Figure 8 shows
that the mass loading of particulate below 2 microns diameter at the baghouse
inlet was about 7 X 10 mg/ACM at a gas flow rate of about 16,779 dscm/hr.
The ratio of these numbers (7 X 10 3 X 3888 t 7 X 10 ** X 16,779) indicates
that less than five percent of the particulate entering the baghouse was
particulate at the ESP outlet; therefore, greater than 95 percent of the
particulate entering the baghouse resulted from condensation in the mixer-
cooler. Figure 9 presents the baghouse inlet data as cumulative percent vs.
particle diameter. From Figure 9, it can be seen that the mass median
diameter of the particulate entering the baghouse (primarily as AS2O3) was
1.2 |jm, and 75% of the particulate mass was less than two microns.
Figure 10 presents the collection efficiency vs. particle diameter data
for the baghouse and shows a minimum in collection efficiency at about 5
microns of approximately 99.8%. The bimodal distribution shown in Figure 10
494
-------
lO^x
<
N lO^i
M
-1 104-:
lb3-
10P-
MG/ACM - MILLIGRAMS/ACTUAL CUBIC METER
GR/ACF - GRAINS/ACTUAL CUBIC FOOT
irlO3
::1&
\
a
"101
::1CP
J-IO"1
h—i i 11 ml 1—i m ihH 1—k i mill
10"1 10° 101 10P
PARTICLE DIAMETER (MICROMETERS)
Figure 5. Average ESP inlet cumulative size distribution.
495
-------
99.93
UJ
>
H
I
0.2"
0.01
h—> j «i m>| i—i i 1111H 1—II 11 ml
1O*"1 10° 101 10^
PARTICLE DIAMETER (MICROMETERS)
Figure 6. Average ESP inlet cumulative percent versus
particle diameter.
496
-------
ioV
*
103--
1CP
MG/ACM - MILLIGRAMS/ACTUAL CUBIC METER
GR/ACF - GRAINS/ACTUAL CUBIC FOOT
tIO1
»1CP
¦LlO"1
—i—i < nnH 1—mm nil a—i i 111 ni
10"1 1CP 101 10*
PARTICLE DIAMETER (MICROMETERS)
Figure 7. Average ESP outlet cumulative size distribution.
497
-------
10^1
\
104::
~
_i
£ 103::
M
H
<
MG/ACM - MILLIGRAMS/ACTUAL CUBIC METER
GR/ACF - GRAINS/ACTUAL CUBIC FOOT
rrl(f
101 g
::10P
!SJ
M
-LlO"1
lO^H 1—i i 11 ml 1—i i 11 nil •—m n ml
1CT1 10P 101 10s
PARTICLE DIAMETER (MICROMETERS)
Figure 8. Average baghouse inlet cumulative size distribution.
498
-------
99.99
99.0
99.5
10
5
2
1
0.5
0.E
o?ol
0.01
I
10
H—I I I Mill 1—I I I Hill 1—I I I I lll|
1CP
101
10^
PARTICLE DIAMETER (MICROMETERS)
Figure 9. Average baghouse inlet cumulative percent versus
particle diameter.
499
-------
lOSr
10^1 r
10"1 ¦ r
10 ¦ ~
10"3::
10
-A.
irO.O
"30.0
499.0 ^
UJ
H—I I I Mil
H—1 I I I I l-H 1—MM
!r99.9
M
U
~
h-
ori nn
• r99.939
10"1 1CP 101
PARTICLE DIAMETER (MICROMETERS)
Figure 10. Baghouse particulate collection efficiency.
500
S8'
-------
is not abnormal for baghouse performance. In other studies, the increased
penetration around 0.5 ym has been attributed to the lack of inertial or
diffusional collection mechanisms, and the increase around 5.0 ym has been
attributed to seepage1. The mass loadings from all the impactor runs used
in the data reduction programs are presented in Table 4. Based on these
data, the average particulate collection efficiency of the baghouse exceeds
99.9%. Data from the mass sampling trains and the trace element trains also
indicate an average efficiency above 99.9% for the baghouse.
Impactor data for the precipitator show a particulate collection effi-
ciency of 98.3%. Trace element data (collected independently by Radian) showl
a somewhat lower precipitator efficiency of 97.1%. However, all data agree
that the overall particulate collection efficiency of the "hot ESP/cold
baghouse system" exceeds 99.9%.
GAS SAMPLING
Two types of modified EPA Method 5 sampling trains were used at each
sampling site. These two trains differed from each other only in the number
of impingers and the solutions in them. An illustrative gas sampling train
is shown in Figure 11.
The trace element train absorbed basic and neutral species in the nitric
acid impingers and acidic species in the sodium hydroxide impinger solutions.
The gas was then scrubbed in 6% H2O2 and dried by silica gel.
The sulfur oxide train contained only five impingers. The first con-
tained 80% isopropanol for absorption of SO3. The second and third contained
6% H2O2 for absorbing SO2, the fourth was empty, and the fifth contained
silica gel.
Both trains contained in-stack and out-of-stack filters which were used
to collect particulate (in-stack) and vaporous arsenic trioxide which con-
densed in the probe after removal of other particulate (out-of-stack).
After sampling, extensive cleaning of the sampling train was required
to remove AS2O3 which had condensed in the probe and glassware. The nozzle
was rinsed with NaOH solution to dissolve AS2O3. Since AS2O3 is condensable
and adheres to walls of sampling equipment, a rigid washing protocol was
used to assure maximum recovery from the sampling trains.
Solid (particulate) and gaseous (volatile) components from the trace
element sampling train were analyzed quantitatively for four elements:
arsenic, antimony, lead, and selenium. A summary of the gas stream measure-
ments and an overall distribution of the four trace elements are shown in
Table 5.
The grain loading data in Table 5 show a particulate removal efficiency
of about 97.1% for the ESP and 99.9% for the baghouse based on in-stack
measurements. Lead and selenium analyses show only about a 90% overall
removal, but the amounts present at the ESP inlet were very low. Antimony
removed was about 99.5%, but the emission from the baghouse of 0.011 Kg/hr
501
-------
TEMPERATURE TEMPERATURE
SENSOR
10% NaOH 6% N2O2 SENSOR
IMPINGER IMPINGERS IMPINGER /T\
10% HNO3
IMPINGERS
DRY
HEATED
BOX
TEMPERATURE
SENSOR
STACK WALL
FILTER
n 00
o
S-TYPE
PITOT TUBE
SILICA GEL
DESSICCANT
VACUUM
LINE
PITOT TUBE
GAUGE
ICE BATH
FILTER HOLDER
TEMPERATURE SENSORS
BY-PASS
VALVE
VACUUM GAUGE
ORIFICE
MAIN VALVE
PUMP
DRY GAS METER!
ORIFICE GAUGE
Figure 11.
Trace element gas sampling train.
