EPA-625/7-76-001
                     ENVIRONMENTAL POLLUTION CONTROL
                          PULP AND PAPER INDUSTRY
                                  PARTI
                                   AIR
                  U.S. ENVIRONMENTAL PROTECTION AGENCY
                             Technology Transfer
                               October 1976
                                           ""•*"

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                              ACKNOWLEDGMENTS

This process design manual was prepared for the Office of Technology Transfer of the U.S.
Environmental Protection Agency. Coordination and preparation of Part I of this manual,
Air Pollution Control, was carried out by EKONO, Inc., Bellevue, Washington, and EKONO,
Oy, Helsinki, Finland.

. .  . U.S. EPA reviewers were James  C.  Herlihy of the U.S. EPA Division of Stationary
Source Enforcement, Washington, D.C., James A. Eddinger of the U.S. EPA Emission Stand-
ards and Engineering Division, Research Triangle Park, N.C., Gene Tucker of the U.S. EPA
Industrial  Environmental Research Laboratory, Research Triangle Park, N.C., and George
S.  Thompson, Jr., of the U.S. EPA Office of Technology Transfer, Cincinnati, Ohio.
                                     NOTICE

The  mention of trade names of commercial products in this publication is for illustration
purposes and  does  not  constitute endorsement or recommendation for  use  by the  U.S.
Environmental Protection Agency.  This manual is presented as a helpful guide to the user
and should in no way be  construed as a regulatory document.

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                                PART 1:  AIR

                                 CONTENTS

Chapter                                                                   Page

         ACKNOWLEDGMENTS                                            ii

         CONTENTS                                                      iii

         LIST OF  FIGURES                                               ix

         LIST OF  TABLES                                                xv

         FOREWORD                                                     xix

   1     INTRODUCTION                                                 1-1

         1.1    Gaseous and Particulate Emissions from Kraft Pulp and Paper Mill
                 Process Sources                                           1-2
         1.2    Gaseous and Particulate Emissions from Sulfite Pulp  and Paper
                 Mill Process Sources                                       1-9
         1.3    Power Boilers                                               1-10
         1.4    References                                                 1-13

   2     DIGESTER  GASES                                               2-1

         2.1    Batch Digesters                                             2-1
         2.2    Continuous Digester Gases                                    2-11
         2.3    References                                                 2-14

   3     EVAPORATION GASES                                           3-1

         3.1    Black Liquor Properties                                      3-1
         3.2    Evaporator Types                                           3-2
         3.3    Evaporator Gas Scrubbing                                    3-8
         3.4    General Evaporator Air Pollution Abatement Programs            3-11
         3.5    In-plant Controls                                            3-13
         3.6    References                                                 3-14
                                      in

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                            CONTENTS - Continued

Chapter                                                                  Page

   4     NONCONDENSABLE GAS TREATMENT                           4-1

         4.1    Gas Stream Characteristics                                   4-1
         4.2    Gas Handling Systems                                       4-6
         4.3    Gas Treatment Systems                                      4-14
         4.4    System Economics                                         4-19
         4.5    References                                                4-20

   5     CONDENSATE TREATMENT                                     5-1

         5.1    Condensate Components                                    5-1
         5.2    Digester Condensates                                       5-2
         5.3    Evaporator Condensates                                     5-4
         5.4    Condensate Chlorination                                    5-9
         5.5    Condensate Stripping                                       5-9
         5.6    References                                                5-20

   6     BROWN STOCK WASHER GASES                                6-1

         6.1    Displacement Washing                                      6-1
         6.2    Diffusion Washers                                          6-2
         6.3    References                                                6-5

   7     STORAGE TANK VENT GASES                                  7-1

         7.1    Storage Tank Vent Gas Composition                          7-1
         7.2    Storage Tank Vent Gas Treatment                            7-1
         7.3    References                                                7-1

   8     TALL OIL RECOVERY  GASES                                  8-1

         8.1    Batch TaU Oil Recovery                                     8-1
         8.2    Continuous Tall Oil Recovery                                8-2
         8.3    References                                                8-3
                                       IV

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                             CONTENTS - Continued

Chapter                                                                   Page

   9     BLACK  LIQUOR  OXIDATION                                     9-1

         9.1    Weak Black Liquor Oxidation—Air                             9-1
         9.2    Strong Black Liquor Oxidation—Air                            9-11
         9.3    Agitated Air Sparging                                        9-16
         9.4    Combination Systems                                        9-16
         9.5    Molecular Oxygen Systems                                    9-18
         9.6    Process Effects                                              9-27
         9.7    Air Pollution Effects                                         9-30
         9.8    Oxidation Tower Emissions                                   9-33
         9.9    Process Economics                                          9-34
         9.10  References                                                  9-43

  10     RECOVERY BOILER DESIGN AND  OPERATION                   10-1

         10.1  General Conditions                                          10-1
         10.2  Combustion of Black Liquor Dry Solids                         10-15
         10.3  Different Recovery Boiler Designs                              10-28
         10.4  Process Variables                                            10-40
         10.5  Diverse Obnoxious Compounds                                10-51
         10.6  Direct Contact Evaporation                                   10-51
         10.7  Flue Gas Scrubbing for Gaseous Emissions                       10-56
         10.8  Collection of Particulate Matter from Recovery Boiler Flue Gas    10-61
         10.9  Economy of Recovery Boiler Operation                         10-74
         10.10 References                                                  10-77
               Appendix 10-1                                              10-81
               Appendix 10-2                                              10-82

  11     LIME BURNING AND  LIME  DUST HANDLING                    11-1

         11.1   Rotary Lime Kilns                                           11-1
         11.2   Fluidized Bed  Calciners                                       11-2
         11.3   Particulate Emission Control                                  11-3
         11.4   Gaseous Emission Control                                     11-6
         11.5   Oxygen Addition                                            11-9
         11.6   Process Economics                                           11-9
         11.7   Lime Dust Handling                                          11-10
         11.8   References                                                  11-11

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                            CONTENTS - Continued

Chapter                                                                  Page

  12     SMELT DISSOLVING TANK                                      12-1

         12.1   Smelt Dissolving Tank Particulate Matter Emissions              12-1
         12.2   Smelt Dissolving Tank TRS Emissions                         12-2
         12.3   References                                                12-4

  13     EMISSIONS OF OXIDES OF NITROGEN, HYDROCARBONS, AND
           WATER VAPOR                                              13-1

         13.1   Nitrogen Oxides                                           13-1
         13.2   Water Vapor                                              13-4
         13.3   Organic  Compounds                                        13-6
         13.4   References                                                13-6

  14     AIR  POLLUTION  CONTROL IN SULFITE PULP MILLS            14-1

         14.1   Sulfite Pulping Processes                                    14-1
         14.2   Digester Gases                                             14-4
         14.3   Washer Gases                                              14-7
         14.4   Evaporator Gases                                          14-7
         14.5   Combustion Gases                                         14-10
         14.6   Acid Preparation Gases                                      14-12
         14.7   SSL Recovery Boilers                                       14-16
         14.8   SSL Recovery Systems                                      14-20
         14.9   Problem of Nitrogen Compounds for Ammonium-Based Pulping   14-27
         14.10 References                                                14-30

   15     OTHER  PROCESS SOURCES                                     15-1

         15.1   Bleach Plant Gases                                         15-1
         15.2   Wastewater Treatment                                      15-3
         15.3   Odor Problems from Diffuse Sources                          15-6
         15.4   References                                                15-8

   16    POWER BOILERS                                               16-1

         16.1   Supply Patterns                                           16-1
         16.2   Combustion Parameters                                     16-2
                                       VI

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                            CONTENTS - Continued

Chapter                                                                Page

         16.3   Boiler Types                                              16-6
         16.4   Particulate Emissions                                       16-14
         J6.5   References                                               16-20

  17     PROCESS MONITORING                                        17-1

         17.1   Source Measurements                                       17-1
         17.2   Gaseous Monitoring                                        17-1
         17.3   Particulate Monitoring                                      17-24
         17.4   Odor Measurements                                        17-31
         17.5   Mobile Laboratories                                        17-35
         17.6   Economics                                               17-39
         17.7   References                                               17-41

APPENDIX A - GLOSSARY OF SYMBOLS                                  A-l

APPENDIX  B - CHEMICAL FORMULAS                                   B-l
                                     vn

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                                LIST OF FIGURES

Figure No.                                                                 Page

  2-1      Batch Digester Flow Sheet                                         2-2
  2-2      Kraft Batch Digester Blow Steam Flow                              2-3
  2-3      Kraft Batch Digester Blow Gas Flow After Condensing and Without
              Equalization                                                   2-4
  2-4      Vapor Pressure of 0.01 M H2S vs. pH at Various Temperatures         2-5
  2-5      Vapor Pressure of 0.0IN CH3SH vs. pH at Various Temperatures       2-6
  2-6      Blow Heat Recovery Control System                                2-7
  2-7      Odor Compounds in Relief  Gas After Turpentine Condenser as a
              Function of Condenser Outlet Temperature                        2-8
  2-8      Vapor Pressures of Methanol, Water, a-Pinene & j3-Pinene              2-12
  2-9      Continuous Digester Flow Sheet                                    2-13

  3-1      Vapor-Liquor Equilibrium for H2S Over Black Liquor                 3-2
  3-2      Multi-Effect Vacuum Evaporation Plant Flow Sheet                   3-3
  3-3      Evaporation Plant Direct Contact Condenser With Water Ring Pump    3-5
  3-4      Evaporation Plant Two Stage Surface  Condenser  With Water Ring
              Pump                                                         3-7
  3-5      Multiple Effect Back Pressure Evaporation Plant Flow Sheet           3-9
  3-6      Flash Evaporation Plant Flow Sheet                                 3-10
  3-7      Multiple Effect  Single Stage  Thermocompressor  Evaporation Plant
              Flowsheet                                                     3-11
  3-8      Hotwell Gas Scrubber for 100 Metric Tons Per Hour (40 gpm) Evapo-
              ration Plant for H2 S—Separation of 95% or More                   3-12

  4-1      Vaporsphere Flow Equalization Gas Holders                         4-7
  4-2      Floating Cover Flow Equalization Gas Holders                        4-8
  4-3      Packed  Bed Scrubber For Noncondensable Gas Handling System       4-11
  4-4      Liquid Condensate Trap for Noncondensable Gas Handling System     4-13
  4-5      Safety Devices for Noncondensable Gas Handling Systems             4-14
  4-6      Unsteady State Flow System  for Batch Digester Noncondensable Gas
              Incineration                                                   4-15
  4-7      Steady  State Flow System for Continuous Digester Noncondensable
              Gas Incineration                                               4-16
  4-8      Noncondensable Gas Incineration System                            4-17

  5-1      Evaporation Plant Surface  Condenser With Water Jet Condenser and
              Water Ring Pump                                               5-5
                                         IX

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                          LIST OF FIGURES - Continued

Figure No.                                                                    Page

  5-2      Evaporation  Plant Surface Condenser With Recirculated Water  Jet
              Condenser and Water Ring Pump                                  5-6
  5-3      Contaminated Condensates Air Stripping Plant Flow Sheet             5-11
  5-4      Contaminated Condensates Steam Stripping Plant Flow Sheet           5-13
  5-5      Stripping Efficiency  for Different  Steam-Condensate Ratios With 10
              Theoretical Plates                                                5-15
  5-6      Condensate Stripping in an Evaporation Plant                         5-16
  5-7      Simplified Flow Sheet for Kraft Process With Condensate Segregation,
              Stripping, & Reuse                                              5-19

  6-1      Vacuum Washers Flow Sheet                                        6-2
  6-2      Pressure Washers Flow Sheet                                        6-3
  6-3      Batch Diffusion Washers Flow Sheet                                 6-4
  6-4      Continuous Diffusion Washers Flow Sheet                            6-5

  8-1      Batch Tall Oil Plant Flow Sheet                                     8-1
  8-2      Continuous Tall Oil Plant Flow Sheet                                8-3

  9-1      Collins Porous Plate Diffuser Weak Black Liquor Oxidation System      9-2
  9-2      Trobeck-Ahlen  Multiple  Sieve Tray Weak Black  Liquor Oxidation
              System                                                         9-5
  9-3      Packed  Tower Systems for Weak Black Liquor Oxidation               9-8
  9-4      Agitated Air Sparging System for Black Liquor Oxidation              9-10
  9-5      Champion Unagitated  Air Sparge  Strong  Black  Liquor Oxidation
              System                                                         9-13
  9-6      Operating  &  Performance Data for Single Stage Strong Black Liquor
              Oxidation  System                                               9-14
  9-7      Champion Two Stage Unagitated  Strong  Black  Liquor Oxidation
              System                                                         9-16
  9-8      Western Kraft Pipeline Reactor Strong Black Liquor Oxidation System   9-17
  9-9      Two Stage Combination Weak & Strong Black Liquor Oxidation With
              Oxygen                                                        9-20
  9-10     Owens-Illinois System for  Two Stage Weak  & Strong Black Liquor
              Oxidation With Oxygen                                          9-23
  9-11     Effect of Production Rate on Capital & Operating Costs for Weak &
              Strong Black Liquor Oxidation With Oxygen                       9-36
  9-12     Operating Costs for Weak Black Liquor Oxidation With Oxygen         9-39

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                         LIST OF FIGURES - Continued

Figure No.                                                                   Page

  10-1      Heat Values vs. Oxygen Demand for Complete Combustion of Lignin
             and Carbohydrates                                              10-4
  10-2      Heat Values vs. Oxygen Demand for Complete Combustion for Some
             North American Dry Solids                                       10-5
  10-3      Flow  Diagram  for  Black Liquor  Through  Recovery Boiler-North
             American System                                                10-8
  10-4      Flow Diagram for Black Liquor Through Fluidized Bed Reactor With
             Waste Heat Recovery Boiler                                      10-9
  10-5      Flow  Diagram for Black Liquor Through Recovery  Boiler Scandi-
             navian System—Low Odor System                                 10-11
  10-6      Principle Design of Air Distribution to Recovery Boiler Furnaces        10-15
                                             o
  10-7      Air Supply According to Gotaverken Angteknik AB Design             10-17
  10-8      Modern Kraft Recovery Unit from Babcock & Willcox                 10-18
  10-9      Modern Kraft Recovery Unit from Combustion Engineering            10-19
  10-10    Equilibrium  Diagram for Condensed Phases and  Gases for a Sodium
             Based Black Liquor                                              10-21
  10-11    Distribution  of  Sodium  and  Sulfur in a Black  Liquor Recovery
             Furnace as a Function of Temperature                             10-23
  10-12    Equilibrium  Diagram for a Na2CO3 -Na2 S System                      10-25
  10-13    Effect of NaCl Addition on the  Melting Point of a Synthetic Pulp        10-26
  10-14    Babcock & Wilcox Recovery Boiler With Cyclone Evaporator            10-30
  10-15    Cyclone Evaporator                                                10-31
  10-16    Combustion Engineering Recovery Boiler With Cascade Evaporator      10-32
  10-17    Open View of Cascade Evaporator                                   10-33
  10-18    Combustion Engineering  Recovery  Boiler With  Cascade  Evaporator
             Ace System                                                    10-34
  10-19    Combustion Engineering Recovery Boiler With Laminaire Air Heater
             & Complete Multiple Effect Evaporation L.A.H. System              10-35
  10-20    Recirculation Air Heater for Scandinavian Recovery Boiler              10-37
  10-21    Typical Gotaverken Angteknik Recovery  Boiler                       10-38
  10-22    Heat Value vs. Oxygen Demand at Complete Combustion              10-45
  10-23    Viscosity of  Black Liquors                                         10-47
  10-24    Adjustable Air Ports                                               10-50
  10-25    British Columbia Research Council Design for H2S Absorption Scrub-
             ber
  10-26    Glitsch-Weyerhaeuser Desigi
  10-27    Formation of Visible Plume
                                        10-58
n for a TRS Scrubbing System             10-59
 Through Condensation of Water Vapor     10-60
  10-28    Electrostatic Precipitator Size as a Function of Collecting Efficiency   10-67
                                        XI

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                          LIST  OF  FIGURES - Continued

Figure No.                                                                    Page

  10-29    Capital Cost for Recovery Boiler Department                         10-75
  10-30    Flue Gas Energy Losses                                             10-76

  14-1     Principal Flow Diagram for Spent Liquor Collection in Blow Pits, 2
              Stage Displacement Wash                                         14-6
  14-2     Principal Flow Diagram  for Spent Liquor Collection in Rotary Washer
              Plant (3 2-Zone Filters)                                           14-8
  14-3     Principal  Flow Diagram for  Spent Liquor Evaporation in  Multiple
              Effect Vacuum Plant, 5  Effect, 7 Bodies                           14-9
  14-4     Flow Sheet for Calcium Base Raw Acid Preparation                   14-13
  14-5     Acid Bisulfite Fortification System Flow Sheet                       14-14
  14-6     Flow Sheet for Magnesium Base Raw Acid Preparation                 14-15
  14-7     Babcock  &  Wilcox  Process for  Magnesia Base  Sulfite  Chemicals
              Recovery                                                       14-21
  14-8     STORA Process for Sodium Base Sulfite Chemicals Recovery           14-23
  14-9     SCA Process  for Sodium Base Sulfite Chemicals Recovery              14-25
  14-10    Tampella Process for Sodium Base Sulfite Chemicals Recovery         14-26

  15-1     Equilibrium Solubility of Chlorine Dioxoide in Water                  15-3
  15-2     Equilibrium Solubility of Sulfur Dioxide in Water                     15-4

  17-1     Gas Sample Handling & Conditioning System for Externally Located
              Continuous Gaseous Monitoring System                            17-3
  17-2     Continuous Source Monitoring System for Reduced Sulfur Emissions
              With Coulometric Titration                                       17-9
  17-3     Total Reduced Sulfur Monitoring With an Electrochemical Membrane
              Cell Detector                                                   17-10
  17-4     Continuous Conductivity  Monitor for Measuring Sulfur Oxide Emis-
              sions from Sulfite Mill Sources                                    17-12
  17-5     Electrochemical Transducer  Membrane Cell for Continuous Sulfur
              Dioxide Monitoring                                             17-13
  17-6     Internally Located  Ultraviolet Spectrometer for Sulfur Dioxide Moni-
              toring in Flue Gas Streams                                        17-14
  17-7     Internally Located  Ultraviolet Spectrophotometer for Sulfur Dioxide
              Monitoring in Flue Gas Streams                                   17-15
  17-8     Rotating Syringe Instrument Calibration Procedure                   17-22
  17-9     Permeation Tube Instrument Calibration Procedure                   17-23
  17-10    ASME Batch Particulate Sampling Train                             17-26
                                         XII

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                         LIST OF FIGURES - Continued

Figure No.                                                                   Page

  17-11    EPA Batch Particulate Sampling Train                               17-26
  17-12    Multistage Cascade Impactor for Particle Size Distribution Determina-
              tion                                                           17-28
  17-13    Membrane Filter System for Particle Size Distribution Determination    17-28
  17-14    Continuous  Monitoring of Particulate Emissions With a Conductivity
              Cell Detector                                                   17-29
  17-15    Continuous  Monitoring of Particulate  Emissions With a Sodium Ion-
              Specific Electrode                                              17-30
  17-16    Swedish Dynamic Dilution System for Odor Level Evaluation           17-35
  17-17    Scentometer Dilution System for Odor Threshold Evaluation           17-36
  17-18    Gaseous Sampling System for  Sulfur Gas Analysis in NCASI Mobile
              Laboratory                                                     17-37
  17-19    Particulate and Gaseous Sample Handling Systems for ITT-Rayonier
              Mobile Laboratory                                              17-39
                                        xin

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                                LIST OF TABLES

Table No.                                                                   Page
                                                                              D

   1-1     External Control Techniques for Gaseous and Particulate Matter Emis-
              sions From Kraft Pulp Mill Sources                               1-3
   1-2     Typical Gas Characteristics for Kraft Pulp Mill Processes               1-4
   1-3     Typical Reduced  Sulfur Gas Concentrations from  Kraft Pulp  Mill
              Sources                                                        1-5
   1-4     Typical Reduced Sulfur  Gas Emission Rates from  Kraft Pulp  Mill
              Sources                                                        1-6
   1-5     Typical Concentrations and Rates  for SOX and NOX Emission from
              Kraft Pulp Mill Combustion Sources                              1-7
   1-6     Typical Concentrations  and Emission  Rates  for Particulate Matter
              from Kraft Pulp  Mill Sources (After Control Devices)               1-9
   1-7     Typical SO2 Emission Rates from Sulfite Pulp Mill Sources            1-10
   1-8     Uncontrolled Air  Pollutant Emissions  from  Fuel  Combustion in
              Auxiliary Power Boilers                                         1-12

   3-1      Effect  of Condenser  Type on Reduced Sulfur Gas  Emission from
              Evaporator Noncondensable Gases                                3-14

   4-1      Gas Flow Rates from a Batch Digester                               4.3
   4-2     Typical Ranges in Digester and Evaporator Noncondensable Gas Flow
              Rates                                                          4.3
   4-3      Flammability Limits in Air for Compounds Present in Kraft Noncon-
              densable Gas Stream                                            4-5
   4-4     Dilution Requirements With Air to  Avoid Explosions for  Digester
              Noncondensable Gas Streams                                     4-5
   4-5      Flame Propagation  Speeds for Air-Mercaptan Mixtures                 4-6
   4-6      Dimensions of  Flow Equalization Devices in  Kraft Noncondensable
              Gas Handling Systems                                            4-9
   4-7      Batch Digester  Blow Gas Flow Rates fpr Sizing Noncondensable  Gas
              Flow Equalization Devices                                       4-10
   4-8      Piping Systems for Kraft Noncondensable Gas Handling Systems        4-12
   4-9      Gas Flow Rates to Burning Devices from Noncondensable Gas Han-
              dling Systems                                                  4-16
   4-10    Capital  Costs   for  Installation of  Noncondensable  Gas Handling
              Systems                          -                              4_20

   5-1      Main Components of Typical Kraft Mill Condensates                   5-2
   5-2      Typical Kraft Mill Condensate Compositions, Mean Values for  10 Mills  5-3
                                       xv

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                          LIST OF TABLES - Continued

Table No.                                                                    Page

   5-3     Typical Kraft Mill Condensate Characteristics for 17 Mills              5-4
   5-4     Calculated Evaporator Condensate Flow, Sulfur, Methanol, and BOD
              Distribution for Liquor Sequence 3-4-5-2-1                         5-7
   5-5     Chlorine  Demand of Reduced Sulfur  Compounds for Oxidation to
              Sulfur                                                         5-9

   6-1     Flows, Composition and Sulfur  Release of Vacuum Washer Foam
              Tank Vent Gases                                                6-3

   8-1     Bath Tall Oil Recovery Plant TRS Components and Typical Concen-
              trations                                                        8-2
   8-2     Tall Oil  Recovery Noncondensable  Gas Flows and Reduced Sulfur
              Emissions                                                      8-2

   9-1     Design and  Operating  Parameters for Porous Plate  Diffuser Black
              Liquor Oxidation Systems                                       9-3
   9-2     Design and  Operating Parameters for  Trobeck-Ahlen Multiple Sieve
              Tray Weak Black Liquor Oxidation Units                          9-6
   9-3     Design and Operating Parameters for Weyerhaeuser Concurrent Flow
              Packed Tower Weak Black Liquor Oxidation System                9-9
   9-4     Operating and Performance  Data for Agitated Air Sparged Weak Black
              Liquid Oxidation Systems                                       9-11
   9-5     Design Criteria of Plug Flow Reactor Systems for Black Liquor Oxida-
              tion With Molecular Oxygen                                      9-19
   9-6     Oxygen Requirements for Weak Black Liquor Oxidation               9-21
   9-7     Effect of Weak Black Liquor Oxidation on Liquid Chemical Composi-
              tion                                                           9-24
   9-8     Effect of Weak Black Liquor Oxidation on Sulfur Gas Emissions from
              Kraft MiU Process Sources                                       9-25
   9-9     Effect of Black  Liquor Oxidation  on Sulfur Gas Emissions During
              Direct Contact Evaporation                                      9-31
   9-10    Effect of Weak Black Liquor Oxidation on Malodorous Sulfur Gas
              Emissions from Evaporator Noncondensable Gases                  9-32
   9-11    Reduced Sulfur Emissions from Black  Liquor Oxidation Tower Vents
              Using Air                                                      9-34
   9-12    Estimated Capital Costs for  Black Liquor Oxidation Systems           9-35
   9-13    Approximate Annual Operating Costs for Black Liquor Oxidation Sys-
              tems Using Air                                                 9-37
                                       xvi

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                          LIST OF TABLES - Continued

Table No.                                                                    Page

   9-14    Effect of Number of Stages on Annual Operating Costs for  Strong
              Black Liquor Oxidation With Air                                  9-37
   9-15    Estimated  Capital Costs for Black  Liquor Oxidation Systems Using
              Molecular Oxygen                                               9-40
   9-16    Operating Costs and Operating  Credits for  Black Liquor Oxidation
              With Oxygen                                                    9-41
   9-17    Typical Ranges in Operating Variables, Reduced Sulfur Emissions and
              Cost Factors for Black Liquor Oxidation Systems                   9-42

  10-1     Black Liquor Dry Solids Content                                    10-2
  10-2     Black  Liquor Combustion Products in Caloric Bomb Using Modified
              Technique                                                      10-6
  10-3     Black Liquor Combustion Products in Recovery Furnace               10-6
  10-4     Operating  Conditions for American and Scandinavian  Black  Liquor
              Concentration                                                   10-12
  10-5     Capital Cost  Comparison of American and Scandinavian Black Liquor
              Concentration Practices                                          10-13
  10-6     Annual Operating Cost  Comparison of American and Scandinavian
              Black Liquor Concentration Practices                              10-13
  10-7     Recovery Furnace Exhaust Gas Properties for Direct and Indirect Con-
              tact Evaporation Systems                                         10-52
  10-8     Effect of Black  Liquor Oxidation on Sulfur Gas Emissions  During
              Direct Contact Evaporation                                       10-54
  10-9     Effect of Black Liquor pH on H2S Emissions During Direct Contact
              Evaporation                                                    10-55
  10-10    Average  Particulate Emissions from  Recovery  Boiler Electrostatic
              Precipitators in the United States                                 10-71

  11-1     Energy Requirements for Lime Mud Calcining Systems                 11-2
  11-2     Operating Characteristics for Particulate Liquid Scrubbers Employed
              on Kraft Lime Kilns                                             11-4
  11-3     Particulate Collection Efficiencies for  Liquid Scrubbers on Kraft Pulp
              Mill Lime Kilns                                                 11-5
  11-4     Gaseous Emissions from Kraft  Pulp Mill Lime Kilns                    11-6
  11-5     Capital and Operating Costs for Lime Kiln Particulate Scrubbers        11-10

  12-1     Smelt Dissolving Tank Particulate Matter Control Devices              12-2
  12-2     TRS Emissions from Smelt Dissolving Tanks                          12-3
                                        xvn

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                           LIST OF TABLES - Continued

Table No.                                                                     Page

  13-1      Effect of Flame Temperature on Nitric Oxide Equilibrium Concentra-
              tion and Reaction Time                                           13-2
  13-2      Nitrogen  Oxide  Emissions from  Kraft Pulp  Mill  Process  Sources   13-3
  13-3      Nitrogen Oxide Emissions from Power Boilers                         13-4
  13-4      Water Vapor Emissions to the Atmosphere from  Kraft Pulp Mill
              Sources                                                         13-5

  14-1      Main Sulfite Pulping Processes                                       14-2
  14-2      Sulfite Process Chemicals, Price, Combustion and Recovery             14-2
  14-3      Scandinavian Sulfite Pulp Mill Emissions                              14-3
  14-4      Typical Potential American Sulfite Pulp Mill Emissions                 14-4
  14-5      Nitrogen  Oxides Emissions from an Ammonium Base Sulfite Recovery
              Furnace                                                         14-29

  15-1      Odorous Gases from Diffuse Kraft  Pulp Mill Sources                   15-7

  16-1      Overall National Distribution  of  Energy Sources for the Pulp and
              Paper Industry                                                   16-1
  16-2      Characteristics of Fuel  Burned in Power  Boilers at Pulp and Paper
              Mills in the United States                                         16-3
  16-3      Uncontrolled Air Pollutant Emission Factors for Fuel Combustion in
              Power Boilers for the Pulp and Paper Industry                      16-4
  16-4      Particulate  Emission Characteristics from Selected U.S. Power Boilers   16-15
  16-5      Typical  Particle Size Distribution of Fly Ash from  Coal- and Wood-
              Fired Power Boilers                                              16-16

  17-1      Selective Prescrubbing Solutions for Sulfur Gas Separation             17-6
  17-2      Approximate Ranges in Calibration  Factors  for Sulfur Gases  With
              Coulometric Titrator                                             17-8
  17-3      Operating Characteristics of Gas Chromatographic Detectors            17-21
  17-4      Odor Threshold Levels for Malodorous Sulfur Compounds             17-32
  17-5      Odor Intensity Level Evaluation Scale                                17-33
  17-6      Approximate Capital Costs for Continuous Gaseous  Stack Monitoring
              Instrumentation                                                 17-40
  17-7      Approximate Capital Costs for Continuous Particulate Stack Monitor-
              ing Instrumentation                                              17-40
                                         xvm

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                                    FOREWORD

The formation of the United States Environmental Protection Agency marks a new era of
environmental awareness in America. The Agency's goals are national in scope and encom-
pass broad responsibility in the area of air and water pollution, solid wastes, pesticides, and
radiation.  A  vital part of EPA's national pollution control effort is the constant develop-
ment and dissemination of new technology.

It is now clear that only the most effective design and operation of air, water, and solids
waste control facilities, using the latest available techniques, will be adequate to meet the
future  air and water quality objectives and to ensure continued protection of the nation's
environment.  It is essential that this new technology be incorporated into the contempo-
rary design of pollution control facilities to achieve maximum benefit of our  pollution con-
trol expenditures.

The purpose  of this manual is to provide  the pulp  and paper  industry engineering commu-
nity with a new source of information for use in the planning, design, and operation of pres-
ent and future control facilities.  It is recognized that there are a number of design manuals,
manuals  of standard practice,  and design guidelines currently available in the field that ade-
quately describe and interpret current engineering practices as related to traditional environ-
mental control design concepts.  It is the intent of this manual  to supplement this existing
body  of knowledge by  describing  new pollution control  methods and by  discussing the
application of new techniques for more effectively removing a broad spectrum of contami-
nants from air and water discharges.  This manual contains two parts; the first describes air
pollution control, while the second presents water and solid waste pollution control for the
pulp and paper industry.

Much of the  information presented is based on  the evaluation and  operation of pilot,
demonstration, and full-scale  plants. The design criteria  thus generated represent typical
values.  These values should be used as a guide and should be tempered with sound engineer-
ing judgment based on a complete analysis of the specific application.

This manual will be updated as warranted by the advancing state-of-the-art to include new
data as they become available and to refine design criteria as additional full-scale operational
information is generated. Part I  of this manual, Air Pollution Control, is presented herein.
Part II, Water and Solids Pollution  Control, is currently in preparation and will shortly be
available for inclusion into this manual.
                                         xix

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                                     CHAPTER 1

                                  INTRODUCTION
From about 1965 to 1975,  air pollution technology in the United States pulp and paper
industry has undergone major advancements in the design of pollution abatement systems for
controlling gaseous and particulate emissions. These technological advances are particularly
significant  in  the  kraft  pulping  segment of  the pulp and paper  industry  because
reduced-sulfur  gas  and particulate emissions have  been reduced by changes within  the
production cycle itself.

A major difficulty confronting the design engineer has been the unavailability of the bulk of
this  emerging practical process design information in  a centralized  source, such as this
manual.  The design approach emphasized in  this  manual combines process  control
techniques to minimize the formation of air pollutants with treatment methods for removal
of air pollutants from process streams and flue gases.

The  information presented in this  manual is not limited  to  North America technology, but
also  includes technology on  internal process control in  the Scandinavian pulp and paper
industry.  Particular stress is  placed on explanation of the  chemical and  physical processes
that  generate air pollutants in specific unit operations so that the advantages and limitations
of both internal and external  process control methods can be understood.

Most of the concepts presented in  this  manual have been demonstrated in  actual field
installations. Some of these  design concepts may be superseded by  a  rapidly advancing
technology; others will endure. Future design approaches must properly consider all aspects
of the relationship between the process and air pollution. In particular, the design engineer
must be cognizant of the  rapidly  increasing cost of energy and its direct relationship to
certain air pollution control  measures. Additionally, deviations from  design performance,
which become  more important as  emission limitations become more stringent, will be less
tolerable.

Chapter 2 through 13 emphasize the  air pollution problems of the kraft  or sulfate pulping
process. Chapter 14 emphasizes the sulfite process.  Chapter 15 and 16 discuss air pollution
sources in pulp mills that are not a direct part of the pulping process. Chapter 17 discusses
process monitoring of air pollutants.
                                         1-1

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1.1  Gaseous and  Particulate Matter  Emissions  from Kraft Pulp and  Paper  Mill Process
     Sources

The United States pulp and paper industry includes more than 360 mechanical and chemical
pulp mills of all types. Further, the industry  plans to build  approximately 40 new mills
during the 1970's.

The industry has made a major  contribution to our country's effort to control excessive air
pollution from its facilities and has in the past cooperated extensively with government
regulatory agencies in dissemination of constantly emerging new control technology (1, 2,
3,4,5,6,7).

The  atmospheric  emissions from the kraft process include both gaseous and particulate
materials. The major gaseous emissions are malodorous reduced sulfur compounds, such as
hydrogen sulfide  (H2S), methyl  mercaptan (CH3SH), dimethyl sulfide (CH3SCH3), and
dimethyl disulfide (CH3SSCH3); oxides of sulfur (SOX); and oxides of nitrogen (NOX). The
particulate  matter emissions are primarily sodium sulfate  (Na2S04) and sodium carbonate
(Na2C03) from the recovery furnace, and sodium compounds from the lime kiln and smelt
tanks.

Both H2S  and the  organic  sulfides are extremely  odorous and are detectable at  a
concentration of only a few parts per billion. Thus odor control is one of the principal air
pollution problems in  a kraft pulp mill. Gas volumes released  per unit  of production vary
considerably between individual process units. Most kraft pulp mill flue gas streams contain
appreciable amounts of water vapor.

A summary of the major external control  techniques for  gaseous and particulate emissions
from specific kraft pulp mill sources is presented in Table 1-1. Specific  applications are
described in appropriate sections of the manual.

      1.1.1   Reduced Sulfur

The major gaseous emissions from kraft pulp mill sources are the malodorous reduced sulfur
compounds, organic nonsulfur  compounds, oxides  of  sulfur  and oxides of nitrogen. The
malodorous sulfur gases emitted from kraft pulp mill sources  all have extremely low odor
threshold levels of between 1 and 10 parts  per billion (ppb) by  volume (8). The most
common reduced  sulfur  compounds emitted  from  kraft pulp  mill sources  are H2S,
CH3SCH3 , and CH3SSCH3 ; other alkyl sulfur compounds can be emitted in small quantities
from certain wood species.

The major potential sources for the reduced sulfur gas emissions to the atmosphere include
digester blow  and relief gases, vacuum washer  hood and  seal tank vents, multiple-effect
                                         1-2

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                                   TABLE 1-1
             EXTERNAL CONTROL TECHNIQUES FOR GASEOUS AND
             PARTICULATE MATTER EMISSIONS FROM KRAFT PULP
                                 MILL SOURCES
           Emission Source
Gaseous Control
    Particulate
     Control
       Digester gases

       Washer vents
       Evaporator gases
       Condensate water

       Condensate Stripper Vent
       Black Liquor Oxidation
         Tower Vent
       Tall Oil Vent
       Recovery Furnance
       Smelt Tank
       Lime Kiln

       Slaker Vent
       Bleach Plant
       Paper Machines
       Power Boilers
Incineration
Condensation
Incineration
Incineration
Scrubbing
Condensation
Steam stripping
Air stripping
Incineration
Incineration

Scrubbing
Scrubbing
Scrubbing
Scrubbing

NA
Scrubbing
Incineration
Adsorption
Condensation
NA
NA (not applicable)

NA
NA
NA

NA
NA

NA
Precipitators
Scrubbing
Filtration
Scrubbing
Scrubbing
Precipitators
Scrubbing
NA
NA
Cyclones
Precipitators
Scrubbing
evaporation hotwell vents, recovery furnace flue gases following direct contact vents, smelt
dissolving tanks, slaker vents, black liquor oxidation tank vents, lime kiln exit vents and
wastewater  treatment operations. Summaries of  values on  variations in gas flow rates,
malodorous sulfur gas concentrations, and emission rates per unit production for the kraft
process units are presented  in Tables 1-2 to  1-4. These values  are based on a variety of
sources.
                                        1-3

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                                TABLE 1-2
    TYPICAL GAS CHARACTERISTICS FOR KRAFT PULP MILL PROCESSES
                                     Process Offgas Characteristic
Flow Rate*
m3/t
(ft3 /ton)
3-6,000
(96-190,000)
0.3-100
(10-3,200)
0.6-6
(20-200)
1,500-6,000
(48,000-190,000)
300-1,000
(9,600-32,000)
0.3-12
(10-400)
500-1,500
(16,000-48,000)
6,000-12,000
(190,000-380,000)
500-1,000
(16,000-32,000)
1,000-1,600
(32,000-51,000)
12-30
(400-1,000)
Temperature
°C
<°F)
65-100
(150-210)
25-60
(80-140)
75-150
(170-300)
20-45
(70-110)
55-75
(130-170)
80-145
(180-290)
70-80
(160-180)
120-180
(250-360)
70-110
(160-230)
65-95
(150-200)
65-75
(150-170)
Moisture Content
%
30-99
3-20
35-70
2-10
15-35
50-90
30-40
25-35
35-45
25-35
20-25
  Emission Source
Digester, Batch:
  Blow Gases
  Relief Gases


Digester, Continuous


Washer Hood Vent


Washer Seal Tank


Evaporator Hotwell


BLO Tower Exhaust


Recovery Furnace


Smelt Dissolving Tank


Lime Kiln Exhaust


Lime Slaker Vent
*At standard conditions of dry gas (21.1°C & 760mmHg (70°F & 29.92 in Hg))
Flow in cubic meters per metric ton and (cubic feet per short ton)
                                    1-4

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                                    TABLE 1-3
        TYPICAL REDUCED SULFUR GAS CONCENTRATIONS FROM KRAFT
                               PULP MILL SOURCES
                                      Concentration (ppm by volume)
          Emission Source

        Digester, Batch:
          Blow Gases

          Relief Gases
 H,S
CH3SH     CH3SCH3   CH3SSCH3
0-1,000     0-10,000    100-45,000   10-10,000
0-2,000    10-5,000    100-60,000   100-60,000
        Digester, Continuous     10-300    500-10,000  1,500-7,500   500-3,000
        Washer Hood Vent
        Washer Seal Tank
  0-5
  0-2
  0-5
 10-50
 0-15
10-700
 0-3
1-150
        Evaporator Hotwell    600-9,000   300-3,000   500-5,000    500-6,000

        BLO Tower Exhaust       0-10        0-25        10-500       2-95

        Recovery Furnace        0-1,500      0-200       0-100        2-95
          (after direct contact
          evaporator)

        Smelt Dissolving Tank     0-75        0-2          0-4          0-3

        Lime Kiln Exhaust       0-250       0-100        0-50        0-20

        Lime Slaker Vent         0-20        0-1          0-1          0-1


Both oxides of sulfur (SOX) and oxides of nitrogen (NOX)  can be emitted in varying
quantities from specific sources in the kraft chemical recovery system. The major source of
sulfur  dioxide   (SO2) emissions  is  the  kraft  chemical  recovery  furnace,  because  of
combustion  of sulfur-containing black liquor fuel. Under certain conditions, somewhat
similar  quantities of sulfur trioxide (S03) can be released to the atmosphere, particularly
when residual fuel oil is added as an auxiliary fuel (9). Lesser quantities of S02 can also be
released from the lime kiln and smelt dissolving tank. Trace quantities of sulfur oxides may
also be released from  other kraft mill sources. Oxides of nitrogen can be formed in any fuel
combustion process by the reaction between oxygen and nitrogen at elevated temperatures.
                                       1-5

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                                   TABLE 1-4
     TYPICAL REDUCED SULFUR GAS EMISSION RATES FROM KRAFT PULP
                                 MILL SOURCES
                             Emission Rate, kg sulfur per metric ton of air dried pulp
       Emission Source

    Digester, Batch:
      Blow Gases

      Relief Gases

    Digester, Continuous

    Washer Hood Vent

    Washer Seal Tank

    Evaporator Hotwell

    BLO Tower Exhaust

    Recovery Furnace
      (after direct contact
      evaporator)

    Smelt Dissolving Tank

    Lime Kiln Exhaust

    Lime Slaker Vent
H2S
0-0.1
0-0.05
0-0.1
0-0.1
0-0.01
0.05-1.5
0-0.01
0-25
CH3SH
0-1.0
0-0.3
0.5-1.0
0.05-1.0
0-0.01
0.05-0.8
0-0.1
0-2
CH3SCH3
0-2.5
0.05-0.8
0.05-0.5
0.05-0.5
0-0.05
0.05-1.0
0-0.4
0-1
CH3SSCH3
0-1.0
0.05-1.0
0.05-0.4
0.05-0.4
0-0.03
0.05-1.0
0-0.3
0-0.3
0-1
0-0.5
3-0.01
0-0.8
0-0.2
0-0.01
0-0.5
0-0.1
0-0.01
0-0.3
0-0.05
0-0.01
The major constituent formed in nitric oxide (NO), a small portion of which can be oxidized
to form nitrogen dioxide (N02):  together they are classified as total oxides  of nitrogen.
Nitrogen oxide emissions from  kraft pulp mill process sources, such as the recovery furnace
and lime kiln, are normally lower than for most other fuel combustion processes. This is
primarily due to the large quantities of water present in black liquor and lime and which act
as a heat sink to suppress the flame temperature. Larger quantities of oxides of nitrogen can
be formed, however, when auxiliary fuels such as natural gas or fuel oil are added to the
recovery furnace.
                                        1-6

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                                    TABLE 1-5
     TYPICAL EMISSION CONCENTRATIONS AND RATES FOR SOX AND NOX
                FROM KRAFT PULP MILL COMBUSTION SOURCES

                          Concentration (ppm by vol.)        Emission Rate, kg/t*

  Emission Source       S02     S03   NOx(asN02)   S02   S03   NOx(asN02)

Recovery Furnace:
  No Auxiliary Fuel    0-1,200   0-100       10-70       0-40    0-4       0.7-5

  Auxiliary Fuel
     Added           0-1,500   0-150      50-400      0-50    0-6       1.2-10

  Lime Kiln Exhaust    0-200      --       100-260      0-1.4    --       10-25

Smelt Dissolving         0-100      --          --        0-0.2
  Tank

*kilograms per metric ton of air dried pulp.


A summary of concentrations and emission rates for oxides of sulfur and oxides of nitrogen
for specific kraft pulp mill sources is presented in Table 1-5. The information included is
based upon a variety  of industry sources. The extreme variations  in operating conditions
that occur in the industry, including  operating combustion temperature and type of fuel,
account for the broad ranges in these data.

     1.1.3   Organic Compounds

Organic compounds other  than those containing sulfur can also  be emitted in varying
quantities  from  several different  kraft  pulp mill process sources.  The major  types of
materials that can be released to the atmosphere include terpenes, hydrocarbons, alcohols,
phenols and other organic compounds liberated from wood. Additional organic compounds
can  be produced when organic  materials are applied as coatings to paper sheet or can be
induced when spent caustic solutions are used as chemical make-up for the process.

The primary significance of these materials is that they may either act  directly as odorant
gases or as liquid  particulate carriers for odorous  sulfur gas molecules, particularly the
terpene compounds. The olefinic hydrocarbons or terpenes may undergo photochemical
reactions in polluted atmospheres.
                                        1-7

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The major  potential process sources of  organic, nonsulfur  compound  emissions to the
atmosphere  include digester  blow  and relief  gases, multiple-effect evaporator  noncon-
densable gases, brown stock washer hood and seal tank vents, black liquor oxidation tower
vents, black liquor storage tank  vents, direct-contact evaporator exhaust gases, digester and
evaporator condensate water vents and wastewater treatment facilities, and paper  machine
coating and drier vents.

Major process variables that affect emissions of these compounds to the atmosphere include
the wood species being pulped or the type of coating material applied, the respective organic
or condensate stream temperature, the volatility of the respective organic compounds, and
the type  and effectiveness of any air pollution control device.

     1.1.4   Particulate Matter Emissions

The major  potential process sources  of  particulate emissions from  the  kraft  chemical
recovery system are the recovery furnace, the smelt dissolving tank, and the lime kiln. The
recovery furnace is the largest potential particulate  emission source. The major chemical
constituent in the recovery boiler particulate emissions is Na2S04, with smaller quantities
of Na2C03  and sodium chloride  (NaCl) also present. The  smelt dissolving tank vents and
lime kiln exhaust gases are also sources of varying quantities of particulate matter consisting
primarily of carbonate, hydroxide, sulfate and chloride salts of calcium and sodium. Particle
sizes from these sources can range from 0.1 /zm (4 X 1CT6 in) to greater  than 1000 /u (4 X
10~2 in)  in  diameter for uncontrolled emissions and from 0.1 to 10 Mm (4 X 10~6 to 4 X
1CT4 in) in diameter where these  sources have high efficiency particulate control devices.

The two major types of particulate matter control devices  employed for kraft recovery
furnaces  are electrostatic precipitators (ESP)  following  cyclone or  cascade-type  direct
contact evaporators, and venturi-type evaporator-scrubbers in a one-or two-stage configura-
tion. Low pressure drop secondary wet scrubbers have been employed to supplement older
and less  efficient primary particulate collection  devices at several existing mills to alleviate
particle fallout in areas adjacent to the plant premises. Packed tower or showered  mesh
demister scrubbers  are employed  for particulate  control on  smelt dissolving tank exhaust
gases, while venturi or cyclonic  scrubbers are normally used for particulate control on lime
kiln or fluorosolid  calciner exhaust  gases.  The amount of particulate matter  emitted from
kraft pulp mill process sources depends both on  the process operating conditions and on the
types and collection efficiencies of any control devices employed.

A summary of typical ranges in particulate concentrations and emissions rates from kraft
pulp mill process sources is presented in Table 1-6.
                                          1-8

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                                    TABLE 1-6
             TYPICAL CONCENTRATIONS AND EMISSION RATES FOR
           PARTICULATE MATTER FROM KRAFT PULP MILL SOURCES
                           (AFTER CONTROL DEVICES)

        Emission Source            Concentration             Emission Rate
                                g/scm       (gr/scf)        kg/t       (Ib/ton)

      Recovery Furnace:
        After Electrostatic
        Precipitator            0.06-1.1     (0.03-0.5)     0.5-12      (1.0-24)

        After Venturi
        Evaporator             0.9-2.3      (0.4-1.0)       7-25       (14-50)

      Lime Kiln                0.07-1.1     (0.03-0.5)    0.15-2.5     (0.3-5.0)

      Smelt Dissolving Tank     0.04-2.3     (0.02-1.0)    0.01-0.5    (0.02-1.0)

1.2 Gaseous and Particulate Emissions from Sulfite Pulp and Paper Mill Process Sources

The primary emissions from the sulfite pulping process are S02 and particulate  matter. In
special cases  of  burning  alkaline  sulfite  liquor  in  recovery  furnaces under reducing
conditions, H2S  emissions  may  also  occur. Otherwise, there are practically no organic
reduced sulfur compounds produced in the sulfite process. Nitrogen oxides are emitted from
various combustion sources, particularly from the  recovery furnace of ammonium-based
mills.

    1.2.1  Sulfur Dioxide

Various process sources within the sulfite mill can emit SO2. The  main sources are the
digester blow pits,  multiple-effect evaporators, and liquid burning  or chemical recovery
systems. Minor process sources include pulp washers and the acid preparation plant. Typical
values of S02  emission rates are listed in Table 1-7.

    1.2.2  Particulate Matter

The recovery furnace is the significant process source of particulate matter in a sulfite pulp
mill. Potential particulate matter emissions depend  greatly on the degree of recovery of
sulfite waste liquor, as well as on the degree of control of particulate matter.
                                        1-9

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                                     TABLE 1-7
                       TYPICAL S02 EMISSION RATES FROM
                           SULFITE PULP MILL SOURCES

           Emission Source       	Emission Rate, kg/t (lb/ton)*	

                                   Uncontrolled            Controlled**

           Blow Pit:
             Hot blow             30-75(60-150)            1-2.5(1-5)
             Cold blow              2-10(4-20)           0.05-0.3(0.1-0.6)

           Evaporators               1-30 (2-60)           0.025-1 (0.05-2)

           Recovery Process       80-250 (160-500)          6-20 (12-40)

           Washers                  0.5-1 (1-2)

           Acid Preparation          0.5-1 (1-2)

            *Per mass, t (ton), of air dried pulp.
           **Alkaline scrubbing of gases.

     1.2.3   Nitrogen Oxides (NOX)

In ammonium-based  sulfite pulp mills, combustion of the spent sulfite liquor will result in
emissions of nitrogen oxides from the  recovery furnace. Emissions from one such system
ranged from 4.7 kg/t  to 11.8 kg/t (9.4 to 23.6 lb/ton).

1.3  Power Boilers

The  pulp and paper  industry is a major energy consumer in the United States, accounting
for  about  2.2  percent  of the total  national  energy consumption. This amounts  to
approximately 1.6 X  101 8 J per year (1.5 X  1015 BTU/yr) of which approximately one-half
is  associated with  the  manufacture of  kraft  pulp  and paper  (10).  Typical energy
consumption requirements for a kraft pulp mill are about 28.7 GJ per metric ton of air dried
pulp (30  million  BTU/ton)  of which 50 to 60 percent  can  normally be supplied by
combustion of the black liquor solids (11).

For  mills  employing  on  site  debarking,  an  additional  20-30 percent  of  the energy
requirement can be supplied by the burning of waste wood in bark boilers. As a result, it is
normally necessary for kraft pulp  mills to  obtain about 5  to 30 percent of their energy
requirements by burning supplementary fuels such as coal, fuel oil,  and natural gas. The

                                       1-10

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total energy that must be supplied by supplementary combustion of coal, oil, gas, or wood
in kraft pulp and paper mill power boilers can range from less than 0.9 million to more than
5.4 GJ per metric ton of pulp (0.75 to 4.5 million BTU per ton of pulp).

The exact energy requirements for the auxiliary fuel burning of coal, oil, gas, or wood will
vary between individual  mills  depending  on their respective energy  balances, physical
characteristics,  and availability  of fuels  in each local area.  The  major air pollutants  of
possible concern from auxiliary fuel burning operations include particulate matter from coal
and wood, sulfur oxides from coal and fuel oil, and nitrogen oxides from coal, oil, gas, and
wood.  Available particulate matter control devices  for coal and wood-fired power  boilers
include electrostatic precipitators, liquid scrubbers, fabric filters, and mechanical cyclones.
A summary of uncontrolled air pollutant emissions from auxiliary fuel combustion in power
boilers in the pulp and paper industry is presented in  Table 1-8 (12).
                                         1 11

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                                    TABLE 1-8
            UNCONTROLLED AIR POLLUTANT EMISSIONS FROM FUEL
                  COMBUSTION IN AUXILIARY POWER  BOILERS

                  	Air Pollutant Emission Rate; kg/106 kj (lb/106 BTU)	

  Air Pollutant     Bituminous Coal*  Residual Fuel Oil**   Natural Gas4"  Waste Wood++
Particulate Matter     0.38 (0.95)
Sulfur Oxides
  (asSO2)

Nitrogen Oxides
  (asN02)

Hydrocarbons
  (asCH4)
 0.84(2.1)
 0.39 (0.98)
0.007 (0.02)
                   0.024(0.060)     0.005(0.01)    1.50(3.75)
      0.46(1.1)           --        0.16(0.40)
      0.23(0.58)       0.16(0.40)     0.43(1.1)
                       0.17(0.43)     0.11(0.28)
Carbon Monoxide     0.021 (0.053)
 *Based on average heating value of 25.7 MJ/kg coal (11,000 BTU/lb)
**Based on average heating value of 41.9 GJ/m3 oil (150,000 BTU/gal)
 +Based on average heating value of 39.1 MJ/m3 natural gas (1,050 BTU/ft3)
++Based on waste wood heating values as follows:
         Item

    Moisture Content

    Heating Value

    Heating Value
   Units

 % by mass

 MJ/kg

 BTU/lb
Dry Basis

    0.0

   18.6

 8000
Wet Basis

   50.0

    9.3

 4000
                                                   0.11(0.28)
                                        1-12

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 1.4  References

     1.4.1   Cited References

  1.  Hendrickson, E. R., Robertson, J.  E., and Koogler, J. B., Control  of Atmospheric
     Emissions in the Wood Pulping Industry, Volumes I, II, III. Final Report, Contract No.
     CPA 22-69-18, U.S. Department of Health,  Education,  and Welfare, National Air
     Pollution Control Administration, Raleigh, North Carolina,  March 15, 1970.

  2.  Proceedings  of the Symposium on Recovery of Pulping Chemicals: Helsinki, Finland,
     May 13-17, 1968,  Finnish  Pulp and Paper  Research Institute, EKONO, Helsinki,
     Finland, 1969.

  3.  Cooper, H. B. H., and Rossano, A. T., Jr., Odor Control Technology for Kraft Pulp Mills.
     Report prepared for U.S. Environmental Protection Agency Odor Control Technology
     Manual, University of Washington, Seattle,  Washington, August, 1970.

  4.  Proceeding of  the International Conference on. Atmospheric Emissions from Sulfate
     Pulping, April 28, 1966, Sanibel Island, Florida. Hendrickson, E. R. (ed.) Sponsored by
     USHPS, University of Florida, and National Council for Air and Stream Improvement.
     Deland, Florida, E. 0. Painter and Printing  Co., 1966.

  5.  Atmospheric Emissions from the Pulp and Paper Manufacturing Industry.  EPA-450/1-
     73-002. September  1973. (Also published as  NCASI  Technical Bulletin  No. 69,
     February, 1974.)

  6.  Field Surveillance and Enforcement Guide-Wood Pulping Industry, U.S. EPA Contract
     No. 68-02-0618. Prepared for EPA, Research  Triangle Park, N.C., October 15,  1973
     (Revised Draft).

  7.  Galeano, S. F., and Leopold, K. M.,^4 Study  of Emissions of Nitrogen Oxides in the Pulp
     Mill. Tappi, 56:74-76, March 1973.

 8.  Wilby,  F.V., Variations  in Recognition Odor  Threshold of a Pond.  Journal of Air
     PoUution Control Association, 19:96-100, February 1969.

 9.  Maksimov, V.  F., Bushmelav, V. A., Torf, A. I., and Lesohhin,  V.  B., Testing the
     Turbulent Flow Venturi Apparatus, Bumazhnaya Proyshlennost, 40:14-15, May 1965.

10.  Personal communication  with Dr. Ronald Slinn, American  Paper Institute, New York,
     New York, November, 1973.

11.  Miller, R. R., One Pulp and Paper Company's View of the Energy Crisis. Tappi, 57:
     62-64, February, 1974.
                                       1-13

-------
12.  Compilation  of Air  Pollutant  Emission  Factors. U.S.  Environmental Protection
    Agency,  Office of Air Programs, Research Triangle Park, N.C. Publication No. AP-42
    (Revised). February 1972.

    1.4.2  Additional Reading

1.   Britt, K. W. (ed.), Handbook of Pulp  and Paper Technology.  New York,  Reinhold
    Publishing Company, 1964.

2.   Wenzyl,  H., Kraft Pulping:  Theory and Practice. New York, Lockwood Publishing
    Company, 1967.

3.   Whitney, R.  P.  (ed.),  Chemical  Recovery in  Alkaline Pulping Processes,  Tappi
    Monograph Series No. 32.  Technical Association of the Pulp and Paper Industry, New
    York, New York, 1968.

4.   Turpentine Recovery Systems. Pulp Chemicals Association, New York, 1972.

5.   Tall Oil Recovery. Pulp Chemicals Association, New York, 1968.
                                       1-14

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                                    CHAPTER 2

                                 DIGESTER GASES
The digestion process is the third most important source  of odor pollution in the kraft
process. Black liquor combustion and weak black liquor evaporation are first and second,
respectively.  Gases from  the  digester contain  organic sulfur compounds  from reactions
between components of the wood and the sulfide of the white liquor. These gases also
contain some H2S, turpentine, and traces of methanol (CH3OH), ethanol (CH3CH2OH),
and acetone (CH3COCH3), as well as displaced air.

The composition  and quantity  of  digester gases will differ  between batch digesters and
continuous  digesters.  Variations will also occur with  the  type  of  wood,  the  sulfide
concentration in the  white liquor, the final cooking temperature, and the cooking time.
Most of these factors are determined by production requirements and vary widely with the
production schedule.

The basic method for minimizing odor pollution in the digester area is to effect adequate
condensation and  to  contain the relief and blow gases. As a  further step, contained gases
may be incinerated. In the following sections important factors in using these methods are
discussed.

2.1  Batch Digesters

Batch digesters often  present air pollution problems because of surges of gas flow produced
during blowing.  These surges can temporarily  overload the  condensing  or heat recovery
system. Economics of production demand that a plant be operated at full capacity. Often,
the capacity of the batch  digesters is limited by their condensing systems, so that full-scale
production can overload  these systems. This problem  can be  compounded since  liquid
carryover  from  short  term overloads can increase fouling of the  condensing systems; thus,
further lowering the capacity and making the systems inadequate for even less than  design
flows. The condensing system for batch digesters must be designed for peak flows and all
components in the  heat recovery system must be kept in good operating condition. To allow
proper operation,  the system  must  have enough temperature  and  pressure  difference
measurements to enable the operator to judge the condition of the system. Figure 2-1  shows
a typical batch digester and its blow heat recovery system. Point sources of odor release are
indicated in this figure.
                                        2-1

-------
l\s
                                 SEPERATOR
                                                                        SECONDARY
                                                                        CONDENSER
                  12     CHIPS
                  13
                  14
                  15
                                                                                      11  HOT WATER
                                                                                      10  COOLING WATER

                                                                                      (?) RELIEF GAS
                                                                                          TURPENTINE


                                                                                          FOUL CONDENSATE
                                                                                          BLOW GAS
                                                                            HEAT
                                                                            EXCHANGER
                                                                                      5   HOT WATER
                                                                                      4   WARM WATER
                                                                                      ^D  BLOW CONDENSTATE BLEED

                                                                                      2   FRESH WATER
                                                                                      fT)  PULP - LIQUOR
                                       POINTS OF POSSIBLE ORDER RELEASE ARE ENCIRCLED BY Q
                                                        FIGURE 2-1
                                               BATCH DIGESTER FLOWSHEET

-------
     2.1.1   Blow Gases

Batch digester blow gases are a major cause of air pollution. As a rough guide to the volumes
that can be expected, a single blow will produce approximately one ton of steam per ton of
air  dry  pulp (90 percent absolute  dry fiber).  Blowing a  200m3 (7100  ft3) digester to
atmospheric pressure from 0.64 MPa (78 psig) produces about 20 t (22 tons) of pulp and
34,000 m3  (1,200,000 ft3) of steam within a 20 minute period (Figure 2-2). Effective
condensation of this volume of steam requires a blow heat recovery system with adequate
capacity and a control system which reacts quickly but retains stability.

The vent from the blow heat accumulator (flow 6 in Figure 2-1) might typically show flow
histories as illustrated in Figure 2-3. These flow histories were recorded with a pitot tube.
Normally, about 90  percent of the volume is steam, and  the  noncondensable portion is
about 3 m3 per metric ton of pulp (96 ft3/ton). Typical flow and composition ranges were
given in Tables 1-2 and 1-3. A successful blow gas treatment program starts with an efficient
blow heat recovery system (section 2.2.1.1) followed by a flow equalization system and a
gas  incineration system.

Blow gas flows, such  as those shown in Figure  2-3,  cases  2 and  3,  will overload  any
reasonable  flow equalization  system and  thus seriously degrade  the entire  blow gas
treatment program. Loss of recoverable heat is also probable in these two cases.
                                                                          •  2500
                                                                                   <
                                                                                   UJ
                                                                          -  1000  en
                                                                                   00
            0
                                TIME, MINUTES
                                   FIGURE 2-2
               KRAFT BATCH DIGESTER  BLOW STEAM FLOW (1
                                       2-3

-------
    20 -I
TO
"o  10 -

 9  ,
      0
   40 -
    30 -
     0
             0
              23
              5
                V( t ) dt = 550m3/blow
           MEAN - 1400 m3/h
            23
             5  V( t )  dt = 300m3/blow
                                                 25
5       10       15       20
       TIME, MINUTES
CASE  I.    NORMAL OPERATIONS
       0
CASE 2.
                                         20
                                  25
              10        15
            TIME, MINUTES
MALFUNCTION OF BLOW HEAT RECOVERY HEAT  EXCHANGERS
    50 -


 .c  40 -
 \
 WE
 °0  30 -


 9  20 -
 U_


    10 -
     0
              23
              < v(t ) dt = 4500m3/blow
             \-J
          ME AN -11,700 mVh
                          20
                                                  25
     CASE 3.
               5        10        15
                      TIME, MINUTES
               INCREASING MALFUNCTION OF HEAT EXCHANGERS

                           FIGURE 2-3
      KRAFT BATCH DIGESTER BLOW GAS FLOW AFTER CONDENSING
                  AND WITHOUT EQUALIZATION (2)
                               2-4

-------
Final odor gas treatment by white liquor scrubbing is not very efficient because of the
dominance of  nonionizable organic sulfur  compounds.  Only H2S and  CH3SH  can  be
efficiently recovered by alkaline scrubbing (Figures 2-4 and 2-5).

Therefore, incineration of the blow gases (section 4.3) after proper gas flow equalization
(section 4.2.2) is recommended.

         2.1.1.1   Blow Heat Recovery

The  blow heat recovery  system, as shown  in Figure 2-1, and its proper operation will
significantly affect the further  treatment of both blow gases and blow condensates. Factors
that  may affect the batch digester air pollution abatement program are final blow pressure,
blow tank  drop separation, primary blow steam condenser,  secondary blow steam con-
denser, blow heat accumulator and blow heat recovery heat exchanger.

The  blow gas flow is directly  proportional to the final blow pressure (i.e., the higher the
final blow  pressure,  the  more violent the blow gas  flow,  and the  more difficult it is to
condense and collect  the  blow  gases for treatment). Decreasing the final pressure, however,
              I
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              X
              q
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              UJ
              cr
              LU
              cr
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                    700
                    600
                      0
                                       6      8     10
                                       pH at  25°C
                                   FIGURE 2-4
                  VAPOR PRESSURE OF 0.01 M H2S VS. pH AT
                         VARIOUS TEMPERATURES  (4)
                                       2-5

-------
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-------
Usually it is a simple temperature control of the water flow (See Figure 2-6). This control
system is too slow for the abrupt beginning of the blow and, when made more sensitive to
compensate  for speed, it often becomes unstable. No realh  good solution to this control
problem exists. One solution to minimize blow gas treatment problems is to install a feed
forward control. This control can be a simple  on-commarid to open the condenser water
valve at a specified time before the digester blow valve is opened. This results in lowering  the
temperature of the hot  water produced in the blow heat recovery system (1). Condenser
outlet temperature should be kept around 90° C  (194° F).

A secondary condenser is very common as a backup for the primary condenser. If properly
controlled it will accommodate temporary  overloads  on the  primary condenser, thus
facilitating treatment  of noncondensable blow  gases.  It may be  either a  direct contact
condenser or a surface  condenser  (the latter gives a more  concentrated condensate). A
suitable condenser outlet temperature (see Figure 2-7) should be around 50° C (122° F).

For  the  blow  heat  accumulator,  the  two determining factors  are sufficient size  to
accommodate large blows and long pauses between blows, and an internal construction that
will cause the zone  between hot- and cooled-water  to be  as sharp and  undisturbed as
possible. Frequently,  hot water does find its way down to the bottom of the accumulator,
enters the condenser, and thereby drastically reduces the condensing capacity. Subsequent
gas handling problems occur.
                                BLOW
                                 GAS
SECONDARY
CONDENSER
      BLOW
      STEAM
   PRIMARY
  CONDENSER
    BLOW
 CONDENSATE
                              HOT WATER
                              WARM WATER
                                                                      COLD WATER
                                   FIGURE 2-6
                   BLOW HEAT RECOVERY CONTROL SYSTEM
                                        2-7

-------
             20  -
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 o>


o
              15  H
                                               o
                                               X
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6
          LJ
          o
          z
          o
          o
              10 -
     5 -
                                   x
                                   o
              0
                50       60        70        80       90

                  CONDENSER   OUTLET  TEMPERATURE, °C

                    •      HYDROGEN SULFIDE
                    o      METHYL  MERCAPTAN
                    X      DIMETHYL SULRDE
                    a      DIMETHYL DISULFIDE
                               FIGURE 2-7
          ODOR COMPOUNDS IN RELIEF GAS AFTER TURPENTINE
               CONDENSER  AS A FUNCTION OF CONDENSER
                       OUTLET TEMPERATURE (2)
The internal construction should consist of adequate baffles to disperse the hot condensate
in top layers that are as even as possible. The accumulator should also be evenly insulated to
avoid local cold spots where liquid will cool and flow down, mixing the contents of the
accumulator.
                                   2-8

-------
Experiences from mills show that installed heat exchanger  surfaces,  which are initially
sufficient, may  later prove too small. This change  occurs  because surfaces gradually get
fouled, especially on the blow heat side, by fiber and liquor carryover. These dirty surfaces
reduce heat  transfer and increase pressure drops, thus reducing flows and further decreasing
heat transfer. Spiral heat exchangers have an advantage over plate exchangers in decreasing
fouling by fiber  and liquor carryover, but they are more difficult to clean. Another factor,
which will reduce heat transfer,  is a reduced hot water demand (flow 5 in Figure 2-1) or a
higher warm water temperature (flow 4 in Figure 2-1). Reduced heat transfer means a higher
temperature in  the accumulator  bottom  and reduced  condensation of the blow gases.
Injecting fresh cool water (flow  2 in Figure 2-1) to the primary condenser with controlled
temperature counteracts the reduction in  condensation. This method is used extensively,
but it has the great drawback of increasing substantially the blow condensate bleed (flow 3
in Figure 2-1).

The  blow condensate  bleed has to be treated, for  instance,  by steam stripping, and the
stripping equipment investment and the stripping steam consumption will directly affect the
size  of the  bleed.  Thus,  if  the  blow heat recovery  capacity is insufficient,  a trade-off
between air  pollution abatement and water pollution abatement must be made. A successful
odor abatement  program should include monitoring of temperature and pressure drops over
the blow heat recovery heat exchangers to check their performance.

         2.1.1.2   Improving Blow Heat Recovery

The most common approach to  blow gas treatment is to use a blow gas collection and flow
equalization system followed by gas incineration.  Any  malfunction in the  blow heat
recovery  chain  will adversely affect  the  blow  gas treatment, primarily the  blow gas
collection and flow equalization system, as clearly demonstrated in Figure 2-3.

Mills with odor abatement problems caused by inadequate blow heat recovery can improve
their situation by:

     1.   Improving the blow heat recovery control system through:

         a.    More extensive instrumentation of condensing, heat accumulating, and heat
              exchanging systems, and

         b.   Better tuning of the control  system.

     2.   Increasing the blow heat recovery capacity through installation of:

         a.    More heat exchanger surface,

         b.   More pump capacity for the  primary blow steam condenser,

                                         2-9

-------
          c.   More pump capacity for the heat exchanger circulation,

          d.   Secondary blow-steam condenser,

          e.   Better baffles in the accumulator, and

          f.   More accumulator volume.

     3.    Increasing the flow equalization capacity through installation of more gas (up to
          10 times more) accumulator volume.

     4.    Decreasing  the  blow heat recovery system load through prolonging the blow
          period, thereby  decreasing pulp production.

An odor control program for the blow heat recovery system should consider recommenda-
tion 1, above, first, and then the others. Recommendation 4 should only be used as a last
resort.

     2.1.2  Relief Gases

The purpose of the digester relief is to remove air and other noncondensable gases during
operation and to reduce  digester pressure  before  blowing. Relief takes place more or less
continuously during  digestion, as well  as during the final deliberate blow  pressure relief
before blowing. For softwood, the amount of steam in the continuous relief is about 180 kg
(400 Ib),  and  in the final relief around 90 kg (200 Ib) (5). In modern batch digesters, the
final blow pressure  is more efficiently and swiftly reduced by introducing cooler weak black
liquor into the upper part of the digester.

The relief is usually passed through a  surface condenser, thus producing  hot water. When
pulping softwood,  the relief condensate will contain turpentine, which is recovered in a
separating vessel.  The noncondensable relief gases flows and compositions are presented in
Table 1-2 and Table 1-3. Hardwood usually  produces more relief gas than softwood. A
normal softwood value is around 1 m3 per metric ton of pulp (32 ft3/ton).

The relief gases do  not present a  major problem because of the relatively small volume and
even flow as compared to blow gases. As with the blow gases, the recommended treatment
is incineration (section 4.3).

     2.1.3  Turpentine Recovery

Turpentine recovery takes place with turpentine-containing softwoods. Recovery is through
gas relief to the turpentine recovery system. Typically about 270 kg  steam/t (540 Ib/ton) of
pulp is relieved to and condensed in the recovery system. The condensate is separated into

                                        2-10

-------
one turpentine fraction and one  underflow of contaminated condensate in the decanter.
This condensate is one source  of odor in a kraft mill and can be treated by steam stripping
(section 5.5).

The  amount of odor compounds in the relief gas can be  decreased  by  decreasing the
condensate  outlet temperature  (Figure  2-7),  which  also produces  a lower  hot water
temperature from the condenser.

Decreasing  condenser  outlet  temperature also means greater  safety  in  collecting  and
handling relief gases  when there is a possibility of contact with air, since lower condensing
temperature correspondingly produces less turpentine in the relief gas (Figure 2-8).

2.2  Continuous Digester Gases

Continuous  digesters present a much smaller pollution problem than batch digesters because
contaminated condensates and odorous gases flow at a regular rate. Treatment capacity can
be designed  for the mean flow, without the need for peak flow equalization as  with batch
digesters.  Continuous digester  odor gases do  not differ significantly in composition from
batch digester odor gases (see Table 1-3).

The  amount of noncondensable gases released from the digester itself varies according to
how the flash steam  is used (Table 1-2). A rather typical downflow of a continuous digester
arrangement is presented in Figure 2-9. Other types of  continuous digesters exist, but they
do not  differ very much with respect to air pollution generation. Some types may have a
relief vent from the  top of the digester to the turpentine recovery system. Most existing
units have a countercurrent wash  zone in the bottom of the digester. The wash liquor (flow
17 in Figure 2-9) temperature  is 75-80°  C (167-176° F). This gives a pulp-liquor (flow 1 in
Figure 2-9)  temperature of 80-85° C (176-185° F). This so-called  "cold blow" produces
very minor odor emissions  from the blow tank. Thus, the major gaseous odor release from
the digester will be the flash steam. This steam can be used and condensed in many different
ways  yielding  noncondensable gases, which may be collected and  incinerated.  The total
amount of flash steam is about 0.8 ton per ton of pulp.

     2.2.1   Flash Steam

The  spent liquor from  the continuous  digester is drawn off  and expanded, or flashed,
usually  in two stages. The flash steam  from  the  primary  flash tank is usually used to
impregnate the chips in the presteaming vessel. The presteaming vessel relief, which contains
the noncondensable gases from the primary flash steam and from the presteamed chips, then
passes  to a turpentine recovery  system.  The  amount of primary  flash steam  to the
presteaming vessel is 0.5-0.6 ton per ton of pulp.  The secondary flash steam  amount  is
0.2-0.3 ton per ton of pulp, and it may be used for various purposes.
                                         2-11

-------
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-------
 14  STEAM
 15  COND.
 13  WHITE LIQUOR


 12  CHIPS

 (D VENT
 11  HOT WATER
 10  COOLING WATER


 (D TURPENTINE



 (D FOUL CONDENSATE

(l6) FLASH STEAM
                                                               (10) WEAK LIQUOR

                                                                (?) PULP+LIQUOR


                                                               fvT) WASH LIQUOR
               POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY (_J

                                    FIGURE 2-9
                      CONTINUOUS DIGESTER FLOW SHEET

         2.2.1.1   Turpentine Recovery

The relief from the presteaming vessel is brought to a condenser. The amount of condensate
is about 0.1 ton per ton of pulp. When pulping soft woods, the condensate is separated in
the decanter into a turpentine fraction (flow 8 in Figure 2-9) and an odorous water fraction
(flow 7 in Figure  2-9), which is typically treated by steam stripping. The  amount and
composition of the noncondensable gases (flow 9 in Figure 2-9) varies between values given
in Tables 1-2 and 1-3. The amount of odorous compounds can be decreased by decreasing
the condenser outlet temperature, as earlier  shown in Figure 2-7. The remaining odorous
gases  may  be  collected  and incinerated  (section 4.3). For safety aspects,  see section
4.1.3.
                                       2-13

-------
         2.2.1.2  Flash Heat Recovery

Primary flash heat recovery takes place in the presteaming vessel, as described previously.
The secondary flash steam (flow 16 in Figure 2-9) may be utilized in different parts of the
kraft mill.  Some possibilities 'are,  using it  as additional impregnating steam, routing it to
the turpentine  recovery system (increasing  the turpentine and hot  water  output),
condensing it in a separate flash steam condenser for hot water production, using it as a
partial  steam source for a black-liquor evaporation plant, and using it as a partial steam
source  for a contaminated condensate steam-stripping column.

Wherever the flash steam is  condensed,  its noncondensable components will remain and
must be collected and treated, preferably by incineration. If treatment does not occur, part
of the  digester odor is simply transferred to another release point,  such as the evaporation
plant.

2.3 References

1.  Kock, P. A., Treating Kraft Digester Waste Gases. M.S. Thesis. Chemical Engineering
    Department, Helsinki Technical University, Finland. September 12, 1972. (Swedish).

2.  Kekki, R.,  Kraft  Mill  Odor Abatement  by  Condensate Stripping and  Waste Gas
    Incineration.  M.S.  Thesis,  Wood Industry Department, Helsinki Technical University,
    Finland. September 18, 1969. (Finnish).

3.  Sarkanen. K. V., Hrutifiord, B. F., Johanson, L. N., Gardner, H. S., Kraft Odor. Tappi,
    53:776-783, May 1970.

4.   Martin, G. C., Fiber Carryover with Blow Tank Exhaust. Tappi, 52:2360-2362, Decem-
    ber 1969.

5.   The  Finnish  Paper Engineers' Association (SPY). The  Pulping  of Wood.  Helsinki,
     Frenckellin Kirjapaino Oy, 1968 (Finnish).

6.  Weast, P. C. (ed.), Handbook of Chemistry and Physics, 47th edition. Cleveland, The
     Chemical Rubber Co., 1966. p. D105-D138.
                                         2-14

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                                     CHAPTER 3

                               EVAPORATION GASES
The evaporation of black liquor is  one of the three major  malodorous  gas  producing
processes in a kraft mill; the other two are black liquor combustion and wood digestion.

For the final evaporation of black liquor to combustion strength, there are two different
methods in current use. One  method is indirect evaporation with steam; the other is direct
contact evaporation with hot flue gases from the recovery boiler. This latter method causes
an air pollution problem in the combustion of  black liquor and is discussed in Chapter 10.

The evaporator  gases discussed here are only the noncondensable gases generated  from
indirect steam evaporation of black liquor. Evaporator gases will contain sulfur and organic
compounds boiled off from the black liquor, plus displaced and leaked air.

The amount and composition of the evaporator gases vary widely depending on black liquor
properties, such  as originating wood, pH, and sulfide concentration,  and evaporation plant
properties, such as temperature level, equipment type, and plant condition.

The most  significant difference between evaporator and digester gases is that the dominating
compounds in  the  evaporator  gases  are  H2S  and  CH3SH. instead  of  organic  sulfur
compounds. This feature  makes odor abatement  and sulfur recovery through white liquor
scrubbing  entirely feasible. Incineration is another possible control technique.

3.1  Black Liquor Properties

The one black liquor property that has the greatest effect  on the evaporation plant odor
release  is  its sulfide  concentration.  Together  with  pH  and  temperature, the sulfide
concentration determines  the quantity  of H2S liberated from black liquor and  eventually
vented. This relationship is shown in Figure 3-1, which also suggests that  a black liquor with
a certain sulfidity will generate more H2S when it has a higher dry solids  concentration (i.e.,
after more evaporation).

Bringing the  sulfide concentration down to zero will obviously eliminate the H2S release
from the evaporation  plant; this reduction in sulfide concentration also  is accomplished by
oxidizing  the weak, black  liquor with air or oxygen  before evaporation (see Chapter 9).
Black liquor from rotary drum vacuum  filters will evolve less odor from diffusion washers,
since measurable oxidation of the sulfur content occurs in the filters (3).
                                         3-1

-------
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                               TEMPERATURE   =  90-I30°C
                               PRESSURE       =  600-1200 mm Hg
                               pH               =  12
                                •  DIRECT  MEASUREMENTS
                                O  INDIRECTLY  CALCULATED
40
      Na2S  IN  LIQUOR, g  Na2S/Kg  LIQUOR

                                    FIGURE 3-1
      VAPOR-LIQUOR  EQUILIBRIUM FOR  H2S OVER BLACK  LIQUOR  (1,2)

 The evaporation plant odor release will also depend on black liquor temperatures and pH,
 but these factors are fixed by process conditions.

 The amount of  noncondensable odor gases from the evaporation plant will be greater if the
 black liquor originates from hardwood digestion than from softwood (3).

 3.2 Evaporator Types

 There are different ways of evaporating spent black liquor indirectly with steam. The most
 important ones, from an air pollution point of view, are discussed in the following sections.

     3.2.1  Multiple-Effect Vacuum Evaporation

 This is the dominant evaporation system. It can have many stages, usually 3 to 7. Each stage
 may  have  multiple  bodies. It  is  normally  equipped with  condensate  flashing, liquor
 preheating, hot  water generating, degassing, tail steam condensing and vacuum generating
 systems.  Falling film, rising film, or forced circulation evaporation can be used. The most
 important features of condensate and gas treatment are shown in a generalized flow sheet in
 Figure  3-2. The flows of condensate from the different stages are shown separately (flows
 2-5 in  Figure 3-2) to illustrate condensate treatment  (section  5.3.2). Actually, they are
 flashed in series through stages 3  to  5.  Stages 2 to  5 are vented  to the vacuum system and
 stages 4 and 5 are equipped with liquor preheaters that serve as vent vapor condensers. The
                                        3-2

-------
           20
          22
                                                                                                       18 WARM WATER
                                                                                                       17 FRESHWATER
                                                                                                       16 FRESHWATER
             STEAM
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                        LIQUOR
 WEAK
LIQUOR
 (12)

SOAP
                          I
                          @
                         SOAP
                                                                                       (6+7+81)

                                                                                     TAIL CONDENSATES
                                 POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY

                                                          FIGURE 3-2
                                  MULTI-EFFECT VACUUM EVAPORATION PLANT FLOW SHEET

-------
tail-end condensing  systems is  shown  with a surface  primary condenser, a generalized
secondary condenser (section 3.2.1.2), and a generalized vacuum device (section 3.2.1.3).
There are many  different ways of feeding and circulating the liquor, but the one shown is
very common.

The odorous evaporator gases are usually released from the vacuum device (flow 9 in figure
3-2) through a duct from the hotwell.

The odorous evaporator condensates are usually released as secondary condensates (flows 2
to 5 in Figure 3-2) and as tail condensates from the hotwell (flows 6 to 8 in Figure 3-2). To
some extent, a trade-off is achieved between the release of odorous gas and condensate.

The range of flow and composition of the evaporator  gases from the  hotwell is given in
Tables 1-3 and  1-2,  respectively. A  discussion of some of the factors  that influence  the
gaseous emission and the condensates from the hotwell follows.

         3.2.1.1  Evaporator Effect—Venting

The evaporation  plant noncondensables generated in each effect must be vented so that they
cannot reduce the condensing heat  transfer coefficient nor  increase corrosion  inside  the
plant. Venting can be  done in several ways, but the two principal systems are single-stage
venting and two-stage venting.

Single-stage  venting is usually done by bleeding off a certain adjusted amount of vapor from
each effect to a central vent duct. In a vacuum evaporation plant, the  vent duct usually goes
to the  secondary condenser, then to the vacuum pump, and the gases are vented  out of the
system from the  hotwell or from  the steam jet ejector.

When  the  weak  liquor  is fed  into  the evaporation plant, there  is  a large boil-off of
compounds  with high vapor pressure in the feed effect (Number 3 in Figure 3-2). This vapor
condenses in the next stage leaving  large amounts of noncondensable vapor in that vapor
space. Even in the next liquor evaporation,stage, boil-off is substantial, and noncondensables
are carried over to  the following vapor space.

The application of two-stage venting can be advantageous. The vapor spaces of the first one
or two stages are vented following the liquor feed stage. These vapor flows amount to 5 to
15  percent of the  total vapor flow and are vented to separate condensers. These condensers
can serve as preheaters for liquor or  water and will, in their turn, be vented to the common
duct leading to the vacuum system. In this  way the heat transfer  coefficients of  the
evaporator surfaces stay high, and the vapor is condensed in two fractions (for instance,
flows 5a and 5b in Figure 3-2). The fraction "a" will be large (say 85 percent of the vapor),
but will contain only  around 40 percent of the  biochemical oxygen demand (BOD) and
                                         3-4

-------
sulfur compounds. The fraction "b" will be small (about 15 percent of the total), but will
contain  the  remaining 60  percent of the BOD and sulfur  compounds. This will greatly
facilitate subsequent treatment of the condensates (section 5.3.2).

         3.2.1.2  Evaporator Condensers

Evaporator condensers are mainly of two types, surface condensers and direct contact
condensers. Both will also be barometric condensers (i.e., they are equipped with barometric
legs or tail pipes).

The direct contact condenser, shown in Figure 3-3, is small in size, efficient, hard to plug or
foul,  and  inexpensive.  It will,  however,  add  water to the  condensates and more
noncondensable gases  will dissolve in this water depending on  the  water  temperature.
Therefore, there is a larger  volume of contaminated water to treat at correspondingly larger
costs (section 5.3.1). High fresh water temperatures in the summer will require large water
                 DIRECT CONTACT
                 CONDENSER
                        A
                LAST
                STAGE
         VAPOR 1
         COND. 2
                                       HOTWELL
                                                      8 FRESH WATER
                                                         VENT
                                                '  WATER RING
                                                I  VACUUM PUMP
                                                I
                                                I
                         SECONDARY     TAIL
                         CONDENSATE    CONDENSATE
                     POINTS OF POSSIBLE ORDER  RELEASE ARE ENCIRCLED BY  Q
                                   FIGURE 3-3
          EVAPORATION PLANT DIRECT CONTACT CONDENSER WITH
                               WATER RING PUMP
                                       3-5

-------
flows to maintain  sufficiently low pressure in the evaporator plant tail end. On the other
hand, there will be  less noncondensable gases to treat.