-------
TABLE 4. IMPACTOR
Sample
Location Run# Date Time, min
ESP Inlet
RLI-3
9/19/78
4
ESP Inlet
KLI-6
9/20/78
4
ESP Inlet
RLI-7
9/20/78
4
ESP Inlet
rli-8
9/20/78
4
Bghs Inlet
RLI-12
9/21/78
4
Bghs Inlet
RLI-13
9/21/78
4
Bghs Inlet
RLI-14
9/21/78
4
Bghs Inlet
RLI-16
9/25/78
4
Bghs Inlet
RLI-17
9/25/78
4
Bghs Inlet
RLI-18
9/25/78
4
Bghs Inlet
RLI-19
9/25/78
4
Bghs Inlet
RLI-21
9/26/78
4
Bghs Inlet
RLI-22
9/26/78
4
Bghs Inlet
RLI-23
9/26/78
4
Bghs Inlet
RLI-24
9/26/78
4
Bghs Outlet RLO-1
Bghs Outlet RLO-2
Bghs Outlet RLO-3
Bghs Outlet RLO-5
Bghs Outlet RLO-6
ESP Outlet RLO-8
ESP Outlet RLO-9
9/19/78 360
9/20-21/78 960
9/20-21/78 960
9/25/78 480
9/25/78 480
9/26/78 4
9/26/78 4
Baghouse efficiency = 99.99+%
ESP efficiency = 98.3%
LOADINGS
Impactor Temp.
Mass Loading
gr/SCF mg/SCM
600
316
18.73
580
304
114.82
555
291
50.36
560
293
58.82
244
118
5.25
230
110
5.07
225
107
5.34
220
104
4.56
220
104
5.04
225
107
5.14
230
110
5.80
228
109
4.55
243
117
5.19
228
109
5.97
243
117
6.70
228
109
0.00024
226
108
0.00029
226
108
0.00028
228
109
0.00027
228
109
0.00029
575
302
1.017
575
302
1.079
4.29x10
2.63x10
1.15x10
1.35x10
1.20x10
1.16x10
1.22x10
1.04x10
1.15x10
1.18x10
1.33x10'
1.04x10'
1.19x10'
1.37x10'
1.53x10'
0.550
0.657
0.642
0.621
0.667
2.33x10:
2.47x10:
-------
TABLE 5. GAS STREAM MEASUREMENTS AND
TRACE ELEMENTS DISTRIBUTION
ESP
Inlet
ESP
Outlet
Baghouse
Inlet
Baghouse
Outlet
Temperature
°C
°F
380
716
327
621
116
240
113
236
Velocity
m/ sec
ft/sec
21.3
70.0
21.9
71.9
16.5
54.0
16.9
55.3
Grain Loading
grams/dscm
grains/dscf
35.2
15.4
1.01
0.442
11.40
4.98
0.00094
0.00041
Flow Rate
dscm/hr
dscf/hr
3899
137,680
3888
137,314
16,779
592,526
19,817
699,833
Trace Elements (total)
As-Kg/hr
-lb/hr
Pb-Kg/hr
-lb/hr
Sb-Kg/hr
lb/hr
Se-Kg/hr
-lb/hr
179
152*
105*
393
334
232
0.14
0.05
0.32
0.11
-
2.1
0.40
—
4.7
0.88
-
0045-
0.0017
-
010
0.0037
_
0.20
0.44
0.015
0.034
0.011
0.025
0.0004
0.0009
* Not measured simultaneously.
504
-------
(0.025 lb/hr) is also very low. Overall arsenic removal for the ESP and
baghouse combined was above 99.9%.
A more detailed analysis of the solid and vapor distribution for arsenic
is shown in Table 6. This table shows that most of the arsenic entering the
precipitator is in a volatile form and relatively little is caught by the
ESP. However, most of the arsenic (as AS2O3) entering the baghouse (after
the gas stream has been quench-cooled) occurs as particulate and this is
efficiently collected in the cold baghouse. Collection of particulate
arsenic in the baghouse was greater than 99.95%; however, overall arsenic
collection efficiency in the baghouse was only 99.8% due to passage of vola-
tile AS2O3. The total arsenic emissions from the system including both
particulate and volatiles were less than 0.2 Kg/hr (0.5 lb/hr).
CONCLUSIONS
The test program conducted at Campbell Red Lake Mines confirmed that
their system using a hot electrostatic precipitator, ambient air quenching
to condense AS2O3, and a cold baghouse is an effective technology for control
of arsenic emissions from a nonferrous smelter. By comparison to most non-
ferrous smelters, the Red Lake installation is small. However, the control
technology for arsenic has been shown to be highly efficient and would likely
be applicable to other nonferrous installations. Major considerations would
involve scale-up and sizing of the dilution mixer-cooler and an economic
evaluation of the baghouse system for handling increased gas volumes.
ACKNOWLEDGMENTS
This test program was sponsored by the Metals and Inorganic Chemicals
Branch of EPA's Industrial Environmental Research Laboratory, Cincinnati,
Ohio, John O. Burckle, Project Officer. Assistance in arranging the test was
provided by Environment Canada's Air Pollution Control Directorate, Ottawa,
Ontario, W. A. Lemmon, Chief, Mining, Mineral and Metallurgical Division.
The gas sampling and analysis were conducted by Radian Corporation,
Austin, Texas under the direction of Dr. John C. Terry.
We would like to give special thanks to Stewart Reid, General Manager,
Ken Dickson, and Scott Roberts, Campbell Red Lake Mines Ltd., Balmerton,
Ontario, Canada who were very cooperative in providing access to the plant
and in making modifications required to conduct the test program.
REFERENCES
1. D.S. Ensor, R.G. Hooper, and R.W. Scheck, "Determination of the Fractional
Efficiency, Opacity Characteristics, Engineering and Economic Aspects
of a Fabric Filter Operating on a Utility Boiler," EPRI FP-297, November
1976.
2. Goodfellow, H. and Gellender, M., "Arsenic Poses Tricky Recovery Task,"
Canadian Chemical Processing, pp. 26-27, February 1978.
505
-------
TABLE 6. SOLID AND VAPOR DISTRIBUTION OF As
(Trace element trains)
Phase
SOLID
gm/dscm
lb/dscf
Kg/hr
lb/hr
VAPOR
gm/dscm
lb/dscf
Kg/hr
lb/hr
TOTAL
Kg/hr
lb/hr
ESP ESP Baghouse Baghouse
Inlet Outlet Inlet Outlet
1.21 2.68xl0~2 6.20 1.8xl0_3
7.58x10"5 1.67xl0~6 3.87xl0_lt l.lxlO"7
4.7 0.10 104 0.02
10.4 0.23 229 0.07
44.6 39.0 7.82xl0~* 9.14x10'
2.78x10"3 2.43xl0-3 4.89xl0"6 5.71x10
174 152 1.31 1.81x10'
383 334 2.90 3.99x10
179 152 105 0.20
393 334 232 0.44
506
-------
Goodfellow, H. et al, "Arsenic Removal from Roaster Off-Gases," Fourth
International Clean Air Congress (1978).
507
-------
CONTROL OF SALT LADEN PARTICULATE EMISSIONS
FROM HOGGED FUEL BOILERS
By:
Michael F. Szabo
Richard W. Gerstle
PEDCo Environmental, Inc.
11499 Chester Road
Cincinnati, Ohio 45246
and
Larry Sims
U.S. Environmental Protection Agency
Region X
1200 Sixth Avenue
Seattle, Washington 98101
ABSTRACT
This paper presents the results of an evaluation of excessive salt
emissions from sixteen hogged fuel boilers in Washington, Oregon, and Alaska.
Logs transported or stored in sea water, absorb substantial amounts of salt
which is noncombustible, and is emitted as a fine particulate when the hogged
fuel is burned, contributing to opacity and particulate emissions. Control
measures considered are fuel pretreatment, combustion modifications, use of
conventional particulate control devices {electrostatic precipitators,
fabric filters, and wet scrubbers), and several novel particulate control
devices. The best available control technology appears to be a mechanical
collector - fabric filter combination; some electrostatic scrubber type de-
vices have also shown excellent collection capability.
508
-------
CONTROL OF SALT LADEN PARTICULATE EMISSIONS
FROM HOGGED FUEL BOILERS
INTRODUCTION
In some areas of the country, principally the Pacific Northwest, wood
fuel that is hogged,* and fired in industrial boilers, comes from logs that
have been transported or stored in sea water, from which the bark may absorb
substantial amounts of salt. There are 16 salt emitting installations in
Washington, Oregon and Alaska, or about 4 percent of the total hogged fuel
boiler population in these states, as shown in Table 1.
TABLE 1. HOGGED FUEL BOILERS WITH EXCESSIVE SALT EMISSIONS (1)
Total no. of
hogged fuel
boilers
No. of boilers
emitting salt
% of
total
Approximate heat
input of boilers
emitting salt,
109 J/h (106 Btu/h)
Washington
98
10
10
1,051 (995)
Oregon
318
2
0.6
362 (343)
Alaska
10
4
40
786 (744)
Total
426
16
4
2,199 (2082)
Salt emission from hogged fuel boilers are a significant problem,
principally because the salt portion of the particulate is primarily sub-
micron in size. Particles larger than 5 ym are deposited in the nasal
cavity or naso-pharnyx. The smaller particles, however, are deposited in the
lungs, including over 50 percent of the particles from 0.01 to 0.1 v>m diame-
ter that penetrate the pulmonary compartment. The tendency of particulates
to penetrate the respiratory systems and be captured is mainly a function of
their geometry rather than their chemical properties.
Health effects of the captured fine particles depend largely on their
chemical or toxic qualities, except for long fibrous material, whose physical
qualities also provide potential for irritation of tissue. Because of the
many unknown factors, it is unwise to generalize concerning health effects of
fine salt particulates from hogged fuel boilers.
*A term derived from the machine that processes the mixture of wood and bark
before it is burned, and is called a "hog" or "hogger".
509
-------
The other important aspect of the problem is that salt-laden particulate
emissions, after passing through a conventional mechanical collector, which is
not efficient in collecting fine particulates, are usually in violation of
both particulate emission and opacity regulations. A plume from these opera-
tions is often highly visible and aesthetically objectionable in the com-
muni ty.
Depending on the salt fraction of the fuel and the type of control
device used, salt particles can constitute 30 to 90 percent of stack emissions
from these boilers. Firing of auxiliary fuels such as No. 6 fuel oil, con-
currently with the hogged fuel can also contribute to stack opacity. Although
emissions from these boilers violate regulations set forth in the State Imple-
mentation Plans (SIP's) and local regulations, available ambient air data
obtained near several of these plants have not shown violations of ambient air
quality standards.
In responding to the salt emissions problem, some companies have cited
potentially high costs and technical problems associated with control of these
emissions with secondary control equipment. The smaller companies, particu-
larly, regard control of salt emissions as a financial burden.
The purpose of this paper is to examine the salt emission problem, and
define the technical problems, and compliance prospects of various control
techniques.
CHARACTERISTICS OF SALT LADEN HOGGED FUEL
The properties of wood residues and bark fuels can vary so widely that no
standard specification is possible.
Many species of wood can be used as fuel, but some are better than
others. Wet cedar bark, for example, is stringy and difficult to reduce in
size. By comparison, dry Douglas fir bark is considered a very desirable
fuel.(1)
Other types of wood fuel are used to supplement the hogged fuel. Among
these are sawdust, wood chips and shavings, wood waste material, and sander-
dust. Sludges and spent liquors from wastewater treatment processes are
burned occasionally, more as a means of disposal than for any advantageous
fuel characteristics. At some installations where downtime of the hogged fuel
boiler is critical to plant operations, the system incorporates provisions for
the burning of an alternative fuel, such as No. 6 fuel oil.
The supplemental fuels are used in various quantities and combinations
depending largely upon the function and design of the plant. Except for
sludge and spent liquors, the supplemental fuels have more desirable burning
characteristics (e.g., lower moisture content and higher heating value per
equivalent mass) than does the hogged fuel. The supplies of hogged fuel,
however, are usually more dependable; also, it is available in much larger
quantities and at comparatively lower costs.
510
-------
Salt Content of Hogged Fuel
In the northwest coastal areas of the United States, logs transported by
water can be subjected to a salt or saline environment for long periods of
time (sometimes weeks or months), especially if they are stored in saltwater.
Such storage allows ample time for the deposition of the various salts and
other chemicals present in seawater upon and in the bark of the logs.
Hogged fuel that is transported or stored in salt or saline waters and
used in power boilers has been shown to contain anywhere from 0.09 to 1.2
weight percent salt (as sodium chloride, NaCl). The primary factors influ-
encing salt content are length of time the logs remain in saltwater and
concentration of salt in the water. These factors, in turn, are affected by
plant location with respect to the wood source, log supply and demand, type
of storage (i.e., dry or wet, fresh or salt water), and, in tidal estuaries
or rivers, the amount of mixing of seawater with freshwater. MacLean and
MacDonald {2) found that the average salt content of bark from hemlock after
6 months flotation was 1.44 to 3.13 percent, of which roughly half was ab-
sorbed in the first 3 weeks.
Since the concentrations and proportions of the various salt fractions
in seawater vary greatly, it is difficult to predict the total emissions of
particulate and salt from hogged fuel boilers.
The effect of mixing of fresh and salt waters is evident when comparing
the salinity at high and low tides in a coastal river. Salinity at low tide
would tend to be lower because there is a greater flow of fresh water past a
given point as the tide moves out. At high tide, the reverse is true. Typi-
cal values are 9800 ppm NaCl at low tide and 13,000 ppm at high tide.
Salt Measurement Techniques
Salt content of wood fuel or flue gas can be expressed as a percentage
of NaCl, of equivalent chloride, or of total sea salt. This can lead to con-
fusion in interpretation of data, since no standard method is consistently
applied to determinations of salt content and the method of analysis often is
not given. The amounts of salt reported as NaCl, all chloride salts, and
total sea salts can differ significantly.