One interesting variation of a water ring pump circuit of an indirectly cooled condenser is
one  in which white liquor is circulated instead of contaminated condensates. The cooled
white liquor scrubs the noncondensable gases and absorbs H2S and CH3SH, first in the jet
condenser, and then in the water ring vacuum pump. Fresh white liquor is supplied, and the
overflow is returned to the process. Using this circuit  saves the investment in  an  entire
scrubber (4).

          3.2.1.2  Surface Condensers

The surface condenser has some advantages over the direct contact condenser. First, it does
not generate more condensate than the vapor condensed. Furthermore, it can be used to
produce clean warm water. It is also easier to control. Condensation  can be controlled at the
desired degree of subcooling of the condensate. Controlled temperature subcooling is used
quite efficiently when the condensation is divided between one primary and one secondary
surface condenser (Figure 3-4). One can design and dimension the condenser and control the
cooling  water flows so finely  that about  85  percent  of the vapor  from the  last stage will
condense  in  the primary  condenser without any subcooling anywhere  on the condenser
surfaces. Lack of subcooling will allow minimum  dissolving of noncondensable gases and
low boiling organics in the condensate, and the condensate will be  quite low in sulfur and
BOD and can be reused without further treatment. The remaining 15 percent of the vapor is
sent to the secondary condenser, where it is condensed and subcooled as much as feasible.
In this way, a substantial enrichment of noncondensable sulfur compounds and low boiling
organics occurs in  this condensate, which may then be treated (e.g., by steam stripping). By
applying the concept of partial condensation in two steps (section 3.2.1.1), one condensate
stream can be divided into a large, rather clean part and a small, highly contaminated part.
The small part can be treated at significantly reduced cost. The importance  of this  design
concept will become more evident in the discussion of condensate treatment (see  section
5.3.2).

          3.2.1.3  Evaporator Vacuum Pumps

Two  main types of vacuum devices are  in common  use, the  water  ring type of vacuum
pump, which is typical in Scandinavian countries, and the steam jet ejector type, which is
typical  in  North America. In some installations other  types  are used, such as  water jet
ejector vacuum pumps. All vacuum pumps that allow noncondensable gases to contact fresh
cooling water will  shift part of the odorous components from the hotwell vent gases to the
hotwell condensates.
                                         3-6

-------
               PRIMARY
               SURFACE
               CONDENSER
          LAST
          STAGE
   VAPOR 1
   COND. 2
                                                            9  WARM WATER
                                                            8  FRESH WATER
                                                  SECONDARY
                                                  SURFACE
                                                  CONDENSER
                                                               VENT
                                                     WATER RING
                                                     VACUMM PUMP
                 SECONDARY    SURFACE      SECONDARY SURFACE
                 CONDENSATE  CONDENSATE   & PUMP CONDENSATE

                    POINTS OF POSSIBLE ORDER RELEASE ARE ENCIRCLED BY
                                  FIGURE 3-4
           EVAPORATION PLANT TWO STAGE  SURFACE CONDENSER
                           WITH  WATER  RING PUMP
         3.2.1.4  Evaporator Flash Steam Feed

One  source of primary steam  for the evaporation plant is the secondary flash tank of a
continuous digester.   Depending on  the  flash steam  capacity  and its  pressure and
temperature, it may meet all or part of the steam demand of a particular evaporation plant.
In the latter case, it may be supplied to the first evaporator effect or to a alater one. From
an air and water pollution point of view, the important thing is that all noncondensable and
low boiling compounds flashed from the spend liquor in the flash tank pass over to the
evaporation plant and add to the noncondensable gas and contaminated condensate released
there. This release will be richer in organic compounds and turpentine than ordinarily is the
situation in the evaporation plant. Thus, sulfur recovery with white liquor scrubbing will  be
less effective,  and more caution must be observed when treating the noncondensable gases
(sections 4.1.3 and 4.2.3).
                                       3-7

-------
          3.2.1.5  Condition of the Evaporator

According to Table 1-2, the amount of gases from the hotwell may vary between 0.3 and 12
m3  per metric ton of pulp (10 and 400 ft3/ton). A reasonable value for softwood would be
around 1  m3 of noncondensable gases  per metric ton of pulp (32 ft3/ton). If much larger
values  occur, the vacuum portions of the evaporation plant should be inspected for air leaks.

     3.2.2  Other Evaporation Types

          3.2.2.1  Multiple-Effect Back Pressure Evaporation Plant

The multiple-effect back pressure system (Figure 3-5) is similar to the vacuum evaporation
plant, but uses a higher feed steam and tail steam pressure.

The tail steam is then used in some other process in  the mill. Most of the noncondensable
gases, both sulfurous and organic, will be carried with the  tail  steam and vented elsewhere,
but  can be  collected for treatment at  that point. The higher pressures in this evaporator
stage entail higher temperatures that cause more noncondensables to  evaporate and also
cause  a greater  risk of  liquor turnout on hot spots.  Because of this risk, this type  of
evaporator is currently not in widespread usage.

          3.2.2.2  Flash Evaporation Plant

The flash evaporation column (Figure  3-6) is a fairly recent type of evaporator for  black
liquor  evaporation. In the column, the liquor is  passed down and flashed in stages under
decreasing pressure. The flash steam is used to  preheat the liquor in stages on its way up to a
new cycle of flashing. The evaporator includes a vacuum condensation system and operates
with complete crosscurrent flow. The gases vented from the vacuum system are very similar
to those from an ordinary multistage vacuum evaporation plant.

          3.2.2.3  Thermocompressor Evaporation Plant

A thermocompressor  evaporation  plant  (Figure 3-7)  is  a single-stage  evaporator with
multiple bodies, through which the liquor is passed in stages. Pressures and temperatures are
rather uniform over  the  whole plant, the compressor furnishing the difference of about 15°
C (27° F). All stages are  vented through a liquor preheater.  Vented gases are similar to those
of a back-pressure evaporation plant.

3.3  Evaporator Gas Scrubbing

Although  noncondensable gas  treatment  will  be  extensively  reviewed in  Chapter  4,
evaporator gas scrubbing is discussed here because this equipment can be integrated with the
evaporation plant in different ways.

                                          3-8

-------
co
              STEAM
               20
                           A
A
A
Al
                              POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                                                                           9) VENT
19  BACK PRESSURE
   STEAM

T

....I

fe
w
T



b

T

I

^
P
I
PREH EATER
T
7
T
i
	 i
«-l
w
P
	 fc.
                                                                         (13) SECONDARY CONDENSATES

                                                                         (lO) WEAK LIQUOR

                                                                         frfr THICK LIQUOR
                                                   FIGURE 3-5

                        MULTIPLE EFFECT BACK PRESSURE EVAPORATION PLANT FLOW SHEET

-------
                                                    STEAM
                                                 22
                                •J-—©—*   ©VENT
                                                    THICK LIQUOR

                                                    WEAK LIQUOR
                                           ->   (13) SECONDARY CONDENSATES
             POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                                   FIGURE 3-6
                   FLASH EVAPORATION PLANT FLOW SHEET

Evaporator gas scrubbing with white liquor in a direct contact condenser of vacuum pump
was discussed previously (see section 3.2.1.2). It is also possible to install a scrubber after
the condenser  and before the vacuum pump. The Venemark-design white liquor scrubber
has been used in such an application (5). A small scrubber can be installed over the hotwell.
Such a  scrubber design  is shown in Figure 3-8. Noncondensable gases pass through the
packed column which is washed with white liquor sprays.
                                      3-10

-------
                                                             COMPRESSOR

                                                               21  POWER

                                                               20  MAKE-UP STEAM

                                                               Tl)  THICK LIQUOR

                                                                   VENT

                                                               101  WEAK LIQUOR
                                                               13) SECONDARY
                                                                   CONDENSATES
                                                PREHEATER
           POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY Q
                                  FIGURE 3-7
           MULTIPLE EFFECT SINGLE STAGE THERMOCOMPRESSOR
                      EVAPORATION PLANT  FLOW SHEET
3.4 General Evaporator Air Pollution Abatement Programs

General programs for abating the air pollution of an evaporation  plant include suitable
combinations of the following actions:

    1.   Checking of plant condition, in general, especially for possible leaks;

    2.   Installing condensers for two-stage venting;

    3.   Installing surface condensers for two-stage condensing;

    4.   Collecting contaminated condensates for treatment, such as steam stripping;

    5.   Installing a  direct contact condenser or using vacuum pump scrubbing with white
         liquor;

    6.   Installing a suitable white liquor scrubber for hotwell gases;
                                       3-11

-------
  SPRAY
 NOZZLES^
                SCRUBBED GASES

                      A

                       «	 I.D.  150 mm (6")
             600mm
             600mm,
600mm
             600mm
             600mm
             600mm
                   /  \

                   I.D. 50mm
                   (2")
      WHITE LIQUOR
        RETURN
                      I.D. 600mm (24")

                      MIST ELIMINATOR



                      25mm (l")  RASCHIG  RING
                                             I.D. 150mm (6")
                           HOTWELL  GASES
                           MAX. FLOW - lOOOmVhr
                                     (590ft3/min.)
                             COOLED WHITE LIQUOR
                             MAX. FLOW - 5O mVhr
                                       (3Oft3/min.)

                          FIGURE 3-8

HOTWELL GAS SCRUBBER  FOR 100  METRIC TONS PER HOUR (440 GPM)
   EVAPORATION PLANT  FOR H2S - SEPARATION OF 95% OR MORE
                             3-12

-------
     7.    Collecting hotwell gases and incinerating them; and

     8.    Installing a weak black liquor oxidation system.

3.5  In-Plant Controls

Malodorous sulfur gases can be released from black liquor during multiple-effect evaporation
by  liquor heating and by the stripping action of the steam (6).  The H2S release during
multiple-effect evaporation is influenced by the inlet Na2S concentration, the black liquor
pH,  and the type  and degree of treatment in  the weak black  liquor upstream  of the
evaporators (7).  The potential liberation of organic sulfur compounds is influenced by their
concentration in the incoming  black  liquor which varies with wood species, the liquor
temperature,  and  the  degree  and  type  of treatment upstream  of  the multiple-effect
evaporators.

Major variables  upstream of the multiple-effect evaporators which can  affect sulfur gas
emissions include the  wood  species  pulped,  the  pulping  conditions, the type of  pulp
washing, and the possible use of weak black liquor oxidation.

The organic  sulfur  emissions from pulping certain  hardwood species are greater than for
most softwoods, particularly at high white liquor sulfidity levels. Vacuum drum washing of
pulp results in a stripping or organic sulfur compounds and in oxidation of a portion of the
Na2S. Diffusion washing is done in the absence of air and does not involve oxidation of the
Na2S or evolution  of the organic sulfur gases. Inlet liquor concentrations of sulfur com-
pounds  to the multiple-effect evaporators are expected to be higher following  diffusion
washing of pulp than following drum washing.

The pH of the black liquor is an important variable affecting the liberation of H2S and to a
lesser extent CH3SH. Both gases are  slightly acidic in nature, with greater ionic dissociation
in aqueous solution favored by  increased pH.  Increasing the black liquor pH  above 12.0
helps to reduce  H2S emissions, lignin  precipitation as a cause of evaporator plugging, and
the tendency for evaporator scaling and corrosion (8) (9). Addition of caustic soda to weak
black liquor in controlled quantities can raise the pH to the required levels.

Weak black liquor oxidation with either air or oxygen can reduce sulfur gas emissions from
multiple-effect evaporator noncondensable  gases.  Reid  (10)  and  Galeano  (11)  report
reductions in H2S emissions of  70 and 99 percent from  multiple-effect evaporators  after
weak black liquor oxidation wjth air and oxygen, respectively.

Malodorous  sulfur  compounds  emitted  from  the black  liquor during multiple-effect
evaporation must end up in  either the  noncondensable gas stream or the condensate liquid.
The  type  of condenser employed  has  a  definite effect on  the distribution of sulfur
compounds between these two. Because of its scrubbing action, the use of the barometric
                                        3-13

-------
jet condenser results in a greater portion of the sulfur compound emissions ending up in the
condensate rather than in the gas stream, as shown in Table 3-1 (6).
                                  TABLE 3-1
      EFFECT OF CONDENSER TYPE ON REDUCED SULFUR GAS EMISSION
              FROM EVAPORATOR NONCONDENSABLE GASES (6)*

  Condenser Type      H2S     CH3SH     CH3SCH3     CH3SSCH3     Total
                     kgS/t     kgS/t       kgS/t         kgS/t        kgS/t

  Surface             2.28       0.49         0.09          0.21          3.07

  Barometric          0.06       0.07         0.05          0.01          0.19

                    Ib S/ton   Ib S/ton      Ib S/ton       lb S/ton      Ib S/ton

  Surface             4.50       0.97         0.18          0.42          6.07

  Barometric          0.12       0.13         0.10          0.02          0.37

  *kg S/t = kilograms of sulfur per metric ton of air dried pulp
  lb S/ton = pounds of sulfur per short ton of air dried pulp


3.6  References

 1.  Venemark, E.,  Black  Liquor  Evaporation, Part 2. Svensk  Paperstidning, 61 (20):
    881-887, October 31, 1958. (Swedish).

 2.  Arhippainen, B. and Jungerstam, B. Kraft Liquor Evaporation. In: Proceedings of the
    Symposium  on Recovery of Pulping Chemicals. Helsinki, Finland, May 13-17, 1968.
    Finnish Pulp and Paper Research Institute and EKONO Oy, Helsinki, Finland,  1969, p.
    132.

 3.  Sarkanen, K. V., Hrutfiorod, B. F., Johanson, L. N., and Gardner, H. S., Kraft Odor.
    Tappi, 53: 776-783, May 1970.

 4.  Ronnholm, A. A. R., Reducing Evaporation Plant Pollution and its Treatment. Paperi
    ja Puu, 54 (11): 715-730, 1972.

 5.  Swedish patent 226 789, Stockholm, Sweden.
                                      3-14

-------
 6.  Hendrickson, E. R., Robertson, J.  E., and Koogler, J. B.,  Control of Atmospheric
    Emissions in  the Wood Pulping Industry, Vols.  I, II, III. Final Report Contract No.
    CPA  22-69-18, U.S. Department of Health, Education, and  Welfare, National Air
    Pollution Control Administration, Raleigh, North Carolina, March 15, 1970.

 7.  Douglass, I. B., Sources of Odor in the Kraft Process: Odor Formation in Black Liquor
    Multiple Effect Evaporators. Tappi, 52: 1738-1741, September 1969.

 8.  Berry, L. R., Black Liquor Scaling in Multiple Effect Evaporators. Tappi, 49: 68A-71A,
    April 1966.

 9.  Cry, M. E., and Harper,  A. M., Multiple  Effect Evaporator Project. Pulp  and  Paper
    Magazine of Canada, 61: T247-T249, April 1960.

10.  Reid, H. A., The Odor Problem at Maryvale. Appitta, 3(2):479-500, December 1949.

11.  Galeano, S. F.,  and Amsden, C. D., Oxidation of Kraft  Weak Black Liquor with
    Molecular Oxygen. Tappi, 53:  2142-2146,  November 1970.
                                       3-15

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                                    CHAPTER  4

                      NONCONDENSABLE GAS TREATMENT
Digester and evaporator noncondensable gases characteristically have relatively low volume
flow  rates  and  high  malodorous  sulfur  compounds  concentrations. Organic  sulfur
compounds, such as  CH3SH, CH3SCH3, and CH3SSCH3, are  emitted  from digesters in
varying quantities from  0.25 to 2.5 kg of sulfur per metric ton of pulp (0.5-5.0 Ib sulfur per
ton of pulp). The primary sulfur compounds that are emitted from multiple-effect evapora-
tors are H2S and CH3SH in quantities ranging from 0.1 to 1.5 kg of sulfur per metric ton of
pulp (0.2-3.0 Ib sulfur per ton of pulp). Unless properly controlled, these gas streams can
cause intense odor pollution at low elevations near the mill.

4.1  Gas Stream Characteristics

The noncondensable gases from digesters and evaporators are relatively high concentration,
low volume streams. Their odor levels are easily reduced by chemical or thermal oxidation.
Major  parameters  of  the  gas stream  which  must  be  considered in  the design  of
noncondensable gas handling and treatment systems include temperature,  moisture content,
flow rate and variability, sulfur gas  concentrations, organic material concentrations,  and
flammability limits.

     4.1.1   Process Sources

The  major noncondensable gas streams collected and treated in kraft pulp  mills are the
blow  and relief gases from batch and continuous digesters, and the hotwell and condenser
vents from multiple-effect evaporators.  The gas collection systems for kraft pulp mills must
be individually designed  to connect all the various gas sources.

Batch  digesters  normally make up  the largest  single volume source and give  rise  to the
greatest variations in flow rates of any  of the noncondensable gas streams. Batch digesters
are normally vented through the  relief system at the  condenser vent and the turpentine
decanter  vent.  The  relief gases are low volume gas streams that flow on a more  or  less
continuous basis during  the 3- to 5-hour cooking period. The relief gases normally contain
large amounts of terpenes, in addition to the sulfur compounds released during the cook,
and can pose an explosion hazard.

The batch digester is normally vented to the blow tank at the end of each kraft cook over a
10- to 20-minute period, with  the resultant release of large quantities of steam, inert gas,
organic compounds, and malodorous sulfur compounds. The volume  of gas released depends
on the digester volume, the  gas temperature and  moisture content, the degree of vapor
                                       4-1

-------
condensation effected, the amount of air volume in the system, and the amount of inert
gases present in the digester. The volume of gas to be handled occurs in a large single surge
at the end of each cook. This feature normally requires the use of a flow equalization device
to avoid upsetting the operation of the burning device and to assure maximum thermal
operating efficiency. The  use  of sufficiently-sized blow heat condensers may obviate the
need for an equalization device.

Continuous  digesters normally are nearly constant flow rate devices (except during periods
of process upset), requiring no flow equalization devices. Continuous digesters are normally
vented at the steaming vessel relief, at both  condenser and turpentine decanter vents, from
the blow tank vent after the condenser system  and sometimes from the top of the digester
unit itself.  The exact venting system arrangement varies  among individual digesters. The
total amounts of gas and the terpenes and organic sulfur compounds liberated are  not
normally as large for continuous as for batch  digesters.  Also,  the amounts are  somewhat
dependent on the digester blow temperatures. The result is that the point of emission for
these materials is often transferred from the digester  system to the brown stock washer and
multiple-effect evaporator  sections.

Noncondensable gases from multiple-effect evaporators differ in character from  vent gases
from  digesters in that they  contain  smaller  amounts  of  terpenes and organic sulfur
compounds. The gases are made up primarily of H2S and  CH3SH liberated from the black
liquor  during the evaporation  process.  Evaporator noncondensable gases are  normally
collected from the hotwell and condenser vents and can vary between mills. Normally, the
noncondensable  gas  flow rates  are considerably larger  from indirect  contact (surface)
condensers than from direct contact (jet) condensers.

    4.1.2   Flow Rates

The gas flow rates  for individual process  streams are subject to wide variations among
individual mills, depending on production rate, process operating variables, and the degree
of  condensation  for heat recovery. The  single most  important  flow rate variable for
noncondensable gas streams is the peak flow rate for batch digester blow gases, particularly
during periods of condenser  malfunction. Some  type of pressure relief system  for batch
digester  blow gas systems should be provided during periods of condenser  malfunction,
which  are usually indicated by gas  temperatures well above normal. A summary  of  one
mill's gas flow rates from  a batch digester to a  gas holder under conditions of average flow,
maximum flow, and condenser upset is presented in Table 4-1.

Major variables affecting the overall gas flow rates from noncondensable gas streams are the
process unit type and operating conditions,  the production capacity for the particular unit
from which the stream is  vented, and the degree and type of vapor condensation employed.
A summary  of typical ranges for noncondensable gas flow rates is presented in Table 4-2.
                                         4-2

-------
                    TABLE 4-1
       GAS FLOW RATES FROM A BATCH
                 DIGESTER* (1)**
  Operating
  Condition
 Average
 Maximum
 Upset
   Digester Blow
       Basis
    45   (1,600)
   113   (4,000)
  3,540   (125,000)
  Pulp Produc-
    tion Basis
              m3/blow  (ft3/blow)    m3/t  (ft3/ton)
    4 (128)
   10 (320)
  320 (10,250)
  *Gas flows at actual stack conditions.
 **Pulp cooking capacity is 11.4 t/cook (12.6 ton/cook).
                   TABLE 4-2
      TYPICAL  RANGES  IN DIGESTER AND
  EVAPORATOR NONCONDENSABLE GAS FLOW
                     RATES
  Source
  Process Stream
Digester
Batch Blow
              Batch Relief
              Continuous
Evaporator     Surface Condenser
              Jet Condenser
   Flow Rate
     m3/t
    (ft3/ton)

   475-6,350
(15,200-203,500)

     0.3-95
   (10-3040)

     0.6-6
    (20-200)

     0.6-13
    (20-420)

     0.3-3
    (10-100)
                       4-3

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Flow rates of noncondensable gas streams to subsequent treatment devices following flow
equalization normally range from 5  to 35 m3/h (3 to 21 cfm) with average values of 10 to
15 m3/h (6 to 9 cfm) reported (2). Higher flow rates are reported when there is insufficient
heat exchanger condensation capacity, when flow equalization is not employed, or when
there is significant air leakage into the gas handling system.

     4.1.3   Safety Considerations

The design and operation of a system for thermal oxidation of the noncondensable gases
require measures to prevent explosions. Four factors that must be considered  are variations
in  gas  flow  rate,  passage of  entrained  moisture droplets  into  the  burning  device,
flammability limits  of sulfur and organic compounds in the gas stream in the inlet piping,
and flame propagation speeds as compared to gas flow velocities in the piping. Of particular
importance is the presence of terpene compounds in the noncondensable gas handling and
burning systems.

Excessive variations in flow rates of noncondensable gases entering the air inlet of  a burning
device can blow out the flame. An additional factor which should be considered is that the
net retention time of the gas  stream in the burning device is reduced during flow surges.
Solutions to this problem are to employ large surge tanks with large condenser capacities, to
use flow equalization devices, and to provide  pressure relief vent systems.

Entrained moisture droplets can create hazards when entering the combustion  zone. First,
the droplets can cool the flame and may even extinguish it. Second, the evaporation of
water can result in a large increase in gas volume as the water changes to steam. These surges
in volume can cause an unstable  operation in the burning device and  lead to a flameout. A
flameout may allow  an  explosive mixture  of flammable materials to accumulate in the
combustion unit, resulting in an explosion when the system is reignited.

Explosive limits must be  considered in the design of noncondensable gas handling systems,
both  before and  after dilution  with primary air. A summary  of flammability limits for
materials commonly present in noncondensable gas streams is listed in Table 4-3  (3).

Terpenes have the lowest explosive limits of any of the compounds listed and are normally
the most critical component in  noncondensable gas streams for purposes  of design for
explosive safety.

Work by Ginodman (4),  Coleman  (5), and DeHaas and Hansen (6) indicates  that it is
necessary to dilute  noncondensable gas streams by a sufficient amount in the primary air
inlet before they enter combustion devices. Dilution must occur at a fast enough rate so that
the nearly oxygen-free noncondensable gas stream can pass the lower explosive limits of the
most critical  material (normally terpenes) without  an explosion.  A  summary of  the
                                        4-4

-------
                                  TABLE 4-3
               FLAMMABILITY LIMITS IN AIR FOR COMPOUNDS
                 PRESENT IN KRAFT NONCONDENSABLE GAS
                                  STREAM (3)

                                           Flammability Limits
                Material                Lower               Upper
                                         Concentration, % by vol.

               H2S                      4.3                45.0
               CH3SH                   3.9                21.8
               CH3SCH3                 2.2                19.7
               Terpene                   0.8
 necessary  dilution  requirements  for batch digester relief gas alone, and for combined
 digester blow and relief gas is presented in Table 4-4.

 The primary danger of explosions exists in the primary air inlet of the combustion device
 immediately following introduction of  the noncondensable  gases. The presence of large
 quantities  of terpene compounds tends to make digester relief gases a greater explosive
 hazard than blow gases.
                                   TABLE 4-4
                DILUTION REQUIREMENTS WITH AIR  TO AVOID
                EXPLOSIONS FOR DIGESTER NONCONDENSABLE
                           GAS STREAMS  (4) (5) (6)

                Gas Stream                 Volume Dilution Required
                                               Air/Gas Ratio

               Relief only                           50/1

               Relief & blow                         20/1
An  additional consideration regarding potential explosions involves  flame  propagation
speeds of air-gas mixtures. Data collected by Ghisoni (7) on flame propagation speeds for
air-mercaptan mixtures are listed in Table 4-5.
                                     4-5

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                                    TABLE 4-5
                    FLAME PROPAGATION SPEEDS FOR  AIR-
                           MERCAPTAN MIXTURES (7)

               Mercaptan Concentration                Flame Velocity
                      % by Vol.                        m/s (ft/sec)

                         18.9                          0.55 (1.8)
                         22.8                          0.46 (1.5)
                         23.1                          0.40(1.3)
                         23.7                          0.37(1.2)
                         25.5                          0.18 (0.6)
                         25.7                          0.15(0.5)

Gas velocities in the noncondensable gas piping and the primary air inlet must be greater
than any flame propagation speeds to prevent  damage to  process units  from  possible
explosions. Maintenance of gas velocities of at least 1 m/s (3 ft/sec) at all times and the use
of flame arrester devices in the noncondensable gas line should minimize the danger of
explosions from excessive flame speeds.

4.2  Gas Handling Systems

The  major  parts  of a  noncondensable gas handling  system are condensation,  flow
equalization, liquid scrubbing, piping, safety control, and air inlet sections.

Removing moisture and turpentine upstream of the burning device is particularly important
in preventing possible flameouts. Sulfur compounds can be absorbed into alkaline liquids for
white liquor makeup. The noncondensable gas piping system must be designed with safety
devices to prevent explosions and must also have design  features to permit the rapid dilution
of the noncondensable gas stream before it enters the combustion unit.

     4.2.1   Vapor Condensation

The  primary purpose of the condensers in digester and evaporator gas handling systems is
heat recovery from the gas stream by water vapor condensation.

The  condenser  systems also remove a portion of the organic vapors, such as terpenes and
sulfur compounds. The vapors so removed also require,  in turn, treatment of the condensate
waters for odor removal. The terpenes can be recovered by flotation and  decantation from
the condensed water. After collection, they can be sold as byproduct turpentine or used as
an auxiliary fuel in the lime kiln.
                                        4-6

-------
The condensers also serve  to reduce the volume of gas to be treated by cooling the gas
stream and thereby condensing out a portion of the contained water vapor. Cooling the gas
stream to between 75 and 85° C  (167  and 185° F)  is normally desirable to remove the
major portion of the water and reduce the gas  volume. Where no flow equalization devices
are employed, the use of large heat transfer  surface  areas  in condensers is particularly
important for blow gas streams from batch digesters. Surface condensers are more desirable
from a water pollution reduction standpoint than comparable spray or jet condensers since
they produce a lower volume of more highly concentrated contaminated  condensate.  This
type of condenser is normally larger in surface area  and, therefore, has a higher capital cost.

    4.2.2   Flow Equalization

The  two  major devices employed  for noncondensable gas flow  equalization are  the
vaporsphere and the floating cover gas holder. DeHaas and Hansen (6) describe the use of a
vaporsphere for collection of batch  digester noncondensable gases.  The  vaporsphere is a
spherical device with a flexible fabric diaphragm attached around the epicenter of the sphere
as shown in Figure 4-1 (2). The diaphragm consists of a mylar film sandwiched between two
layers of cotton canvas that can be fabricated by local tent and awning manufacturers. The
useful life of such mylar canvas diaphragms is as long as 27 months (8).
      Diaphragm
Sliding
Weight
    Vacuum
   Pressure
     Relief




[Off
On



»
•


Flow
Control


         Relief  Gases
                                   FIGURE  4-1
           VAPORSPHERE  FLOW  EQUALIZATION GAS  HOLDERS  (2)
                                       4-7

-------
Several operating and safety features are necessary to assure safe and reliable operation of
the vaporsphere.  The system contains a counterweight connected to the diaphragm which
moves up and down as the gas flows  in and out of the vaporsphere. The diaphragm  is
weighted to provide a slight positive pressure on the system at all times. Automatic flow
controls  on the  system  are  used  to prevent damage to the diaphragm and to prevent
excessive  air  leakage. A pressure and  vacuum relief system  prevents damage  to the
diaphragm from either excessive gas flows from  condenser malfunctions or  from excessive
suctions. The condenser section and all associated piping to the vaporsphere must be sealed
to prevent the possibility of air leakage and a potential explosive mixture from forming (6).
Also, the  inlet  and  outlet  gas streams must  be vented  separately for  the full flow
equalization effect to be obtained from the vaporsphere.

Floating  cover gas holders also are used for flow equalization in kraft pulp  mills. They
consist of two vertical cylindrical tanks with the upper located inside the lower and a water
seal to prevent gas leakage, as shown in Figure 4-2.  Gas enters the inside cover cylinder
above  the  water  seal, causing the cover to be displaced upward when gas enters and
downward as it exits. Gas is withdrawn  from a separate exit pipe to achieve the full degree
of flow equalization for the system.

The  floating cover gas holder must have several  operating and safety features to assure its
reliable operation. A pressure and vacuum  relief system  is added to prevent  the shell
cylinder  from being damaged or dislodged by excessive pressures or  vacuums. These are
         Blow
By-pass
 Vent
                                                          Vacuum
                                                         Pressure
                                                          Relief
                                                          System
                                                                Handling
                                                                 System
                                    FIGURE  4-2
          FLOATING  COVER  FLOW  EQUALIZATION GAS  HOLDERS (2)
                                        4-8

-------
 connected to a hydrostatic water seal with an overflow to the drain to prevent the water seal
 from being lost. A bypass vent allows venting of excessive gas surges such as those caused by
 condenser malfunctions. Ping pong balls are  placed in the space above the water  seal
 between the shell and cover cylinders to maintain alignment and smooth operation.

 Vaporsphere and floating cover gas holders can  normally  be constructed  of mild steel.
 Capital costs for these systems, including all appurtenances, are about $20,000 to $35,000.
 A summary of the dimensions and construction materials for flow equalization devices is
 presented in Table 4-6 (2).
                                    TABLE 4-6
              DIMENSIONS OF FLOW EQUALIZATION DEVICES* IN
                   KRAFT NONCONDENSABLE GAS  HANDLING
                                   SYSTEMS  (2)

                      Type of    	Gas Holding Dimensions
             Mill     Unit**
             A         VS
             B         VS
             G         VS
             H         VS
             E         FC

              *Mild steel used at all five mills as material of construction.
             **VS = vaporsphere; FC = floating cover.


Values for digester blow gas design flows for sizing flow equalization devices are reported by
Blosser and Cooper (2) and DeHaas and Hansen (6). A summary of digester blow gas volume
flows for varying operating conditions is listed in Table 4-7.

     4.2.3   Liquid Scrubbing

Liquid scrubbing of the noncondensable gas stream is added for purposes of organic mist
removal,  gas stream cooling, and sulfur recovery. DeHaas and Hansen (6) report that a liquid
scrubber  is needed to prevent turpentine mist droplets from reaching the burning device and
causing periodic flameouts. The contact of the noncondensable gas stream with sufficiently
cool scrubber liquor results in additional cooling of the gas stream, reduces its volume, and
removes additional water vapor.
Diameter
m (ft)
8.2 (27)
8.2 (27)
8.2 (27)
6.6 (21)
8.5 (28)
Height
m (ft)
..
--
-
..
4.6 (15)
Volume
m3 (cu ft)
170 (6,000)
170 (6,000)
170 (6,000)
142 (5,000)
283 (10,000)
                                       4-9

-------
                                                  TABLE 4-7
BATCH DIGESTER BLOW GAS  FLOW RATES FOR SIZING NONCONDENSABLE GAS FLOW  EQUALIZATION DEVICES
  Operating Condition


Normal

Condenser Malfunction

No Heat Recovery
                                   Gas Flow/Digester Blow4
   Average                  Range
             m3/blow (cu ft/blow)
  42  (1,500)

 425  (15,000)
 14-113 (500-4,000)

285-570 (10,000-20,000)
                                                        Gas Flow/Unit Production*
                              Average                Range
                                        m3 /t (cu ft/ton)
 4 (125)
47 (1,500)
  2-5 (70-160)
31-62 (1,000-2,000)
1,130  (40,000)    5,660-18,400 (200,000-650,000)    1,250 (40,000)    625-2,000 (20,000-64,000)
*Gas flows are at stack conditions.

-------
The use of alkaline scrubbing liquids, such as sodium hydroxide (NaOH) solution or white
liquor, results in removal of the acidic sulfur compounds, such as H2S and CH3SH, from the
noncondensable gas for return to the chemical makeup system. It is not possible, however,
to remove significant quantities of organic sulfur  compounds, such as CH3SCH3 and
CH3SSCH3, from  the noncondensable gas stream under normal circumstances. Reduced
sulfur gas emissions from digester gas sources can range from below 0.25 to above 2.5 kg per
metric ton of pulp (0.5 to 5.0 Ib per ton of pulp); evaporator noncondensable gas  sulfur
emissions are normally 0.025 to 0.25 kg per metric ton of pulp (0.05 to 0.5 Ib per ton of
pulp). Sulfur compounds from  the digesters are primarily CH3SCH3 and CH3SSCH3 ; while
sulfur compounds from the evaporators are primarily H2S with lesser  amounts of CH3SH.

The alkaline scrubbers generally used are packed bed scrubbers employing a countercurrent
flow of liquid and gas. The usual packing for these devices is gravel, stone, or one-inch thick
packing rings (2). The liquid solution employed for scrubbing is either caustic soda or white
liquor for the return to the chemical makeup system, or water  for subsequent discharge to
the sewer. A typical scrubber system is illustrated in Figure 4-3 (2).
                                  FIGURE 4-3
              PACKED  BED SCRUBBER FOR  NONCONDENSABLE
                           GAS  HANDLING  SYSTEM
                                       4-11

-------
     4.2.4   Piping Systems

Design of the piping for a noncondensable  gas handling system requires consideration of
such items  as materials for construction, explosion hazard  safety, and gas flow  pressure
drop. Noncondensable gas handling systems are normally constructed of mild steel, but 304
or 316 stainless steel has been used in some  applications to inhibit corrosion. Constructing
noncondensable  gas piping  systems  to  obtain a minimum velocity of  1 m/s (3 ft/sec)  is
usually necessary to  minimize the likelihood of flame propagation through the pipe. High
velocities also result  in  increased pressure drops through the piping, particularly  for lines
longer than  30 m (100 ft). An auxiliary  fan or a  large diameter pipe should be installed to
minimize pressure drop  across the system and to allow use of the inlet draft of combustion
air systems. Most mills  have installed 10 cm (4 in) diameter pipe for noncondensable gas
piping systems; while 7  cm (3 in) and 20 cm (8 in)  have been used in a few mills. The larger
diameter piping  has found its  greatest  use  at  mills employing  continuous  digesters.  A
summary of piping system gas velocities, pipe diameters, and materials of construction is
presented in Table 4-8.


                                    TABLE 4-8
             PIPING SYSTEMS FOR  KRAFT  NONCONDENSABLE  GAS
                             HANDLING  SYSTEMS (2)

           Mill     Diameter     Gas Velocity      Materials of Construction
                    cm  (in)     m/s  (ft/sec)

           A        10  (4)      1.5   (4.8)            Mild steel
           B        10  (4)      1.6   (5.3)            Mild steel
           C        20  (8)      3    (9.9)            316 Stainless
           D        20  (8)     13    (41.7)           304 Stainless
           E        10  (4)      1.7   (5.7)            Mild steel
           F        10  (4)       -    -             Mild steel
           G        20  (8)      7    (23.1)           Mild steel
           H         7  (3)      1    (3.3)            Mild steel


     4.2.5   Safety Considerations

In designing noncondensable gas handling systems, specific safety features are needed  to
assure a minimum explosion hazard, prevent liquid  entrainment, and assure stable operation
of combustion devices.  Condensate traps to remove water  are placed at low points in the
piping system at  intervals of approximately  15 m (50 ft)  and just upstream of the air inlet.
A typical  liquid condensate trap  design is illustrated in Figure 4-4  (2). The packed bed
scrubbers also serve the function of moisture removal to prevent upsets of burner operation,
                                         4-12

-------
                                     •O—1--
                                I
                                       10"
                                     n
                                          \
                                               Liquid
                                    FIGURE  4-4
              LIQUID CONDENSATE  TRAP FOR NONCONDENSABLE
                             GAS HANDLING SYSTEM
or flameouts, by liquid droplet entry to the combustion zone, and to eliminate false signals
to the flameout control.

Flame arresters of the leaf or grid type are commonly added to the noncondensable gas line
to prevent the passage of any flame fronts to the process units. The flame arrester normally
is added to the noncondensable gas line immediately upstream at the point of introduction
to the primary air ducts, as shown in Figure 4-5 (2). One or more additional flame arresters
can be installed in long piping systems to gain further protection from possible explosions.

Two features are added to provide emergency venting of excess gas pressures during possible
explosions. Rupture discs are added to noncondensable gas lines at  approximately 30 m
(100 ft)  intervals. These devices have diaphragm discs  set to explode at certain bursting
pressures. To cope with power failures, an emergency vent release is normally placed in the
noncondensable  line connected to the flameout control  for the combustion  device. A
continuous recording device with input from an orifice meter provides a useful record of
system flow rate.

Design for safe confluence  of the  noncondensable gas  stream with the  primary air to the
combustion device requires consideration of inlet draft,  gas velocity, and physical features.
A  damper is normally  placed in the  primary air  inlet  upstream of the point  of gas
introduction to provide an  inlet vacuum of 2.5 to 7.5 cm H20 (1 to  3 inches of water)
necessary to maintain gas flow. Otherwise,  an auxiliary fan is needed in the noncondensable
gas handling system. The gas is added through a horizontal pipe  placed across the primary
                                        4-13

-------
      Flow  Record
      And Control
 Flame
  Out
Control
                                                                      Damper
                               Flame
                            Arrester
          Primary
          Air Fan
                                  FIGURE 4-5
    SAFETY  DEVICES  FOR NONCONDENSABLE GAS HANDLING SYSTEMS

air duct with holes evenly spaced along the pipe on the downstream side to achieve even
distribution and rapid dilution of the gas to well below explosive limits. Gas velocities in the
primary air inlet duct after mixing with air are normally above 9m/s (30 ft/sec).

4.3 Gas Treatment Systems

The major techniques for treatment of malodorous sulfur gases to prevent their emission to
the atmosphere are thermal oxidation and liquid absorption. The major types of devices for
thermal incineration of noncondensable gases are the lime kiln  and catalytic furnaces, but
limited use is also made of other combustion systems. The liquid  solutions employed to date
are acidic  chlorination bleaching effluent and caustic solutions.  Thermal oxidation in lime
kilns provides a positive means for destruction of malodorous sulfur gases. Liquid scrubbing
can also prove effective as an alternate treatment system or as a  pretreatment technique for
safety considerations or for sulfur recovery.

    4.3.1  Lime Kiln Incineration

Incineration of digester and evaporator noncondensable gases in lime kilns provides a positive
method for odor control, without the necessity of constructing an additional combustion
unit, and  also  allows heat recovery. The noncondensable gases normally are added to the
primary air inlet of the lime kiln at a dilution of at least 50 to 1 and an air inlet velocity of
                                       4-14

-------
at least 9 m/sec (30 ft/sec). A flameout control on the lime kiln is connected to a three-way
emergency bypass vent in case of power failure.

Gases are incinerated in the kiln at maximum temperatures of about 1200 to 1400° C (2200
to 2550° F) to achieve complete oxidation of the  sulfur compounds present. The S02
formed is largely collected  by the lime, as calcium sulfite (Ca2 S03) instead of being emitted
to the atmosphere. To date, mills employing  noncondensable gas incineration have not
observed any  adverse  effects  on either jeburned  lime  quality or causticizing system
operation from the  burning of these gases.

The basic layout of the burning system for incinerating noncondensable gases consists of the
collection  piping and burning sections. Systems where batch digester gases are collected
require a flow equalization device or suitably large  condenser capacity. The layout for a
typical unsteady state system for batch digester blow and relief gases is illustrated in Figure
4-6;  a steady state system for  a continuous digester  is  illustrated in Figure  4-7 (2). A
combination system for gases from both batch  and  continuous digesters, plus  evaporator
gases, is illustrated in  Figure 4-8. The noncondensable gas flow rates, their sources, and
burning devices employed for eight existing installations are presented in Table 4-9.
             Scrubbing
Flow Record
    Control
                                Flame Out
                                 Control
    Flow
Equalization
Auxiliary
   Fan
(Optional)
       Rupture
         Discs
     Liquid
                                                                 Flame
                                                              Arresters
    Tank
          Condenser
          Scrubber
                                  Effluent
                                   FIGURE  4-6
UNSTEADY STATE FLOW SYSTEM  FOR  BATCH DIGESTER  NONCONDENSABLE
                            GAS  INCINERATION (2)
                                       4-15

-------
        Steaming
Vessel
Flash
Tank
Blow
Tank *
Digester
Relief
Evaporator
Vent
Digester _



T
i

now necoraer
\
\
o
1— f_l
"T f\ II i
W "
Auxiliary
Fan (Optio
                        Flame  Out
                          Control
                                          By-pass
                                            Vent
      Condensate
       Aeration
Discs
          Flame
        Arresters
  Cyclone or
Condensate  Traps
                               FIGURE  4-7
        STEADY  STATE FLOW SYSTEM  FOR  CONTINUOUS DIGESTER
                 NONCONDENSABLE  GAS  INCINERATION  (2)
                                TABLE 4-9
  GAS FLOW RATES  TO BURNING DEVICES FROM NONCONDENSABLE GAS
                          HANDLING SYSTEMS  (2)
Mill    Daily Pulp Production
           t (ton)/day

A          363  (400)
B          999  (1,100)
C          499  (550)
D          499  (550)
E          545  (600)
F         1,135  (1,250)
G          499  (550)
H          145  (160)
    Gas Flow Rate*   Sources Included**   Burning Device
     m3/h  (cfm)
        42 (25)
        48 (28)
       357 (210)
     1,490 (875)
        51 (30)

       850 (500)
        17 (10)
   BD,TD
   BD,EH,TD
   CD
   CD,CS
   BD,CD,EH
   BD,CD,EH
   BD
   BD
Lime Kiln
Lime Kiln
Lime Kiln
Lime Kiln
Lime Kiln
Lime Kiln
Cat. Furn.
Cat. Furn.
 *Reported at stack conditions.
**Source Code:
    BD - Batch digester (both blow and relief).
    CD - Continuous digester (both blow and relief).
    CS - Condensate stripping.
    EH - Evaporator hotwell.
    TD - Turpentine decanter.
                                   4-16

-------
           Gases  From  Multiple
            Effect Evaporators
          Gases  From  Turpentine
                 Condenser
 Relief Gases Fi
             •rorn
  Kamyr Digester
Blow &  Relief
Gases  From 4
Batch Digesters

8  Asplund
Defibrator Off
     Gases
                                         Floating  Cover
                                           Gas Holder
   Heat
Accumulator
                    Entrainment
                     Separator
                                       FIGURE  4-8

                    NONCONDENSABLE GAS  INCINERATION SYSTEM

-------
     4.3.2   Other Incineration Systems

Both catalytic and auxiliary furnaces can be used for incineration of noncondensable gases,
but they require  a separate combustion unit and the heat cannot normally be recovered.
Coleman (5) and  DeHaas  and Hansen  (6) report on the use of an auxiliary furnace for
combustion of noncondensable gases. The piping  system is the same as for lime kilns. The
system burns the  gases at 850 to  1000° C (1560 to 1830° F), but does not recover any of
the heat produced.  The use  of  the auxiliary  furnace has been discontinued because of
control  maintenance problems  and  unstable  burner  operation  caused  by water  and
turpentine mist entrainment.

A  second installation employs an auxiliary  furnace for incineration of batch digester and
evaporator noncondensable gases  without the use of a flow equalization device. The system
employs both blow heat condensers with large heat transfer surface areas and surge tanks of
large capacity.  Large air dilution is required along with careful gas flow rate control; one
explosion with the system did occur (9).

Catalytic  furnaces are  employed for  noncondensable  gas  incineration at two mills.
Noncondensable gases are diluted with air,  as in lime kiln systems, and are incinerated at
400° C  (750° F) in the presence of porcelain rods coated with an alumina-platinum catalyst.
Shortcomings of the  systems  include incomplete oxidation  of organic sulfur compounds,
requirements for considerable maintenance of automatic controls, and frequent replacement
of the catalyst cells if these  are allowed to contact water droplets (2).

Other methods of incineration of incineration of noncondensable gases have been reported.
Ghisoni (7) reports on  the burning of  digester and  evaporator noncondensable gases in a
natural  gas-fired power boiler; Lindberg (10) describes the addition of noncondensable gases
to the  primary air inlet of  a kraft recovery furnace.  Adams (11)  describes the use of
recovery furnaces,  auxiliary furnaces, and  waste wood  burners  for  incineration of
noncondensable gases.

     4.3.3   Liquid Scrubbing Systems

The two major types of scrubber liquids employed, to date, are caustic soda and acidic
chlorination bleaching effluent. Chase (12) describes the use of alkaline scrubbing to remove
H2S from  evaporator noncondensable  gases for return to the chemical  makeup  system.
Alkaline solutions do not have any great affinity for nonpolar organic sulfur gases and do
not achieve any significant removal of them.

Morrison (13) describes a system  where batch digester blow  and  relief gases are incinerated
in a lime kiln. A backup system employs addition of the digester gases to the dropleg of the
acidic chlorination bleaching  stage when the burning system is not in use. The excess
                                        4-18

-------
chlorine converts  the  sulfur compounds to elemental sulfur, sulfonyl chlorides, sulfoxy
compounds, and  oxidized terpenes.  The system appears to oxidize the sulfur compounds
sufficiently to prevent their release to the atmosphere.

     4.3.4   Standby Systems

The maximum  degree of  control  of  non-condensable gases is achieved by  providing
alternative combustion units for the  incineration of these gases. If lime kiln incineration is
practiced, the alternative  could be routing  the gases to a  second lime kiln, a boiler or a
furnace. Such a  system offers  the  possibility for continued noncondensable gas control
during periods when the primary incineration device is out of service.

4.4  System Economics

     4.4.1   Capital Costs

The capital costs for a noncondensable gas handling system depend on these parameters:

     1.  Diameter, length, and materials of  piping;

     2.  Number, type, and size of safety appurtenances; and

     3.  The possible use of flow equalization gas holders and auxiliary fans.

The flow equalization gas holder is basically a fixed cost item of $25,000 to $50,000, based
on type and size of unit, and is independent of the amount of pulp production. Piping and
safety  devices add an additional $25,000 to  $75,000  or more to the cost of the system. The
use of auxiliary furnaces instead of lime kilns  increases  the cost of the system. A summary
of installed capital costs for noncondensable  gas handling systems is presented in Table 4-10.

     4.4.2   Operating Costs

The major  operating cost variables for noncondensable gas handling systems are those for
the additional electric power for increased  kiln air draft or auxiliary fans,  and for system
maintenance. Combustion of noncondensable gases in the lime kiln achieves additional fuel
savings  by reducing oil or gas requirements. In addition, turpentine can be burned in the
lime kiln to reduce fuel requirements. Morrison (13) reports system maintenance costs of
$2,100 per year; while fuel savings of $1,200 result from burning  noncondensable gases in
the lime kiln. Operating costs for noncondensable gas incineration are estimated to be $0.01
to $0.05 per metric ton of pulp produced, with an average of $0.03  per metric ton ($0.01 to
0.05 per short ton, avg. $0.03/short ton).
                                        4-19

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                                   TABLE 4-10
                    CAPITAL COSTS  FOR INSTALLED  NON-
                   CONDENSABLE GAS  HANDLING SYSTEMS

                 Burning Device
                   or System               Installed Capital Cost
                                       S/daily t        (I/daily ton)

               Lime Kiln                110-165         (100-150)

               Auxiliary Furnace          165-220         (150-200)

               Catalytic Furnace          192-275         (175-250)

4.5 References

 1. Personal communication with Mr. Andrew F. Reese, Fibreboard Corporation, Antioch,
    California, February 1970.

 2. Blosser, R. 0., and Cooper, H. B. H.,  Current Practices in Thermal Oxidation of
    Noncondensable Gases in the Kraft Industry. Atmospheric Pollution Technical Bulletin
    No. 34. National Council of the Paper Industry for Air and Stream Improvement, Inc.,
    New York, New York, November 1967.

 3. Perry,   J.  H.   (ed.).  Chemical Engineers  Handbook,  3rd.  edition.  New  York.
    McGraw-Hill Book Company, 1950. p. 1585-1586.

 4. Ginodman, G. M., Purification of Waste Streams from Sulfate Cellulose Manufacture.
    Bumazhnaya Promyshlennost, (Moscow) 22 (7): 16-22, November-December 1947.

 5. Coleman, A. A., The Combustion of Noncondensable Blow and  Relief Gases  in the
    Lime Kiln. Tappi, 4L166A-168A, October 1958.

 6. DeHaas, G. G., and Hansen, G.  A., The Abatement of Kraft Mill Odors by Burning.
    Tappi, 38:732-738, December 1955.

 7. Ghisoni, P., Elimination of Odors in a Sulfate Pulp Mitt. Tappi, 37:201-205, May 1955.

 8. Hansen, G. A., Odor and Fallout Control in a Kraft Pulp Mill. Journal of Air Pollution
    Control Association, 12:409-412, September 1962.
                                      4-20

-------
 9.  Personal  communication with  Mr.  Dwayne J. Clark, Simpson  Lee Paper  Company,
    Everett, Washington, 1972.

10.  Lindberg, S., How One Swedish Mill Destroys Air and Water Pollutants. Pulp and
    Paper, 41 (3):35-39, January 16, 1967.

11.  Adams, D. F., A Survey of European Kraft Mill Odor Reduction Systems. Tappi,
    48:83A-85A, May 1965.

12.  Chase, S.  J., Control of Air Pollution at the Champion Paper  and Fibre  Company.
    (Paper Read to  the  Semi-Annual Technical Meeting  of the Air  Pollution  Control
    Association. Houston, December 3, 1956.)

13.  Morrison,  J. L.,  Collection  and  Combustion   of Noncondcnsable  Digester and
    Evaporator Gases. Tappi, 52:2300-2301, December 1969.
                                     4-21

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                                   CHAPTER 5

                           CONDENSATE  TREATMENT
Besides noncondensable air polluting gases the kraft process produces condensates that are
contaminated with different compounds. These compounds can generate both air and water
pollution.

A general course of action is to decrease the amount of condensates, collect them, and reuse
them; if they are not reusable, the procedure then is to strip them of their contaminating
compounds and oxidize these compounds to less harmful forms.

5.1  Condensate Components

The  kraft process produces two main condensates,  namely  digester and  evaporator
condensates. Both contain compounds that fall into two broad classes:

     1.    BOD producing compounds that mainly generate water pollution, and

     2.    Odorous compounds that mainly generate air pollution.

Characteristically, the first class of compounds are volatile (boiling points between 56 and
150° C (133 and  302° F)), chemical oxygen  demanding, and partly toxic (turpentine).
Typical of the second class of compounds is that they are volatile (boiling points -59 to
117° C (-75 to 243° F) ), reduced sulfur containing, odorous, chemical oxygen demanding,
and toxic.

Although this  section  of  the manual  focuses  chiefly  on the second class, the odorous
compounds, it is advantageous and sometimes  necessary to briefly touch upon the first class,
the BOD  compounds. They may be treated with the same methods, keeping in mind that
they are less volatile than the odorous substances.

The  main components of typical  kraft mill  contaminated eondensates are  enumerated in
Table 5-1.

Components 1 to 3, especially CH3OH, are  mainly responsible for the  BOD load of the
condensates; components 4 to 8 are mainly responsible for the toxicity of the condensates;
and components 5  to 8 are mainly responsible for the odor of the condensates. Because of
the entrained black liquor, the condensates are, for the most part, on the alkaline side. All
these components will exist in kraft mill condensates in varying amounts depending on place
of origin  (digester or evaporator), pulp  raw  material (wood  species), operating practices
                                        5-1

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                                    TABLE  5-1
    MAIN  COMPONENTS OF TYPICAL KRAFT MILL  CONDENSATES (1, 2, 3)

                                                                            Odor
No.        Component         Boiling Point        BOD        Sulfur       Threshold
~~                            ^C     CFT        kg/kg         %            ppb

1          CH3OH             64.7  (148.5)        1.00          0          100,000
2          CH3CH2OH         78.5  (173.3)        1.23          0            10,000
3          CH2COCH3         56.5  (133.7)        0.67          0          100,000
4
CH3OH
CH3CH2OH
CH2COCH3
Turpentine
(Pinene)
H2S
CH3SH
CH3SCH3
CH3SSCH3
64.7
78.5
56.5
154

-59.6
7.6
37.5
117
(148,5)
(173.3)
(133.7)
(309)

(-75)
(45.7)
(99.5)
(243)
1.00
1.23
0.67
—

0.60
0.07
0.31
0.61
0
0
0
0

94
67
52
68
D          H25               -3V.0  (-(?>)          U.OU         V<4           0.4-5
6          CH.SH              7.6  (45.7)         0.07         67           0.4-3
                                                                            1-10
                                                                            2-20
 (such as sulfidity and cooking time), type of equipment (continuous or discontinuous), and
 condition of equipment (such as capacity and age).

 Because of the  difficulty in separating air and water pollution aspects of condensate
 treatment, both are covered together.

 5.2  Digester Condensates

 Digester condensates will vary in flow and composition. The quantities of condensates are
 especially different for batch digesters and for continuous digesters.

     5.2.1  Batch Digester Condensates

 The amount of turpentine decanter condensates from batch digesters is fairly similar in
 different mills. The condensates may vary depending on digester pressure relief method. A
 typical composition is given in Table 5-2 and flow and load range in Table 5-3.

 The amount of batch digester blow condensates, or the overflow or  bleed from the blow
 heat accumulator (flow 3  in Figure 2-1) is heavily dependent on the amount  of additional
 fresh water (flow 2 in Figure 2-1) that is injected into the condenser (blow  steam, direct
 contact)  to boost vapor condensation (section 2.1.1.1). Additional water  means  more
 condensate to  treat. The importance of a  properly dimensioned and efficiently working
 blow heat recovery system is again emphasized.
                                        5-2

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                                    TABLE 5-2
   TYPICAL KRAFT  MILL  CONDENSATE  COMPOSITIONS,  MEAN  VALUES FOR
                                   10  MILLS  (2)
  Condensate
  Compound
 H2S
 CH3SH
 CH3SCH3
 CH3SSCH3
 Total S
 CH3OH
 CH3CH2OH
 CH3COCH3
 Total BOD
Turpentine
 Decanter
   mg/1
     90
    250
    400
    130
    550
   6,500
   1,600
    160
    860
Means: Digester
     Blow
     mg/1

       60
       80
       70
       50
       180
     4,300
       500
       40
       490
Evaporator
  Effects
   mg/1
     40
     10
      5
      5
     51
 10,000
     60
      6
  1,070
Evaporator
  Hotwcll
   mg/1

    100
     40
       7
     15
    135
   1,000
     40
     10
   1,060
The  volume of blow  condensates will diminish if the blow heat recovery system works
without heat exchanging (i.e., if the direct contact condenser is fed with cool or warm water
from the outside and  the hot contaminated water in the blow  heat accumulator is used
directly for pulp washing. Such a system, however, increases the odors from the pulp in the
washing area. This pulp will require bleaching. Therefore, such a system will probably not
meet future air pollution requirements.

     5.2.2  Continuous Digester Condensates

The continuous digester relief and flash condensates are rather similar in composition to the
corresponding'condensates from batch digesters.

The continuous digester flash condensate will be found in different parts of the kraft mill,
depending on where the flash steam is used. Usually the black liquor from the continuous
digester is expanded or flashed in two stages, and the steam from the primary flash tank is
used in the presteaming vessel (Figure 2-9).  Most of the noncondensable and low boiling
compounds end up in the turpentine recovery system and are vented from that system. The
secondary flash steam  may be put into the presteaming vessel,  into  a  condenser for hot
water  generation,  or  into  an  evaporation  plant  as primary  steam  (section  3.2.1.4).
Consequently, the flash condensates will end up in those places,  perhaps mixed with other
condensates. Large variations in the amounts of flash and turpentine condensates actually
emanating from the digester area will consequently occur.
                                       5-3

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                                    TABLE  5-3
 TYPICAL KRAFT  MILL CONDENSATE CHARACTERISTICS  FOR  17 MILLS (4)
   Characteristic
 Terpentine
  Decanter
                                              Condensates
  Digester
   Blow
   Evaporator
     Effects
  Evaporator
   Hotwell
Flow, m3/t (gal/ton)
  Max.                0.3(72)         4(960)       6(1,440)
  Mean               0.15(36)        2(480)*     6(1,440)
  Min.                0.08(19)        0.9(216)     6(1,440)
                                               16 (3,840)
                                                7(1,680)*
                                                1.5 (360)
Sulfur, kg/t (Ib/ton)
  Max.
  Mean
  Min.

BOD 7, kg/t (Ib/ton)
  Max.
  Mean
  Min.
0.08 (0.16)
0.05 (0.10)
0.01 (0.20)
5(10)***
1(2)
1 (1)***
0.72(1.44)
0.36 (0.72)
0.10(0.20)
2(4)
 2.25 (4.5)**
 0.30 (0.6)
 0.10 (0.2)**
15 (30)**
 6.5(13)
 5(10)**
0.60(1.2)
3.5 (7)
  *Values exceeding 1 m3/t (240 gal/ton) pulp indicate fresh water addition to blow heat recovery system
   or condensing system.
 **Includes evaporator hotwells.
***Includes digester blow.


5.3  Evaporator Condensates

The evaporation  plant Condensates are usually divided into primary condensates emanating
from the first stage, secondary  condensates emanating from the other stages, and hotwell
condensates.  The hotwell  condensates include condensates from primary and secondary
condensers  and from the vacuum pulp (refer to Figure 3-2). If the primary steam is  pure
back pressure steam, the primary condensates are clean (with no leaks) and may be returned
to the power station feedwater system.

If the  primary  steam  is wholly or  partly flash steam  the  primary condensate will be
contaminated  and  need  treatment.  The  hotwell condensates  will  be  usually  more
contaminated than the secondary condensates. Fairly typical values  for composition and
amounts of the different condensates are given in Tables 5-2  and 5-3. The composition of
the condensates  varies  greatly  depending on wood species cooked, sulfidity, evaporation
temperatures, and condensers.
                                        5-4

-------
     5.3.1  Evaporator Condensate Quantity Reduction

The quantity of combined evaporator condensates should be about 7.5 m3 per metric ton of
pulp (1800 gal/ton) without direct contact evaporation and with a dry solids yield of 1.5 kg
per kg of pulp (1.5  tons/ton). The total evaporator condensates flow, however, is usually
larger (see Table 5-3), because fresh water is added (see Figure 5-1). Surface condensers will
reduce the amount of condensate (see Figure  3-4). To eliminate the addition completely, a
heat exchanger can be installed and the hotwell condensate circulated through it for cooling
and reuse in condenser  and vacuum pulp (Figure 5-2). By feeding white liquor into the loop
there will be a simultaneous gas scrubbing and condensate return to the white liquor system
(section 3.2.1.2).
                                   WATER JET
                                   CONDENSER
                   BAROMETIC
                   SURFACE
                   CONDENSER
              LAST
              STAGE
      VAPOR 1
      COND. 2
                     9  WARM WATER
                     8  FRESH WATER
                                                              VENT
                                                   I  WATER RING
                                                   !  VACUUM PUMP
                      SECONDARY
                      CONDENSATE
TAIL
CONDENSATE
              POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                                  FIGURE 5-1
EVAPORATION  PLANT SURFACE  CONDENSER WITH WATER  JET CONDENSER
                           AND WATER  RING PUMP
                                      5-5

-------
                                          WATER JET
                                           CONDENSER
               SURFACE
               CONDENSER
         LAST
         STAGE
 VAPOR 1
 COND. 2
                                                                 9 WARM WATER
                                                                 8 FRESH WATER
                                                                   VENT
                                                      WATER RING
                                                      VACUUM PUMP
                                          CIRCULATION
                                          PUMP
                  SECONDARY
                  CONDENSATE
TAIL
CONDENSATE
             POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY Q
                                  FIGURE 5-2
     EVAPORATION PLANT  SURFACE CONDENSER WITH  RECIRCULATED
              WATER JET  CONDENSER AND WATER  RING PUMP
    5.3.2  Evaporator Condensate Segregation

A natural and very effective way of facilitating contaminated condensate treatment is to
separate the condensates at their point of origin and group them according to their sulfur
and BOD content. This procedure applies especially well to the evaporation plant, where
most of the sulfur tends to concentrate in the tail end, and most of the BOD after the weak
liquor feed.

The use  of segregation is illustrated on an  evaporation plant with 5 stages such as that
presented in Figure 3-2. The liquor flow is 3-4-5-2-1 and stages 4 and 5, following the liquor
feed stage, are equipped with two-stage venting through liquor preheaters (section 3.2.1.1).
About 15 percent of the vapor to the  stage condenses in the preheater.  The evaporation
plant condenser is divided into two surface condenser stages. About  13 percent of the vapor
                                       5-6

-------
 is condensed in  the  second  stage and 2 percent in a Mater-ring vacuum  pump  that has
 indirect cooling.  Under these circumstances the distribution of flows, sulfur, CH3OH, and
 BOD will be approximately as presented in Table 5-4.
                                   TABLE 5-4
   CALCULATED EVAPORATOR CONDENSATE  FLOW, SULFUR,  METHANOL,
     AND BOD  DISTRIBUTION FOR LIQUOR SEQUENCE 3-4-5-2-1 (4,5,6,7)

    Point of Release
 (No. refers to flows in
       Fig. 3-2)             Temp.      Flow(a)     Sulfur(b)    CH3OH(C)   BOD(d)


2-Effect, condensate      115 (239)      21.0          1           11
2-Effect, vent

3-Effect, condensate      100 (212)      18.0          1           2         2
3-Effect, vent
4— Effect, condensate
4— Effect, preheater
4-Effect, vent
5— Effect, condensate
5— Effect, preheater
5— Effect, vent
6— Primary condenser
7— Secondary condenser
8— Vacuum pump
85 (185)
80 (176)

70 (158)
65 (149)

55 (131)
40 (104)
30 (86)
15.3
2.7

17.0
3.0

19.6
2.9
0.5
3
3

2
2

12
49
27
24
33
8(0
6
12
4(e)
3
12
7
22
32
8
6
11
4
-------
By combining various effect condensates, for example, from effects 2 to 5, with those from
the primary  condenser, 90.9 percent of the  condensates with  19 percent of the total
reduced sulfur (TRS), and  35 percent of the total BOD at a mixing temperature of 86° C
(187° F) are  obtained without  flashing. This flow may be reused within the process. The
remaining 9.1 percent of the condensates will contain 81 percent of the sulfur and 65
percent of the BOD.  This  small flow can be rather easily treated, for instance by steam
stripping. Then, by combining flows from 2,3, and 5 with those of the primary condenser,
75.6 percent of the  condensates  are obtained that  contain only  16  percent  of  the total
sulfur and 13 percent of the total BOD. Other liquor sequences, vaporization distributions,
and preheater locations will change the distribution of contaminants.

Condensate segregation  requires special  piping arrangements and usually causes a small
increase in the primary steam consumption in the evaporation plant because of incomplete
secondary condensate flashing.

Segregation  of the hotwell condensates is needed solely for  odor abatement.  Hotwell
condensates include those  from the primary condenser, the secondary condenser, and the
vacuum pump. These  condensates alone contain 88 percent of the TRS in 23 percent of the
condensate volume.

     5.3.3  Weak Black Liquor Oxidation

An effective way to eliminate odors is to oxidize the weak black liquor.

By oxidizing the weak black liquor, the sulfide is converted to thiosulfate and the CH3SH to
CH3SSCH3.  Therefore, H2S and CH3SH  are not  liberated  in  the  evaporation process.
Consequently the condensates will require little, if any, treatment for odor abatement. The
oxidation can be quite effective, and different systems using air have been developed. One
example is the British Columbia  Research  Council  (BCRC)  system that works as a weak
black liquor  and odorous gas oxidation  system. This  particular  system uses  bleach plant
effluent containing rest chlorine as one gas scrubbing and oxidation agent, and so obviously
it a bleach plant must be present.

There  are, however,  some drawbacks to weak liquor oxidation. During evaporation  and
storage the elemental sulfur generated tends to reconvert to H2S, and the CH3SSCH3 to
CH3SH. Such reconversions will nullify to some extent  the previous oxidation effort (8).

Another difficulty, especially with resinous softwoods, is the foaming of the weak liquor in
the  oxidation process.  Its foaming tendency is a function  of its concentration. At high
concentrations  the foaming decreases. Black liquors that  can be  oxidized as  thick liquors
may be impossible to  oxidize as weak liquor, even up to 23 percent dry solids (8).
                                         5-8

-------
 Furthermore, black liquor oxidation, weak or thick, will decrease the heat value of the dry
 solids by an average of 523  MJ per metric ton of pulp (0.45 X 106 BTU short ton of pulp).
 Also, the  oxidizing air strips off odorous compounds from the liquor. Current trends are
 toward oxidation of thick black liquor just before its evaporation by direct contact with the
 recovery boiler flue gases. For more information on weak black liquor oxidation, see section
 9.1.

 5.4  Condensate Chlorination

 It is possible to deodorize condensates from digesters  and evaporators by mixing elemental
 chlorine (C12)  in the condensates. Since the chlorine demand of reduced sulfur compounds
 is high (Table 5-5) this is an  expensive method unless there is an inexpensive source of
                                   TABLE  5-5
                  CHLORINE DEMAND OF REDUCED SULFUR
                COMPOUNDS FOR  OXIDATION TO SULFUR (9)
            Compound    pH
            H2S
            H2S

            CH3SH
            CH3SH

            CH3SCH3
            CH3SCH3

            CH3SSCH3
            CH3SSCH3
4
8

4
8

4
8

4
8
         C12 Demand
              Redox Potential
                                kg Cl2/kg Compound      (+) Volts
9.2
5.6

6.7
5.1

2.2
2.4

1.9
2.2
0.25
0.26

0.45
0.61

0.44
0.62

0.58
0.70
chlorine available. For instance, in some mills bleach plant effluent containing rest chlorine
is mixed with odorous condensates, and the combined effluent has neither H2 S nor chlorine
odor (10).

5.5  Condensate Stripping

Stripping the contaminated condensates has proved an efficient and economical way of
removing odor  and BOD. Condensate stripping  is becoming  the  dominant  treatment
                                       5-9

-------
method. The two principal ways are air stripping and steam stripping. To facilitate stripping,
certain preconditioning techniques should be utilized.

     5.5.1   Condensate preconditioning

After the  volume of condensates to be  stripped has been reduced to the minimum feasible
through using  reduced fresh  water  input and condensate  segregation,  there  are  other
conditions to be met.

A  higher  temperature  of the  released contaminated  condensate  will  aid  treatment by
stripping.  Air stripping will be  more effective since the  volatiles will have a higher  vapor
pressure. Similarly, steam stripping will  require less steam for heating the condensates to the
boiling point.

The  condensates should have as low a pH as possible. A low pH means a high volatility for
the ionized  sulfur compounds,  H2S and  CH3SH (Figures 2-4 and  2-5); whereas, high pH
(greater than 9) results in low volatility as well as  foaming. High pH is the result of black
liquor entrainment  and,  thus, may indicate insufficient drop and  mist separation in the
vapor flows or equipment overloading.  The points to watch for high pH values are the blow
tank and the evaporator drop separators.

     5.5.2  Air Stripping

A  very simple condensate treatment is to  strip the condensates in a multistage column (that
is, a column with  multiple trays) with a large countercurrent  flow of air or flue air (Figure
5-3) (11).

Use  of an atmospheric vent (see flow  5  in Figure 5-3) for air or flue gas stripping simply
translates a water pollution  problem into an  air pollution problem. Although part of the
TRS will  be oxidized while feed and air or flue gas mix in the column, most of the TRS will
pass out  through  the vent. With small feeds and air flows the vent can  be connected to a
boiler furnace or to an incinerator, but for air flows equal to thousands of cubic meters per
hour this  is  hardly feasible.

Experience  with working units (7)  indicates that column features and  performance are
approximately as follows:

      1.  Tray type, bubble cap;

      2.  Liquid feed rate, Ffee(j m3/h (Ffee(j  gpm)

      3.  Tray number, 10-20;
                                         5-10

-------
        STRIPPING
        COLUMN
                                                 ©VENT

                                                 1) FOUL CONDENSATES
                                                  2 REST ACID
                             BLOWER
                                                  4  AIR
                                            •>    1  CLEAN CONDENSATES
               POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY Q

                                 FIGURE 5-3

               CONTAMINATED CONDENSATES AIR  STRIPPING
                            PLANT FLOW SHEET
      4.  Tray distance, 500 mm (20 in)


      5.  Column diameter, 130\/Ffeec( mm (2.44\/Ffee(j in)

      6.  Air flow, 18 X Ffeed;


      7.  Feed temperature, 65-70° C (149-158° F);


      8.  Feed pH, below 9;

      9.  TRS of feed, 130-320 g/m3


    10.  BOD (CH3OH) of feed, 390-1030 g/m3


    11.  TRS removal, 80 percent; and


    12.  BOD removal, 0-10 percent.


Tripling the air flow will increase TRS removal to 84-94 percent and BOD removal to 10-15
percent.
                                     5-11

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The eondensate feed must be as hot as possible and the pH must stay below 9. If pH rises
above  9,  reduced stripping  efficiency  and  serious foaming problems  will follow. The
temperature of the ambient air will affect the feed temperature, but  this effect is small,
about 6° C(ll° F).

If the  pH of the condensates cannot be brought below 9, rest acid must be added to the
eondensate feed (see flow  2  in Figure 5-3).  Rest acid is obtained, for example, from the
manufacture of bleach plant  C1O2, and  added at the feed pump suction. A control loop
from after the pump can be set to keep the feed  pH at a suitable level of about 7-8. Air or
flue gas stripping will not remove more than 95 percent of TRS nor more than 15 percent of
BOD at reasonable gas flow rates.

     5.5.3  Steam Stripping

Steam  stripping was first used about 25 years ago at Skoghall, Sweden (9). This steam
stripping  plant is  still working and has been modernized. Steam stripping  research was
carried out in  pilot plants (7, 12) and, subsequently, put into full-scale operation in a num-
ber of kraft mills.

          5.5.3.1   Separate Steam Stripping Plants

Steam  stripping seems to be the most feasible method of purifying contaminated kraft mill
condensates. Steam stripping can be divided into two categories: TRS removal and BOD
removal. Characteristic features of stripping TRS  are strong pH dependency and low steam
consumption,  which is about  2 percent of the  eondensate feed for a 90 percent TRS
removal. Characteristic features of stripping BOD  are much lower pH dependency and high
steam consumption, which is about 20 percent of the feed for a 90 percent BOD removal.

A typical steam stripping  plant is  shown in Figure 5-4. Condensates (flow 3)  will enter
a storage tank with level alarms. Rest acid addition (flow 2) is controlled to keep pH at 7-8
after leaving the tank to enter the  preheater/primary condenser. Condensates then receive
heat from outgoing stripped  condensates  in a heat exchanger.  Then they pass a direct
contact steam  injector  with temperature  control to insure  that column feed temperature is
sufficient and  constant. Condensates next pass down the column  countercurrently to fresh
steam  injected at  the bottom  of the column. Condensate feed is kept constant by flow
control subject to alarm from  the storage tank-level monitor. Stripping steam flow also is
kept constant, but the flow control set point will follow feed flow in a fixed ratio that can
be adjusted. This ratio will essentially be equal to the steam/condensate ratio, taking into
account steam consumed for heating the condensates to the boiling point.

The  vapors  rise through a  fortifier section of the column countercurrently to eondensate
feedback  and  enter the primary condenser/preheater, where most of the vapor condenses.
                                        5-12

-------
 SECONDARY
 CONDENSER '
PRIMARY
CONDENSER
   HEAT
   EXCHANGER
                       FIGURE  5-4
8 VENT


7 FRESH WATER

6 WARM WATER
                                         5 TURPENTINE PHASE
                                        4  STEAM
                                        2 REST ACID
                                        1  CLEAN CONDENSATES
    CONTAMINATED  CONDENSATES  STEAM  STRIPPING
                  PLANT  FLOW SHEET
                          5-13

-------
The  noncondensables and the rest  of the water vapor pass to the seeondary condenser,
which serves as a final condenser and gas cooler. The cooling is controlled by the outgoing
gas temperature, which will stay constant. Adjusting the set point of the controller means
adjusting the temperature and the dew  point of the noncondensable gases from the vent
(flow 8 in Figure 5-4). By adjusting the dew point, these gases can be kept at a humidity
level (40%) that will greatly diminish the explosion risk.

The column condensates flow to a separator where eventually turpentine and other organic
compounds may form an oily layer to  be  drawn off with a  level control and  piped  to
incineration (flow 5 in Figure 5-4). The underflow  is pumped back to the column fortifier
section top under level control. To give a very general idea of how to dimension and what to
expect from a steam stripping column, the following data are given:

     1.    Tray type, bubble cap;

     2.    Steam flow rate, Fsteam in metric tons per hour (Fsteam short ton/hour)

     3.    Tray number, 10-20;

     4.    Tray distance, 500 mm (20 in);

     5.    Column diameter, 780\/Fsteam mm (29.2\/Fsteam in)

     6.    Feed pH, below 9;

     7.    TRSoffeed,210g/m3

     8.    BOD of feed, 300 g/m3

The  removal percentage as a function  of the steam/condensate ratio  is approximately as
shown in Figure 5-5. A lower pH of the feed would improve these removal rates for H2S,
especially at low steam/condensate ratios. After the contaminants have been stripped out of
the condensate, and the main part of the water vapor condensed, the remaining gases are
usually incinerated.  Such gases may be  incinerated  in either the lime kiln, a separate
incinerator, or possibly the recovery boiler.

          5.5.3.2  Evaporation Steam Stripping Plants

Stripping of the condensates is usually  performed in a separate stripping column with fresh
steam. One way to reduce the high treatment costs involved  is to combine this stripping
with the evaporation plant, using "secondary steam" from the black liquor evaporation step.
Such an arrangement is shown in Figure 5-6.
                                         5-14

-------
       >-
       o
o
LU-
LL
LU
       LU
       DC
            100
            80  -
            60
            40  -
            20  -
              0
                 0246          8         10

                              STEAM/CONDENSATE,  AS  %

                                  FIGURE 5-5
        STRIPPING  EFFICIENCY FOR  DIFFERENT STEAM-CONDENSATE
                    RATIOS WITH  10 THEORETICAL  PLATES

The stripping column is placed on top of the second effect, similarly to the alcohol stripping
in a sulfite spent liquor evaporation plant.  All dirty condensates are piped to the stripping
column, where the stripping efficiency is 95 percent. The CH3 OH stripped off is withdrawn
through the liquor preheater of the third effect together with a small amount of steam and
is condensed to form an 8-10 percent CH3OH water solution. The CH3OH is recovered in a
small stripping column and  destroyed by burning. Alternatively, it can be recovered and
sold. The total BOD-removal of the condensates by this stripping arrangement is estimated
at 90 percent. If the stripping column is placed on top of the second evaporator effect, the
heating  surface of the evaporation plant  has to be  increased  by approximately 7 to  8
percent. This increase is necessary to compensate for the pressure loss in the column and the
lower  condensation  temperature in the  third  effect. The steam  consumption  of the
evaporation plant increases by 5 to 7 percent.
                                      5-15

-------
en

t—>
ON
                  STRIPPING
                   COLUMN
            STEAM
                                                                                          HOT WATER,
	 	 	 "W
(N^l ^

A

1



•Mm
A

1








^*^^J 1 ^
COLD WATER
JL
t
                                                                                      GASES TO DESTR.
              PRIM.
              COND.
               FOUL CONDENSATE
                FROM DIGESTER
                                                                               CONDENSATE TO REUSE
                                                  FIGURE 5-6

                             CONDENSATE STRIPPING  IN  AN EVAPORATION PLANT (13)

-------
By using the stripping arrangement shown in Figure 5-6, the evaporation plant will produce
thick liquor and  clean, practically distilled water  for  reuse, plus concentrated volatile
organics as a byproduct.

     5.5.4   Condensate Finishing

Condensate  finishing may be feasible when very high removal efficiencies of greater than 99
percent are  required. Instead of stripping with  an excessively large steam/condensate ratio,
enough stripping steam is first used to decrease TRS by 90 or  95 percent, and  then the
stripped  condensate is treated  with ozone, chlorine, or activated carbon (14). Condensate
stripped  of TRS  may  be  finished in a biological treatment plant where most of the
remaining traces of TRS and  BOD are removed. Biological condensate  finishing can be
accomplished if the mill already has a biological treatment plant  for its other wastewaters.

     5.5.5   Condensate Reuse

Condensate  reuse  includes reuse of both treated and untreated  condensates. When using
treated condensates that  have been stripped of their TRS and  BOD, there are no reuse
problems since the treated condensate is clean, hot, distilled water.

When reusing condensates that  have been stripped of their TRS components only, the BOD
components will come out  from the process at  some other point. If the components are
recirculated, they  will build up within the process unit until such  a level is reached that
discharge takes place. The best way to treat such condensates is to direct the discharge of
BOD components into the recovery boiler through the black liquor.

When using untreated contaminated condensates, the TRS must be removed since it will
either go down the drain to the receiving waters or it will start circulating and building up
until it is discharged to the atmosphere and/or to the water. A possible solution is to put the
TRS back into the white liquor, where it will provide part of the sulfide and decrease the
sulfur make-up  demand.

         5.5.5.1  Pulp Washing

Untreated contaminated condensates have  been used for pulp washing, especially in mills
with direct  blow  heat  recovery (see section 5.2.1). But such an arrangement causes air
pollution from washer  hoods,  and water pollution from the  pulp screening, if  an
open-screening  system  is  used. A closed-screening system  will reintroduce  the  water
pollutants, raise their level, and  increase air pollution.

Before contaminated condensates  are  used for pulp washing,  they  should be  stripped
completely  for open-screening systems, and at least  of  their  TRS components for
                                         5-17

-------
closed-screening systems. A stripped condensate is a very good washing liquid, namely hot
distilled water. The total condensate amount, without adding fresh water, will  be about
7.5-8.5  m3/per metric  ton  of pulp (1800-2000 gal/ton). The amount will depend upon
whether there is  continuous or batch digestion. Of this amount, about 8 m3 per metric ton
of pulp (1900 gal/ton)  can  be used for washing pulp, thus meeting the whole wash water
demand.

         5.5.5.2   Lime Kiln Flue Gas Scrubbing

Using contaminated condensates as scrubbing liquids in lime kiln flue gas scrubbers amounts
to stripping them with  hot  gases while they reciprocally wash particulates from  the gases.
The result is increased air pollution caused by TRS. To use stripped condensates in lime kiln
or flue gas scrubbers offers no advantage, since ordinary fresh water will do the same job
while additionally picking up heat from the hot gases. Stripped condensates are hot distilled
water and  should  be used  in proper  applications, such as washing pulp in the washing
department or in the bleach plant. Fresh water demand for  a lime kiln scrubber is about
4 m3 per metric ton of pulp (960 gal/ton).

         5.5.5.3   White Liquor Liquid Makeup

In the digester, the liquid/dry wood ratio can be trimmed by black liquor recirculation. The
necessary liquid makeup to the white liquor can be supplied at three points, namely in the
smelt dissolving tank, in the mud washers, and in the white liquor itself.

The  smelt dissolving tank requires liquid at about a rate of 1.5 m3 per metric ton  of pulp
(360 gal/ton). The liquid should preferably be cool water, since the heat input to the smelt
dissolver is already  so  great that addition of hot water would cause vaporization  of both
water and volatile and odorous substances. Condensates, therefore, are not suitable for smelt
dissolver liquor makeup  unless they are first stripped of TRS and cooled.

The  mud washers need hot water and usually 50 percent of the white liquor liquid demand,
approximately 1.5 m3  per metric ton  of pulp (360 gal/ton), enters the process this way.
Contaminated condensates will cause odor problems at the mud  washer, but TRS stripped
condensates may be used although BOD components then will be fed back into the process.

If the mud washer system is highly effective, that is, if it will work with less wash water than
the  1.5 m3  per  metric ton of pulp (360 gal/ton), the remaining makeup liquid can be
supplied directly to the white liquor. Treated or untreated condensates can be used since
odorous compounds will, for the most part, be absorbed by the alkali of the white liquor.
The BOD level will rise,  of course.
                                        5-18

-------
 A plan for condensate segregation, stripping, and reuse is suggested in Figure 5-7. This plan
 will require about 1.5 m3 per metric ton of pulp (360 gal/ton) of fresh water makeup and
 about 1 m3- per metric ton of pulp (240 gal/ton) of hot water makeup for the pulping cycle.
 Seven percent of the combined digester and evaporator condensates, containing 91 percent
 of the combined TRS and 73 percent of the combined BOD, are  stripped and used for
 washing pulp and lime mud. Seventeen percent of the BOD is returned  to the white liquor.
 Ten  percent of the BOD  and 9  percent of  the  TRS  are  returned to pulp washing.
 Approximately 50 percent of these  contaminants will remain with the pulp and get washed
 out during screening; the rest will return to the black liquor.
                (2 t D S )    WOOD
                                          PULP (09 t D S )

                                             55m'
             1 m1
        3 kg BOD
        OSkgTRS      3m'

                1 5m'
               1 5mJ
          Fresh Water
                       1    I    I   I   I   I    I   I   T
                       15   15  13  02 13 02  13  02   - m'
                       115   100  85  80 70 65  55  40   30 °C
                       01   02  22  32 06 11  04  14  0 8 kg BOD
                                        -010503 kg TRS
                1 3mJ
                1 Om'
                                        06m'
                                               1 6r
                 02m'
                                               02m'
1 Om'
                                  FIGURE 5-7
SIMPLIFIED  LIQUID FLOW SHEET FOR  KRAFT  PROCESS WITH CONDENSATE
                     SEGREGATION, STRIPPING,  &  REUSE
                                       5-19

-------
5.6 References

 1. Nylander,  G.,  Report  on  Forest  Industry  Waste  Waters.  Svensk Pap_perstidning
    67(15):565-572, August 1964 (Stockholm).

 2. Leornados, G., Kendall,  D., and Barnard  N., Odor Threshold Determinations of 53
    Odorant Chemicals. Journal of Air Pollution Control Association, 19:91-95, February
    1969.

 3. Wilby,  F. V.,  Variation  in Recognition Odor Threshold of a Panel. Journal of Air
    Pollution Control Association, 19:96-100, February 1969.