In this regard the Council of Forest Industries of British Columbia has
stated the following (3):
"The salt content of particulate matter is commonly expressed as equiva-
lent sodium chloride. This value is obtained by multiplying the chlor-
ide ion concentration by 1.58, the factor representing the weighted
average of the ratios of chloride salt molecular weight to the chloride
content of molecule. Actually, the value so obtained does not represent
sodium chloride (which would be obtained by multiplying chloride by
1.66) but the total chloride salts. A further correction actually is
required for accuracy because chloride salts represent about 90 percent
of sea salts and so the factor to get total sea salts from chloride
content is 1.76. This latter correction rarely has been made."
511
-------
This statement emphasizes the importance of identifying the basis upon
which salt concentrations are presented and the need for a standard reference
method.
Particle Size Distribution of Salt Emissions
The size distribution of the salt particles strongly affects the removal
efficiency of control equipment. Particulate emissions from boilers burning
hogged fuel with no salt can be handled adequately by conventional control
devices. With salt-laden hogged fuel, however, the size range of the parti-
culate emissions is much smaller; and the size fractions below 1 ym typically
are composed of more than 50 percent salt.
In an analysis of the effect of salt (NaCl) on overall mean particle size
in emissions from hogged fuel boilers, the salt fraction reduced the overall
mean particle size by a factor of approximately 10 (4) (see Figure 1). As
this figure shows, the mean particle size of the nonsalt fraction is about
17.5 ym. The mean particle size of the salt fraction is about 0.23 ym. The
overall mean particle size, about 2.15 ym, indicates the extent of size
reduction caused by salt in the emissions. Some data show that the overall
mean particle size of the salt-laden particulate can be as low as 0.2 ym.
Comparison of particle size data from the Weyerhauser North Bend plant
with data from Crown Zellerbach's Port Townsend plant and data from British
Columbia verifies the reduction in mean particle size due to salt particles.
All of these impactor measurements were taken at the outlet of a multiclone or
cinder collector. The reduction in the overall mean particle size can result
in dust even finer than the composite shown in Figure 1, depending on the size
distribution of the nonsalt fraction of the dust.
PREVENTIVE CONTROL TECHNOLOGY TO REDUCE SALT EMISSIONS
Handling and Pretreatment of Fuel
Among various preventive methods for reducing salt particulate emissions
from hogged fuel boilers, the most obvious is not to transport the logs via
saltwater; in most cases this is not possible. Transport by flat rafting
rather than in bundles will reduce the salt content of hogged fuel, as will
reducing the duration of storage in salt water (see Figure 2).
TYPICAL BUNDLE
TYPICAL FLAT RAFT
WATER
LINE
Figure 2. Storage of logs by flat raft and bundle.(5)
512
-------
100
80
O ASH ONLY
A SALT AND ASH
~ SALT ONLY
COMPOSITE SIZE DISTRIBUTION
40 -
* = 17.5u"i
UJ
x = 2.1 5uin
og ¦ '12.6 X
CC
Q_
0.9
0.8
0.7
0.6
0.4
0.3
i « 0.23 ui"
og « 12.6
0.2
95
98
30 40 50 60 70
80 85 90
20
10
2
5
PERCENT LESS THAN STATED PARTICLE SIZE
Figure 1. Particle size distribution at Weyerhauser Co.,
North Bend, Oregon, plant (composite from 11 tests).
513
-------
About half of all salt absorbed in 6 months is absorbed in the first 2 or 3
weeks of contact with seawater.
Hydraulic deparking of logs with fresh water is reported to have reduced
salt content in one instance but not in another (5).
Bark pressing can reduce moisture content by 50 percent and remove sub-
stantial quantities of salt in the process. This can result in lower opacity
and particulate emissions because of increased boiler efficiency and fewer
fine particles of salt. The disadvantage, however, is serious water pollution
from the bark pressate wastewater.
Other methods of predrying fuel do not reduce the salt content, but do
reduce combustible emissions by improving boiler efficiency.
Combustion Modifications
Combustion modifications can increase boiler efficiency and thereby
reduce salt emissions because salt is not combustible.
About 65 to 85 percent of the salt in bark passes out of the chimney as
particulate emissions; the remainder is retained in the boiler, on the grates,
or on boiler surfaces, reducing efficiency of the boiler and disrupting uni-
form air flow through the fuel beds. Plant engineers therefore should try to
reduce both the salt and moisture content of the bark at reasonable cost. The
resulting improvement in fuel quality will reduce consumption of auxiliary
fuel and boiler cleaning requirements. Additional research is needed into
ways to slow the quenching effect on vaporized salt leaving the combustion
zone in hogged fuel boilers. This would generate large particles of salt that
would be more easily collected by a secondary control device. In most opera-
tions, however, additional control measures will be necessary to reduce salt-
laden particulate emissions to acceptable levels.
SECONDARY CONTROL TECHNOLOGY TO REDUCE SALT EMISSIONS
Four types of conventional control devices are used to reduce particulate
emissions from hogged fuel boilers: mechanical collectors, electrostatic
precipitators (ESP's), wet scrubbers, and fabric filters. Only the latter
three can remove submicron salt particles. Mechanical collectors, nonethe-
less, are used on salt-laden hogged fuel boilers and ESP's are not.
The following sections on each control device briefly summarize appli-
cability for controlling salt emissions, performance on hogged fuel boilers
(salt-free and salt-laden fuel), operation and maintenance problem, and dis-
posal techniques for collecting salt/particulates. The best available method
of collecting salt/particulates is also discussed.
514
-------
Mechanical Collectors
Particulate emissions from hogged fuel boilers have traditionally been
controlled by mechanical collectors. Although mechanical collectors generally
do not meet applicable emission regulations when controlling salt-laden
particulate, they are still important as part of a total system of particulate
control. These collectors are usually designed for dust loadings of 2 to 11
g/nr (1 to 5 gr/scf) (6). Particle size distribution of ash from combustion
of nonsalt bark is normally 30 to 40 percent less than 10 micrometers.
Pressure drop across the mechanical collector ranges from 3.8 to 6.4 cm (1.5
to 2.5 in. water) for best collection efficiency.(6)
With ash from salt-free bark, mechanical collectors can sometimes achieve
collection efficiencies of 85 to 90 percent with outlet loadings of 172 to 258
kg/J (0.4 to 0.6 lb/10 Btu). A collection efficiency of 30 to 70 percent is
more realistic estimate. Even under optimum conditions opacity may be a
problem, especially with stoker-fired boilers, which respond slowly to upset
conditions. Opacity can be limited to 20 percent during normal operation, but
may exceed 20 percent during boiler upsets.
With salt-laden particulate, the collection efficiency of the mechanical
collector is considerably lower because of the fines of the salt particles;
overall efficiencies of 30 to 50 percent would be expected. Test data from
Weyerhauser's North Bend plant confirm this observation. (7)
Operation and maintenance problems associated with cyclone system re-
moving particulates from flue gas of hogged fuel boilers are principally due
to plugging, leakage of air into the system, and changes in gas flow to the
cyclone.