 4. EKONO Oy, Helsinki, Finland, files.

 5. Arne, H. G., and Bergkvist, S., Methanol Distribution in an Evaporation Plant. Svensk
    Papperstidning, 77(10):380-382,1973 (Stockholm).

 6. Jonsson, S. E., Black Liquor Evaporation. Svensk Papperstidning, 74(7): 191-196, April
    15,1971 (Stockholm).

 7. Backstrom, B., Hellstrom,  H., and Kommonen, F., Purification of Malodorous Sulfur
    Containing Condensates from Turpentine Separation, Digester Blow and Spent Liquor
    Evaporation at  the Oy Kaukas Ab,  Kraft Mill. Paperi ja Puu, 52(3):113-120, 1970
    (Helsinki).

 8. Murray, F. E., The Oxidation of Kraft Black Liquor. Pulp & Paper Magazine of Canada,
    64:82-86, January 5, 1965.

 9. Ruus, L., Report on Forest Industry Waste  Waters. Svensk Papperstidning, 67(19):751-
    755, October 15, 1964 (Stockholm).

 10. Lindberg, S.,How Uddelholm Destroys Air and Water Pollutants at the Skoghall Works.
    Pulp and Paper  Magazine of Canada, 69(7):T178-T183,  April 5, 1968.

 11. Morgan, I. P., and Murray, F. E.,/4 Comparison of Air and Steam Stripping as Methods
    to Reduce Kraft Pulp Mill  Odor and Toxicity from Contaminated Condensate. Pulp &
    Paper Magazine of Canada, 73(5):62-66, May 1972.

 12. Matteson, M. I., Johanson, L. N., and McCarthy, J. L., SEKOR II; Steam Stripping of
     Volatile Organic Substances from Kraft Pulp Mill Effluent Streams. Tappi, 50:86-91,
    February 1967.
                                        5-20

-------
13.  Study  of Pulp and Paper Industry's Effluent Treatment.  EKONO Oy. Prepared  for
    FAO Advisory Committee on Pulp and Paper. Session 13, Rome, May 15-16, 1972.

14.  Hansen, S. P., and Burgess, F. I., Carbon Treatment of Kraft Condensate Wastes. Tappi,
    51:241-245, June 1968.
                                     5-21

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                                    CHAPTER 6

                         BROWN STOCK WASHER GASES

After mixing with the original liquor and added black liquor, the pulp passes from the blow
tank to the washing plant. The washing process is a minor source of air pollution compared
to combustion, evaporation, and digestion. Generally, the washing produces large quantities
of ventilation air slightly contaminated with organic sulfur compounds through contact with
the black liquor. The amount  of air and of sulfur compounds will mainly depend on the
type of washing process and equipment and to a minor degree on wood, sulfidity, pH, tem-
perature,  and  other factors  that are generally determined by the production process. The
two main washing processes are displacement washing and diffusion washing.

6.1  Displacement Washing

Displacement washing usually takes place on  rotary drum filters using air for maintaining
the pressure difference over  the washed pulp sheet. Thus, the hot black liquor is exposed to
large quantities of air. This exposure will have two effects. One effect, for example, is that a
portion of the reduced  sulfur in the black liquor will be oxidized and will decrease the sub-
sequent odor generating capacity of the black liquor during evaporation. The other effect is
that part of the reduced  sulfur will become volatile and will contaminate the air.

     6.1.1  Vacuum Washers

The most common type of kraft pulp washer  is the vacuum washer (Figure 6-1). The two
main points of odor release to the atmosphere are the hood vent (flow 5 in Figure 6-1), with
large flows of slightly contaminated air  and the foam tank vent (flow 6 in Figure 6-1), with
smaller flows of more polluted air. Tables 1-2,  1-3, and 1-4 summarize typical emissions
from these sources.

Because of their large volume and low concentration of odorous components, the only prac-
tical way  of treating washer gases, especially those from the hood vent, is to incinerate them
in an existing boiler. For example, the gases from the washer vents are used as part  of the
combustion air in an auxiliary furnace or a black liquor recovery boiler  at several U.S. and
Swedish mills. Proper safety precautions must be designed into the system. These precautions
include condensate traps,  rupture disks, flame arresters and flame control, and emergency
vents. The smaller flow  rate of the foam tank vent gases allows incineration in the lime kiln.
The sulfur content of foam tank gases varies with their source, that is, whether softwood or
hardwood, see Table 6-1.

Some mills use  contaminated condensates from blow heat recovery accumulators or evapo-
rators for washing. This practice will  increase the odor release from the washers significantly.
The hood vent TRS emission may increase 5 to 15 times, and the foam tank vent TRS emis-
sion may increase from 20 percent to 4.5 times  when changing from fresh hot water wash to
condensate wash (1). The abatement method is to strip the condensates before use.
                                        6-1

-------
                                                                  5) VENT
                        HOOD
                                                                  4  WASH
                                                                  3  WASH
                                                                  2  WASHED PULP
                                                                     PULP
                                                                 (e) VENT
                                                                 (lo) WEAK
                                                                     LIQUOR
         POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY (J)

                                   FIGURE 6-1
                        VACUUM WASHERS FLOW SHEET
     6.1.2   Pressure Washers

Pressure washers have  closed hoods with blowers circulating air for maintaining the pressure
difference across the pulp sheet (Figure 6-2). The hood vent (flow 5 in Figure 6-2) and the
foam tank vent (flow 6 in Figure 6-2) will both have a small flow rate and a high sulfur
concentration  compared  to  the vacuum  washer hood vent.  Black liquor oxidation and
release of reduced sulfur to the atmosphere will be less, too.

Odor abatement  is much easier  than  for  a vacuum washing plant and  can best  be
accomplished by incineration of the vent gases in the lime kiln.

6.2  Diffusion Washers

Diffusion  washing usually takes place in a closed reactor, and ideally there is no air involved.
Therefore, black  liquor oxidation and odor release are very  small, when compared with
displacement washers..
                                        6-2

-------
                      TABLE 6-1
    FLOWS, COMPOSITION AND SULFUR RELEASE OF
     VACUUM  WASHER FOAM TANK VENT GASES (2)
   Wood Species     Flow
      Pine
      Birch
(ft3/ton)

   64
 (1930)
  65
 (2080)
           Component
  Sulfur
kg/t  (Ib/ton)
H2S
CH3SH
CH3SCH3
CH3SSCH3
Total
0
0.03
0.06
0.04
0.13
0
0.06
0.12
0.08
0.26
H2S
CH3SH
CH3 SCri3
CH3SSCH3
Total
0
0.02
0.13
0.13
0.28
0
0.04
0.26
0.26
0.56
POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                      FIGURE 6-2
           PRESSURES WASHERS  FLOW SHEET
                                                5) VENT


                                                4  WASH
                                                3  WASH


                                                2  WASHED PULP


                                                ?) VENT
                                           *->•  (lO)  WEAK LIQUOR
                         6-3

-------
     6.2.1   Batch Diffusers

Batch diffusers can still be found in old mills, but in many cases have been replaced by
vacuum washers (Figure 6-3). Because of the batch operation there is liquor-air contact
during blowing, washing, and emptying of the diffuser. Consequently, there is some black
liquor oxidation and odor release, symbolized by flow 5 in Figure 6-3. The sulfur release
will be difficult to determine and abate. Continuous venting to the lime kiln is probably tHe
best method of gas treatment.
1
! +
DIFF.
WASH
1


i
i
i ._ ._

DIFF.
WASH
2
1

• 	 1 	 -^ \^/ •"-'•"
1 "3 WASH
DIFF.
WASH
3
	 ^ 2 WASHED PULP

,.-,..,,,.,, k (10) WP^^ 1 lni ir>R
                                    FIGURE  6-3
                   BATCH DIFFUSION  WASHERS FLOW SHEET

     6.2.2   Continuous Diffusers

Continuous  diffusers have been  integrated with continuous digesters for the past decade.
They are also now made as separate washers (Figure 6-4). The washing process is closed-off
to minimize air  infiltration. Thus, the oxidation of  the black liquor and the subsequent
release of reduced sulfur, symbolized by flow 5 in Figure 6-4, are kept at a minimum. The
reduced sulfur is rather easily contained and is incinerated in the lime kiln.

Scrubbing of washer gases  with  an alkaline solution, such as white liquor, is not generally
practiced since the washer gas TRS is predominantly nonionizable sulfur compounds.
                                        6-4

-------
                      CONTINUOUS DIFFUSION WASHERS FLOW SHEET

                     	,	>.      ®VENT
                                   I
                                   I
                                           ->     @ WEAK LIQUOR

                                                    3 WASH
                                                     2  WASHED PULP
                                                     1  PULP 4- LIQUOR
            POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                                 FIGURE  6-4
              CONTINUOUS DIFFUSION WASHERS  FLOW SHEET
6.3  References

1.   Atmospheric  Emissions  from  the  Pulp  and  Paper  Manufacturing  Industry.
    EPA-450/ 1-73-002. September 1973. (Also published  as NCASI Technical Bulletin
    No. 69, February 1974.)

2.   Kekki, R.,  Kraft  Mill Odor  Abatement by  Condensate  Stripping and Waste Gas
    Incineration. M.S. Thesis, Wood Industry Department, Helsinki Technical University,
    Finland. September 18, 1969 (Finnish).
                                     6-5

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                                    CHAPTER 7

                          STORAGE TANK  VENT  GASES
All storage tanks, especially black liquor tanks that hold sulfide-containing liquor,  are
potential  air polluters when vented to the atmosphere. All of the storage tanks of a mill
together,  however, are a minor source of air pollution, even smaller than the vacuum washer
hoods.

7.1  Storage Tank Vent Gas Composition

The liquor in the tank is usually hot,  and the tank vent will emit water vapor contaminated
mainly with organic reduced sulfur compounds.  The composition of the noncondensable
gases  is rather similar  to that of the washer gases. The main factors influencing release of
odorous gases are liquor sulfide concentration, pH, and temperature. These factors are fixed
by  external circumstances (i.e., the general production process), and cannot be changed
except within a very narrow range.

7.2  Storage Tank Vent Gas Treatment

The fact that storage tanks are dispersed over the mill, thus constituting several air pollution
sources, makes their treatment quite difficult. An effective program is to connect all storage
tank vents to a central duct leading to an  incinerator or an oxidation tower. This system is
applied in one Scandinavian mill (1) and at  least one U.S. mill (2).

One way of reducing the release of odor from black liquor storage tank vents is to use weak
black liquor oxidation. (See Chapter 9.)

7.3 References

1.  Air Pollution Abatement Problems  of the Forest Industry. Statens Naturvardsverk
    (Sweden). Publication  1969: 3, July 1969.

2.  Michigan Department  of Natural Resources, Division  of Air Pollution Control.  Staff
    Activity Report dated  April 24, 1973.
                                         7-1

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                                    CHAPTER 8
                          TALL  OIL  RECOVERY GASES
When pulping softwood with an alkaline process, recovery of tall oil can be quite profitable.
Soap  is  skimmed from weak,  intermediate,  and strong  black  liquor storage tanks  and
evaporators, and from black liquor oxidation plants. To liberate  the fatty acids from their
sodium salts, the  soap is acidified with sulfuric acid (H2S04). The H2S04 will also displace
other  weak acids present, such as H2S  and CH3SH,  creating  a potential air  pollution
problem. The two  factors with the greatest influence on the odor release are the soap
washing efficiency and the recovery mode (batch or continuous). Tall oil recovery gases are
minor odor sources in the kraft mill.

8.1  Batch Tall Oil Recovery

A typical batch tall oil plant flow scheme is depicted in Figure 8-1. The major odor emission
point is the boiler vent (flow 5 in Figure 8-1). When H2S04 is mixed with soap, the sulfide
of the  residual black liquor, which has not been washed away, reacts to form H2 S. Thus,
there is first  a sudden surge followed by  a gradual decrease in evolution of H2S. Another
surge of H2S may follow when the brine is neutralized with white liquor. The major part of
the noncondensable flow  is air. The H2S  concentration may vary between zero and 2
percent. Other TRS components and some typical concentrations are  given in Table 8-1.
Flow rates and sulfur emission rates of TRS are given in Table 8-2.
                                                                 7  ACID
SOAP FROM
PULPWA
BLACK
LIQUOR
BLACK
LIQUO
SH
i
r
SOAP
TANK
R

r


i
MIX
TANK

SOAP
WASH
TANK
t

TO EVAP.

k





fA
BOILER
1



/
s








1
TA
Ol
WA



LL '
SH _



6 WATER
->• © VENT
4 WHITE LIQUOR
- — 3 6 1 bAM
	 ^ (2) TALL OIL
— > (l) SALT SOLUTION
TO EVAPORATOR
                                   FIGURE  8-1
                    BATCH TALL OIL  PLANT  FLOW SHEET
                                        8-1

-------
                                  TABLE 8-1

                     BATCH TALL OIL  RECOVERY PLANT
                       TRS COMPONENTS AND TYPICAL
                           CONCENTRATIONS  (1,2)

               TRS Component                Typical Concentration
                 H2S
                 CH3SH
                 CH3SCH3
                 CH3SSCH3
g/m3
8.7
0.3
0.2
0.03
(gr/cu ft)
(3.8)
(0.13)
(0.09)
(0.01)
                                 TABLE 8-2
               TALL OIL  RECOVERY NONCONDENSABLE GAS
                       FLOWS AND REDUCED SULFUR
                              EMISSIONS  (1)(2)

              Condition        Gas Flow           TRS Emission
                           m3/t (cu ft/ton)     kg of S/t (Ibs/ton)

              Maximum      20    (641)          0.66    (1.32)

              Mean          11    (352)          0.15    (0.30)

              Minimum        1    (32)           0.01    (0.02)

Since most of the TRS is H2S, one treatment method often used is to duct the boiler vent
(flow 5 in Figure 8-1) directly to a white liquor scrubber.

8.2  Continuous Tall Oil Recovery

A typical continuous  tall oil plant flow scheme  is depicted in Figure 8-2. The major odor
emission point is the reactor vent (flow 5 in Figure 8-2). The flow is much smaller but more
concentrated in TRS than for  the batch process. There are no surges of H2S, hut a continu-
ous flow. The best treatment method is to apply a white liquor scrubber to the vent.
                                    8-2

-------
ACID
WHITE LIQUOR
SO
BLK
LIQ.
B
L
^P
"^
SOAP
TANK
LK.j
• • »J
IQ.
fe.
F

SOAP
WASHER
i
i A

w
k.
w

ACID
MIXER

^1


.r" 4
REAC-
TOR
CENTRI- TALL
F FUGAL F1 OIL
SEP WASH
t
V
6 WATER

(jT)VENT
	 ^ (?)TALLOIL
^ rT^CJAIT
                POINTS OF POSSIBLE ODOR RELEASE ARE ENCIRCLED BY
                                  FIGURE  8-2
                 CONTINUOUS TALL OIL PLANT FLOW SHEET
                                                                       SOLUTION
8.3 References

1.   Air Pollution Abatement Problems of the Forest Industry.  Statens Naturvardsverk
    (Sweden). Publication 1969:3, July 1969.

2.   EKONO Oy, Helsinki, Finland, files.
                                      8-3

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                                    CHAPTER  9

                           BLACK  LIQUOR OXIDATION
Black  liquor  oxidation  is  extensively applied to facilitate  odor  control and chemical
recovery in kraft pulp mills. Its immediate purpose is oxidation of Na2 S to innocuous salts
to prevent the release of H2S. Black liquor oxidation can be performed on either weak or
strong black  liquor by air or molecular oxygen. Overall reviews of black liquor oxidation
practices have been prepared by Collins (1), Landry (2), Hendrickson (3), and Blosser and
Cooper (4).

9.1  Weak Black Liquor Oxidation — Air

Weak black liquor can be oxidized with air to decrease reduced sulfur emissions from both
multiple-effect and direct-contact evaporation  systems. Systems in extensive use for weak
black liquor  oxidation include sieve tray towers (5),  porous carbon  black  diffusers (6),
packed absorption towers (7), vertical-slat falling-film packed towers (8), and agitated air
spargers (9). Rotating fluid  contactors (10) and pressurized vessels (11) are used to a lesser
extent for weak liquor oxidation.

A previous  survey by Blosser and  Cooper (4)  indicates  that it is  possible  to obtain
consistently high efficiency of Na2 S oxidation with the porous diffuser (Collins) and the
sieve tray tower (Trobeck),  provided that sufficient gas-liquid interfacial contact areas and
air flow rates are used and that  liquor and Na2S loadings are kept sufficiently low. Less
effective performance is observed for packed towers because of inadequate liquor retention
times and for agitated air spargers because of frequent mechanical breakdowns.

Most weak black liquor oxidation systems, employing air in the United States, are located in
the Pacific Northwest, upper Midwest, and Northeast, -where highly resinous  pine wood
species  are not pulped. Blosser and  Cooper (4) report the following problems with weak
black liquor oxidation systems using air:

     1.   Excessive foaming when pulping highly resinous pine wood species,

     2.   Incomplete  Na2S  oxidation efficiency caused by  improper  or under-design of
         systems,

     3.   System overload  caused  by increased pulping capacity without expansion of
         existing facilities, resulting in inadequate liquor retention time, and

    4.   Inability to  achieve effective oxygen mass transfer from air into the black liquor.
                                         9-1

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     9.1.1  Porous Plate D iff users

Collins (12) reports  on the development of  a  two-stage system for weak  black  liquor
oxidation  with air  in  consecutive aeration and deaeration steps (see Figure 9-1).  The
aeration stage  employs passage of black liquor across a series of horizontal porous plates
arranged  vertically in a parallel  flow  arrangement. Air is blown consecutively  in series
through the porous plates in a crossflow configuration to create gas-liquid interfacial contact
by the generation of foam. The liquor depth on  the plates is normally 10 to 20 cm (4 to 8
in).

The  air-foam-liquor mixture flows off  the  plates for  deaeration in a retention tank to
provide for gas-liquid phase separation. Foam is dissipated by mechanical foam  breakers
atop the deaeration tank, and liquid droplets entrained in the exit gas stream are returned to
the deaeration tank via a cyclone separator, as shown in Figure 9-1. Recent studies by Van
                                       TO
                                   ATMOSPHERE
                  FOUR STAGE
                   OXIDATION
                     TOWER
      BLOWER
                                                             MOTOR
                                                               FOAM BRAKER
                                                  FOAM
                                                  TANK

                             BLACK
                            LIQUOR
                             PUMP
                                 FIGURE  9-1
          COLLINS  POROUS PLATE DIFFUSER WEAK  BLACK LIQUOR
                       OXIDATION SYSTEM  (14)
                                        9-2

-------
Donkelaar (13) and Shah and Stephenson (14) indicate that the Na2S concentration in weak
black liquor can be reduced to below 100 mg/1 by a porous plate diffuser oxidation system.
Design and  operating parameters calculated from  published  data for the two mills are
presented in Table 9-1. The systems are not normally suitable for highly resinous pine black
liquors because of excessive foaming.
                                    TABLE 9-1
   DESIGN AND OPERATING PARAMETERS  FOR POROUS  PLATE  DIFFUSER
                 BLACK  LIQUOR OXIDATION SYSTEMS  (13, 14)
             Parameter

 Production, t (ton)/day
 Liquor flow, m3/h (gpm)
 Air flow, m3 /h (cfm)
 Power, kW (hp)
 Plate area, m2  (ft2)
 Na2S loading, kg/m2/h (Ib/ft2/hr)
 Liquor loading, m3 /m2 /h (gal/ft2 /hr)
 Air loading, m3/m2/h (cu ft/ft2/hr)
 Power loading,  kW/t/day (hp/ton/day)
 Na2S inlet, g/1
 Na2S outlet, g/1
 Oxid. effic., %Na2S
 Oxygen ratio, act./theor.
Mill A (13)
Mill B( 14)
450 (500)
205-227 (900-975)
25,600 (16,500)
186 (250)
650 (7,000)
2.4-2.9 (0.5-0.6)
0.30-0.35 (7.4-8.6)
39 (128)
0.34 (0.41)
8.0-10.0
0.01-0.05
99+
7.7
180 (200)
68-108 (300-475)
13,700 (8,850)
75 (100)
312 (3,350)
1.2-2.4 (0.2-0.5)
0.22-0.35 (5.3-8.5)
44 (144)
0.34 (0.41)
2.3-6.6
0.01-0.10
98-99+
14.2
     9.1.2   Sieve Tray Towers

The Trobeck-Ahlen, Bergstrom-Trobeck, or Lundberg weak black liquor oxidation systems
consist of countercurrent flows of air and liquor through a series of perforated sieve trays
(15). The system is arranged in consecutive aeration and deaeration stages for consecutive
air-liquor  contact  and air-liquor  separation in series.  The  aeration stage consists of  a
seven-stage vertical sieve tray absorption tower where black liquor can be added at the third,
fifth or seventh tray (numbered upward). The liquid is allowed to cascade downward from
one tray to the next through a series of overflows and downcomers.

Air is introduced upward through the bottom of the tower and passes upward through the
sieves, generating small gas bubbles and  foam to facilitate interfacial gas-liquid contact. The
liquor is drained from the oxidation tower into a deaeration tank. Sufficient retention time
allows entrained gas bubbles to separate from the liquid. Mechanical foam breakers are used
                                       9-3

-------
for foam control. The exhaust air stream from the oxidation tower passes through a cyclone
separator for removal of entrained foam  and liquor droplets. The system is illustrated in
Figure 9-2.
  b
Extensive experience has been  obtained with sieve  tray  systems for  weak black liquor
oxidation in Europe,  Canada,  and the  United States since 1942. Yemchuk  (16)  and
Kacafirek (17) report  Na2S oxidation efficiencies from 85  to 95 percent with parallel
operation of Trobeck-Ahlen weak black liquor oxidation towers. Sylwan (18) reports Na2S
oxidation efficiencies of up to 98  percent for a single tower Trobeck weak black liquor
oxidation unit.

The higher oxidation efficiencies reported by Sylwan are at least partially the result of the
greater air-to-liquor flow ratio of 36 m3 air/m3 black liquor (4.8 cu ft air/gal black liquor)
as opposed  to the lower values reported for  the other systems of 15.0 m3 air/m3 black
liquor (2.0 cu ft air/gal black liquor).

Blosser and  Cooper (4) report on extensive experience with use of multiple tray-type black
liquor oxidation towers at several kraft pulp mills in the United States. Relatively high Na2S
oxidation efficiencies of 96 to 99 percent are noted for all units without excessive foaming
when the pine furnish  is  20 percent or less.  A summary of results is presented in Table 9-2.

Two recent installations, employing Trobeck-Ahlen weak  black liquor  oxidation systems,
have been modified to prevent excessive foaming at mills pulping substantial quantities of
pine wood species. Rippee (19) reports on a system for foam control at a western U.S. kraft
mill pulping approximately 40 percent pine wood species (primarily ponderosa and western
white). The liquid collected from the cyclone separator is drained to the tall oil recovery
system instead of the deaeration tank. A portion of the strong black liquor is recycled to the
inlet of the weak liquor oxidation tower to increase the solids concentration from 11 to 18
percent by  weight. Approximately 20 percent of the liquor  entering the  tower is strong
black liquor,  resulting in a substantial  decrease  in foaming problems. Na2S  oxidation
efficiency for the system has averaged 96 percent over an extended period.

Robinson (20) reports on modification of the design of a Trobeck oxidation unit to control
foam at a southern U.S. kraft mill pulping about 70 percent pine wood species. A portion of
the air is introduced at the tangential inlet of the cyclone separator to act as a piston for
controlling  the foam layer. The liquid stream passes through a series of deaeration tanks for
further foam suppression and soap  recovery. The Na2S oxidation efficiencies have averaged
approximately 98 percent over an extended period.
                                          9-4

-------
Ol
                            T
        BLACK
        LIQUOR
        INLET
                     STORAGE
                    ^ TANK
                                                        T
                                         —£Xh->
—00-
                                                 Cr
       (r
         AIR
        INLET
T
                                         CYCLONE
                                        SEPARATOR
                       OXIDATION
                         TOWER
                                                                         A  1  A
DEAERATION
                                               BLACK
                                               LIQUOR
                                               EXIT
                                              FIGURE  9-2

              TROBECK-AHLEN  MULTIPLE SIEVE TRAY WEAK BLACK LIQUOR OXIDATION SYSTEM (15)

-------
                                                    TABLE  9-2
  DESIGN AND  OPERATING PARAMETERS  FOR  TROBECK-AHLEN  MULTIPLE  SIEVE  TRAY WEAK BLACK LIQUOR
                                              OXIDATION UNITS  (4)
           Parameter

Liquor flow, m3/h (gpm)
Air flow, m3/h (cfm)
Plate area, m2 (ft2)
Na2S loading, Kg/m2/h (Ib/ft2/hr)
Liquor loading, m3/m2/h (gal/ft2/hr)
Air loading, m3/m2/h (cu ft/ft2/hr)
Na2 S inlet, g/1
Na2 S outlet, g/1
Oxid. effic., %Na2S
Oxygen ratio, act./theor.
Mill A
MillB
                                                                                     MillC
MillD
73 (320)
8,500 (5,000)
72 (780)
6.3 (1.3)
1.0 (24.6)
118 (388)
5.6
0.03
99
14.3
79 (350)
8,500 (5,000)
62 (665)
8.3 (1.7)
1.3 (31.5)
137 (451)
6.5
0.2
97
12.7
98 (430)
8,500 (5,000)
62 (665)
11.2 (2.3)
1.6 (38.6)
137 (451)
7.0
0.3
96
16.8
68 (300)
8,500 (5,000)
63 (672)
12.2 (2.50)
1.1 (26.7)
135 (444)
11.3
0.1
99
19.0

-------
     9.1.3  Packed Towers

 Packed towers are used for weak black liquor oxidation with air at kraft pulp mills in the
 Pacific  Northwest with varying and somewhat limited effectiveness. The basic principle of
 operation  is to  provide gas-liquid contact by  providing  a  large interfacial  surface area
 through the use of packing instead of foam. Two different types of packed tower systems
 for weak  black liquor oxidation with air have been developed,  one  by the Weyerhaeuser
 Company (21) (22) and the other by the British Columbia Research Council (23) (24).

 The Weyerhaeuser packed tower  weak  black  liquor  oxidation system uses concurrent
 downward vertical contact of air and black liquor in the tower to control foaming (25). The
 tower  is packed with conventional packing materials to provide for gas-liquid  interfacial
 contact area as shown in  Figure 9-3. Design  and operating  parameters for several actual
 systems are presented in Table 9-3 (4).

 The weak  black  liquor oxidation system of British Columbia Research  Council introduces
 air and liquor concurrently at the top of parallel towers using successive layers of vertically
 layered packing sheets. The sheets allow wetting of  both sides, and the spaces provide for
 lower pressure drops and resultant lower power requirements than do conventional packed
 tower  configurations.  Murray  (25) reports that vertically layered asbestos packing sheet
 allows effective gas-liquid  contact. West (26)  observes  that Na2S oxidation efficiency ap-
 proaches 100 percent at relatively low loading rates below 140 kg Na2S  per hour per m2  of
 packing area (29 Ib/hr per ft2) when two oxidation towers are located in a  series arrangement.

 Packed towers are relatively simple devices to construct and operate, and they have  minimal
 foaming problems and low horsepower requirements. They  can be  successfully operated
 when followed by a deaeration tank to prevent air entrainment in pumps. Most installations
 do not provide  sufficient retention time for nearly complete oxidation of Na2 S to occur,
 have insufficient capacity for the pulp production rates involved, and are  subject to plugging
 with pulp, particularly from continuous digesters.

     9.1.4   Agitated Air Sparging

Agitated air spargers are used for weak black liquor oxidation at a few kraft pulp mills. The
air sparger  system employs a completely mixed tank containing weak black liquor with air
 dispersed through a turbine aerator at  the bottom of the tank.  A rotary agitator, located
immediately above the aerator, shears the  air into small bubbles to maximize the gas-liquid
interfacial contact area. Additional power is provided to break  the air  into small  bubbles
instead  of  providing additional fan capacity and excess air. A typical unit is diagramed in
Figure  9-4.
                                         9-7

-------


—5?—
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FORCED
DRAFT
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T EXHAUST

BLACK
LIQUOR
INLET


PACKED
TOWER

4

AAA















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(jAi




DEAERATION
TANK
BLACK
fe. i i ssi I/-M-I

EXIT
                   WEYERHAEUSER  SYSTEM  (21)
                            AIR HEADER
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     BLOW-
     TANK
    HOT
   WATER
ACCUMULATOR
VEI!JT DILUTION
•=n J - w^
r i



	 -&L
FORCED
DRAFT
FAN
^ 	
T



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4



I
INLET
PACKED
TOWERS
AIR
VENT
                        OXIDIZED BLACK LIQUOR
                          TO EVAPORATION
    BRITISH COLUMBIA  RESEARCH  COUNCIL SYSTEM (23)


                        FIGURE 9-3

  PACKED TOWER SYSTEMS FOR WEAK BLACK LIQUOR OXIDATION
                            9-8

-------
                                                    TABLE 9-3
   DESIGN AND OPERATING PARAMETERS FOR  WEYERHAEUSER CONCURRENT FLOW PACKED TOWER WEAK
                                     BLACK LIQUOR  OXIDATION SYSTEM (4)
              Parameter

Liquor flow, m3/h (gpm)
Air flow, m3/h (cfm)
Packing area, m2 (ft2)
Packing volume, m3  (cu ft)
Tower height,  m (ft)
Na2S loading:
  Surface area, kg/103 m2/h (Ib/103ft3/hr)
  Volumetric, kg/103 m3 /h (lb/103 ft3 /hr)
Liquor loading:
  Surface area, m3/103m2/h (gal/103 ft2/hr)
  Volumetric, m3/103m3/h (gal/103ft3/hr)
Air loading:
  Surface area, m3/m2/h (cu ft/ft2/hr)
  Volumetric, m3/m3/h (cu ft/cu ft/hr)
Na2 S inlet, g/1
Na2S outlet, g/1
Oxidation efficiency, %
Oxygen ratio,  act./theor.
     Mill A

   102 (450)
 6,100 (3,600)
12,900 (139,000)
 1,300 (46,000)
   7.0 (23)
     MillB

    68 (300)
17,000 (10,000)
13,950 (150,000)
 1,050 (37,200)
  11.0 (36)
     MillC

   238 (1,050)
11,400 (6,700)
22,000 (237,000)
 2,410 (85,000)
  13.1 (43)
     MillD

   127 (560)
10,800 (6,400)
35,000 (376,000)
 3,370 (119,000)
  11.0 (36)
9.8 (2.0)
97.2 (6.0)
7.9 (0.2)
79 (0.6)
0.47 (1.54)
4.7 (4.7)
1.20
0.03
98
33.0
31.2 (6.4)
415.0 (25.9)
4.9 (0.1)
55 (0.4)
1.22 (4.00)
16.2 (16.2)
6.40
1.30
80
25.0
86.4 (17.7)
790.0 (49.2)
10.8 (0.3)
99 (0.7)
0.52 (1.70)
4.7 (4.7)
8.00
1.60
80
3.9
32.2 (6.6)
334.0 (20.8)
3.6 (0.1)
38 (0.3)
0.31 (1.02)
3.2 (3.2)
8.90
2.50
72
6.2

-------
                                 FOAM
                               BREAKER
             BLACK
             LIQUOR
             INLET
       AIR
                                                              EXHAUST
                                                                 AIR
. AGITATOR
          BLACK
      •>  LIQUOR
           EXIT
                          TURBINE
                          AERATOR
                                 FIGURE  9-4
    AGITATED  AIR SPARGING  SYSTEM FOR  BLACK  LIQUOR OXIDATION
Two major problems noted with agitated air sparging units for weak black liquor oxidation
are foaming caused by either low weak liquor solids content or by a wood furnish of greater
than 10 percent pine (27) and mechanical breakdown of the liquid agitator system.

Methods employed for foam control are:

     1.    Placement of mechanical foam breakers at the top of the tank to break up stable
         foam, requiring 0.08 to 0.16 kW per metric ton per day (0.1 to 0.2 hp per short
         ton per day).

     2.    Use of chemical or kerosene defoamer in the black liquor,

     3.    Recycling of strong black liquor to the  weak black liquor to increase  the net
         liquid viscosity, and

     4.    Reducing retention time with resultant lowered oxidation efficiency. A summary
         of performance data for two units is presented in Table 9-4 (4).
                                     9-10

-------
                                    TABLE 9-4
         OPERATING AND PERFORMANCE DATA FOR AGITATED  AIR
           SPARGED WEAK BLACK LIQUID  OXIDATION  SYSTEMS (4)

                   Parameter                      Mill A             Mill B

     Liquor flow, m3/h(gpm)                     243  (1,070)       262 (1,150)
     Airflow, m3/h(cfm)                        8,500  (5,000)      4,250 (2,500)
     Tank volume, m3 (ft3)                       200(7,060)       200(7,060)
     Retention time, min.                          49                 46
     Na2S loading, kg/m3/h(lb/ft3/hr)             3.64  (0.23)        8.75 (0.55)
     Liquor loading, m3/m3/h (gal/ft3/hr)           1.2  (9)            1.3 (10)
     Air loading, m3/m3/h (ft3/ft3/hr)             42.5               21.2
     Na2S inlet, g/1                                3.0                6.7
     Na2S outlet, g/1                              0.2                0.2
     Oxidation efficiency, %                        93                 97
     Oxygen ratio, act./theor.                      8.0                2.6
     Power required:
       Agitator, kW/t/day (hp/ton/day)            0.14  (0.17)        0.10 (0.12)
       Foam breakers, kW/t/day (hp/ton/day)      0.16  (0.19)        0.10 (0.12)
     9.1.5   Rotating Fluid Contactors

A limited amount of experience has been obtained with the Ashcroft dual vortex contactor,
primarily in polishing weak  black liquor to upgrade existing units. The system introduces
weak black liquor tangentially into an axially downward flowing pipe of small diameter. Air
is introduced to the liquid tangentially at the point where the small pipe connects with a
wider pipe, resulting in large shearing forces and rapid gas-liquid mixing (28).

Preliminary results indicate that the system is suitable for weak black liquor oxidation with
nonresinous wood species, but excessive foaming occurs for pine black liquor. One system
employed to upgrade the performance of an existing weak liquor oxidation system feeds air
into an Ashcroft unit at 5,000 m3/h (3,000 cfm) with the black liquor flowing at 227 m3/h
(1,000 gpm).  Installation of the  unit results in increasing the overall Na2S oxidation
efficiency during weak black liquor oxidation from the former 50 to 60 percent to between
95 and 99 percent.

9.2  Strong Black Liquor Oxidation—Air

Strong black liquor oxidation with air, following multiple-effect evaporation, is employed at
mills that pulp substantial quantities of resinous pine) wood species to alleviate potential
foaming  problems,  particularly in the southeastern United States. Strong black liquor

                                       9-11

-------
oxidation reduces malodorous sulfur gas emissions from the direct contact evaporator and
counteracts  the tendency for  Na2S to reform. The major types of systems employed for
strong black .liquor oxidation are single- and two-stage unagitated air sparging. Agitated air
spargers, plug flow reactors, and dual vortex contactors also are used for strong black liquor
oxidation to a limited extent.

     9.2.1   Single Stage Unagitated Air Sparging

The  major technique employed for strong black liquor oxidation, to date, is a completely
mixed unagitated air sparging tank. Hawkins (29)  (30) describes the development of the
original  system  at the Pasadena,  Texas, mill  of the Champion Paper Company. The system
employs aeration and deaeration tanks  arranged in series. The black liquor is oxidized by
sparging with air in  a single stage aeration  tank, followed liy separation of entrained air
bubbles from the liquid in the deaeration tank. The aeration tank consists of a cylindrical
section mounted between two conical sections, as shown in Figure 9-5.

Air is introduced in the' bottom of the  cylindrical section of the aeration tank through a
series of eight radially branching sprayer arms in a Christmas tree arrangement. The sprayer
arms are constructed of 20 cm (8 in) pipe with 19 nozzles for air outlet per branch and are
connected to a  central header. Each nozzle is 3.8 cm (1.5 in) in  diameter. The air is caused
to deflect downward against a deflector plate to achieve a fanning air curtain effect.

Air is introduced to  the central header at a rate of 10,000 m3/h (6,000 cfm). The exhaust
air is drawn through a  cyclone  separator to remove entrained foam  and liquor droplets,
which are then  returned to the aeration tank. The black liquor is introduced to  the top of
the tank through a 15 cm (6 in)  diameter pipe located  above the liquid level. The liquor is
sprayed in from the bottom of the pipe through a series of 2.5 cm (1 in) holes evenly spaced
along the pipe  to obtain  even liquid  distribution  and  to  provide a means for controlling
foam. The black liquor is  withdrawn from the bottom of the tank and  passed to two
deaeration tanks for gas-liquid separation to prevent pump malfunctions.

A Na2S oxidation efficiency of 97 to 98 percent occurs at an inlet concentration of 30 g/1
and at a liquid retention time of 2.5 hr. The process results in reducing H2 S emissions by 90
percent  from  the recovery furnace  following  direct evaporation as compared  to  H2S
emissions from unoxidized black  liquor.

Blosser  and  Cooper (4) have prepared an extensive review of strong black liquor oxidation
practices at  kraft pulp mills. It  is necessary to provide 3- to 5- times the stoichiometric
amount of air for oxidation of Na2S  to sodium thiosulfate (Na2S203). Air-to-liquid flow
ratios must  normally be greater  than 110 to 190 m3 air/m3 strong black  liquor (15 to 25
ft3/gal of black liquor.) It is normally necessary to maintain a minimum liquor depth of 2.7
to 3.6 m  (9 to  12 ft) with a  minimum  liquor retention  time of 120 to 150 minutes to
                                        9-12

-------
                           EXIT GAS
OJ
BLACK
LIQUOR
INLET
BLACK
LIQUOR
EXIT
LEGEND
1  HEAVY LIQUOR STORAGE TANK
2  LIQUOR INLET NOZZLES
3  OXIDIZED LIQUOR OUTLET
4  BLOWER
5  AIR SPARGER
6  CYCLONE SEPERATOR FOR
  AIR DISCHARGE
7  OVERFLOW
8  HEAD TANKS
                                              AIR
                                              COMPRESSOR

                                                      FIGURE 9-5

                 CHAMPION  UNAGITATED  AIR SPARGE STRONG BLACK LIQUOR OXIDATION SYSTEM (29)

-------
provide for effective oxidation of the Na2 S. The product of the oxygen ratio (actual oxygen
addition rate to theoretical oxygen addition rate) and liquor retention time in minutes must
be 600 or greater to reduce exit Na2S levels to less than 0.5 g/1 in the oxidation tower exit
liquid. The product of oxygen ratio and retention time must be greater than 1,000 to reduce
Na2 S levels in the exit liquor  to less than 0.2 g/1, as shown in Figure 9-6. These findings
point to  the possible development  of two-stage systems for strong black liquor oxidation
with air.

Morgan (31) (32) reports on the results of a study on a new strong black liquor oxidation
system. He finds that the efficiency of strong black liquor oxidation is a function of liquor
height, air flow rate, inlet Na2S concentration, and  retention time.  The rate of Na2S
oxidation increases with increasing  Na2S concentration and  air  flow rate, decreasing
retention time, and decreasing liquor height. Na2S oxidation efficiencies of up to 99 percent
are observed, but  oxygen transfer efficiency  is relatively low so that  large quantities of
excess air (3.5 times theoretical or more) are required to achieve high degrees of oxidation.

Foam control is particularly a problem during strong black liquor oxidation with air at kraft
pulp mills in the southeastern United States. Several control methods are available. Effective
soap removal during multiple-effect evaporation upstream of the oxidation unit removes
 CD
  i
 CO
  CM
  D
 CO
 UJ
 rr
                    200
400
600
800
1000
                           SUPPLY
                OXYGEN 	;	 X  RETENTION  TIME (minutes)
                           REQ T
                                   FIGURE  9-6
      OPERATING & PERFORMANCE DATA FOR  SINGLE  STAGE STRONG
                    BLACK LIQUOR  OXIDATION SYSTEM  (4)
                                        9-14

-------
foam-producing materials (33). Cyclone separators in the exhaust gas line remove entrained
foam and liquor droplets. Provision of adequate height  above the liquor level allows for
foam dissipation. Two to 5 m (6 to 15 ft) is the normal minimum required.

Mechanical and chemical methods of foam dissipation also are employed. Mechanical foam
breaking requirements for strong black liquor oxidation systems are  aeration tank only—0 to
0.022 kW per daily metric ton of pulp (0 to 0.027 hp/ton pulp per day) and aeration plus
deaeration tank—0 to 0.007 kW per daily metric  ton of pulp (0.0 to 0.1 hp/ton pulp per
day). Chemical defoaming agents used include diesel oil and kerosene in dosages from 50 to
250 1/m3  (50 to 250 gal/1,000 gal) strong black liquor, primarily on an intermittent basis.
Mechanical  foam breaking has  not  proved suitable  in many  cases  where  there  are
considerable  foaming problems. The cost of chemical defoaming runs as high as $0.55 to
$1.65/t ($0.50 to $1.50/ton) of pulp.

     9.2.2  Two Stage Unagitated Air Sparging

Limitations in achieving Na2S oxidation efficiencies above 99  percent to  reduce exit
concentrations of Na2S  to less  than  0.1 g/1  with single  stage  units  have  led to  the
development of multiple-stage strong black liquor oxidation systems. Padfield (34) reports
on  efforts to upgrade the single stage strong black  liquor oxidation system from the former
97-98 percent Na2S oxidation efficiency to  a desired 99-100 percent. A second aeration
tank  of design similar to the first  placed  after  the initial aeration tank provides  for
additional Na2 S oxidation. About a 60 minute retention time is provided in each  oxidation
tank  with an additional 30 minutes of deaeration in the bottom of each tank to  facilitate
liquor pumping. The exit  liquor from the second stage oxidation tank is sent directly to the
direct-contact evaporators to prevent reversion to sodium sulfide.

The system has two air blowers for adding air to the black liquor,  one for each tank. The
blower  for the first oxidation tank provides 10,400 m3/hr  (6,000 cfm)  of air at 143 kP#
total  pressure (1.41 atmospheres or 6 psig) with  a  187 kW (250 hp) motor. The blower for
the second oxidation tank provides air at 8,500 m3/hr (5,000 cfm) at the same pressure and
using the same power. Two 19 kW (25 hp) mechanical foam breakers are required, one on
each tank. Total power requirement for the two  stage system is thus 410 kW (550 hp) for a
pulp production of 770 Mg (850 tons) per day,  or about 0.53 kW  per daily metric ton of
pulp (0.65 hp/ton/day). The system is illustrated in  Figure 9-7 (34).

The average  overall Na2S  oxidation efficiency for the system averages 99.95  percent,
resulting in an average Na2S concentration in the exit black liquor of 0.02 g/1. Emissions of
H2 S from the recovery furnace are about 2 ppm by volume as compared to 50 ppm with the
original  single stage system.
                                        9-15

-------
 BLACK
 LIQUOR
  EXIT
                                                       STRONG BLACK
                                                      LIQUOR STORAGE
BLACK
LIQUOR
INLET
                                    FIGURE  9-7
        CHAMPION  TWO  STAGE  UNAGITATED STRONG BLACK  LIQUOR
                             OXIDATION  SYSTEM (34)
9.3  Agitated Air Sparging

The completely mixed agitated air sparging units employed for strong black liquor oxidation
are of the same design  as those described in section 9.1.4 for weak liquor systems. One unit
introduces 10,400 m3air/h (6,000 cfm) into a black liquor flow of 80 m3/h (350 gpm) at 48
percent solids using 150 kW (200 hp)  agitated turbine aerator (4). The aeration  tank has a
diameter of 9.5 m (31 ft) and a height of 7.9 m (26 ft), and is operated at a liquid depth of
3.7 m (12 ft). Five 19 kW (25 hp) mechanical foam  breakers are installed on the top of the
unit, but have severe maintenance problems. No definitive results regarding efficiency of the
unit  are available because of the lack of sufficient operating experience. The system requires
frequent maintenance because of agitator and foam breaker malfunctions.

9.4  Combination Systems

Tobias and Robertson  (35) describe the development of an agitated concurrent plug-flow
reactor  system  for oxidizing of strong black liquor with air. The primary purpose of the
system  is  to  polish the  strong black  liquor following multiple-effect  evaporation  to
counteract  the reversion to Na2S of oxidized weak  black liquor in an existing system. The
system employs one  oxidation unit  on the exit pipe of the strong black liquor storage tank,
and  the other on the strong black  liquor recirculation line, as shown in Figure 9-8. Each
reactor  is located so that  a pipe with a central  axially located baffle forms two mixing
chambers, each with  an agitator. Two parallel pipes are located immediately upstream of the
agitators in both chambers, with holes drilled so as to obtain a fanning air curtain effect.
                                        9-16

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                                                60% HEADER

                                                40% HEADER-
        BLACK LIQUOR INLET
                                   RECIRCULATION LINE
                                    IN-LINE
                                    OXIDIZER
                                                CYCLONE
                                                 L.C.V.
RETENTION TANK    STRONG BLACK
(RECIRCULATION)     LIQUOR TANK
                                                                                FURNACE
                                                                               EXIT GASES
                                                                             CYCLONE
                                                                            SEPARATOR
                                       FIGURE  9-8

   WESTERN KRAFT PIPELINE  REACTOR  STRONG BLACK LIQUOR OXIDATION SYSTEM (35)

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The  air bubbles are sheared into smaller bubbles by the rotating action of the agitators to
increase gas-liquid interfacial contact area, turbulence, and mixing.

Preliminary results  indicate that the Na2S oxidation efficiency for the system is virtually
100  percent with both units operating at inlet concentrations of 3 g/1 or less. The oxidation
efficiency is practically  100 percent with the system on the storage-tank exit line alone at
Na2S concentrations of 1.3 g/1 or less. The system also  appears to  have higher oxygen
utilization  efficiencies than conventional air sparged units. No major problems with foaming
are observed, but the wood furnish of the mill is primarily oak hardwoods. The two phase
flow system reduces liquor line plugging but  also causes substantial line vibration, requiring
secure fastening of the pipes. When the system was installed at a second mill, the operation
of the recovery furnace bed became  unstable, possibly because of sulfate formation with
resultant pH reduction and subsequent lignin precipitation from the black liquor.

Martin  (36) describes  reversion  phenomena  during multi-effect evaporation and  also
describes a secondary strong  black liquor  oxidation  system installed downstream of an
existing high efficiency weak black liquor oxidation unit. The system injects 680m3/h
(400 cfm) of air into the bottom of an existing storage tank through a series of eight vertical
pipes connected to a central header, where the black liquor flow is about 68 to 80 m3/h
(300 to 350 gpm). Results indicate that the Na2S concentration of the black liquor entering
the direct  contact evaporator could be maintained  at below 0.005 g/1  with the inlet Na2S
concentration below 1.5 g/1.

9.5  Molecular  Oxygen Systems

The  severe  foaming  problems  accompanying  oxidation of  weak  black  liquor  with
atmospheric oxygen (air) at kraft pulp mills pulping pine woods in the southeastern United
States led to considering use of molecular oxygen as an alternative. The foaming problems
of air  oxidation  can   be  alleviated  by  using  molecular  oxygen  because  the  inert
nitrogen-argon  diluting medium is no longer present. The major drawback  to using oxygen
for black liquor oxidation has been its cost, but recent trends are toward lower oxygen
prices, thus making it a more competitive alternative. Additional possible uses of oxygen
include pulp bleaching, chemical pulping, wastewater treatment, and addition to recovery
furnace firing zones for odor control.  Such applications may lead to further increases in
oxygen consumption at  a given mill.

     9.5.1   Preliminary Studies

Early laboratory investigations of Na2S oxidation  in black liquor with molecular oxygen
have been made by Bergstrom and Trobeck (37), Venemark (38), Ricca (39), Sakhuja.and
Bosu (40), and Miller (41). Major findings of these laboratory studies are that the sulfide
oxidation occurs in more than one rate-limiting regime, that Na2S203 is the major reaction
                                         9-18

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product, but that varying amounts of Na2 S04 are formed. The reaction rate is influenced
by temperature and catalyzed by the presence of organic constituents in the black liquor.
Cooper (42) reports on weak and strong black liquor oxidation with molecular oxygen in a
series of pilot scale experiments.  Design criteria are listed in Table 9-5 for plug flow reactor
systems.
                                    TABLE 9-5
          DESIGN CRITERIA OF PLUG  FLOW  REACTOR  SYSTEMS FOR
          BLACK  LIQUOR OXIDATION WITH MOLECULAR OXYGEN  (42)
                 Criteria

       Performance:
         Na2S efficiency, %
         Na2S outlet, g/1
       Liquor:
         Reynolds number
         Velocity, m/s (ft/sec)
         Temperature, °C (°F)
         Liquid pH
         Solids, % by wt.
       Oxygen:
         Oxygen ratio, Act/theoret.
         Total pressure, atm
         Partial pressure, % purity
       Retention time (Liquor basis):
         Piping section,  seconds
         Storage tank, minutes
     Weak
  Black Liquor
   Oxidation
      99+
   0.01-0.02

    100,000
1.5-4.5 (4.9-14.8)
77-88(170-190)
   12.0-12.6
     15-17

     1.1-1.3
     3.0-4.4
    90-99.5

     60-120
     15-45
     Strong
  Black Liquor
    Oxidation
       99+
    0.01-0.02

  10,000-20,000
 1.0-3.0(3.3-9.8)
110-115(230-239)
    12.0-13.0
      48-51

     1.2-2.5
     3.0-5.1
     90-99.5

     60-180
      30-60
Both weak and strong black liquor oxidation systems employ plug flow reactors followed in
series by tall storage tanks, as shown in Figure 9-9.

     9.5.2   Weak Black Liquor Oxidation

Weak black liquor oxidation of lightly  resinous pine black  liquors with oxygen allows
stabilization of Na2 S without causing the excessive foaming as with air. The process also has
a lower  capital cost than comparable air units because it can be carried out in a plug flow
reactor  within the piping of an existing mill. The  use of oxygen would  probably not
                                       9-19

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           OXYGEN
WEAK BLACK (
LIQUOR FROM V
WASHERS
J
( '

(

c


4,-J

                                          EXHAUST
                                            GAS
to
o
                                         WEAK
                                        LIQUOR
                                        STORAGE
"i    r
  i    i
  i    i
  !    i
 I   i
                                                                          ••i
                                                                           EXHAUST
                                                                             GAS
                                                           STRONG
                                                           LIQUOR
                                                          STORAGE

                                                           L.J    i ___
                                                                                       1
                                                                                            STRONG  BLACK
                                                                                            LIQUOR TO
                                                                                            DIRECT CONTACT
                                                                                            EVAPORATORS
                                                      MULTIPLE
                                                       EFFECT
                                                     EVAPORATORS
                MOLECULAR OXYGEN

                WEAK BLACK LIQUOR

                STRONG BLACK LIQUOR

                                                     FIGURE 9-9

                 TWO STAGE COMBINATION WEAK &  STRONG BLACK  LIQUOR OXIDATION WITH  OXYGEN

-------
 overcome the problems of Na2 S reversion in oxidized black liquor unless the process were
 carried out at a temperature above 130° C (266° F).

 Kosaya (43) describes the first reported use in 1956 of molecular oxygen for weak black
 liquor oxidation  at the Kehra  pulp mill in the Soviet Union. The  system was originally
 installed to reduce the H2S emission from multiple-effect evaporators, with the aim of
 improving the water quality and reducing odor. The pulp mill was a kraft mill that did not
 use recovery furnace flue gas for direct contact evaporation. The system injects oxygen into
 the black liquor upstream of the multiple effect evaporators through a "dosing apparatus."
 Results indicate that it is possible to achieve "essentially complete" oxidation of the Na2S
 in the black liquor at an inlet concentration of 7.5 g/1 within "several minutes of retention
 time" in the pipe at  a temperature of 70° C  (158° F). No problems with foaming were
 observed, and absorption of oxygen in the pipeline reactor was virtually complete.

 The study indicates that  the oxygen consumption is about 12 percent greater than the
 stoichiometric  amount for conversion  of Na2S to Na2S2O3, as shown in Table 9-6. An
 overall finding is that oxidation of weak black liquor with molecular oxygen is potentially
 attractive from an economic standpoint.
                                    TABLE 9-6
             OXYGEN REQUIREMENTS FOR  WEAK BLACK LIQUOR
                                 OXIDATION* (43)

                       Dimension               Theoretical       Actual

             m3 of 02/tof Na2S                   315           370
             (cuftofO2/tonof Na2S)            (10,080)      (11,850)

             kg of O2/tof pulp                      24            58
             (Ib of 02 /ton of pulp)                 (48)          ( 1 16)

             m3 of O2/m3 of black liquor**          2.1            2.5
             (cu ft of O2 /gal of black liquor)        (0.28)         (0.33)
              *Black liquor conditions: 15% solids, 7.54 g Na2S/l.
             **Based on black liquor flow of 8 m3/t pulp (1900 gal/ton pulp).
Freedman (28) reports on the oxidation of weak black liquor with molecular oxygen in
1970, when an Ashbrook rotating cyclonic fluid contactor was  used  to provide contact
between the oxygen and black liquor. The system introduces black liquor tangentially into a
                                        9-21

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small diameter pipe and redirects its flow axially to form a centrifugal cyclonic liquid flow
pattern.  Oxygen is then introduced tangentially into the liquid at the point of concentric
expansion into a larger diameter pipe. The violent mixing of the fluids causes a large number
of small oxygen bubbles to form. Collectively, these bubbles will provide a large interfacial
contact area between the gas and the liquid.

Pilot studies, using the contactor for oxidation of weak black liquors with oxygen, indicate
that effective  oxidation  of Na2S can  be achieved with liquors from  both  northern
hardwoods and  southern  pine  woods. Excessive foaming does not occur  in either case
because of  the absence of  the argon-nitrogen diluent from the  incoming gas stream.
Essentially complete oxidation of sodium sulfide can be obtained within 15 to 60 seconds
for hardwood black liquors, and within about 60 seconds for southern pine black liquors.
Oxygen usage efficiencies of 85 to 90  percent  are  obtained in both cases, making the
technique economically favorable. The exothermic oxidation also warms the black liquor by
up to 10°  C (18° F), thus decreasing evaporator stream heating requirements.

Galeano and Amsden (44) report on an extensive study of using molecular oxygen tor weak
black liquor oxidation at  the Owens-Illinois, Inc. kraft  pulp mill in Orange, Texas, where
southern pine wood species are primarily pulped. Basically, the system introduces molecular
oxygen into a flowing stream of black liquor in a two-step pipeline reactor with a retention
time of 15 to 40 seconds, followed by tank storage with a retention time of 8 to 12 hours.

Oxygen gas  enters the 61 cm (24 in) diameter horizontal inlet pipe at a total pressure of
0.45 MPa  (SOpsig) at a flow rate of 850  to  I,360m3/h (500 to 800 cfm) through an
injector oriented perpendicular to the liquid. It is broken up into small bubbles under high
turbulence by a venturi effect. The black liquor enters the inlet pipe at a flow rate of 227 to
386 m3/h (1,000 to 1,700 gpm) with a Na2S concentration of 9 to  12 g/1, a solids content
of 12 to 15 percent, and a temperature of 93 to 100° C (200 to 230° F). After contacting
the oxygen, the black liquor flows through a 30 m (100 ft) length of 25 cm (10 in) diameter
pipe, where the oxygen is  absorbed and the Na2S oxidized.  The  initial contact  section
provides for a liquid Reynolds number of 100,000 to 200,000 in highly turbulent  flow. The
system is diagramed in Figure 9-10.

Galeano and Amsden observed Na2 S oxidation efficiencies of 85 to 98 percent (94 average)
without excessive foaming with an oxygen usage efficiency of 75 to 95 percent (91 average).
About 90 percent  (of the total 94 percent)  of the Na2S is oxidized to  Na2S203 within the
pipeline reactor section. There  is a subsequent  conversion of about 15  percent  of the
Na2 S2 O3  to Na2 S04, as listed in Table 9-7.

Not all the  sulfur  is accounted for by the above chemical analyses alone, indicating either
the formation of other products, such as polysulfide, sulfite, or  polythionate ions, or loss by
gasification. The initial Na2S oxidation reaction is  extremely rapid. Additional changes in
                                        9-22

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                      OXYGEN
                     WEAK
                     BLACK
                     LIQUOR
                  t
MULTIPLE  EFFECT
 EVAPORATORS
GO
                       OXYGEN
                  V	
                     STRONG
                      BLACK
                     LIQUOR
                                                     EXHAUST
                                                      GAS
                                                       t
                                       V	
                                                   STORAGE
                                                    TANK
                                                                 EXHAUST
                                                                   GAS
                                                                   T
                                                                 STORAGE
                                                                  TANK
                                                                  OXYGEN
                                                               STRONG BLACK
                                                               LIQUOR TO
                                                               DIRECT CONTACT
                                                               EVAPORATOR
                                                    FIGURE 9-10
                       OWENS-ILLINOIS SYSTEM  FOR TWO  STAGE WEAK & STRONG  BLACK LIQUOR
                                             OXIDATION WITH  OXYGEN  (45)

-------
Na2S
g/1
4.56
0.45
0.19
Na2S2O3
g/1
0.80
4.35
4.61
Na2S04
g/1
0.53
0.70
0.63
Total
g/1
5.89
5.50
5.43
                                    TABLE  9-7
          EFFECT OF WEAK BLACK LIQUOR  OXIDATION  ON  LIQUID
                         CHEMICAL COMPOSITION  (44)

                                  Chemical Component
          Location       Na2S     Na2S2O,    Na2S04    Total     Account
       Reactor inlet       4.56       0.80        0.53       5.89       100.0

       Reactor outlet     0.45       4.35        0.70       5.50        93.7

       Storage outlet      0.19       4.61        0.63       5.43        92.8


black liquor during the oxidation process are a rise of 0.1  to  0.3 pH units because of
thiosulfate formation, and a temperature rise of  5  to 8° C (10  to 15° F) caused by the
exothermic oxidation.

Galeano  and Amsden  observed  several benefits of weak black  liquor oxidation with
molecular oxygen as compared to lack of oxidation, namely:

     1.   H2S emissions during direct contact evaporation declined by  95 to 99 percent
         during multiple-effect evaporation,

     2.   Na2 S04 makeup requirements declined by 30 kg/t pulp  (60 Ib/ton pulp),

     3.   Evaporator condensate water quality improved, and

     4.   Tall oil yield increased by about 15 percent.

The  effects of weak  black liquor  oxidation on sulfur gas emissions from  various process
sources are listed in Table 9-8.

The  mill is located adjacent to an  oxygen pipeline where the low  oxygen  cost of $9.35/t of
02 ($8.50/ton) results in a maximum net mill operating cost of $0.06 to $0.09 per  metric
ton  of pulp ($0.05  to  $0.08/ton). Projected capital costs for similar oxygen generating
installations  are $55 to $83 per daily metric ton of  pulping capacity ($50 to $75 per daily
short ton).
                                        9-24

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                                    TABLE 9-8
  EFFECT OF  WEAK BLACK LIQUOR OXIDATION  ON SULFUR GAS  EMISSIONS
                   FROM KRAFT MILL PROCESS SOURCES  (44)
           Source
 Cyclone evaporator exit
 Weak black liquor storage
   tank vent
 Evaporator noncondensable
   gas vent

 Evaporator condensate
    H2S
    RSH
RSR + RSSR
    S02

    H2S
    RSH
    RSR

    H2S
    RSH

   Sulfide
    BOD
                                                  Concentration
 Constituent    Unoxidized
   191
   161
    20
   243

   553
   760
  1,274

  1,208
»500

    72
   863
                Oxidized
                                           ppm, by vol.    ppm, by vol.
 10
  5
  0
241

  3
375
280

  4
300

 28
530
          Reduction
  95
  97
100
   1

  99
  50
  78

  99
>50

  60
  38
     9.5.3   Strong Liquor Oxidation

Very little  work has been devoted to the use of molecular oxygen for strong black liquor
oxidation. Galeano and Amsden (45) describe the use of strong black liquor oxidation with
oxygen for polishing to counteract reversion to Na2S from the previously oxidized weak
black liquor. The system introduces the oxygen into the strong black liquor in an expanded
pipe section at two places. The black liquor then flows downward to take advantage of the
gas bubble "holdup" phenomenon with its attendant increased effective retention time, and
then through a horizontal pipe  section to a strong black liquor storage  tank.  The liquid
Reynolds number in the pipeline reactor ranges from  10,000 to 20,000 in the liquid at 53
percent solids, at a temperature of 120° C (250° F), and with a liquid retention period of 30
to 50 seconds.

Initial  studies indicate potential problems with black liquor cooling, lignin oxidation, and
incomplete  oxygen  mass transfer into  the black  liquor, resulting in incomplete Na2S
oxidation. Little Na2S  reversion in the oxidized weak black liquor is observed, possibly
because of the high inlet temperatures, 93 to 110°  C (200 to 230° F), of liquor from the
continuous  digesters. The liquid retention time in each pipeline reactor stage is  about one
                                      9-25

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minute. The result is a reduction in the Na2S concentration in the strong black liquor from
1.5 g/1 to between 0.05 and 0.10 g/1.

     9.5.4   Digester Injection

It is possible to introduce oxygen into the kraft recovery system at the digester at the end
of the cook. Oxidation of the Na2S and sodium mercaptide (NaHS) at the end of the cook
prevents gaseous  conversion  and release of  these  compounds into the atmosphere. The
technique has the advantages of performing the oxidation in an enclosed reactor, thus
assuring  complete use  of the oxygen added,  and performing  the operation  at high
temperatures to assure maximum reaction  rates and minimal possibilities for reversion to
Na2S. The technique can also result in oxidation of organic sulfur compounds to prevent
their  discharge  and  to  limit  emissions  from  the  digester, washers,  multiple-effect
evaporators, tall oil and storage tank vents, and the direct contact evaporators.

Possible digester corrosion and excessive oxygen consumption can possibly occur because of
competing  side reactions, such as sulfate  formation and  lignin  oxidation and  potential
degradation in pulp quality. Two studies investigated addition of molecular oxygen to kraft
mill digesters at the ends of cooks to  oxidize the Na2S  and NaHS present in the black
liquor.

Fones and  Sapp (46) report on the addition of oxygen to a kraft digester to oxidize black
liquor  in 1960. Oxygen is added at the ends  of successive cooks to  a  pressure vessel
containing  pulp alone, black liquor alone, and a mixture of pulp and black liquor. For the
test  cook with  pulp  alone, the lignin  content of the pulp is reduced by oxidation, and its
brightness is increased.

When oxygen is injected into a digester containing both pulp and black liquor at the end of
the cook, the Na2 S is rapidly and completely oxidized. The high reaction temperatures and
pressures result in the formation of substantial quantities of Na2 SO4 with a resultant drop
in  liquid pH and a sharp increase in the amount of oxygen consumed. In addition,  the
oxidation process also reduces the bursting strength and brightness of the pulp, and its lignin
content. Oxidation of the black liquor with oxygen at the end of the cook is uneconomical
because of  the excessive oxygen requirements and the detrimental  effects on pulp quality.

Tests were  made to determine the effect of oxygen addition  on the oxidation of Na2S in a
pressurized vessel containing only black liquor. Oxygen was added to black liquor alone in a
digester at  a total pressure of 0.8 MPa (100 psig) and 150° C (302° F) in a series of stages to
recirculated black liquor.  The oxygen then reacted with the Na2S and other constituents
until the pressure returned to its initial value in a time span of four hours. Results indicate
that the initial reaction product was Na2S2O3, but substantial quantities of Na2SO4 were
also present causing a resultant drop in liquid pH. Because oxygen consumption was about
                                        9-26

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 double that for stoichiometric  conversion  of Na2S to  Na2S2O3, the  process  proved
 prohibitively expensive with oxygen costing $22/t ($20/ton).

 Kringsted and McKean (47) describe oxidation of Na2S and NaHS in black liquor at the end
 of a draft cook  in the presence of pulp. Their results indicate that by oxygen addition to a
 kraft digester at the end of a cook, the Na2S concentration in black liquor can be reduced
 by 90 percent and NaHS by 99 percent but twice the stoichiometric amount of oxygen is
 required.

 An additional finding is that no  sodium  polysulfide or Na2S03 could  be  detected in the
 black liquor during the oxidation period, indicating that it might be possible to reduce  or
 eliminate  potential problems caused by reversion to Na2 S. No measurements for Na2 S04
 were made during the tests.

 The  process does not affect the pulp strength or yield, but does make the pulp easier to beat
 and  also reduces its brightness. Oxygen addition to a digester at the end of a kraft cook can
 effectively oxidize Na2 S and  NaHS, and  can  minimize potential problems  caused by
 reversion.  Of particular importance is that the relatively  long retention time of about 20
 minutes at a high temperature of approximately 180° C (356° F) can result in the possible
 oxidation   of substantial amounts  of Na2S203 to Na2S04,  and  also  of  lignin.  The
 nonselective  oxidation  process can require  excessive quantities of oxygen in assuring
 oxidation  of the Na2 S and NaHS, thus making the process economically unattractive.

 9.6  Process Effects

Black liquor oxidation influences the  operation of the kraft chemical recovery system  in
other ways in   addition to reducing malodorous sulfur gas emissions.  Major  influences
include:

     1.   Improvement  in  multiple-effect evaporation  through reduced scaling  on  heat
         transfer surfaces,

     2.   Reduced corrosion rates of metal evaporating surfaces,

     3.   Possible increases in tall oil yield,

     4.   Reduced chemical makeup requirements for Na2SO4 and calcium oxide (CaO),

     5.   Possible increases in green and white liquor sulfidities, with resultant effects on
         pulp yield and quality, and

     6.   Possible effects on black liquor heating values.
                                        9-27

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     9.6.1   Evaporator Scaling

Berry (48)  finds that the H2S evolved from unoxidized  weak black  liquor causes the
formation of  an iron sulfide  (FeS) scale,  which inhibits the rate  of heat transfer  in the
multiple-effect evaporators. The precipitation  of lignin on the heat transfer surfaces  can be
minimized by maintaining a  sufficiently high liquid pH, thus alleviating another potential
source of evaporator scaling.  Any Na2S04  formed during weak liquor oxidation becomes
less soluble  as liquid temperature increases  and can cause a scaling problem during the latter
stages of evaporation at higher solids concentrations.

     9.6.2   Corrosion During Evaporation

Weak black liquor  oxidation can  substantially reduce the  corrosion  of  multiple-effect
evaporator surfaces, particularly in  the vapor shell sections.  Von Essen (49) reports that the
corrosion of evaporator surfaces is caused primarily by the formation of FeS from H2S in a
moist atmosphere on metal surfaces. The rate of corrosion is reduced by about 85 percent
by reducing the inlet Na2S concentration in black liquor to less than 3 g/1. Cyr and Harper
(50) report  that the average life of multiple-effect evaporator tubes is substantially increased
by weak black liquor oxidation.

     9.6.3   Tall Oil Yield

Weak and strong black liquor oxidation  can increase tall oil yields for byproduct recovery
by  either physical  or chemical mechanisms  (33).  Foaming,  however, could be a serious
problem with air oxidation  of pine black liquors  where  large  quantities of tall  oil are
obtained. Galeano and Amsden  (44) report that tall oil yield is increased by approximately
15 percent  as the result of weak black liquor oxidation with molecular oxygen. Also, there
is no apparent decrease in tall oil quality and there are increases in yield of as much as 7.5 to
12.5 kg per metric ton of pulp (15 to 25  Ib/ton). Rippee (19) reports an approximate  10
percent increase in tall oil yield at a western U.S. kraft pulp mill achieved by oxidizing weak
black liquor with air and recycling strong black liquor for foam control. At a 1972 National
Council symposium  (51), black  liquor oxidation  was reported to result in  increased tall  oil
yields. But the quality of the tall oil decreased, possibly because of its oxidation.

     9.6.4  Chemical Recovery

Black  liquor  oxidation reduces the sulfur losses from kraft process  sources, resulting in
decreased Na2S04 and lime chemical makeup requirements. Galeano and Amsden (44, 45)
report that  to maintain a given  sulfidity level the Na2SO4 makeup rate can be reduced  by
15  to 30 kg per metric ton of pulp (30 to 60 Ib/ton) of pulp. To provide a portion of the
sodium makeup  requirements, NaOH  can  be used. The lime makeup  requirement  can  be
reduced by 1.0  to  1.5 kg per metric ton of pulp (2-3 Ib/ton), and the total  lime  mud
                                         9-28

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processing rate can  be reduced by 5.0 to  7.5 kg per  metric ton of pulp produced (10 to
15 Ib/ton).  Specific chemical  savings  for maintaining  particular sulfidity levels will vary
between mills.

     9.6.5  Liquor Sulfidity

A major effect of black liquor oxidation on the kraft pulping process is the increase in green
and  white liquor sulfidities. The reason for the increased sulfidity levels is the retention of
additional sulfur in  the recovery system; less H2S is lost to the recovery  furnace flue gas
stream  during the direct contact evaporation (52).  The increase in green liquor sulfidity
results  in a decrease in both lime makeup and lime burning requirements,  with a resultant
decrease in fuel requirements at the lime kiln (44).

Black liquor oxidation can result  in increased white liquor sulfidity  levels of two  to five
percent or more  for a given Na2S04 chemical makeup  rate, or a reduction of 15 to 30 kg
per  metric  ton  of pulp  (30  to 60 Ib/ton) to maintain a  given sulfidity level  (14). The
increased white liquor sulfidity levels can increase the rate of delignification during digestion
of wood chips. The  greater sulfidities also can increase  the emission of malodorous organic
sulfur  compounds from  digester blow and  relief gases, brown stock washer vents, and
multiple-effect evaporator gases (53).

Shah and Stephenson (14) observed that the installation of weak black liquor oxidation
results in a 2- to 5-percent increase in white liquor sulfidity for given chemical makeup rates.
Increases in pulp quality and  yield were observed  during  the  digestion process, but they
probably are offset, at least in part, by increased corrosion caused by higher Na2S levels.
The  increased sulfidity may cause increased malodorous sulfur gas emissions from the kraft
recovery system because of the higher sulfur circulation rates.

One drawback to the usage of black liquor oxidation is that  it can cause the inlet Na2S
concentrations in the black liquor to rise sharply, thus overloading the oxidation capacity of
the  existing  system. Ritchey  (54) reports  that installation of a strong liquor  oxidation
system  using air at an existing mill resulted in about a 50 percent increase in the inlet Na2S
concentration in  the unit, that is, from about 40 to 60 g/1. The equivalent sulfidity increase
was  approximately 10 percent, resulting in a reduction in Na2S oxidation efficiency from
the expected 98 percent to about 80 percent. To maintain proper sodium-sulfur ratios in the
recovery system it is necessary to add NaOH as part of the sodium makeup.

     9.6.6  Energy Balances

Black liquor oxidation affects energy conservation in the kraft recovery system.  Lime kiln
fuel  requirements are reduced with increasing liquor  sulfidity levels (see section  9.6.5).
Oxidation of black liquor results in a reduction in net heating value of black liquor because
                                         9-29

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of the oxidation of both the, reduced sulfur compounds and the organic lignin fuel materials.
Losses in heating value can range from 2 to 6 percent (55). The oxidation of black liquor
can act to maintain cleaner heat transfer surfaces, thus reducing steam requirements during
multiple-effect evaporation. With air oxidation there is a slight evaporation of water, which
acts  to concentrate the solids. But  the  effect can be offset for  air by liquor  cooling,
particularly for strong black liquor.

The  oxidation of kraft black liquor  increases  the  efficiency of heat transfer by reducing
multiple-effect evaporator scaling. Weak black liquor oxidation with air increases the liquor
solids concentration by 1  to 2 percent by weight  because of the water  evaporated, thus
reducing  steam and fuel  requirements. Freedman  (28) observes that weak  black liquor
oxidation with pure oxygen results in a warming of the black liquor by 10° C  (18° F),
thereby reducing the sensible heat requirements during multiple-effect evaporation. There
are several potentially  detrimental effects. Miller (41)  observed during a series of pilot scale
studies that  oxidation with molecular oxygen caused  a cooling  of strong black liquor.
Tomlinson and Douglas (22) note that the endothermic heat requirements for reduction of
Na2S203 and Na2S04 in the kraft recovery furnace can result in slightly lower heat release
and  lower subsequent  steam generation for use in  digestion and evaporation sections. The
decreased heat release could cause possible increases in supplementary fuel costs.

Roberson (56, 57) observed that weak black liquor oxidation with air reduced its heating
value by about 2 percent, with a resulting decreased heat release. Lindholm and Stockman
(55) find that weak black  liquor  oxidation with  molecular oxygen  reduced  the liquor
heating value by 2.0 and 3.6 percent for Na2S oxidation efficiencies of 90  and 100 percent,
respectively.  The loss  in heating value is caused primarily by oxidation of organic matter,
resulting  in  increased  oxygen consumption and decreased heat availability.  Cooper (58)
observes that black liquor oxidation with oxygen can decrease the heating value of  weak
black liquor by 2 to 4 percent, and strong black liquor by 6 to 8 percent.

9.7  Air Pollution Effects

Weak black liquor oxidation has several beneficial effects on reducing malodorous sulfur gas
emissions from several sources,  including the  multiple-effect evaporator noncondensable
gases,  tall oil vents, storage tank vents, and the direct  contact evaporator of the  recovery
furnace.   Weak black  liquor oxidation  also  can   result in improvement  of evaporator
condensate water  quality to facilitate  process water reuse. Strong black  liquor oxidation
results in a reduction  in malodorous sulfur gas emissions from direct contact evaporation,
but  does  not provide any of the benefits for multiple-effect evaporation as does weak  black
liquor oxidation. Black liquor oxidation with air creates an additional source of odorous gas
emissions. These emissions are practically negligible when molecular oxygen systems are
employed.
                                         9-30

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     9.7.1   Direct Contact Evaporation

Direct contact evaporation is used at most kraft pulp mills to concentrate black liquor from
50  percent solids  to 60-65 percent solids.  Direct contact  evaporation  may  liberate
malodorous sulfur gases on contact of liquor with recovery furnace flue gas because the C02
of the flue gas may acidify the lignin sufficiently to generate H2S from Na2S. Black liquor
oxidation can minimize reduced sulfur gas emissions from the direct contact evaporator by
oxidizing Na2S to stable products in the liquid.

Major variables affecting  the emission rate of malodorous sulfur gases  are inlet Na2S and
NaHS concentrations, liquor pH, and degree of gas-liquid contact. Murray (59) observes that
increasing the  black liquor pH above  12.0 causes a substantial decrease in H2S emission
from  recovery  furnaces  following  direct contact  evaporation. Increasing the  degree of
gas-liquid  contact by increasing the pressure  drop  at high liquor firing rates can increase
total reduced sulfur (TRS) emissions from the direct-contact evaporator.

The primary variable affecting H2S emission from the direct-contact evaporator is the Na2S
concentration of the incoming black liquor. Oxidation of the Na2S and NaHS to innocuous
products in either  weak  black  liquor upstream or strong  black liquor downstream can
substantially  reduce  malodorous  sulfur gas  emissions.  This  reduction  is  particularly
significant for the multiple-effect evaporator noncondensable gases, tall oil vent gases, and
recovery furnace flue gases, and also evaporator condensates. The substantial reduction in
sulfur gas emissions from direct-contact evaporation is shown in Table 9-9.

Due  to  their growing relative significance, organic sulfur  emissions from direct  contact
evaporation will also be  important to consider as more  stringent emission standards  are
adopted. Methyl mercaptan (CH3SH) is oxidized in the presence of oxygen to CH3SSCH3,
which has a lower odor threshold level. CH3SCH3  and CH3SSCH3 can be removed by the
                                    TABLE  9-9
           EFFECT OF BLACK LIQUOR  OXIDATION ON SULFUR GAS
           EMISSIONS DURING DIRECT  CONTACT EVAPORATION  (60)

                                                 Liquor
          Sulfur Compound         Unoxidized               Oxidized
                                kg S7t     (IbS/ton)       kg S/l   (IbS/ton)

          H2S                 2.50-15.00  (5.0-30)       0.05-1.00 (0.1-2.0)
          CH3HS               0.15-1.00  (0.30-2.0)     0.02-0.10 (0.04-0.2)
          CH3SCH3             0.02-0.08  (0.04-0.16)    0.01-0.03 (0.02-0.06)
          CH3SSCH3            0.05-0.15  (.1-.3)         0.01-0.08 (0.02-0.16)
                                        9-31

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stripping action of the flue gases. The degree of removal depends on the inlet concentrations
in the liquid, the liquid and gas temperatures, and the pressure drop across the direct con-
tact evaporator. Polishing with oxygen at temperatures of 120 to 140° C (250 to 285° F)
may  be necessary to oxidize the organic sulfur constituents  in strong black liquor and,
therefore, prevent their release to the atmosphere.

     9.7.2  Evaporator Noncondensable Gases

Weak black liquor oxidation has substantially reduced H2S and CH3SH emissions from
evaporator noncondensable  gases, according to Douglass (61), Reid (62, 63), and Galeano
(44). The sulfur gas emissions are reduced by increasing black liquor pH and decreasing inlet
Na2S and NaHS concentrations. The effect of weak black liquor oxidation on malodorous
sulfur gas emissions from multiple-effect evaporator noncondensable gases is presented in
Table 9-10 (60).

                                  TABLE 9-10
              EFFECT OF WEAK BLACK LIQUOR OXIDATION ON
                MALODOROUS  SULFUR GAS EMISSIONS  FROM
                 EVAPORATOR  NONCONDENSABLE  GASES (60)

                                                Liquor
        Sulfur Compound          Unoxidized                Oxidized
                              kg S7t    (flTS/ton)        kg S/t    (IbS/ton)

        H2S                  0.05-1.40 (0.10-2.8)      0.00-0.01 (0-.02)
        CH3SH               0.05-0.50 (0.10-1.0)      0.05-0.10 (0.1-0.2)
        CH3SCH3             0.01-0.02 (0.02-0.04)     0.01-0.04 (0.02-0.08)
        CH3SSCH3           0.00-0.01 (0.0-0.02)      0.02-0.05 (0.04-0.10)

     9.7.3  Evaporator Condensate Waters

The oxidation of weak black liquor results in reducing the amounts of malodorous sulfur
gases liberated for subsequent absorption in the evaporator condensate waters. As reported
by  Kosaya (43) and Chisoni (64), two European mills employ weak black liquor oxidation
primarily to control odor levels in evaporator condensate waters and to gain, as a side effect,
control of water pollution.

Three additional studies have been made regarding the effect of weak black liquor oxidation
on  evaporator condensate water quality. Shah and Stephenson (14) find that the installation
of  a weak black liquor oxidation system reduced the BOD of the condensate water by 27
percent and substantially reduced  the odor level. The  liquid pH is raised from 6.5 to 9.0,
making the evaporator condensate  water suitable for process reuse as brown stock washer
                                       9-32

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 makeup water.  Galeano  and Amsden (44) note a 38  percent  reduction in BOD  and 60
 percent  reduction  in  sulfide ion concentration  with weak black liquor  oxidation  using
 molecular oxygen at a  southern U.S. kraft mill.

 Turner and Van Horn (65)  find that CH3OH contributes 64 percent of the total BOD of
 evaporator condensate water from evaporating unoxidized black liquor  at a southern kraft
 mill.  Galeano and  Amsden (44) speculate that one reason for the decrease in BOD of the
 condensate water is the partial oxidation of CH3OH to formaldehyde (HCHO) during weak
 black liquor  oxidation, along with the reduction in sulfide ion  concentration. Weak  black
 liquor oxidation may promote reuse  of evaporator condensate water by reducing the
 treatment required  for recycling.

     9.7.4   Tall Oil Vent Gases

 The oxidation of  weak black liquor reduces the inlet Na2 S concentration of the liquid
 supplied to the tall oil recovery system;  the H2S liberated during acidulation of the tall oil
 proportionately  declines.  Little information is available  on sulfur gas emissions from tall oil
 processing.

     9.7.5   Storage Tank Vents

 Black liquor  oxidation can reduce the emission of H2S and CH3SH from storage tank vents.
 Galeano and  Amsden (44) report reductions of 99 percent for H2S, 50 percent for CH3SH,
 and 78 percent for organic sulfides from the storage tank vents following weak black liquor
 oxidation with oxygen.

     9.7.6   Ambient Air Quality

 Hendrickson  and Harding (3) observe that the installation of parallel strong black liquor
 oxidation systems  substantially reduces the  malodorous  sulfur gas  concentrations in the
 ambient air near  the mill. The ambient level of odorous gas concentrations, as measured by
 "reducible sulfur" concentration (a general indicator of H2S levels), was 50 to 75 percent
 lower after installation of black liquor oxidation facilities.

 9.8  Oxidation Tower Emissions

 Black liquor oxidation facilities can reduce sulfur gas emissions from the direct contact and
 multiple-effect evaporators, tall oil vents,  and storage tanks. The use of weak or strong black
liquor oxidation with air does provide an additional source of reduced sulfur compounds to
the atmosphere.   Reduced sulfur  emissions from tank vents where black liquor is oxidized
with molecular oxygen are negligible in  comparison to those from  air oxidation systems
because the nitrogen-argon diluent is no longer present to cause any stripping.
                                        9-33

-------
Primary factors affecting sulfur gas emissions from black liquor oxidation tower vents when
air is used for oxidation are the inlet concentrations in the black liquor, the temperature of
the black liquor, the height of the black liquor in the tank, and the air flow rate per  unit
volume. Sulfur gas emissions tend to increase for higher liquid temperatures and for greater
air flow  rates because of the greater volatility of warmer  gases  and the stripping action of
the air. Sulfur gas emissions also tend to increase with increasing inlet concentrations in the
incoming black liquor. Use of hardwood species or contaminated  condensates for brown
stock washer  makeup can be the cause of increased inlet concentration.

The  primary  sulfur gas constituents present in black liquor oxidation tower exhaust are
organic sulfur compounds, such as  CH3SCH3  and CH3SSCH3. In  addition, other volatile
organic  nonsulfur  constituents can be stripped  from the black liquor, such as terpenes,
alcohols, and hydrocarbons. The emissions from weak and strong black liquor oxidation
systems are summarized in  Table 9-11 (66).
                                   TABLE  9-11
             REDUCED SULFUR EMISSIONS FROM  BLACK LIQUOR
                  OXIDATION  TOWER VENTS USING AIR (66)

      Condition         Weak Black Liquor             Strong Black Liquor
                    kg S/daily t   (Ib S/daily ton)     kg S/daily t  (Ib S/daily ton)

      Average           0.06          (0.11)            0.05          (0.10)

      Minimum         0.01          (0.02)            0.005         (0.01)

      Maximum         0.11          (0.22)            0.09          (0.18)
Large volumes of exhaust gases with high moisture content are generated during black liquor
oxidation. These gases can be incinerated in large boilers that have sufficient combustion
capacity,  provided that safety precautions are taken to accommodate the wet gases. Hisey
(67) describes a system at a kraft pulp mill  in South Africa where exhaust gases from the
black liquor oxidation tower are added to the primary air inlet of the recovery furnace. This
is the only installation known to date to employ this technique.

9.9  Process Economics

Two recent surveys of capital  and operating costs were  made  for weak and strong black
liquor oxidation systems that use air (4) (56). The  weak  systems normally have higher
capital costs than the strong ones for equivalent production because of greater liquid volume
                                        9-34

-------
at lower solids concentration. The strong liquor oxidation  systems normally have higher
operating costs than the weak ones of equivalent production because a greater amount of
energy is required for oxygen mass transfer into the viscous strong liquor.

     9.9.1   Air Oxidation Systems

Calculations by Roberson (56, 57) verify the higher capital  and operating costs for strong
black liquor oxidation systems than for equivalent weak ones. Roberson's design features
for a hypothetical processing  mill were an inlet Na2S concentration of 6.0 g/1 in the weak
black liquor or the  equivalent in strong  liquor  concentration,  and a  Na2S oxidation
efficiency of 99 percent across the system. Results  of the calculations are shown in Figure
9-11.

Blosser  and Cooper (4) present a compilation of capital and operating costs for existing
weak and strong black liquor oxidation systems from data supplied by individual mills. Capi-
tal cost figures are adjusted to a base of December 1968 from reported values to correct for
the effects of inflation and are listed in Table 9-12.
                                   TABLE 9-12
               ESTIMATED CAPITAL COSTS  FOR BLACK  LIQUOR
                              OXIDATION SYSTEMS

              Unit Description
      Location           Type              Installation Cost*        Reference
                                         $/dailyt (I/daily ton)

      Weak        Packed tower            440-660 (400-600)            4
                  Multiple tray            440-880 (400-800)            4
                  Agitated sparger         330-385 (300-350)            4
                  Rotating fluid            55-165 (50-150)            28

      Strong      Unagitated sparger       300-770 (275-700)            4
                  Agitated sparger**       440-660 (400-600)            4
                  Plug flow reactor**         6-55 (5-50)              35

       *Adjusted basis of December 1968.
      **Estimated values.
Annual operating costs for black liquor oxidation systems depend on a number of variables.
The major  component expense  is for electric power, but equipment maintenance, interest
                                       9-35

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 A.  CAPITAL COST
       500
                      6.00 gm/
                       50.0 % by wt
                             500                1000
                       PRODUCTION - TONS PULP/ DAY
 1250
 B.  OPERATING  COST
     50,000
                      500        1000        1500
                       PRODUCTION - TONS PULP/ DAY
2000
                          FIGURE 9-11

EFFECT OF PRODUCTION RATE ON CAPITAL & OPERATING COSTS  FOR
   WEAK & STRONG BLACK LIQUOR  OXIDATION WITH OXYGEN (56)
                             9-36

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on invested capital, and depreciation must also be included. A summary of direct annual
operating costs for black liquor oxidation systems is presented in°Table 9-13.

                                 TABLE  9-13
          APPROXIMATE ANNUAL OPERATING COSTS FOR  BLACK
                 LIQUOR OXIDATION SYSTEMS USING AIR
      Black Liquor
      Weak
      Strong
    Power Requirement
 kW/daily t     hp/daily ton     S/daily t

 0.16-0.49       0.2-0.6        11-55
                Operating Cost*
 0.41-0.82
0.5-1.0
22-220
S/daily ton

  10-50

  20-200
      *Does not include operating cost credits.
Operating variables for systems include inlet Na2S concentration; liquor depth in the tank;
auxiliary facilities, such as foam breakers and agitators; and the possible need for chemical
addition for  foam control or pH adjustment. An additional factor is whether a system is
single- or multiple-staged. Sheppard (68) reports a 30 percent decrease in annual operating
costs in converting from a single- to a double-stage strong black liquor oxidation system, as
listed in Table 9-14.
                                 TABLE 9-14
     EFFECT  OF NUMBER OF STAGES ON ANNUAL OPERATING COSTS
         FOR STRONG BLACK LIQUOR OXIDATION WITH AIR (68)
     Number
       of
      Stages
Flow
                 m3/h    (cfm)
                76,500 (45,000)
  Power
                kW    (hp)
                41,000 (24,000)      1,870  (2,500)

                35,800 (21,000)      1,810  (2,430)
  Annual Operating Cost
  $/dailyt  (I/daily ton)
               2,720  (3,650)      100.50     (91.15)
                                  69.25     (62.81)

                                  67.00     (60.77)
                                     9-37

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     9.9.2   Molecular Oxygen Systems

The  capital  costs  for  black liquor  oxidation  systems  using  molecular oxygen  are
considerably  lower than  for  air  oxidation  systems.  This is because  molecular oxygen
processing can be carried out within the piping of an existing mill without constructing large
separate tanks.  Galeano and Amsden (44) estimate the cost for two stage weak black liquor
oxidation system at $55 to  $83 per metric ton of installed daily capacity ($50 to $75 per
ton/day) and the cost for strong black liquor polishing at $83 to $165 per metric ton of
installed  daily capacity ($75 to $150  per ton/day).  Cooper (58) estimates capital costs for
black liquor oxidation with oxygen,  depending on the materials of construction used, as
listed in Table 9-15.

The  primary factors affecting operating costs for strong black liquor oxidation with oxygen
are the inlet Na2S concentration and the oxygen unit purchase cost. Normally, 15  to  25
percent excess oxygen must be added to provide for maximum Na2 S oxidation efficiency.
An additional  5 to  10 percent more oxygen than that required for  weak black liquor
oxidation alone must also be added for strong black liquor polishing. The effect of oxygen
unit costs  and  inlet  Na2 S  concentration  on unit  operating costs for  weak black liquor
oxidation with oxygen is presented in Figure 9-12 (66). A summary of calculated ranges in
operating cost  credits and debits for black liquor oxidation with oxygen is presented in
Table 9-16  (69).

A summary of  estimated operating parameters, reduced sulfur emissions, and  capital and
operating cost  values for  weak and strong black liquor  oxidation with air and oxygen is
presented in Table 9-17.
                                         9-38

-------
     1.50
O-

O

(A
 I
o
o
QC
UJ
Q_
O
     1.00  -
0.50 -
    0.00
                3,000 Ib DSATP
                2,040 gal BL/TP
                15% So I ids
                Oxygen Ratio = 1.20
                Na2S Concentrations
                    gram/liter
                                    10
                                             15
                                                      20
                                                               25
                                                                        30
                                                                           35
                                 OXYGEN PR ICE-DOLLARS/TON

                                 FIGURE  9-12

        OPERATING  COSTS FOR  WEAK BLACK LIQUOR  OXIDATION
                              WITH  OXYGEN  (66).
                                      9-39

-------
 Reactor Section
                                                       TABLE 9-15
                       ESTIMATED CAPITAL COSTS  FOR BLACK LIQUOR OXIDATION  SYSTEMS
                                              USING MOLECULAR OXYGEN*
Weak Liquor System
                                                                                            Strong Liquor System
Carbon Steel
 Stainless Steel
                                           Carbon Steel
                                                           (All costs in I/daily t (I/daily ton))
                                                                                                            Stainless Steel
Liquid pumping
Oxygen injection
Piping section**
Storage tank
Total
  2-3 (2-3)
  1-2 (1-2)
  2-6 (2-5)
11-22 (10-20)
16-33 (15-30)
  2-3 (2-3)
  1-2 (1-2)
13-17 (12-15)
28-55 (25-50)t
44-77 (40-80)
1-2 (1-2)
1-3 (1-3)
3-6 (3-5)
6-11 (5-10)
1-2 (1-2)
1-3 (1-3)
6-11 (5-10)
14-28 (13-25)1"1"
                                           11-22 (10-20)
                                                                                                           22-44 (20-40)
 *Cost data from Popper, H. (ed.). Modern Cost Engineering & Techniques. New York, McGraw-Hill Book Co., 1970. p. 80-178.
**Liquid retention times of 30 to 60 seconds.
 'Liquid retention times of 15 to 30 seconds.
ft
 '
  Liquid retention times of 30 to 60 seconds.

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                                   TABLE 9-16
     OPERATING COSTS AND OPERATING CREDITS FOR BLACK  LIQUOR
                        OXIDATION WITH OXYGEN (69)
Cost or
 Credit
Cost
Cost
Cost
          Item
Oxygen:
  Weak Black Liquor
               Strong Black Liquor
Electric Power
Operating and Maintenance
       Total Costs
Credit
Credit
Credit
Credit
Tall Oil Yield
Sodium Sulfate
CaO
Kiln fuel savings
        Amount
      10-60 kg 02/t
   (20-120 Ib02/ton)

     0.5-7.5 kg 02/t
    (l-151bO2/ton)

  0.04-0.12 kW pert/day
(0.05-0.15 hp per ton/day)

      0.5-1.Oh/wk
        0-13kg/t
      (0-26 Ib/ton)

        0-15 kg/t
      (0-30 Ib/ton)

      0.5-2.5 kg/t
      (1-5 Ib/ton)

         3-10%
       Total Credits
Cost/Credit
   Range
$/t ($/ton)
 0.11-1.65
(0.10-1.50)

 0.01-0.28
(0.01-0.25)

 0.01-0.03
(0.01-0.03)

 0.01-0.02
(0.01-0.02)

 0.14-1.98
(0.13-1.80)

 0.0-1.10
 (0.0-1.05)

 0.0-0.33
 (0.0-0.30)

 0.01-0.06
(0.01-0.05)

 0.03-0.11
(0.03-0.10)

 0.04-1.66
(0.04-1.50)
                                      9-41

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to
                                                         TABLE 9-17
      TYPICAL RANGES IN  OPERATING VARIABLES, REDUCED SULFUR EMISSIONS  AND COST FACTORS FOR BLACK
                                             LIQUOR OXIDATION SYSTEMS (69)
              Item

Operating Variables:
  Oxygen reqm't., act/theor.
  Power reqm't., kW/daily
               (hp/daily ton)
  Na2 S to direct contact evaporator,
    g/1
  Na2S oxidation efficiency, %

Reduced Sulfur Emissions:
  Evaporator gases, kg/t
  Tall oil vent, kg/t
  BLO tower vent, kg/t
  Recovery furnace, direct contact
    evaporator, kg/t

Economic Factors:
  Capital cost, $/daily t
             (I/daily ton)
  Annual operating cost, S/daily t
                     (I/daily ton)
                                                Air Oxidation Systems
                                                                                 Molecular Oxygen Systems
WBLO*
Only
4-6
0.2-0.8
(0.3-1.0)
0.3-1.5
97-99
0.05-0.5
0.05-0.1
0.05-0.25
0.1-1.5
330-880
(300-800)
11-110
(10-100)
SBLO**
Only
5-8
0.4-1.2
(0.5-1.5)
0.02-0.1
95-99
0.5-5.0
0.5-0.75
0.05-0.15
0.05-1.0
550-880
(500-800)
27-220
(25-200)
WBLO&
SBLO
7-10
0.6-1.5
(0.7-1.8)
0.01-0.5
99-99"
0.05-0.5
0.05-0.1
0.05-0.3
0.05-0.5
660-1100
(600-1000)
55-275
(50-250)
Digester
Only
2-3
0.1-0.2
(0.1-0.3)
0.05-0.5
90-99
0.025-0.25
0.025-0.1
0
0.025-0.5
11-55
(10-50)
33-550
(30-500)
WBLO
Only
1.1-1.3
0.02-0.08
(0.03-0.10)
0.3-1.5
98-99
0.05-0.25
0.025-0.1
0-0.01
0.05-1.5
27-137
(25-125)
11-330
(10-300)
SBLO
Only
1.4-1.7
0.02-0.04
(0.02-0.05)
0.1-1.0
96-98
0.05-5.0
0.5-0.75
0.-0.005
0.1-1.5
55-220
(50-200)
22-440
(20-400)
WBLO &
SBLO
1.2-1.5
0.04-0.12
(0.05-0.15)
0.01-0.10
99-100
0.05-0.25
0.05-0.1
0.025-0.25
0.05-0,25
55-165
(50-150)
17-385
(15-350)
      *WBLO = Weak Black Liquor Oxidation.
     **SBLO = Strong Black Liquor Oxidation.

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9.10 References

  1.  Collins,  T. T., The Oxidation of Sulfate Black Liquor and Related Problems. Tappi,
     38:172A-175A, August 1955.

  2.  Landry, J., Black Liquor  Oxidation  Practice and Development—A Critical  Review.
     Tappi, 46:766-772, December 1963.

  3.  Hendrickson,  E. R.,  and Harding, C. J., Black Liquor  Oxidation as a  Method  of
     Reducing Air  Pollution from  Sulfate Pulping.  Air Pollution  Control Association,
     14:487-490, December, 1964.

  4.  Blosser,  R. 0., and Cooper, H. B. H., Survey of Black Liquor Oxidation Practices in the
     Kraft Industry. NCASI Atmospheric Pollution Technical Bulletin  No. 39. National
     Council  of the Paper Industry for Air and Stream Improvement, Inc., New York, New
     York, December 1968.

  5.  Trobeck, K. G.,  The  B T System for Soda and Heat Recovery  in Sulfate Pulp Mills.
     Paper Trade Journal, 133(15):40-48, April 20, 1960.

  6.  Collins,  T. T., The Oxidation of Sulfate Black Liquor and Related Problems. Tappi,
     38:172A-175A, August 1955.

  7.  Bialkowsky, H. W., and Dellaas, C. G., Stabilization of Doughs Fir Kraft Black Liquor.
     Paper Mill News, 74(35): 14-22, September 1, 1951.

  8.  Wright,  R. H., and Klinck, R. W.,  What Port Alberhi Has Done to Control Kraft Mill
     Odors. Paper Trade Journal, 139(41):22-24, October 1955.

  9.  Walther, J. E., and Amberg, H. R., Odor  Control in the Kraft Pulp Industry. Chemical
     Engineering Progress, 66:73-80, March 1970.

10.  Freedman, H. L., A Different Approach  to  Oxidation  of Bhick Liquor. Paper Trade
     Journal, 154(23):50-51, June 8, 1970.

11.  Tirado, A. A.,  Guevara, M. V., and Banduni,  J. S., Oxidation of Black Liquor Under
    Pressure. Air Pollution Control Association, 12:34-37, January 1962.

12. Collins, T. T., Some Aspects of Oxidizing Sulfate Black Liquor. Paper Trade Journal,
     130(3):37-40, January 19, 1950.

13. Van Donkelaar, A., Air Quality Controls in  a Bleached Kraft Mill.  Pulp and Paper
    Magazine of Canada, 69(18):69-73, September 20, 1968.

                                        9-43

-------
14.  Shah, I. S., and Stephenson, W. D., Weak Black Liquor Oxidation: Its Operation and
     Performance. Tappi, 51:87A-94A, September 1968.

15.  Trobeck, K. G., Some Data on the Oxidation of Black Liquor. Paper Trade Journal,
     135(l):27-31,July4, 1952.

16.  Yomchuk, E. M., Oxidation of Black Liquor at the Great Lakes Paper Company, Ltd.
     Pulp and Paper Magazine of Canada, 71(14):45-50, July 17, 1970.

17.  Kacafirek, S., Kubes, J., and Racek, J., Practical Experience with Oxidation of Black
     Liquor at the Steti Pulp Mill. Papir a Celuloza, 23:194-196, July 1968.

18.  Sylwan, 0., Practical Results  Obtained with Black Liquor Oxidation.  Paper Trade
     Journal, 137(10): 14-16, September 4, 1953.

19.  Personal communication with Mr. Joseph Rippee, Potlatch Forests, Inc., Lewiston,
     Idaho, November 1970.

20.  Personal communication with  Mr. Wayne Robinson,  Eastex, Inc.,  Silsbee, Texas,
     December 1972.

21.  Tomlinson, G. H., Tomlinson, G. H., Jr., Swartz, J. N., Orloff, H. D., and Robertson, S.
     H., Improved Heat and Chemical Recovery in the Alkaline Pulping Processes. Pulp and
     Paper Magazine of Canada, 47:71-77, August 1946.

22.  Tomlinson, G. H.,  and Douglas, H. R., A Progress Report on the Secondary Recovery
     of Heat and Chemicals in the Alkaline Pulp Mill. Pulp and Paper Magazine of Canada,
     53:96-104, March 1952.

23.  Wright, R. H.,  The Effect of Packing Type on  the Rate of Black Liquor Oxidation.
     Tappi, 36:85-88, February 1953.

24.  Wright, R. H., and Klinck, R. W., What Port Alberni Has Done to Control Kraft Mill
     Odors. Paper Trade Journal, 139(41):22-24, October 10, 1955.

25.  Murray, F.  E.,  A Study of Black Liquor Oxidation in Towers Packed with Asbestos
     Cement  Plates. The  Canadian Journal of  Chemical Engineering, 36(2):69-72, April
     1953.

26.  West, W. B., Improving Black Liquor Oxidation Efficiency  of Packed  Towers. Tappi,
     43:192A-194A, October 1960.
                                       9-44

-------
 27,  Scott,  C. W., Weak Black Liquor Oxidation to Reduce Air Pollution  with  Foam
     Concentration of Soap and  Increased Soap Recovery. 'Southern  Pulp and  Paper
     Manufacture, 33(l):26-27, January 10, 1970.

 28.  Freedman,  H. L., A  Different Approach to Oxidation of Black Liquor. Paper Trade
     Journal, 153(23):50-51, June 8, 1970.

 29.  Hawkins, G., Black  Liquor Oxidation  at Champion's  Texas Mill has Unusual  Twist.
     Paper Trade Journal,  146(10):38-39, March 5, 1962.

 30.  Hawkins, G., Air Pollution Control at Champion Papers, Inc., Pasadena Mill, Texas.
     NCASI Atmospheric Pollution Technical Bulletin No. 26, National Council of the
     Paper Industry for Air and Stream Improvement, New York, New York, August 1965.

 31.  Morgan, J. P., Sheraton, D. F., and Murray, F. E., The Effect of Operating Variables on
     Strong  Black Liquor Oxidation.  Pulp and Paper Magazine  of Canada, 71(6):48-51,
     March 20, 1970.

 32.  Morgan, J. P., Sheraton, D. F., and Murray, F. E., The Effect of Operating Variables on
     Strong Black Liquor Oxidation. Paper Trade Journal, 154(1):41, January 5, 1970.

 33.  Ellerbe, R.  W., Why, Where, and How U.S. Mills Recover Tall Oil Soap. Paper  Trade
     Journal, 157(26):40-43, June 25, 1973.

 34.  Padfield,  D. H., Control of Odor From Recovery Units by Direct Contact Evaporative
     Scrubbers with Oxidized Black Liquor. Tappi, 56:83-86, January 1973.

 35.  Tobias,  R.  C.,  Robertson,   G.  C.,  Schwabauer,  D.  E.,  and  Dickey,  B.,  A
     Non-Conventional Strong Black Liquor Secondary Oxidation Treatment. (Presented at
     the West  Coast Regional Meeting of the National Council of the Paper Industry for Air
     and Stream Improvement.  Portland, November 4, 1970.)

 36.  Martin, F.,  Secondary Oxidation Overcomes Odor  from Kraft Recovery. Pulp and
     Paper, 43:125-126, June 1969.

 37.  Bergstrom, H., and Trobeck, H. G., Analysis of Black Liquor. Svensk Papperstidning,
     39(22):554-557, November 30, 1939. (Stockholm)

38.  Venemark,  E.,  On   the   Oxidation  of  Black  Liquor.   Svensk   Papperstidning,
     59(18):629-640, September 1956. (Stockholm)

39.  Ricca, P.  M., A Study  in the Oxidation  of Kraft Black Liquor.  Ph.D. Dissertation,
     University of Florida, Gainesville, February 1962.

                                       9-45

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40. Sakbuga, L., and Bosu, S., Studies on the Fixation of Sulfide Sulfur in Sulfate Black
    Liquor. Indian Journal of Technology, 6:149-152, May 1968.

41. Personal communication with Mr. Roy L. Miller, St. Regis Paper Company, Pensacola,
    Florida, February 1972.

42. Cooper, H.  B. H., and Rossano, A. T., Jr., Black Liquor Oxidation with Molecular
    Oxygen in a Plug  Flow Reactor.  Tappi,  56:100-103, June 1973.

43. Kosaya, G.  S., Black Liquor  Oxidation with Oxygen. Bumayhnoya Promyshlennost,
    31:15, June 1956.

44. Galeano, S. F.,  and Amsden, C. D., Oxidation  of Weak Black Liquor with Molecular
    Oxygen. Tappi, 53:2142-2146, November  1970.

45. Owens-Illinois Odor Reduction Oxidation System Made  Available. Southern Pulp and
    Paper Manufacturer, 36(3):38, March 10, 1973.

46. Fones,  R. E., and Sapp, J. E., Oxidation of Kraft Black Liquor with Pure Oxygen.
    Tappi, 43:369-373, April 1960.

47. Kringstad,  K. P.,  McKean, W. J.,  Libert, J., Kleppe, P. J., and Laishong, C., Odor
    Reduction  by In-Digester  Oxidation  of Kraft Black  Liquor with Oxygen.  Tappi,
    55:1528-1533, October 1972.

48. Berry, L.  R., Black Liquor Scaling in Multiple Effect Evaporators. Tappi, 49:68A-71A,
    April 1966.

49. Von Essen, C. G., Corrosion Problems in  Sulfate Pulp Mills. Tappi, 33:14A-32A, July
     1950.

50. Cyr,  M. F., and Harper, A. M., Multiple Effect Evaporator Project. Pulp and Paper
    Magazine of Canada, 61:T247-T249, April 1960.

51. NCASI and Members Host Symposium on  Black Liquor Oxidation. NCASI Monthly
    Bulletin, 10:1-3, February-March 1972.

52. Staidl,  J. A., and Schmitt, M. G., Some Practical Aspects of the Chemistry of Sulfur in
    the Kraft Recovery Process. Paper Mill and  Wood Pulp  News, 61(44): 12, October 29,
     1938.

53. Sarkanen, K. V., Hrutfiord, B. I., Johanson, L.  N., and Gardner, H. S., Feature Review
    Kraft Odor. Tappi, 53:766-783, May 1970.

                                       9-46

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 54.  Personal communication with Mr. Rick Ritchey, Southland Paper Mills, Inc., Lufkin,
     Texas, October 1972.

 55.  Lindholm,  I.,  and Stockman, L., Heat Evolution during  Black Liquor Oxidation.
     Svensk Papperstidning, 65(19):755-759, October 15, 1962.

 56.  Roberson, J. E., How Does Recovery Odor Control Affect a Kraft Mill Energy Balance.
     Pulp and Paper, 43:151-154, November 1969.

 57.  Roberson, J.  E.,  The Effect  of Odor Control on a Kraft  Mill Energy Balance.  Air
     Pollution Control Association, 20:373-382, June 1970.

 58.  Cooper, H. B. H., Black Liquor  Oxidation with Molecular Oxygen in a Plug Flow
     Reactor. Ph.D. Dissertation, University of Washington, Department of Civil Engineer-
     ing, Seattle, Washington, August 1972.

 59.  Murray, F. E., and Rayner, H. B., Emissions  of Hydrogen Sulfide from Kraft Black
     Liquor during Direct Contact Evaporation. Tappi, 48:588-593, October 1965.

 60.  Hendrickson,  L.  R.,  Roberson,  J. E., and  Koogler, J.  B.,  Control of Atmospheric
     Emissions  in the Wood Pulping Industry. Volume I. Final Report, Contract No. CPA
     22-69-18, U.S. Department of Health, Education, and Welfare, National Air Pollution
     Control Administration, Raleigh, North Carolina, March 15, 1970.

 61.  Douglas, I. B., Sources of Odor in the Kraft Process, III. Odor Formation in Black
     Liquor Multiple Effect Evaporators. Tappi, 52:1738-1742, September 1969.

 62.  Reid, H. A., The Odour Problem at Maryvale. APPITA, 3(2):479-500, December 1949.

 63.  Reid, H.  A., Soda Recovery  and Losses in Kraft  Pulping.  APPITA, 4(3):338-360,
     December 1950.

 64.  Ghisoni, P., Elimination of Odors in a Sulfate Pulp Mill. Tappi, 37:201-205, May 1954.

 65.  Turner, B. G., and Van Horn,  J.  1., Identification of Volatile Compounds in Kraft Mill
     Evaporator Condensates. (Tappi Southeastern Section Meeting, Atlanta, March 1969.)

 66.  Blosser, R. 0., Miscellaneous Sources and Tre'nds in Kraft Emission Control: Overview.
     Tappi, 55:1189-1191, August 1972.

67.  Hisey, W. 0.,  Abatement  of Sulfate Pulp Mitt Odor and Effluent Nuisances.  Tappi,
     34:1-6, January 1957.
                                       9-47

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68. Sheppard, M., Design of a High Efficiency Heavy Black  Liquor Oxidation System.
    (Presented at the Tappi Environmental Conference, San Francisco, May 15, 1973.)

69. Cooper, H. B.  H., Recent Developments and Future Trends in Black Liquor Oxidation.
    (Presented at the Tappi Environmental Conference, San Francisco, May 15, 1973.)
                                        9-48

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                                   CHAPTER 10

                 RECOVERY BOILER DESIGN AND OPERATION


 10.1   General Conditions
                                                                                  ^
     10.1.1   Process Parameters Outside Recovery Boiler

 Recovery boilers  are  used  for  the combustion of  spent liquor from  the  following
 sodium-base pulping processes:

     1.  Sodium sulfate (kraft)

     2.  Sodium bisulfite, and

     3.  Neutral sulfite semi-chemical (NSSC) in combination with kraft.

 The kraft process can be  used for making paper  grade pulp or, when combined with a
 prehydrolysis stage, this process can be used to make dissolving pulp. The pH of the bisulfite
 process can  range from acidic to  alkaline at the finishing stage of the cooking. The NSSC
 process also has a certain pH range.

 The amount of black liquor and its characteristics depend to a great extent  on the type of
 pulping used (owing to differences in the cooking liquor and the pulp yield), the species of
 wood used for pulping and the site of the growth of the wood.

 The part of the wood no longer present in the pulp after cooking is converted to:

     1.  Organic and inorganic components in the dry solids in the black liquor;

     2.  Volatile compounds, such as terpenes, CH3OH, C02 ;

     3.  Soap and waxes; and

     4.  Water.

The organic part of the black liquor dry solids can  be  classified  as lignin derivatives and
carbohydrates. Lignin contains more carbon and less oxygen than the carbohydrates. The
hydrogen content is almost the same for both. The  compositions of softwood lignin and
hardwood lignin are slightly different. The heat  values and oxygen demand for combustion
differ for the different compounds (1) (2).
                                        10-1

-------
Part of the  volatile compounds, the soaps and  the  waxes may be present in black liquor
when it reaches the recovery boiler for combustion. Part of these components may also have
been stripped or skimmed from the black liquor.

The chemicals in the cooking liquor are present in the black liquor as the inorganic fraction.
A  portion of the inorganic fraction is involved  in reactions with  the  organic material
dissolved in the black liquor  during the cooking and another portion of the inorganics is
involved in reactions after the cooking. Only the  sulfur present as sulfide ion (S~2), that can
be converted to thiosulfate ion (S203~2) in the brown stock washing or in an  oxidation
plant, is discussed here.

The pulp yield calculated after the cooking on a bone dry pulp basis can vary considerably
for different processes, grades of pulp, and species of  wood.

The majority  of recovery  boilers are used  for burning kraft liquor. Some are used for
bisulfite liquor  or for cross-recovery for kraft and neutral sulfite pulping liquor.  The
characteristics of bisulfite and NSSC liquors can  deviate considerably from the normal kraft
liquors.

The dry solids content of  the black liquor from the kraft pulping process can vary over a
wide range. Some extreme values are shown in Table 10-1.
                                    TABLE 10-1
                    BLACK LIQUOR  DRY SOLIDS CONTENT

                  Process                      Dry Solids Content*
                                             kg/t              (Ib/ton)

              Linerboard pulp               1,100              (2,200)
              Paper-grade pulp               1,000              (2,000)
              Dissolving pulp                2,250              (4,500)

              *Based on air-dried ton of production.
Changes in  the reduction ratio (ratio of sulfide sulfur to total sulfur) and the causticizing
efficiency (molar ratio of NaOH to NaOH plus Na2C03) in the white liquor must be
monitored since these changes influence the dry solids composition.
                                         10-2

-------
The heat value of the black liquor dry solids varies with the composition and is higher for
black liquors of high lignin content. The amount of air needed for complete combustion is
also higher  for liquors of high lignin  content. The heat value and the air  required for
complete  combustion  are  related  linearly.  The  deviations,  however, from this  linear
relationship are, in many cases, of the same magnitude as the amount of excess air that can
be used for  combustion in a recovery boiler. The deviations from the linear relationship are
increased  by variations in  the reduction, causticizing efficiency, and residual alkali  in the
black liquor. Figure 10-1 shows the bomb heat values  (BHV)  of softwood and hardwood-
lignins and an average pulp carbohydrate composition as functions of the oxygen required
for complete  combustion. Figure 10-2 shows the BHV for some  North American black
liquors and the maximum observed deviations from the indicated linear relationship.

The amount of air needed for the combustion and the amount of flue gas can easily be
determined  if the elemental analysis of the black liquor dry solids and the water content in
the black liquor are known. The elemental analysis should represent the same stage  in the
process as the figure for the dry solids fired. One method of calculation of air and flue gas
flows is shown in Appendix 10-1.

The BHV  should be  determined  for  the black   liquor without  drying it  in advance.
Investigations show that  considerable changes take place in the heat  value of the black
liquor dry solids if it is dried to a powder before it is put in the bomb (3) (4). A black  liquor
sample of 60-70 percent solids is placed in the bomb and a small amount of paraffin oil is
added on top  to increase the heat  generation, to get complete combustion without residual
carbon in the ash, and to prevent the dry solids  ash from splattering out of the sample
holder. The air should  not be purged from the bomb as the very small amounts of nitric
oxides formed during combustion  are useful as a catalyst for oxidizing all sulfur to sulfate.
The heat value determined in this way has a much smaller variance than the heat value
determined  according to the method of drying the  black liquor to a powder. By relating the
BHV  to the black liquor, the error introduced by the method of determining the  Black
liquor to dry  solids is bypassed. Comparisons of heat values before and after black  liquor
oxidation  should be made using sodium as a reference element because of the change  in dry
solids  concentration.   Comparisons  of heat  values  before and after a direct  contact
evaporator can be made by adding another reference element to the black liquor as sodium
is packed up with ash in the flue gas.

The combustion products in the bomb are in the fully oxidized stage if made with the
modified procedure. Two cases can be considered depending upon the  molar ratio S/Na2.
This ratio  is  normally less  than  0.4 for -kraft liquor,  but could be more  than  1 for
sodium bisulfite liquor.  The combustion products are shown in Table 10-2.

The influence of the  formation  of carbonic acid (H2CO3) and the uneven distribution of
H2S04 in different condensate drops may be neglected.
                                         10-3

-------
CD
UJ
CD
   I2,000_
   II,OOO_
   IQOOQ.
   9,000 _
   8,000 _
   6,000_
   5,000
                                         SOFTWOOD LIGNIN-
             STRAIGHT LINE THROUGH  ORIGIN
                                 HARDWOOD LIGNIN
                CARBOHYDRATES
          30"
401
50'
60"
                      OXYGEN FOR  COMPLETE COMBUSTION
                            IO"3 Ibmoles/lb

                           FIGURE 10-1

 HEAT VALUES VS. OXYGEN DEMAND  FOR COMPLETE COMBUSTION OF
                  LIGNIN AND CARBOHYDRATES
                               10-4

-------
     8000 -
  m
     7000 -
  UJ
  X
  CD
  S
  O
  CO
     6000 -<
     5000
                                 Straight Line  Through  Origin
          25                   30                  35                  40
                     OXYGEN DEMAND FOR COMPLETE  COMBUSTION
                                  IO"3 Ibmoles/lb

                                FIGURE  10-2
 HEAT VALUES VS.  OXYGEN DEMAND FOR COMPLETE COMBUSTION  FOR
                   SOME  NORTH AMERICAN DRY  SOLIDS
The combustion products from the components of the dry solids in the boiler differ from
those of the bomb. For the different S/Na2 ratios, the combustion products are shown in
Table 10-3.

The compounds  given in parentheses in Table 10-3 should be kept  low in quantity for
acceptable operating conditions. The flue gas might, during incomplete  combustion, also
contain H2S, CO, H2CH4, and CH3SH.
                                    10-5

-------
                                  TABLE  10-2
            BLACK LIQUOR COMBUSTION PRODUCTS  IN CALORIC
                     BOMB USING MODIFIED TECHNIQUE

                               	S/Na2	
             Phase               Less than 1.0            Greater than 1.0

         As ash                 Na2S04,Na2C03             Na2S04
         As gas                       C02                    C02
         Ascondensate            H2O, H2C03             H20,H2S04
                                  TABLE 10-3
     BLACK LIQUOR COMBUSTION PRODUCTS IN RECOVERY FURNACE

                                                  S/Na2
          Phase                   Less than 1.0               Greater than 1.0

As smelt                         Na2S,Na2C03               Na2S, Na2C03
                          (Na2SO4, Na2S203, Na2Sn)    (Na2SO4, Na2S2O3, Na2Sn)
Asgas                          C02,(S02),H20             C02,S02,H20
As dust entrained in the gas        Na2S04, Na2C03            Na2SO4, Na2CO3
As condensate                         nil                         nil
 Chlorides in the dry solids will form sodium chloride in the smelt and the dust, as well as
 hydrogen chloride (HC1) and free chlorine (C12) in the gas.

 Corrections have to be made for the actual combustion products in the recovery boiler and
 applied to the calculations for the furnace design and steam generation.

 The amount of flue gas and the release of heat can vary considerably, as calculated per ton
 of pulp. Furthermore, the ratio between flue gas  and heat release does not have a constant
 value, but has to be considered during design of the  recovery  boiler to produce the correct
 temperature in the furnace and to produce the correct temperature of the superheated
 steam generated in the boiler, These facts complicate the design and operation of recovery
 boilers. The American method of stating the recovery boiler capacity in pounds of dry solids
 per  day is less inexact than the  old European method of rating in tons of pulp per day.

 Turpentine and soap are formed during the cooking, especially in the pulping of softwood.
 The turpentine is released in the vapor form but the soap is dissolved in the liquor. Some of
 the soap is skimmed from the black liquor and used for the production of tall oil, which is a
                                      10-6

-------
valuable byproduct. The remaining soap in the black liquor adversely affects the operation
of the evaporation  plant. The soap has a high heat value and will cause complications in the
furnace if it is not efficiently removed.

     10.1.2  Comparison of American and Scandinavian Liquor Concentration Methods.

The concentration  of the black liquor from  the  brown stock washing  depends on the
equipment in the washing department. The black liquor concentration is increased with the
number of theoretical exchange units in the  browrn stock washing and with higher losses of
pulp, if such  can be allowed. The black liquor concentration normally ranges between 15
and 18 percent dry  solids. It is not possible to burn black liquor at this low concentration. It
must be concentrated to at least 55 percent, and normally 60-65 percent dry solids, before
injection into the recovery boiler. The dry solids concentration refers to the solids present
when the black liquor leaves the evaporation unit and does not include the chemicals from
the ash hoppers or dust collectors. The hopper and precipitator ash mixed with the black
liquor increases the  black liquor concentration by 3-4 percent.

The flow of the  inorganic and  organic compounds within a North American black liquor
recovery boiler is shown in Figure 10-3.  This figure assumes the use of a dry ash conveyor
system for the boiler lube  bank, the economizer, and the electrostatic precipitator.  The
normal  practice in  North America has been, however, to use the black liquor in the ash
hoppers and the electrostatic precipitator to convey the ash back to the furnace. Figure 10-3
illustrates the change in the ratio between inorganic and organic contents of the black liquor
within the recovery  boiler department. It does not, however,  show the changes in the
composition of the black liquor that take place in the direct contact evaporator.

It is possible to use a fluidized bed incinerator  for the combustion of black liquor, in which
case a lower  concentration, down to 35 percent, can be  injected into the reactor.  The
chemicals are, however, not recovered in a suitable form for further processing, and the heat
recovery as steam is much decreased. These facts make the use of fluidized bed  apparatus
for  the  recovery  of kraft black liquor uneconomical except for very small plants. Figure
10-4 shows, in principle, the flow of the inorganic and  organic contents through such a
system.

The concentration of the black liquor before injection into  the recovery boiler furnace can
be effected in several ways. Most older recovery boilers in  the United States and Canada use
a multiple-effect evaporation plant to concentrate the brown stock wash liquor to about 50
percent. Direct contact evaporators are then used for concentrating up to about 65 percent
solids before injection into the boilers. Three different types of direct contact evaporators
have been used. These are cascade evaporators, cyclone evaporators, and venturi scrubbers.
The flue gases are normally cooled from 400° C (750° F) to an exit temperature of 110 to
150° C (230 to 300° F).
                                         10-7

-------
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                                                        STEAM TO MILL
SECONDARY
   DUST
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                                                                  FLUE
                                                                  GAS
                                    WASTE HEAT
                                      BOILER
                PRIMARY DUST
                 COLLECTOR
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                     FIGURE 10-4

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      REACTOR WITH WASTE HEAT  RECOVERY BOILER

-------
The  use  of  multiple-effect evaporation alone to achieve 60-65 percent concentration has
been practiced  at all pulp mills in Sweden and about half of the pulp mills in Finland. The
concentration of the black liquor at furnace injection has  been  increased from about 50
percent in 1940 to 60-65 percent today. Figure 10-5 illustrates the  flow of the inorganic and
organic contents through such a system.

The  two different approaches  to the evaporation of  the  black liquor  both  have their
advantages.  The,  Scandinavian method of concentrating  the  black  liquor  exclusively by
multiple-effect  evaporators for injection into the boiler will give  greater steam generation.
The American method of employing recovery boilers with direct contact evaporators are less
expensive to install.  They are  also  self-compensating for capacity  at  overloading  of the
recovery  boilers. The exit temperature of flue gas  at  the  direct contact evaporator will
increase  at  increasing load. This increase  means that more heat  is available for the final
concentration of the  black liquor. The additional heat is normally sufficient to compensate
for the attendant lower dry solids concentration from the multiple-effect evaporator, which
also is  a result of overloading.

Overloading a Scandinavian-type recovery  boiler normally means that the black liquor from
the evaporation  plant will have  a lower solids concentration; however, the overall heat
transfer  coefficients will increase at lower concentrations, and the influence of  the boiling
point  rise on heat requirements will, therefore, be reduced. This reduction will limit the
decrease, in  concentration to  some extent. The steam-generation  rates of  the Scandinavian
system and of the American system would be approximately the same  if the dry solids
concentration after evaporation are the same. Furthermore, the flue gas flow in the furnace
would be higher with the Scandinavian system since the final concentration of the black
liquor  is made  in the furnace. This fact tends to increase the vertical velocity of the flue gas
and  to increase  carryover of dry particles from the black liquor  and  their  subsequent
combustion in  suspension. The emission of H2S in the flue gases is lower than  with direct
contact evaporators operating with no oxidation of the black liquor, even under  the adverse
conditions  of overloaded  boilers. There may also be a certain decomposition  of  the dry
solids  in the direct  contact  evaporators, which will decrease the steam generation  rate.
Investigations,  however, of the decomposition reactions taking place in the direct contact
evaporators are not conclusive. Decreases in heat value of  the dry solids by as much as 6
percent have been reported (5).

The lower capital cost for the American system and the possibility of increasing  the load  on
the  recovery boilers while  maintaining  combustion  conditions as  allowed by the self-
compensating characteristics  of the  direct contact evaporators were very attractive to the
Scandinavian pulp engineers. Investigations were made on the economics of the American
and Scandinavian recovery boiler designs and operation. The feasibility  of the Scandinavian
design was  investigated, and the additional investment was, in most  cases, found attractive
because  of  higher energy efficiencies. The fuel and power  prices in Scandinavia, however,
                                         10-10

-------
                                                        STEAM TO MILL
                                                 ^COMPENSATION FOR INCREASED
                                                 EVAPORATION
MIZER
ELECTROSTATIC
PRECIPITATOR
                BOILER
                 BANK
                                            > \\v\\V\\V\-V
                          \X\\\\\\\\\\\\\\\\\\
                                                                         FLUE GAS
                                                                      BLACK
                                                            INORGANICS ( LIQUOR
                         FIGURE 10-5

FLOW DIAGRAM FOR BLACK LIQUOR THROUGH RECOVERY BOILER
          SCANDINAVIAN SYSTEM-LOW ODOR SYSTEM

-------
were considerably higher than in most parts of the United States and Canada. An updated
comparison of the conditions on the marginal investments is given in Tables 10-4, 10-5, and
10-6.

The flue gas from an American recovery boiler with a direct contact evaporator has a higher
water vapor content than the  flue gas  from a  Scandinavian recovery boiler with a large
economizer. The higher  moisture condition is very advantageous for the operation of the
electrostatic precipitators traditionally used for collection of dust from flue gas. A higher
voltage can be used  with higher water vapor content, and the handling characteristics of the
dust are better at a  lower temperature and a higher water content. (The favorable tendency
of cooling has a practical  limit at about  115° C  (240° F), and the dust absorbs so much

                                   TABLE 10-4
   OPERATING  CONDITIONS  FOR  AMERICAN AND SCANDINAVIAN  BLACK
                           LIQUOR  CONCENTRATION*
  Assumed Operating Conditions

Dry solids, kg/hr (Ib/hr)
Fuel heat value, cents/million Btu
Steam heat value, cents/million Btu
Cost of electric power, cents

Heat consumption by back pressure
  power generation, Btu/kWh
Black liquor concentration at:
  Washers, %
  Evaporators, %
  Direct evaporators, %
  Furnace
Heat consumption in s-effect
  evaporation, Btu/lb
Exit flue gas temperature, °C (°F)
Flue gas C02 content afterboiler, %
Marginal investment costs for:
  Back pressure turbine, $/kW                     -                       60
  Power boiler (steam generating
     section alteration included),
     $/lb steam                                   6                        6

*0il and power prices and investment costs based on data available in Sweden, September, 1973.
                                             American
                                             Practice

                                         45,000 (100,000)
                                                85
                                                95
                                           120/kW month
                                             +0.6/kWh

                                               4,000

                                                17
                                                50
                                                63
                                                63

                                                235
                                             150 (300)
                                                18
  Scandinavian
    Practice
45,000 (100,000)
       85
       95
 120/kW month
    +0.6/kWh

     4,000

       17
       63

       63

      235
    150 (300)
       18
                                       10-12

-------
                                  TABLE 10-5
   CAPITAL COST COMPARISON OF AMERICAN AND  SCANDINAVIAN  BLACK
                     LIQUOR CONCENTRATION PRACTICES
           Item
American Practice
Scandinavian Practice
 Direct contact evaporator
 Larger evaporator plant
 Larger recovery boiler
 Larger back pressure turbine
 Larger power boiler

 Total
 Difference in capital cost
    200,000
    150,000
    350,000
                               600,000
                               600,000
                                30,000
    1,230,000
                  880,000
                                 TABLE  10-6
        ANNUAL OPERATING COST COMPARISON OF AMERICAN  AND
        SCANDINAVIAN BLACK  LIQUOR  CONCENTRATION PRACTICES

                                                Conversion from North American
                                                    to Scandinavian Practice
              Item
            Increases
         Decreases
Cost of evaporation
Evaporator maintenance
Recovery boiler steam value
Recovery boiler maintenance
Power generation value
Turbine maintenance
Maintenance of power boiler

Total
Total decrease in annual operating cost

Gross margin:  ||||~ X 100% = 29%
             73,300
              6,000

              8,000

                600
          324,900

           15,800

            3,000
             87,900                343,700
                        255,800
                                    10-13

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humidity  from  the  gas  thai handling the  dust becomes  extremely difficult.)  Higher
migration velocities could, therefore, be observed in  the precipitators operating after the
direct contact evaporator than those after  a  recovery boiler using a large economizer for
cooling   due  gas  and  using the  black liquor  which  was  fully  concentrated  in  the
multiple-effect  evaporators.   The  characteristics  of  different  types  of  direct  contact
evaporators are discussed in section 10.3, Boiler Design, and in more detail in section 10.6,
Direct Contact Evaporation.

10.2  Combustion of Black Liquor Dry Solids

     10.2.1   Arrangement of Combustion

The two main objectives in operating a recovery boiler are  to recover the  chemicals in the
reduced state  (that is, the sulfur should be present as sulfide and not sulfate) and to  recover
the heat to generate  steam for the process. The value of the chemicals has normally been
much higher than the value  of the heat in the dry solids. It has, therefore, been standard
practice when fuel has been inexpensive to  reduce the cost of the very expensive recovery
boiler installation  by  accepting a low heat recovery and producing the necessary steam in a
separate power boiler.

The combustion within a recovery  boiler has  to be separated into two zones because of the
two different objectives of  the  recovery boiler operation. The first  zone has to be
maintained  under reducing conditions less than the  stoiehiometric  amount of air.  The
products of this zone are a discharge of the  chemicals in the molten state with the sulfur
present mainly as  sulfide and a discharge of the organic matter as a gas having considerable
heat value.  The second zone of the combustion starts with the  addition of secondary air.
The amount  of secondary air corresponds to the  amount of additional air  theoretically
needed for complete combustion of the gas, plus an excess of about 10 to 20 percent.

Figure 10-6 is a schematic drawing of two recovery boiler furnaces. The left hand figure is
the Babcock & Wilcox (B&W) design. The nozzles for the primary air (item 4 in the figure)
and the  secondary air (item  5 in  the figure) are shown. The right-hand  figure shows the
primary  and  secondary  air  supply,  according  to  another manufacturer,  Combustion
Engineering (CE). In the Babcock &  Wilcox  (B&W)  design, some of  the  secondary air is
introduced  at a  higher level (item 6 in the figure) and is normally referred to as tertiary air.

The black  liquor  is  sprayed in rather small drops over the cross section  of the  furnace
through  the flue gases. The intention is to dry the black liquor droplets to a concentration
where the  heat value of the char material with the residual moisture is sufficient to keep a
reasonably  stable  combustion going. The  liquor  spray nozzles can be placed as item  3  in
Figure  10-6 indicates. The  black  liquor dry solids  are collected in  the bottom of the
                                         10-14

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  BABCOCK a WILCOX
          .Steam
                          12
COMBUSTION ENGINEERING
         fSteam
         i    """
      r-rt
 LEGEND
 I.  Furnace
2.  Smelt Spouts
3.  Black Liquor  Spray Nozzles
4.  Primary Air Supply
5.  Secondary  Air Supply
6.  Tertiary Air  Supply
7.  Position of  Char Bed Burners  for Oil  or Gas
8.  Normal  Configuration of  Char  Bed
8*.  Same  at  Low  Primary  Air Flow and  Pressure
9.  Screen Tubes
10.  Superheater
II.  Boiler  Tube  Bank
12.  Exit to Economizer

                                  FIGURE 10-6
              I
                      SECTION A-A
                 PRINCIPLE DESIGN OF AIR DISTRIBUTION TO
                        RECOVERY BOILER  FURNACES
                                      10-15

-------
recovery  boiler furnace in a char bed. The air burn* off the dry solids from the top of the
char bed, which normally has the form indicated in the left-hand part of Figure 10-6.

The primary air is supplied around the circumference of the recovery7 boiler furnace. The air
supply should be  distributed with  a fairly constant ratio  to the local supply of black liquor
dry solids. The primary  air is normally preheated to  a  temperature  of 150° C (300° F).
Some boilers with higher primary air temperatures have been made. The primary air jets will
penetrate a certain distance  into the furnace depending on  the size  of the  jets and their
velocity.  The oxygen in the air will be consumed in the combustion during the penetration
of the jet into the furnace. The partial pressure of the oxygen in the  border  zone between
the flue gas atmosphere and the char bed will, therefore, vary over the wrhole extension of
the air jet. The temperature will also vary. The oxygen partial pressure and the temperature
are the two main parameters that determine the combustion conditions in a limited spot.
The combustion  conditions and the equilibrium  conditions  will, therefore,  vary over the
entire cross section of the furnace. One Scandinavian manufacturer (Gotaverken) introduces
some  of the primary air at a higher level and with a wider spacing between the nozzles and
at a considerably  higher pressure than for the conventional primary air supply. This is called
high primary air  and the arrangement is shown  in the left-hand part of Figure 10-7. The
conventional  design is shown in the right-hand part of the figure. The char bed will form as
indicated  in the  figure. Other Scandinavian  manufacturers are using  secondary  air for the
same purpose as high primary air (6).

The flue gases formed during primary and secondary air combustion cool by radiation to the
furnace walls to about 870° C (1600° F) before they enter the convection heating surfaces
(i.e.,  the  screen  tubes, the  superheater, and the  lube  boiler bank).  The flue gas  has a
temperature of about 450° C (850° F) after the  tube  boiler  bank.  The flue  gases  are then
further cooled in  an economizer to about 400° C  (750° F) in the American system  with the
direct  contact evaporator and to about  160° C (320°  F) in the Scandinavian system. Two
typical American recovery  boilers of modern design with economizers instead of direct
contact evaporators are shown in Figures 10-8 and 10-9  for B&W and for CE, respectively.

The reason for having a recovery boiler furnace large enough to allow the flue gases to cool
to about 870° C  (1600° F) before they enter the convection surfaces is to allow complete
combustion of entrained  organic  particles before they can reach the coder  tube surfaces.
The deposits from  organic  matter have been shown to be much more difficult to remove
from the heating  surfaces than normal tube deposits.

Burners for oil or natural gas to supply heat to the lower part of the furnace are arranged as
indicated in Figure 10-6, item 7. These burners are used to:
                                        10-16

-------
    HIGH SECONDARY AIR
    LOW SECONDARY AIR
    HIGH PRIMARY AIR
    LOW PRIMARY AIR
                                       CLC    C   F.  [   C   f.   C  [
                                    NEW SYSTEM  —  OLD SYSTEM
                                 FIGURE 10-7
    AIR  SUPPLY ACCORDING TO GOTAVERKEN ANGTEKNIK AB  DESIGN

     1.   Bring the recovery boiler up to operating temperature and to supply heat for
         drying the injected black liquor drops,

     2.   Supply additional heat to increase the temperature in the furnace to stabilize the
         combustion during disturbances,

     3.   Smelt down the bed at shutdowns  to allow inspection of the lower part of the
         walls and the bottom of the furnace, and

     4.   Supply heat for additional process steam generation when the black-liquor supply
         is not sufficient.

The first three uses of the burner are necessary for the operation of  the boiler. The fourth
use is only justified when operating the recovery boiler so that enough heat is supplied for
process steam generation. It coincides in this case with the second use mentioned. Steam in
excess of the normal rating should be generated by  load carrying oil or gas burners placed at
                                      10-17

-------
                                                       Furnace
                                                       Slag Screen
                                                       Tertiary Air Ports
                                                       & Windbox
                                                       Black Liquor
                                                      /Oscillator Burner
Oscillator Burnerv
                                                       Secondary Air Ports
                                                     /& Windbox
   Pin Stud
Upper Limit
       II
                Smelt Spouts
                     &Hood
                                             Primary Air Ports
                                             & Windbox
                             Dissolvings j Dissolving TankT,]   L
                               T,«L, '-van.    A „;*-,*„„  u_TTi   r
                                    FIGURE  10-8

       MODERN  KRAFT  RECOVERY UNIT FROM  BABCOCK & WILCOX



a higher level in the recovery boiler. This position will also tend to keep the temperature of

the superheated  steam  at  the correct value since the ratio  of flue gas flow to  steam

generation is considerably lower for oil and gas than for black liquor.


The operation of  auxiliary  fuels  in  the  recovery boiler  presents an  explosion hazard.

Detailed instructions about the installation and operation of such burners have been made

by  the Black Liquor Recovery Boiler Advisory Committee (BLRBAC) (7) (8). Instructions

about emergency shutdown procedures  are also given by BLRBAC.
                                         10-18

-------

                  	SECTIONAL SIDf ELEVATION
                         FIGURE 10-9
MODERN KRAFT RECOVERY UNIT FROM  COMBUSTION ENGINEERING
                            10-19

-------
     10.2.2   Reactions in the Primary Air Zone

A standard for operating the recovery boiler to obtain the proper reactions from the process
was developed by observing  the effects  of disturbances in the combustion process and of
changes in operation parameters.

A typical example illustrating this  state of knowledge is as follows: The combustion in the
hearth could, at local spots, have an insufficient rate of combustion and a low temperature,
resulting in a condition sometimes called "black out." The remedy was to use a compressed
air lance to blow the black char from the wall and reignite the local spot. Heavy corrosion
was often observed  at places where black out conditions were common. Air lancing was
believed  to be the reason for this corrosion. Compressed air lances were also used al places
with normal  combustion without any resulting corrosion. It was then deduced that the
black out conditions themselves were responsible for the corrosion, but the mechanism was
not understood. Many of the problems that arise in attempting to achieve stable combustion
without  forming local  black  out  conditions are  caused  by the  necessity  of reaching a
compromise between two competing design requirements in providing sufficient  reaction
surface between fuel and primary air. The furnace cross section  should  have a relatively
small size with the present arrangement to allow suitably  high temperatures in the primary
combustion  zone. At the  same time the cross section  of the furnace  has  to exceed a
minimum size to avoid high vertical gas velocity. A  high vertical velocity would tend to
entrain burning char pieces and black liquor drops. This carryover would result in buildup of
deposits on the heating surfaces in the boiler following the  furnace,  and a higher dust load in
the gas.

The very complicated combustion conditions for  sodium-based pulping liquors have been
investigated  by  Bauer and Borland (9)  by means of thermodynamic  calculations of the
equilibrium conditions  at different temperature levels for a  certain composition of black
liquor. Bauer  and Dorland used the negative logarithm of the partial  pressure of oxygen
(rO = -logj o 02) as the independent variable and negative logarithms of the partial pressures
of the possible compounds (-logi 0 P)  in the flue gas as the dependent variable in determining
equilibrium  conditions. This  method  was introduced  by Sillen and Andersson (10) for
calcium bisulfite and magnesium bisulfite  liquor.  The possibility  of complete emission of
sodium or sulfur or  both to the flue gas  under certain furnace conditions and black liquor
compositions  exists (see Figure 10-10). The  upper portion of  Figure  10-10 shows an
equilibrium  diagram for condensed phases and  partial pressures  of gases for  727° C
(1,340° F), and the lower portion shows the same diagram  for 1,127° C (2,061° F). The
kinetics of the possible reactions that can  occur were not taken into account in preparing
these diagrams.  Rosen  has  extended  the black  liquor equilibrium  studies to  include
variations in the black liquor water  content and in the pressure (11).
                                        10-20

-------
                                                                     1000°K
   32   30   28   26   24   22   20    18   16   14    12
                       24  22   20   18    16    14   12    10    8   64
  1400°K
                    Upper Diagram - NaOH and Na2Q cannot exist at
                                  this or higher temperatures.

                    Lower Diagram - The partial pressure of metallic
                                  sodium (Na and Na2) is becoming
                                  more and more  important as the
                                  temperature increases, and will
                                  lead to high fly ash losses.
                               FIGURE  10-10

EQUILIBRIUM  DIAGRAM FOR CONDENSED PHASES AND  GASES  FOR A
                  SODIUM  BASED  BLACK  LIQUOR (9)
                                   10-21

-------
The previously cited work and other observations indicate that the release of sodium from
the burning char  of  black  liquor  is mainly  a  function  of the temperature and  the  gas
conditions in  the border  zone between  the  bed  and  the flue  gas. The total sulfur,
concentration and the sodium/sulfur ratio control the sulfur release to the flue gas. Figure
10-11 shows the condition^ for A.normal_kra.ft liquor. Sodium is used to represent the base
in the process. Small amounts of potassium are normally present in the wood and also in  the
makeup chemicals. Some potassium will, therefore, be present in the black liquor, but it will
react in the same way as sodium. It has a lower boiler point and its compounds have a lower
melting point.  The influence  of the potassium can normally be neglected  except for mass
balances.

The  amount  of sodium which is  distributed to the flue  gas increases sharply with  the
temperature,  following a curve similar to a vapor pressure curve.  The release of sodium is
also influenced by the primary air velocity, which determines the diffusion conditions. The
amount oTsodium in th"e flue gas will increase considerably with increasing air velocity, and
this relationship j5S£jois_iojndicate_ that the-release of sodium could be by evaporation. The
rate of evaporation would depend on the  diffusion conditions in the border zone between
the flue gas atmosphere and the bed. The sodium  to sulfur ratio in the snuvlt_is_strgngly
dependent on the temperature. The release of  sulfur to  the flue gas will, therefore, be a
function of the temperature,  the  Na2/S  ratio  in the black liquor, and,  possibly, of  the
absolute sulfur content in the black liquor.

Large variations in the Na2/S  ratio in the smelt have been observed during firing of bisulfite
liquor  of  constant Na2/S ratio. Variations in the smelt sulfidity have been observed to a
lesser extent  when firing kraft black liquor. The curves for sodium and sulfur release shown
in Figure  10-11 are made for normal contents of sodium and sulfur, that is, 18 percent and
3.5 percent, respectively. At very low temperatures, 700° C (1,300° F), all the sulfur in the
black liquor  is released to the flue gas. This actually takes place at the shock pyrolysis as
shown in the  research for the Billerud-SCA process. Release  of  sulfur  to the flue  gas
atmosphere decreases at increasing temperatures, but eventually increases  again until total
release  occurs above 1,540° C (2,800° F). There  should, consequently, be  combustion
conditions with a  sufficiently  high temperature and velocity to evaporate enough sodium to
combine with all the sulfur released to the flue  gas, provided that the molar S/Na2 ratio is
less than 1.0 in the dry solids.

The  sulfur may be present in the flue gas over the bed as S, H2 S  or SO2.  A  low temperature
favors the presence of S and H2S. Elemental sulfur can also be formed by the combustion of
H2S with 02 depending on the temperature and  molar ratio. Elemental sulfur in the nascent
state seems to be  responsible for much of the observed corrosion on recovery  boiler tubes
(12) (13). Observations and analyses made in the last few  years show a good correlation
between the factors  that influence the temperature  in the furnace  and  the release of H2S
andSO2 (14).
                                         10-22

-------
    UJ
       oO
    °=u.
    UJ O
    gco
    b <
    UJ
    en
                 S02  IN  FLUE GAS
                                                   at  high S-content
  I
  I
  I
•••k
                                                '   NA2C03 IN  'DUST
                                ABSOLUTE  TEMERATURE

                         S: Normal sulfur content in dry solids  (3.5%)
               	   S-' High sulfur content in dry  solids  (5.0%)
                                 FIGURE  10-11
DISTRIBUTION OF SODIUM AND SULFUR IN A BLACK  LIQUOR  RECOVERY
               FURNACE AS  A  FUNCTION  OF  TEMPERATURE
H2S is probably present in the flue gas in the primary air combustion zone at about 50 to
200 ppm  if the temperatures  in  the  border  zone between the bed and the flue gas
atmosphere are optimal for the operation. H2S concentrations up to 15,000 ppm have been
observed in this region when black out conditions were present. Small amounts of CH3SH
have been  found in the primary air combustion zone, but no measurable amounts (<1 ppm)
have been  found at the entrance to the screen tubes  when there was sufficient secondary air
supply.

The sodium probably  evaporates from the bed as elemental sodium and reacts with oxygen
to form Na2O  within a very short distance from the bed. The Na20 reacts with CO2 to
form Na2CO3  at the high temperature in the furnace. Part of  the sodium compounds
sublime to dust from the vapor phase. The black liquor droplets and agglomerates will pass
through the primary air combustion zone. The volatile compounds will be partially stripped
from the black liquor  during the drying. Decomposition of  the black liquor dry solids and
pyrolysis and possibly combustion of the smallest fraction of the drops will start before the
drops  have reached  the bed. The small  amounts of mercaptans, organic sulfides, and
                                     10-23

-------
hydrocarbons  that are found  may be  stripped during the  drying  or the beginning of
pyrolysis and are not necessarily products of the combustion itself.

Black liquor drops too small to fall downward can be carried upward by the flue gas and
burn in the secondary  air combustion zones. The temperature in the reaction zone can then
reach much higher values than at normal combustion, and  a higher than normal release of
sulfur can take place.
                                                         X
The smelt is collected  at the bottom of the bed and discharged through the smelt spouts to
the dissolving tank. The melting point depends on the smelt composition (see Figure 10-12).
The melting point increases considerably with contamination of calcium and decreases writh
the presence of potassium and chlorides (Figure 10-13). The  melting point and  the heat
transmission rate determine the thickness of the smelt layer on the tubes.

     10.2.3   Secondary Air Combustion

The final combustion starts immediately after the introduction of secondary air. The total
amount of primary and secondary air must for most  boilers be more than 110 percent of the
theoretical air (that is, stoichiometric  air). It should, on the other hand, be less than 125
percent to avoid the possible formation of sticky dust, which  has  a great tendency to foul
the heating surfaces in  the economizer and/or the  collecting plates in the electrostatic
precipitator. The two air limits correspond to 2 and 5 percent excess 02 in  the flue gas for
an average kraft liquor.

By definition, the secondary air is the  difference between the total air and the primary air.
Soni(^ investigations indicate that the primary air flow should be between 60  and 70 percent
of the theoretical  air. The remaining air to be used as secondary air would be as a minimum
40 percent, and, as a maximum, 65 percent of the theoretical air. The secondary air should be
supplied in the  furnace so that  it mixes with   the gas coming from the primary air
combustion zone below. The primary air combustion zone gas may have wide variations in
its demand  for oxygen to complete the combustion, depending on local variations in the
bed. It is fortunate that it has been possible to achieve practically complete combustion in
the recovery boilers without using very large amounts of excess air.

The secondary air is  supplied according  to two different methods  by the main North
American manufacturers of recovery boilers, as shown in Figure 10-6. The (low) secondary
air is, in the B&W boilers, placed  only  a  few feet above  the  primary air  nozzles, and it
controls, in many cases, the height of  the bed in the center of the furnace. This air supply
has, therefore,  a mixed function, and acts  along the walls as secondary air to complete the
combustion of the gas, but in the center of the furnace as primary air to burn off the bed.
The sum of the  primary air and  (low) secondary air is  normally not sufficient to give
complete combustion. The tertiary air  in B&W boilers is supplied above the spray nozzles to
                                         10-24

-------
                                                           -2I56°F
           o

           LU
           o:
           LU
           CL
           LU
  1600
I564°F
                                                              I380°F
                          /                              >
                   1300
              Na2C03 % 100

                      COMPOSITION  OF SMELT  % BY WEIGHT

                                  FIGURE 10-12
          EQUILIBRIUM DIAGRAM FOR A  Na2CO3-Na2S  SYSTEM  (15)

give a reasonable amount of excess air to complete the combustion. The nozzles for the air
cannot be controlled, and the velocity of the air is, therefore, proportional to the flow and
to the absolute  temperature  of  the  air. This means that changes in the load or in the
distribution will influence the velocity of the different air jets and the resulting turbulence
which is presumed necessary for complete combustion.

The design of CE boilers according to Figure 10-6 uses a tangential air supply to produce a
rotary movement of the gas in the furnace. Only four big nozzles are used, one in each wall.
Each nozzle is divided into compartments which can be shut off individually to control the
velocity in the air jet when the flow is changed. The tangential air supply was previously
placed high in the furnace and the rotary movement tended to  load one side of the super-
heater and boiler tube bank more than the opposite side. The air nozzles have recently been
moved downwards in the furnace. This change should increase  the  temperature in the
                                       10-25

-------
              u.
              o
                    °F
                   1500-


                   1400-
                   ,300-
              CD
              ?   1200-
              _
              UJ
                    1100-
    SMELT
COMPOSITION
    Na2S
    NcuCO
50%
25%
25%
                                        50
          100
                              WEIGHT % NaCI  ADDED
                              TO SYNTHEIC  SMELT
                                  FIGURE 10-13
                EFFECT OF NaCI  ADDITION ON THE  MELTING
                        POINT  OF  A  SYNTHETIC PULP
primary air combustion zone slightly, by radiation, and decrease the sideways influence on
the furnace temperature distribution.

The design of a typical Scandinavian recovery boiler, such as shown in Figure 10-7, has the
secondary air supply split  between two levels. The low secondary air is placed above the
primary air nozzles but below the spray nozzles,  and on all four walls of the furnace. Most
of the secondary is normally supplied at this level. The rest of the secondary air is supplied
in the upper level of boilers as high secondary air or tertiary air. The high secondary air is
supplied on two or four boiler walls.

All air nozzles are adjustable in the modern Scandinavian design. This feature allows control
of the  air velocity with total independence of the flow and allows distribution of the air
between different levels. It also allows adjustment of the air distribution in response to the
liquor spray pattern. This has proved very  valuable, especially in cases where the boiler load
was low during the start-up period of mill operation.

     10.2.4   Formation of Particulate Matter

The flue gas from a black  liquor recovery boiler  contains large amounts of dust. The dust
load varies between 40 and 75 kg per metric ton of dry solids (80 and 150 Ib/ton). The dust
                                       10-26

-------
is formed by the release of sodium from the bed to the flue gas over it. The amount of
sodium released does not seem to  depend on the sodium content in the black liquor dry
solids.

The sodium and sodium salts in the black liquor evaporate at a rate dependent upon their
partial vapor pressures and the diffusion conditions. A few feet above the bed, solid salts are
present as Na2C03 even at very high partial pressures of S02 • The Na2C03 reacts later with
S02 to form Na2S03, which is then oxidized to Na2S04 by the excess oxygen in the flue
gas.
                                                           •f
The excess oxygen and the content of S02 in the flue gas shpw a corresponding decrease in
this temperature field according to complete analyses using gas chromatography on samples
taken before the screen tubes, before the superheater, after the superheater, and after the
boiler tube bank (6).

The dust will  contain only Na2SO4 after the boiler if the content of S02  is in excess of
what is necessary for the stoichiometric conversion of Na2CO3 to Na2S04, and S02 will be
present in the flue gas at the  exit of the economizer. In the  opposite case (too little S02),
Na2CO3 will still be present in the dust after the boiler, and the S02 content will be zero or
very low at the exit.

The partial pressures of H2SO4  and S03  in the flue gas are  a function of  the reaction
conditions. The formation of S03 and H2 S04 can be increased considerably if fuel oil with
a high content of vanadium pentoxide (V2 05) is fired in the char bed oil burners or in load
carrying oil burners. The S03 and H2 S04 are probably absorbed on the dust particles, and the
dust gets sticky. Dust containing Na2C03  can also form sticky dust through adsorption of
S03.
                                                                 /
Rather  high concentrations of H2S04 and S03  were found when firing of sodium sulfite
liquor. The SO2  content in the flue gas can be as high as 0.5 percent, that is, more than 10
times that of kraft boilers. Corrosion has occurred in the tube bank on such boilers, but not
on kraft boilers. Some instances of corrosion in the last part of the economizers for the low
odor type of recovery boiler might have been caused by SO3 and H2S04>-

Dust sampled  from 400° C (750° F) down to 150° C  (300°  F)  has been analyzed,  and
differences in the crystalline structure were found. These differences may explain some of
the differences in bulk weight and handling characteristics of dust  in  the  electrostatic
precipitators at different temperatures, even when no sticky dust was observed.

Very small  amounts of Na2S  (0.2 percent) have been found in both large and small dust
particles. Large particles of the range 5-15 jitm (2.0-5.9 X 10"4 in) could be disintegrated ash
from the combustion of very  small drops or agglomerated sublimation products. Smaller
                                       10-27

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particles, down  to  1 jitm  (3.9 X 10 5 in), were analyzed,  and the smaller  ones  of  these
probably could only have  formed by sublimation. The smaller particles had about the same
Na2S content, which  was probably a product  of reaction  between Na2C03 and H2S or
elemental sulfur.

The  dust containing Na2S has no odor at normal atmospheric conditions. H2S, however, is
generated when the dust is exposed to CO2 and water or water vapor. This fact can explain
the occurrence of kraft odor at some distance from the mill when there is almost no odor at
the mill site.

NaCl may be present in the dust if the black liquor contains chlorides. The HC1 and chloride
in the flue  gas  will probably react with Na2S04  in the  lower temperature range. It is
possible to collect a substantial amount of HC1 by scrubbing the flue gas with water. Sodium
chloride condenses at  lower temperature than Na2S04, and this dust is, therefore, likely to
have a smaller diameter than the Na2S04 particles in the dust.

Iron can  be found in  the dust deposits  on  the  tubes at  what would  be an  alarming
concentration if it were caused by corrosion in the  boiler. Iron compounds exist, however,
with a  relatively high partial pressure at  the temperature  of the primary and secondary
combustion  zone and their evaporation is possible, with condensation taking place on  the
cooler surfaces in the boiler and economizer.

10.3  Different Recovery Boiler Designs

The present design  of recovery boilers was developed in North America essentially by B&W
and  CE. Manufacturing overseas  was  by allied  companies or  licensed boiler  manufacturers
and  to some extent by independent boiler manufacturers. The main differences between  the
two American boiler types are in the  air supply, the spraying of the black  liquor, the smelt
discharge, and the design of the  direct contact evaporator for the final concentration of  the
black liquor.

Another approach was often  chosen in Scandinavia, with no  direct contact evaporators for
recovery boilers. The black liquor  was evaporated  to the final concentration in multiple-
effect evaporators, and large economizers were used for cooling the flue gas before  the
electrostatic precipitators. The reason  for this approach was the high cost of fuel and power.

Changes in  the American design were necessary to produce odor free operation. The high
ratio of the value  of the recovered chemicals to recovered heat,  and the very low price of
fuel in  North America made it economically feasible to run  the recovery boilers with
incomplete  combustion. The deficit in steam generation was made up by  oil- or natural
gas-fired boilers  at a comparatively low capital cost. The incomplete combustion sometimes
                                        10-28

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caused very high emissions of H2S. The direct contact evaporators were another source of
H2S, which was formed by the reaction:

                       2NaHS + C02 + H20 -» Na2C03 + 2H2S

The latter source could, however, be eliminated by oxidation of the black liquor, that is, by
converting the Na2S to Na2S203  by air or  molecular oxygen in contact with the black
liquor ahead of the multiple-effect evaporation. This practice, however, would not eliminate
the generation  of H2S in the boiler  itself.  It  would instead have a  slight effect  in the
opposite direction by  decreasing the heat value  of the black liquor. This would lower the
temperature in the primary air combustion zone and move the chemical equilibria toward
forming more H2S and S, and less S02, from the bed.

The two North American manufacturers proceeded in  different ways  to achieve odor-free
operation in the  cases where complete oxidation was  not used. B&W adopted the design
with a large economizer instead of the direct contact  evaporator. CE has also tried other
approaches  to  the problem. The  Scandinavian  manufacturers have generally  kept their
design, with small refinements.

     10.3.1   Babcock & Wilcox (B&W) Recovery Boilers

B&W recovery boilers  used  two different types of direct contact evaporators prior to the
low odor era. These were cyclone evaporators and venturi evaporator-scrubbers.

A B&W recovery  boiler with a cyclone evaporator is shown in Figure 10-14. The primary air
is supplied through air nozzles placed around the circumference of the boiler at an almost
constant height over  the  slightly inclined bottom of the furnace. The air  nozzles are
arranged in groups of normally 4-5 nozzles at the same height. The air flow pressure in each
group can be controlled by a damper. The smelt  is discharged at the front wall through
watercooled smelt spouts.

Secondary air is supplied at the distance  of about 2 m (6 ft) above the primary air nozzles
(measured at the center) by a smaller number of nozzles. Secondary air is supplied on all
four walls. Spray nozzles for the black liquor are placed in the front and rear walls above the
secondary air level.

The nozzles can be tilted up and  down, and they  also have a sideways swinging motion.
Tertiary air is supplied above the  spray  nozzles for the black liquor. The air to the three
windboxes is supplied  by one or two forced draft fans and steam coil air heaters with the
pressure adjusted  to the demand for the secondary or tertiary air. The air is throttled to the
primary air windboxes to  adjust the pressure  to  the correct level.  Measurements of the air
flows to the different windboxes are made after the steam coil air heater.
                                        10-29

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                         SOOT
                        BLOWERS
                          GREEN-LIQUOR
                          RECIRCULATINS
                             PUMPS
BLACK-LIQUOR
  PUMPS
  CYCLONE
 EVAPORATOR
RECIRCULATING
   PUMPS
                                   FIGURE 10-14
                        BABCOCK &  WILCOX  RECOVERY
                     BOILER  WITH CYCLONE EVAPORATOR
The furnace is arranged with a rather large arch in the rear wall to distribute the gas in the
superheater.  Screen tubes  are  used ahead  of the  superheater to  adjust  the  flue gas
temperature to give the correct temperature of the superheated steam. The flue gas then
passes the boiler tube bank and the economizer before exiting to the cyclone evaporator.

The  cyclone  evaporator is  shown  in  Figure  10-15.  The flue gas enters the  cyclone
tangentially near the bottom.  Black liquor is sprayed across the gas inlet. The liquor drops
are separated from  the gas on its helical path to the outlet on the top of the cyclone. Black
liquor is recirculated to nozzles at the top for wetting the walls. This wetting flushes the
black liquor drops  and dust,  which  have separated from the gas, to the sump  tank in the
bottom  of  the  separator.   Some control  of  the  exit  gas  temperature  or the  final
concentration of the black liquor can be exercised by adjusting the black liquor flow to the
spray in the gas inlet. The operation resembles a low pressure drop venturi scrubber. The
liquor is heated in  a direct steam heater before it is sprayed into the furnace. The heater
steam condensate dilutes the black liquor to a lower concentration than it had when leaving
the direct contact evaporator.

The dilution of the black liquor is disadvantageous for heat economy, and it represents an
explosion hazard if  the black liquor injection into the furnace is interrupted temporarily and
                                        10-30

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                          MECHANICAL
                            POWER
                            STRAIN
                               NER.
                                                       WALL-WETTING
                                                         NOZZLES
                          UOUOR TRANSFER
                           TO SALT CAKE
                             MIX TANK
                                                         CYCLONE
                                                        EVAPORATOR
                           FLUE
                          GASES
                                   FIGURE  10-15
                             CYCLONE  EVAPORATOR


the steam control to the direct heater should malfunction. Reheating of the black liquor
using indirect heat exchangers has been done in Finland.

B&W introduced a boiler design with a large economizer to meet the demand for low odor
generation (see Figure 10-8). The main difference from the previous design is that the height
of the  furnace was increased  resulting  in a proportional increase in the retention period for
the combustion with tertiary air. Also, a large economizer was supplied for cooling of the
flue  gas  to  about 200° C  (400° F) before the  exit to the electrostatic precipitator.  The
economizer  consists of  long vertical tubes with  a number of baffles arranged  to  give
substantial crossflow for the flue  gas. This arrangement increases the gas  velocity  and
achieves  better heat transmission than parallel flow. The crossflow arrangement seems to
create some pockets with low gas velocity and to create some problems with the ash disposal
with sootblowing, which is done with conventional retractable sootblowers.

     10.3.2   Combustion Engineering (CE) Recovery Boilers

The CE recovery boilers were equipped with cascade evaporators before the requirement for
reduced odor levels. A typical design is shown in Figure 10-16. The primary air was supplied
                                        10-31

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                                                                - SECTION '* -A' -
                                  FIGURE 10-16
            COMBUSTION  ENGINEERING RECOVERY BOILER  WITH
                            CASCADE EVAPORATOR
at all four walls with primary air nozzles about 1 m (3 ft) above the bottom of the furnace.
All primary nozzles were placed in a common windbox. The smelt was discharged through
smelt spouts about 0.3 m (1 ft) above the bottom decanting hearth, which is horizontal.

Black liquor spray nozzles  were  placed about 6 m (20 ft) above the primary air nozzles on
all four walls, with several  nozzles per wall. The nozzles could be  tilted up and down, but
not swung sideways.

Except for the very first such boilers built, the supply of secondary air has been introduced
at a distance of about 2.4  m (8  ft) above the level of the liquor sprays. The secondary air
nozzles were divided  into sections which could  be shut off individually to allow for adjust-
ment to the secondary air  flow.  Oil burners were, in some cases, placed in the secondary air
windboxes for generation of steam in excess of that generated from black liquor combustion.
                                       10-32

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The furnaces were built originally without a nose at the rear.wall under the superheater. A
nose was  added around  1960 to distribute the flue gases through the superheater. The
superheater  was made of panels with a side spacing of 30cm (12 in) to improve  the
operating conditions of the superheater.

The gas passed the screen tubes, the superheater, the boiler tube bank, and the economizer
to one or two cascade evaporators of the double rotor type. A cascade evaporator is shown
in Figure 10-17. The rotors are constructed of tubes between the end disks and carry black
liquor up  through the  flue gas pass.  The gas velocity and the temperature determine the
diffusion conditions and the evaporation rate to  the flue  gas.  The time of exposure to the
gas is considerably longer on the outer part of the  rotors than on the inner part and may
cause overdrying  of black liquor at the outside  of the rotors. This can be decreased by
increasing  the speed of the rotors. The residence time in the cascade is rather long (about
four  hours)  compared  to  the  cyclone  evaporator.  The   variations in black liquor
concentration caused by changes in the flue gas  conditions should be damped by the large
volume  in the cascade.  The  concentration  is normally controlled  by  bypassing  the
economizer with more or less flue gas through a damper or by diluting the black liquor with
weak black liquor.

The demand for low emissions was met by CE with the introduction of the air cascade (see
Figure  10-18). The direct contact with the flue gas was eliminated by using the combustion
                                  FIGURE  10-17
                    OPEN  VIEW OF CASCADE EVAPORATOR
                                       10-33

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 LEGEND
 —*   ROUTING  OF  AIR
 .Tj   ROUTING  OF  FLUE GASES
 I.  LAMINAIRE  AIR HEATER
    (COOLING FLUE GAS)
 2.  AIR CASCADE EVAPORATOR
 3.  FORCED DRAUGHT  FAN
 4.  SECONDARY  AIR
 5.  PRIMARY AIR
                                FIGURE 10-18
            COMBUSTION  ENGINEERING RECOVERY  BOILER WITH
                    CASCADE EVAPORATOR ACE SYSTEM
air for the direct contact evaporation. This system is called ACE. The air was heated to
400-425° C (750-800° F) by rotating air heaters of the Ljungstrom type. This system had
the disadvantage that the air to the furnace carried all the evaporated water vapor from the
black  liquor and so increased the humidity ratio by about 0.1 kg H20/kg dry  air (0.1 Ib
H2 O/lb  dry air). The increase in the partial pressure of the primary air  increases the
endothermic reaction between CO and H2O to form CO2 and hydrogen (H2) at the contact
with  the  bed.  The temperature in the primary air combustion zone would, consequently,
decrease, and tend to increase the emission of sulfur to the gas and increase the ratio of H2S
to S02. The lower temperature also would tend to decrease the reduction of the smelt.

A more recent development is shown in Figure 10-19. Recovery boilers of this type use large
rotary air heaters to cool  the flue gas to a  temperature suitable for the  electrostatic
precipitators. The combustion air is heated to 315° C (600° F) to increase the temperature
in the furnace, which increases the release of sodium, reduces the emission of sulfur from
the bed, and decreases the ratio of H2S to SO2. This system is called LAH. The forced draft
fans are placed after the heaters and are common for both primary and secondary  air. This is
disadvantageous in increasing power consumption and in decreasing the accuracy of air
measurement.
                                     10-34

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  LEGEND
I.
2.
3.
4.
        ROUTING OF AIR
        ROUTING OF FLUE GASES
       LAMINAIRE  AIR  HEATER
       INDUCED  DRAUGHT  FAN
       SECONDARY AIR
       PRIMARY AIR
                                 FIGURE 10-19
   COMBUSTION  ENGINEERING RECOVERY BOILER  WITH LAMINAIRE AIR
   HEATER &  COMPLETE  MULTIPLE  EFFECT  EVAPORATION  LA.H. SYSTEM
The final evaporation stage is made in a special type of concentrator following the multiple
effect evaporator to reach 65 percent dry solids before the hopper and filter ash are added.
The influence on the emissions at the high air temperature was very favorable. The theory
predicts a greater evaporation of sodium and a decrease in the H2S/S02 ratio above the bed.
The results of the operation seem  to verify this prediction; however, difficulties were
encountered with the  cleaning  of the air heater because of narrow spacing between the
regenerator elements.  Frequent  water  washing  was  required.  The air heaters were
dimensioned for  operation with 80  percent boiler load on one air heater,  while the other
heater was  out for washing. The water washing may prove to be too troublesome for the
operation and may cause  some extra corrosion problems similar to those experienced in
Scandinavian water washing of gilled tube economizers used before 1960.

CE now has designed recovery boilers (see Figure 10-9) that include large vertical steel tube
economizers. The economizers are built in two or three passes with an open space between
two consecutive passes. The flue gas passes downward parallel to the tubes and upward in
the empty space between the tubes.  The  pressure drop is relatively low even with high gas
velocities, and the steam used for sootblowing can be kept at a minimum. Finned tubes are
often used with this design to enlarge the heat transmission surfaces of the tubes. This type
of economizer has previously been used in Scandinavia with very good results.
                                      10-35

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     10.3.3   Scandinavian Recovery Boilers

The design of recovery boilers in Sweden and Finland beginning with the first injection-type
boiler included an economizer followed by an air heater for the cooling of the flue gases to
130° C (260° F). Such cooling was achievable when the heating surfaces were "technically
clean." This design was chosen because of high fuel prices. The large dust load of the gas
meant  that the  economizer and  air  heater  often had to be water  washed, even with
continuous shot  cleaning. Washing periods between  5 and  21 days were standard. The
differences in the cleaning cycles were  probably due to acid dust caused by high excess air;
however, the influence of acid dust was not known at the time. The water washing damaged
the economizers and  they  had, on the  average, to  be  rebuilt every 10-15 years. The
recirculation air heater, Figure 10-20, was used to decrease the corrosion of the air healers
in the "cold corners." The introduction of the hot precipitator in 1957 (placed between the
boiler tube bank and the economizer) solved the problem of keeping the economizers clean
without water washing, but  it was more difficult to keep the precipitator  in operation for
long periods without shutdowns for cleaning. The long steel tube economizers were chosen
after 1967, when the precipitators had to be designed for high collecting efficiency for
particulate emissions and  trouble free operation.

A typical Scandinavian recovery boiler is  shown in Figure 10-21.  It is manufactured by
Gotaverken Angteknik AB (G.V.), Sweden, a licensee of B&W, England. Other  manufac-
turers of  recovery boilers in Scandinavia  include Svenska  Maskinwerken AB, Sweden, Oy
Tampella AB, Finland, and A. Ahlstrom Osakeyhti5, Finland.

A comparison  between  the North American  (Figures 10-14  through  10-19) and  the
Scandinavian standards (Figures 10-20  and 10-21) are shown  in the preceding illustrations.
The outstanding features  in the Scandinavian design are as follows:

     1.    The primary and secondary air flows are handled by separate fans and air heaters,
          A, B.

     2.    The suction  ducts to the fans are conveyed from the top of the building to give
          straight ducts,  C,  D, for  an  accurate measurement of the gas flow. This also is
          advantageous for ventilation  since all recovery boilers in Sweden and Finland are
          built indoors because of the cold climate.

     3.    The  primary  air  is  split between low primary  air, E,  in  the conventional
          windboxes  and high  primary air, F, in windboxes at a higher position and  at a
          higher pressure through a booster fan (Gotaverken).