Electrostatic Precipitators (ESP's)
Although ESP's are used widely in controlling particulate emissions from
combustion sources, they are rarely used on boilers fired with hogged fuel.
Several successful installations on boilers that burn salt-free wood have been
put into operation in the last 5 or 6 years.
From the precipitation standpoint, ash from hogged fuel boilers poses two
potential problems that must be accounted for in design:(1) 1) since it is
carboneceous, the ash has a lower resistivity than coal ash and tends to lose
its charge quickly, and 2) the ash is prone to reentrainment because of its
low density, coupled with flake-like particle shape. When the ash is salt-
laden, the submicron salt particles require a conservative design, i.e., a
larger collecting surface.
Performance of ESP's on boilers burning salt-free hogged fuel has re-
portedly been excellent. Where an ESP was retrofitted downstream of a cyclone
at a paper mill, the average outlet loading was 0.04 g/m3 (0.018 gr/acf);
guarantee was 0.06 g/m3 (0.025 gr/acf).
515
-------
The only reference to performance of an ESP on a boiler burning salt-laden
hogged fuel comes from the Council of Forest Industries of British Columbia.(3)
Test data from a pilot unit installed at the Victoria Sawmill Division of B.C.~
Forest Products, Ltd., showed that performance varied from excellent [0.05 g/m
(0.02 gr/dscf)] to poor [0.7 g/m3 (0.3 gr/dscf)].(3) The program showed that
the ESP could reduce salt fume to an acceptable level when it was maintained
and operated within precise limits. When control of operation was not virtually
perfect, however, the performance deteriorated markedly.
Expected operation and maintenance problems would be the possibility of
fire from "char" build-up on precipitator walls, and especially in the hoppers.
Reentrainment is also a problem with bark ash because of its low resistivity.
Pilot scale tests have reportedly been performed on a wet-ESP controlling
salt laden particulate emissions from a hogged fuel boiler in Alaska, but data
from this installation were not available for this paper.
Wet Scrubbers
Wet scrubbers are applfed to a number of hogged fuel boilers to control
particulate emissions. Most are installed downstream of a multiple cyclone
collector. In salt-free applications under normal operating conditions the
outlet loadings range from 0.05 to 0.14 g/m3 (0.02 to 0.06 gr/scf) at pressure
drops of 15 to 38 cm (6 to 15 in.) water.(8)
Only one boiler fired with salt-laden hogged fuel is currently equipped
with a venturi scrubber, Crown Zellerbach, Port Townsend plant; this scrubber
is installed downstream of a multiclone.(9) The reported outlet loadings are
0.16 to 41 g/m3 (0.07 to 0.18 gr/scf), depending on the salt content of the
fuel. A salt concentration greater than 1 percent will limit the performance
of this venturi scrubber, which operates at a pressure drop of 38 to 51 cm (15
to 20 in.) water. The salt content of particulate emissions from the scrubber
is reported to be 50 to 70 percent or more. Opacity of the stack effluent is
estimated at 35 percent. This scrubber would require a pressure drop greater
than 51 cm (20 in.) water to Achieve the roughly 80 percent efficiency needed
to meet the 0.23 g/m3 (0.10 gr/dscf) particulate emission regulation, assuming
a typical inlet concentration of 1.14 g/m3 (0.50 gr/dscf), and greater than 1
percent salt in the fuel.
Using unpublished impactor data for this installation, an approximate
fractional efficiency curve was prepared, and is shown in Figure 3. Overall
mass efficiency from these data is approximately 68 percent. The character-
istic sharp increase in penetration as particle size decreases, is evident in
this curve.(10)
Common maintenance problems for venturi scrubbers include plugging of
lines, nozzles, and pumps and erosion/corrosion and scaling of internal com-
ponents.
Fabric Filters
Fabric filters are used on four hogged fuel boiler installations: Simpson
Timber Co. in Shelton, Washington; Long Lake Lumber Co. in Spokane,
516
-------
OVERALL MASS EFFICIENCY = 68%
.1
10
5
1.0
PARTICLE SIZE, ym
Figure 3. Approximate penetration as a function of
particule size for a venturi scrubber operating
on a salt-laden bark/oil fired boiler.(10)
517
-------
Washington; Georgia Pacific Corp. in Bellingham, Washington; and the most
recent installation, which started operation in the latter part of June 1979,
at Scott Paper Co. in Everett, Washington. All of these installations except
Long Lake Limber use logs that have been transported in sea water. Table 2
summarizes design parameters for these four installations.
Plant engineers estimate the overall mass efficiency of the Simpson
Timber Co. baghouses at 90 to 95 percent, stating that a number of bags are
always broken, due to improper installation, faulty construction, etc.
Outlet mass loadings were tested at 0.03 g/sm (0.01 gr/dscf), for the smaller
unit and 0.092 g/sm^ (0.09 gr/dscf) for the larger unit.(11) Opacity did not
exceed 5 percent (mostly only heat waves were visible).
The Georgia Pacific Facility started up in February, 1979, and is reported
to be operating well with no major problems. The plant was only recently
tested for compliance, and plant officials unofficially estimate outlet loading
to be approximately 0.05 g/m^ (0.02 gr/dscf) with zero visible emissions. This
is well within emission limitations.
Operating data are not available for the Scott Paper Co. facility since it
only recently began operation.
The Long Lake facility reports an overall mass efficiency of 99 percent.
However, operation at this facility has been marred by two fires.
Although no fractional efficiency data are available, the dramatic re-
duction in opacity with installation of the baghouses indicates that they are
efficient collectors of submicron salt particles. Fractional efficiency data
from tests of fabric filters in other industries support this view.
Disposal of Salt-Laden Ash
The disposal of salt-laden ash from secondary collectors is a problem
sometimes overlooked in evaluating the salt emissions problem. Landfilling of
the ash is complicated by the presence of salt, which presents a potential
leachate problem. Return of a slurry containing salt and ash to a bay or ocean
is unacceptable because of the ash content, and the ash cannot be sold for
cement or asphalt production because of the salt.
Acceptable disposal practices include providing an impermeable liner for
an ash pond or diluting the salt concentration by mixing the ash slurry with
other plant wastewater streams or with municipal wastewater and processing the
ash slurry by conventional wastewater treatment. Sometimes, the sludge is
returned to the boiler. Wet treatment seems to be more effective in terms of
efficiency and economy. Either method, however, increases capital and opera-
ting costs.
NOVEL FINE PARTICULATE CONTROL DEVICES
Experience at plants with salt particulate emissions shows that, except
for fabric filters, the conventional control devices cannot consistently
518
-------
TABLE 2. DESIGN PARAMETERS FOR BAGHOUSES ON
HOGGED FUEL BOILERS
Simpson
Timber3
Long Lake
Lumber
Georgia-
Pacific3
Scott
Paper3
Volume flowrate, acfm
130,000
25,000
180,000
260,000
Inlet gas temperature,
°F
500
400(200)b
440
420
A/C ratio, acfm/ft^
4.5
4.0
4.0
3.9
Bag cleaning method
Pulse jet
Pulse jet
Pulse jet
Pulse jet
Pressure drop, in. ^0
9-9.5
5.8-6.8
not > 7
N.A.