     4.    The  secondary air is split between low secondary air, H,  on all  four walls  below
          the sprayer level and high secondary air, I, (Tertiary air)  above  the sprayer level.
                                        10-36

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o
CO
                 ELECTROSTATIC
                 PRECIPITATOR
               Arrangement
               tator
                                   'SCRUBBER
                                 I.D. FAN
                                ECONOMIZER
                              AIR PREHEATERS
of' hot  precipi-
                                                                       rELECTROSTATIC
                                                                       I  PRECIPITATOR
                                     IR PREHEATERS A D FAN

                              '-RECOVERY BOILER
Arrangement  of "warm"  pre-
cipitator
                                                                                                         7500Fi FLLE GAS
                                                                                                                r	,. Feed Wat«t to
                                                                                                                380° Boiler  Drum
                                                                                                 FLUE GAS^750°F
                                                                                                                220°F
                                                                                                                    ECONOMIZER
                                                                                                          420°{F
                                                                                                                340°F
                                                                                                        300°

I«S°F
0°iF i
                                                           iFead Water,
                                                             Inlet   [--

                                                           ECONOMIZER
                                                                                         AIR PRE-
                                                                                         HEATERS
                                                                                                  220°
                                                                                                                                     295°F
                                                                                                                                     n
                                                                                                                                   1 250°F
                                                                                                                                      ll 300°F
                                                                                                          85°F
                                               PRIMARY-SECONDARY
                                                     AIR
                    PRIMARY-SECONDARY
                         AIR
Arrangement  of  heating  surfaces
in economizer-air  preheaters.  Left,
old system-,  right,  new system
(after 1956)
                                                                 FIGURE  10-20
                              RECIRCULATION AIR  HEATER  FOR  SCANDINAVIAN  RECOVERY BOILER

-------
       PRIMARY/
        AIR  V
c
( )
l) FD FANS A
A,
AIR
HEATERS
B,
D AIR DISTRIBUTION
f FLOW METERS
T\ SECONDARY
/) AIR
A2
p rc
B2
. r
* L
BOOSTER FAN
/C\ . r
^G ' ^


HIGH
i. SECONDARY
AIR
LOW
t. SECONDARY
AIR
F
>i HIGH PRIMARY
AIR g-
LOW PRIMARY
                                   AIR
                                                0
                   FIGURE 10-21

TYPICAL GOTAVERKEN ANGTEKNIK RECOVERY BOILER
                      10-38

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     5.   The cooling of the flue gases after the boiler tube tank is made with a long vertical
          steel tube economizer in several passes.

     6.   The  gas  flow is  from the top of the tube baffle down  to  the bottom, which
          facilitates the ash transport at sootblowing (same as CE).

     7.   The flue  gas is normally cooled to 160° C (320° F). This has been found to be the
          economic temperature for steam generation, the cost of the economizer and cost
          of the precipitator, and the power consumption in the induced draft fan.

     8.   The  ash  handling in the hoppers of the recovery boiler  tube bank  and  the
          economizer banks,  and  in  the precipitator,  is of  the  dry type.  Drag chain
          conveyors of  a special design (Redler) are used. Rotary valves are used for sealing
          of the gas  passage  at the discharge of  the  ash.  The ash from  the hoppers is
          conveyed to the mixing tank with vertical dust chutes.

     9.   To avoid leakage of water vapor up into  the dust chutes,  which causes clogging,
          small screw conveyors are inserted before the mixing tank and equipped with air
          jet seals.

     10.3.4  Rebuilding of Old Recovery Boilers to Low Odor Design

An existing recovery boiler with a direct contact evaporator can be  rebuilt to the low odor
design by installing a feed  water economizer,  possibly combined with an air heater. This
addition will  eliminate  the emission  of  odorous  compounds  from  the  direct contact
evaporator, but not from the furnace when the furnace is overloaded.

The air supply system can be  revised  to achieve better control of the air supply and thus
increase the capacity for complete combustion of black liquor dry  solids.

The evaporation plant has to be equipped with a concentrator to increase the  concentration
of the black liquor to the  recovery boiler department to  62-65 percent. The electrostatic
precipitator capacity must be increased to correspond to the reduced  emission of particulate
matter allowed by regulations.  The higher black liquor concentration will tend to decrease
the flue gas flow, but the higher  flue gas temperature increases the velocity of flow.  The
result  is  often an increase  of about  10 percent in   required   precipitator capacity.
Furthermore, the migration velocity of the dust is lower primarily because of the lower
humidity  of the flue  gas. The  dust collecting .efficiency  has to be  upgraded to meet the
regulations.  A  new full  size precipitator or at least an additional  precipitator has to be
included in the rebuilding program. The additional dust collecting capacity may be achieved
with a scrubber if there is a bleaching plant which requires hot water  and if the stack plume
caused by the lower exit gas temperature is acceptable.
                                        10-39

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These additions can be erected with the old equipment in operation, and they can be ducted
to the existing small  economizer  and the ID fan  in a few days, if local conditions are
favorable. The  costs for lost  production caused by downtime  will then  be limited  to
acceptable figures.  This approach  is probably  the  least expensive  way to get low odor
operation (16).

The  usage of  an existing recovery  boiler for bark-  and oil-firing after suitable  changes is
possible  if  a  new  modern recovery boiler  is built in  combination with an increase  in
production. The losses are then restricted to the special equipment for black  liquor firing
and the green liquor system. The life expectancy for a boiler converted to oil  and/or bark
firing is, in many cases, greater than if the boiler were kept on-line as a recovery  boiler.

10.4   Process Variables

     10.4.1   Objectives of the Recovery Boiler Process

The main purposes of the recovery boiler operation in the ordinary kraft process are:

     1.   Recovery of chemicals in the black liquor for dissolving into green liquor with all
          sulfur in the reduced state,

     2.   Generation of steam for the process in the  mill,

     3.   Generation of steam with high pressure and high temperature to allow  generation
          of power for the mill to the maximal extent,

     4.   Control of the  combination to avoid emission of malodorous gases, S02, and
          particulate matter,

     5.   Allowing the use of the most inexpensive makeup chemicals available, and

     6.   Reliable operation with minimal capital, operating, and maintenance costs.

Exceptions  to  these general objectives are possible under certain circumstances. An example
is cross-recovery between  bisulfite or NSSC  and kraft. The flue gases can, in  this case, be
scrubbed for recovery of S02  for production of the cooking acid for the bisulfite or NSSC
process.  The recovery process should then be operated to give sufficient S02  in the flue gas.
This operation will decrease the sulfidity of the green liquor (and the white liquor from the
recausticizing), and normally this decrease is favorable for a mill with small sulfur losses.

Some  of the objectives given above are contradictory. For example, the generation of steam
with a high pressure and temperature for the generation of power increases low  capital and
                                         10-40

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 maintenance costs. The capital cost increases with pressure arid temperature and so do the
 dangers  of corrosion and  increased maintenance costs.  Another contradictory pair are
 avoidance  of  emission of S02 and low operational costs. Sulfur dioxide emissions can be
 avoided  by evaporation  of  sufficient amounts of sodium  to combine with the SO2 in the
 gas, but this increased evaporation will increase the dust load in the gas and possibly increase
 the  demand  for  sootblowing steam.  It  will  also  increase  the size of the electrostatic
 precipitators needed.

 Clearly then,  the design  of the recovery boiler  must be the result of a balance between the
 different objectives and conditions, not only in the recovery boiler but also in the different
 departments which  are  influenced  by  or which influence the operation  of the recovery
 boiler. Some  of the operation and design  parameters 'for  normal kraft  recovery boiler
 application are discussed in the following section.

     10.4.2   Firing Rate for Dry Solids

 Combustion of a  given amount of black liquor solids produces a certain heat input in the
 recovery furnace. The combustion requires sufficient air for  complete combustion and will
 produce  an amount  of flue gas that depends on the  DS composition  and amount of excess
 air. The  temperature in the furnace will depend on four factors, the  heat input from the dry
 solids, the preheated combustion air, the  furnace dimensions and the operating pressure.
 The first depends  on the production, on the type of wood species used for the mill, on the
 pulping yield, on the efficiency of the  brown stock  washing, and on  possible oxidation of
 the black  liquor.  The second factor in the heat input  can be varied from ambient air
 temperature up to 150° C (300° F) by  heating the air with steam of 0.45 MPa and 1.1  MPa
 (50 and  150 psig)  pressure. The air can be heated further by high pressure steam from the
 boiler  to 230-260° C (450-500° F), depending on  the pressure. This will give  a slightly
 higher temperature in the furnace for improved operation or to compensate for a decrease in
 load. Heating  the combustion air from the flue  gas with indirect air  heaters  will increase'the
 cost of  the  air ducts and  would have an  adverse  effect at load changes  (i.e., the air
 temperature  would   decrease  when  the  black  liquor  heat input  decreased).  The air
 temperature will, on the  other hand, increase with increased load; in which case decreasing
 its temperature, in some cases even to the extent of using ambient-temperature air, can offer
 a favorable solution.

The  furnace   dimensions must be  chosen with  some regard to the  flue  gas flow from
combustion of the dry solids. The amount of dry flue gas  from the organic content in the
dry solids  can vary considerably with different species and probably also with the rate of
growth and the age of the wood and chip storage time. Changes in the inorganic content and
the causticizing efficiency and reduction will cause  only relatively small corrections.  The
furnace dimensions, therefore, depend not  only  on the total  amount of dry solids but also
                                        10-41

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on the composition of the black liquor. These factors should be taken into consideration if
carryover of sprayed liquor droplets are to be avoided.

Sufficient retention time in the furnace and excess air are needed to  achieve complete
combustion before  the flue gas enters the cooling tube surfaces. On the other hand, the
oxygen content should not be too high to avoid the possible formation of sticky dust.

These  factors  and the conditions for the chemical reactions mean that it is impossible to
operate a recovery boiler  over a wide  firing range.  Incomplete combustion and further
generation  of  H2S  will result if the boiler is overloaded, and a low reduction of the smelt
and a high emission of SO2 will result at a low load with low temperature in the furnace.
The latter  case can be avoided to some extent by increasing the heat input by firing an
auxiliary fuel such as oil or natural gas in the char bed burners.

The average firing  rate at which complete combustion can be achieved without frequent
peaks  above 1 ppm H2S (with the black liquor  concentrated to at least 62 percent dry
solids) is, according to present observations, as follows:

    Dry solids load per unit of furnace cross section, standard liquor conditions

     1.  American black liquor, 14,700 kg/m2/day (3,000 Ib/sq ft/day)

     2.  Scandinavian black liquor, 13,200 kg/m2 /day (2,700 Ib/sq ft/day)

This  difference is  of great importance when the results from  combustion in different
countries are compared based upon economy and emissions of odor and particulate matter.
There are,  however,  a  number  of cases where higher loads have been  obtained with
satisfactory combustion  results. These results might  depend on  a greater attention  to the
operation of the boiler than can be expected as an average.

A great number of black liquor analyses have been investigated, both from North America
and Scandinavia. The oxygen demand and the flue gas from  the  dry solids (the vapor from
the black liquor has been excluded) at theoretical complete combustion without excess air
varies considerably  more for North American liquors than for Scandinavian:

      Variation in Oxygen Demand and Flue Gas for Kraft Black Liquor Dry Solids

      1.   North American, ±19 percent

      2.   Scandinavia, ±8 percent
                                         1042

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Normally there is a rather consistent linear relationship between the oxygen demand and the
BHV  for all fuels and organic compounds. The BHV is, however, determined with complete
combustion to the oxidized state for  all the elements and with the water vapor in the
condensed  state.  The  heat that  is released in the recovery boiler is less  than the BHV.
Corrections have to be  made  for the  reduction  of  the sulfur  compounds, the heat of
evaporation for the water in the black liquor,  the water formed by the combustion of the
hydrogen in the  dry solids, and the heat of fusion for the smelt. It is, therefore, unlikely
that  there   should  be a  consistent relation between the heat  input  steam and the  air
consumption, or between flue gas flow and the amount of black liquor dry solids fired.

The possibilities of reducing emissions  by varying  the  dry  solids firing rate in a recovery
boiler are limited. The rated capacity for firing is determined  after  observation of the heat
value  and characteristics of the black liquor.  The  primary air flow for the rated capacity
must  be  determined to give the correct  temperature in the recovery boiler furnace in order
to achieve the right proportions of sodium and S02  in the flue  gas. This procedure probably
leaves a  rather small range of firing rates at the maximum load within which  complete
combustion without release of H2S can be achieved. This range is about 5 to 10 percent and
limits the  possibility  of variations of  the operation on the  upper side. Temporary load
changes of greater magnitude must be covered by tank storage for black liquor and green
liquor.

The range of load changes  is also limited on the lower side. The relative amount of primary
air  and/or the temperature  of the primary air has to be increased at a lower firing rate of dry
solids to avoid emission of S02. Such corrective measures might allow operation down to
80% of the  normal firing rate. The temperature in the primary air combustion zone will be
too low at  still lower loads without additional supply of heat  by firing of oil or natural gas
in the char bed burners. High emissions of H2S (and elemental sulfur) can also take place in
the primary air zone  at low temperatures. Results  from test runs indicate that if the H2S
release in the primary air zone is considerably above 200 ppm, even locally,  it will be very
difficult to achieve an H2S concentration below 1 ppm in the exit from the boiler. It should
be  possible to get down  to 65% of the  normal  dry  solids  firing rate  with acceptable
combustion and emissions with a  substantial firing of additional  fuel like oil or gas in the
char bed  oil burners to increase the temperature in the primary  air  zone.

Operation at reduced load of black liquor dry solids with additional  fuel requires an even
distribution of the additional heat input. Concentration of the  heat supply from one or two
burners can distort the configuration of the bed and cause carryover of organic material and
combustion in  suspension. Means to  control the air  to fuel ratio  for the  additional fuel
should be provided.

The possibilities  of operating  recovery boilers at low load  have not been investigated
thoroughly. The high investment cost for  the recovery boiler makes low load conditions
                                        10-43

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rather unusual. The minimum load which can  be recommended without causing unstable
conditions in the primary air combustion zone, and without additional heat input from oil
or natural gas, would be  80-100% of the maximum  load with respect to S02 emissions and
65-100% with respect to H2S emissions.

     10.4.3   Black Liquor Characteristics

The  most important characteristics  of  the  black liquor  in  the  design and  operation of
recovery boilers are:

     1.   Heat value and oxygen demand for complete combustion,

     2.   Viscosity and surface tension,

     3.   Organic/inorganic  ratio and S/Na2 ratio, and

     4.   Boiling point rise.

The  heat value  and  oxygen or air demand  for  complete  combustion  are  of utmost
importance for both the proper  design and proper operation of a recovery boiler. Attempts
to achieve a usable formula  for the heat value as a function of the elemental analysis have
met  with only limited success. The best solution seems to be to conduct laboratory cooks to
the desired degree of delignification and bomb tests on the test liquor.

The  variability in the black  liquor composition with different wood species  is related to the
relative proportion of  lignin and carbohydrates, and the type of carbohydrates, present in
the  different  species (17, 18).  Appendix 10-2 shows the composition and  the BHV for
softwood lignin, hardwood lignin, and carbohydrates, and the oxygen demand required for
complete combustion.  The  composition and heat values have then been approximated for
dry solids with a normal content of chemicals.

The, amount of heat that will be released in a  reducing atmosphere, that is, the "efficient
heat value in reducing atmosphere," and the "resulting heat value  of black liquor" (the
total released heat used for steam generation) assuming no losses from the flue gas, have also
been calculated. The values  are shown in Figure 10-22 as a function of the oxygen demand.
Three  lines through the  origin are shown. The  deviation from a linear relation  is shown as
a percent. The BHV show the smallest deviations; however, these deviations are of the same
magnitude as the allowable variations in the excess oxygen. The two lower lines in Figure
 10-22, which show the heat released in a reducing atmosphere and the total heat available
for steam generation, show  a very good linear relationship, but the carbohydrates have great
deviations from the lines through the origin.
                                        10-44

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                                                                See App. 10-2
 A= softwood lignin
 B= hardwood lignin
 C= carbohydrates
B, > Ditto, adjusted  for normal
                 content  of  inorganics
    11,000-
    10,000 H
    9pOO-
    8,000-
GO

U   TJOOO
u   6,000-
    5,000-
                                       Bomb heat value
                                    ,-,  Efficient  heat value  in  reducing
                                    ~r  atmosphere
                                    ^  Resulting  heat value of black liquor
                                     -  Deviation from linear relation
                                       between heat value and oxygen
                                       demand and  vice  versa  in  percent
                                       of  factual value
                                 40
                                         50
                                          3
                       OXYGEN DEMAND, I0  Ibmoles/lb DRY SOLIDS
30
                               FlGUKb

    HEAT VALUE VS. OXYGEN  DEMAND AT  COMPLETE  COMBUSTION
7o
                                    10-45

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Variations in the lignin content in the wood, the amount of extractives and differences in
the cooking process, variations in climatic conditions and the  age  of the trees  before
harvesting, and other variables might explain the great variations in the composition of the
black liquor dry solids from different mills. Consideration of both the variations in the dry
solids content and the BHV when comparing various firing rates seems to be more accurate
than considering dry solids content alone.

The viscosity of black liquor as a function of the temperature is shown in Figure 10-23 (19).
Very great variations were observed between different mills using the same species of wood.
The viscosity is of great importance for the spray pattern. Liquor from eucalyptus pulping
has, as an example, such high viscosity that the black liquor concentration is often  limited
to 60 percent to achieve a reasonably good, stable distribution of spray.

The surface  tension seems to have little influence on the spray pattern. Investigations made
with additives  to decrease the  surface tension in  order  to  improve the spraying showed
without significant changes.

The inorganic/organic ratio affects the heat value per pound of dry solids and complicates
the operator's efforts to keep the load of the recovery boiler constant. The variations in the
inorganic-organic ratio cause  variations in the smelt flow and green liquor production. Such
variations may decrease the  reduction, as a high inorganic  content often is a result of a
temporarily low causticizing efficiency and reduction.

The  molar  ratio S/Na2, along with the  concentration  of sulfur and the temperature,
probably determines the release of sulfur from the bed. The ratio should, therefore, be kept
as nearly constant as possible. In some plants, both the supply  of makeup salt cake and also
the spent acids from the C1O2  generation and the tall oil plant are added continuously at a
constant rate. The ash transport from the ash hoppers in the boiler and the economizer can
be equalized to some extent by adjusting the sequence of the sootblowers to give  a more
steady supply of ash from  different parts. The scraper conveyors used in Scandinavia, as
compared to the common black liquor flushing system used in North America, tend to give
a steady supply of ash.

The boiling point rise is a function of the inorganic/organic ratio and the concentration. The
black liquor is normally sprayed into the furnace with a temperature near the boiling point.
Changes in the  spray pattern will  follow  a  decrease  in  the  boiling point  rise if the
temperature is kept at  a constant value. An increase in the fraction of fine drops, which will
cause carryover, will be the result of spraying at a temperature above the boiling point.

The  boiling  point  rise is  sometimes  used  for  determination  of the black liquor
concentration.  This  method gives  inaccurate results because of the  variations  in the
inorganic/organic ratio.
                                         10-46

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                  1000
              c
                          * Where DS is dissolved solids
                                  FIGURE  10-23
                      VISCOSITY OF BLACK  LIQUORS (19)

     10.4.4   Black Liquor Oxidation

Black liquor oxidation was discussed in Chapter 9. Oxidation of the black liquor with air
converts the sulfide sulfur to thiosulfide. The importance of this oxidation to the recovery
boiler process is that  remaining sulfides react with C02 in the flue gas to  form H2S on
contact. H2S is also produced from sulfides on contact in a wet precipitator and a wet ash
conveying system. This latter release, however, is insignificant compared to that from direct
contact evaporation. The generation of H2S on direct contact with the flue gas is totally
eliminated with  complete  oxidation; however, oxidation tends to increase slightly  H2S
generation in the furnace.

The decrease in  heat value  of  the dry  solids caused by  black liquor oxidation leads to  a
decrease in the temperature in the furnace. The lower temperature decreases the ratio of
H2S and sulfur to S02, but increases the total sulfur release.
                                       10-47

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Investigations of the influence of oxidation are difficult because not only the sulfide sulfur
but also  the organic matter is  oxidized. Some volatile  compounds are stripped from the
liquor, and the addition of oxygen increases the weight  of the dry solids. Dry solids with 4
percent sulfur content increase in weight at complete oxidation by 2.5 percent.

Water is evaporated from the black liquor on contact with the air during oxidation. The heat
consumption for  this evaporation  is  quite  high,  about 3270 kj/kg, (1400 BTU/lb) as
compared to 486 kj/kg (210 BTU/lb) in a conventional 6-effect evaporator. Furthermore,
the black liquor is cooled  in the direction of the wet bulb temperature, and this heat has to
be replaced in  the evaporating plant. The loss in heating valve caused by the oxidation of
Na2S to Na2S2O3  must be replaced in the furnace.

The  loss in heat value per unit mass of original  dry solids almost doubles if the oxidation
efficiency is increased from 90 to near 100 percent (20). The heat losses in steam generation
can be calculated as  116 kj per kg of dry solids (50 BTU/lb) per each percent sulfur for 90
percent oxidation and 209 kj per kg of dry solids (90 BTU/lb) per each percent sulfur for
100 percent oxidation. Allowances are  made for the evaporation in the oxidation towers as
compared to evaporation in a 5-effect evaporator. This heat loss might gain in importance
with the increased sulfidities  resulting from  recovering  the malodorous gases and  with
increased fuel prices.

The oxidation decreases the air consumption by 1-2 percent but the heat input has a greater
influence on the furnace  temperature. The primary air temperature must be increased to
compensate for the  lower heat value and to keep the reaction condition constant in the
primary air zone with use of fully oxidized liquor.

     10.4.5  Air Distribution and Air Temperatures

The  air  distribution,  the air temperatures, and  the  spraying pattern  are  the main
independent variables by  which the operation .of the recovery boiler  can be changed. The
specific load, expressed as dry solids per unit cross section of the furnace, depends on the
dimensions of the recovery boiler and the production of the pulp mill. The heat value of the
dry solids of the black liquor depends on the type of pulp and the species of wood used for
production.

The total amount of air which can be used for the combustion under the present technology
has both lower and  upper limits. The lower limit is set by the condition that at  least 10
percent excess  air has to be used  to avoid unburned matter in the flue gas and formation of
H2S. The upper limit of about 20-25 percent excess air is necessary to avoid the  possible
formation of S03 in the flue gas, and the resultant sticky dust caused by S03 adsorption on
the dust. A certain  amount of the total air must be used for the primary air combustion
zone  to give  the proper  oxygen  concentration in  this zone  and  to give the   correct
                                        10-48

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temperature. With the present recovery boiler designs, the correct°amount for primary air is
approximately 60 to 70 percent of the theoretical amount of air for complete combustion
without  excess  air. The  reduction  of the sulfur compounds in the smelt  is probably
enhanced by a low oxygen concentration and a high  temperature, within certain limits. An
increase  in  the  amount of primary air increases both  the oxygen concentration and  the
temperature. The  two parameters  of  oxygen  and temperature  are evidently  coupled
together, and some combination of them must exist that is the most favorable.

The air is currently heated either by steam from the backpressure turbine at  pressures of
1.13 MPa (150 psig) and 0.45 MPa (50 psig) or with feedwater recycled from the top of the
economizer  through air heaters and fed back to the inlet of the economizer (Figure 10-20).
The most economic choice depends on the prices of fuel and power at  the specific site. The
steam air heaters  give an  increased flow  of high pressure steam through  the  turbine and
increase the back pressure power generation. The recirculation air heater permits  a lower
capital cost for the economizer for the same  flue gas exit temperature. This latter system is a
very clean and easily operated arrangement, but limits the air temperature  to about 150° C
(300° F). Possible ways to increase the air temperature are either to return to using the
indirect type of air heater with cooling of the flue gas by air or to use a rotary Ljungstrom
type air heater. The Ljungstrom type air heater requires rather extensive ducting of the hot
air  from the heaters to the furnace. Direct  heating by firing auxiliary  fuel after the steam
coil or recirculation air heaters is possible. The water  vapor from the additional fuel will
change the equilibrium conditions for the  endothermic water gas reaction, which is only
partly compensated by the increase in the partial pressure of CO2. This method works the
same way as firing oil in the furnace with the char bed burners, which is favorable at low
load conditions.  The primary air velocity is increased,  tending to increase the evaporation of
sodium.

Advantages in raising the primary air temperature are:

     1.    Velocity for the same flow of air increases and therefore improves the  sodium
          evaporation from the bed,

     2.    Release of sulfur from the bed decreases,

     3.    Distribution  of the released sulfur changes in  the direction of more S02 and less
          H2S, and

     4.    Decreasing the primary air flow is possible, and consequently, more air is available
          for the secondary air supply.

Adjusting the velocity and  distribution  of the primary air flow within the furnace is
probably valuable so  that the flow  corresponds to  the black liquor distribution. B&W
                                        10-49

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designed a system with an adjustable  nozzle, which was used around 1960 in the United
States, but it was  abandoned.  This system  has, however, been used in Scandinavia (see
Figure 10-24) as standard equipment.

The difference between the total air flow-and the primary air flow is used as a computation
of secondary air to complete the combustion. Assuming an average figure of 115 percent of
the theoretical air as the total combustion air, and 70 percent used as primary air, only 45
percent is available to complete the combustion. This small amount has  to cover the whole
cross section of the furnace and effects good mixing of the combustible gas with the air. The
suppliers of recovery boilers all  use different  ways of achieving complete combustion. They
all seem to  work provided  that  a reasonably low amount  of H2S is  present in the gas
mixture coming from the primary air combustion zone.

Jones, Brink, and  Thomas suggest that another method of controlling  the  combustion
conditions is to  supply oxygen to the  air (21, 22). This tends to reduce the volume of the
combustion chamber and to allow the temperature to be increased to the desired level. But
                                                       PRIMARY AIR
                                                       PORT REGISTER
                                                         PRESENT  ADJUSTABLE
                                                               DESIGN
                                  FIGURE 10-24
                            ADJUSTABLE AIR PORTS
                                       10-50

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there are some economic drawbacks. The amount of flue gas per unit of steam generated is
less. The superheating of the steam, which is of extreme importance for the back pressure
power generation, is  considerably more expensive and more complicated than with the
present type of superheaters, because the flue gas temperature  at  the  entrance  to the
superheater must be kept low enough to avoid very difficult slagging problems.

The greatest problem with oxygen addition is probably that of achieving a sufficient mixing
of the gas from the primary air combustion zone with secondary air.

10.5   Diverse Obnoxious Compounds

CH3SH,  CH3SCH3, CH3SSCH3, carbonyl sulfide (COS), and carbonyl hydrogen sulfide
(COSH) were  measured in  addition to H2S, S02, and S in the lower part  of the recovery
boiler in the gas phase. They were not found to any measurable extent, however, in the
upper part of the furnace in the neighborhood of the screen tubes where the boilers were
operated with 2 percent excess  oxygen. Some tests indicate that  H2S is present in rather
high concentrations when CO and H2 are also present in high concentrations. A very good
correlation seems to exist between H2S and H2  content. Both H2S and S02 emissions can
be  controlled  to a satisfactory degree by  applying  present technology to  the  design  and
operation of recovery boilers as previously discussed (23).

The obnoxious compounds,  according to available test results, present no problem  in a
reasonably  well operated recovery boiler. CH3SH, CH3SCH3 , and CH3SSCH3 can, however,
be stripped from the black liquor in the direct contact evaporators.

The NOX emission from recovery boilers are probably of minor importance because of the
low  combustion temperatures  that are  reached  even  locally  in  comparison  to   the
temperatures which are reached  in  the flames of  oil and pulverized coal burners.  The
relatively high NOX content which has been found after combustion of ammonium bisulfite
(NH4HS03) liquor, is probably caused by oxidation of the monatomic nitrogen produced
by the cracking and combustion of ammonia.

10.6  Direct Contact Evaporation

Direct contact evaporation of black liquor is performed at  most kraft pulp mills  in  the
United States for concentrating the liquid from  40-50 percent to 60-70 percent solids (ash
from precipitators and ash hoppers not included) to facilitate combustion  in the recovery
furnace. The normally used direct contact evaporators are the cascade, cyclone, and venturi
evaporators. The direct  contact evaporator can  act as an air pollution source,  but also in
some aspects as an air pollution control device. A development by CE was the ACE system
in which the  black liquor is concentrated by the combustion  air in a direct contact
evaporator. The air is preheated by Ljungstrom air heaters to 400-425° C (750-800° F). The
                                       10-51

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water vapor  and any  compounds  released from the black  liquor are carried with  the
combustion air to the furnace. This arrangement is called an indirect contact evaporator.
The  relative  suitability  of direct  contact evaporation,  as  opposed  to multiple-effect
evaporation of black liquor, is based on operation and capital  costs, flexibility with respect
to capacity, and environmental aspects.

     10.6.1   Specific System Characteristics

Direct contact evaporation can be performed in one of three different types of systems. One
of the systems, the high pressure drop venturi scrubber with strong black liquor as the liquid
medium, has  the dual function of water evaporation and particulate removal. The venturi
type evaporators are gradually  being  complemented with additional  dust  collectors or
replaced because of their inability to meet particulate air pollution emission standards. The
different types of direct contact evaporation systems employed are shown in Figures 10-14
through 10-17. The  operating characteristics of recovery boiler flue gases, as indicated by
moisture content, gas temperatures, and particulate loadings for the different types of direct
and indirect contact systems for concentrating strong  black liquor, are presented in Table
10-7.

                                   TABLE  10-7
     RECOVERY  FURNACE EXHAUST GAS  PROPERTIES  FOR DIRECT AND
                 INDIRECT CONTACT EVAPORATION SYSTEMS
                                      Direct Contact Evaporators
          Property

Pressure drop, in. w.g.

Flue gas temperature at:
         economizer exit, °C
                       (°F)
          evaporator exit, °C
Flue gas moisture content at:
       economizer exit, g/m3
                     (gr/cf)
        evaporator exit, g/m3
                     (gr/cf)
       precipitator exit, g/m3
                     (gr/cf)
Cascade
  2-4
Cyclone
  2-4
Venturi
 15-30
                                          Indirect
ACE
 2-4
315-370
(600-700)
150-163
(300-325)
315-370
(600-700)
132-163
(270-325)
315-370
(600-700)
88-110
(190-230)
315-370
(600-700)
—
—
5-9
(2-4)
2-7
(1-3)
0.1-1.1
(0.05-0.5)
7-11
(3-5)
5-9
(2-4)
0.1-1.1
(0.05-0.5)
7-11
(3-5)
0.9-1.0
(0.4-0.8)
—
—
7-11
(3-5)
—
—
0.07-0.2
(0.03-0.10)
                                        10-52

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 The cascade- and cyclone-type direct contact evaporators are low pressure drop gas-liquor
 contact  devices that are used for concentration of strong black liquor by evaporation.
 Electrostatic precipitators, located downstream of the direct contact evaporator,  provide
 particulate emission control and chemical recovery with these low pressure drop systems.
 Cascade  evaporators employ a  rotating cylindrical drum with attached  tubular wheels
 perpendicular to the direction of the gas stream for gas-liquid contact and are normally used
 with  recovery boilers manufactured by CE. Cyclone evaporators are basically low pressure
 drop  cyclonic scrubbers for gas-liquid contact and are normally used with recovery boilers
 manufactured by B&W.

 The direct contact evaporator serves several functions besides the concentration of black
 liquor to 60-70 percent solids. These functions include:

      1.    Reducing  the  inlet gas temperature to the electrostatic precipitator, where the
          lower gas temperature results in a reduced volume  flow rate, which allows a
          smaller precipitator to be constructed at a lower capital cost;

     2.    Reducing  the  inlet particulate loading to the electrostatic precipitator by 20-40
          percent by weight, primarily by scrubbing of large particles emitted from the
          furnace, as reported by Hisey (24);

     3.    Absorbing about 75 percent of the S02 emitted from the recovery  boiler (25,
          26), and nearly all of the sulfur trioxide (27); and

     4.    Absorbing H2S emitted from the recovery boiler (28, 29, 30) under conditions of
          high black liquor pH  and low sodium sulfide concentration in the strong black
          liquor.

     10.6.2  Air Pollution Control

Direct contact  evaporation  has  the  potential  for  liberating substantial  amounts  of
malodorous sulfur gases from the black liquor  or for absorbing sulfurous gases generated
from  the recovery furnace. Considerable amounts of H2S and lesser amounts of CH3SH can
be released from the black liquor by  acidifying Na2S and sodium mercaptide (CH3SNa) by
the action of acidic flue gas constituents, such as C02, S02 , and SO3 . Organic sulfur gases,
such  as  CH3SCH3 and CH3SSCH3, and malodorous  organic nonsulfur compounds can be
evolved from the heating of the black liquor by contact with recovery boiler flue gas (31).

Major variables affecting the potential  for release of malodorous sulfur compounds from
black liquor during  direct  contact evaporation include inlet liquid composition, liquid pH
and alkalinity levels, inlet  liquor and flue  gas temperatures, and the degree of gas-liquid
contact. Recent studies  (28, 32) indicate a substantial increase in reduced sulfur emission
                                        10-53

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levels to  the stack, if sodium sulfide concentration is increased in the strong black liquor
entering the direct contact evaporator. Therefore, a high black liquor oxidation efficiency
must be obtained to reduce the Na2S to between 0.01 and 0.1 g/1 to minimize malodorous
sulfur gas generation during direct contact evaporation. See Table 10-8 (33).
                                   TABLE 10-8
    EFFECT  OF  BLACK LIQUOR  OXIDATION  ON SULFUR  GAS EMISSIONS
                DURING DIRECT  CONTACT  EVAPORATION (33)

Sulfur Gas         	Unoxidized Liquor	        	Oxidized Liquor	
                     kg/t             Ib/ton              kg/t              Ib/ton

   H2S            2.5-15             5.0-30.0           0.05-1.0            0.1-2.0
 CH3SH          0.15-1.0           0.3-2.0            0.025-0.10          0.05-0.20
 (CH3)2S          0.025-0.075       0.05-0.15         0.005-0.025         0.01-0.05
(CH3)2S2         0.05-0.15          0.10-0.30         0.005-0.075         0.01-0.15
These findings have been verified in subsequent studies conducted by Martin (34).

Murray and Rayner (29) have shown that the liquid pH of the incoming  black liquor can
have a considerable impact on H2 S emissions during direct contact evaporation. Increasing
the liquid pH reduces the rate of H2S  formation at any given level of sodium  sulfide and
results in a reduction in the amount of H2S generated in the direct contact evaporator. See
Table 10-9 (33).

Two  additional variables are liquid alkalinity  and gas-liquid contact. The presence of
substantial proportions of carbonate ion in the black liquor at high pH of 12.0 or more gives
the liquid a large potential buffering capacity against pH reductions caused by contact with
C02 from the recovery furnace flue gas. Increasing the  degree of gas-liquid contact may
result in either a substantial increase or decrease in malodorous gas emissions, depending on
the characteristics of the black liquor and the flue gas temperatures.

Under certain conditions, the direct contact evaporator can act as an air pollution control
device to  absorb H2S from the recovery boiler combustion  zone flue  gases (28, 30, 32). A
particular advantage in direct contact evaporation, where high degree black liquor oxidation
is practical, is absorption of H2S  from the combustion  zone during periods of recovery
boiler upset.
                                        10-54

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                                    TABLE  10-9
               EFFECT OF BLACK LIQUOR pH  ON  H2S EMISSIONS
                  DURING DIRECT CONTACT  EVAPORATION (4)

                                               H? S Concentration*
           pH           Na2S          Inlet         Outlet          Change
           _____


          12.6           14.2            11              35             +24
          12.3           18.3            27            122             +95
          12.1           15.3            24            180           +156

          *Computed in ppm by volume at 0 C and 760 mm Hg (32  F and 1.0 atm).
     10.6.3   Complete  Multiple-Effect Evaporation  and  Indirect Contact  Evaporation
             Comparison

Indirect contact evaporation is used to concentrate black liquor from 50 to 60-65 percent
solids to eliminate the possibility of odorous gas release during direct contact evaporation.
Systems  employing complete multiple-effect evaporation and indirect contact evaporation
have been installed at new kraft mills in the United States and Canada (35, 36).

The multiple effect evaporation to the dry solids concentration used for injection into the
boiler eliminates a potentially large source of malodorous sulfur gas emissions resulting from
direct contact evaporation. The technique has proved successful in  extensive experience in
Scandinavia  (37). Multiple effect evaporation to virtually eliminates the recovery furnace as
a source of  malodorous gas emissions. It is then only necessary to operate the recovery
boiler so  as  to  minimize sulfur gas  emission. The Scandinavian system also has a lower
moisture content in the exit gases from the recovery boiler than systems using direct contact
evaporators, therefore reducing plume opacity caused by condensed water droplets.

Indirect contact evaporation can also reduce sulfur emissions, but by introducing the water
vapor from the air cascade into the furnace, the combustion equilibrium conditions change
and  the  temperature  decreases in the primary air  combustion  zone. The air-cascade
evaporation  system  may  cause  corrosion  and  particulate  plugging of  the  rotary heat
exchanger.

Direct contact evaporators tend to have greater heat economy than  air cascade evaporators,
but lower heat economy than complete use of additional multiple-effect evaporators (38).
The  difference in  heat  economy grows  in  importance with  increasing fuel  costs. An
                                       10-55

-------
additional factor in direct contact evaporation is that high degree black liquor oxidation can
reduce the heating value of black liquor by as much as 5 to 10 percent (39).

10.7   Flue Gas Scrubbing for Gaseous Emissions

Flue gas treatment with absorption  of  the  malodorous  gases would be  an economical
method of converting old recovery boilers to meet present standards for  air pollution if it
were not  possible to rebuild  the plant to the low odor  system. This equipment can be
erected after  the precipitator during the operation  of  the  boiler and connected in a
reasonably short time if space  is available. Any changes requiring a long downtime for the
recovery boiler are prohibitive because of production loss. With fuel prices below $1.42 per
million kj ($1.50 per million BTU), it is  almost never feasible to rebuild an existing unit to
gain better heat economy even  though a new design would be feasible for a new mill.

In an existing recovery boiler,  it may not be possible to operate without release of S02 and
H2S. The  existing direct contact evaporators can be used to absorb SO2. The absorption of
H2S may  also be possible, but in most cases insufficient to meet the regulations regarding
emissions of this gas.
                6
Methods and equipment have  been designed for the absorption of H2S. Three of these
systems are commercially available. Only the last type has been installed in more than one
mill. These systems are:

     The B. C. Research Council absorption scrubber,

     The TRS Weyerhaeuser absorption scrubber, and

     The Gadelius-Misubishi absorption scrubber.

The B. C. Research Council's (B.C.R.C.)  method uses  a rather concentrated  Na2CO3
solution  for  absorption  of the  H2S  (29,  40).  The  principle is that  the  carbonate
concentration should be in equilibrium with the CO2 partial pressure in the gas to avoid
excessive loading with Na2CO3 and to avoid an increase in the lime consumption in the
recausticizing department. The  carbonate solution is sprayed at the top  of a packed tower,
and the liquor moves downward countercurrent to the flue gas.

The liquor is extracted at the bottom of the scrubber and is pumped to an oxidizing unit
where Na2S is converted to Na2S203. Enough iron compounds are normally available in the
remaining dust after the precipitator to act as a catalyst for a rapid conversion to Na2S203.
This conversion was observed during the operation of scrubbers for generation of hot water
by heat recovery from flue gas (41).
                                        10-56

-------
The method is the property of B.C.R.C., and SF Products Canada, Ltd., manufactures and
markets  the  equipment. It has been discussed in combination with heat recovery for
generating hot water for the bleach plant, in which case the feasibility of the installation
seems satisfactory, especially at  the present level of fuel prices.  The pressure drop on the
flue gas side  depends on the concentrations of H2S and  CH3SH in the flue gas, as these
determine the  number of  exchange units in  the packed tower.  One application  was
calculated at 750 Pa water (3 inch w.g.)  pressure drop for reduction from 600 ppm to
5 ppm, and the same pressure drop for the heat recovery section.  Compressed air is used for
the oxidation unit. A bleed-off of the scrubbing liquor and fresh alkali make-up is necessary
to maintain the correct liquor composition.  A typical arrangement is shown in Figure 10-25.

The TRS System was developed  and patented by the Weyerhaeuser Company (42) and uses
a  Nalco  water  solution of chelated ferric  chloride in a  proprietary  formulation.  The
absorbed H2S forms elemental sulfur, and a special packing  is used to avoid plugging. This
package was developed  and is marketed by Fritz W. Glitsch & Sons, Inc.

The TRS-scrubber absorbs up to  99  percent of the H2S and collects about 85 percent of the
particulate matter. The resulting slurry of colloidal sulfur, salt, and other materials is passed
through a thickening and washing operation for recovery of the sulfur. Soda ash and air are
used to keep the solution  neutral and  to oxidize the Nalco water solution. The solids
concentration is kept at about 20 percent by bleeding to the green liquor. The sulfur, salt,
and Nalco solution are recovered. Several  scrubber units  are used in parallel; one can be
taken out for cleaning, or the  packing elements can be removed  for cleaning.  The pressure
drop with all units in  operation on the  gas  side  is about 1250 Pa water (5 inch w.g.). A
typical arrangement is shown in Figure 10-26. The above  method  has the advantage of
removal of sulfur from the kraft  process. Such removal  may become a necessity in the
future to  keep the sulfidity of the white liquor in a range  such that corrosion effects on
equipment remain tolerable.

The Gadelius-Mitsubishi method  uses one or two stages of spray  nozzles in  countercurrent
arrangement to  the flue gas flow for the absorption  of H2S in Na2C03 or NaOH. Fresh
alkali is used for the makeup of the absorbing solution. Several units are in operation in
Japan. This scrubber has very low pressure drop on the gas side,  about 245 to 375 Pa (1 to
1.5 inch w.g.). None of these methods will have a high absorption efficiency  for CH3SH,
CH3SCH3, or CH3SSCH3. The absorption efficiency for the other sulfur compounds must,
therefore, be  quite  great  in  some cases to bring the TRS down  to  the limits of the
regulations.  This fact  should  be considered at the design  stages for flue gas  scrubbing
equipment.

Other methods for removing reduced sulfur compounds, such as  scrubbing with an alkaline
suspension of activated carbon,  were suggested after laboratory  studies of absorption  (43,
44). The recently recognized importance of both increasing  the furnace temperature in the
                                       10-57

-------
                          f HS-"Free"  Flue Gas
o+ni-lc 1 *-
Possible Heat
Recovery Section — >
i
Demister^^^
Qnrnv NAT? l^i
Park inn
AKcnrhor fe
MUoUl UCl w
Flue Gas —
From R.B. """""•*
Water For
J.

*•»,.
^

to
s
/

Cooling 8 Saturation ]
--J
^^ Cir
Possible S0o+— w
/
-^
\
T
^-
^t\
viy
cu lation
Pump
Alkali
Make Up+
(Water)
<
i
^
t
f
\
4
(Air-02)Exit
\
V
Oxidizer
Compressed
I4 Air
o— o— o— o
^-
* "
Discharge To
Pulp Mill
Dust Recovery Section
Na2C03
Na2S203
Na2S
NaSH
                          FIGURE 10-25
         BRITISH COLUMBIA RESEARCH COUNCIL DESIGN FOR
                   H2S ABSORPTION  SCRUBBER
                              10-58

-------
       Furnace
              V
                   Cyclone
                   Evaporator
p
Ol
                   SALT CAKE
                   RECOVERY
                                                     ABSORPTION  TOWERS
                                                                                      T
                                                   : Spent  Scrubbing Liquid
                                                           •
                                                   	.*






Catalyst
Regenera-
tion tanks






* Oo
       STACK
Return
Scrubbing
Liquid
                                                   REGENERATION TANKS
               SULFUR
               RECOVERY
                                 TNa204
                               MAKE-UP

                                                FIGURE 10-26
                         GLITSCH-WEYERHAEUSER DESIGN FOR A TRS SCRUBBING SYSTEM
       RECYCLE

-------
   PSYCROMETRIC  CHART.
          Flue Gas Conditions
         Mixing Lines
  Ambient Air
   Conditions
    HUMIDITY
      RATIO
         DRY  BULB  TEMPERATURE


            The mixing  lines  show extention

Flue  Gas Condition          I         I

Ambient Air  Condition       3        3a
Condensation
            Yes       No       No
       From (T) to (2)

             FIGURE 10-27

FORMATION OF VISIBLE PLUME THROUGH
   CONDENSATION  OF WATER VAPOR
 la

 3a

No
                               10-60

-------
primary air combustion zone to reduce the emissions of sulfurous gases (9, 11, 13, 14); and
increasing  the  residual  alkali  to  reduce  the formation  of  some  of the malodorous
compounds, possibly needs further examination since this method can probably reduce the
emissions of malodorous gases. These changes probably arc easier to adopt operationally
than an absorption system.

SO2 can be removed from the flue gases by any of several designs of simple scrubbers. Two
of  these are  the SF Scrubber-Modo  System and  the  Warkaus scrubber (45),  which is a
double venturi scrubber arranged with the gas and scrubbing liquid in  parallel flow with
practically  no pressure drop on the gas side.

One major problem, when using a scrubber installation,  is  that the  flue gas  becomes
saturated and  the visibility of the plume from the stack increases considerably. This problem
is of less importance if  the scrubber is used also for hot water generation for  the bleach
plant. The humidity of  the flue gas is then reduced, therefore, reducing the plume. The
plume can  be virtually eliminated if cold water is available to reduce the humidity ratio still
further. (See Figure 10-27.) Cooling of the flue gas adversely  affects the plume buoyancy.

10.8  Collection of Particulate Matter from Recovery Boiler Flue Gas

     10.8.1   General Conditions

Formation of the particulate  matter in the  flue gas was explained  briefly in 10.2.4. The
amount of dust, its particle size distribution, and its handling characteristics depend on the
reaction conditions in the recovery boiler.

The dust load varies between 40 and 75 kg per metric ton of dry solids (80 and 150 Ib/ton).
This gives  a range from  140 to  680 kg per metric ton of  pulp (280 to 1360 Ib/ton) for
extreme  combinations of  operating conditions of  recovery boiler operation and cooking
yield.

The dust concentration in the flue gas calculated on a dry gas basis at standard conditions of
15.5° C and 760 mm Hg  (60° F and 29.92 in. Hg) will vary correspondingly  from 9 to
25 g/m3 (4 to 11 gr/dscf). Only the dust that requires high efficiency collection of small size
particles is  included  in these figures. Coarse dust from the sootblowing, which will  settle by
sedimentation at normal velocities in a settling chamber, is  not included since this will be
separated in the distribution  plenum  before the collector elements  and will  not likely
present any burden to the dust collector.

The dust load can  increase above the given figures if part of the dry  solids is  burnt in
suspension. The dust from the  kraft process consists chiefly of Na2S04  and Na2CO3. The
concentration of Na2C03  depends mainly on  the ratio between sodium and sulfur in the
                                        10-61

-------
flue gas. Traces of NaCl and Na2C03  are normally  found at mills operating with sealogged
wood. The concentration of NaCl can be considerable. The concentration of dust in the gas
                                                                                    o
might also be increased by pulping of sealogged wood. The  dust particles can contain small
amounts of Na2S.

The present trend of increasing the temperature in the combustion zone to achieve an exit
flue gas from the boiler that is virtually free from H2S and S02  also probably increases the
dust content in the flue gas, as compared to recovery units of similar, but earlier design.
Increases of about 40 percent are possible for a new unit as compared to an old unit at the
same  mill and operating with  the same kind  of liquor. The most important data for the
design of a dust collector are the gas flow, the dust load, and the particle size distribution.
Design data from the recovery  boiler manufacturer  should be used  with appropriate safety
margins for the dust collector to cover upset conditions.

The small particle size of the dust makes mechanical dust collectors unsuitable for cleaning
of flue gas from  recovery boilers.  Baghouses are  not suitable  because of the handling
characteristics of the dust. Electrostatic precipitators are, therefore,  used for recovery boiler
installation.  The economic  collection efficiency  based on the price of the  recovered
chemicals,  capital and operation costs, and payoff time are estimated at between 92 and 97
percent for  different conditions. Such a collection efficiency  generally will not  control
particulate emissions to the degree needed to meet air pollution control regulations.

Wet scrubbing of  the flue gas  with water or with thin black liquor, which was previously
oxidized at an efficiency of 99 percent  or more and stripped  of  methyl mercaptans  and
organic sulfides, can provide sufficient collection efficiency to  meet air pollution control
regulations. This alternative  is, however, less  attractive as  the chemicals are present in an
aqueous solution. The  discharge back into the process causes  an increase of the inactive
chemicals,  operational  difficulties, and slightly increased losses in other departments. This
method may, however,  prove the most economical for an old mill under certain conditions.

A combination of an electrostatic precipitator with about 95 percent collection efficiency,
followed  by  a scrubber to  achieve  a total  efficiency of  99.5  percent, is economically
favorable if used in  combination with heat recovery from the flue gas. The application of a
tail end scrubber  must  be investigated  thoroughly  for particle  size distribution. The
collection  efficiency is high for dust  larger than 1 (j.m (3.9 X  1(T5 in), even with  a  low
energy scrubber (probably because of  the hygroscopicity of  the dust), but decreases rapidly
with decreasing particle size.

     10.8.2   Electrostatic Precipitators

The functioning of an electrostatic precipitator is based on  movement of charged particles
of dust  in  an electrostatic  field. The emission electrodes are  given  a negative potential
                                         10-62

-------
ranging from  30,000  to  80,000 volts depending  upon  operating  conditions. They emit
electrons that charge the dust particles, and at the same time they form, together with the
grounded collecting electrodes, an electrostatic field.

The high tension  negative current  is achieved with a transformer and a set of rectifiers,
normally forming  one unit and  monitored by a spark rate control unit, designed to give a
certain  number of  flashovers per second. The theory  and technology  for  electrostatic
precipitators are comprehensively treated by White (46) and Oglesby (47).

Development work to increase the reliability of the operation and to decrease the capital,
operating, and maintenance  costs is still in progress. The engineer must consider not only
the influence  of  the  gas and  dust characteristics,  but also the process to which the
precipitator has to  be applied. The dust and  gas characteristics for the recovery  boiler
process are more favorable for the precipitator operation with the North American recovery
boiler system than with the Scandinavian system. This difference is mainly because of higher
water vapor content in the  flue gas from American direct  contact  evaporators. This fact
should  be recognized  when new methods for achieving a high black  liquor  dry  solids
concentration  are  considered,  or  when a  change  from  sootblowing with  steam  to
sootblowing with compressed air is considered.

The negatively charged dust particles  move to the grounded collecting electrodes (plates),
where  they  transfer  a  certain part of their charge,  depending on the resistivity of the dust.
The dust particles  are  kept on the plate by the electrostatic  field and the remaining charge.
The dust is removed  from the plates by rapping the  plates. The acceleration in the surface of
the plates during  the  rapping must be sufficient  to dislodge the dust from the plates by
shearing action. The dust falls downwards, mainly following the plates like a  web at ideal
conditions, and is collected  at the bottom of the precipitator.  The ideal rapping  system
should  dislodge the whole  dust layer with one  shock  wave passing through the plates.
During  the shock  waves,  some flakes  of dust that are near the plates retain little charge.
These flakes are shaken loose and are very easily entrained in the gas flow. The  dust layer
should  be allowed  to build up  to a  certain thickness between  rappings to minimize
reentrainment.

The discharge  or  emitting  electrodes collect dust particles with  a positive charge and
therefore  need rapping. The dust collected on  the emitting electrodes can vary between
needle-like deposits to thick  layers of dust if they are evenly  distributed. The former type is
often found if the dust contains large amounts of chlorides, and the latter if the normal
Na2S04-Na2C03  dust is  sticky. The acceleration  during the shock waves at the rapping
usually  is about 20-40  times the acceleration due to the earth's gravity (g's) for normal dust.
It is very difficult to  get an even distribution of the  acceleration forces. The acceleration
must be increased  to above 200 g's if sticky  dust is generated at combustion. High stresses
are caused above  200 g's  on the  components of the emitting and  collecting system and
                                        10-63

-------
decrease the  periods  between  major maintenance. These  figures relate to  parts  of  the
collecting plates where the acceleration is at  a  minimum. Values considerably higher  can
normally be measured at the points where the rapping forces are applied.

The  dust, collected  at the bottom of the precipitator chambers, is discharged in different
ways. The dry bottom design used prior to the 1950's in North America was changed to a
wet  bottom  design  because of required  maintenance of the dry bottom conveyors. The
Scandinavian  precipitators used dry and  wet  bottom designs at the  start, but changed to
mechanical  conveyors after  experiencing  corrosion  with  the wet  bottom design. The
difficulties with the conveyors  were eliminated by using a heavier steel bolted chain with
larger  pitch  (about  15 cm  (6  in)) and  bearings of graphite  for the shafts to  reduce
maintenance.  Screw conveyors  for the transverse  transport at the end were changed to
Buhler conveyors of the same type as used for cement kilns.

The  American trend has been toward dry bottom design with the low odor concept, but the
problems  associated  with  heat  distribution  and  heat  insulation  seem  to  have been
overlooked. The bottom of the  precipitator is  heated only by the collected dust as the ideal
gas velocity below the electric fields is zero. The same conditions are valid for the walls of
the precipitator  chamber at the sides of the electric fields that are heated only by radiation
from the collecting plates nearest the walls. Any gas passing below the  fields and at the sides
of the  fields will decrease the collecting efficiency by "sneakage" and is almost intolerable
in high efficiency precipitators.

The  collection efficiency is easily calculated assuming an even gas distribution (i.e., the same
velocity in all parts of a cross section of the precipitator fields), an even dust distribution in
the  gas entering the  precipitator, and an  instantaneous mixing of the gas over the cross
section. Deutsch's formula gives  the collection efficiency as:

                                 i? = 1 - exp  (- wL/Rc)                        Eqn. 10-1

where:

     17 = the collecting efficiency of the precipitator

     w = the overall migration velocity of the dust in m/s (ft/sec),

     L = the total effective length of the electric fields in m (ft),

     R = the distance between the emitting wire and the collecting plates in m (ft),

     c = the gas velocity  in m/sec (ft/sec).
                                         10-64

-------
Deutsch's formula  does not take into account the possible variation in w with the velocity,
the possible reentrainment of particles,  or the influence of particle  size distribution. This
formula is valuable to calculate the influence of limited changes (±15 percent) in gas load on
precipitator efficiency.

Attempts were made to refine the Deutsch formula to apply it to calculating the difference
in performance  of particles  of  different  size and  therefore  evaluate results from test
precipitators. Deductions based  mainly on results  of operation with sodium sulfate, dust
from pulverized coal firing, and certain metallurgical processes (48) show that the collecting
efficiency is:

                                 17 = 1 - exp (- wL/Rc)k                       Eqn. 10-2

where the K (dimensionless exponent) normally varies between 0.5 and 0.8.

Equation 10-2 gives less increase in the collection efficiency for a given relative increase in
collecting surface than the original Deutsch's formula.

The  above formulas can be written in another form that might  give a better illustration of
the relationship between precipitator size and the gas flow versus the collection efficiency.
Observing that the velocity, c, is:
                                             Q
                                         c = jrii                               Eqn. 10-3

                                     V = A X B X  L                            Eqn. 10-4

where:

     G = gas flow, m3/sec (ft3/sec)

     A = effective height of fields, m (ft)

     B = effective width of fields, m (ft)

     V - effective volume of electric fields, m3 (ft3 ).

Equations 10-1 and 10-2 can be written as:

     17 = 1 - exp (-wV/RG)                                                    Eqn. 10-5

     r? = 1 - exp (-wV/RG)k                                                   Eqn. 10-6
                                         10-65

-------
In addition:

                                  B = 2 X R X n                              Eqn. 10-7

                                  C=2XnXAXL                          Eqn. 10-8

where:

     n = number of passages in the field,

     n + 1 = number of collecting plates in each field, and

     C - total effective collecting plate area, m2 (ft2).

Equations 10-5 and 10-6 can be rewritten as:

                                 T? = l-exp(-wC/G)                         Eqn. 10-9

                                 r) = 1 - exp (- wC/G)k                      Eqn. 10-10

Based on a  reasonably large number  of  tests from  precipitators  on Scandinavian type
recovery boilers,  equation 10-10  is  recommended  for  use with  k = 0.7 for  estimating
purposes. For American low odor units that have a higher exit flue gas temperature, 205° C
or higher as compared to  160° C (400° F vs. 320° F), a value of k = 0.6 is recommended.
The relative  changes in precipitator size with  varying  collection  efficiencies are shown in
Figure 10-28.

Comparative tests with different designs of components for electrostatic precipitators must
be executed with very accurate control of the flue gas and dust conditions for the  tests to be
of real value.

The following  design parameters are available  for the design of electrostatic precipitators.
These are:

     1.   Flue  gas flow,

     2.   Flue  gas temperature,

     3.   Flue  gas composition of dry gas,

     4.   Flue  gas moisture content,
                                         10-66

-------
o

UJ
u.
LL
UJ
O
UJ
O
O
                                                     HOT  PRECIPITATOR
                                                     GAS  TEMPERATURE
                                                          380 °C
WARM PRECIPITATOR

GAS TEMPERATURE
    99.4
        20        40    60   80 100        200   300     500       1000  1500



                                FILTER  PLATE AREA, m2


                                      CONDITIONS


                             GAS  VOLUME  =   5400 m3/ t D.S.

                             CO            =   16-17% by VOLUME

                             SOLIDS CONC.  =   60%




                                FIGURE 10-28


           ELECTROSTATIC  PRECIPITATOR  SIZE AS A  FUNCTION

                     OF COLLECTING  EFFICIENCY (37)
   5.   Flue gas dust content,



   6.   Particle size distribution of dust,



   7.   Operating pressure for precipitator,



   8.   Variations in above  data  because of temporary overloading, sootblowing, and

        possible additional gas flows as from dissolving vent stacks, and mix tank vent

        stacks, and
                                    10-67

-------
     9.    Maximum  possible  variations  because  of  changes  in the  pulp  production
          parameters, such as species and pulp yield.

The  above list represents more  than what was  previously available for the design of a
precipitator.  The demand for accurate information, however, has increased because of the
high efficiency currently required. A  safety margin above the boiler manufacturer's figures
for factors such as gas flow and temperature, should be allowed to accommodate the errors
caused by the measurement errors.

The following parameters can be determined for the electrostatic precipitator itself and the
regulations for particulate emission. These are:

      1.  Dust collection  efficiency,

      2.  Number of parallel precipitator chambers,

      3.  Number of electric fields,

      4.  Number of transformer-rectifiers,

      5.  Gas inlet arrangement,

      6.  Type of rapping and rapping frequencies,

      7.  Maximum gas velocity in the precipitators,

      8.  Type of dust discharge, wet or dry bottom,

      9.  Material of shell for precipitator chambers, and

     10.  Heating of shell and heat insulation.

The dust collection efficiency must be chosen with some consideration for the degeneration
that often takes  place  in the physical  condition of  an electrostatic  precipitator. The
alignment of the emitting electrodes with  respect to the collecting plates is important. The
alignment is easily upset during the exchange of the high voltage insulators if poor guidance
is  provided for the adjustment. Because the collection efficiencies can decrease  if the
alignment is faulty, performance guarantees, over a  two-year period from startup, are often
requested. Curves for changes in efficiency as function of various parameters, such as gas
flow and gas temperature, should be included in the guarantees.
                                         10-68

-------
The number of  fields  and chambers must be determined, taking into consideration the
decrease in collection deficiency, if and  when  one field goes out of operation, causing a
subsequent change in outlet dust concentration.

The number of  transformer-rectifiers must be  determined for possible variations  in the
specific dust load, temperature, and moisture content at the precipitator inlet. Varying dust
loads in different chambers distort the distribution  of the current to different chambers if
the chambers are coupled in parallel. The more dust in the gas, the lower the voltage and the
lower the emission (i.e., the chamber with  the highest dust concentration will determine the
voltage on all parallel fields).  But this chamber will emit less current than the other parallel
chamber(s). The  efficiency  of the fields will, therefore, decrease, especially in the field with
the high dust load. This will disturb the next field downstream. The first two fields in each
chamber should, therefore, have separate  transformer-rectifiers, but coupling the third or
later fields in parallel to the same transformer probably is justified if the dust concentration
is very low in these fields.

The gas inlet arrangement should allow isolating the precipitator for maintenance, achieving
a good gas distribution,  and avoiding buildup of dust in the lower part of the inlet plenum,
which will eventually disturb the gas distribution.  Extension of the  bottom conveyor to
cover the bottom of the inlet duct is a good solution if sufficient baffling is arranged to
avoid sneakage of gas below  the fields. Guide vanes and  gas distribution plates  must be
rapped efficiently to give satisfactory performance for a prolonged period.

The rapping  mechanisms should be sufficient to clean the emitting electrodes and collecting
plates  without being actuated for extended  periods.  The main part of the  dust  layer
probably is discharged in the beginning of the rapping and extended rapping tends to break
away  flakes  of dust, causing "snowflaking." Tests show large increases in dust losses with
increases above the optimum frequency of rappings.

The gas velocity  in  an  electrostatic precipitator must  be limited  to avoid snowflaking. A
velocity not  exceeding 1.0 m/s (3.5 ft/sec) is recommended for American design (47). This
normally is the practice  also for Scandinavian precipitators, except where the precipitator is
followed by a scrubber for heat recovery, in which case higher velocities are acceptable.

The dust discharge can be accomplished with black liquor pumped over  the bottom  either
continuously or intermittently. Some designs  use impellers in the bottom to stir  the dust
into the black liquor. The black liquor is then discharged to the cascade evaporator or to a
mixing tank. The gas velocity between the baffles that prevent sneakage of gas below the
electric  fields is very low. Here the gas acts, to  some extent, as a revolving gas volume at
approximately  the wet  bulb  temperature. Corrosive conditions are easily reached in the
lower part of the  fields, especially if the black liquor  contains a high  concentration of
                                        10-69

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chlorides. Such conditions adversely affect the lifetime of the emitting electrodes and, in
some cases, have also damaged the lower parts of the collecting plates.

Using a dry  bottom avoids these difficulties, but proper chains and scrapers with sufficient
strength and stability must be provided. The transverse conveyor, preferably at the inlet end
of the  precipitator,  should have sufficient capacity  to  accommodate the  uneven feeding
from the different fields. Drag  chain  conveyors of the same type as normally used for
cement kilns, give  outstanding service.  If screw conveyors are used, they  must have a
relatively large diameter and  sufficient stiffness to avoid vibrations. The troughs for the
transverse conveyor should not have more depth than is the necessary minimum.

The  reason  for  minimizing  the depth  is to  keep  the temperature  high  and  avoid
condensation of water vapor into the hygroscopic dust. The troughs must be well insulated.
The  dust is  preferably discharged via  a rotary  valve to seal  against air leakage into the
precipitator chamber. (The induced  draft fan operates better if it is placed on the clean side
of the  precipitator.) Air leakage can  cause  local corrosion,  which  can disturb the gas
distribution.

The  rotary valve should be isolated from the mixing tank by a short screw conveyor with an
air  screen to prevent diffusion of  water vapor  up into the  rotary valve and  the dust
conveyor. This arrangement has proved reliable in avoiding plug-ups of the rotary valve and
dust conveyors. The necessary amount of air for the air screen is less  than 1.7 m3/min
(60 ft3/min) per precipitator chamber, and this air normally passes out through the vent on
the mixing tank.

The  normal precipitator housing in North America consists of tile walls and concrete roof
and  little heat insulation. This design is susceptible to cracking with subsequent air leaks and
inside corrosion. The concrete housing used in Scandinavia was designed to avoid cracks by
proper consideration  of heat conduction. This design was used  as long as the gilled cast iron
economizer  was still in use. The precipitator must be water washed at intervals of a few days
to two  weeks. The temperature of the flue gas to the precipitator  can be as low as 110  C
(230° F) for about one work shift immediately following a water washing. Hot precipitators
with steel plate  chambers placed  between  the boiler outlet and the economizer that
operated  at  400° C (750° F) or slightly lower, came into  use in 1957. This  design was
adopted  for  use  with  the current long vertical tube economizer. The exit flue gas
temperature is around 160° C (320° F). Experience seems to indicate that a steel plate shell
can  be  used down to 127° C (260° F)  if the precipitator is equipped with very  good heat
insulation of at least 10 cm (4 in) of mineral wool above the  top  of the stiffeners for the
shell. All stiffeners and flanges in ducts must be insulated.

Heating of the shell is probably not necessary if good heat insulation is supplied. Heating of
precipitators was standard practice before starting  of the operation in Scandinavia, but has
                                         10-70

-------
 now been abandoned. No corrosion damage has been observed, to date. The present practice
 in  the  U.S.A. and  Canada  of  not  covering the  stiffeners with sufficient insulation is,
 however, damaging,  and if it is economically feasible to use a heated chamber instead of
 good insulation, the method should be used.

 Dust emission data  from recovery boiler electrostatic precipitators used presently in the
 U.S.A.  are given  in  Table 10-10. The values are average values of decile groups. The  table
 includes information from precipitators on 87 recovery boilers.
                                   TABLE  10-10
                   AVERAGE  PARTICULATE  EMISSIONS FROM
                  RECOVERY  BOILER  ELECTROSTATIC PRECIPI-
                      TATORS IN THE UNITED  STATES  (49)

               Emission decile                  Average Emission rate
                                             kg/t              Ib/ton

               First (lowest)                   1.1                2.1
               Second                        1.7                3.3
               Third                          2.4                4.8
               Fourth                        3.4                6.8
               Fifth                          6.2               12.4
               Sixth                          8.5               17.0
               Seventh                        9.2               18.4
               Eighth                        14.2               28.4
               Ninth                         23.2               46.3
               Tenth (highest)                37.6               75.2
The  first decile  represents dust concentration of  0.11 g/m3 (0.05 gr/dscf), the  fifth
0.46 g/m3 (0.2 gr/dscf), and the tenth 2.3 g/m3 (1 gr/dscf). The collection efficiencies are
99 percent, 95 percent, and 80-90 percent, respectively.

     10.8.3   Liquor Scrubbing of Recovery Boiler Flue Gas

The  cyclone evaporator and the venturi  evaporator scrubber were originally used for the
recovery of heat and for concentration of the black liquor to a level suitable for firing in the
furnace. The cyclone evaporator was a low energy type of scrubber that collected only the
coarsest size fractions of the dust in the flue gas. The venturi evaporator scrubber has high
energy requirements but is capable of collecting finer dust than the cyclone evaporator. The
dust  collecting efficiency was decreased  by  the high viscosity of the concentrated black
                                       10-71

-------
liquor used in these types of scrubbers.  This  liquor cannot  be atomized into sufficiently
small drops for efficient dust collection. The capital cost for the recovery boiler department
was  considerably decreased by the introduction of the venturi  evaporator scrubber, as the
venturi scrubber and its associated enlarged induced fan compared favorably with alternative
types of heat recovery equipment with a  low  efficiency electrostatic precipitator. The
operating costs  were probably  never favorable because of high power consumption  and
rather large losses of Na2S04. A combination  of high interest rate for the capital and a low
price for electric power  can, however, sometimes justify the choice of the venturi evaporator
scrubber, if air pollution control is not an overriding factor. These types of scrubbers were
discussed previously in section 10.6.

To achieve high collection efficiencies for the particulate matter in the flue gas, liquors with
low viscosity must be used, such as thin black liquor or water.  A scrubber using thin black
liquor discharges its liquor with the dissolved dust to a multiple-effect evaporation plant.
The  increased load of Na2S04  and/or  Na2C03 in  the  black liquor  charged  to  the
evaporation plant increases the fouling rate. A comparison between a scrubber using thin
black liquor and  a precipitator should, therefore,  consider the changes needed  in the
evaporation plant.  Such changes include increasing the heating surfaces to compensate for
the increased  fouling rate and to  accommodate the increased  boiling point rise. Another
factor to consider is the cleaning of the evaporation plant by boiling out with water and its
attendant disturbances  in the operation. Oxidized liquor must be used to avoid emission of
H2S, and so allowance must be made for the  corresponding heat loss from oxidation.

Water can be used in the scrubber; the dust  is then collected and discharged as  a water
solution. This solution  cannot be concentrated very much before crystallization occurs. The
recovery of the  chemicals will, therefore, have an influence on  the evaporation plant of the
same magnitude as when black liquor is used as the scrubbing liquid. The emission  of H2S,
however, is avoided if the dust is reasonably free from Na2S.

Existing precipitators which were designed for insufficient collection efficiency to meet the
present regulations for  emission of particulate  matter, or which have degenerated to a lower
collection efficiency because of design  deficiencies and/or inadequate maintenance, can be
retrofitted with scrubbers. The dust amount collected in these scrubbers is reasonably small,
and  influence on the recovery  cycle is  less than if the scrubber installation were to collect
the total dust load of the flue gas.

Using flue gas scrubbers in many cases  provides an attractive solution to increasing the total
collecting efficiency of an existing recovery boiler plant. But the space requirement for an
additional precipitator may make it almost impossible to use an  additional precipitator.
Operational problems can very likely occur  if  long horizontal ducts are used to convey the
gases from an existing plant to an additional electrostatic  precipitator. Using a  flue  gas
scrubber can  then provide a practical  solution that also results in a capital  cost saving. A
                                         10-72

-------
venturi type of scrubber is used in most such cases but it has a.relatively low pressure drop
of 1500-2500 Pa (6-10 in. w.g.).

Another solution was practiced in a number of mills in cold climates. A low energy type flue
gas scrubber can be used for recovery of heat from the recovery boiler flue gas to produce
hot  water for the bleaching plant (and  possibly  also for the brown stock washing). A
scrubber of this type is placed after  the  electrostatic precipitator to collect some of the
remaining dust from the flue gas.  Combinations of electrostatic precipitators of 95 percent,
or even lower, collecting  efficiency with a scrubber,  recovering heat for producing  hot
water, have  given  a combined collection  efficiency of 99.5 percent  and have operated
successfully for several years (41,  50). This is a  very economical combination, as the cost of
the scrubbers is justified by both heat recovery for heating water and  saving in the cost of a
smaller precipitator.  Mills  in  warm  or hot climates where hot water is usually available in
abundance cannot use this approach  effectively. The decrease  in the generation of back
pressure power is an important factor in an economic evaluation, because it can reduce the
value of the saving in fuel by up to 50 percent.

Use  of scrubbers in existing mills  to increase the  total collection efficiency of the existing
precipitators  should  be  carefully  evaluated.  The particle  size  distribution after  the
precipitators  is important to  the  collecting efficiency of  the  scrubber.  The collection
efficiency declines rapidly  with decreasing particle size. The same precipitator design  can
have great variations in amounts of  fine dust at the exit at operating conditions which seem
very similar when based on the data for the electric current and voltage in the precipitator
fields. Therefore, new equipment should  be guaranteed by  the  supplier covering both  the
efficiencies of the precipitator and the scrubber.

The  efficiency of the scrubbing depends  mainly  on the contact surface  area and relative
velocity between the water drops and the  dust  particles. The atomization of the liquid and
the relative velocity between the gas and  the liquor drops can be achieved in either or both
of two  ways, acceleration of the gas and applying  the  liquor through high pressure
atomization nozzles. Pressures  up to 10.3 MPa (1,500 psig) are  sometimes used for  the
atomization of water. Similar drop size is  achieved by using steam or compressed air with a
pressure  of about 0.79 MPa (100 psig). Atomizing by accelerating the gas  to high velocities
can avoid clogging of the nozzles if  the liquid is recirculated. Atomization of the liquid by
using gas velocity consumes much more power  than atomization by high  pressure nozzles.
One  particularly interesting design is the use of co-current water sprays. The  impact from
the water drops will reduce the pressure drop in the venturi throat, and designs are available
in which the impact from  the water drops compensates for the pressure drop  of the gas in
the scrubber (45).
                                        10-73

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10.9   Economy of Recovery Boiler Operation

Recovery of the chemicals and heat from the black liquor dry solids is of vital importance to
the economics  of pulp production. The great variations in climatic conditions and in pulp
yields  for different species of wood make it very  difficult to give a general picture.  The
prices  for electric  power also vary considerably and, consequently, affect the economic
feasibility of back  pressure power generation. Power generation favors feed water heating
with  steam  from  the  turbine  extractions.  The correspondingly  higher  feed  water
temperature might  make changes necessary in the  arrangement of the economizer and air
heater  and in the most economic exit gas temperature.

The changes in the  design and operation of recovery boilers, as a result of the last few years'
accumulation of research data, might make them more economical. Diagrams for the capital
cost for recovery boilers, for electrostatic precipitators, and for the complete recovery boiler
department are given  in Figure 10-29. The price level is for January 1974. The price for the
boiler  is given for steam conditions of 4.2 MPa (600 psig), 400° C (750° F) for  low
back-pressure  power  generation  and  8.4  MPa (1,200  psig) and  480° C  (900° F) for
reasonably high back-pressure power generation.

Two flue gas temperatures, 160° C (320° F)  and  250° C  (480° F), were stated for the prices
for electrostatic precipitators. The lower price for the precipitator at 160° C (320° F) must
be  compared   with  an  increase  in  price  for  an air heater-economizer. The handling
characteristics of the dust are, however, much better at the lower temperature.

Rather high exit gas temperatures were used in the U.S.A.  and  Canada  as compared to
Scandinavia. They reflect the low fuel prices in North America. Figure 10-30 shows the  heat
loss per °F per year in the exit gas for varying firing  rates. The value of the heat loss per  year
for the difference in exit flue gas temperature between American and Scandinavian practice
is shown for varying fuel prices in cost per million BTU of steam.
                                        10-74

-------
Capital Cost
   106 $


     20
Price Level Jan. 1974
"Average Black Liquor Dry Solids"
  A. Recovery Boiler, delivery & erection, 400°F exit gas.
       Low odor with economizer.
  B. Electrostatic Precipitator, collecting efficiency, 99.5%
  .C. Complete R.B. Department, including building ventilation,
       instrumentation, electric power supply, feed water
       treatment, black & green liquor systems connected to
       mill, steam lines to back pressure turbine, stack, &
       dissolving vent stack condenser.
     15  -
     10  -
      5  -
                                  Dry Solids, 10*lb/d
                                                                             Steam
                                                                             psig/°F
                                                                             1200/900
                                                                              600/750
                                                                             Flue Gas

                                                                             400° F
                                                                             320° F
                                         FIGURE  10-29
              CAPITAL COST FOR  RECOVERY  BOILER  DEPARTMENT
                                              10-75

-------
Btu/°F,Year
  4000  _
  X10*
  3000   -
   2000  -I
   1000  -
                       "Average Black Liquor Solids"
                         Black liquor concentration 62%, excess air, 02 at exit,
                         400-320°F represents exit flue gas temperatures
                         according to the American & Swedish design practice
                         respectively.
  $/Year
(400-320)°F
                                     Dry Solids, 106 Ib/day

                                       FIGURE  10-30

                              FLUE  GAS  ENERGY  LOSSES
         80
         x 103
                                                                                      -   70
                                                                                       -   60
                                                                                       -   50
                                                                                       -  40
                                                                                       -  30
                                                                                       -  20
                                                                                       -  10
                                             10-76

-------
 10.10   References

  1.  Rydholm, S. A., Pulping Processes. New York, Interscience Publishers, 1965, p. 777.

  2.  'Passinen, K. Chemical  Composition of Spent  Liquors, p. 183. Gullichsen, J.  Heat
      Values  of Pulping Spent  Liquors,  p. 211. In: Proceedings of the Symposium  on
      Recovery of Pulping Chemicals. Helsinki, Finland, May 13-17, 1968. Finnish Pulp and
      Paper Research Institute and EKONO Oy, Helsinki, Finland,  1969. p 000.

  3.  Vegeby, A., Scandinavian Practices  in the Design and Operation of Recovery Boilers.
      Tappi, 49:103A-109A, July 1966.

  4.  Alhojarvi, J., Summary Report on the Properties of Spent Liquors. In: Proceedings of
      the Symposium on Recovery of Pulping Chemicals. Helsinki, Finland, May 13-17,
      1968.  Finnish  Pulp and Paper Research  Institute,  EKONO, Helsinki, Finland, 1969,
      p. 167.

  5.  Venemark, E., Svensk Papperstidning (Stockholm), 59(18):629-640, 1956. (Swedish).

  6.  Vegeby, A.,  Unpublished Investigation for Institute for Vattenoch Luftvoadsforskning,
      Stockholm, Sweden.

  7.  Safe Firing of Auxiliary Fuel in Black Liquor Recovery Boilers. Black Liquor Recovery
      Boiler Advisory Committee. April 1967.

  8.  Emergency  Shutdown Procedure approved for Black Liquor Recovery Boilers, Black
      Liquor Recovery Boiler Advisory Committee. April 17, 1968.

 J}'.  Bauer, F. W., and Borland, R. M., Canadian Journal of Technology, 32:91, 1954.