Bag fabric
Teflon-coated
fiberglass
Nomex
Teflon-coated
fiberglass
Teflon-coated
fiberglass
Precollector
Mechanical
cinder
collector
None
None
None
Material handling
system
Screw
conveyor
Screw
conveyor
Screw
conveyor
Water
sluice
3
Metric conversions: acfm x 0.028 = m /min ?
acfm/ftz x 5.09 = m3/s per m
(°F - 32)/1.8 = 6C
in. H20 x 2.54 = cm H20
aSalt-laden hogged fuel boilers.
^Operating temperature.
N.A. - not available.
519
-------
reduce emissions enough for compliance with particulate and stack opacity
regulations. Thus, there is a need for more effective control, especially in
the submicron particle size range. This section evaluates some novel or
promising particulate control devices now being tested for control of salt
emissions from hogged fuel boilers.
Classification of Novel Control Devices
Three categories of wet scrubbing(12) are considered promising for control
of the submicron salt particles from hogged fuel boiler emissions: foam
scrubbing, flux force condensation, and electrostatic scrubbing. Of these,
electrostatic scrubbing appears to offer the greatest potential for full-scale
application to salt-laden hogged fuel boilers.
Dry scrubbers also are considered here as a novel device, since there have
been two pilot tests and one full-scale test on three salt-laden hogged fuel
boilers.
Electrostatic Scrubbing-This device embodies the principles of an elec-
torstatic precipitator and a scrubber. The basic idea is to augment the
collection processes associated with spray scrubbing and electrical collection
forces. This involves the use of electrostatically charged water droplets or
charged pollutant particles or both.(13,14) The scrubber may be of the spray
or venturi type. Particulate collection efficiencies higher than 99 percent
have been achieved with some devices.(14,15) The initial costs are somewhat
higher than those associated with conventional methods, but lower than those of
the foam and force-flux/condensation scrubbers. Operating costs are reported
to be much lower than those of conventional methods.(16,17, 18) It appears
that devices based on electrostatic scrubbing would provide good control over
opacity of stack emissions and would be more effective than the other novel
devices in controlling particulate emissions. Performance of the various types
or electrostatic scrubbers indicates that these units should be tested further
in full scale applications. Table 3 summarizes the various types of electro-
static scrubbers now available, and results of pilot and/or full scale tests on
salt-laden hogged fuel boilers.
Dry Scrubbers-The dry scrubber is a recently developed system that uses a
moving bed of granular material (media) instead of water droplets to capture
particulates. The dirty media is shaken at the bottom of the unit, and parti-
culates fall into a storage bin. The cleaned media is then conveyed back to
the top of the unit. The following advantages are claimed for the unit:(l)
1. It requires no water supply.
2. The particulate is removed dry.
3. Because there is no corrosion potential, mild steel can be used.
4. The unit can be small and lightweight because of high-velocity
throughput.
This device can be effective when temperature and natural collection are
well controlled. Officials of the Simpson Timber plant at Shelton rejected the
dry scrubber in favor of a baghouse because the scrubber would not eliminate
520
-------
TABLE 3. NOVEL FINE PARTICULATE CONTROL DEVICES APPLIED TO
BOILERS BURNING SALT-LADEN HOGGED FUEL3
Manufacturer/
unit
Commercially
available
Base
equipment
Particles
charged
Water
droplets
charged
Results of outlet
emission tests
g/sdm3(gr/sdcf)
Ceil cote Co./
ionizing wet
scrubber
Yes
Wet
scrubber
Yes
No
0.005-0.042 (0.002-0.018)b
<20% opacity
TRW Inc./
charged droplet
scrubber
Yes
Wet
scrubber
No
Yes
Data not available
Pollution Control
>-* Systems, Inc.,/
UW electrostatic
scrubber
Yes
Wet
scrubber
Yes
Yes
0.11 (0.05)
20% opacity
Union Carbide
Bendix Division/
APS electrotube
Yes
Wetted-
wall pipe
precipitator
Yes
No
0.004-0.10 (0.002-0.04)
near 0% opacity
Air Pollution
Systems/electro-
static scrubber
Yes
Venturi
scrubber
Yes
No
Data not available
aAll devices have undergone pilot-scale field tests on boilers burning salt-laden hogged
fuel; information from some of the tests is sparse.
^After two stages of the collector.
-------
the stack plume. A full-scale installation of a salt-laden hogged fuel boiler
at Port Gamble, Washington, was shut down because of scrubber operating prob-
lems. These problems reportedly concerned cake build-up at the discharge of
the moving bed and blinding of the screen that contains the moving bed.
On a more recent full-scale test on a salt installation in Canada, Com-
bustion Power Company (CPC) has solved this problem by providing a steeper
angle for the discharge cone. However, opacity limits are still being exceeded.
BEST AVAILABLE CONTROL TECHNOLOGY FOR CONTROLLING SALT EMISSIONS FROM HOGGED
FUEL BOILERS
Operating data on boilers fired with salt-laden hogged fuel indicate that
the best available conventional control system is a fabric filter preceded by
a mechanical collector. A mechanical collector is needed to remove cinders
that would increase the chance of fire in the fabric filter.
Simpson Timber Company's fabric filter installation at Shelton has opera-
ted successfully for over 2 years in compliance with both grain loading and
opacity regulations. The dramatic decrease in opacity is evidence of the
ability of the fabric filter to capture submicron particles of salt. This
ability to maintain high levels of collection efficiency in the submicrometer
particle size range has been demonstrated in fabric filter applications in
other industries. Some maintenance problems were encountered but plant offi-
cials do not regard them as excessive. Likewise, operation at the Georgia
Pacific facility has reportedly been excellent with low outlet levels, and no
visible emissions.
Use of a venturi scrubber on the salt-laden hogged fuel boiler at Crown-
zellerbach's Port Townsend mill has been partially successful in that opera-
tions comply with the applicable emission regulation when the salt content of
the fuel is below 1 percent. At higher salt levels the scrubber cannot remove
enough of the additional fine salt particles to achieve compliance. More
efficient removal of submicron salt particles would require additional pressure
drop over the present maximum of 51 cm (20 in.) water. Opacity has not been
measured but is estimated by plant officials at 35 percent. Plant engineers
report that the Port Townsend scrubber has not required excessive maintenance
although trouble areas may become apparent with more operating time.
Electrostatic precipitators have not been applied to boilers burning salt-
laden hogged fuel, probably because ESP manufacturers will guarantee outlet
grain loading but not visible opacity. For compliance with visible opacity
regulations when the boiler fuel is salt-laden, an ESP may be too large to be
economically competitive with a fabric filter or wet scrubber. Sizing of the
ESP would have to accommodate the "worst case" in terms of salt content in the
fuel.
In the novel device category the University of Washington (U of W) elec-
trostatic scrubber, the Ceilcote ionizing wet scrubber (IWS), and the APS
522
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electrotube show the best performance characteristics. A recent full scale
test of the U of W scrubber on a salt-laden hogged fuel boilers showed outlet
emission levels of 0.114 g/sdm3 (0.05 gr/scf) and 20 percent opacity. Data
from pilot tests of the Ceilcote IWS and the APS electrotube show outlet
loadings of 0.09 g/dsm3 (0.044 gr/dscf), and lower.
The dry scrubber marketed by Combustion Power Company (CPC), and recently
redesigned with an electrostatic cage installed in the gravel bed of each
scrubber, may be able to reduce opacity levels that have been one of the
shortcomings of the dry scrubber, but further testing of this device is re-
quired.
CONCLUSIONS
Although the data base is small, test operations have demonstrated that
salt-laden particulate emissions from hogged fuel boilers can be controlled to
comply with stringent particulate and opacity regulations, using currently
available control technology. The best available control technology is use of
a fabric filter, preceded by a mechanical collector, to collect cinders and
lessen the chance of fire. Pretreatment techniques such as hydraulic debar-
king, various drying systems or combustion modifications, do not reduce salt
emissions significantly. Bark pressing can reduce moisture content by 50
percent and remove substantial quantities of salt as a result, but the bark
pressate presents a serious water pollution control problem.
Electrostatic scrubbers, categorized as novel control devices, have shown
the potential for effective collection of salt-laden particulate emissions from
hogged fuel boilers, and additional full scale installations of these devices
are likely in the future.
Acknowledgement
This paper is based on work done by PEDCo Environmental, Inc. (19) for
U.S. EPA Region X, under Contract No. 68-01-4147, Task No. 51, Mr. Michael
M. Johnston, Task Manager.
523
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REFERENCES
1. Boubel, R.W. Control of Particulate Emissions from Wood-Fired Boilers.
EPA 340/1-77-026. 1977.
2. MacLean, H., and B.F. MacDonald. Salt Distribution in Sea Water Trans-
ported Logs. State of Alaska, Information Report VP-X-45.
3. Council of the Forest Industries of British Columbia. The Basis for
Requesting a Variance on Marine Salts as Polluting Particulates in Stack
Gases from Hog Fuel Fired Boilers. Submitted to the Pollution Control
Branch, Department of Lands, Forests, and Water Resources, Government of
British Columbia, Victoria, B.C. September 1974.
4. Whyte, J.E. Hog Fuel Boiler Emissions - Particle Size Via Cascade Im-
pactor Tests. Weyerhauser Co., North Bend, Oregon. Technical Report No.
046-4504. August 1976.
5. Leman, M.J. Special Environmental Problems Originated by Burning Bark
from Saltwater-Borne Logs. Proceedings of a Conference on Wood and Bark
Residues for Energy, Oregon State University. February 1975.
6. Horzella, T.E., and H.R. Newton. Controlling Air Pollution from Hogged
Fuel Boilers. Pulp and Paper. February 1974. pp. 71-75.
7. Whitman, J.L., and H. Burkitt. Compliance Alternatives for Stack Emis-
sions from the Hog Fuel Boilers at North Bend, Oregon. Prepared for
Weyerhauser Company, Corporate Engineering. December 1977.
8. Mick, A.H. Wood Waste Fired Boilers: Wet Scrubber Technology. Georgia
Pacific Corporation. Presented at the 69th Annual Meeting of the Air
Pollution Control Association. Portland, Oregon. June 27-July 1, 1976.
9. Cupp, S.J. Operating Experience with a Boiler Firing Salt Water Borne
Hogged Fuel. Crown Zellerbach Corporation. 1978.
10. Szabo, M.F. and R.W. Gerstle. Operation and Maintenance of Particulate
Control Devices in Kraft and Pulp Mill and Crushed Stone Industries.
PEDCo Environmental, Inc. EPA 600/2-78-210. October 1978.
11. CH2M Hill, Inc. A Study of Particulate Emission Discharges from the
Boiler Baghouse Installation. Prepared for Simpson Timber Co., Shelton,
Washington. April 1976.
524
-------
12. Mcllvaine, R.W. Fine Particulate Scrubbing-New Problems and Solutions.
EPA-600/2-77-193. September 1977.
13. University of Washington Electrostatic Droplet Scrubber. In Evaluation
of Eight Novel Fine Particle Collection Devices. EPA-600/2-76-035.
February 1978.
14. Pilat, M.J., G.A. Raemhild, and D.C. Harmon. Fine Particle Control with
UW Electorstatic Scrubber. EPA 600/2-77-193. September 1977. pp. SOS-
SIS.
15. Keller, F.R. Fluidized Bed Combustion Systems for Energy Recovery from
Forest Products Industry Wastes. Presented at Meeting of Forest Products
Research Society, Denver, Colorado. September 1975.
16. Deardorff, D. Wet Wood Waste as a Viable Fuel Supply. Presented at
meeting of Forest Products Research Society, Denver, Colorado. September
1975.
17. Jasper, M., and P. Koch. Suspension Burning of Green Bark to Direct-
Fired High-Temperature Kilns for Southern Pine Lumber. Presented at
meeting of Forest Products Research Society, Denver, Colorado. September
1975.
18. Guidon, M.W. Pilot Studies for Particulate Control of Hog Fuel Boilers
Fired with Salt-Water Stored Logs. Georgia Pacific Corp., Bellingham,
Washington. November 1977.
19. Szabo, M.F., S. Kothari, C. Doolittle, and R. Gerstle. Control of Salt-
Laden Particulate Emissions From Hogged Fuel Boilers. PEDCo Environmental,
Inc. Final Report. U.S. EPA contract No. 68-01-4147, Task No. 51. May 1979.
525
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AUTHOR INDEX
AUTHOR NAME
PAGE
Ariman, T.
111-222
Bacchetti, J. A.
1-529
Bernstein, S.
11-125
Bibbo, P. P.
11-219
Bickelhaupt, R. E.
1-154
Blackwood, T. R.
IV-312
Bloomfield, D. P.
111-14-5
Brackbill, E. A.
I11-472
Brines, H. G.
1-351
Brookman, E. T.
IV-274
Brown, J. T. (Jr.)
II1-439
Buchanan, W. J.
11-168
Burckle, J. 0.
II1-484
Bush, J. R.
IV-154
Carlsson, B.
II1-260
Carr, R. C.
1-35, III-270
Chang, C. M.
11-314
Chapman, R. A.
1-1
Chmielewski, R.
III-l
Cooper, D. W.
I11-127
Cowen, S. J.
IV-424
Cowherd, C. (Jr.)
IV-240
Czuchra, P. A.
III-104
Darby, K.
1-15
Daugherty, D. P.
IV-182
526
-------
AUTHOR NAME
PAGE
Dennis, R.