 10. / Sillen,  L. G.,  and Andersson, T., Solid-Gas Equilibria of Importance in  Burning
~~~'    Conventional Ca or Mg  Sulfite  Waste Liquor.  Svensk  Papperstidning (Stockholm),
      55:662, 1962. (Swedish).

 11.  Rosen, E., Calculations for the Gasification of Spent Cooling Liquors. Royal Institute
      of Technology. Stockholm, Sweden.  1962.

 12.  Davidsson, S., and  Stelling, 0., Corrosion of Carbon Steel in Black Liquor Recovery
      Boilers. Royal Institute of Technology. Stockholm. 1968 (in Swedish).

 13.  Stelling,  0.,  and Vegeby, A., Corrosion on Tubes in Black  Liquor Recovery Boilers.
      Pulp and Paper Magazine of Canada 70(10):T236, August 1969.
                                        10-77

-------
14.  Lang, C. J., DeHaas, G. G., Gommi, J. V., and Nelson, W., Recovery Furnace Operating
    Parameter Effects on 502 Emissions. Tappi, 56:115, June 1973.

15.  Timmerman,  J., Physio-Chemical Constants of Primary  Systems in Concentrated
    Systems, Vol. 3. New York, Interscience Publishers, 1960.

16.  Wilson, A. W.,  "Big Sky" Mill  Weathers Montana Pollution Battle.  Pulp and Paper,
    45(9):77-81, August 1971.

17.  Rydholm,ibid, p. 610.

18.  Annergron, G. D., Haglund, A.,  and  Rydholm, S. A. Reported by K. Passinen, Ref. 2
    above.

19.  Hultin, S. O., In:  Proceedings  of Symposium on Recovery of Pulping Chemicals:
    Helsinki, Finland,  May  13-17,  1968.  Finnish Pulp  and Paper Research Institute,
    EKONO, Helsinki, Finland, 1969. p. 167.

20.  Lindholm, I., and  Stockman, L., Heat Evolution During Blade  Liquor Oxidation.
    Svensk Papperstidning (Stockholm), 65(19):755, 1962.

21.  Jones, K.  H., Thomas, J. F., and Brink, D. L., Control of Malodors from Pulp Mills by
    Pyrolysis. Journal of Air Pollution Control Association, 19:501-504, July 1969.

22.  Brink, D. L., Thomas,  J. F.,  and Jones, K. H., Malodorous Products from  the
    Combustion  of  Kraft  Black Liquor:  HI.  Rationale  for  Controlling Odors.  Tappi,
    53:837-843, May 1970.

23.  Brosset, C., Chalmbers Institute of Technology, Sweden. Personal communication.

24.  Hisey, W. O., Abatement of Sulphate Pulp Mill Odor and Effluent Nuisances. Tappi,
    34:1-6, January 1951.

25.  Harding, C.  L,  and Galeano, S.  F., Using Weak  Black Liquor for Sulfur  Dioxide
    Removal and Recovery. Tappi, 51:48A-51A, October 1968.

26.  Galeano, S.  F., and Harding, C. L, Sulfur  Dioxide Removal and Recovery from Pulp
    Mill Power Plants.  Journal of Air Pollution Control Association, 17:536-539, August
     1967.

27.   Maksimov, V. F.,  Bushmelov,  V. A., Torf, A. L, and Lesokhin, V. B., Testing the
     Turbulent Flow Venturi Apparatus, Bumazhnaya  Promyshlennost, (Moscow), 40:14-
     15, May 1965.

                                      10-78

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28.  Blosser, R. 0., Cooper, H. B. H., Duncan, L., Tucker, T. W., and Megy, J. A., Factors
     Affecting Gaseous Sulfur Emissions in the Kraft Recovery Furnace Complex. Paper
     Trade Journal, 153(21):58-59, May 26, 1969.

29.  Murray,  F.  E., and Rayner, H. B., Emission of Hydrogen Sulfide From Black Liquor
     During Direct Contact Evaporation. Tappi, 48:588-593, October 1965.

30.  Walther, J.  E.,  and Amberg,  H. R., The  Role of the Direct Contact Evaporator in
     Controlling  Kraft Recovery Furnace Emissions. Pulp and Paper Magazine of Canada,
     72:65-67, October 1971.

31.  Sarkanen, K.  V., Hrutfiord, B. F., Johanson, L. N., and Gardner, H. S.,  Kraft Odor.
     Tappi, 53:766-783, May 1970.

32.  Blosser, R. 0., Cooper, H. B. H., Duncan, L., Tucker, T. W., and Megy, J.  A., National
     Council of the Paper Industry for Air and Stream Improvement. New York. Technical
     Bulletin. December 31, 1969.

33.  Hendrickson,  E. R., Roberson, J. E., and Koogler, J.  B., Control of Atmospheric
     Emissions in the Wood Pulping Industry, Volumes I, II, HI. Final Report, Contract No.
     CPA22-69-18,  U.S.  Department of  Health,  Education,  and Welfare, National  Air
     Pollution Control Administration. Raleigh, North Carolina, March 15, 1970.

34.  Martin, F.,  Secondary Oxidation Overcomes Odor  from Kraft Recovery.  Pulp and
     Paper, 43:125-127, June 1969.

35.  Clement, J.  L., and Elliot, J. S., Kraft Recovery Boiler Design for Odor Control. Pulp
     and Paper Magazine of Canada, 70(3):47-52, February 7, 1969.

36.  Hochmuth,  F. W., An  Odor Control System for Chemical Recovery Units. Pulp and
     Paper Magazine of Canada, 70(8):57-66, April 18, 1969.

37.  Air Pollution Problems  of the Swedish  Forest Industries. Statens  Naturvardsverk
     (Sweden). Publication 1969:3.  1969.

38.  Vegeby, A. ibid.

39.  Arhippainen, B., and Jungerstam, B., Operating Experience of Black Liquor Evapora-
     tion to High Solids Content.  Tappi, 52:1095-1099, June 1969.

40.  Oloman, C., Murray, F. E., and Risk, J. B., Selective Absorption of Hydrogen Sulfide
     from  Stack Gas. Paper Trade Journal, 153(7):92-94, February 17, 1969.
                                      10-79

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41.  Vegeby, A., Canadian Pulp and Paper Association, Montreal, Canada. Technical Paper
     T242. January 23-26, 1968.

42.  Murray, J. S., Tappi Environmental Conference, San Francisco, May 15,  1973.

43.  Bhatia, S. P., de Souza, T. L.-C., and Prahaco, S., Removal of Sulfur Compounds from
     Kraft  Recovery  Stack  with  Alkaline  Suspension  of Activated  Carbon.  Tappi,
     56:164-167, December 1973.

44.  Teller, A. J., and Ambert, H. R.,  Considerations in the Design for TRS and Particulate
     Recovery from the Effluents of Kraft Recovery Furnaces.  Preprint TAPPI Environ-
     mental Conference, May, 1975.

45.  Jafs, D.,  Recovery  of Heat and Chemicals from Flue Gas Using the  Warkaus Venturi
     System. Papper Och Tra (Helsinki), 48(6):337-342, June 1966. (In English)

46.  White, H., Industrial Electrostatic Precipitation. Reading, Addison-Wesley, 1963.

47.  Oglesby,  S. Jr., A  Manual  of  Electrostatic Precipitation Technology. Final Report.
     Contract  No. CPA 22-69-73, United States Department of Health, Education,  and
     Welfare,  National Air Pollution  Control Administration,  Raleigh, North  Carolina,
     August 1970.

48.  Berg, B. R., Development  of New Horizontal-Flow, Plate-Type Precipitator for Blast
     Furnace Gas Cleaning. Iron and  Steel Engineer, 36:93-100, October 1959.

49.  Atmospheric Emissions from the Pulp and Paper Manufacturing Industry. EPA-450/1-
     73-002. September 1973.  (Also published as  NCASI  Technical Bulletin No. 69,
     February, 1974).

50.  Soderstrom, J., In:  Minutes from  the  Swedish  Steam  Users Associations Recovery
     Boiler Conference.  1971.
                                       10-80

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                             APPENDIX  10-1. AIR AND  FLUE GAS  QUANTITIES AT  COMBUSTION OF BLACK  LIQUOR  DRY SOLIDS
O
CO
Assume:  Reduction of smelt is 100 R%
         No formation of Na2Sn or Na2S203
         Dust losses after precipitator negligible.

         Dry Air is T-=B  parts 02 and 1 - r^7 "'
         Humidity ratio of air is w Ib/lb dry air
         Volume ratio of moist air/dry air m = 1  + Tfr w
         Black liquor D.S. composition is:
            carbon, C                       100 c%
            hydrogen, H                    100 h%
            sodium, Na                     100 n%
            sulfur, S                        100 s%
            oxygen, 0 (as difference)         100 o%
            inert oxides                     100 i%
            c+h+n+s+o+i=l

Theoretical complete combustion without excess air
                                                                                             NOTE:  Neither vapor from the water in the black liquor or air from soot-
                                                                                                     blowing has been included. No air leakages have been considered.
                                                                          " (includes C02 , argon, etc.)
                                                                          29
                              Product
                Quantity
               moles/lb D.S.
                                     Ur, moles/lb D.S.
                                                                                                                     Flue gas, moles/lb D.S.
                                                                             Air
                                                            C02
                                    H20
                                                                                                      N2
                              Na2C03      45~98"3107:
                                 1.5 e
Corr.
for02
1.5 X 4.77 me
                                                                        -4.77 mf
                                                                     1.5 X 4.77 I
                                                                     -4.77 (m - 1) f
                                                                                                                              -3.77 f
                                                        3 ?
        The flue gas flow Fz, at z% by volume oxygen in dry gas is Fz =
                                                                                                     100
                                                                                                                                                       Total
C02
H20
Na2S
Na2S04
h ,
2.016 " b
R 32.07
a 4.77 ma
0.5 b 0.5 X 4.77 mb
2 d 2 X 4.77 md
a 4.77 (m - 1) a
n + n ^ Y A. 77 tm IMK

2 X 4.77 (m - 1) d
3.77 a
0.5 X 3.77 b
2 X 3.77 X d
4.77 ma
0.5 b + 0.5 X 4.77
2(4.77 m - 1) d
mb
1.5 X 3.77 Xe      1.5 (4.77 m - 1) e
                                                                        - (4.77 m - 1) f
                              At theoretical complete combustion without excess air
                                     Air consumption, A = 4.77 m (a + 0.5b + 2d + 1.5e - f) moles/lb D.S.
                                     Dry flue gas flow, Fdry = 4.77 a + 3.77 (0.5b + 2d + 1.5e - f) moles/lb D.S.
                                     Total flue gas flow, Ftot = 4.77 m (a + 0.5b + 2d + 1.5e - f) + 0.5b - 2d - 1.5e + f moles/lb D.S.
                                     Water vapor, Fv = 4.7-7 (m - 1) (a +0.5b + 2d + 1.5e - f) + b moles/lb D.S.
                                                 	100 a	 „,

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APPENDIX  10-2.   HEAT   VALUES  VS.  OXYGEN  DEMAND   FOR  COMPLETE
COMBUSTION
Variable
Analysis
Carbon, C
Hydrogen, H
Oxygen, 0
Bomb heat value (X)
Oxygen demand for complete
combustion
Analysis adjusted to "normal"
content of inorganics
Carbon, C
Hydrogen, H
Oxygen, 0
Sodium, Na
Sulfur, S
Bomb heat value (0.734 X)
"Efficient heat value in reducing
Softwood
Units Lignin
A
% 64
C/ f.
/o u
% 30
Btu/lb 11340
10-3lbmole/lb 58.8
A,
% 47.8
O/ A A
/C Q.Q
% 25.6
% 18.2
% 4.0
Btu/lb 8300
Btu/lb 7340
Hardwood
Lignin
B
60
6
34
10620
54.3
Bj
44.8
4.4
28.6
18.2
4.0
7800
6840
Carbo-
hydrates
C
46
6
48
7560
38.2
c,
34.6
4.4
38.8
18.2
4.0
5550
5890
 atmosphere" (reduction for
 sulfide, and heat of evaporation
 for vapor from combustion of
 hydrogen)

"Resulting heat value of black
 liquor" (above value adjusted
 for heat of evaporation of
 water in black liquor)

Oxygen demand for complete
   combustion
   Btu/lb           6390
10"3lbmole/lb       44.2
5890
 40.8
3640
 29.1
     Note: See Figure 10-23 regarding linear correlation.
                                     10-82

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                                   CHAPTER  11

                  LIME BURNING  AND  LIME DUST  HANDLING
The causticizing of liquor, commonly called green liquor, from the smelt dissolving tank by
addition of lime or calcium oxide (CaO) results in the generation of a lime mud or calcium
carbonate (CaC03) sludge. The lime and sludge  are then washed and calcined at elevated
temperatures in either a rotary kiln or a fluidized bed calciner to recover calcium oxide. This
oxide can then be reused for reclaiming additional  white liquor, the  chemical solution for
digesting pulp. The  normal auxiliary fuels used as heat sources for lime mud burning are
natural gas'and residual fuel oil. The two major potential air pollutants from lime mud
burning are the gaseous emissions and the particulate emissions of entrained lime dust from
the burning zone. The gaseous emissions are H2S from the lime mud and, possibly, organic
sulfur compounds from the scrubbing water.

11.1  Rotary Lime Kilns

     11.1.1   Design Features

The  rotary kiln is the  most commonly employed  device for lime mud reburning in kraft
pulp mills. The device is an open-ended inclined  cylinder that is rotated so that lime mud
added at the upper end gradually passes to the lower end and drops out into a bin  as dry
lime. Fuel and air flow countercurrently to the lime from the lower end of the kiln. The kiln
exhaust gases normally pass through a mechanical cyclone collector for  lime dust recovery
and finally through a liquid scrubber for particulate control (1).

Rotary lime kilns employed in kraft pulp mills can range  from about 2.4 to 4.0 m (8 to 13
ft) in diameter and from about 30 to 120 m (100 to 400 ft) in length. They are designed to
burn 36  to  360 t (40 to 400 tons) of  lime (as dry CaO) per day (2). The lime kilns  are
normally inclined at a slope  of about  ten degrees from the horizontal plane and  can be
supported by  two-  to four-supports, depending  on their length. The  lime kilns must be
designed with a number of auxiliary components, including a lime mud feed system, hot lime
conveying system, air inlet and preheating system, gas exhaust system, kiln rotation system,
and  instrumentation systems (3).  Major  kiln design variables include kiln  length,  kiln
diameter, rotation speed  and  angle of incline,  which  influence  solids retention time,
gas-solids contact area, and temperature.

     11.1.2   Operating Parameters

Lime mud at 55 to  65 percent solids  and with  sodium content of  0.1 to 2.5 percent by
weight (as Na20) enters at the upper  end of the  lime kiln  and  passes through successive
                                        11-1

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stages of water evaporation, mud preheating, and lime calcination. Temperatures in the kiln
vary from 150 to  260° C (300 to 500° F), at the upper or wet end, to 1200 to 1300° C
(2200 to 2400° F) at the hottest part of the calcination zone near the lower or dry end.
Energy requirements  for the lime kiln operation are for water evaporation, preheating and
calcining the lime  mud, and power to rotate the  lime kiln, drive the air fans and flue gas
fans, pump  the scrubber liquid, and convey the lime mud and reburned lime. The major
types of fuels burned in lime kilns are natural gas  and residual fuel oil; turpentine and coal
may also be used. Fuel requirements for lime kilns and fluidized bed calciners are listed in
Table 11-1.

Two of the major design variables affecting particulate emissions from lime kilns are kiln
length and diameter.  These variables can affect the amount of particles swept from the kiln
exhaust gases by governing the gas velocity and the  gas-solids contact area.
                                    TABLE 11-1
             ENERGY REQUIREMENTS FOR  LIME MUD CALCINING
                                  SYSTEMS (3) (4)

                 Rotary Kiln                         Fluid Bed Calciner
           kj/t             (BTU/ton)             kj/t             (BTU/ton)

      2.3-4.7 X106*     (2-4 X106)*        2.1-2.5 X 106       (1.8-2.0 X 106)*
      9.3-17.4 X 106**   (8-15 X106)**      8.0-9.2 X 106       (7-8 X 106)**
       *per metric (t) or short ton (ton) of pulp.
      **per metric (t) or short ton (ton) of lime, as CaO.
H2S emissions from the lime kiln are affected by the Na2S  content of the lime mud
(particularly the aqueous phase) and by the presence of Na2S in the scrubber wash water.
The use of digester and evaporator condensate as lime kiln scrubber water can result in the
stripping of organic sulfur compounds  into the exit flue gas. The presence of sufficient
excess air in the kiln can reduce the concentration of H2S in the exhaust by providing an
oxiding atmosphere sufficient for H2S conversion to S02 (5).

11.2   Fluidized Bed Calciners

Fluidized bed  calciners are alternatives to rotary lime kilns for the calcination of lime
mud to lime. The lime mud is first washed to reduce soluble sodium compounds to  a sodium
content of 0.1 to 0.5 percent by weight (as Na20) and then dried on a vacuum filter. The
dried lime mud  at 55 to  65 percent solids is then suspended in the flue gas from  the
                                        11-2

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fluidized beds at a temperature of about 150° C (300° F) to evaporate the water. The solids
are then passed through a two-stage cyclone system to recover the dried lime mud solids and
then fed  into a bed  of fluidized lime pellets formed by  calcination. The bed is kept in
suspension by  the  action  of an  air fan located  below the cooling chamber from which
reburned  lime is removed. Natural gas or fuel  oil is injected  into the suspended bed and
burned  to provide the  heat  necessary  for  the calcination reactions to  take place at  a
temperature of about 825 to 875° C (1500 to 1600° F). The entrained particles and the
combustion gas products then pass out from the calciner to entrain the wet mud and pass
through the two-stage cyclone system and a venturi scrubber for particulate removal (5).

Fluidized  bed calciners are employed at several kraft pulp mills and have lime burning rates
ranging  from 23 to 136 t per  day as CaO (25-150  ton/day). Fuel requirements for fluidized
bed  calciners  are generally lower  than for  lime  kilns  because of the small combustion
chambers  used  which  have smaller radiation  heat losses. Electricity requirements for
fluidized bed calciners, however, are generally greater than for rotary kilns because of the
energy required for suspending the bed and operating the venturi scrubber. Major operating
variables affecting fluidized  bed  reactor operation are  the mud drying temperature, the
calcination zone temperature, the excess air level,  the bed fluidization level, and the sodium
content of the lime mud.

Major operating variables affecting particulate emissions from the calciner unit are the air
sweep velocity  and the lime feed rate. Variables affecting gaseous emissions from fluidized
bed calciners are the same as those affecting reduced sulfur emissions from rotary lime kilns.

11.3  Particulate Emission Control

The major means of  controlling particulate emissions  from  lime kiln and fluidized bed
calciner exhaust  gases  are liquid scrubbing, using either an impingement or  venturi-type
scrubber,  and, recently, electrostatic precipitation. The scrubbing devices are usually placed
following  a mechanical cyclone collector used  either for removal of the  larger lime dust
particles,  as with lime kilns, or for predrying the lime mud for fluidized  bed calciners.
Particulate inlet loadings to scrubbing devices from lime kilns can range from 7 to 35 g/m3
(3 to  15 gr/cu ft)  at standard  conditions  of 21.1° C,  1.0 atmosphere,  dry gas (70° F,
29.92 in. Hg, dry gas). The dust losses constitute about 1 to 5 percent of the total dry solids
load to the kiln.  Particle  size measurements for the above mass concentrations are not
reported with these  data, but the lime particles generally comprise the larger sizes and
sodium  particles the  smaller ones (4). Comparable data  are not available for fluidized bed
calciners,  but it is necessary to use a two-stage cyclone  mud  drying system  to avoid
overloading the venturi scrubber.
                                         11-3

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     11.3.1   Scrubbing Systems

The  major types of scrubbers employed on lime kilns, to date, are the impingement and
venturi types, with cyclonic scrubbers also employed, but to less extent. Impingement type
scrubbers were extensively employed in the past for particulate scrubbing on lime kilns and
have the advantages of relatively low pressure drop and scrubber shower rate, with resultant
reduced operating costs. The devices are limited, however, in their maximum scrubber slurry
water solids concentrations due to possible scrubber plugging. In addition, they normally
have lower  particulate removal  efficiencies  because of less efficient gas-liquid contact.
Impingement type scrubbers have higher capital costs than  venturi scrubbers  on  similar
installations basically due to their larger size and greater complexity.

Venturi scrubbers are commonly used on lime kilns at the newer kraft pulp mill installations
primarily because of higher particulate removal efficiencies than achievable  by the older
impingement type scrubbers. Venturi scrubbing systems can operate with slurry water solids
concentrations of up to 30 percent  by weight without excessive plugging. A summary  of
operating characteristics for kraft lime-kiln scrubbers is presented in Table 11-2 (4).
                                  TABLE  11-2
         OPERATING CHARACTERISTICS FOR PARTICULATE LIQUID
             SCRUBBERS EMPLOYED.ON KRAFT LIME  KILNS (4)
                 Parameter

       Shower rate ratio, 1/m3
                        (gal/103 ft3)
       Slurry solids, % by wt.
       Pressure drop,  mm Hg
                            (inH20)
       Power required,* kW per t/day
                     (hp per ton/day)
       Power required,** kW per t/day
                     (hp per ton/day)
                                                    Scrubber Type
Impingement

  0.54-2.0
   (4-15)
     1-2
    9-13
    (5-7)
0.041-0.049
 (0.05-0.06)
  0.13-0.16
 (0.16-0.20)
  Venturi

 1.73-3.21
  (13-24)
   10-30
   19-28
  (10-15)
0.082-0.099
(0.10-0.12)
 0.27-0.34
(0.33-0.42)
        *per mass of pulp.
       **per mass of lime.
                                       11-4

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     11.3.2  Performance Characteristics

A number of studies were conducted to determine particulate collection efficiencies of lime
kiln scrubbers.  Stuart and  Bailey (6) report that venturi scrubbers were able to  achieve
96-97  percent particulate removal from lime kiln exhaust gases at pressure drops of 1.7 to
2.8 kPa (7 to 10 in water); while Landry and Longwell (7) report that venturi scrubbers can
achieve particulate removal  efficiencies of 98-99 percent at pressure drops of 2.4 to  3.7 kPa
(10  to 15 in water). A  series of studies were conducted on a joint basis by the National
Council for Air and Stream  Improvement and the U.S. Environmental Protection Agency to
establish the particulate collection efficiencies of 66 existing lime kiln scrubbers. Venturi
scrubbers were  able to produce  consistently higher  particulate collection efficiencies than
the impingement scrubbers,  as shown in Table 11-3 (2).
                                    TABLE  11-3
   PARTICULATE COLLECTION EFFICIENCIES  FOR  LIQUID SCRUBBERS ON
                       KRAFT  PULP MILL  LIME KILNS  (2)

                                 Impingement Scrubbers         Venturi Scrubbers
         Parameter              Average         Range        Average        Range

Inlet concentration,* g/m3         27.38        8.00-33.96       18.60       5.85-31.83
                   (gr/cuft)     (11.94)      (3.50-14.81)       8.11      (2.55-13.88)
Outlet concentration,* g/m3         1.78        0.99-3.56         0.73       0.27-2.29
                   (gr/cuft)      (0.78)      (0.43-1.56)       (0.32)     (0.12-1.00)
Removal efficiency, % by wt.      92.2         86.8-96.9        94.8        85.5-99.1
Emission rate,** kg/t                1.78        1.14-2.09         1.01       0.33-2.60
                 (orlb/ton)      (3.55)      (2.28-4.18)       (2.02)     (0.66-5.19)

 *Concentrations are reported at standard  conditions of 21.1 C and 760 mm Hg (70 F and 29.92 in Hg),
  dry gas.
**Emission rates are based on an air-dried ton of pulp basis (i.e., 10% moisture, by weight).
 Information developed  during the study indicates that high pressure drop venturi scrubbers
 can achieve significantly lower particulate levels than reported in Table 11-3 (2). Particulate
 concentrations  at  standard   conditions  of  between  0.02  and  0.11 g/m3  (0.01  to
 0.05 gr/cu ft), corresponding to emission rates of 0.01 to 0.05 kg per air dried metric ton of
 pulp (0.02 to 0.1 Ib/ton), were measured. Very little information exists regarding particulate
 emission control following fluidized  bed calciners. Erdman (8) reports on a high pressure
 drop venturi scrubber following a two-stage cyclonic mud drying system. The pressure drop
 through the venturi is 5.4 kPa  (22 in water). Although the dust carryover from the calciner
                                         11-5

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section  is  12  percent, the  scrubber  emits  a  particulate  concentration  of  0.16 g/m3
(0.07 gr/cu ft),  which corresponds to an emission rate  of 0.24 kg per metric ton of pulp
(0.49 Ib/ton).

11.4  Gaseous  Emission Control

Lime mud  calcining in rotary kilns or fluidized bed reactors can emit H2S, organic sulfur,
S02, and nitrogen  oxides  to  the  atmosphere. The gaseous emissions result either from
materials entering the calcining unit system or  from materials entering the kiln. The major
process  operating variables affecting gaseous emissions  include excess air level, operating
temperature, and solid and gas-phase retention times.

Major input material properties affecting gaseous emissions include  the respective Na2S
contents of the input lime mud and scrubber water,  .organic  sulfur levels in the inlet
scrubber water, and the moisture content of the lime mud. The major  design variable
affecting gaseous emissions from the calcining system are the length and, to a lesser extent,
the diameter for rotary kilns, and the diameter and height for fluidized bed calciners.

A  summary of gaseous emissions  from rotary lime kilns and  fluidized  bed calciners is
presented in Table 11-4.
                                   TABLE  11-4
       GASEOUS EMISSIONS  FROM  KRAFT PULP  MILL LIME KILNS  (2)
 Gaseous
Constituent
H2S
CH3SH
CH3SCH3
CH3SSCH3
TRS
S02
    Concentration
Average         Range
   ppm, by volume
  108
   14
   27
    5

   34
0-500
0-90
0-245
0-11

0-140
         Emission Rate
Average
                                     Range
                                                         kg sulfur per t pulp
                                                        (Ib sulfur per ton pulp)
0.24 (0.48)
0.03 (0.07)
0.02 (0.05)
0.01 (0.03)
0.31 (0.63)
0.14(0.28)
                  0-1.88(0-3.76)
                  0-0.17(0-0.34)
                  0-0.22 (0-0.43)
                  0-0.10(0-0.20)
                  0-2.37 (0-4.73)
                  0-1.11 (0-2.20)
                                        11-6

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     11.4.1   Lime Mud

The most important gaseous emissions from lime reburning systems are malodorous reduced
sulfur compounds. Hydrogen sulfide can be volatilized from the Na2S present in the lime
mud by contact with CO2 from the flue gas. Above a threshold Na2S concentration of 0.2
percent by weight, the generation of H2S  is directly proportional to the residual Na2S
content of the lime mud. This linear  relationship is similar to that  for direct contact
evaporation (9). The amount of H2S released can be  controlled by reducing the residual
Na2S level by more  efficient lime mud washing. It is not  normally feasible, however, to
reduce the residual sodium content in the lime  mud  to  less than  0.1  percent by weight
because of possible mud ring formation.

Prakash and Murray (5) report that H2S emissions from the  lime mud occur primarily from
Na2S dissolved in the aqueous portion and not from the solid portion of the lime mud. The
H2S emissions can be reduced by drying the mud to a solids  concentration of 70 percent by
weight or more before burning.

     11.4.2   Scrubbing Water

The  scrubber water can be a source of both H2S and  organic sulfur compounds emissions
from  kraft mill calcining units equipped with  scrubbers.  The presence  of Na2S in the
scrubber water can result in the release of  H2S by contact with C02 if the liquid pH is
sufficiently low. The emission rate of H2S and organic sulfur compounds increases with the
inlet Na2 S and organic sulfur concentrations, with rising liquid- and gas-phase temperatures,
and with an increasing degree of gas-liquid contact, as represented by the scrubber pressure
drop. The potential for organic sulfur release is particularly great if untreated digester or
evaporator condensates are used as lime kiln scrubber makeup water.

Caron (11) reports that using lime mud wash water instead of fresh water for lime kiln
scrubbing results in stripping of 0.10 to 0.22 kg sulfur per metric ton of pulp (0.2 to 0.4 Ib
sulfur/ton) as compared to an absorption of only 0.035 kg sulfur per metric ton of pulp
(0.07 Ib  sulfur/ton) with  fresh water. Normally, fresh water should be employed as the
scrubbing medium to avoid the  stripping  of odorous  gases. If  condensate waters are
employed, steam stripping should be employed prior to the scrubbing operation.

One U.S. mill has significantly reduced TRS emissions from a lime kiln venturi scrubber by
adding sodium hydroxide to the  scrubber water to raise the pH. The  scrubber  water is
recycled to.the causticizing system (10).

     11.4.3   Combustion Variables

The  major combustion variables that can  affect reduced sulfur emissions from lime mud
calcining operations are the excess air level, the temperature profile, and the mud retention

                                        11-7

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time in the kiln. Caron (11) reports that the TRS emissions from the combustion zone are
minimized at excess oxygen levels of four percent by volume or greater. Though no definite
patterns have been established, the kilns that have cooler wet-end temperatures tend to have
relatively higher reduced sulfur emissions because the sulfur compounds can be volatilized
without burning. Sufficient retention time must be provided at temperatures above 760° C
(1400° F) to oxidize the reduced sulfur compounds.

Walther and Amberg (12) report that shorter lime kilns tend to  have lower reduced sulfur
emissions  than  longer lime kilns, though no definite correlation could be established. The
probable reason is that  short lime  kilns must operate at higher  average temperatures
throughout than the long kilns to achieve an equivalent degree of calcination. The result is a
more complete oxidation of reduced sulfur compounds. An additional factor is that the
evolution of Na2S at low temperatures in oxidizing atmospheres promotes H2S formation;
its evolution at higher temperatures promotes S02 formation (13).

Limited data  indicate that  reduced sulfur  emissions  from  fluidized bed  calciners are
minimal.  This  may be   due to the  relatively long  retention  time  at uniformly  high
temperature which provides for efficient oxidation of the sulfur compounds (14). One test
shows an emission rate of less than 0.01 kg sulfur per metric ton of pulp (0.02 Ib/ton). Flash
drying of the mud tends to minimize H2 S formation in the fluidized bed units.

The  burning of digester  and evaporator noncondensable gases in the lime kilns brings an
additional source of sulfur compounds to the units. The conversion of these materials to
S02  is essentially complete because they are added with the primary air  at the hot end of
the lime kiln and so have sufficient retention  time for complete combustion to take place
(15). The addition of green liquor dregs with  the lime mud to the cold end of the lime kiln
can substantially increase the reduced sulfur emissions, because these materials are normally
contaminated with Na2S from the green liquor. There is also insufficient retention time at
high enough temperatures for complete oxidation to take place.

     11.4.4   Sulfur and Nitrogen Oxides

The  concentrations of sulfur oxides in lime-kiln exhaust gases are normally minimized
because the CaO can act as an efficient adsorption and reaction medium to form CaSO3 and
CaS04. Long kiln length, with sufficient oxygen and high calcination efficiencies, promote
efficient S02  removal. To date,  no  adverse effects on  lime kiln operating efficiency were
traced to  the  sulfur released by  the burning  of either residual fuel oil or noncondensable
gases. In a limited series  of tests, it was not possible to measure the presence of S02  in the
exhaust gases  of  a fluidized bed calciner, probably because the calcining and flash drying
provided a two-stage removal system.
                                         11-8

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 Galeano and Leopold (16) report that the lime kiln is the only major process source where
 significant quantities of  nitrogen oxides  can be measured.  The primary reasons for the
 presence of oxides of nitrogen are that there is sufficient excess air at a temperature of 1200
 to  1300° C (2200 to 2400° F) to promote the reactions.  The amounts of nitrogen oxides
 formed in the fluidized bed calciners are probably significantly less than from rotary kilns
 because of the lower operating temperatures of 825 to 875° C (1500 to 1600° F). To date,
 no  specific tests have been conducted to  determine the amount of oxides of nitrogen in
 fluidized bed calciner exhaust gases.

 11.5   Oxygen Addition

 Modecular oxygen can be  added to  the  combustion  air of a lime kiln to control  H2S
 generation from the lime mud  in the combustion zone.  The oxygen must be added together
 with the primary air to the firing zone at the dry end of the lime kiln. This practice will
 promote  effective mixing  with  the  combustion gases and will provide  for  complete
 oxidation  of any  H2S released from the mud. Precautions should be taken to assure that
 overheating of the kiln does not occur in localized areas. Such overheating could damage
 refractory  materials, interfere with kiln operation, or result in increased emissions of oxides
 of nitrogen.

 There is very limited field experience, to date, with the addition of oxygen to lime kilns for
 reducing H2S  emissions.  Singman (17) reports that oxygen addition  to  lime  kilns can
 substantially increase the lime mud throughput rates for previously  overloaded lime kilns
 without  excessive lime losses.  An addition of '0.454kg (1 Ib)  of oxygen results in a net
 decrease in lime makeup rate  of 1.8 kg (4 Ib) as CaO and,  consequently,  a considerable
 savings  in operating costs  for causticizing. Decreases in H2S  emissions may result for
 relatively short lime kilns, particularly where higher temperatures are maintained at the wet
 end. Additional process variables that would affect H2S emissions in addition to kiln length
 include kiln diameter, mud washing efficiency and  inlet sulfide level, mud firing rate and
 solids concentration, and gas velocities at different locations in the kiln.

 11.6   Process Economics

 The primary economic factor to consider for effective air pollution control of lime-calcining
 systems is the installation of  devices for particulate control.  The respective capital and
 operating costs for impingement and venturi scrubbing  devices are presented in Table 11-5
 (4).

Gaseous emission  control does not normally require substantial capital investment unless
flash drying of lime mud must be instituted. Maintenance of sufficient excess air, proper
washing of lime mud, and the use of fresh water normally are sufficient to minimize gaseous
                                         11-9

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                                   TABLE  11-5
              CAPITAL AND  OPERATING  COSTS  FOR LIME  KILN
                         PARTICULATE  SCRUBBERS (4)

                                                      Scrubber Type
                   Cost Item                    Impingement        Venturi

        Capital cost,* I/daily t pulp                  30-33             22-27
                      (I/daily ton pulp)           (27-30)           (20-25)
                           I/daily t lime            99-110            71-93
                       (I/daily ton lime)           (90-100)          (65-85)

        Annual operating cost,** $/t pulp             3-7               7-11
                           ($/ton pulp)            (3-6)             (6-10)
                                $/tlime            11-22             22-38
                            ($/tonlime)           (10-20)           (20-35)

         *Based on 1966 data.
        **Based on 0.9 cents/kWh.

emissions. Addition of NaOH  to the particulate scrubber makeup water to minimize H2S
emissions by increasing liquid pH levels will increase operating costs.

11.7  Lime Dust Handling

A minor source of fugitive particulate emissions from the causticizing system of a kraft pulp
mill consists of lime dust releases from storage tanks and bins, and conveying and transfer
facilities. Activities where lime dust is  loaded or unloaded, dumped, or transferred are
particular problems because of  the dryness of the material that is handled. The lime dust is a
localized emission source within the immediate area of the causticizing plant.

The three major approaches to control fugitive  lime dust emission are to:

     1.   Confine the potential emission sources to prevent air leakage,

     2.   Wet the dust to prevent its becoming airborne by wind or by transfer operations,
         and

     3.   Use special air pollution control equipment.
                                        11-10

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The  first  method is effective in limiting  the  potential sources or fugitive emissions  by
effective housekeeping. It  also facilitates the  subsequent installation of  particulate  air
pollution  control equipment. Wetting the dust is  effective  in  controlling fugitive dust
emissions, but  it can  make the  lime difficult to handle if overdone.  It generally  is not
recommended as an effective technique for controlling lime dust emissions.

The  third method for  controlling dust emissions from lime storage and transfer facilities
involves the  use  of particulate  control techniques, such as centrifugal separation,  liquid
scrubbing, and  fabric filtration.  Keeping the dust as dry as possible facilities its recovery;
therefore, liquid scrubbing is undesirable. Centrifugal collectors are not advantageous in that
they require high pressure drops and tend to have low collection efficiencies.

The  only known installation for particulate  lime  dust recovery  from storage facilities
employs a fabric filter baghouse for  a 907 metric ton per day (1000 ton/day)  kraft pulp
mill. The air  vents from the lime storage tanks are vented into a central duct and passed to a
baghouse with a design flow rate of 4,400 m3/h (2,600 cfm) at a maximum temperature of
290° C  (550° F). The  filter bags have a total  surface area of 121  m2  (1,300ft2) with a
cleaning cycle of once  each 20 minutes. The filter bags are made of a siliconized glass cloth
with a  design ratio of air  to cloth  filter area of 22.8 to 27.4 m3/h/m2  (1.25  to 1.50
cfm/ft2).

Total capital cost for  the system  was $5,900  in 1966. The fans have  a total capacity of
13.4 kW (18 hp) with a resultant direct  annual operating cost  of $1,080 per year. The
system can recover 225 to  450 kg/day of lime (500 to 1,000 Ib/day), which is equivalent to
an annual savings of $1,800 to $3,600, based on a lime price of $22/t ($20/ton) (18, 19).

11.8  References

 1.  Libby,  E. C. (ed.)., Pulp and Paper  Science and  Technology,  Volume I, Pulp. New
     York. McGraw-Hill Book Company, 1962. p. 211-227.

 2.  Atmospheric Emissions from the Pulp and Paper Manufacturing Industries. Cooperative
     NCASI-USEPA Study Project,  Publication No.  EPA-450/1-73-002,  United  States
     Environmental Protection Agency, Research Triangle Park, North Carolina, September
     1973.

 3.  Kramm, D. J., Selection and Use of the Rotary Lime Kiln and Its Auxiliaries—II. Paper
     Trade Journal, 156(35):25-31, August 21, 1972.
                                        11-11

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 4.  Taidor, C. E., Lime  Kilns  and Their Operation.  Proceedings of the International
    Conference on Atmospheric  Emissions from Sulfate Pulping. Hendrickson, E. R. (ed.).
    DeLand, Florida. E. 0. Painter Printing Company, April 28, 1966. p. 244-251.

 5.  Prakash, C. B., and Murray, F. E., Studies on H2S Emission during Calcining. Pulp and
    Paper Magazine of Canada, 74:99-102, May 1973.

 6.  Stuart,  H.  H., and Bailey, R. E., Performance  Study of a  Lime  Kiln and Scrubber
    Installation. Tappi, 48:104A-108A, May 1965.

 7.  Landry, J. E., and Longwell, D. H., Advances in Air Pollution Control in  the Pulp and
    Paper Industry. Tappi, 48:66A-70A, June 1965.

 8.  Erdman, A.,  Fluidized Bed  Burning of Kraft Mitt  Lime Mud.  Paper Trade Journal,
    154(36):40-42, September 7, 1970.

 9.  Van Donkelaar, A. J., Air Quality  Control in a Bleached Kraft Mill. Pulp and Paper
    Magazine of Canada, 69(18):69-73, September 20, 1968.

10.  How  a  Mead  Kraft Mill Operates  without Air  Environment Problems.  Paper Trade
    Journal, 158:26-29, April 8, 1974.

11.  Caron, A. L.,  Suggested Procedures for the Conduct  of Lime Kiln Studies to Define
    Minimum  Emissions  of Reduced  Sulfur  Through Control of Kiln  and  Scrubber
    Operating Variables. Special  Report No. 71-01, National Council of the Paper Industry
    for Air and Stream Improvement, Corvallis, Oregon, January 1971.

12.  Walther, J. E., and Amberg, H. R., Odor Control in the Kraft Industry.  Chemical
    Engineering Progress, 66:73-80, March 1970.

13.  Collins, T.  T.,  The  Oxidation of  Sulfate  Black  Liquor. Paper  Trade Journal,
    130(3):37-40, January 19, 1950.

14.  Personal Communication with Dr. Hal B. H. Cooper,  Texas  A&M University,  College
    Station, Texas, June 1974.

15.  Blosser, R. O., and Cooper, H. B. H., Current Practices in Thermal Oxidation of
    Noncondensable  Gases in the Kraft Industry. Atmospheric Pollution Technical Bulletin
    No. 34, National Council of  the Paper Industry for Air and Stream Improvement, New
    York, New York, November 1967.
                                       11-12

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16.  Galeano, S. F., and Leopold, K. M., A Survey of Emissions of Nitrogen Oxides in the
    Pulp Mill. Tappi, 56(3):74-76, March 1973.

17.  Personal Communication with Mr. Thomas L. Singman, Union Carbide Corporation,
    Tarrytown, New York, August 1973.

18.  Personal communication with Mr. James B. Ellis, Fibreboard Corporation, Antioch,
    California, September 1973.

19.  Personal communication with Mr. Andrew F. Reese, Fibreboard Corporation, Antioch,
    California, March 1970.
                                       11-13

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                                    CHAPTER 12

                            SMELT  DISSOLVING  TANK
The smelt dissolving tank is a large vessel located below the recovery furnace. A molten
mixture,  primarily  of sodium sulfide  and sodium carbonate  (smelt), is  continuously
removed from the floor of the recovery furnace. The smelt is mixed with water in the smelt
tank to produce green liquor. The smelt tank is an open, agitated vessel, covered by a hood
from which large volumes of steam are emitted when the molten smelt and water mix. The
smelt tank, along with the recovery boiler and lime kiln, is one of the main particulate
matter sources in the Kraft pulp mill.  In addition, the smelt tank can be a source of TRS
emissions.

12.1   Smelt Dissolving Tank Particulate Matter Emissions

Particulate matter consisting of both dissolved and undissolved NaOH, Na2C03, and Na2S is
emitted from the  smelt tank with the  rising flow of gases. Table 1-6 indicates that typical
smelt  dissolving  tank  particulate  emissions are 0.01-0.5 kg  per  metric ton of pulp
(0.02-1.0 Ib/ton)  following control devices. The majority  of  the smelt tanks that  are
controlled  use simple mist eliminator pads to filter the particulate matter from the escaping
vent gases (1).

The mist  eliminator pads consist of fine wire mesh screens, approximately 30 cm (1 ft)
thick.  Droplets condense from the gas  on the  wire  mesh and  are  washed back into  the
dissolving tank by water sprays. As may be seen from Table  12-1, typical pad efficiency for
particulate matter removal is about 70-90%.

A higher collection efficiency can be achieved by following the mist eliminator with a spray
or  packed  tower scrubber.  Alternatively,  some  mills  have low  pressure drop  venturi
scrubbers   15-20 cm (6-8 in)  of water,  cyclone spray scrubbers or  packed towers  for
particulate  control  without  mist  eliminator pads.  One such  packed  tower  gave 98%
collection efficiency (Table 12-1).

It is possible to combine the vent gases from the smelt tank  with the main flue gases from
the recovery boiler prior  to  a recovery boiler particulate collection device. One expected
difficulty  with this approach would be the effect of the water vapor content of the smelt
tank yejit  gases on recovery boiler particulate matter collection efficiency in an electrostatic
precipitator. A second potential problem would be the likelihood of H2S formation when
the Na2 S entrained in the smelt tank vent gases come into contact with the recovery boiler
C02.
                                        12-1

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                                    TABLE  12-1
               SMELT DISSOLVING TANK  PARTICIPATE. MATTER
                              CONTROL DEVICES (1)

                                    Collection Control
       Control Device                   Efficiency               Emission Rate
                                          Percent              kg/t     (Ib/ton)

     Pad entrainment                        71.8               0.026    (0.052)
     Separator                              77.2               0.075    (0.15)
                                           77.8               0.32     (0.63)
                                           90.2               1.2       (2.3)
                                           93.4               0.6       (1.2)
                                           70.8               0.79     (1.58)
     Pad plus shower scrubber                96.2               0.21     (0.41)
     Pad plus packed scrubber                91.9               0.60     (1.20)
     Packed scrubber                        98.4               0.025    (0.05)
Another  combined treatment method for smelt tank vent and recovery boiler gases would
be an electrostatic precipitator and scrubber combination. As discussed in Section 10.8.1, a
tail end scrubber can offer significant benefits from a heat recovery standpoint. In cases
where "snowing" from the precipitator occurs, a tail end scrubber can also offer significant
particulate matter collection advantages. Introduction of the smelt dissolving tank vent gases
after the precipitator and prior to the scrubber is possible, with additional heat recovery
benefits.

Particulate matter control, as such, has not yet been required for smelt tanks in Scandinavia,
although some mills use indirect condensers to recover heat from the smelt tank vent gases.
The  condensate  is returned to the smelt tank and the warm water produced is used for
washing.  The feasibility  of designing for heat recovery  in combination with  particulate
control is-enhanced by increasing fuel costs.

12.2  Smelt Dissolving Tank TRS Emissions

The  presence of  some  reduced  sulfur compounds  (Na2S)  in  the  smelt proper, and
occasionally some reduced sulfur gases from the recovery furnace, can cause TRS emissions
from the smelt tank vent. The amount of such gases  is highly variable, reported to range
from the equivalent of 0 to 1.85 kg H2S per metric ton of pulp (0-3.7 Ib/ton)  (1). Variables
that effect the  TRS emission rate  are the sulfide content of the particulate  matter in the
vent gases, the turbulence in the dissolving tank, the type of solution used in a scrubber, if
present, and the pH of the scrubber liquor (2). The  effect of some of these variables is
                                        12-2

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illustrated in Table 12-2. This table, prepared from a NCASI special study in 1970-1971 (2),
indicates that the most effective TRS control was wet scrubbing with fresh water.
                                  TABLE 12.2
             TRS EMISSIONS FROM SMELT DISSOLVING TANKS (2)
                   TRS
Mill
II
III


IV



V



VI


VII

VIII
IX
X


XI
XII

XVII
kg/t
0.005
0.06
0.005
0.005
0.02
0.02
0.04
0.055
0.005
0.01
0.0005
0.0005
0.005
0.005
0.0005
0.01
0.015
0.005
0.0005
0.01
0.005
0.0005
0.0005
0.005
0.005
0.005
(Ib/ton)
0.01
0.12
0.01
0.01
0.04
0.04
0.08
0.11
0.01
0.02
0.001
0.001
0.01
0.01
0.001
0.02
0.3
0.01
0.001
0.02
0.01
0.01
0.001
0.01
0.01
0.01
Control Device
None
Packed Tower
Packed Tower
Spray
Showers
Showers
Demister
None
Demister
None
Demister
None
Demister
None
Demister
Showers
Showers
Demister
None
Demister
Demister
Demister
Packed Tower
Demister
Demister
Showers
                                                            Scrubbing Solution
                                                         Weak Wash and Contami-
                                                           nated
                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water

                                                         Fresh Water

                                                         Fresh Water

                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Contaminated Condensate

                                                         Fresh Water
                                                         Fresh Water
                                                         Fresh Water
                                                         Weak Wash
                                                         Weak Wash and Contami-
                                                           nated Condensate
                                                         Weak Wash and Contami-
                                                           nated Condensate
                                                         Fresh Water
                                      12-3

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Weak wash from  the lime  mud clarifier,  lime mud washing  filtrate and  evaporator
condensates are used as well  as fresh water for scrubber liquor. Increased organic sulfides
were found when evaporater  condensates were used. A small amount of hydrogen sulfide
was liberated from lime mud clarifier supernatant in some systems, probably because of
acidification of the scrubbing liquor (3).

Recently, addition of caustic to smelt tank scrubbers has been suggested for improved TRS
control (4). In one mill, caustic will be added to clean evaporator condensates to be used as
smelt tank scrubber liquor. The scrubbing liquor will then be used for lime mud washing (5).

12.3  References

1.    Atmospheric Emissions  from  the  Pulp  and Paper Manufacturing Industry. EPA
     450/1-73-002. September 1973. (Also published as NCASI Technical BuUetin No. 69,
     February 1974.)

2.    Factor Affecting Emission of Odorous Reduced Sulfur Compounds from Miscellaneous
     Kraft Process Sources. NCASI Technical Bulletin No. 60, March 1972.

3.    Blosser, R.  0. Miscellaneous Sources and Trends  in Kraft Emission Control: Tappi,
     55:1189-91, 1972.

4.    Anon.  How a Mead  Kraft Mill Operates  Without Air Environment Problems.  Paper
     Trade Journal, 158:26-29, April 8, 1974.

5.    Testimony of W. A.  Wrase, S. D. Warren Company before  the Air Pollution Control
     Commission, Department of Natural Resources, State of Michigan. May 17, 1974.
                                       12-4

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                                   CHAPTER  13

           EMISSIONS OF OXIDES OF NITROGEN, HYDROCARBONS,
                              AND WATER  VAPOR
Additional pollutants from the kraft pulp mill include oxides of nitrogen and hydrocarbons.
Oxides of nitrogen can be emitted from combustion sources, such as the recovery furnace,
the lime kiln of the chemical recovery system, and the power boilers. Hydrocarbons and
other organic nonsulfur compounds can be emitted in varying quantities from the digester,
washers, evaporators, and direct contact evaporators. Both types of pollutants may have
potential significance to photochemical air pollution when acting together. Water vapor is
also  emitted in varying quantities from all kraft pulp mill sources. Condensation of water
vapor into a  visible plume  may present some hazard  if the plume  restricts visibility
adversely, such as across a highway, airfield, or harbor.

13.1  Nitrogen Oxides

Nitrogen oxides are formed by the reaction of atmospheric nitrogen and oxygen at elevated
temperature when  fuels are burned. Nitrogen oxides also  can form from the oxidation of
nitrogen which is  present as a trace  constituent  in fuels. The major oxides of nitrogen
formed  during  combustion processes are nitric oxide (NO) and nitrogen dioxide (NO2).
Their chemistry of formation is as follows:
                                  N2 +02

                                 2NO + 02 -» 2N02
The nitrogen oxides present in exhaust gases from combustion processes normally are 90 to
95 percent by volume NO, and 5 to 10 percent by volume NO2 .

The primary variables affecting the rate and degree of formation of nitrogen oxides during
combustion  processes are the flame temperature and the oxygen content of the gas in the
flame zone.  The degree of nitrogen oxide formation tends to increase exponentially with
temperature above about 1300° C (2400° F), particularly when the oxygen concentration in
the combustion zone is two percent by volume or greater (1). The rate of NO formation
increases sharply with temperature, as shown in Table 13-1 (2).

The  major variables that can  influence the nitrogen oxide  formation  during thermal
oxidation processes include the flame temperature, oxygen content in the flame zone, fuel
nitrogen content,  and combustion unit configuration.  Flame temperature is increased by
                                       13-1

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                                   TABLE 13-1
             EFFECT  OF  FLAME  TEMPERATURE ON NITRIC  OXIDE
           EQUILIBRIUM  CONCENTRATION  AND REACTION TIME  (2)

               Gas*
           Temperature            NO Concentration          Reaction Time
           °C      (°F)            ppm, by volume                sec

           2430  (4400)                19,000                  0.004
           1930  (3500)                14,000                  0.090
           1760  (3200)                4,000                  0.7
           1430  (2600)                   500                 21.0
           1090  (2000)                   10                162.0

           *Reactant gases: 77% N2, 15% 02, and 8% inerts.
increased fuel heating value, but decreases with increasing fuel moisture content. The heat
release rate and  combustion  volume  available  for  heat release also  influence  flame
temperature. Both black liquor and lime mud tend to have moisture contents of 30 to 40
percent by weight, which act to inhibit increases in flame temperature because the water
acts as a heat sink. The fuel nitrogen content of both of these fuels is relatively low, 0.1 to
0.5 percent by weight for black liquor and negligible for lime mud.

The fuel nitrogen  content is the primary variable limiting NO emissions at temperature
below 1100° C (2000° F);  while flame temperature becomes dominant above 1300° C
(2400° F) (3).

Galeano and Leopold  (4) report that nitrogen oxide concentrations from kraft pulp mill
sources are relatively low when compared to power boilers and are higher from lime kilns
than from recovery furnaces, as listed in Table 13-2.

Nitrogen  oxide emissions in  both cases can be minimized by operating combustion units at
minimum  excess air, minimum flame temperatures, and maximum fuel moisture contents.

    13.1.1   Kraft Recovery Furnaces

The configuration of the combustion unit is an important factor in determining potential
nitrogen oxide emissions from kraft pulp mill combustion sources. Kraft recovery furnaces
have relatively large volume rectangular combustion chambers and relatively low heat release
rates of  170 to 340 MJ/h per m3 (4,500 to 9,000 BTU/hr per ft3), which act to inhibit high
flame temperatures. The fuel has a relatively low heating value of 14 to 16 MJ per kg dry
solids (6000-7000 BTU/lb). Black liquor has a high moisture content of 30 to 40 percent by

                                       13-2

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                                   TABLE 13-2
      NITROGEN OXIDE EMISSIONS FROM  KRAFT PULP MILL  PROCESS
                                  SOURCES (4)
  Process Unit
Recovery furnace
Lime kiln
                                         NOX Concentration
                                               Emission Rate
    Temperature
   °c          F
Average    Range
   ppm, by vol.
1010-1230 (1850-2250)     32       0-53
1260-1390(2300-2530)    200     113-260
Average      Range
 kg N0x/t (Ib/ton)
                     3(6)
                    18(36)
            0-5 (0-10)
          10-25 (20-50)
weight that inhibits flame temperature rise. Flame temperatures are about 980 to 1260° C
(1800 to 2300° F) for normal operation in a kraft recovery furnace.

In addition, the endothermic reduction of Na2S04 to Na2S in the smelt bed of the recovery
furnace also acts as a heat sink to inhibit excessive flame temperatures. Another built-in
control method is that the air flow is  split between  primary and  secondary zones, and
sometimes a tertiary zone. A reducing atmosphere above the smelt bed also acts  to remove
oxygen from the combustion zone to inhibit the formation of nitrogen oxides. Introducing
tangential air into the furnace can act to spread out the flame front and, in that way, also
inhibit the increase in gas temperature. As a result of inhibiting both flame temperature and
oxygen level in the combustion zone, the nitrogen oxide levels normally range from 0 to 50
ppm by volume.

     13.1.2  Lime Kilns

Lime kilns are used to burn a mud containing CaC03 to recover CaO by addition of natural
gas or fuel oil.  The fuel  and air are added  countercurrently with the mud at opposite ends of
a long cylindrical rotary kiln, resulting  in a maximum temperature of  1260 to 1430  C
(2300  to 2600° F) at the fuel addition end of the kiln and  a temperature of 200 to 315° C
(390 to 600° F) at the  lime mud addition end of the kiln. The lime mud contains 30 to 40
percent water  by weight.  Most of this water has been evaporated  by the time the mud
reaches the hot end  of the kiln and  does not exert as great a suppressing effect on flame
temperature as does the black liquor  in a recovery furnace. The firing zone is narrower and
is concentrated with less  lateral turbulence in the  lime kiln, producing higher  flame
temperatures. As a result, higher nitrogen oxide concentrations exist in flue gases from lime
kilns than from recovery furnaces in kraft pulp mills.

     13.1.3  Power Boilers

Steam  is required in pulp mills for process and space heating, driving equipment, and
generating electricity. Although significant quantities of steam are  generated in recovery
                                        13-3

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furnaces, conventional industrial boilers supply much of the steam required by a pulp mill.
Power boilers are the largest sources of nitrogen  oxides in pulp mills. Table 13-3 lists some
values of nitrogen  oxide concentrations from kraft  and NSSC pulp mill power boilers.
Nitrogen oxide emission factors for auxiliary power boilers were presented in Table 1-8.
                                    TABLE  13-3
           NITROGEN OXIDE EMISSIONS  FROM  POWER BOILERS  (4)
                                     NO,
            Unit              Average    Range         Temperature       Excess Air
                                 ppm, by vol.          °C        fF)         %

Pulverized coal, front end,        375    310-445   1425-1480(2600-2700)
  86,000 kg/h (190,000 Ib/hr)
Pulverized coal & bark,           205    150-280   1200-1425 (2200-2600)
  32,000 kg/h (70,000 Ib/hr)
Gas fired, 45,000 kg/h            436    325-535                              33
  (100,000 Ib/hr)
Gas fired, 100,000 kg/h           190    161-232    870-980  (1600-1800)      46
  (220,000 Ib/hr)
Bark fired, 122,000 kg/h          123    101-145   1040-1140(1900-2080)      89
  (270,000 Ib/hr)
 13.2   Water Vapor

 Water vapor is emitted in varying quantities from all kraft pulp mill sources. Water vapor can
 be considered as an air pollutant under certain circumstances, such as when it acts to reduce
 visibility in highly humid  or cold atmospheres or  acts to modify climate or rainfall. Water
 vapor  emissions also  represent a potential  loss of heat energy that might otherwise be
 recovered by condensation to reduce overall plant energy consumption.

 Crabtree (5) reports that  the  potential loss  of water vapor to the atmosphere from kraft
 pulp mill operations is about 5 to 8 t of water per metric ton of air dried pulp produced (5
 to 8 ton/ton). The major potential sources of water vapor released to the atmosphere are the
 recovery furnace and the paper machines, as shown in Table 13-4 (5).

 The loss of water vapor from kraft pulp mill sources is affected by process operating factors,
 water  reuse and recycling practices,  and the relative efficiencies of heat recovery systems.
 The use of efficient  heat recovery  systems for  digester  blow  gases and multiple-effect
 evaporators tends to reduce the amount of water vapor released to the atmosphere. The use
                                        13-4

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                                    TABLE 13-4
    WATER VAPOR EMISSIONS  TO  THE ATMOSPHERE FROM  KRAFT PULP
                                MILL SOURCES (5)

      Source                  Moisture Content               Water Emission Rate
                                  %,byvol.                kgH20/t    (Ib H20/ton)

Batch digester                       30-99                 100-250      (200-500)
Multiple evaporator                  50-99                 500-1000   (1000-2000)
Recovery furnace                    25-35                2200-2500   (4400-5000)
Smelt tank                          35-45                 150-250      (300-500)
Limekiln                            25-35                 350-750      (700-1500)
Paper machines                       5-15                1700-2500   (3400-5000)
     Total                                                5000-7250   (10000-14500)

of indirect contact surface condensers also minimizes the amount of water used in  the
cooling processes. The  use of cooling towers  for heat dissipation tends to provide an
additional source of water vapor emissions to  the atmosphere  if water is recycled. Fuel
savings in reduced steam usage can range from $0.55 to $3.30 per metric ton of pulp ($0.50
to $3.00/ton) by effective use of heat transfer equipment and water reuse practices (6).

Evaporation of process  fuels to high solids concentrations prior to burning results in less
water vapor emission to the, atmosphere. Evaporation of black liquor to 65 to 70  percent
solids by forced circulation evaporation can improve the heat economy of the kraft recovery
system. Concentration of lime mud to 68 to 70 percent solids prior to firing in the lime kiln
reduces the water vapor release rate from the firing zone.  The kiln  exhaust gas from a
scrubber must be cooled to below 65° C (150°  F) to achieve an actual reduction in water
vapor emission from  the lime kiln that is sufficient to  compensate for possible water
evaporation in the liquid scrubber.

The  two major techniques  employed, to date,  for controlling local water vapor levels in
ambient air are condensation  to prevent  its release to the atmosphere and dispersion
resulting from high stack gas velocity or tall  stacks. Dispersion of moisture plumes by means
of elevated  discharges may be particularly necessary  for kraft pulp mills  located near
highways,  airports, harbors or populated areas  to alleviate potential fog. The presence of
large quantities of particulate matter  in ambient  air can act as condensation nuclei .to
accelerate the formation of fog and  inhibit its dispersal where an additional source of water
vapor already exists.

Shumas and Hansen (7)  describe a system where flue  gases from the recovery furnaces and
bark-fired power boilers at  a 1,135 metric ton per day (1250 ton/day) kraft pulp mill are
piped 0.8 km (0.5 mile)  to  a stack 62 m high (200 ft) and discharged 258 m (850 ft) above
                                        13-5

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the valley floor. The purpose of the stack is to discharge the moisture plume above the level
of the normal winter inversions in the narrow valley so as to alleviate any potential fog
formation problems, such as interference with aircraft landings and takeoffs. Plume rise
above the top of the stack ranges from 90 to 210 m (300 to 700 ft), depending on stack gas
flow rate and temperature and the season of the year.

Flue gases from the respective combustion units are passed through a venturi scrubber with
a flow rate of 510 m3/h  (2250 gpm) and are cooled to 50° C (125° F). The gas then flows
through a wood stove cylindrical duct 4.9 m (16 ft) in diameter at a flow rate of 620 m3/s
(1.3 million cfm) with a  fan discharge gauge pressure ranging from 6.3 to 31.8 cm of water
(2.5 to 12.5 in water).

The capital cost,  including  all engineering fees, for  the system is about $6,500,000 to
$7,000,000. The system  is powered by two parallel turbine drive booster fans  of 2050 kW
(2,750 hp) each. At an electric  power cost of $67/kW/year ($50/hp/year), the operating
costs are computed as $275,000 per year for the fans.

A system recently was placed in operation at a kraft pulp mill to alleviate moisture plume
problems from three parallel paper machine drier vents near a major interstate highway (8).
The exhaust gases total 230,000 m3/hr (135,000 cfm) with a temperature of 46° C (115° F)
and a moisture content of 10 to 15 percent by volume. The gases are collected and passed
through a central fan of  about 112 to 150 kW  (150 to 200 hp), then through a cylindrical
metal stack 25 m high (80 ft) with an exit diameter of approximately 2 m (6 ft), resulting in
an exit gas velocity of 22-30 m/s (75-100 fps). The capital cost for  the system is about
$80,000;  the  annual direct  operating cost approximately $10,000,  based on $67/kW/year
($50/hp/year).

13.3  Organic Compounds

Organic compounds other than those containing sulfur are also  emitted in varying quantities
from several types of processing units in kraft pulp mills. For  a discussion of nonsulfurous
organic compounds, see section 1.1.3.

 13.4  References

 1.  Smith, W. S., and Gruber, C. W., Atmospheric Emissions from Coal  Combustion: An
     Inventory Guide. U.S. Public Health Service Publication No. 999-AP-24, Cincinnati,
     Ohio, 1966.

 2.  Ermenc,  E.  D.,  Controlling  Nitric Oxides   Emissions.  Chemical  Engineering,
     77(12): 103-105, June 1,1970.
                                         13-6

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3.   Bartok, W., Crawford, A. R., and Skipp, A., Control of Nitrogen Oxide Emissions from
     Stationary Combustion Sources. Combustion, 42(4):37-40, October 1970.

4.   Galeano, S. F., and Leopold, K. M., A Survey of Emissions of Nitrogen Oxides in the
     Pulp Mill. Tappi, 56:74-75, March 1973.

5.   Crabtree, V. F., Abatement Procedures Presently in Use or Feasible: Other Operational
     Sources. In: Proceedings of the International Conference on Atmospheric Emissions
     from Sulfate Pulping, April 28, 1966,  Sanibel Island, Florida, Hendricksen, E. R. (ed.).
     Deland, Florida, E. 0. Painter Printing Co., 1966. p. 252-264.

6.   Tomlinson, G. H., Science of Wood Pulping. Tappi, 44:133A-142A, January 1961.

7.   Shumas, F. J., and  Hanson, G. A., A Unique Solution  to Punching  through  the
     Inversion Layer. (Presented at 1973 Tappi Environmental Conference. San Francisco.
     May 15, 1973).

8.   Personal communication with  Mr.  David E. Mansfield,  Western Kraft Corporation,
     1973.
                                       13-7

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                                    CHAPTER  14

              AIR POLLUTION  CONTROL IN  SULFITE PULP  MILLS
Even though many technological similarities exist between the sulfite and the sulfate or
kraft process, the air pollutants generated in each of these processes are quite different. The
sulfite process usually operates with acidic S02 solutions and, therefore, S02  is the main
gaseous air pollutant. In certain special cases of alkaline sulfite liquor burning  in recovery
boilers where chemical reduction may take place, H2S may also be emitted. Organic reduced
sulfur compounds are not produced.  Since the odor threshold is roughly one thousand times
higher for S02  than for reduced sulfur compounds, odors generated in the sulfite pulp mill
will  generally be much less  than in the kraft process. The sulfite process will also emit
particura^e matter from spent liquor burning.

Even within \he sulfite industry itself, there are many differences in S02 and particulate
emissions because of differences in cooking liquor bases, acidities, and recovery methods.

14.1   Sulfite Pulping Processes

     14.1.1   General

Sulfite  pulping  is  practiced with  numerous  modifications.  Cooking can be  done with
different bases, namely calcium (Ca), magnesium (Mg), ammonium (NH4), and sodium (Na).

Cooking can be done:

     1.   At different pH levels, from acidic to alkaline, subject to the solubility constraints
         imposed by the base used;

     2.   In several stages with changes in cooking liquor between;

     3.   With a high yield of pulp and with additional mechanical defibration (the neutral
         sulfite semi-chemical processes).

The  main sulfite  pulping processes,  pH range,  principal  cooking liquor composition, and
bases are presented in  Table 14-E  This table  also lists the  allowable pH range for the
different bases.

The potential for release of S02 into  the atmosphere increases with decreasing pH.
                                         14-1

-------
                                  TABLE 14-1
                      MAIN SULFITE  PULPING PROCESSES
 Cooking Method     pH Range
 Acid bisulfite
 Bisulfite
 Neutral sulfite
 Alkaline sulfite
 Two-stage
 Two-stage
 Three-stage
     1-2
     2-6
     6-9
    10-12
 6-8 and 1-4
  2 and 5-8
4 and 2 and 8
    Cooking Liquor

M(HS03)2,S02
M(HS03)2
MS03,MC03
MSO3,M(OH)2,MS
MS03,M(HS03)2,S02
M(HSO3)2,S02,MS03
M(HSO3)2,S02,MS03
 Base Alternatives

= Ca, Mg, (NH4)2,Na2
= Mg,(NH4)2,Na2
= (NH4)2,Na2
= Na2
= Na2
The cost of chemicals and the possibility of chemical recovery provide more information on
the emission potentials of the different sulfite processes. Based on different chemical prices,
combustion products and recoveries of chemicals, calcium-based processes have the highest
air pollution potential, ammonium-based processes next, and magnesium- and sodium-based
processes are the least polluting, see Table 14-2.
                                 TABLE 14-2
    SULFITE PROCESS CHEMICALS, PRICE, COMBUSTION  AND RECOVERY


Base
Ca
Mg
NH4
Na
Relative
Price
Per Mole
1/7
2/3
3/4
1

PH
Range
1-2
1-6
1-9
1-12
                              SSL* Combustion
                               Dust or Smelt

                             CaS04,CaO,CaC03
                                   MgO
                                   None
                               Na2S,Na2C03
                                    Gas
                                  Products

                                    S02
                                    S02
                                S02,N2,NOX
                                  S02,H2S
                                  Chemicals
                                   Feasibly
                                  Recovered

                                     None
                                  MgO + SO 2
                                     SO2
                                 Na2S,NaCO3
*5pent Sulfite Liquor.
 Many sulfite mills continue to dispose of their spent sulfite /iquor (SSL) to receiving waters.
 These are primarily calcium-based mills, but some magnesium-based mills and a number of
 NSSC mills using ammonium or sodium practice this disposal technique. These mills are
 usually  old  and though they.will have less  S02  emission than similar  mills  with  SSL
 recovery,  their water pollution  contribution through SSL dumping  is so great that major
                                      14-2

-------
mill modifications, or in some cases,  closures, may be  required to comply  with  water
pollution regulations.

    14.1.2  Atmospheric Emissions from Different Sulfite Processes

Typical  S02 and  dust emissions for different Scandinavian sulfite processes are given in
Table  14-3.  Examples of the typical  S02  and particulate emissions from United States
sulfite pulp mills are presented in Table  14-4.
                                  TABLE 14-3
             SCANDINAVIAN SULFITE PULP MILL  EMISSIONS  (1)

                                             Mill Number
Parameter
Base
Cooking pH
Cooking yield, %
SSL collection, %
SSL evaporation
SSL combustion
Chemicals recovery
Sulfur losses,
kgS02/t
Washing loss
Spills
Condensates
Gaseous loss
Flue gases
Flue gas dust
Sulfur make-up,
kgS02/t
Total S02 emission,
kgSO2/t
Flue gas dust,
kg dust/t
Dust collection, %
Dust emission,
kg dust/t
1_
Ca
1-2
50
0
No
No
No


210
0
0
10
0
0
220

10

0

0
0

2
Ca
1-2
50
70
Yes
Yes
No


56
20
26
10
83
25
220

93

90

80
18

_3_
Mg
1-2
55
85
Yes
Yes
Yes


20
10
15
10
15
0
70

25

70

99
1

4
Mg
3-5
55
85
Yes
Yes
Yes


40
10
110
2
25
0
87

27

100

99
1

5
Na
3-5
50
90
Yes
Yes
Yes


15
10
10
2
20
0
57

22

20

97
1

6
Na
3-5
75
0
No
No
No


100
0
0
2
0
0
100

2

0

0
0

7
Na
3-5
75
85
Yes
Yes
Yes


16
10
8
2
14
0
50

16

10

97
0

8^
NH4
3-5
75
0
No
No
No


100
0
0
2
0
0
100

2

0

0
0

9_
NH4
3-5
75
80
Yes
Yes
No


16
10
8
2
74
0
100

76

0

0
0

                                       14-3

-------
                                    TABLE 14-4
     TYPICAL POTENTIAL AMERICAN  SULFITE PULP MILL  EMISSIONS (2)
                                                     Mill Number
1
Ca
acid sulfite
45
No
No
No
75
—
—
—
2.5
0
2
Mg
acid sulfite
43
Yes
Yes
Yes
8.5
2
2.5
9.5
—
1
3
Mg
bisulfite
50
Yes
Yes
Yes
8.5
2
2.5
10**
—
1
4
NH4
acid sulfite
50
Yes
Yes
No
1.5*
2
5
200
0.25
1.5
           Parameter

Base
Type
Cooking yield, %
SSL evaporation
SSL combustion
Chemicals recovery
S02 emissions, kg SO2/t
  Digester blow
  Washers and screens
  Evaporators
  Boiler
  Acid towers
Particulate matter emissions, kg/t
 *Packed tower absorber with 91% S02 removal.
**Venturi absorber after recovery boiler.


In the following sections, the air pollutants generated in the basic sulfite process operations,
from  digestion to cooking acid preparation, are discussed, and the effects of different bases
and acidities are treated separately for each basic operation.

The general treatment  for SO2  emissions is scrubbing with alkaline solutions. The recovery
of S02 is a function of ptj and gas film resistance, and usually exceeds 90 percent.

Different dust separation methods can be used for collecting particulate matter. The three
most  widely used are cyclones, scrubbers and electrostatic precipitators.

14.2  Digester Gases

Sulfite cooking is performed both in batch digesters and in continuous digesters, the former
being more prevalent.  The continuous digester causes no air pollution  problem;  the batch
digester may  release substantial amounts of S02 depending on the blow  system and cooking
acid pH.
                                        14-4

-------
     14.2.1   Batch Digesters

Acid bisulfite cooking in batch digesters creates high SO2 pressures in the digester; bisulfite
and neutral sulfite cooking have correspondingly lower pressures. Digester relief and blow
cause S02 emissions.

S02  and cooking liquor are relieved from the digester to the acid preparation accumulator
system  in several  steps,  top  relief,  side  relief, high-pressure  blowdown, and,  finally,
low-pressure blowdown. All the S02 relief is recovered in the acid preparation system (section
14.7). For bisulfite and neutral sulfite cooking, the S02 pressure in the digester is lower and
a less elaborate relief system will return the S02  to acid preparation. The digester relief will,
therefore, usually not constitute an air pollution problem for any base or for any pH range.

The digester  can be emptied in two ways.  The usual North  American practice is to relieve
the digester down to a certain blow pressure, (.14-.28) MPa (20-40 psig), and them blow the
contents into the blow pit. This procedure is called hot  blow  (Figure 14-1).

The European practice is to relieve the digester down to atmospheric pressure and then flush
the contents  into the  blow pit by pumping spent liquor into the digester. This procedure is
called cold blow.

Hot  blow will release considerable amounts of S02 when the spent  liquor  is flashed,
especially for acid bisulfite cooking. The S02 emission  from the blow pit can  amount to
about 30 kg  S02  per  metric ton of pulp (60 Ib S02/ton) (2, 3). The blow gas is roughly 95
percent water vapor, 3 percent S02, and 2 percent C02 (2).

One  feasible  treatment for the SO2  emission  is to  scrub  the blow gas with an alkaline
solution of the base and return the solution to acid preparation.  S02 recovery efficiencies
are reported to be as  high as 97 percent for this type  of treatment (3). This method works
well with Na and NH4 bases. Magnesium and  calcium  require cumbersome slurry scrubber
systems.

Another method  is to scrub the blow gases with cold water to recover heat and S02; the
SO2  is  then  steam stripped from the water solution. This method is  reported as  being
insufficient for SO2 recovery (3).

Hot blow also  may be performed in ordinary blow tanks with appropriate heat recovery
systems. This is usually practiced with bisulfite  or neutral sulfite cooking, either of which
has much less potential for S02  release.

Blow pit emissions for bisulfite or neutral sulfite cooking are about 10 kg per metric ton of
pulp (20 Ib S02/ton) (2).  Cold blow will show considerably smaller S02 release than hot
                                        14-5

-------
                                                              LU
                                                              O
                                                              <

                                                              CO
                                                              CM

                                                              CO
                                                              O
                                                              m
                                                              o
                                                              LU
                                                              O
                                                           LU
                                                           DC

                                                           (D
LU
O
                                                              LU Q-
                                                              Q_ CO
                                                              CC
                                                              O
                                                              DC


                                                              Q


                                                              O
ADDITIONAL DIGESTERS  AND BLOW PITS
                                                              DC
                                                              CL
  TTT
                     14-6

-------
blow. With acid bisulfite cooking, the S02  emission from the blow pit with cold blow will
be 2-10 kg S02 per metric ton of pulp (4-20 Ib S02/ton).

Again, the most  practicable treatment is  scrubbing  with alkaline solution of the base.
Calcium and magnesium base are not generally suitable for this application.

With bisulfite or  neutral sulfite cooking, the cold blow pit S02 emission is less than 2 kg
SO2 per metric ton of pulp (4 Ib/ton), making scrubbing impractical.

     14.2.2   Continuous Digesters

The continuous digester is essentially  a completely closed system cooking with magnesium,
ammonium, or sodium  base in the pH  range 4-6. As with a kraft continuous digester, a
continuous diffusion washing stage is  integrated with  the digester.  These features make the
continuous sulfite digestion process a very  clean one,  with negligible emissions of S02. The
S02  containing gases from the  presteaming vessel relief and from the flash evaporators are
returned to the cooking acid preparation system.

14.3   Washer Gases

There  are many different sulfite pulp washing systems. For example, pulp may be washed
by displacement in the digester or in the blow pit. Additional washing in rotary drum filters
or in continuous diffusers is usually necessary (Figure  14-2)  because  of current  water
pollution regulations for SSL collection efficiency and the high cost of the sodium- and
magnesium-based  chemicals. Typically, the S02 emissions from washers and screens are 8 kg
S02  per metric ton of pulp (16 Ib S02/ton) for bisulfite cooking and even less for bisulfite
and neutral sulfite cooking. Because of the large gas flows involved in vacuum rotary drum
washing, S02  recovery through scrubbing is  not  practicable. For washing methods with
smaller vent gas flows, scrubbing may be used.

14.4   Evaporator Gases

The evaporation of SSL results in the release of S02  from the liquor. The more acidic the
cooking liquor, the more SO2 is liberated during SSL  evaporation.  The evaporation systems
in use are the  same  as those used  for kraft black liquor, and the multiple-effect vacuum
evaporation plant is the most commonly used for both. Each evaporation effect is vented to
the vacuum  system, and the vacuum system vent becomes the main emission point for
evaporator gases.  The hotwell is the main emission point for evaporator condensates. These
condensates contain S02 and can contribute to the evaporation plant S02 emission (Figure
14-3).
                                        14-7

-------
         WASH  WATER
                                                                                                    VENT
CO
                                         FILTER NO. I
                                          /jT.x
-------
         EFFECT NO. I
        (BODY AandB)
   NO. 2
(BODY AandB)
NO. 3
NO. 4
NO. 5
                                                                             SURFACE
                                                                             CONDENSER
 STEAM
                                                                                         VENT
                                          FIGURE 14-3

PRINCIPAL FLOW DIAGRAM FOR SPENT LIQUOR EVAPORATION IN MULTIPLE  EFFECT VACUUM PLANT,
                        5 EFFECT, 7 BODIES (SPARE BODIES NOT SHOWN)

-------
For bisulfite and neutral  sulfite cooking, the evaporator  gases contain small  amounts of
S02, usually less than 1 kg SO2  per metric ton of pulp (2 Ib SO2/ton). With acid bisulfite
cooking, the evaporator gases contain  20-30 kg SO2  per metric ton  of  pulp (40-60 Ib
SO2/ton)  (1).  This  significant  SO2  emission can be treated and  recovered by several
methods. The feasibility of each method depends on what other results can  be achieved
simultaneously by that particular method.

A  widely practiced  method  of eliminating evaporator  SO2 emissions is  to return the
evaporator gases to the acid preparation plant to recover the SO2.

Another method is to scrub the  gases with an alkaline solution of the base and then return
the solution to acid preparation. This procedure, however, requires a greater investment and
is difficult with calcium and, to a lesser degree, with a magnesium base.

If the sulfite mill has a weak liquor fermentation plant, either for CH3CH2OH or yeast, the
SSL must be neutralized prior to its use in fermentation. This can  be done by stripping it in
a separate   evaporation stage or  by  adding  base. After fermentation, the SSL can be
evaporated with insignificant SO2 emissions.

The weak SSL can be steam stripped in the evaporation  plant, as mentioned  before. The
S02 stripped off is returned to acid preparation, and subsequent evaporation of  the stripped
SSL will yield very little S02.

The weak SSL  can be neutralized by adding base, CaO, MgO, ammonia (NH3),  or NaOH, to
the liquor before evaporation. This practically eliminates  gaseous SO2 emissions from the
evaporation  plant.  SSL  neutralization,  however, is mainly proposed  to  reduce  the
evaporation condensate BOD by hydrolyzing acetic acid (CH3COOH) so that the resulting
nonvolatile  acetate will proceed with the liquor to combustion. Full-scale neutralization
trials  began in Sweden  in 1972, but, thus  far,  continuous  operation over long periods of
time has been difficult to maintain. Acid bisulfite cooking with magnesium base requires the
addition of about 35 kg MgO per metric ton of pulp (70 Ib MgO/ton) to the SSL before
evaporation to raise the pH to 6.5 (4).

14.5   Combustion Gases

All sulfite  process spent liquors can be burned. Calcium, magnesium and ammonium-base
SSL can be burned about any supporting combustion fuel if evaporated to 55 percent dry
solids,  atomized  thoroughly, and burned  with preheated  air in a  separate  combustion
chamber ahead of the main furnace. During combustion of  any SSL, emissions of S02 and
particulates of base oxides, carbonates,  and sulfates can occur, depending on the particular
prevailing equilibrium conditions.
                                        14-10

-------
Sodium-based SSL is subject to either oxidizing or reducing combustion. In the latter case
combustion is in a black liquor recovery boiler and H2S emissions are possible. The majority
of particulate matter emissions will be Na2S04. Most of the sodium and sulfur is recovered
in the smelt.

     14.5.1  Calcium SSL Combustion Products

Collecting calcium SSL with 70 percent efficiency and burning it produces about 80 kg S02
and 90 kg  dust per metric ton of pulp (160 Ib and 180 Ib/ton) (Table 14-3). The dust is
about  60 percent CaSO4  and 40 percent CaO, which includes some CaC03 and traces of
other salts.

The  flue  gases are  usually  passed  through  a  cyclone  dust  separation  system  with
approximately 80 percent collection efficiency. The dust is stored in a pile or dumped into
receiving water, because no economic use has been  found for it so far  (5). Better removal
can be achieved by electrostatic precipitators, but they  are too expensive  for recovering a
virtually  worthless dust.

An SSL  collection efficiency of 70 percent is very low, and water pollution  abatement
requirements may demand a smaller washing loss.  Dumping calcium sulfite  ash into the
water also may not be permitted.

SO2  can  be removed from the flue gases by scrubbing them with CaO or CaCO3  slurry. But
even if all the CaO of the dust from a 100 percent flue gas dust separation could be used for
this, the  SO2 emission would still be around 40 kg/t (80 Ib SO2/ton). A byproduct of this
control method is about 250 kg per metric ton of pulp of a 50 percent sludge (500 Ib of
sludge/ton), for which no use is known.

More  effective  S02  removal requires additional lime and  produces more sludge. At this
point,  changing the base might be a more feasible alternative.

     14.5.2  Magnesium SSL Combustion Products

Collecting magnesium SSL with 85 percent efficiency and burning it produces around 15 kg
SO2/t  plus 70 kg magnesium  oxide (MgO)/t for acid bisulfite cooking and 25 kg  S02/t plus
100 kg MgO/t for bisulfite cooking  (30 plus 140 Ib/ton and 50 plus 200 Ib/ton respectively).
These figures assume MgO recirculation (Table 14-3).

These  emissions, however, are  effectively treated  in a chemicals  recovery  system that
separates the MgO dust by cyclones and  scrubbers, hydrolyzes the  MgO  into magnesium
hydroxide slurry, and scrubs the S02 from the flue gases with this slurry. The net result is a
                                       14-11

-------
small dust emission and moderate  S02  emission,  which .can  be further  decreased  by
additional scrubbing stages as discussed in section 14.9.1.

     14.5.3   Ammonium SSL Combustion Products

Collecting ammonium  SSL  with 85 percent efficiency and  burning it produces  around
130kg  S02/t (260 Ib SO2/ton) for bisulfite cooking. During  combustion, the NH3  is
converted to water and nitrogen (N2),  and  consequently  the base is lost. Scrubbing SO2
from the flue gas with fresh ammonia solution has been successfully demonstrated in  at least
two  American ammonium-based mills   (6, 7). The nature of any air pollution  problems
related to the nitrogen compounds emitted from ammonium-based pulping is discussed in
section 14.10.

     14.5.4   Sodium SSL Combustion Products

Sodium SSL  burning may yield various combustion products depending primarily on the
method of burning.  Usually the SSL is burned in a reductive recovery boiler similar to a
black liquor recovery boiler. Most of the inorganic dry solids are then transformed into a
smelt of primarily  Na2S and Na2CO3. Part of the inorganic dry solids are entrained in the
flue gases as Na2S04. It, however, is efficiently recovered by electrostatic precipitators and
returned to the furnace to be reduced to Na2S. The dust emission, therefore, stays low and
usually it is about 1  kg Na2SO4/t (2 Ib  Na2S04/ton). The SO2 emission may vary widely
and is a function of the SSL sulfidity (5).

The  treatment of the emissions from sodium SSL combustion depends on  the particular
chemicals recovery system chosen. At present  there are  10  major sodium  SSL  chemical
recovery systems with additional "customized" variations (IPC, Mead, Stora, Sivola, Western
Precipitation, AST, SCA-Billerud, Tampella, ITT-Rayonier, and cross-recovery with  a kraft
mill). Both the CO2  and  SO2 in the flue gases may be  used, and depending on the system,
S02  emissions will  be about 6-20 kg S02/t (12-40 Ib S02/ton).

Combustion under reducing  conditions similar to the conditions in a  black liquor recovery
boiler, especially to those in a pyrolysis  reactor (AST and  SCA Billerud), will also produce
H2S.  (H2S generation  and abatement in a black liquor recovery boiler were discussed in
section 10.2.) H2S from a pyrolysis reactor is usually converted to sulfur in a Glaus  reactor
or oxidized to S02  in a separate furnace.  H2S emissions  are negligible.

14.6  Acid Preparation Gases

Preparation  of the cooking acid occurs in  two  steps, preparation  of the  raw  acid and
eventual fortification to cooking acid strength. With chemical recovery, raw acid preparation
is  simultaneously the recovery process. Acid preparation gases pose a minor  S02  emission
problem.

                                       14-12

-------
     14.6.1  Calcium Cooking Acid Preparation

With calcium base, S02 derived from burning of sulfur or pyrite is passed countercurrently
to water in  towers packed with limestone (Figure 14-4). The vent to the atmosphere will
emit some S02,  but this emission is rather  minor and usually amounts to less than 1 kg
S02/t  (21b  S02/ton). Typical values  are 0.2kg S02/t (0.4 Ib S02/ton) with a 20° C
(68° F) water temperature (1,8). Higher water temperatures might increase the emission of
S02 by decreasing absorption efficiency.

The raw acid is fortified to  cooking acid strength with digester relief and blowdown; the
vent from the acid fortification system is connected to the acid tower (Figure 14-5).

     14.6.2  Magnesium Cooking Acid  Preparation

Magnesium raw acid is prepared in the chemicals recovery system (Figure 14-6). The raw
acid may be  used directly for bisulfite cooking or fortified for acid bisulfite cooking (Figure
14-5).  In both cases,  the only emissions will  be with the flue gases from the chemicals
recovery system. About 15 to 25 kg S02/t (30 to  50 Ib  S02/ton) will be emitted. This
amount can be decreased by additional recovery stages.
                                 LIMESTONE
VENT
 WATER
                                                                   WATER
                                                                   RAW ACID
                                                                   VENTS FROM
                                                                   EVAPS & ACID
                                                                   FORTIFICATION
                                  FIGURE 14-4
         FLOW  SHEET  FOR  CALCIUM BASE  RAW  ACID PREPARATION
                                     14-13

-------
                                                        Vent To Acid Tower
                                                        Or Recovery System
                                                        Low-Pressure Slowdown

                                                        High-Pressure Slowdown
                                                        Cooking Acid


                                                        Side Relief Acid
                                                   •«    Raw Acid
                      FIGURE  14-5

ACID BISULFITE  FORTIFICATION SYSTEM FLOW SHEET

-------
       VENTS  FROM
        EVAPS.
         FLUE GAS
         1% S02
f"
K^
Ul
          RETURN
          Mg(OH)2
          SLURRY
                                                                                            WATER
                                                                                            MAKE-UP
                                                                                             S02
RAW ACID TO
FILTERS S
EVENTUAL
FORTIFICATION
                                                 FIGURE  14-6

                           FLOW SHEET FOR MAGNESIUM BASE RAW ACID PREPARATION

-------
     14.6.3   Ammonium Cooking Acid Preparation

Ammonium  cooking acid  preparation  systems are  similar to those  described  above.
Sulfur dioxide, either from sulfur burning or from flue gases, is absorbed in an NH3 solution.
Depending on pH, there is a potential for NH3 emissions. Performance runs  with spray
tower absorbers using liquid circulation  have shown  that when the pH dropped below 6,
almost 10 percent of the S02 feed was lost through the tower vent, while NH3  losses were
negligible. When the circulation pH  rose above  7.6, NH3  losses were substantial, but S02
losses went down to 0.3-0.6 percent. A pH of 7.1 kept both S02 and NH3 losses below 1
percent,  but  caused a large  plume from the absorption tower vent  (1). Fine pH control is
required  to maintain SO2 and NH3 emissions at acceptable  minimums.

     14.6.4   Sodium Cooking Acid Preparation

There are as many systems for sodium cooking acid preparation as there are  for sodium
recovery. These recovery systems are covered in section 14.7.

14.7  SSL Recovery Boilers

     14.7.1   Design Parameters

Spent sulfite liquor is, as a rule, incinerated to recover both heat and cooking chemicals in
reusable  form. Whenever chemicals are recovered, their recovery is  considered the primary
purpose  of the incineration  process, especially for magnesium- and  sodium-based processes
where the base is recovered. In  the  calcium-based  process, recovery of the  base is not
feasible,  and in the ammonium-based processes, only S02 can be recovered.

Since sodium bisulfite recovery was previously  discussed  under kraft processing, magnesia
(MgO) recovery is discussed in this section.

There are a variety of SSL combustion systems. To help  in evaluating different systems, a
short survey of combustion fundamentals and requirements will be presented.

Burning  intensity, which is the heat  input rate per  unit of furnace volume, depends mainly
on temperature.  The theoretical combustion temperature is a function of SSL heat value,
combustion air temperature, and fuel-air ratio. The actual combustion temperature depends,
in addition, on the heat and mass exchange  with  surrounding zones.

Stable  ignition  requires  transport  of  energy  to  the  ignition zone.  This is normally
accomplished by  recirculation  of hot furnace  gases or hot air. The recirculation  can be
external or internal.
                                         14 16

-------
Complete combustion is  essential for the recovery of  chemicals  and heat.  A necessary
condition for complete combustion is sufficient residence time in the high temperature zone
of the furnace. This requires sufficient furnace volume and controlled flow pattern.

Since magnesium  sulfate (MgS04)  cannot be economically converted  into  magnesium
bisulfite, a low sulfate content of the combustion product is important relative to recovery.
The  formation of  sulfate depends on factors,  such  as combustion excess air, combustion
temperature according  to the chemical  equilibrium, the  residence time  of the ash in
intermediate  temperature  zones,  and  different  catalysts.  Not  all  sulfate  formation
mechanisms are fully understood.

Additional design  parameters  are  set   by   requirements  of  boiler  availability  and
controllability. Therefore, the most important factors  become heat  surface fouling  and
partial load combustion control.

     14.7.2   SSL Recovery Boiler Systems

Burning  of SSL is usually carried out in power-type boilers  with  some  modifications.
Ash-free  auxiliary fuels can  normally be burned in the same furnace or boiler without great
disadvantage, if the auxiliary fuel input is not too high when compared to the SSL input.
Possible disadvantages of using auxiliary fuel are dilution of the  gases to the S02 recovery
system and an increased percentage of sulfate  due to temperature conditions and catalytic
conversion. SSL boilers commonly have a low furnace heat release rate, which requires a  low
furnace exit temperature, widely spaced heat exchange surfaces, and effective  heat surface
cleaning equipment. Similar features exist in successful kraft recovery boiler designs.

SSL  burning in a fluidized bed reactor was practiced, to some extent, for all soluble bases.
This system is normally supplied as an integrated recovery unit and not in combination with
a power boiler. Environmentally, it does not differ very much from a conventional recovery
unit. The actual burning  of SSL can be carried out in a variety of different burning systems.
The most important systems are:

     1.    Small precombustion chambers, called Loddby furnaces;

     2.    Large furnaces, called Lurgi-Lenzing-Steinmuller (LLS);

     3.    Small furnaces of Babcock & Wilcox (B&W) for liquor burning;

     4.    Furnaces, designed  by Tampella, that  incorporate features  of  both LLS  and
          Loddby furnaces; and

     5.    Fluidized bed reactors by Copeland and Dorr-Oliver.
                                        14-17

-------
Loddby furnaces are small precombustion chambers that are usually horizontal and  are
attached to a main boiler furnace. The refractory walls of the cylindrical chamber heat  the
tangentially-introduced  combustion  air  to  a  high  temperature.  External  recirculation
stabilizes the flame in the muffle. The burnout of the flame occurs in the main furnace. The
flame  temperature   at  the muffle  exit  is  approximately  175° C (350° F) below  the
theoretical combustion temperature.  The dust from a Loddby furnace usually has a smaller
mean particle size than dust from most other combustion systems. Loddby furnaces have
good  controllability  and, if a few of these chambers are used jointly, they are suitable  for
partial load operations. In some Scandinavian  mills, SSL is considered a good pressure
control fuel for recovery-power boilers.

Lurgi-Lenzing-Steinmuller  (LLS) furnaces are large units  that are integrated  with a steam
boiler  main  furnace, but are separated by a verticle membrane tube wall that has a screen
passage in the lowest part.  Excess cooking of the furnace is prevented by covering the tube
walls  with a thin refractory  layer.  The LLS furnace exit temperature is  about 305° C
(550° F) below the  theoretical combustion temperature.  Effective  ignition  and mixing is
accomplished by spraying the SSL countercurrently to the air flow.

Babcock & Wilcox furnaces for liquor burning are uncooled or slightly cooled. They can be
similar to the LLS furnace. Unlike other systems, B&W uses self-stabilizing liquor burners
that have  their own air registers. Auxiliary fuel is burned in the same or in different burners.
The controllability of a B&W system is good.

Tampella furnaces are combinations  of LLS furnaces and  Loddby furnaces. Only Ca-based
liquor has been fired thus far.

Fluidized  bed  reactors, by Copeland  and Dorr-Oliver, are used to some extent for SSL
treatment. They provide  an  environment  for  combustion that  gives a high intensity of
reaction at comparatively low temperatures.

     14.7.3   Operation Parameters

Because the recovery boiler is normally followed by chemical recovery, boiler  operation has
a greater direct influence on the chemical cycle than on the gas and solids emission from the
plant.

Excess air and the use of auxiliary fuel, for an existing installation, are the most important
operating parameters. The firing rate is  also important, although large variations are  not
normally expected.

For calcium-base SSL without recovery, the excess  air  has  little or no influence  on the
amount of S02 or S03 emitted from the boiler. High excess air can somewhat improve the
                                        14-18

-------
formation  of  sulfate and, thus,  reduce the amount of SO2 and S03. Even a complete
sulfatizing  of the base will bind only about 65 percent of the sulfur. A normal value is about
35 percent. Using too much excess air can make heat recovery uneconomical.

The  dust content of the flue gases from calcium-base  SSL  firing is often approximately
11.5g/m3  (5 gr/cu ft)  when  continuous  sootblowing is  practiced.  With  intermittent
sootblowing,  the  concentration  can  be  as low as 8-9.2 g/m3  (3.5-4 gr/cu ft); during
sootblowing, it is correspondingly higher.

The  use  of auxiliary fuel dilutes the flue gases to lower concentrations, but increases the
costs of  dust  separation and eventual S02 removal. The use of  wood refuse as fuel in the
same boiler improves the binding of S02 and S03 to the ash.

Variation of excess air to Mg-base SSL liquor firing changes the sulfate content of the ash.
Low excess air should be used to avoid sulfate formation so  that all sulfur is converted to
S02  in the flue gas for more efficient recovery. In practice, a part of the  Mg, usually less
than 10  percent, leaves the boiler as sulfate with a correspondingly slight decrease in the
concentration of S02.

Only natural gas and oil can be used as auxiliary fuels. Even a high sulfur heating oil dilutes
the flue gas S02 concentration, and no appreciable extra SO2 absorption is expected from
oil use. An excessive use of auxiliary  fuel may increase the sulfatizing of the base  or the
formation of weakly soluble combinations.

Soot formation  must be  avoided, especially when burning auxilliary fuels.  The soot will
discolor the MgO ash and may also discolor the cooking acid made from the  ash. Ultimately,
the soot may discolor the pulp.

The  operating parameters seem  to  have  a  great  influence  on  the results of firing
ammonia-based SSL. Different problems were reported, such as pluggage of boilers, boiler
corrosion, and the emission of smoke,  often referred to as blue haze. According to one mill
plant, by installing and  using glass fiber-bed demisters on an S02  absorption system (6), the
blue  haze is eliminated.

In tests by EKONO, the pH of the ash correlates strongly  with the excess air of  liquor
burning.  This  correlation is probably caused by formation of S03 during the combustion.
By combining ammonia-base  SSL and bark-firing, an alkaline ash reaction with a low excess
air is possible. Where sulfur recovery  is not feasible, an alkaline ash is the only way to
prevent S03 formation and subsequent direct discharges  of it into the atmosphere.
                                        14-19

-------
14.8  SSL Recovery Systems

The  combustion  gases  for  bases  that  are feasibly  recoverable are usually treated in a
chemical recovery system. The efficiency of the recovery process is also an indication of the
amount emissions. Because of high costs, SO2 recovery has been limited to 50-80 percent
and dust separation to 95-99 percent, both depending on the base and the recovery system.
These  recovery  efficiencies, however,  can be  increased with higher  investment.  Dust
emissions below  1 kg/t and  S02  emissions  below  10kg  S02/t  (2 Ib  and 20 Ib/ton,
respectively) have been  adopted as emission regulations for new sulfite mills (9, 10). These
regulations will effectively phase  out calcium-based sulfite mills  in the future and will
require more efficient recovery systems for the other bases.

All sodium recovery systems that displace H2S from dissolved sulfide solutions with flue gas
C02  or sodium bicarbonate  (NaHC03) will have H2S as an intermediate gas in the system.
The  H2S is usually converted to sulfur  in either a Glaus reactor or oxidized to SO2 in a
furnace with almost 100 percent efficiency in both cases. Under normal circumstances, H2S
is not emitted into the atmosphere.

The  recovery  of soluble base chemicals and S02  is carried out using a few well known
chemical reactions.  Nevertheless, the number of systems offered is large. Also some plants
use their own methods, which have not yet been described in the literature.

     14.8.1   Magnesium Base

The basic reaction in all  systems is an absorption of flue gas S02 into Mg(OH)2 to produce a
bisulfite solution.

In the B&W process (see Figure 14-7), the recovered MgO together with the make-up MgO is
slaked  with  steam addition.  The  hydroxide  is  circulated   in  absorption  Venturis  in
countercurrent flow.  The  hydroxide reacts with recirculated Mg(HS03)2 to form  MgSO3.
In the recirculating  acid, the sulfite reacts with SO2 to form new Mg(HSO3 )2. The extent of
SO2  absorption in each stage is governed by acid concentration, SO2  concentration in the
gas, temperature, recirculation rate, and gas velocity. The use of venturi absorbers has made
much  higher  monosulfite  concentration possible than  produced  with packed towers.
Combustion gases discharged to the recovery system stack can be expected to contain  less
than 250 ppm S02 by volume.

The  system as practiced in  Veitsiluoto,  Finland, has no dust  separation  and no separate
slaking  system. The  initial  circulation  solution is formed in the  first  tower,  somewhat
simplifying the  system. The design of the  dust absorption system  depends  on close
operational tolerances, but seems to work satisfactorily.
                                        14-20

-------
                                                       FLUE GAS
FLUE GAS
I
STEAM
SULFUR
SULFU
BURNE

ff£ '
R
R

	 T
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RAW ACID
i WARM WATER

                                   FIGURE 14-7

   BABCOCK  & WILCOX PROCESS FOR MAGNESIA  BASE SULFITE CHEMICALS RECOVERY

-------
Systems with only one absorption unit have been offered by several manufacturers, among
them Ahlstrom, Svenska Flaktfabriken, and Tampella. The reactors used are more complex,
but can be considered as multistage reactors. Marbles have been introduced to increase the
reaction surface, without significant effect.

In most  cases, some heat is recovered from  the gases  in a  cooling venturi before the
absorption units.

The recovery plant stack is practically the only emission point. Excessive emissions of S02
usually occur  only when  the pH control of the absorption units does not work properly.
High MgO emissions are normally the result of separator malfunction at plant overload.

     14.8.2   Sodium Base

All recovery methods produce  a  solution of Na2SO3 or Na2CO3, or both,  with the least
possible concentration of undesired sulfur combinations.  To obtain this result, the sulfur
combinations are removed by carbonization with C02. The carbonate or bicarbonate is then
converted to sulfite by sulfitizing with S02. In some processes the sulfides and bisulfides are
removed by crystallization of the carbonate. A few modern recovery systems are:

     1.    The STORA process,

     2.    The SCA-Billerud process,

     3.    The Tampella process,

     4.    The Sivola-Lurgi process, and

     5.    The Institute of Paper Chemistry method.

In the STORA process, the clarified green liquor from the recovery boiler, containing Na2S
and Na2CO3, is carbonated with recirculated CO2. Na2S is converted to H2S. The H2S is
stripped from the  liquor by CO2  as the carrier  gas. The H2S reacts  in a Glaus reactor with
S02 to form  elemental sulfur (S). The CO2 is  recirculated to  the process. In this way  the
green liquor  is converted to  NaHC03.  The NaHC03  is then  reacted  with NaHS03  to
produce Na2S03. Part of the  sulfite solution  absorbs SO2  in an absorption tower and is
returned as bisulfite.

Depending on process requirements, the cooking liquor is made from suitable  proportions of
the sulfite and bisulfite solutions. When a higher pH is needed as for semichemical pulping,
some bicarbonate  is bypassed to  the sulfite solution.  The principal process components of
the STORA process are illustrated in Figure 14-8.
                                        14-22

-------
f"
to
CO
                         S02  ABSORPTION

                         FROM FLUE GAS
STRIPPING OF H2S AND

PRODUCTION OF Na2S03
             WATER
             FLUE GAS
             GREEN  fr
             LIQUOR
             DREG£
PRODUCTION OF ELE-
MENTAL SULFUR
(CLAUS REACTOR)
ABSORPTION  AND
STRIPPING OF S02
                                                                                           ELEMENTAL
                                                                                           SULFUR
.
I 	 £


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-------
In the  SCA-Billerud  process, evaporated spent liquor is sprayed  into a  stream  of hot
combustion gases containing a small excess of air, but still insufficient for the combustion of
the organic substances in the  liquor.  A pyrolysis of the liquors takes place, resulting in a
combustible gas and a powder. The gas contains almost all the sulfur as H2S, as well as other
combustible components, including H2, CO, and some hydrocarbons. The powder contains
all the sodium, mostly as carbonate with a minor  part as sulfate,  as well  as carbon. The
process is illustrated in Figure 14-9.

The gas and powder mixture is cooled in a reactor  boiler to a temperature  well above the
dew point of the gas, which produces high pressure steam. Separation of the powder from
the gas occurs in both dry and wet separators. The gas is further cooled by condensing water
vapor in a scrubbing tower, which enriches the gas before burning. The H2S in the pyrolysis
gases is  converted to S02 in a boiler. The SO2 in the flue gases is then absorbed by a soda
lye solution before being  exhausted.

The powder separated from the gas is mixed with water, and the soluble salts are leached
out to produce the Na2CO3 solution used in the absorption just described. The remaining
carbon can be burned separately.

To produce cooking  liquor,  the product from the S02 absorption is fortified with makeup
sodium  and sulfur in the acid making plant.

Like the STORA process, the Tampella process  (Figure 14-10) is very flexible and  can be
applied  to various cooking methods. The essential part of the Tampella process consists of
three  subprocesses: precarbonation, bicarbonation-crystallization evaporation, and prepara-
tion of bicarbonate by carbonation. The starting point is combustion of the black liquor in a
reducing atmosphere. In  the Tampella recovery  process, a smelt consisting mostly of Na2S
and Na2C03 is formed in a conventional way.

The  smelt is dissolved  in  a strong green liquor,  which,  depending on the need, can be
converted to H2S, a  high sulfidity smelt solution, a low sulfidity smelt solution, an Na2CO3
solution, or monohydrate crystals free from sulfide and thiosulfate. The H2S can be burned
to SO2 for the preparation of Na2SO3  solution, or converted, by means of  the Glaus
Process, to elementary S for  the production of polysulfide liquor (11).

To prepare polysulfide  liquor,  a high sulfidity  solution (88-92  percent) can be used by
combining evaporation  and mechanical separation of carbonate crystals.  Low sulfidity
carbonate solution can be used for preparing NSSC liquor  or kraft white liquor. Crystal
Na2C03 can be employed for the neutralization of sulfite and  black liquor and also for
preparing low-yield Na2 S03  cooking liquor.
                                        14-24

-------
to
01
                  PRODUCTION OF
                      H2S
DUST   SEPARATION
(DRY)    (WET)
                                                                                                 ABSORPTION OF
           OIL
           SPENT
           LIQUOR
                   RECIRCULATION OF
                   CARBON TO EVAPORATION
                   PLANT
STEAM
                                                          LEACHING OF
                                                          CHEMICALS
                                                                                                           FORTIFICATION
                                                          FIGURE  14-9

                              SCA  PROCESS  FOR  SODIUM BASE SULFITE  CHEMICALS  RECOVERY

-------
to
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                   S02 ABSORPTION FROM
                        FLUE GAS
           WATER
           FLUE GAS
           GREEN
           LIQUOR
           DREGS
                       r
                           t
PRODUCTION OF
     NoHS
                                                          1
STRIPPING OF
    H2S
PRODUCTION OF
    NaHCOs
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                                                                                               FORTIFICATION
                                                           FIGURE  14-10

                           TAMPELLA PROCESS FOR  SODIUM BASE SULFITE CHEMICALS RECOVERY

-------
In the Sivola-Lurgi process, green liquor, containing mainly Na2S and Na2CO3, reacts first
in the precarbonation tower with the flue gases from the recovery boiler and then with C02
in the main carbonation stage. The purpose is to  convert  the active sodium of the green
liquor to NaHC03  and to liberate  the  sulfide as H2S in the  carbonation  stage. The
bicarbonate  obtained is then split thermally in a decomposer, producing Na2C03 and C02
for carbonation.

The  major  portion  of  the carbonate reacts  with NaHS03 in a reactor  and  produces
Na2SO3  and C02. The  reactor and  the  decomposer are built one above the  other and
connected with a clock bottom. Portions of the Na2S03 obtained are used for cooking
liquor preparation  and for sulfitation. In sulfitation step  Na2SO3  reacts with S02  to
produce bisulfite. The S02  is obtained by burning  elementary sulfur and H2S in the same
furnace.

The  method of the Institute of Paper Chemistry is probably the simplest. On the other
hand, the products of the reaction can be controlled only within certain limits.  The green
liquor is piped after clarification to the sulfitation tower, where it is treated with S02 from
a sulfur furnace. The formation of thiosulfate cannot be prevented, but it can be kept at a
reasonable level, 7-15 percent in the cooking liquor.  The Na2S04  content can-be kept below
5 percent. As a final product, a mixture of Na2C03 and Na2S03 is obtained, which, after
makeup, can be reused in the NSSC process.

     14.8.3   Ammonium Base

Since this base cannot be recovered, recovery is limited to SO2. S02 recovery can take place
in the same  type of equipment as for the magnesium base. For the operation of a multistage
venturi absorber system, pH control is essential for proper functioning (see section 14.7.3).

14.9  Problem of Nitrogen  Compounds from Ammonium-Based Pulping

Ammonium  base SSL contains about 3 to 10 percent N2 by weight on a dry solids basis
(11). Nitrogen in spent sulfite liquor is present primarily in the form of ammonium salts,
which can result in the emission of  ammonium sulfite and ammonium sulfate containing
particles from  the recovery  furnace. The gaseous forms of N2 that can be emitted from the
recovery  furnace  include diatomic nitrogen gas (N2), NH3, and nitrogen oxides.  Amine
compounds  can theoretically be  emitted from the ammonium-base sulfite recovery furnace,
particularly in strongly reducing atmospheres.

Variables affecting the generation and relative amounts of the respective individual nitrogen
compounds  which can be  formed include the nitrogen  content  of the fuel,  the flame
temperature, the flame zone retention time, and the configuration of the combustion unit.
                                        14-27

-------
The flame temperature is influenced by the heating value of the fuel, the moisture content,
the heat release rate, and the combustion volume of the furnace.

Blakeslee and Burbach (12) report that nitrogen oxide emissions tend to increase with the
nitrogen content for the coal- and oil-fired power boilers. Little information is available re-
garding the effect of the fuel nitrogen content of ammonium base SSL; however, there are
indications that at the lower flame temperatures, particularly  below 1200° C (2200° F), the
primary gaseous products are N2 and NH3 .  The nitrogen present in the fuel is not present in
the diatomic elemental form, but as NH4 ion.

Flame  temperature is  an important variable affecting the  relative amount of nitrogen oxide
emissions formed during the combustion of SSL.  Palmrose  and Hull (13) report that the
flame temperature during ammonium base SSL combustion is about 980 to 1315° C (1800
to 2400° F).  The flame temperature is observed to  increase as the solids content of the
spent liquor increases, because less water is present from the fuel being evaporated to act as
a heat sink. Heat release rates for ammonium base sulfite recovery furnaces generally are
about  1.9 to  8.8 GJ/h per cubic meter (50,000 to 230,000 BTU/hr per  cubic foot) (14).
Heating values are generally between  16 and 21 MJ per kg  of dry solids (7,000 to 9,000
BTU/lb). Loddby-type furnaces normally have relatively large combustion volumes  so that
excessively high temperatures are not found when burning ammonium base SSL.

Studies by Bartok, Crawford, and Skopp (15) indicate the effect of fuel nitrogen content on
nitrogen oxide emissions. Tests run on  fuel oil-fired power boilers indicate that the amount
of No formed is directly proportional to the nitrogen content of the fuel over a range of 0.2
to 1.0 percent nitrogen  by weight.  The nitrogen  content of the fuel appears to  be the
predominant  factor affecting NO emissions  at temperatures below 1300° C (2370° F).
Flame  temperature appears  to be the  predominant factor affecting NO  formation above
1100° C (2000° F), particularly  at 02  concentrations above  2.0 percent by volume in the
flue gas.

Because of the lower nitrogen content of fuel oil (0.2 to 1.0 percent by weight) compared
to ammonium base SSL (3.0 to 10.0 percent by weight), the results of the Bartok et al. study
are not directly applicable to recovery furnaces. They do tend to indicate the general trend
of increasing NO emissions in direct proportion to fuel nitrogen content. This relationship is
applicable particularly at the lower flame temperatures observed in ammonium base sulfite
recovery furnaces because of the large combustion  volumes and the high moisture contents
of the fuels.

A very limited  amount  of measurements  have been made  regarding NO emissions for
ammonium base sulfite recovery furnaces.  Results of one series of tests indicate that total
nitrogen oxide concentrations ranged from 200 to  500 ppm  by volume as N02 with peak
values observed of up to 1000 ppm (16). A summary of results is presented in Table 14-5.
                                       14-28

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                                   TABLE  14-5
    NITROGEN OXIDES* EMISSIONS  FROM AN AMMONIUM BASE SULFITE
                           RECOVERY FURNACE (16)

              Concentration                       Emission Rate**
Condition    ppm by volume   kg N0x/t (Ib/ton)    kg NOX/106 kj  (Ib NOX/106 BTU)

Average           350            8.3    (16.6)          0.36            (0.84)
Minimum         200            4.7     (9.4)          0.22            (0.51)
Maximum         500           11.8    (23.6)          0.52            (1.21)

 *Nitrogen oxides are computed as equivalent nitrogen dioxides.
**Based on flow rate 625 m3/h per daily metric ton of capacity (24,300 ft3/h per short ton) and a liquor
  heating value of 22.7 MJ/kg (9750 BTU/lb).
Concentrations as reported are substantially greater than the 25 to 75 ppm of NOX for kraft
and  magnesium base sulfite  recovery  furnaces,  indicating the  possible role  of nitrogen
content of the fuel in causing NOX emissions.

Nitrogen oxide emissions  can  be controlled by one or any combination of the following
methods:

     1.    Evaporating to  the  minimum solids content (about  50 percent)  necessary to
          support combustion and to provide efficient heat recovery, with the remaining
          water acting as a heat sink during combustion;

     2.    Operating  furnaces at relatively  low firing  rates to minimize the rate of heat
          release;

     3.    Modifying  the type of combustion unit in which the SSL is burned, as by shifting
          to the use of fluidized beds; and

     4.    Operating the furnace at minimum excess air  consistent with efficient combustion
          in an oxidizing atmosphere.

It is also  possible to reduce the nitrogen  content of  the fuel by reducing the amount of
NH3 added to the chemical makeup. The  effect  would be to reduce both the NH3 and
nitrogen oxide emissions from the  combustion process. The NH3 produced could be
removed in the acidic  scrubbing  solution during  passage  through the  series of liquid
scrubbers. The danger in reducing the NH3 makeup is that  it  could alter  the pulping
conditions adversely  and reduce  the  pH  of the liquid  scrubbing  solution,  resulting in
increased emissions of S02.
                                       14-29

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14.10  References

 1.  EKONO Oy, Helsinki, Finland, files.

 2.  Hendrickson, E. R., Roberson, J. E., and Koogler, J. B., Final Report, CPA 22-69-18,
    U.S. Department of Health, Education, and Welfare, National Air Pollution Control
    Administration, Raleigh, North Carolina, March 15, 1970.

 3.  Johnson,  W. D., and Gansler, H., How Rayonier Cuts  Blow Pit Emissions with
    Chemical Scrubbing System. Pulp and Paper, 45(13):54-56, December 1971.

 4.  Axelsson, 0., and Wahlund, L. G.,  Occurrence of Volatile BOD Compounds  in the
    Sulfite  Process  with  Spent Liquor Neutralization  and Condensate Reuse.  Svensk
    Papperstidning, 75(8):287-296, April 30, 1972 (Stockholm).

 5.  Simmons, T., Svensk Papperstidning, 67(7):286-293, April 15, 1964 (Stockholm).

 6.  Guerrier, J. J., Cooperative Effort Solves Small Mill's Air Problem. Pulp and Paper,
    48:62-64, March 1974.

 7.  Copeland, G. C., and Wheeler,  C. M., A Progress Report on the Copeland Recovery
    System at  the  Franconia Paper Corporation,  Lincoln, New Hampshire. (Presented  at
    the TAPPI New Hampshire Section Spring Meeting, Lincoln, New Hampshire, April 23,
    1970.)

 8.  Air Pollutants  Abatement Problems of the Forest Industry. Statens Naturvarsdverk
    (Sweden). Publication 1969: 3, July 1969.

 9.  Alaska Administrative Code, Title 18, Chapter 50.060 (1 and 2) (1973).

10.  Oregon Administrative Regulations, Chapter 340, Section 25-360 (2) (1973).

11.  Clement, J. L., and Sage, W. L., Ammonium Base Liquor Burning and Sulfur Dioxide
    Recovery. Tappi, 52:1449-1456, August 1969.

12.  Blakeslee,  C.  E.,  and  Burbach, H.  E., Controlling NOX Emissions  from  Steam
    Generators. Journal of Air Pollution Control Association, 23:37-42, January 1973.

13.  Palmrose, G. V., and Hull, J.  H., Pilot Plant Recovery  of Heat and Sulfur from Spent
    Ammonia-Base Sulfite Pulping Liquor. Tappi, 35:193-198, May 1952.
                                      ]4-30

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14.  Saiha, E., The  Tampella Process.  In: Proceedings of the Symposium on Recovery of
    Pulping Chemicals. Helsinki,  Finland, May  13-17, 1968. Finnish Pulp and  Paper
    Research Institute and EKONO Oy, Helsinki, Finland, 1969.

15.  Bartok, W., Crawford, A. R., and Skoop, A., Control of Nitrogen Oxide Emissions
    from Stationary Combustion Sources. Combustion, 42(4):37-40, October 1970.

16.  Waddington, E. B., ITT-Rayonier, Inc. Shelton, Washington. September 1973. Personal
    communication.
                                      14-31

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-------
                                    CHAPTER  15

                            OTHER  PROCESS  SOURCES
15.1   Bleach Plant Gases

Bleach  plant gases differ significantly  from other gaseous emissions from  a kraft  mill.
Depending on  the bleaching system and the bleach chemical preparation,  the  gaseous
effluents may include C12, chlorine dioxide  (C102), and S02. All of these gases have about
the same  odor  threshold of about 0.1-1 ppm. The bleach plant belongs  to the minor air
pollution  sources in a kraft mill; its influence is usually restricted to the mill area itself or to
the immediate surroundings. The noticeable effects of the bleach plant gaseous emissions are
mainly odor and corrosion of nearby objects.

     15.1.1   Bleach Plant C12  Emissions

If a bleach plant has a chlorination stage, there are C12 emissions from the C12 bleach tower
vent and from the hood  vent of the succeeding washing stage. If the washer is of a vacuum
rotary drum  type,  the total C12  emission is about 0.5 kg C12  per metric  ton of pulp (1 Ib
Clj/ton) and the total vent flow is about 900 m3/t (29,000 ft3/ton) (1).

For a pressure filter or a continuous diffuser,  the  vent gas flow is much smaller and more
concentrated in C12. The C12 emission is much smaller, too. One feasible way of eliminating
the C12  emission is to scrub  the  C12-containing gases with an NaOH solution. This, in turn,
can be  used  as a part of the  hypochlorite solution needed in the  hypochlorite bleaching
stage. By  eliminating  C12 emissions, C12 is  saved. Profits are based on C12 savings and on
scrubber investment and operational costs.

     15.1.2   Bleach Plant C102 Emissions

If the bleach plant  has two C1O2 stages, C102  is emitted from both the washer hood vents
after the bleach towers and the C102 manufacturing process itself.

If the washers are of  a vacuum rotary drum type, the total C1O2  emission is about 0.3 kg
C102/t  (0.6 Ib  ClO2/ton) and  the corresponding  vent flow  is approximately  800m3/t
(26,000 cu ft/ton) (1). Pressure washers or continuous diffusers will give smaller flows and
emissions.

The washer hood vent gases  are  too diluted for C102 recovery. Therefore, C102 destruction
through scrubbing with an alkaline hydrogen peroxide (H202) solution is  the only feasible
treatment. The reaction is:
                                         15-1

-------
                  2 G1O2 + 2 NaOH + H202 = 2 NaC102 + 02 + 2 H2O

The sodium chlorite (NaC102) solution must be eliminated with the mill wastewaters, since
oxidation to C1O2  with C12 is hardly feasible. The H202 is expensive, and yearly scrubbing
costs may amount to almost 30 percent of the scrubber investment.

In every  C102  manufacturing process, the last step consists of an absorption stage where the
C102 gas is run countercurrently to a stream of cool water to produce a C102 solution. The
C102 concentration is  usually  between 5 and  10 g/1. The C1O2 loss or emission from this
absorption tower is directly proportional to the water temperature (Figure 15-1). At a water
temperature of 20° C (68°  F), the C102  emission is  about 0.2 kg C102/t (0.4 Ib C102/ton)
and the tower vent flow is about 60 m3/t (1900 ft3/ton) (1).

In warm climates, where the fresh water temperature is approximately 20° C (68°  F), it may
be feasible to recover  more C102  in  the absorption tower simply by cooling the water.
Cooling the water from 20° C (68° F) to 4° C (39° F) will decrease the C102 emission from
the absorption tower  by 50 percent (Figure  15-1). In a colder climate, as in Canada or
Scandinavia, cooling is not feasible. Here, alkaline H202 scrubbing must be used.

     15.1.3  Bleach Plant SO2 Emissions

A bleach  plant may need SO2 in either of two situations:

     1.   If it has a Mathieson process for C1O2  manufacture, or

     2.   If it  has  a final SO2 treatment stage (to destroy C12 and hypochlorite rests and to
         adjust the pH of the pulp).

This treatment requires an absorption tower for manufacturing S02 water, and  the tower
will emit some S02. Gas flows are probably  about 25 m3/t  (800 ft3/ton) and the  S02
emission is approximately 1.5 kg S02/t (3 Ib/ton) (1).

In warm  climates,  SO2 recovery through  cooling  of the  fresh water may be possible.
Reducing the fresh water temperature from 20° C (68° F) to 4° C (39° F) will decrease the
S02 emission by 75 percent (Figure 15-2).

In colder climates, fresh  water cooling is not feasible. Tower vent gases may be scrubbed
with an alkaline solution in combination with C12 emissions from the bleach plant, but then
the C12 is no longer recovered as hypochlorite (section 15.1.1).
                                       15-2

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           o>
          LU
          O
          O
           CM
          O
          O
                   0     20   40   60    80    100   120    140   160
                          CI02    PARTIAL  PRESSURE, mm Hg
                                  FIGURE  15-1
              EQUILIBRIUM SOLUBILITY  OF  CHLORINE DIOXIDE
                                  IN  WATER (2)
15.2  Wastewater Treatment

     15.2.1  Process Variables

Release  of malodorous gases from liquid effluent streams during transport or wastewater
treatment operations has been frequently overlooked as an air pollution source. The major
sources  of odorous gases in liquid effluent streams are from the digester and evaporator
condensate waters, with additional contributions from black liquor spills. Odorous gases can
be released to  the atmosphere from free liquid surfaces during  open channel flow, from
manholes and  sewer vents (particularly if there is a change in liquid  elevation),  from
pumping stations, from liquid recycling, and from wastewater treatment operations (4).
                                       15-3

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                01       0.2      0.4   0.6  0.81.0      2.0       4.0   6£>  8.010

                     SO    CONCENTRATION,  % BY  VOLUME
                                   FIGURE  15-2
        EQUILIBRIUM  SOLUBILITY OF SULFUR  DIOXIDE IN WATER (3)
The major  physical and chemical processes that can result in evolution of odorous gases
from liquid effluent streams include:

     1.   Volatilizing of dissolved gases by increasing liquid temperature;

     2.   Increasing the degree of liquid turbulence by agitation, pumping, or other means;
         and

     3.   Releasing acidic gases, such as H2S and CH3SH, by a reduction in liquid pH.

Specific processes that  can result in increased evolution of odorous gases from wastewater
streams include:

     1.   Contact with  hot flue gas streams, as during liquid scrubbing;
     2.
Liquid pH changes resulting from contact with acidic flue gas streams containing
C02,  acidic liquids such as chlorination effluent waters, or spent acid  streams
from tall oil acidulation or C102 manufacture; and
     3.   Agitation of liquid streams by mechanical aeration, diffuser aeration with gases,
         or other means.
                                       15-4

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Process variables affecting odorous gas release include inlet concentrations of the particular
constituents, liquid temperature, pH changes, degree of agitation, physical  configurations,
02 content of the liquid, and biological activity.

The  major chemical  constituents that  produce odors from pulp and paper  mill effluent
waters are organic sulfur and organic nonsulfur compounds. A major potential problem can
exist with kraft pulp  mills, where organic sulfur compounds, such as CH3SH and CH3SCH3
formed during the cooking operations, can be released into digester condensate waters. They
can also be released to a lesser extent from black liquor into evaporator condensate waters
or from  black liquor spills  from storage tanks or overloaded multiple-effect evaporator
systems.

Organic nonsulfur compounds, such as terpenes and other wood extractive materials, can
also be evolved from effluent waters in either kraft or sulfite pulp mills. Terpene compounds
can be evolved from digester condensate waters in kraft pulp mills in the form of droplets that
can absorb  sulfur gases from subsequent transport downwind from treatment  processes.
Overloaded  or  incompletely mixed biological  waste  treatment facilities  with  sludge
accumulations or sections with inadequate dissolved oxygen (DO) levels can be  the source of
odorous gases, such as organic acids formed by anaerobic fermentation. Agitation of liquids,
where  these conditions exist, can  result  in  the  liberation of substantial quantities  of
malodorous gases to  the atmosphere. If these gases are associated with particulate matter,
such as  bacterial  cells or  liquid droplets, they  can  be transported for  long distances
downwind.

     15.2.2   Treatment Methods

The  two major approaches  to minimize  malodorous gas generated from  liquid effluent
streams  are  by  inplant treatment and by  modification  of the wastewater  treatment
facilities. The  major inplant treatment methods to reduce odorous gas release to the liquid
effluents are by air or steam stripping of digester and evaporator condensate waters. (These
methods are discussed in Chapter 5  of the manual for kraft pulp mills and  Chapter 14 for
sulfite  pulp  mills.) An additional  inplant treatment method involves effective housekeeping
to minimize black liquor spills, as well as other spills, by accidents or carelessness, systems
for segregation and containment of spilled liquids, and enclosure of flowing liquid effluent
streams.

One step that can be taken to prevent release of odorous gases from effluent waters is to en-
close tanks, as well as the biological aeration process. This step is feasible for activated sludge
aeration processes, that it would normally involve a prohibitive expense if applied to aerated
stabilization basins, because of the extensive surface areas involved. A modification to the
activated sludge process was recently developed that employs molecular oxygen (O2) as the
oxygen supply and uses four enclosed aeration tanks in  series (5).  The system results in a
                                         15-5

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small 02-rich gas stream that is easily incinerated in a lime kiln or other combustion unit.
The process also has potential as a pretreatment system, which is upstream of a conventional
aerated stabilization basin, to biologically oxidize or stabilize rapidly reactive materials, such
as terpenes or organic sulfur compounds.

Addition of specific  oxidizing gases to the liquid stream will reduce odor levels in pulp and
paper mill effluent waters.  Gases that have possible application for deodorization of liquid
effluent waters include C12, C102, 02, and ozone (O3). Alferova, Panova, and Titova (6)
describe a two-stage system, developed in the Soviet Union, for treatment of digester and
evaporator condensate waters by  aeration  and   chlorination  in  series.  The  combined
condensate liquid is first aerated in a multiple tray tower with air and then contacted with
acidic  chlorination  bleachery  effluent for further oxidation  of the  odorous compounds
present. Chlorine dioxide  has been applied for effluent treatment  odor control in  the
petroleum industry, and can have application in the pulp and paper industry (7).

The 02 can also be used for odor control by augmenting dissolved oxygen levels in existing
secondary wastewater  treatment  facilities.  Murray and Rayner (8) report that CH3SH is
easily oxidized  in the presence of 02 to CH3SSCH3. The CH3SSCH3  can then be oxidized
to sulfones, sulfonic acids, and other relatively innocuous products, but long retention times
are required. Any CH3SCH3 is not easily oxidized in the presence of 02 except at very high
temperatures not normally found in wastewater treatment.

Two  recent studies  describe  the  effects of O2  addition  to liquid  effluent streams  for
augmenting dissolved oxygen levels  in aerated stabilization basins and receiving waters. A
sidestream  oxygenation system installed at  the Brewton, Alabama, kraft pulp mill, where
4.5 t (5 tons) 02/day were added  to the liquid effluent, resulted in an increase in dissolved
oxygen levels and a decrease in odor from the aerated stabilization basin (9). Amberg (10)
describes similar results when sidestream oxygenation is employed for  augmenting dissolved
oxygen levels in receiving waters.

15.3   Odor Problems from Diffuse Sources

A complete inventory of the odorous emissions from a kraft pulp mill has, so far, comprised
measurements in stacks and other  point sources only. In addition, however, odorous sulfur
compounds are emitted from sources of a diffuse nature, such  as leaking process equipment
and settling basins; these must also be taken into  account. Expansion of odor elimination
systems for large point sources will increase the relative importance of the diffuse sources. A
method for quantitative assessment of  these emissions was developed and is described as
follows (11):

     1.  The  method  uses  simultaneous  sampling  and  flow  measurement  in  a
         suitable  number  of  sections on  the  lee side of  the   diffuse  source.  The
                                         15-6

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         total  emission  is  calculated  by  summation  of  the  contributions  from
         individual sections.

    2.   The instrumentation consists of a fixed analysis unit (a gas chromatograph with a
         flamephotometric  detector)  and  portable  units  that  are  all  capable  of
         simultaneous gas sampling and flow measurement.

The method was tested  at a kraft pulp mill in a survey of the diffuse  emissions from the
settling basin, as well as from the digestery, recovery furnace, and limew ashing building. The
results of the measurements are given in Table 15-1.
                                   TABLE 15-1
     ODOROUS  GASES  FROM  DIFFUSE  KRAFT PULP MILL SOURCES (11)
   Diffuse Emission Source


Settling Basin Measurement I


Settling Basin Measurement II


Settling Basin Measurement III


Digestery I (fir)


Digestery II (birch)


Recovery furnace I


Recovery furnace II


Limewash building
                                          Emitted Compound-Emission Rate
H2S

1,900
(4.2)
1,700
(3.8)
1,800
(9.0)
2
(0.004)
3
(0.007)
4
(0.009)
10
(0.022)
1
(0.002)
CH3SH
g/
600
(1.3)
600
(1.3)
600
(1.3)
10
(0.022)
10
(0.022)
10
(0.022)
40
(0.088)
—
CH3SCH3
h (Ib/hr)
60
(0.13)
70
(0.15)
80
(0.18)
60
(0.13)
40
(0.088)
10
(0.022)
20
(0.044)
—
CH3SSCH3

400
(0.88)
400
(0.88)
600
(1.3)
40
(0.088)
20
(0.044)
30
(0.066)
60
(0.13)
—
                                        15-7

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15.4  References

 1.  EKONO Oy, Helsinki, Finland, files.

 2.  Haller, I. F., and Northgraves, W. W. Tappi, 38:199, April 1955.

 3.  The Finnish Paper  Engineers' Association (SPY). The Pulping of  Wood. Frenckellin
     Kirjapaino Oy, Helsinki, 1968 (Finnish).

 4.  Sableski, J. J., Odor Control in Kraft Pulp Mills: A Summary of the State of the Art.
     U.S. Department of Health, Education, and  Welfare, Public Health Service, National
     Center for Air Pollution Control, Cincinnati, Ohio, May 10, 1967.

 5.  Grader, R. J., South, W.  D., and  Djordjevic, B.,  The Activated Sludge Process  Using
     High Purity Oxygen for  Treating Kraft Mill Wastewaters.  Tappi,  56:103-107,  April
     1973.

 6.  Alferova, L. A., Panova, V.  A., and Titova, G. A.,  Deodorization of Effluents from  the
     Manufacture of Kraft Pulp. Bumazhnaya Promyshlennost (Moscow, USSR), 38, (6),
     5-8, June 1963.

 7.  Samsel, J. J.,  and Hawkins, E. A., Waste Water  Treatment at Texaco's Puget Sound
     Refinery.  Proceedings of the  American Petroleum  Institute, Division of Refining,
     40(III):302-308, 1960.

 8.  Murray, F. E., and  Rayner,  H. B.,  The Oxidation  of Dimethyl Sulfide with Molecular
     Oxygen.  Pulp and Paper Magazine, 69(9):64-67, May 3, 1968.

 9.  02 and 03 - Rx for Pollution. Chemical Engineering,  77(4):46-48, February 23,  1970.

10.  Amberg, H. R., Wise, P. W., and Aspitarte, T.  R., Aeration of Streams with Air and
     Molecular Oxygen. Tappi,  52:1866-1871, October  1969.

11.  Institutet for Vattenoch Luftvardsforskning Bulletin, 2(2):6-7, 1973 (Stockholm).
                                         15-8

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                                  CHAPTER  16
                                POWER BOILERS
 16.1   Supply Patterns

 The pulp and paper industry is a major energy consumer in the United States with a total
 process energy  consumption  requirement  in 1970  of about  1.6 X 1018 J  per year
 (1.5 X 101S BTU  per year). (See Chapter 1.) Based  on 1972 figures compiled  by the
 American  Paper Institute, about 33 percent of the industry's total energy requirement is
 supplied  by  combustion of its waste  pulping liquors and  an additional  7 percent  by
 combustion  of wastewood and  bark.  The remaining 60 percent  of the  total  energy
 requirement must be supplied by purchase of auxiliary fossil fuels, such as  coal, oil, and
 natural gas or by purchase of electricity. A breakdown of the total national process energy
 supply patterns for the pulp and paper industry is presented in Table 16-1 (1). Considerable
 variations in energy supply for mills in different regions of the United States can exist.
                                   TABLE  16-1
 OVERALL NATIONAL DISTRIBUTION OF ENERGY  SOURCES FOR THE  PULP
                         AND  PAPER INDUSTRY (1) (2)
   Energy Source
Pulping liquors
Waste wood
Bituminous coal
Residual fuel oil
Distillate fuel oil
Natural gas
Purchased electricity
Other sources
Total
Overall Energy Consumption
% of Total*
33
7
11
20
2
21
5
1
100
MJ/yr**
522 X 109
110 X 109
174 X 109
317 X 109
32 X 109
332 X 109
79 X 109
16 X 109
1582 X 109
(BTU/yr)**
(495) X 1012
(105) X 1012
(165) X 1012
(300) X 1012
(30) X 1012
(315) X 1012
(75) X 1012
(15) X 1012
(1500) X 1012
 *Based on 1972 data.
**Based on 1970 data.
                                      16-1

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16.2  Combustion Parameters

To  obtain usable energy from fuel, fuel is normally burned in air to release heat and then
the heat energy is harnessed by using it to generate steam. The kinetic and thermal energy of
the steam, in turn, can be used in any variety of ways. To make the fuel burn satisfactorily,
it must be mixed with air in the proper proportions in a suitable furnace. To convert water
to steam at appropriate pressure, the water normally flows through tubes, which are warmed
on their outside by the hot gaseous products of combustion. Once steam forms, it is further
heated to the desired degree of superheat by channeling it through another set of similar
tubes, called the superheater. The  types and  amounts  of air pollutants emitted from the
combustion process depend on the characteristics of the fuel burned, the configuration of
the combustion unit, and the operating parameters of fuel and air supply.

     16.2.1   Fuel Characteristics

The major parameters affecting air pollutant  formation during fuel combustion in power
boilers are the physical state, heating value and  moisture  content of the fuel, as well as its
ash, sulfur, and nitrogen contents. Coal and wastewood are the primary types of solid fuels
employed; both require extensive materials handling facilities for the unburned fuel before
combustion and  for the noncombustible ash  following  combustion. Both distillate and
residual liquid fuel oils are employed in the pulp and paper industry. Heavy residual fuel oil
is more commonly used and has a high air pollution potential. Natural gas is the primary
gaseous fuel  employed in the pulp and paper industry; it is easiest to burn and has the
lowest air pollution potential of all the fuels. Each of  the different fuels requires a different
type of combustion unit because of differences in burning characteristics.

The heating value of a fuel determines the amount that must be burned to generate a given
amount of usable energy. The heating value  varies with the kind of fuel employed, the
location from which it was derived, and its moisture content. Heating values for bituminous
coal and wastewood can vary substantially among those derived from different locations, as
can their ash and sulfur contents. A summary of typical heating values, moisture contents,
sulfur  and ash contents of  coal, oil, gas,  and wood  fuels employed in  pulp and paper
manufacture are presented in Table  16-2 (3).

Major fuel  consumption  parameters affecting air pollutant emissions from power boilers in
the pulp and paper industry include sulfur and ash contents of the fuels. The SO2 emissions
from fuel combustion are directly proportional  to the sulfur content. The sulfur content is
normally  significant in  bituminous  coal and residual fuel oil. The  major  methods for
minimizing air pollution from fuel combustion of S02 in power boilers are the substitution
of low sulfur oil, coal, or natural gas for high sulfur fuels and the construction of tall stacks
for dilution of ground level S02 concentrations.
                                         16-2

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                                   TABLE 16-2
  CHARACTERISTICS  OF  FUEL BURNED IN  POWER BOILERS AT PULP AND
                   PAPER  MILLS IN THE  UNITED  STATES  (3)
                                                       % by Wt. Content of
               Heating Value (As Fired)
                Moisture
 Fuel     Units    Average
(Range)
      Ash
Average (Range)
Sulfur
Coal     MJ/kg        31       (24-34)
         Btu/lb    13,500   (10,500-14,700)

Oil       MJ/1         41       (34-43)
         Btu/gal   149,000  (122,000-155,000)

Gas      MJ/m3        38       (37-40)
        Btu/cuft     1,030     (1,000-1,070)
               10 (7-15)    8.1 (3.5-15)  1.9 (0.5-10)
                0.5 (0-1)   0.1 (0.01-0.2) 1.8 (0.1-3.5)
                0.2 (0-1)
      Neg.
 Neg.
Wood     MJ/kg         11       (9.3-13)        25(10-60)   2.9(0.1-20)   0.1(0-0.2)
          Btu/lb      4,600     (4,000-5,500)
The  particulate  emissions from fuel combustion generally are proportional  to  the  ash
content of the fuel and are  significant for coal and wood combustion. The ash material
contained in the fuel can be  removed from the boiler as fly ash in the exhaust gases or as
bottom ash in  solid form or  as liquid slag, depending on the  ash fusion temperature.
Normally  80 to 95 percent of the ash material is emitted from the  boiler as  fly ash.
Unburned or partially burned  fuel aslo can  be emitted from the combustion  chamber as
particulate matter. Devices normally used for particulate emission control on coal-fired
power  boilers in the pulp and paper industry are electrostatic precipitators or  mechanical
collectors;  mechanical collectors and  liquid  scrubbers  are  most commonly used  for
wastewood- and bark-fired power boilers.

Recent investigations  show that other constituents in fuels may also  contribute to air
pollution. Martin and Berkau (4) report that the nitrogen content of fuels can contribute to
the formation of nitrogen oxides, though the reaction between atmospheric N2 and 02 is
predominant for flame temperatures above 1,300°  C (2,400° F). In addition, trace elements,
particularly in bituminous coal and residual fuel oil, may be significant air pollutants, such
as the  nonmetals chlorine (Cl), fluorine (Fl), and phosphorus  (P), and the heavy metals,
such as beryllium (Be), mercury (Hg), lead (Pb), cadmium (Cd), zinc (Zn), arsenic (As), and
selenium (Se). A summary of air pollutant emission factors from specific fuel combustion
processes for industrial power boilers is presented in Table 16-3 (5).
                                        16-3

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                                                    TABLE 16-3
  UNCONTROLLED AIR POLLUTANT EMISSION  FACTORS FOR  FUEL  COMBUSTION IN POWER BOILERS FOR THE
                                         PULP  AND PAPER INDUSTRY (5)
      Fuel Burned

Bituminous Coal
  Pulverized General
  Pulverized Wet Bottom
  Pulverized Dry Bottom
  Pulverized Cyclone
  Spreader Stoker

Fuel Oil
  Residual-Tangential
  Residual-Horizontal
  Distillate-Tangential
  Distillate-Horizontal

Natural Gas
  Process Boilers
  Gas Turbines

Waste Wood
  No Reinjection
  50% Reinjection
  100% Reinjection

%A = Percent ash in fuel.
%S = Percent sulfur in fuel.
                           Particulate Matter
Sulfur Oxides
Nitrogen Oxides    Hydrocarbons    Carbon Monoxide
kg Part
106 kj
0.34 (%A)
0.28 (%A)
0.37 (%A)
0.04 (%A)
0.28 (%A)
0.023
0.023
0.015
0.015
0.006
0.006
1.50
1.79
2.32
Ib Part
106 BTU
0.80 (%A)
0.65 (%A)
0.85 (%A)
0.10 (%A)
0.65 (%A)
0.053
0.053
0.035
0.035
0.014
0.014
3.500
4.150
5.400
kgS02
106kJ
0.82 (%S)
0.82 (%S)
0.82 (%S)
0.82 (%S)
0.82 (%S)
0.46 (%S)
0.46 (%S)
0.41 (%S)
0.41 (%S)
0.00048
0.00048
0.16
0.16
0.16
lbSO2
106 BTU
1.90 (%S)
1.90 (%S)
1.90 (%S)
1.90 (%S)
1.90 (%S)
1.06 (%S)
1.06 (%S)
0.96 (%S)
0.96 (%S)
0.001
0.001
0.375
0.375
0.375
kgN02
106kJ
0.39
0.64
0.39
1.17
0.32
0.115
0.230
0.123
0.246
0.160
0.183
0.43
0.43
0.43
lbNO2
106 BTU
0.900
1.500
0.900
2.740
0.750
0.267
0.535
0.285
0.570
0.372
0.425
1.010
0.010
0.010
kgCH4
106kJ
0.0064
0.0064
0.0064
0.0064
0.0064
0.0005
0.0005
0.0005
0.0005
0.016
0.016
0.11
0.11
0.11
IbCH4
106 BTU
0.015
0.015
0.015
0.015
0.015
0.001
0.001
0.001
0.001
0.038
0.038
0.250
0.250
0.250
kg CO
106kJ
0.021
0.021
0.021
0.021
0.021
0.0005
0.0005
0.0005
0.0005
0.0005
0.0005
0.11
0.11
0.11
IbCO
106 BTU
0.050
0.050
0.050
0.050
0.050
0.001
0.001
0.001
0.001
0.001
0.001
0.250
0.250
0.250

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     16.2.2  Furnace Characteristics

Each fuel burned must  be fired in a separate type of boiler depending on its physical state
and  burning characteristics. Natural  gas is normally burned for steam generation in package
type boilers of varying  sizes,  where the gas-air mixture can be added in a horizontal
configuration along one side. Burning residual fuel oil in power boilers is similar to burning
in gas-fired boilers where the oil-air mixture is normally  added in either a horizontal or
corner tangential firing configuration.

The  burning of oil requires a more complex fuel delivery  system, normally with pumping,
than gas.  Residual  oil is  also  usually heated in inlet piping by steam tracing to reduce its
viscosity and, therefore, prevent plugging of the fuel lines.  Tangential firing of oil or gas can
result in substantial reductions in nitrogen oxide concentrations as compared to horizontal
firing because  of  the reduced  flame temperatures, greater  air-fuel  mixing, and resultant
lower excess air requirements (6).

Burning solid fuels in power boilers, such as coal or wastewood, requires even more complex
design  than burning oil  in power boilers because of the more complex combustion process
and the more complex  fuel delivery  and ash handling systems. The major types of coal-fired
power  boiler employed in the pulp  and paper  industry are the spreader stoker, the chain
grate stoker (now in limited use), and several varieties of pulverized firing systems, including
dry bottom, wet bottom,  and cyclone type units, which can be horizontally, tangentially, or
vertically  fired. The type and design of boiler used depends on the burning characteristics
and  chemical  composition of  the  coal;  these  include the ash  content,  ash softening
temperature, fixed carbon content, volatile carbon  content, and heating value.  More
thorough  presentations of design  parameters for  coal-fired power boilers are available in
several references (7, 8,  9, 10).

Wastewood- and bark-fired power boilers can burn the wood alone or can be modified to
burn other fuels on an auxiliary basis or in combination.  Wastewood and bark are burned in
power  boilers  on  chain  grates in a radiant Dutch-oven  type boiler or in a horizontal
air-blown suspended firing configuration in a vertical Stirling boiler. Wood handling systems
(including hammermill grinding to a given particle size for suspended firing), bottom and fly
ash handling systems, and underfire  and overfire air controls  must be provided. Major fuel
characteristics  affecting the design  of wastewood-fired  power  boilers include  ash  and
moisture  content,' particle size variations,  and fixed and volatile  carbon content. The
amount of particulate matter swept  from the combustion chamber is normally greater from
the horizontal suspension firing units than from the Dutch-oven type units.

     16.2.3  Combustion Variables

The  most important variable affecting the combustion process is the fuel-to-air ratio, which
can  be varied  between wide  limits even  when the other  variables remain unchanged.

                                         16-5

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Variations in the fuel-to-air ratio can be expressed as variations in the amount of excess air
or as  variations in  the oxygen content  of the flue gas. The  variations  influence the
combustion  temperature,  combustion efficiency,  ignition speed,  flame length, flame
radiation, and flue  gas composition.  Poor combustion  efficiency  increases  particulate
emissions,  while high  combustion  temperature, together with  an  oxygen-rich  furnace
atmosphere, produces oxides of nitrogen.

Appropriate  mixing  of  fuel and  combustion air is essential to achieve fast  and complete
combustion.  The design of firing and furnace has the major influence  on the  mixing, but
mixing also can be controlled to  some extent by operational variables such as velocity and
direction of  fuel, and especially  air streams. If mixing is not sufficient, part of the fuel
remains unburned and results in particulate emissions. The formation of nitrogen oxides is
known to be higher in diffusion flames, where mixing takes place in the flame rather than in
premixed flames (11). Nitrogen oxide formation increases with both flame temperature and
02 level in the combustion zone.

16.3   Boiler Types

The  major  parameters affecting the  design of power boilers to minimize potential air
pollutant emissions include:

     1.  Physical state, chemical composition, and burning characteristics of the fuel,

     2.  Method of firing fuel and mixing,

     3.  Combustion chamber volume and configuration, and

     4.  Other  operating parameters, such as fuel  firing rates, excess air levels, and air
         distribution.

The  type,  size, and  collection  efficiency  requirements for the gas  treatment system are
determined by the gas flow rate and flow conditions from the boiler, the exit gas pollutant
concentrations from the boiler, and  specific parameters, such as particle size. Whatever the
fuel, the air pollutant emission characteristics from power boilers are  very much influenced
by design and operating parameters.

     16.3.1    Gas-Fired Power Boilers

Natural gas-fired power boilers normally are the simplest in design and operation of any of
those used in the pulp and paper industry. They require only piping of the gas to the boiler
and  mixing  it with air to provide  adequate heat for steam  generation; they require no
complex fuel or ash material handling systems.  To this extent, the design and operation of
                                         16-6

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these boilers are essentially the same as gas-fired boilers in the electrical utility industry, but
the units in the pulp and paper industry are normally smaller and generate steam chiefly for
process use instead  of for driving turbines. The only potential air pollutants from natural
gas-fired power boilers are oxides of -nitrogen generated at elevated temperatures during the
combustion process.

The two major types of natural gas-fired power boilers  used in the pulp and paper industry
are horizontal- and tangential-fired units. These units differ only in the method of air-fuel
introduction to the combustion chamber. The horizontal-fired units employ firing along one
side of the furnace; the tangential units use corner-firing. The nitrogen oxide emissions from
horizontal-fired units are normally greater than from tangential-fired units because of their
characteristically higher flame temperatures.

The lack  of radiating particles in a gas flame requires a large furnace volume  to ensure
sufficient  cooling of the  flue gas before entering the other parts of a boiler.  A certain
amount of particle formation is useful and can be achieved with special burner design or
with two-phase firing for boilers designed primarily for burning natural gas. Natural gas can
also be burned in pulp  and paper industry power boilers that are designed primarily for
other fuels and where the  gas can be used either for startup or combination firing. Because
of the relative simplicity of gas firing in combination boilers, the other fuels dominate the
design requirements and  the boiler emissions.

The major operating variables for gas-fired power boilers are the amount of excess air added
and the distribution of primary and secondary air if off-stoichiometric firing is used.  Flue
gas recirculation to the furnace can also be added  to improve combustion efficiency and to
reduce excess air requirements. Operating natural gas-fired power boilers at sufficient excess
air  allows complete combustion, but more important it  also promotes  nitrogen  oxide
formation (12).

The sulfur  and ash  content of  natural gas are negligible, and the  respective SO2  and
particulate matter emissions are insignificant. Complete combustion assures that minimal
emissions  of CO and hydrocarbons will  occur. The major  air pollutants from natural gas
combustion are nitrogen oxides, which  can be minimized by off-stoichiometric firing, use of
minimal excess air, and flue gas recirculation (13).

     16.3.2  Oil-Fired Boilers

Oil-fired power boilers are normally similar in design to natural gas-fired units except that
they  require  more  complex fuel handling  facilities and  longer retention times in the
combustion zone. The two major types of liquid petroleum fuels burned are No. 2 distillate
fuel oil and  No. 6 residual fuel oil.  Both types of oils require fuel storage, pumping and
injection systems, with steam or electrical tracing of fuel lines needed for residual fuel oil to
                                         16-7

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avoid plugging by the highly viscous liquid. Atomization of the oil injected into the furnace
from the firing guns is necessary to provide a spray of fine droplets, assuring complete
combustion of the fuel. The major potential air  pollutants from  fuel oil combustion are
nitrogen oxides and sulfur oxides, and, to a somewhat lesser  extent,  particulate matter.

The two major types  of furnace configurations for burning fuel oil in pulp and paper
industry power  boilers are horizontal  front-fired units and tangential corner-fired units.
These  are  similar in design to those employed for natural gas firing except that longer
retention times  in  the  combustion  zone are  normally   required  to assure complete
combustion of the oil. Atomizing and mixing promote combustion, and thereby reduce the
formation of unburned carbon particles to the practical minimum.

The furnace volume must be sufficiently large enough to prevent the flame from impinging
on  the cooled walls; a cooled flame would inhibit the last steps of the combustion reactions
and result in the emission of carbon soot particles. The volume of the furnace varies directly
with  the released  thermal  energy,  and  the formation of  nitrogen  oxides,  in  turn,  is
proportional to this energy. Because a smaller combustion chamber  causes an increased heat
release rate per unit volume, an increase in flame temperature results, which is accompanied
by an increase in formation of nitrogen oxides.

Major parameters affecting the operation of oil-fired power boilers are:

     1.    Excess air level,

     2.    Air flow distribution pattern,

     3.    Fuel inlet temperature as it affects oil viscosity, and

     4.    Amount of steam  added for atomization.

The excess air level during fuel oil firing must be  carefully  controlled within the maximum
and minimum limits determined by  consideration of both thermal  energy conversion
efficiency and air pollutant emission levels. The lower practical limit for excess air is reached
when carbon particles and combustible gases, such as CO and hydrocarbons, are detected in
the  flue gas.  The upper practical limit for excess air is  defined by the formation  of
significant  quantities of nitrogen oxides and the presence of sulfur trioxide (S03). Sulfur
trioxide formation is accelerated  by the presence of trace metal catalysts, such as vanadium
(V) in the oil. An additional consideration is that excess air levels above the upper limits
result in a decrease in the thermal efficiency of the boiler by increasing the flue gas flow rate
and the flue gas outlet temperature.
                                         16-8

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An important variable in maintaining sufficient atomization is the viscosity of the fuel oil,
which is maintained by keeping the oil temperature within rather narrow limits. Increased
viscosity at low oil temperatures dampens the vibration of oil  droplets and thus reduces the
atomization effect. Too high a fuel temperature tends  to cause coking of the oil on the
hottest parts of the burner and disturbs the atomization process by producing steam and gas
bubbles in the oil before it is atomized (14).

The major potential air pollutants from fuel oil combustion are sulfur oxides, nitrogen
oxides, and particulate matter.

The S02  emissions from fuel oil combustion are essentially a function of the sulfur content
of the fuel, which normally ranges from 0.1 to 0.5 percent by weight for light distillate oils
and from 0.5 to 3.0 percent, or more, by weight for heavy residual fuel oils, depending on
the  source  of  the  crude.  The  major method  for  controlling S02  emissions  from oil
combustion in  the pulp and paper industry, to date, is the substitution of natural gas or low
sulfur oil for high sulfur oil. Varying quantities of S03, accounting for up to 10 percent of
the total sulfur burned, can also be formed during fuel oil combustion. The S03  can be
hydrolyzed in the presence  of water to H2S04  and cause corrosion  on cold metal surfaces.
The amount of SO3 formed depends on the sulfur content of the fuel, the excess air level,
and the possible presence of trace metal catalysts in the oil; high concentrations of up to 50
or 60 ppm by volume have been observed (15).

Particulate matter emitted from fuel oil combustion consists of inorganic material from the
ash content of  the fuel and organic materials resulting from  incomplete combustion. The
organic  matter  from  fuel  oil combustion consists  primarily of unburned carbon  soot
particles  resulting from incomplete combustion of the  oil droplets, approximately 95
percent by weight of the soot particles are less than 10 jum (4 X 10 ~4  in) in diameter. The
carbon content of the particulate matter  can be as high as 58 percent by  weight, and
because  of this high carbon  content,  these  particles  are   not particularly  suitable for
collection by electrostatic precipitation (16).

The emission of particulate matter from oil firing is strongly dependent on the  excess air
level during combustion. Any  major disturbance  in burner operation  is almost certain to
produce a considerable increase in the  particulate matter emissions. The burning of residual
fuel oils with sulfur contents above 1.5 percent by weight can cause the particulate matter
from the boiler.to become saturated with adsorbed  H2S04 droplets when the gas stream
temperature drops below the acid dew point.  The sticky acidic particle deposits  can cause
rapid corrosion of metal surfaces and, therefore, lower the  efficiency of the mechanical
cyclone collectors.

Nitrogen oxide controls for oil-fired power boilers are similar to those for gas-fired units; the
objectives  are  to minimize flame temperatures and  excess air  levels.  Nitrogen oxide
                                         16-9

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formation is controlled  by off-stoichiometric firing with split air distribution, minimum
excess air, and possibly, flue gas recirculation.

     16.3.3   Coal-Fired  Boilers

Coal-fired power boilers are normally more complex in design than either oil- or gas-fired
units because complex solid fuel handling and ash handling systems are required, the fuel
injection system is more  complicated,  and  longer retention times with greater air-fuel
mixture  turbulence levels  in the  furnace   are  generally  required  to assure complete
combustion.  The main type of coal burned at present for auxiliary power generation in the
pulp and paper industry is the  bituminous grade from the eastern or midwestern United
States. The  properties  of  coals can vary considerably  between  individual fields.  The
respective sulfur contents, ash contents,  and heating values all affect potential air pollutant
emissions. The major  potential air  pollutants from coal combustion in pulp  and paper
industry power boilers  are sulfur oxides, particulate matter, and nitrogen oxides.

There are several types of coal-fired power boilers employed in the pulp and paper industry.
These  differ mainly in the  methods by which the  fuel is added  to the furnace,  the
configurations  by which air is added to the  combustion  chamber, and the  boiler bottom
configuration.  The major coal feed configurations  commonly employed in  the pulp  and
paper industry include pulverized firing, spreader  stoker, and chain-grate stoker  units. The
type of unit  employed depends on the amount and properties of the coal to be burned. The
ash content,  ash softening temperature, fixed  carbon and volatile carbon contents, moisture
content,  and, to a lesser extent, sulfur content all influence the type of unit employed. The
sizes and shapes of the furnaces are largely influenced by the heating value and ratio of fixed
to volatile carbon contents in the coal.

The smaller  coal-fired power boilers  in the  pulp and  paper industry  are generally either
underfeed or overfeed  chain-grate stoker units or suspended-firing spreader stoker units. The
larger coal-fired power  boilers  are  generally pulverized-firing, employing either vertical,
horizontally  opposed,  tangential corner, or wall-fired units, with either dry or wet bottoms.
The pulverized units are employed  for  coals with high ash softening temperatures, while
coals with ash softening temperatures below 1,200° C  (2,200° F) are generally burned in
cyclone furnaces, though these are not commonly employed in the pulp and paper industry.
The design  features of the  individual  types  of coal combustion units are  extensively
reviewed in other references (7, 8, 9, 10).

The particulate emissions in the flue gases from coal-fired power boilers tend to increase
with increasing coal ash content and  with increasing ash softening temperature. They  also
basically show the  greatest increase  where pulverized units employ dry bottoms because of
suspension of small particles. Particulate emissions are lowest from pulverized cyclone units
because  most  of the ash  is removed  as slag  from  the furnace bottom. The flue gas
                                         16-10

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temperature  of the furnace section outlet normally must be  maintained below  the  ash
softening temperature of the coal being burned in order to keep particulate matter from
melting and  sticking on tubes in the superheater section. Coal furnaces must generally be
designed to provide sufficient cooling by means of water walls or other methods to prevent
superheater  slagging. The resultant lower  flame zone temperature  also acts to reduce
nitrogen oxide emissions.

Major air operating variables affecting  emissions from coal-fired power boilers include  the
excess air level  and the relative distribution of primary and secondary air (or underfire or
overfire  air). The  excess air level  is significant  in  that increased  CO  and hydrocarbon
emissions are favored by minimum excess air levels. These minimum levels also can increase
the carbon content of the  particulate matter leaving  the boiler.  This  increased carbon
content can adversely affect  operation of the electrostatic precipitator by increasing particle
resistivity levels. High excess  air levels tend to increase the amounts of nitrogen oxides
present. Excessive overfire air levels on chain-grate or spreader stoker units can also exert a
sweeping action to entrain small particles in the exhaust gases.

Fuel-related  operating  variables  influence particulate  emissions from  coal-fired power
boilers.  The  fineness of  pulverized  coal  particles is influenced  by the degree  of grinding,
which is a  factor that may  vary during the lifetime of the grinding elements in  the
pulverizer. If not controlled, pulverizer wear affects the combustion efficiency by causing
progressively larger particles  that increase the carbon content in  the fly ash. In grate firings,
an excessive  air flow can penetrate areas in the fuel layer if its deposit on the grate is not of
uniform thickness. Localized particulate entrainrrient results.

The  particulate emissions from coal-fired power boilers depend on the ash content of  the
coal, its ash softening temperature, the method of firing  employed, and to a lesser extent on
the excess air for organic constituents.  Particulate emissions generally are greatest  from
pulverized dry bottom units, lower from chain  grate and spreader stoker  units, and lowest
from  cyclone furnaces.  The degree  of coal grinding  before firing and the type of  firing
employed  both  influence the  particle  size distribution  in  the  exhaust gases. There is
insufficient  information  regarding  the  particle  size distribution from  pulp and  paper
industry coal-fired units at present.  It  is known, however, that the amount of particulate
matter discharged as fly ash  relative to  the amount of  slagged bottom ash increases with  ash
softening temperature.

S02  emissions from coal combustion are directly proportional to  the sulfur content of  the
coal being burned. Increases in excess air level tend to cause increases in S03 emissions from
coal  combustion. But the presence of calcium and magnesium oxides  in the particulate
matter tends to bind any S03  as CaS04  or MgS04. Present control methods for sulfur
oxides include fuel substitution and the construction of tall stacks. Nitrogen  oxide  levels
from coal combustion tend to increase with combustion zone flame temperatures and excess
                                         16-11

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air levels in the exhaust gases. Nitrogen oxide levels are comparable to those from oil-fired
units, except for cyclone furnaces, where  the high temperatures required for ash slagging
also cause excessive formation of nitrogen oxides.

     16.3.4   Wood-Fired Boilers

Waste wood combustion is often employed in pulp and paper mills because the  material is
readily available from  wood debarking or  associated lumber  and plywood manufacturing
operations. The major  types of materials that can be burned in waste wood boilers are bark
from  mechanical or hydraulic debarking operations and  waste wood materials, such as
sawdust, shavings, slabs, and chips from lumber and plywood manufacture. Major factors
affecting waste wood boiler design and use are the quality and amount of material available
and its burning characteristics. Waste wood materials  are often burned in combination with
other fuels such as oil, gas, coal, or clarifier sludge from wastewater treatment.  The major
potential air pollutant from waste wood combustion is particulate matter; hydrocarbons and
nitrogen oxides are also emitted.

The starting point in designing a waste wood firing furnace is different from that for other
fuels because bark and wood are waste fuels to be  incinerated. The objective of firing waste
wood is to maximize the use of  the  released energy from waste wood  to minimize the
amount of other fuel to be purchased.

Major design  parameters for waste wood  boilers  are heating value of the fuel, moisture
content, ash content, particle size, waste type, and wood species. The moisture  content of
the fuel  is an  important parameter;  an increasing  moisture content value  results in a
decreasing net heating value due to increased water evaporation and a resultant  increase in
the flue gas volume. The moisture content of the fuel  depends on the type of debarking and
the storage time. Mechanical debarking (10 to 30 percent water by weight) results in a lower
moisture content than hydraulic debarking (40 to 60 percent water by weight).

The net heating value for most waste woods is about 17 to 21 MJ per kg of dry wood (7,200
to 9,000  BTU/lb), or  8.4  to 13 MJ/kg (3,600 to  5,400 BTU/lb)  on an as-fired basis. The
heating value for waste woods tends to vary  with  wood species. The presence of extractive
materials, such as terpenes and tall oils, can substantially add to the energy content of the
wood.

The particulate emissions from waste wood combustion vary with ash content and particle
size of the material  being burned.  The ash  content can vary from below 1 to 20  percent by
weight on an as-fired  basis.  The sizes and shapes of the wood particles being burned can
influence the design of grating systems, the  type of firing employed,  and the  relative
distributions of underfire and overfire  air in the furnace. Furnace fouling is often found in
bark firing, especially  when bark  is fired together with other fuels. In some cases, small
                                         16-12

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amounts of minerals, gathered in the fuel  during storing in  sea water  or  on ground, can
lower the ash softening temperature so that it is very sticky iri the furnace and cannot be
easily removed.

Thw  major types of combustion units employed for waste wood burning  employ pile
burning or  suspension burning.  The successive  operations of water evaporation, volatile
carbon distillation and oxidation, and fixed carbon oxidation must occur  in series. Flat grate
Dutch-oven boilers have  been extensively  used  for  waste wood combustion in the past.
These employ both underfire- and overfire-air jets as applied to a stationary grate and do not
normally require extensive pregrinding of the fuel. Thin bed suspension firing is often used
for combination firing of wood and other fuels. It allows higher firing rates and is normally
employed for large  units. A disadvantage is that it  normally requires  pregrinding of the
wood in a hammermill.

Major operating variables affecting the combustion reactions in wood-fired power boilers are
the excess air level, the ratio of overfire to  underfire air, the  combustion temperature, and
the pile or suspension bed thickness to regulate fuel burning rate. The overfire air jets should
be located and operated to produce a minimum of entrained particulate matter.

The operation of a grate firing should be as even as possible for the best efficiency, so that
its  combustion  air can be correctly regulated.  A grate  firing in  a separate combustion
chamber can be operated with substoichiometric air flow, if the furnace temperature must
be kept low, or if the furnace atmosphere must be reducing. If neither of these conditions is
required,  the  grate must have the proper excess  air to burn combustible gases before they
enter the main furnace.

The major  potential air pollutant from  wood-fired  power boilers in the  pulp and paper
industry is  particulate matter, which can result  from either  inorganic ash in the wood or
from incomplete combustion.

The  particulate  matter  from  wood-firing is often large  in size, 5  to 10 /zm  (2 to
4 X  10~4  in) or greater. Its  specific gravity is usually low so that the  use of mechanical
cyclone collectors is not always possible. The electrical properties of the  fly  ash are not very
suitable for electrostatic precipitation because of the high carbon content that causes high
particle resistivity. The minerals in the wood can cause abrasion in the collecting equipment
and ducts resulting in rapid metal wear (3). More complete analysis and classification of the
chemical  composition and physical size characteristics of particulate matter  emitted from
wood-fired power boilers is needed than has been  reported to date.

Potential gaseous emissions from wood-fired power boilers are oxides of nitrogen, oxides of
sulfur,  and hydrocarbons resulting from volatilization  of the wood. The nitrogen oxide
emissions from wood-fired power boilers  are generally lower than for  fossil fuel firing
                                         16-13

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because of the large combustion volumes per unit amount of fuel burned, the normally high
excess air levels of 50 percent or more, and the high fuel  moisture content that results in
low  flame temperatures  of 980  to  1,200° C (1,800 to 2,200° F). Emissions of S02
from  wood-fired power boilers generally are low because the sulfur content of wood is
generally less than 0.1  percent  by weight.  The emissions  of terpenes, hydrocarbons, and
other volatile organic constituents by  distillation and incomplete combustion vary with
wood species, furnace temperature, and retention time. The extent  of these emissions as
potential air pollutants has not been fully described.

16.4   Particulate Emissions

Particulate emissions from power boilers consist of inorganic ash from the fuel and partially
burned or unburned fuel from combustion  processes. Both of these components vary
considerably  with the  type  of fuel  and firing.  The relative amount of organic material
present depends primarily on the excess air level and retention time in  the combustion zone.
The ash contents of most fuels used in the pulp and paper industry normally are less than 5
percent by weight, with the exception of coal, which normally varies from  5 to 15 percent
by weight. The fly ash in coal consists mainly of quartz, aluminum oxides, iron oxides, and
alkali oxides  (17). Smaller amounts of trace metals and other materials also may be present.
The composition of the ash in coal, wood, or oil depends  largely  on the source of the fuel.

The  amount  of unburned carbon in  the form of soot and grit is very dependent on  the
quality of oil firing; while fly ash from the firing of coal and wood normally  contains only a
few percent unburned carbon. This smaller  amount again depends on the quality of firing,
on excess  air addition,  on the ratio of primary  to  secondary air, and especially  on  the
behavior of the fuel on the grate.

Particulate emissions  from power boilers normally do not  contain large quantities of trace
metals or organic hazardous materials (18).

The  control  of  particulate  emissions  from  coal- and wood-fired  power boilers is  not
normally as complex as other processes within the pulp and paper industry. The problems of
control are more similar to those found in the power generation field,  and the approaches to
their  solution  are  the  same. Electrostatic  precipitation is the most efficient method  for
particulate control for coal-fired power boilers-, cyclone collectors offer a less efficient,  but
also  less expensive, way for either coal- or wood-fired boilers. Liquid scrubbers or fabric
filters are  effective for particles  below  5 jum (2 X 10~4  in) and offer  an alternative to
electrostatic precipitation, particularly for wood-fired boilers. Typical particulate emissions
from power  boilers in  the pulp and  paper industry  are presented in Table 16-4 (3),  and
typical particle  size characteristics for coal- and wood-fired power boilers are presented in
Table 16-5 (3).
                                         16-14

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                                                 TABLE 16-4
             PARTICULATE EMISSION CHARACTERISTICS FROM SELECTED  U.S. POWER  BOILERS (3)


Number
of
Boilers
Percent of Fuel Supplied,
Btu Basis
Coal
Oil
Gas
B/W*
Type**
Pressure
Drop
in. of water
Particulate Concentration
Inlet
Outlet
Collection
Efficiency
g/m3 (gr/cu ft) %




ON
1— i
en



18
2
2
16
2
2
2
3
2
100
100
100
0
75
0
73
0
0
0
0
0
46
0
0
16
25
0
0
0
0
0
0
62
0
39
0
0
0
0
54
25
38
11
36
100
C
S
P
C
C
C
C
C
C
2.5
—
2.2
2.7
3.9
2.8
2.5
2.8
0.2
4.28

11.2
7.9

5.3

4.3
3.2
(1.87)

(4.89)
(3.47)

(2.30)

(1.88)
(1.40)
0.85 (0.37)
0.57 (0.25)
0.98 (0.43)
1.05 (0.46)
0.41 (0.18)
0.39 (0.17)
2.75 (1.20)
0.71 (0.31)
0.89 (0.39)
80
—
91
87
—
93
—
84
72
Emission
Rate
kg/h (Ib/hr)
129 (284)
136 (300)
180 (397)
140 (309)
73 (160)
70 (153)
228 (502)
202 (445)
44 (96)
 *B/W = bark and wood waste.
**C = Cyclone.
  S = Liquid scrubber.
  P = Electrostatic precipitator.

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                                            TABLE 16-5
TYPICAL PARTICLE  SIZE  DISTRIBUTION OF FLY ASH FROM COAL- AND  WOOD-FIRED POWER BOILERS (3)

                                         Percent by Weight of Particles in a Given Size Range
Coal-Fired Power Boilers
Particle
Diameter
/im (in)
0-10(0-4 X 10~4)
10-20(4-8 X 10~4)
20-30(8-12 X 10~4)
30-40(12-16 X 10~4)
40-75(16-30 X 10~4)
75-150 (30-60 X 10~4)
150+ (60 X 10-4+)
Pulverized
Coal

25
24
16
14
13
6
2
Cyclone
Furnace

72
15
6
2
—
5
—
Spreader
Stoker

11
12
9
10
12
17
29
Traveling
Grate

_
—
11
—
12
30
47
Underfeed
Stoker

7
8
6
9
8
19
43
Bark-Fired
Power
Boilers

12
10
7
6
14
16
35
 Total                   100           100           100           100            100            100

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     16.4.1   Electrostatic Precipitators

Electrostatic precipitators are used primarily for coal-fired power boilers in the pulp and
paper industry. Electrostatic precipitation is suitable for particulate emission control where
high collection efficiency is required, but only a small pressure drop  can be tolerated. The
capital  cost of electrostatic  precipitators is  relatively high, and electric power is consumed
during  their operation. The particulate  collection efficiencies  are  lower  for  oil-fired or
wood-fired boilers because the carbon content of the fly ash causes considerable  increases in
particle resistivity as compared to coal-fired  units.

The design of  electrostatic precipitators depends on gas flow rate, gas temperature, gas
humidity,  inlet dust loading, and the electrical properties of dust, such as particle resistivity.
The resistivity  essentially  defines the  design  migration  velocity of the  particles to the
collection  electrodes. The gas flow rate determines the design gas velocity, which defines the
necessary width of the precipitator for  operation. The particle size distribution in the inlet
dust has minor influence  on the design of the precipitator, unless a major portion of the
dust  is less than 1 pm (3.9 X 10~5 in) in diameter. At such  small  particle sizes, the
decreasing particle collection efficiency must be compensated for by reduced gas velocity.
The precipitator  reduces  the particle size  distribution by selectively removing the larger
particles.  On the other hand, when a  precipitator failure occurs, the resulting particles
emitted are selectively both larger and heavier. This operating characteristic should be taken
into account when selecting  continuous monitoring equipment (19).

The factor which varies the most in the operation of  an electrostatic precipitator is the gas
velocity or gas flow rate, which varies with boiler load and,  in combination boilers, with
fuel-to-fuel ratio. Smaller  variations in the gas velocity are caused  by changes in the gas
temperature and  in the humidity of the  gas. Temperature  and humidity are treated as
independent variables in determining the efficiency of the  precipitator. The humidity of the
flue gas depends on the moisture and hydrogen content of the fuel. The quality of the fuel
and the firing together cause variations in the unburned carbon content of the  dust, which
influences  the precipitator  efficiency and the inlet dust loading. Electrostatic precipitators
can be applied to the combination firing of coal and bark.

     16.4.2   Cyclone Collectors

Mechanical cyclone  collectors  are  used primarily  for particulate  emission  control  on
wood-fired power boilers,  or as first stage collectors for coal-fired units. In a cyclone
collector, the flue gas with its dust burden  is fed into a centrifugal field, where centrifugal
forces separate dust particles from the flue  gas. Single cyclones or multicyclone equipment
is  used. Multicyclones consist  of a group of  small cyclones arranged  in  parallel.  The
minimum  particle size collected by a single cyclone  is about 10/zm (4 X 10~4 in); while
multiple cyclones will collect 5 jtzm (2 X 1CT4 in) particles. A separated flow with high dust
                                          16-17

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content may further be conducted into a secondary separation with cyclones,  and the flue
gas recirculated to the main collector.

A cyclone collector has a lower capital cost than an electrostatic precipitator and may also
be applicable to coal- and oil-fired power boilers. Because of the energy required to maintain
centrifugal movement, the pressure drop across the cyclone is inevitably high.  Fouling can
be  serious, especially in multicyclone equipment, where gas and dust  must pass through
rather small ducts. Therefore, frequent maintenance of equipment may be required.

The design of a cyclone  collector depends on gas flow rate, gas temperature, gas humidity,
inlet dust loading, density  of particles, and particle size distribution. The abrasive properties
of particles may set special requirements for materials used in construction of cyclones and
associated ductwork; while the  temperature and humidity of gas have their effects on
fouling, particularly for bark char or sander dust combustion (20).

The gas flow rate from a power boiler  defines the size of a single cyclone or the number of
cyclones required in multicyclone equipment, depending on the inlet velocity required for
effective separation of particles from the gas stream by centrifugal forces. This velocity, in
turn, depends on the density and size distribution of particles  in the flue gas stream from
the power boilers. The circulating  movement can be introduced to the  gas  either by
tangential inlet or by radial vanes with axial inlet. The central outlet tube for clean gas may
be  located either straight ahead in a horizontal cyclone or  at the inlet end of a vertical or
inclined cyclone,  causing a  180°  turn and additional separation. A secondary circuit, with a
fan to induce suction, may  be connected into the separated dust outlet.

A single- or monocyclone-collector is applicable for rather coarse dust and small flow rates.
A small boiler firing woodwaste and bark is a typical application. A multicyclone collector
allows  for large flow rates by adding to the number of cyclones. The cyclone size can be
optimized for required separation and pressure drop.

The separation in a cyclone  depends strongly on the centripetal acceleration of the gas and,
therefore,  on the gas flow  rate. This must  be kept  within acceptable limits to achieve
operation between inadequate separation at low velocity and  too high a pressure drop at
high velocity. Flow rate can be  controlled by shutting off certain parts in a multicyclone
collector; otherwise, a low collection efficiency must be expected at low boiler loads.

The particle size distribution may change along with boiler load, different fuels, and other
changes in firing.  Therefore, the collection efficiency of a  cyclonic collector also changes.
There is no general rule to control the particle size distribution during firing except to try to
maintain relatively uniform conditions.
                                         16-18

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     16.4.3  Liquid Scrubbers

Liquid scrubbing is used primarily for particulate collection on wood- and bark-fired power
boilers in the pulp and paper industry. Scrubbing traps the particulate matter entrained in
the gas stream  in liquid for subsequent removal and disposal. Liquid scrubbing has  the
advantages of being able to remove gaseous and particulate  materials simultaneously,  of
removing fine particles below 1 jum (4 X  10 ~s  in) in diameter, of recovering additional
thermal energy by cooling the gas stream, and possibly of improving primary clarifier sludge
settling characteristics. Vertical impingement and venturi scrubbers have been the primary
types of  unit employed, to  date.  Successful installations were reported in Montana by
Effenberger (21), in Texas by Ritchey (22),  and in South Carolina by Pearce (23).

Several  major variables affect the design of liquid scrubbing systems for particulate emission
control on wood-fired power  boilers. Major gas stream variables include the overall gas flow
rate, fuel firing rate, flue gas temperature and moisture content,  and the allowable scrubber
pressure drop  as determined by  fan capacity  characteristics. Major particulate matter
characteristics  include the  total  particulate mass  concentration  and the  particle  size
distribution, particularly  for  particles  less than  10 jum   (4 X 10~4 in)  in  diameter.
Liquid-phase design  parameters include the  makeup  and recycle  shower  rates,  liquid
pumping   capacity,   nozzle  configurations   and  sizes,  and   allowable  slurry  solids
concentrations.  The physical configuration and  gas-liquid contact geometry chosen  are
determined by  the type of scrubber purchased or designed. The materials of construction,
such as stainless steel, must be chosen to avoid corrosion, abrasion, and thermal damage.

The operation  of a liquid scrubber is influenced  primarily by the liquid phase parameters
such as recycle  flow rate, makeup water flow rate, nozzle liquid pressure drop, and slurry