1-494
Dietz, P. W.
II1-429
Donovan, R. P.
1-476
Drehmel, D. C.
IV-170
Durham, M. D.
IV-368
Dybdahl, A. W.
IV-443
Ellenbecker, M.
J.
II1-171, II1-190
Engelbrecht, H.
L
11-279
Ensor, D.S.
111-39
Ernst, M.
IV-30, IV-42
Eschbach, E. J.
11-114
Evans, J. S.
IV-252
Fasiska, E. J.
IV-486
Faulkner, M. G.
IV-508
Fedarko, W.
IV-64
Ferrigan, J. J.
1-170
Finney, W. C.
11-391
Furlong, D. A.
1-425
Garrett, N. E.
IV-524
Gastler, J. H.
IV-291
Gavin, J. H.
II1-81
Giles, W. B.
IV-387
Gooch, J. P.
1-132
Gooding, C. H.
111-404
Grace, D. S.
II1-289
Guiffre, J. T.
1-80
527
-------
AUTHOR NAME
PAGE
Hall, F. D.
III~25
Hardison, L. C.
II1-382
Hoenig, S. A.
IV-201
Hudson, J. A.
1-263
Iinoya, K.
111-237
Isoda, T.
II1-16
Jaasund, S. A.
11-452
Kalinowski, T. W.
111-363
Kallio, G. A.
II1-344
Kearns, M, T.
II1-61
Kelly, D. S.
I-100
Kinsey, J. S.
II1-95
Kolber, A. R.
1-224
Ladd, K. L.
1-317
Lamb, G. E.R.
II1-209
Lane, W. R.
1-410
Langan, W. T.
1-117, 11-256
Larson, R. C.
II1-448
Leonard, G.
11-146
Lipscomb, W. 0.
1-453
Malani, S.
1-570
Marcotte, W. R.
1-372
Martin, J. R.
1-591
Masuda, S.
11-65, 11-334, 11-483
McCain, J. D.
IV-496
McDonald, J. R.
11-93
528
-------
AUTHOR NAME
PAGE
Mitchell, D. A.
III-162
Modi a, J. C.
11-399
Mosley, R. B.
11-45
Mycock, J. C.
1-432
Neundorfer, M.
11-189
Nixon, D.
1-513
Noll, C. G.
11-374
Nunn, M.
11-369
Ondov, J. M.
IV-454
Ostop, R. L.
1-342
Parker, R.
IV-1
Patch, R. W.
IV-136
Patterson, R. G.
IV-84
Pearson, G. L.
1-359
Pedersen, G. C.
I11-416
Petersen, H. H.
11-352
Pilat, M. J.
1-561
Potter, E. C.
1-184
Ranade, M. B.
1-538
Raymond, R. K.
11-173
Rinard, G.
11-31, IV-127
Roehr, J. D.
11-208
Rolschau, D. W.
II1-251
Ruth, D.
11-427, 11-441
Samuel, E. A.
II-l
Schliesser, S. P.
1-56
529
-------
AUTHOR NAME
PAGE
Self, S. A.
II1-309
Severance, R. L.
IV-321
Shale, C. C.
1-390
Smit, W.
1-297
Smith, S. B.
11-502
Spafford, R. B.
1-202
Sparks, L. E.
11-417, IV-411
Stenby, E. W.
1-243
Stock, W. E.
IV-333
Surati, H.
11-469
Szabo, M. F.
I11-508
Tendulkar, S. P.
IV-338
Tennyson, R. P.
111-117
Tsao, K. C.
IV-14
Umberger, J. H.
11-296
VanOsdell, D. W.
11-74
VanValkenburg, E. S.
IV-351
Wang, J. C. F.
IV-396
Weber, E.
IV-98
Wybenga, F. A.
11-242
Yung, S.
IV-217
530
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TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completingJ
1. REPORT NO. 2.
EPA-600/9-80-039c IERL-RTP-1063
3. RECIPIENT'S ACCESSION NO.
4. TITLE AND SUBTITLE
Second Symposium on the Transfer and Utilization of
Particulate Control Technology (Denver, July 1979)
Vol. III. Particulate Control Devices
5. REPORT DATE
Sept. 1980 Issuing Date.
6. PERFORMING ORGANIZATION CODE
7. AUTHOR(S)
F.P. Venditti, J.A. Armstrong, and Michael Durham
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORGANIZATION NAME AND ADDRESS
Denver Research Institute
P.O. Box 10127
Denver Colorado 80210
10. PROGRAM ELEMENT NO.
EHE624
11. CONTRACT/GHANT NO.
R805725
12. SPONSORING AGENCY NAME AND ADDRESS
Industrial Environmental Research Laboratory
Office of Research and Development
U.S. Environmental Protection Agency
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Proceedings; 6/79-6/80
14. SPONSORING AGENCY CODE
EPA/600/13
IS. SUPPLEMENTARY NOTES
IERL-RTP project officer is Dennis C. Drehmel, MD-61, 919/541-2925. EPA-600/7-79-044a
thru -044d are proceedings of the 1978 symposium.
16. ABSTRACT
The proceedings document the approximately 120 presentations at the EPA/IERL-RTP-
sponsored symposium, attended by nearly 800 representatives of a wide
variety of companies (including 17 utilities). The keynote speech for the 4-day meeting
was by EPA's Frank Princiotta. The meeting included a plenary session on enforcement.
Attendees were polled to determine interest areas: most (488) were interested in
operation and maintenance, but electrostatic precipitators (ESPs) and fabric
filters were a close second (422 and 418, respectively). Particulate scrubber interest
appears to be waning (288). Major activities of attendees were: users, 158; manufac-
turers, 184; and R and D, 182. Technical presentations drawing great interest were the
application of ESPs and baghouses to power plants and the development of novel ESPs.
As important alternatives to ESPs, baghouses were shown to have had general success in
controlling coal-fired power plant emissions. When operating properly, baghouses can
limit emissions to^5 mg/cu nm at pressure drops of ^2 kPa. Not all baghouse install-
ations have been completely successful. Both high pressure drop and bag loss have
occurred (at Harrington Station), but these problems appear to be solved.
17. KEY WORDS AND DOCUMENT ANALYSIS
a. DESCRIPTORS
b. 1DENTIFIERS/OPEN ENDED TERMS
c. COSATI Field/Croup
Pollution Scrubbers
Dust Flue Gases
Aerosols
Electrostatic Precipitators
Filters
Fabrics
Pollution Control
Stationary Sources
Particulate
Baghouses
13B 07A
11G 21B
07D
131
14G
HE
18. DISTRIBUTION STATEMENT
Reld^e to Public
19. SECURITY CLASS (This Report)
Unclassified
21. NO. OF PAGES
547
20. SECURITY CLASS fThis page)
Unclassified
22. PRICE
EPA Fo»m 2220-1 (R«v. 4-77) PREVIOUS EDITION IS OBSOLETE ~ U. S. GOVERNMENT PRINTING OFFICE : 1980--657-lG5/Ol59
531
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