solids concentration.  Gas phase pressure drop is usually subject to a certain amount  of
adjustment, but is also influenced by the gas flow rate as determined by fuel firing rates.
The scrubbing  systems described by  Effenberger (21) and  Ritchey  (22)  operate  at  gas
pressure drops  of about 15  to  25 cm (6 to 10 inches) of water  and are able to achieve
particulate mass concentrations at standard conditions of 0.02 to 0.05 g/m3 (0.01 to 0.02
gr/cu ft) or less, corresponding to  particulate mass removal efficiencies of 99 percent  or
greater. Detailed particle size measurements for  these units were not provided  on either
inlets or  outlets to  the respective collectors.  Liquid pH must be controlled to avoid
corrosion, particularly if coal or oil  is burned in combination  with wood. Liquid scrubbers
can prove particularly useful to upgrade the particulate collection efficiencies of existing
mechanical cyclone collectors on wood-fired power boilers.

     16.4.4  Fabric Filters

Fabric filters have not been extensively used for particulate collection on power  boilers in
the pulp and paper industry. They  have the advantages of high removal efficiencies for fine
                                        16-19

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particles of less than 1 ptm (4X10 s in) in diameter with pressure drops of 8 to 15 cm (3 to
6 in) of water. There is not enough operating experience with fabric filtration, to date, on
wood-, coal-, or oil-fired power boilers in the pulp and paper industry to present detailed
design and operating parameters.

16.5  References

 1.  Slinn, R. J., The Paper Industry's Energy: A Survey by the American Paper Institute.
     Southern Pulp and Paper Manufacturer, 37:39, March 1974.

 2.  Miller,  R. R.,  One Pulp and Paper Company's  View  of the  Energy Crisis.  Tappi,
     57:62-64, February 1974.

 3.  Hendrickson,  E. R.,  Roberson, J.  E.,  and Koogler, J. B., Control of Atmospheric
     Emissions in the Wood Pulping Industry, Volume II. Final Report, Contract No. CAP
     22-69-18, U.S. Department of Health, Education, and Welfare, National Air Pollution
     Control Administration, March 15, 1970.

 4.  Martin, G. B., and Berkau, E. E., Combustion Processes and Air Pollution.  (Presented
     at  the  National  Meeting of  the American Institute of Chemical Engineers. Atlantic
     City. August 30, 1971.)

 5.  Compilation of Air Pollutant  Emission  Factors.  U.S.  Environmental  Protection
     Agency, Office  of Air Programs, Research Triangle Park, North Carolina. Publication
     No. AP-42, February  1972.

 6.  Cuffe,  S. T., and Gerstle, R. W., Emissions from  Coal-Fired Plants: A Comprehensive
     Survey. U.S. Public Health Service, Cincinnati, Ohio. Publication No. 999-AP-35. 1967.

 7.  Fryling, G. R.,  Combustion Engineering,  Revised Edition. New York, Combustion
     Engineering, Inc., 1966.

 8.  Steam: Its Generation and Use, 38th Edition, The Babcock & Wilcox Company, New
     York, 1972.

 9.  Caron, A.  L., The  Control  of Particulate  and Gaseous Emissions  from  Coal-Fired
     Stationary  Combustion  Units.  National Council of the Paper  Industry for Air and
     Stream  Improvement. New York. Atmospheric Pollution Technical Bulletin No. 42.
     October 1969.
                                        16-20

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10.  Technology Transfer Process Design  Manual for Pollution Control in the Fossil Fuel
     Electrical Utility Industry. Prepared by Radian  Corporation, Austin, Texas, for  the
     U.S. Environmental Protection Agency, Washington, B.C., 1974. In press.

11.  Lowes, T. M., and Heap, M. P., The Emission of Oxides of Nitrogen from Natural  Gas
     and Pulverized Fuel Flames. International Flame Research Foundation. Ijmuiden, The
     Netherlands, Document No. D 09/a/82, 1972.

12.  Walsh,  R.  T., Boilers,  Heaters and Steam Generators. In: Air Pollution Engineering
     Manual, Danielson, J.  A.  (ed.).  U.S. Public Health  Service.  Cincinnati,  Ohio.
     Publication No. 999-AP-40,1967.

13.  McGuire, W. F., Thompson, P. C., and Smith, L. L., Theory and Application of Nitric
     Oxide Emission Reduction in  Utility Boilers. In: Proceedings of the First Annual
     Symposium on Air Pollution Control  in the Southwest. Texas A&M University, College
     Station, Texas. November 5-7, 1973.

14.  Bagdon, K. M., Viscosimetry of Oil Burner Control.  Instrumentation  Technology,
     19:43-46, February 1970.

15.  Mineur, J., and Hulden, B., The Sulphur  Problems in Oilfired Boilers: A Review.
     EKONO OY. Helsinki, Finland. Publication Series No. 116. 1968.

16.  Smith,  W., Atmospheric  Emissions from Fuel Oil Combustion. U.S. Dept. of Health,
     Education  and Welfare,  Public Health Service Publication No. 999-AP-2. November
     1962.

17.  Tankha, A., Try Fabric Dust Collectors on Small Boilers. Power, 117(5):72-73, August
     1973.

18.  First  Draft Report, Group  of  Experts on Emission Measurement  Techniques  for
     Particulate  Matter  from   Selected Sources.   OECD.  Paris.  Addendum   1   to
     NR/ENV/73.25. 1973.

19.  Cooper, H. B. H., The Particulate Problem: Continuous Particulate Monitoring in  the
     Pulp  and Paper Industry. In: Proceedings of the Symposium on Instrumentation for
     Continuous Monitoring of  Air and Water Quality. Miami University, Pulp and Paper
     Foundation. Oxford, Ohio.  June 20, 1973.

20.  Barron, A., Studies on the Collection of Bark Char Throughout the Industry.  Tappi,
     53:1441-1448, August  1970.
                                       16-21

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21. Effenberger, H. K., Cradle, D. 0., and Tomany, J. P., Control of Hogged Fuel Boiler
    Emissions. Tappi, 56:111-115, February 1973.

22. Ritchey, J.  R., Venturi Wet  Scrubber for Particulate Control on a Bark  Boiler.  In:
    Proceedings of  the First  Annual  Symposium  on  Air Pollution Control in  the
    Southwest. Texas A&M University, College Station, Texas. November 6, 1973.

23. Pearce, A. E., Mechanical Dust Collection  with Secondary Wet Scrubbing as Applied to
    a Bark Fired Power Boiler. In: New Approaches to Particulate Collection at Bark Fired
    Power  Boilers.  National Council  of the  Paper  Industry  for  Air and  Stream
    Improvement, Inc.,  New York. Atmospheric Pollution Technical  Bulletin  No. 51,
    October 1970.
                                      16-22

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                                   CHAPTER 17

                              PROCESS MONITORING
17.1  Source Measurements

To  determine compliance with existing and proposed pollutant emission standards and to
inventory material losses, the types and amounts of gaseous and particulate materials from
pulp and paper mill flue gases must  be  measured. Developing methods for direct and
accurate  measurement  of air pollutant levels for TRS compounds, oxides of sulfur, and
particulate matter is important because of increasingly stringent emission standards in all
levels of government.

Source testing of particulate matter emissions from pulp and paper mill flue gas streams can
be performed by both  batch- and continuous-sampling methods. Batch testing provides an
average  concentration  value for a given time period. Continuous monitoring provides a
record  of instantaneous concentration  values over a prolonged time interval to determine
compliance with  air pollution regulations  and to act as a monitor of process equipment
operation.  Continuous monitoring instruments normally  necessitate a higher capital cost
than batch sampling equipment, but manpower requirements are normally lower once these
systems are placed in operation and maintained  by competent, trained personnel.

The successful operation of continuous  monitoring systems for measuring particulate matter
and  gaseous  emissions   requires  instrumentation  that  is  accurate,  reliable,  stable,
reproducible, of  simple  operation  and low maintenance  requirements,  and subject  to
minimal   interferences.  In  the  design   and   operation  of  continuous   monitoring
instrumentation,  both  the sample handling and detector systems must  be considered. A
suitable system for reducing and reporting the voluminous amounts of data that can  be
generated is also needed.

17.2  Gaseous  Monitoring

The major classes of gaseous pollutants emitted from pulp and paper industry sources that
may require continuous or batch monitoring are malodorous sulfur compounds, oxides of
sulfur, oxides of nitrogen, and organic  nonsulfur compounds. The major malodorous sulfur
compounds of  interest include H2S, mercaptans, dialkyl sulfides, and dialkyl disulfides.
These are commonly classified together  as TRS compounds.

The major oxides  of  sulfur include  SO2  and S03 from  combustion  processes where
sulfur-containing  fuel is burned. Oxides of nitrogen of interest include NO and N02 from
                                        17-1

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combustion  processes.  Organic  nonsulfur  compounds  include aliphatic,  olefinic, and
aromatic hydrocarbons, terpenes, phenols, and other organic compounds from kraft and
sulfite mill sources.

Gaseous monitoring systems normally have similar, and in some cases, common, sample
conditioning systems. The different gaseous constituents are monitored by gas detection
devices that depend on the constituent monitored.

     17.2.1  Sample Conditioning Systems—General

The purpose of the sample handling and  conditioning system is to remove the sample from
the flue gas  and transfer  it to the  detector for subsequent analysis without changing the
concentration or character of  the constituents to be measured. The major elements of the
sample  conditioning system are the sample probe, the liquid scrubber, the preselective gas
separator, the transfer tubing, and the prime mover, as shown in Figure 17-1 (1).

     17.2.1.1   Sample Probe

The sample probe  is located internally within the stack to remove a portion of the moving
gas stream  from the duct into the sample  conditioning system. A straight stainless steel tube
with a 90  degree bend is sufficient for flue gas streams containing negligible quantities of
particulate matter,  such as digester, washer, evaporator, and black-liquor oxidation tower
gases. The  probe  should be aligned  so that any condensate formed drains away from the
stack towards the scrubber.

The particulate matter present in flue gases from recovery furnaces, lime kilns, and smelt
tanks must be removed by one of two types of filtration systems. The system devised by the
National Council for Air and Stream Improvement employs an open end  25 mm (1 in)
diameter tube  packed  with glass wool  for particle removal that  must be cleaned and
repacked periodically to prevent plugging (2).

The system devised by Thoen  (3) is designed to prevent plugging by using a porous ceramic
probe equipped with a compressed air blowback feature that is actuated by solenoid valving
and a timer for 30 second intervals once each 10 minutes.  The ceramic probe is normally
employed in particulate laden  gas streams above their dew points where condensation is not
likely,  such as in power boilers and kraft recovery furnaces. Where significant quantities of
S03  are present, such as from oil-fired power boilers, the sample probe should be heated to
150° C (300° F) or higher  to prevent condensation of sulfuric acid.

         17.2.1.2  Conditioning Device

The major  purpose of the conditioning device is to  prevent condensation of water vapor in
the gas stream. One approach calls for passing the gas stream from the sample probe to the

                                        17-2

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            I
                                                 AMPLIFIER
RECORDER
OJ
INLET
PROBE
LT-H.
/^_- ., "i
fl
— ^
L^ r






— ==^—1
SAMPLE
CONDITIONING





1 	 ,.





DETECTION FLOW FLOW
SYSTEM CONTROL MEASUREMEr*
f-^r
i 11

JT PRIME
MOVER
            T
                                               FIGURE 17-1

                      GAS SAMPLE HANDLING & CONDITIONING SYSTEM FOR EXTERNALLY
                            LOCATED CONTINUOUS  GASEOUS  MONITORING SYSTEM

-------
detector through a line heated to a temperature above its dew point. Maintaining the sample
lines at above  110° C (230° F) is normally  necessary.  Inlet sample probes must also be
heated to 150° C (300° F) or higher to effect evaporation, particularly if there are mist
droplets entrained in the gas stream, such as sulfuric acid droplets.

A second approach is to employ some type of condenser in the sample line downstream of
the probe. The condenser serves  the dual functions of removal of the water vapor and
cooling of the gas stream before it enters the detector. The condenser can effectively remove
the water if the gas stream is cooled to  30° C (86° F) or lower, but it is normally necessary
to acidify the condensed  moisture to a pH  of 2.0 or less by adding H2S04 to minimize
absorption of SO2.

Under  some circumstances the  gas  stream  should be diluted with air upstream of  the
detector to cool the gas stream or to  prevent condensation. To control the air flow rate into
the system accurately, the exact degree  of dilution must be known. This dilution technique
is  not  suitable for extremely low gas  concentrations approaching the sensitivity of  the
detector.

         17.2.1.3   Prime Mover

The remaining  portions  of  the  gas handling  and conditioning system  are  the flow
measurement and control section, the drying section, and the prime mover. A desiccant,
such as silica gel or Drierite, can  be used to protect either the flow meter or the vacuum
pump  downstream  of the detector, but it is not commonly  used. Micrometering needle
valves made of stainless steel  can be  used to control the gas flow rate, since they are able to
resist the corrosive gas conditions. Flow metering is done with a rotometer or orifice flow
meter.

The prime mover can be located either  upstream  or  downstream of the detection  cell
depending on the system in use. A positive displacement vacuum pump that is leakproof and
sealed  may be put in use upstream of the detection unit. Upstream location of the vacuum
pump is particularly desirable if a  considerably larger volume of stack gas is removed from
the duct than is sent to the detector,  or if  the detector must be operated under positive
pressure. Such an arrangement does involve potential problems of particulate plugging and
moisture condensation, as well as sample dilution by air leakage. Location of the prime
mover downstream of the detection  cell allows use of either a mechanical vacuum pump, or
a steam, air, or water aspirator.

          17.2.1.4   Sampling Lines

Sampling lines should be of sufficient diameter to provide  for minimum pressure drop,  but
small enough for minimum retention time. Tubing of 0.635  cm (0.25 in) to 1.27 cm (0.5 in)
                                        17-4

-------
is normally optimum for sample line construction. Wall materials of inert polyethylene or
Teflon should be used to avoid possible losses by physical adsorption or chemical reaction.

Also, electrically heated Teflon tubing is commercially available for sampling lines of up to
60 m (200 ft) in length. Tygon and rubber tubing, plus carbon steel, cast iron, and copper
fittings all react with sulfur compounds and should not be used.

     17.2.2  TRS Monitoring Systems

The major malodorous sulfur compounds of interest for continuous monitoring applications
in kraft pulp mills include H2S, CH3SH, CH3SCH3, and CH3SSCH3. Major sources of H2S
from the kraft process include the recovery furnace, smelt tank, lime kiln, multiple-effect
evaporator, and tall oil vent gases. Major sources of the organic sulfur compounds include
the  digester blow and  relief gases,  brown stock  washer  hood and seal tank vents,
multiple-effect evaporator noncondensable gas and condensate liquid  streams, black liquor
oxidation  tower  exhausts,  and the  recovery  furnace  used following  direct  contact
evaporation.  The organic sulfur compounds can  often create  problems in sample handling
because of their tendency to condense and adhere to tubing walls.

          17.2.2.1  Gas Conditioning

Accurately transferring TRS compounds from the flue gas to the detector is complicated by
the presence of large quantities of  water vapor  and particulate matter in the flue gas at
elevated gas temperatures. Terpenes, S02, and organic sulfur compounds can be interfering
constituents in coulometric detection systems.

It is first necessary to remove sulfur dioxide from combustion sources such as recovery
furnaces, lime kilns, and smelt tanks. A  liquid scrubber is located immediately downstream
of the probe which contains a solution of potassium acid phthalate (KHC8H404). The
purpose of the scrubber is to selectively remove sulfur dioxide from the flue gas, condense
excess water vapor, remove additional particulate matter, and cool the gas stream from stack
to  ambient temperature  conditions.  Two different  types of potassium acid phthalate
scrubbers can be used, the continuous flow and the batch, nonflow types.

The NCASI  continuous flow scrubber employs  a two chamber system containing liquid
absorption and overflow chambers  in series. A  three percent solution of potassium acid
phthalate  passes from a storage  bottle  through  a glass wool filter in the bottleneck to
remove  particulate  matter to prevent plugging (2). The liquid flow rate to the scrubber is
controlled  by a  limiting  flow orifice  of capillary  tubing at a  rate between  0.5 and
1.0 cm3/min, which  requires replenishment  once per week.  Gas is  drawn at a rate of
25 cm3/min through the scrubber of 40 cm3 capacity. A 1.0 m (3 feet) long dropleg is used
to maintain suction and prevent leakage.
                                        17-5

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TRS losses,  as H2S, are  between 0.1 and 0.2 ppm by volume  at a liquid flow rate of
1.0 cm3/min because of the finite solubility of the compounds in the KHC8H404 solution.
However, the TRS losses increase considerably  with increasing liquid flows above this rate.

The scrubbing system developed by Thoen (3) is a batch system employing a  saturated
solution of KHC8H404 which requires replenishment once every 30 days and  which has
lower total reduced sulfur losses once the system reaches equilibrium. As the solution is
depleted, S02 removal  is  less  efficient,  and may cause interferences  in the  detector.
Condensation at water vapor in the system may  also cause malfunctions in the system.

After the S02  scrubber,  the gas stream passes through a system employing preselective
scrubbers where analytical selectivity can be obtained between the  respective reduced sulfur
gas constituents. Thoen (4)  developed scrubbing solutions which could be used for selective
removal of  H2S, H2S  plus CH3SH, and H2S plus  CH3SH  plus CH3SCH3. Respective
constituent  concentrations  are  monitored by differences in instrument readings. The
solutions used are listed in Table  17-1.

                                   TABLE 17-1
  SELECTIVE PRESCRUBBING SOLUTIONS FOR SULFUR GAS  SEPARATION (4)

                                         Scrubbing
     Gases Removed                       Solution                   Concentration
                                                                       % by wt.

          S02                           KHC8H4O4                        3
       H2S + S02                      CdSO4 - H3B03                     1-2
    H2S + S02+RSH                       NaOH                          10
S02 + H2 S + RSH + RSR                    AgNO3                         0.5
An additional  procedure used on certain gas conditioning systems is to pass the flue gases
through a combustion furnace  at a temperature of approximately 815° C (1500° F) to
convert the reduced sulfur compounds  to SO2 upstream of the detector (5). Conversion of
the reduced sulfur  compounds  to  S02  makes it feasible to  use S02 sensitive detection
methods,  such as flame photometry  and ultraviolet spectrophotometry  in addition to
coulometric  titration. An additional advantage of the combustion process is that it converts
potentially interfering  organic compounds, such as terpenes and  olefinic and  aromatic
hydrocarbons to nonreactive C02 and water. The combustion step is  necessary for kraft  mill
sources containing extensive quantities of terpenes, such as digester blow and relief gases,
washer hood and seal tank vents, and black-liquor oxidation tower exhausts.
                                       17-6

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          17.2.2.2   Gas Detection Systems

Available detectors for continuous monitoring of reduced sulfur compound levels in kraft
pulp mill process streams include coulometric titration, electrochemical membrane sensing,
flame photometry, and ultraviolet spectrophotometry. TRS monitoring is complicated by
interference from S02 and by the sensitivity of some detectors to only H2S and S02.  It is
often either necessary or desirable to convert the organic sulfur compounds and H2S to S02
by  oxidation upstream of the detector, particularly when using  detection methods other
than coulometric titration.

Coulometric titration is useful for measuring concentrations of H2S and organic sulfur gases
in kraft pulp mill flue gases; but it is also sensitive to olefinic and aromatic hydrocarbons,
terpenes, acrolein (CH2CHCHO), and N02 (6).

The technique  operates on the principle that the H2S, S02, and organic sulfur compounds
present are oxidized to sulfate ion by the action of a halogen electrolyte. A certain electrical
current is required to maintain a constant concentration of halogen gas generated from the
electrolyte. The current level is proportional to the overall concentration of reactive gases
passing through the detection cell (7). Coulometric titration is sensitive to all  compounds
that can be oxidized by the halogen in varying degrees peculiar to each compound and is not
specific to any  one compound.

A commercially available instrument employs a solution  of  16 percent hydrobromic  acid
(HBr) as the electrolyte with sensing,  generating, and reference electrodes  all located on a
common shaft  (8). The detection cell can be actuated to generate a series of fixed levels of
bromine  gas   (Br2)  by  changes  in   instrument  attenuation setting corresponding to
concentration ranges varying from 1 to 800 ppm full scale as H2S.

Flue gas  is  drawn  into  the  cell where the  presence  of reactive gases is  detected by the
consumption of bromine gas generated as the result of the chemical oxidation reactions.  The
current required for maintaining a constant  bromine level in  the cell increases  as the total
concentration of reactive gases increases. The current is converted into an equivalent voltage
potential  in an amplifier  and the signal is transmitted to  a continuous printout on a 0 to
100 mv recorder.

Calibration of  individual cells for particular gases to be  measured  is  necessary to assure
accurate results. The reason  is that the  response  of the  instrument to the  gases passing
through the cell depends both on their concentrations and on the valence states of the sulfur
atoms  relative  to bromine. Each gas then must be calibrated individually to  determine its
relative response in the cell. The response to particular gases also varies between individual
detection cells  because of differences in  hydraulic characteristics, background noise level,
and quality control. A summary of approximate calibration factors for sulfur gases observed
for a Barton coulometric titrator is listed in Table 17-2 (9).

                                         17-7

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H2S
0.008-0.013
0.030-0.035
0.09-0.12
0.24-0.26
0.7-0.9
1.5-2.5
4.5-9.0
CH3SH
0.014-0.016
0.040-0.045
0.12-0.13
0.35-0.38
1.2-1.4
3.0-3.5
10.0-13.0
CH3SCH3
0.035-0.040
0.090-0.100
0.31-0.37
1.0-1.2
3.5-4.0
11.0-12.0
35.0-40.0
CH3SSCH3
0.030-0.035
0.09-0.12
0.25-0.30
0.5-1.0
2.0-3.0
5.0-8.0
20.0-25.0
S02
0.03-0.04
0.09-0.12
0.30-0.35
0.9-1.0
2.5-3.5
8.0-10.0
25.0-30.0
                                    TABLE 17-2
   APPROXIMATE RANGES IN CALIBRATION FACTORS FOR SULFUR GASES
                       WITH  COULOMETRIC TITRATOR (9)

                          Concentration Factor (ppm by volume/scale unit)
 Attenuation

      0.1
      0.3
      1.0
      3.0
     10.0
     30.0
    100.0
Cells should be calibrated on a weekly basis to assure continued accuracy, particularly for
high concentration levels (above 50 ppm by volume as H2S) because of cell response drift.
The electrolyte solution must be changed at least once per month to avoid depletion. The
system also appears to require frequent maintenance and suffers from particulate plugging in
the cell. Three possible arrangements for the system are illustrated in Figure 17-2.

Electrochemical membrane  cells that are specific for either S02 or H2S can be used for
monitoring reduced sulfur gas emissions. It is  normally necessary to convert the reduced
sulfur compounds to SO2 in an oxidation furnace upstream of the S02 selective membrane
cell. The H2S  selective  membrane cell cannot be used to measure TRS levels in kraft pulp
mill flue gases because of interferences from organic sulfur compounds.

The principle  of operation  is similar  to that of coulometric titration. SO2 operates as the
working material in a specially constructed electrochemical transducer cell of proprietary
design. The detection cell is a totally  enclosed system where the gas sample passes adjacent
to a semipermeable plastic membrane  across  which the SO2  or H2S diffuses into the
electrolyte solution. The gas produces a change in the electrochemical potential across the
cell that is directly proportional to the concentration in the gas stream over a concentration
range from 0.01 to 5,000 ppm by volume.

The gas handling system  for the electrochemical membrane cell is similar to  that of the
coulometric titrator as shown  in Figure  17-3. The  two major differences are  that it is
necessary  to locate  a  leakproof vacuum pump  upstream of the detector  because the
membrane cells must be operated under positive pressure, and a combustion furnace is
needed  to oxidize the  reduced sulfur compounds to SO2 • Parallel detection cells can be
located to record both S02 and TRS simultaneously. The gas flow to the detector cell is
approximately 500 cm3/min.

                                        17-8

-------
 A.  NCASI   SYSTEM(l):
                                          RECORDER
                                          (0-100mv)
                                       CONTROL
                                        MODULE
                                            FLOW
                                             ETER
                                     ^^ I         \\vvvu\\
                                                          DRYING
                                                          TUBE
                                                   VACUUM
                                                   PUMP
                                 DETECTION
                                 CELL
 B.  WEYERHAEUSER SYSTEM(2):
                                         RECORDER
                                         (0-100mv)
SURGE      S02
BOTTLE   SCRUBBER
                                 DETECTION
                                   CELL
                            FIGURE 17-2

CONTINUOUS SOURCE MONITORING SYSTEM FOR  REDUCED  SULFUR
            EMISSIONS  WITH  COULOMETRIC TITRATION
                                 17-9

-------
        C.  HUMBOLT  COUNTY SYSTEM(4):
                                          RECORDER
                                          (0-100mv)
                                        CONTROL
                                        MODULE
                                            FLOW
                                            METER
                                                    VACUUM
                                                    PUMP
                                     LTl
                                  DETECTION
                                   CELL
                               FIGURE 17-2

  CONTINUOUS SOURCE  MONITORING SYSTEM  FOR  REDUCED  SULFUR
          EMISSIONS  WITH  COULOMETRIC TITRATION-CO/V77My£D
                                                           ECORDER

                                                          CONTROL  MODULE

                                                           riROTOMETER
                                                   DETECTION
                                                      CELL
     t
LIQUID   I
SCRUBBER
        IDROPLEG
                                   VACUUM
                                    PUMP
                               FIGURE  17-3
   TOTAL REDUCED SULFUR MONITORING  WITH  AN ELECTROCHEMICAL
                       MEMBRANE CELL DETECTOR
Spectrophotometric methods are used to measure SO2 and must be used with an oxidation
furnace upstream of the detection device to convert the reduced sulfur compounds to S02.
Methods of detecting S02 are discussed in section 17.2.3.


Some reduced sulfur compounds may be monitored by gas chromatographic methods, as
discussed in section 17.2.5.
                                    17-10

-------
     17.2.3   Sulfur Dioxide Monitoring Systems

SO2  can be released to the atmosphere from either kraft or sulfite pulp mill process sources
or from the combustion  of  sulfur-containing fuels in coal- or oil-fired power boilers. Major
potential sources of S02  emissions in the kraft process include the recovery furnace, lime
kiln, and smelt tank. The major potential sources of S02 in sulfite pulp mills include the
digester, evaporator, and acid-making stage vent gases, and the recovery furnace.

A small amount  of the S02  formed can be subsequently  oxidized  to S03, which is
converted to H2S04 mist at temperatures sufficiently below the acid dew point. Normally,
the major constituent to be monitored in flue gas streams in S02. The S02 concentrations
can  be monitored either within a  duct  or  externally  following sample collection and
withdrawal.

          17.2.3.1   Sample Conditioning

The  major constituents making sampling difficult are the large quantities of water vapor and
particulate matter in flue gas streams. The S02  can be easily removed from the inlet gas
stream during condensation of water  because of its relatively high solubility. The S02 can
also  react with certain tubing wall materials, resulting in additional loss of S02. Therefore,
design of the sample handling and conditioning systems must be carefully made to avoid
such' losses.

          17.2.3.2   Detection Systems

Detection systems used for continuous monitoring of S02 include electrolytic conductivity,
electrochemical transducer conversion, coulometric titration, and ultraviolet spectrometry.
The  first three are all located externally from the stack and require prior sample withdrawal,
while  ultraviolet  spectrophotometry  can be  performed either internally  or externally.
Internal location  of the detector eliminates any possible problems caused by air leakage or
moisture condensation that might occur during sample withdrawal. External location is the
only feasible approach for certain detectors and results in fewer problems with particulate
interferences and high gas temperatures.

Miller, Brown, and Abrams (10) describe a continuous monitoring system  for measuring
total sulfur  oxides  (S02 and S03)  in the flue  gas  streams  from digester blowpits and
acid-making  absorption towers  in sulfite pulp mills.  The stack gas is withdrawn from the
duct and passed  through a water scrubber. The soluble gases dissolve and pass through an
externally located conductivity cell. The relatively simple, inexpensive nonspecific detector
is suitable for measurement  of sulfur oxide  levels  in  the  highly moisture-laden sources
because  only negligible  quantities  of  interfering,  conductivity-producing  gases  and
particulate matter are present in the gas stream.
                                        17-11

-------
The gas is withdrawn from the duct at a rate of 3 1/min (0.1 cfm) and mixed with deionized
water at a rate of 1 1/min (0.3 gpm). The gas-liquid mixture passes concurrently downward
through a 5.1 cm (2 in) diameter tower 66 cm (26 in) long for S02 absorption and then into
a separation  chamber where the gas stream is  withdrawn through a  vacuum pump. The
liquid stream containing the dissolved  sulfur oxides is then passed through a conductivity
cell that can measure S02  concentrations in ranges of either 0 to 1,000 or 0 to 10,000 ppm
by volume. A 10 mv recorder is used. The system, as illustrated in Figure 17-4, has proved
successful over extended periods of operation.

Electrochemical transducer membrane cells can  be employed for selective measurement of
S02  concentrations in  combustion unit and process source flue gases.  The principle of
operation was discussed in section 17.2.2.2.

Electrochemical  membrane cells  specific  for  SO2  are   relatively  free  from  chemical
interferences from oxides of nitrogen, S03, water  vapor, and hydrocarbons. They maintain
a relatively stable calibration without substantial drift in response for extended periods with
minimum maintenance. The major operating difficulties are that the presence of particulate
matter tends to plug the membranes, water vapor condensation in the detector causes erratic
response, and the presence of H2S04 mist can  cause severe corrosion. The detection cells
must be replaced about once per year as the electrolyte solutions become depleted.

A continuous S02 monitoring system that uses  an electrochemical transducer cell has been
described in the literature (11). The sample conditioning system consists of a ceramic heated
probe through which gas is withdrawn  at a rate of 500 cm3/min through a heated sample
     FLUE
     GAS
                  GAS FLOW
               —  30  1/min.
  WVTER
&	
                           PACKED
                           TOWER
                                       1-2 1/min.
                             .>  WATER
                                OVERFLOW
I
                                                                    GAS
                                                                    FLOW
                                       CONDUCTIVITY
                                            CELL
                                          r
                        MODULE
                                                          WATER
                                                          FLOW
                                    RECORDER
                                  FIGURE  17-4
     CONTINUOUS CONDUCTIVITY  MONITOR  FOR MEASURING  SULFUR
           OXIDE EMISSIONS FROM SULFITE MILL SOURCES (10)
                                      17-12

-------
line. The gas stream is pulled  by a leakproof stainless-steel vacuum pump with a Teflon
diaphragm  located upstream of  the detector because of the necessity  of operating the
transducer  cells under positive pressure. A refrigerator and  condenser are used to remove
water vapor as it is necessary to maintain a strongly acidic medium to avoid absorportion of
SO2.  The   gas  stream  then  flows into the  detection  cell  for  subsequent analysis,
amplification, and recording. The  system is illustrated in Figure 17-5.

The system has proved successful during extensive field use but requires replacement of the
modular transducer detection cells at intervals ranging from  6 to 18 months, depending on
cell design and S02 level in the flue gas stream.

Coulometric titration can be used for  monitoring S02  emissions from  coal-  and oil-fired
power boilers and sulfite pulp mill process sources where reduced sulfur compounds are not
present. The technique has been described  in  section 17.2.2.2. Coulometric titration is
nonselective for S02 and so organic olefins and other materials can interfere. It  is, therefore,
necessary to add a combustion furnace upstream of the detector in order to oxidize the
olefins from oil-fired boiler flue gases to  prevent their interference.

The method also requires removal of water upstream of the  detector to avoid flooding the
cell.
Ultraviolet spectrophotometry is useful for measurement of SO2 stack concentrations either
internally or externally. The method operates on the principle that the degree of ultraviolet
      t
I — <="^^
fi
"~^^ _
JL

L-i r
C
r^r=*
1 	 ' HEATED y~4
SAMPLE -* 	 *•
LINE VACUUM
PUMP
SAMPLE.
XWDITIONER

1
O/ATtTD
                                                              DETECTION
                                                                 CELL
                                                   RECORDER
                                  FIGURE 17-5
  ELECTROCHEMICAL TRANSDUCER  MEMBRANE  CELL FOR  CONTINUOUS
                     SULFUR  DIOXIDE MONITORING (11)
                                       17-13

-------
radiation absorbed  at some characteristic wavelength by passage through a gas  stream  is
proportional  to the SO2 concentration. The method is relatively specific for S02  if the
proper wavelength ultraviolet  light source is used, and it has only minimal interferences
from water and particulate matter. Three different modes of ultraviolet spectrophotometry
are used in commercially available instrumentation for continuous monitoring of S02 levels
in flue gas streams.

Thoen, DeHaas,  and Baumgartel  (12) describe the use of an internally located ultraviolet
emission spectrometer for continuous monitoring of S02 emissions from a magnesium base
sulfite  recovery  furnace  following  the  absorption  towers.  The  instrument employs  a
detection system where the ultraviolet radiation from a single beam mercury vapor  lamp at a
wavelength of 254 nm is passed across a duct through a cylindrical perforated fiber glass
tube 2 m  (6 ft)  long. Gas molecule interchange into and out  of the fiber glass tube  is
facilitated by a series of holes located at 90 degrees to the direction of flow; this orientation
inhibits large particles and water droplets from entering. The system is illustrated in Figure
17-6.

The system is suitable for measuring S02 concentrations from 80 to 4,000 ppm by volume
with an electronic output of 0 to 10 volts DC. The system has a minimum of interferences,
is resistant to corrosion, and employs no moving parts. The device is relatively insensitive to
low SO2  concentrations, displays sluggish response to rapid changes in concentration,  and
the fiber glass is not suitable for high gas temperatures. The mercury vapor lamp at 254 nm
wavelength does not correspond to the S02 maximum absorbance at 280 nm.
         Light
Control
Recorder
                                                                Module
          0-IOV D.C.
                                   FIGURE  17-6
         INTERNALLY  LOCATED  ULTRAVIOLET SPECTROMETER FOR
         SULFUR DIOXIDE MONITORING  IN FLUE  GAS STREAMS (12)
                                       17-14

-------
Saltzman  (13)  describes  the  use  of  an  externally located  dual  beam  ultraviolet
spectrophotometric analyzer for continuous monitoring of S02 levels in flue gas streams
from  coal- and oil-fired  power  boilers and sulfite  recovery  furnaces.  The gas stream is
withdrawn from the duct through a heated porous ceramic probe of 20 ju (7.9 X  10"4 in)
porosity  to  remove particulate  matter,  as shown in Figure 17-7.  The gas sample  passes
through a heated,  electrically traced  Teflon tube to prevent condensation of water  vapor.
The gas stream then passes through the photometric detection cell at a rate of 1 1/min and is
drawn through an  air aspirator maintained at a constant vacuum. A compressed air purge
located upstream of the detection cell is used to remove particulate  matter from the  probe.

The detection system is a dual beam photometer in which ultraviolet  light of 280 nm is
passed through the sample cell to provide for specific S02  absorbance. Visible light, with a
wavelength of 578  nm, is passed through the reference cell so that it is possible to minimize
the potential interference from NO2. The system is heated to prevent water condensation, is
rugged and durable under field conditions, and can maintain calibration on a stable basis for
extended periods. To prevent small particles from depositing on the  detector cell surfaces, a
glass wool filter is used in the sample line to remove them (14).
       Ceramic
                    Heated
                                   3*   Compressed
                                                                          Recorder
                                        Air  Purge
                                     Filter
       Filter
       1
      Flue
      Gas
                     Line
                                             Sample
                                             Cell
                              Liquid
                               Trap
Reference
  Cell
     Air
     Aspirator
                                  FIGURE  17-7
     INTERNALLY  LOCATED ULTRAVIOLET SPECTROPHOTOMETER FOR
        SULFUR  DIOXIDE MONITORING IN  FLUE GAS  STREAMS  (11)
                                       17-15

-------
An ultraviolet correlation spectrometer, located internally, is used for measuring SO2 in flue
gas streams (15). The system employs a cylindrical slotted probe placed in the gas stream
perpendicular to  the  direction  of  flow.  Ultraviolet  light  at  a series  of  wavelengths
corresponding to the absorbance characteristics of S02 are passed from the light source into
the stack and reflected from a mirror back into the detector.  The degree of absorption of
light is proportional to the S02 concentration in the flue gas and is indicated in ranges from
0 to  1,000 or 0 to 5,000 ppm by volume.

     17.2.4  Nitrogen Oxide Monitoring Systems

Oxides of nitrogen are released to the atmosphere from any combustion process because of
the reaction between 02 and N2 at elevated temperatures. Approximately 90 percent of the
oxides of nitrogen formed is NO  with the remainder being mostly N02. The  potential
sources  of  emissions of  nitrogen oxides to  the atmosphere  are coal-,  wood-,  oil-, and
gas-fired power  boilers, recovery furnaces and lime kilns in kraft pulp mills, and recovery
furnaces and sulfur burners in sulfite pulp mills.

The  possible constituents  to be monitored in flue gas streams include NO, NO2,  and total
oxides of nitrogen.  Continuous monitoring of oxides of nitrogen may be performed either
internally  within the stack or externally from outside the stack by sample withdrawal.
Oxides  of  nitrogen emissions from kraft and sulfite pulp mill  process sources are not
normally as significant as from power boilers because large quantities of water present in the
spent cooling liquors and lime mud inhibit the occurrence of high flame temperatures.

          17.2.4.1  Sample Conditioning

The  gas NO is relatively insoluble in water, but NO2  can be removed during condensation of
water unless strongly acidic conditions are  maintained.  For certain types of detectors, a
heated, electrically-traced inert Teflon  sampling line  can be used if condensation is to be
prevented  and  the  detector  kept  heated.  Prime  movers  employed  can be  either
corrosion-resistant and leakproof vacuum pumps or air,  steam  or water aspirators. The
features for sample handling and conditioning systems for oxides of nitrogen measurements
are similar to those employed for S02  systems, as described  in the previous section.

          17.2.4.2  Detection Systems

The  major detection  systems  employed for continuous oxides of nitrogen measurements
include electrochemical transducer membrane cells, ultraviolet spectrophotometry, infrared
spectrophotometry,  and   chemiluminescence.  Most  detection   systems employed for
continuous oxides of  nitrogen  measurements are located  external to  the stack and,
therefore, require sample conditioning systems.
                                        17-16

-------
 Electrochemical transducer membrane cell detectors are available for measuring either N02
 or total  oxides  of  nitrogen (NO plus N02)  levels in flue gas streams.  The  sample
 conditioning system requires withdrawal of the stack gas sample through a ceramic probe
 for particulate  removal,  and a heated line to prevent moisture condensation.  The gas is
 drawn  by a sealed, leakproof stainless steel vacuum pump via the sample  conditioning
 condenser into the electrochemical membrane detector. The concentration of N02 and/or
 NO is taken as being proportional to the change in electrochemical potential across the cell
 with  a readout  of 0 to 10 mv. Readable concentration ranges are zero to 500, 1,000, or
 5,000 ppm by volume.

 Ultraviolet spectrophotometry is useful for measurement of oxides of nitrogen levels in flue
 gas streams. The  detection principle is the same as for the S02 system described (section
 17.2.2.2) except that the wavelengths for the light beams to the sample and reference cells
 are 436 and 578 nm, respectively, where N02 is the chemical compound being measured. A
 reactor also converts NO to N02 at elevated temperature and pressure. The concentrations
. of N02 alone and N02  plus NO are read sequentially in a timed cycle.

 Infrared  spectrophotometry can also be used  for measuring of oxides of nitrogen. The
 principle  is the  same as for ultraviolet spectrophotometry except that characteristic infrared
 absorption peaks  for NO and NO2  are used. There is only limited field experience with the
 technique, to date.

 Chemiluminescence is  another technique for measuring oxides of nitrogen. Stack gas is
 withdrawn  from the duct, conditioned to remove particulate matter  and water vapor, and
 then passed to a catalytic chemical reactor to form N02  and 02. Light is produced by this
 reaction,  and the intensity of the light is proportional to the inlet concentration of NO. The
 NO concentration alone is determined by letting the sample (containing both NO and NO2)
 bypass the reduction chamber so that N02 is not reduced and, therefore, not detected.

      17.2.5  Gas Chromatography

 Gas  chromatography separates  constituents of  a gaseous mixture  by exploiting  their
 differences in relative affinity for  a given  packing material in a concurrent flow column.
 These differences cause the various constituents of the mixture to pass through the column
 at different rates. The gaseous components can then be individually  analyzed as they pass
 the  column exit by means of a suitable detector. Gas chromatography provides a versatile
 means of analyzing for a wide variety of compounds over a wide range of  concentrations,
 but only  for discrete samples and not on a continuous basis.

 The major elements of a gas chromatograph are the sample collection and handling system,
 the  sample injection system, the carrier gas flow  system, the separation column, and the
 detector. The primary variables in gas chromatography are the sample handling procedure,
                                         17-17

-------
the sample  size,  the  column dimensions, and packing, the column temperature,  and the
detector. Compounds of primary interest for measurement by gas chromatography  are H2S
and the organic sulfur compounds, plus other organic compounds such as hydrocarbons,
terpenes, and alcohols.

          17.2.5.1  Sample Handling

The  major  difficulties  in  sample handling and conditioning  systems are posed by the
presence of excessive quantities of particulate matter, water vapor, organic  and aqueous
mists  and droplets, and  pulp fibers (15). A cyclone separator that uses an enlarged sample
probe  with  a  splatter plate pointed downward  in the gas stream is necessary to  prevent
droplet entrainment in mist-laden gas streams. Porous ceramic or sintered stainless probes or
open tubes packed with glass wool can be used to remove particulate matter upstream of the
sampling lines for gas streams at temperatures above their dew points.

Water vapor is present in flue gas streams from most pulp mill sources in quantities ranging
from  20 to 95 percent by volume. The major  methods  for alleviating possible losses of
gaseous components to  be analyzed are to  heat  the sample probe, sampling lines, and any
collection vessels  by electric tracing to temperatures above the  gas dew points, or to dilute
the source gas with dry gas of known composition to below the level at which water will
condense at  the given temperature.

A possibility that must be  anticipated and allowed for in the design  and construction of gas
sample handling systems is chemical or physical  reaction between the gases to be analyzed
and the wall materials of the sample lines or containers. Teflon and glass are the most nearly
inert wall materials readily available for sampling  lines, while 316 stainless valves and fittings
are  sufficiently  inert  and corrosion-resistant  for  normal use.  Polyethylene and  poly-
propylene can also be used, but are subject to  melting at gas temperatures above 120° C
(250° F).

The two major types of sample handling systems are the batch and continuous types. Batch
systems in use include the use of cylindrical gas collection  flasks and evacuated glass bottles
that   normally must  be  heated to  150° C  (300° F)  or higher  to  prevent moisture
condensation.  Sample injection into the chromatograph normally is made by  glass syringes
of varying  sizes.  It is sometimes necessary to  concentrate samples by  freezeout, solid
adsorption, or liquid adsorption to have sufficient material to perform analyses.

Continuous  sampling systems use withdrawal of  the gas sample  through a heated line from
the source at a relatively high flow rate. Gas samples can be injected into the chromatograph
at frequent  intervals through a  sample loop  and  port assembly.  Dilution of a  given sample
prior to injection into the chromatograph is sometimes necessary when using detectors, such
                                        17-18

-------
as the flame photometric unit. Also, to minimize the retention time in the sampling lines,
withdrawal of a considerably larger volume of gas than that passed through the sample is, at
times, recommended.

          17.2.5.2   Column Technology

Selecting  the  proper  column and packing is  important  to  successful  analysis by  gas
chromatography. Pertinent variables in column technology are the column length and
diameter,  the solid support, the liquid used, and the tubing material. Selection of the proper
column is necessary to  facilitate the separation  of the gaseous components of interest;
different gases have different relative affinities for different packing materials.

Gas chromatographic column materials sufficiently inert and temperature resistant for sulfur
gas analyses include 316 stainless, glass, and  Teflon (16). The sample injection system, the
separation column, and  the detector must be heated to facilitate many of the sulfur gas
separations. Pressure and temperature limits, inertness, and durability must be considered in
selecting tubular column materials. The degree of separation between components that can
be achieved by a column increases with increasing length and decreasing diameter. Column
diameters normally vary from 0.3 cm (0.125 in) to 0.6 cm (0.250 in) with lengths ranging
from  3  to 30 m (10 to 100 feet). Carrier gas flow rates will vary from 30 to 150 cm3/min
(15).  Teflon column temperatures should not exceed 100 to  150° C (212 to 300° F).

Solid phase support materials  must be of sufficient inertness, porosity, uniformity, strength,
and ease  of packing for generalized use. The separation efficiency  of the solid support is
directly proportional  to its porosity and surface  area,  but inversely proportional to its
inertness.  Noninert columns result in tailing of peaks. Normal column solid supports include
Chromosorb G, P, T, and W.

The liquid phase of the  column separates the various components by either vapor pressure
or polarity; in  either case molecules of greater molecular  weight  tend to remain in the
column for longer periods. Liquid phase materials in common use for sulfur gas separation
include the Carbowax  20X  and  1540, polypropylene,   glycol,  binonylphthalate  and
polyphenyl ether and Poropak Q (15).

          17.2.5.3   Chromatographic Detectors

The major chromatographic detection systems  in use  for sulfur gas analyses are thermal
conductivity,  flame ionization, flame photometry, and microcoulometry.  The  detectors
used  must  have  accuracy, stability, sensitivity,  selectivity, durability, rapid  response,
minimum maintenance,  and  freedom from interfering substances. Thermal conductivity
detectors   are   not  specific   or sensitive  enough  for  most  mill  applications,  and
                                         17-19

-------
microcoulometric detectors are unsuitable because they can be easily overloaded at high
concentrations and require frequent maintenance.

A summary of detector characteristics is presented in Table 17-3.

The  two  major detectors  in  use  in pulp mills include the flame  ionization and  flame
photometric units. Both systems are free from water interference, have excellent stability
characteristics,  and both  require hydrogen flames and a nitrogen carrier gas. The  flame
ionization detector is suitable for organic sulfur and nonsulfur compounds, which can be
ionized in flames, but is not sensitive to  the inorganic H2S or S0a. The flame photometric
detector is suitable for H2S and SO2, as well as organic sulfur compounds, over the range
from  5 ppb to 5 ppm by volume  (17).  Samples of higher  concentrations require  either
dilution or the use of very small sample loops of 0.5 cm3 or lower capacity.

     17.2.6   Calibration Procedures

Periodic  calibration of gaseous monitoring instruments is necessary to their  continued
accuracy.  The typical calibration procedure checks the instrument by using it to measure a
gas stream of known concentration. The difference between  the known and the measured
values is the error of the instrument. Methods in use for preparing known gas concentrations
include rotating  syringes,  motor-driven  syringes, flexible fabric  bags,  known cylinder
mixtures,  and  permeation  tubes.  Of these,  all except the  motor-driven syringes are in
common use.

          17.2.6.1  Rotating Syringe

Rotating syringes can be used for calibration of gaseous monitoring instruments over both
ambient and source ranges  of concentration. The  technique employs dilution of pure gas
from  a small syringe  with air in a large syringe. The large syringe is placed in an upright
position and caused to rotate  by the action of  an airstream directed against its vanes. The
flow  rate of gas  mixture from  the  syringe is controlled by a calibrated limiting flow
capillary,  which is usually  a  broken thermometer. The capillary is inserted into the  air
stream to form the required  gas concentration,  which is  then  fed  to the instrument, as
shown in Figure 17-8 (18).

The small syringes can range in size from 0.5 to 10.0 cm3 while the large syringes can be 50
to 100 cm3. Adding pure gas to the large syringe at extremely high inlet  concentrations is
sometimes necessary. The thermometer capillaries are individually  calibrated  by a soap
bubble flow meter; the  flow rate remains constant because  of the constant pressure exerted
by the rotating  plunger.  Flow rates for the  capillaries normally range  from 0.5  to
5.0 cm3/min. Air flow rates for the dilution system can vary from 0.5 to 50 1/min (0.018 to
1.8 cfm).
                                        17-20

-------
71
IS3
                                                       TABLE 17-3
                       OPERATING CHARACTERISTICS  OF GAS CHROMATOGRAPHIC DETECTORS  (15)
          Detector
     Thermal Conductivity
     Flame lonization
     Flame Photometric
     Bromine Coulometric
 Sensitivity
ppm, by vol

  10
   0.5
   0.005
   0.5
Stability
  Good
Excellent
Excellent
  Poor


H2S
Yes
No
Yes
Yes
Gases

S02
Yes
No
Yes
Yes
Analyzed
Organ.
Sulfur
Yes
Yes
Yes
Yes

Organ.
Comp.
Yes
Yes
No
Yes*
   Water
Interference
    Yes
    No
    No
    No
     'Tor oxidizable organic compounds only.

-------
The rotating syringe method provides a versatile and  inexpensive means of providing gas
concentrations  over a wide range  with a minimum of equipment. But the method is not
without  disadvantages. It  does  require prior calibration of the capillary flow rate. The
syringes are subject to sticking in humid atmospheres.  Careful loading is necessary to avoid
errors. And it is necessary to rotate the syringe plunger by directing an air stream against its
attached vertical rotor; otherwise there is no assurance that a constant delivery pressure is
maintained, which  is necessary  to assure a constant  gas flow rate through the  syringe
capillary.

Permeation tubes provide a versatile and accurate means of calibrating gaseous monitoring
instruments. The technique operates  on the principle  of a constant rate of diffusion of a
pure gas through a porous membrane of fixed cross-sectional area  and  thickness at  a given
temperature  (19).  The  permeation tube is placed in  a  chamber  immersed in a constant
temperature bath,  and dilution  air is caused to pass through the system. The gas stream
containing the dilution air plus the pollutant gas at a specified concentration is then caused
to flow from the dilution chamber to the instrument for calibration purposes, as shown in
Figure 17-9.

Permeation tubes  are made of  cylindrical  lengths of Teflon plastic filled with liquefied
pollutant gas to be measured. They can be fabricated or purchased commercially (20). The
constant temperature bath is normally kept at 25 to  27° C (77 to 80° F). The physical
                           Rotating
                            Syringe
                                                                         To

"A U
Mixing 	 t
Chamber 	 '
Instrument
\ -ru


LxnausT
                    Purification
                      Setion
                                   FIGURE  17-8
        ROTATING  SYRINGE  INSTRUMENT  CALIBRATION  PROCEDURE
                                        17-22

-------

Air \
Supply r

_ r—
/ i
L
Constan
Temp. B
n — n
U/ U

t
3th
1
I
1
Permeation Tube
, 	 ^ To Instrument
Mixing f/ — • — w
— L T — - wTo Exhaust
 Dilution
                                  FIGURE  17-9
        PERMEATION TUBE INSTRUMENT CALIBRATION  PROCEDURE


variables for the permeation tube are its diameter, wall thickness, and length. The primary
operating variables are the temperature and the air flow rate. Permeation rates for tubes are
normally calibrated gravimetrically in terms of weight loss per unit time.

Permeation tubes require the use of a constant  temperature bath and are more cumbersome
for field use  and more expensive than rotating syringe systems. Newly developed systems
now available commercially eliminate the use of liquid baths, resulting in a lighter weight,
more compact and less complex unit.

         17.2.6.3  Gas Cylinders

Stainless gas cylinders used for calibration of gaseous monitoring instruments are cylinders
of given volume. Known amounts of gas are inserted in them and pressurized with either air
or nitrogen to produce known concentrations (21). The gas mixture can then be added at a
given flow rate to a stream of flowing air to produce a given concentration. The technique is
suitable for field use and is simple and inexpensive; however, severe losses can occur during
storage for certain gases by adsorption on or reaction with wall materials.

         17.2.6.4  Calculation Procedures

     1.   Rotating Syringe
                                 CLQL_VSQ
                                   Qi

         where:   Cj   =  concentration fed to instrument, ppm by volume
                  CL   =  concentration in large syringe, ppm by volume
                  Qj   =  flow rate of dilution air, cm3/min
                                       17-23

-------
                   QL  =  flow rate from large syringe, cm3 /min
                   VL  =  volume of large syringe, cm3
                   Vs  =  volume of small syringe, cm3

          Note.—Because Qj, QL ,  VL  and  Vg appear only in ratios, any consistent units of
          volume and volume/time may be  used.

     2.    Permeation Tube (20)


                                             (R) (T)
          where:   Cj   =  concentration fed to instrument, ppm by volume
                   MW =  molecular weight of gas, g/mole
                   P    =  pressure of calibration system, mm Hg
                   Qj   =  dilution air flow rate, 1/min
                   R    =  permeation rate, ng/min
                   T    =  water bath temperature, °K

17.3   Particulate Monitoring

Particulate sampling and analysis are necessary for recovery furnaces, smelt tanks, and lime
kilns in kraft pulp mills, for recovery furnaces in sulfite pulp mills, and for all coal-, wood-,
and oil-fired power boilers.  Determination of  particulate mass concentration is necessary to
meet current air pollution regulations, with increasing emphasis being placed on particle size
distribution. Batch techniques are specified for determining particulate mass concentrations,
emission rates, and size distributions. Continuous  monitoring of particulate emissions is
gaining increasing emphasis as more accurate and reliable methods are developed (22).

     17.3.1  Preliminary Considerations

Particulate matter is normally defined as any material emitted into the atmosphere in either
a solid or liquid state, including dusts, fumes, smoke, flyash, soot, tars, droplets, and mists.
Changes in physical state with  temperature  can cause  confusion in  defining particulate
matter for materials such as organic vapors and acid mists. Changes in  chemical form, such
as oxidation of S02  to  H2S04 in  sampling train impingers after collection, can cause
especially serious difficulty in interpreting emission standards.

Particulate matter  must be defined as such for either stack or standard conditions.  The
present definition accepted by the U.S. Environmental Protection Agency is that particulate
matter is material collected on a filter of porosity 0.45 ;um (1.8 X  10~s in) which has been
                                         17-24

-------
heated  to 121° C  (250° F). Other definitions for particulate matter specified  by local
agencies  include,  among  others,  material  collected in  liquid  impingers  at  standard
conditions.

Particles in pulp and paper mill sources can vary from less than 0.01 to greater than 100 /zm
(4X10  7 to 4X10  3 in) in diameter, depending on the source and type of collector used.
Sampling at isokinetic conditions is normally  necessary because of the inertial properties of
particles greater than 5 jum (2 X  10"4 in) in diameter. The normal procedure employed for
maintaining isokinetic conditions within plus or  minus 10  percent  of the  actual stack
velocity  is to locate a pitot tube parallel to the sample probe to measure the gas velocity
continuously and to make periodic corrections as needed.

     17.3.2   Batch Particulate Sampling

Batch  particulate  sampling is used to determine total  particulate  concentrations and
emission  rates  from flue gas streams  and  is necessary for calibration  of continuous
particulate monitoring devices.  Batch sampling provides an  average value for  particulate
concentration  in  a duct  over  a  given time period, but  does  not  provide real-time
instantaneous values. Collection methods for batch particulate sampling include filtration
and liquid impingement.

A low volume system (0.8 to 2.5 m3/h or  0.5 to 1.5 cfm), specified by the American
Society of Mechanical Engineers (ASME), relies on an internally located alundum  thimble
for particulate  collection, as shown in Figure  17-10 (23). The amount of material collected
is then  analyzed gravimetrically. The  standard  specifies collection of  particulate materials
larger than 1 /um (4 X 10" s in) in diameter. The method is relatively simple and inexpensive,
but has several shortcomings. The thimbles tend to pass small particles, particularly during
the initial collection  period, in  direct proportion  to increasing thimble porosity. The
thimbles are  subject to leakage  by  improper  sealing around their gaskets, to  plugging and
washthrough in wet gas streams, and to dust losses during handling following collection.

The Los Angeles County Air Pollution Control District employs  a multiple-stage collection
train for particulate sampling in an arrangement similar to that employed  by the  National
Council for Air and Stream Improvement (24). One sampling train employs  an internally
located  heated  cyclone and alundum thimble  as the primary collection stage, followed by a
series of liquid impingers, and a five  grade thimble for removal of particles of diameter
greater than 0.3 /im (1.2 X 10~6 in).

The vacuum  pump is placed downstream of  the meter to preclude air dilution of the gas
stream. A sampling rate of 0.8 to 1.5 m3/h (0.5 to 0.9 cfm) is employed. Total particulate
matter is then  determined by gravimetric analysis of the material collected. The  method
requires several analyses, and the possibility of sulfate formation from oxidation of SO2 in
the impinger liquid following collection introduces further uncertainty.

                                        17-25

-------
           t
      ALUNDUM
      THIMBLE
                VACUUM
                 PUMP
                             CONDENSER
  GAS
METER
          -^-
           T
                                FIGURE 17-10
                ASME BATCH PARTICULATE SAMPLING TRAIN
The participate  sampling train,  specified by the U.S. Environmental Protection Agency,
employs an externally located two-stage collection system as shown in Figure 17-11 (25).
The primary collection stage consists of a glass cyclone and a 5 cm (2 in) diameter glass fiber
filter of 0.30 to 0.45 /um (1.26 to 1.8 X 10~~6 in) porosity that is heated to 121° C (250° F)
to  avoid  condensation.  The  secondary  collection stage consists  of  a series  of
Greenburg-Smith liquid impingers containing water. The  impingers are used to condense
water  vapor present  so as to prevent flooding the pump and meter and for absorption of
potentially corrosive gases. In addition, the water vapor content of the stack gases is
determined from the condensate collected in the impinger. The vacuum pump follows the
                       xffSTACK   r
                        UWALL
                                FIGURE  17-11
                EPA  BATCH  PARTICULATE  SAMPLING TRAIN
                                     17-26

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impingers and is followed by the gas meter to measure the volume ^sampled. An S-type pitot
tube, located in parallel to the sample  probe, facilitates the maintenance of isokinetic
sampling  conditions.  The  system  provides for efficient particulate  collection,  but is
cumbersome to handle. The small filters are subject to rapid plugging at high loadings, and
large pressure drops tend to  develop after prolonged sampling periods.

     17.3.3  Particle Size Distribution

A major problem in particulate sampling is the determination of particle size distribution in
flue gas streams. Particles can range in diameter from less than 0.01 to greater than 100 /um
(4 X 10~7  to 4 X  10~3 in), with densities ranging approximately from 0.8 to 2.5 g/cm3.
The problem of determining particle  size in gas streams is made  especially difficult by  the
high humidity of the gas, the presence of water droplets, and the tendency of particles to
agglomerate  or coalesce. The major methods employed for particle size determinations in
emissions from pulp and paper mill sources are multistage cascade impaction and membrane
filtration.

Bosch, Pilat, and  Hrutfiord (26) describe  the use of a multistage cascade impactor  for
particle size determination from kraft  recovery furnaces over  the range  0.5  to  20 jum
(2 X 10~5  to 8 X 10~4 in) in diameter. The system is an internally located six-stage cascade
impactor  in  which particles of progressively smaller diameter are collected on  successive
silicone-coated  plates by  passage through multihole  plates with  progressively smaller
diameter holes.  Sampling times  vary from  30 seconds to 30 minutes, depending on  the
particulate loading of the source being measured.

The system must be calibrated for a given dust at a certain flow rate prior to collection to
determine  the  approximate mean particle diameters for individual stages. Particle size
determinations are  made  by  gravimetric weighings  of individual  plates;  the system  of
classification of particles is as percentage by weight within given sizes. Particles smaller than
0.5 jum (2 X  10"5 in) in diameter are collected on a filter following the impaction plates, as
shown-in Figure  17-12.

The National Council  for  Air  and  Stream Improvement has  developed  a  method  for
collection of particles on a membrane filter for subsequent visual counting and sizing by
means of an  optical  microscope (27). The system employs a parallel flow heated collection
train with the membrane filter on one branch and a valving system on the other, as shown in
Figure 17-13. The system  must be  hydrostatically  balanced so that there is  the same
resistance to flow in both  branches. It  is then placed in the stack and heated  to 93° C
(20P° F).  This temperature is high enough  to prevent condensation, but not so high as to
damage the  filter. The gas stream is first  directed through the bypass system  and then
through the  filter  for 20 to 30  seconds  to  collect the particles. The technique is suitable
only for sizing particles on a count basis, which is a time consuming procedure, and does not
                                        17-27

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           I    CASCADE IMPACTOR  ASSEMBLY
           2    WATER BATH
           3    IMPINGERS
           4    THERMOMETER
           5    DRY GAS METER
           6    VACUUM GAUGE
           7    VACUUM PUMP
           8    GAS FLOW RATE REGULATOR

                  FIGURE  17-12
MULTISTAGE CASCADE  IMPACTOR FOR PARTICLE  SIZE
        DISTRIBUTION DETERMINATION (26)
                  FIGURE 17-13
  MEMBRANE FILTER SYSTEM  FOR PARTICLE SIZE
        DISTRIBUTION DETERMINATION  (27)
                      17-28

-------
allow determination of  particle  size  distribution  on a  weight basis. But, the  sample
collection can be done in a short time interval, does not require the extensive precalibration
required for the cascade impactor devices, and does not require prior knowledge of the dust
density characteristics.

     17.3.4   Continuous Monitoring

Continuous particulate monitoring is gaining increasing emphasis for  determination of stack
concentrations and emission rates. Methods now in use for particulate monitoring in the
pulp and paper industry include wet chemical techniques, such as conductivity and specific
ion determinations, and  optical bolometry  and transmissometry. Techniques which may
become applicable in the future, include beta ray attenuation, piezoelectric crystallography,
electronic sensing, optical nephelometry, holography, electric ion capture, and optical lasers,
lidar, and radar (28).

          17.3.4.1  Chemical Methods

Leonard (29) describes a wet chemical system for determining particulate emissions from a
kraft recovery furnace on a continuous basis. The system operates on the principle that the
increase in electrical conductivity of a liquid stream caused by the presence of sulfate ion is
proportional  to the total  particulate loading,  which is  primarily Na2SO4. Flue  gas is
withdrawn from the duct at a predetermined average isokinetic  velocity into a  probe and
contacted concurrently  with  deionized water,  as  shown in Figure 17-14. The  soluble
particulate matter is passed  through a detection cell containing a conductivity probe and the
                                            CONCENTRATION
                                              READOUT
                                  FIGURE  17-14
           CONTINUOUS MONITORING OF PARTICULATE EMISSIONS
                  WITH  A CONDUCTIVITY  CELL DETECTOR  (29)
                                        17-29

-------
gas is  drawn  off  by a  vacuum pump. The  method is simple and  inexpensive, but is
nonspecific because any conductivity-producing substance can cause interference, e.g., SO2
or CO2 from the flue gas.

Tretter has developed a modification to the conductivity method for measuring particulate
concentrations which alleviates conductivity interferences  caused by the presence of SO2
and CO2 in the flue gas (30). The method uses a sodium ion-specific electrode to measure
the sodium content of a water stream as an indicator of total particulate concentration. The
system withdraws the gas sample through a probe and allows the absorption into water in a
concurrent flow condenser by condensation of compressed steam, as shown in Figure 17-15.
The liquid stream is then directed to a detection cell containing the sodium electrode where
a liquid  pH of 8.5 to 9.5  is maintained for maximum sensitivity. The system requires
periodic maintenance, and, for calibration, a correlation between sodium  ion level and total
particulate concentration must be developed by using parallel batch tests.

          17.3.4.2  Optical  Determinations

Optical devices in use for particulate monitoring include internally located bolometers and
transmissometers, where  the degree of light attenuation is  a  function  of the particulate
concentration  in the duct. The  system  directs  a light beam across the duct to  a detector
where the resulting electrical signal is amplified, transmitted,  and printed  out on a recorder.
                                              TO DRAIN
                                  FIGURE  17-15
     CONTINUOUS MONITORING OF PARTICULATE  EMISSIONS  WITH A
                    SODIUM  ION-SPECIFIC ELECTRODE  (30)
                                        17-30

-------
The sensitivity and accuracy of optical stack monitoring devices is affected by path length,
intensity, and wavelength of the light beam, moisture content as a function of temperature
of the stack  gas, particle size distribution, and particle mass concentration of the flue gas
stream. (31).

Gansler describes the use of a bolometer for measuring particulate emissions from a kraft
recovery furnace flue gas (32). The system uses a tungsten lamp with an optical band from
approximately 500 to 2,500 nm. The device is suitable for measuring particle concentrations
at levels below 29/m3 (0.9 gr/cu ft), and provides warnings of precipitator malfunctions.
The device requires extensive prior  calibration by parallel batch tests and also frequent lens
cleaning.

Beutner describes the use of optical transmissometers for continuous particulate monitoring
at a  number  of installations (33). The system employs a  fixed length light beam with a
wavelength of 300  to 800 nm, about the visible region of the spectrum. The gas stream
passes between the beam and the detector.  Fans are used to blow sufficient air across the
surface to keep the lens surfaces clean. The device can be  either permanently mounted or
portable. The system is  effective as an indication of particulate matter primarily of a size
range of 0.1-1.0 nm (4 to 40 X  10~6 in). The  system must be calibrated for the specific
source by an extensive series of batch tests.

Optical particulate  measurement devices are  suitable for application to kraft recovery
furnaces, coal- and wood-fired power boilers, and kraft lime kilns, and possibly ammonium
and magnesium-base sulfite recovery furnaces.

They are particularly useful for providing warnings of possible particulate emission control
equipment malfunctions  and as indicators of combustion unit efficiencies for power boilers.
They are also useful  for emission monitoring  for possible compliance with stack opacity
standards, because optical transmittance is the  property they measure. Their use is limited
for mass emission  monitoring because  long-term correlations with  batch tests must be
determined under stable operating conditions. The lenses of the detectors tend to become
obscure so that they can require  frequent cleaning, and changes in particle size distribution
can alter readings. The devices are probably most suitable for stacks with low particulate
concentrations following high efficiency control devices where the particle size distribution
is relatively uniform.

17.4   Odor Measurements

Measurements of odors  is a  major problem  in the pulp  and paper industry. The major
chemical components that may cause a community odor  nuisance  are reduced sulfur
compounds, such as H2S, CH3SH, CH3SCH3, and CH3SSCH3. These malodorous sulfur
compounds and, possibly,  other  organic  compounds  are  emitted  from the  digesters,
                                        17-31

-------
evaporators, recovery furnace, smelt tank, lime kiln, and wastewater streams in kraft pulp
mills. Under certain circumstances, S02  and organic sulfur compounds from sulfite pulp
mills can also cause a community odor nuisance.

Odor measurement is an  extremely complex task because of variations in the types and
amounts of odorous gases present, variations in response between individuals, and variations
within an individual with  time because of his physical condition, smoking history, time of
exposure,  and prior history of exposures.  Meteorological variables, such as temperature,
humidity,  wind velocity,  and turbulence,  can affect odor responses of individuals (34).
Odorous gas molecules  that are adsorbed on the surfaces of particles can travel for  longer
distances in a concentrated form and bring sharper responses than the odorous gas molecules
alone (35).

     17.4.1  Threshold Levels

The odor threshold levels  for most malodorous sulfur compounds emitted from kraft pulp
mill sources are between 1 and 10  ppb by  volume, as shown in Table  17-4 (36):
                                   TABLE 17-4
                        ODOR THRESHOLD  LEVELS FOR
                   MALODOROUS SULFUR  COMPOUNDS (36)

                   Sulfur Compound                   Threshold
                                                     ppm, by vol

                         H2S                               4
                        CH3SH                              2
                      CH3SCH3                             4
                      CH3SSCH3                             6
                         S02                            3,000
The odorous gas levels for flue gas streams can be anywhere from 1,000 to 100,000 times as
high as those observed at distances of up to 10 km (6 miles) from the kraft pulp mill.

The evaluation  of odors can be made  in terms of their character, intensity, pervasiveness,
and acceptability. The intensity of an odor tends to increase with the logarithm of its
concentration according to the Weber-Fechner  law (37). Odor intensity levels can be rated
on an arbitrary scale in values ranging from one to five, as shown in Table 17-5 (35).
                                      17-32

-------
                                    TABLE 17-5
                       ODOR  INTENSITY LEVEL EVALUA-
                                    TION SCALE

                      Arbitrary                       Intensity
                      Number                          Level

                         0                         No Odor
                         1                         Detectable
                         2                         Faint
                         3                         Noticeable
                         4                         Strong
                         5                         Overcoming
These levels can be used as reference scales for people on a panel in an attempt to provide
some quantitative scale of odors.

     17.4.2   Design Considerations

Major considerations in performing an odor threshold evaluation include the selection of
human subjects, the sample collection methods, and the dilution techniques employed. The
two general approaches to odor level measurement are:

     1.   Direct subjective organoleptic determination of odor threshold levels, and

     2.   Nonselective chemical determination of individual odorant concentrations by gas
         chromatography or other means.

Odor level measurements require careful collection and  careful evaluation of samples.

An odor panel is normally selected as a group of people who have particular sensitivity,
accuracy, speed, and  reproducibility for evaluating the odor threshold  and intensity levels
for the gases being measured. The odor panel goes through a training program for a specific
period, and then  is used to test stack gas mixtures. Panelists may undergo  testing for 15
minutes  at a time with rest periods of at least 30 minutes to avoid olfactory fatigue. It is
normally best to expose panelists to low concentrations before high concentrations to avoid
deadening their response. Prince and Ince find that responses can vary as  much as 40 percent
for individual panelists on any given day and 20 percent for a panel of observers (38).

Gas  sample collection may be by direct  piping of the flue gas to a panel to  minimize
potential losses, but this may not always be feasible. Samples of stack gas or ambient air
                                        17-33

-------
may also be collected in plastic bags, heated evacuated bottles, or in syringes. Odorous gas
losses by  solution in condensed water  vapor, or physical absorption or chemical reaction
with container walls, are to be eliminated normally by the use of inert wall materials such as
Teflon.

     17.4.3  Dilution Techniques

Two  publications by  the National  Council for Air  and Stream Improvement  present
extensive  discussions of systems for dilution of odorous gases to facilitate evaluation of
odorant intensity levels  (39) (40). The major techniques for evaluation of odor threshold
levels are:

     1.    Static progressive dilution using glass syringes,

     2.    Dynamic dilution of odorous gases using odor-free air,

     3.    Dilution of odorous air with odor-free air by respiration, and

     4.    Vaporization of odorous compounds in a continuous flow system.

The American Society for Testing and Materials specifies progressive static  dilutions in glass
syringes. The same is first collected in a 100 cm3  glass syringe (41). The flue gas sample then
is diluted in the  laboratory by adding aliquots of known volume to other syringes of this
same size. The aliquots then are  given to  an odor panel in  order of decreasing dilution
(increasing odorant concentration) to evaluate odor threshold levels and odorant intensities.
The syringes must be cleared after use and dried to avoid contamination of future samples.

Several  systems are available for dynamic dilution of odorous gases by means of odor-free
air  for  subsequent observation  by  a panel. Normally, the odorous gas  sample must be
collected  in some type  of  container  before being  added to  the  dilution system  in a
laboratory. Field  exposure of subjects prior to odor level evaluations can lead to olfactory
fatigue  and resultant insensitivity to odorous gas concentrations. Cederlof (42)  describes a
system  in  which  odorous gas samples are collected from kraft  mill flue  gases in  flexible
plastic bags and  returned to the laboratory where they are diluted with purified  air for
exposure to a panel  inside a hood, as shown in Figure  17-16. The inlet odorant gas flow is
controlled at a constant rate by a limiting flow capillary, while the air flow rate is measured
by a rotometer.

Direct observations of odor levels  in the atmosphere or from flue gas streams by means of a
progressive dilution device known  as  a  "Scentometer" have been made.  The device is
powered by the observer's respiration. The system draws in variable amounts of odorous gas
or air through an adjustable orifice followed by dilution with air which has been  purified by
                                         17-34

-------
                  PLASTIC BAG
                 WITH ODOROUS
                     GAS
r
                                                   FAN
                        ~     CALIBRATED
                       / \ CAPILLARY TUBES
                                                           ODOR
                                                           PANEL
                                 FIGURE 17-16
SWEDISH  DYNAMIC  DILUTION SYSTEM FOR  ODOR LEVEL EVALUATION  (42)


activated carbon, as shown in Figure 17-17 (43). The relative amounts of odorous gas and
purified air are governed by a series of four orifices which provide for dilution factors of 2,
7, 31, and 170 times, respectively. The procedure starts from the greatest degree of dilution
and works progressively toward lesser dilution until the odor threshold is observed.

17.5  Mobile Laboratories

Mobile laboratories provide the capability of making comprehensive evaluations of ambient
and  source  concentrations for gaseous and particulate  materials  in the  field.  Mobile
laboratories  used  for air pollution studies should be equipped with instrumentation for
continuous measurement of sulfur compounds, a gas chromatograph for analyses of specific
compounds, analytical equipment for additional constituents, such as oxides of nitrogen,
and particulate sampling equipment.

Walther and Amberg describe a mobile laboratory mounted in a van that has the capability
of monitoring malodorous sulfur compounds. It uses two gas chromatographs in parallel
equipped with thermal conductivity and flame ionization detectors, respectively (44). The
sample handling system consists of a heated ceramic probe and a Teflon electrically traced
                                      17-35

-------
                                              -NOSEPIECES
                               JLJ



/
PURIFIED /
AIR FOR
DILUTION
ji
!ii
j!'
1
IM
'i1
i ! i
i!i
i i :
' • 1
ili

ft
f
I


\
\
\
I

ODOR

VV

f i

! 1 1
!ft
!j!
V '
1 1
Jji
i!1
: II
iii
ill
i'i
^ CHARCOAL
BED

\
PURIFIED
AIR FOR
DILUTION
^ ^^ HRAni IATFH
SERIES OF
\ ORIFICES
                              ODOROUS AIR
                                FIGURE  17-17
                 SCENTOMETER DILUTION SYSTEM FOR ODOR
                       THRESHOLD EVALUATION  (43)
sampling line for carrying the flue gas from the stack to the instruments. The van proved too
small and had to be placed in a permanent location, thus negating its mobility.

Mulik, Stevens, and  Baumgardner report on a mobile laboratory mounted in a trailer which
employed  flame photometric detectors for measuring reduced sulfur compounds (45).
Samples were taken at a rate of 50 1/min (1.8 cfm) from a source gas through a 0.63 cm'
diameter and 76 m long (0.25 inch by 250 ft) electrically traced Teflon sampling line main-
tained at 182° C (360° F) to prevent moisture condensation. The gas chromatograph em-
ployed a  10-port sample valve with a 10 cm3  (0.6 cu in) sample  loop actuated on a 10-
minute sequence for sample injection. A dynamic dilution system was used for sample dilu-
tion factors covering a range  from  10 to 1,000,000 to 1, depending on the requirements
for optimum sensitivity with the flame photometric detector.

    17.5.1  NCASI System

The National Council for Air and Stream Improvement employs a trailer for mobile emission
sampling of kraft pulp mills (46). The sample system consists of a heated  ceramic sample
probe, a heated, electrically traced Teflon sampling line, and a heated conditioning box for
                                     17-36

-------
dilution of the sample, as shown in Figure 17-18. The gas is withdrawn at a greater rate than
that required  for the instruments in order to minimize retention time in the sample lines.
The instrumentation  consists of two coulometric titrators with  a  furnace  for sulfur  gas
analysis and a gas chromatograph equipped with flame ionization and flame photometric
detectors. Total cost  of the entire system  when constructed was approximately $30,000,
and it required two men for its operation.

The sampling handling system  employs  withdrawal of  the sample  gas  at a rate of
approximately 1 1/min (0.035 cfm) through a probe heated to 150° C (300° F) to evaporate
water droplets to alleviate  droplet entrainment in the sampling lines. The electrically traced
0.63 cm (0.25 inch) diameter Teflon sampling line  33 or 66 m (100  or 200 ft) length is
heated to 110 to 121° C (230 to 250° F) to prevent moisture condensation. The gas stream
then passes into a heated sample conditioning box, where portions of the gas stream are bled
off to either the continuous sulfur monitors or the gas chromatograph, while  the remainder
passes to a large glass carboy acting as a condensate trap to avoid damage to the pump. The
sample gas stream leading to continuous sulfur analyzers can be  diluted by factors from zero
to 25.
            HEATED SAMPLE PROBE

             *'•*&*
             ^^\
                      TO 2ND TITRATOfi **°™!*!L   ^
                                                (RECORDER
                                               I AMPLIFIER
                                 AMPLIFIER   RECORDER
             HEATED
             SAMPLE
              LINE
                        SO SCRUBBER
                  ^
                      DILUTION AIR
                      OR Ml
                          TROGEMl
                                           DETECTION CELL
                                          7-PORTl'
                                          SAMPLlfT
                                                SAMPLE
                                                LOOP
                                        CVAJJYEJ_I_
                                        TT
                                        ft I   "
                                   IIO°C(2
                                     £o*F7
      GAS FLOW •
                           Ea
                           #
    ELECTRICAL	
CHROMATOGRAPHIC DETECTORS
   I. FLAME IONIZATION
  2. FLAME PHOTOMETRIC
   GAS FLOW RATES
 Qs  SAMPLE FLOW     K)-100 ML/MIN.
 0A  DILUTION AIR     I50-24O  "
 QB  INSTRUMENT FLOW  250     "
HEATED SAMPLE CONDITIONING
       BOX
                          VENT -*-KI)| W\-* VENT
                                 t_~J DETECTORS

                                  —I—   GAS
                                      CHROMATOGRAPH
                                 COLUMN
FT
                                 CONDENSER
                                   TRAP
       TOTAL FLOW
                     1000
                                   FIGURE  17-18
          GASEOUS  SAMPLING SYSTEM  FOR SULFUR GAS  ANALYSIS
                      IN  NCASI  MOBILE LABORATORY (46)
                                        17-37

-------
The continuous sulfur analysis system either splits the gas flow into two parallel streams for
simultaneous measurement of both total sulfur and TRS, or passes a single stream through
an S02  selective scrubber containing either KHC8H404 or citric  acid (C6H8O7). The gas
then passes  through  parallel quartz combustion  furnaces heated to 760° C  (1,400° F) for
oxidation of the reduced sulfur compounds to SC>2, which then is measured by means of
parallel  bromine coulometric titrators. The difference between the signals for  total sulfur
(S02 +  H2S + organic sulfur) and TRS (H2S + organic sulfur) is taken as being proportional
to the  S02  concentration of the flue  gas stream. The separation and identification of
specific  sulfur compounds is made by periodic injection of samples of given volume through
the sample loop. The sample size is governed by the volume of the sample loop.  The sample
gas then passes through a separation volume for analysis by either flame ionization or flame
photometric detection.

     17.5.2   Rayonier System

Waddington  describes  the  mobile  laboratory  constructed  by  ITT-Rayonier,  Inc., for
monitoring paper mill emissions and  ambient air  from a 10.7 m (35 foot) truck  trailer (47).
The system has both  particulate and gaseous sampling trains and sample handling systems, as
shown in Figure 17-19. The particulate systems consist of an EPA  train for total particulate
analyses and an Andersen sampler for particle size determination.  The gaseous sampling
system  consists  of parallel heated sampling lines and a  sample  conditioning system to
provide  for  dilutions with air  by factors of 10 to 10,000. The  gaseous instrumentation
consists  of  continuous monitors for reduced  sulfur  compounds,  SO2,  CO, oxides of
nitrogen, and hydrocarbons. The gas  chromatograph is equipped with  flame ionization and
flame photometric detectors.

The sample  handling system for source gas analyses employs withdrawal of  the sample gas
through heated filters located outside the stack through an electrically traced heated Teflon
sampling line of 0.95 cm (3/8 in) diameter 33 or 66 m (100  or 200 ft)  in length. The gas
stream passes through at a rate of 100 to 160 1/min (3.5 to 5.7 cfm) and small portions then
pass out into  a series  of  as many as three consecutive  dilution stages in a heated sample
conditioning box  for  dilutions  from 15  to 1,500  to one. Portions  of the gas streams,
including SO2, NOX, CO, and  TRS, are routed through three separate channels to several
detectors for measuring individual gaseous products.

Ambient monitors are used for S02 and N02.  The gas chromatograph employs parallel
flame photometric and flame ionization detectors with a 10-port sampling valve for sample
injection.

For particulate sampling, two parallel EPA-type sampling trains can be  used on the  inlets
and outlets of particulate emission  control devices. Either can  be fitted  with  Andersen
cascade  impactors  that can   be   internally  or externally  located  for particle  size
determinations. Ambient  particulate measurements can  be made  with several high volume

                                        17-38

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                                                                 AMBIENT SAMPLING TOWER
                                                                  (15'-45' EXTENTION)

100' » 200' UMBILICAL CORD (HEATED TEFLON)
t "DUPLEX FILTER
* — p*h\
IEPA COLLECTION
BOX
DEVICES
' 1 L jr*>w
i 1 C^j "\
EPA COLLECTION BOX
a ANDERSEN CASCADE
IMPACTOR

S ( HEATED GUI
loo' a 200'
UMBILICAL
CORDS FOR
PARTICULATE
SAMPLE
' 	 /

CL
SPAC)
VENT j-'-l DUPLEX FILTERS
ID U 1 METAL BELLOWS
	 	 LJ— y-1— 	 ( H E ATED)
M '
DILUTION PUMPA
DILUTION GAS i- | *

fTION GAS *> 	 . r
EPA MOTOR
CONTROL BOXES
INST

r
r
f
MET METEROLOGICAL \ H
SENSORS a AMBIENT )
SAMPLE FILTERED S
INTAKE (GLASPAC) /
(HEATED GLASPAC)
ATM
GAS PUMPS 1 VENT

f
f
r
r
1

H
H
ATM
VENT
t
f
f
n n n t
rJ
RUMENT
— 00—
n
MODULE
—«
] C
* "%


TEST
TER


                                                                          HI - VOLUME
                                                                          CASCADE IWPACTOR
                                                                       AMBIENT TEFLON
                                                                       SAMPLE LINES
                                   CALIBRATION GAS TO TIP OF GAS
                                      SAMPLING UMBILICAL
                                   FIGURE  17-19
         PARTICULATE  AND  GASEOUS SAMPLE  HANDLING SYSTEMS
                FOR ITT-RAYONIER MOBILE  LABORATORY (47)
suspended particulate samplers carried in the van. A weather station is also carried in the van
for meteorological studies of wind speed, wind direction, and air temperatures. Total cost of
the entire van was about $250,000, and its operation requires three men.

17.6   Economics

An important factor in the design and operation of continuous stack monitoring systems is
the cost of the necessary equipment. Auxiliary equipment, such as recorders, sampling lines,
fittings, and  other appurtenances, must be included  in the total cost. Failure to  properly
account  for  the auxiliaries can cause  variations in  prices  between  different  systems.
Approximate ranges in capital cost  for gaseous monitoring equipment  are listed  in Table
17-6, and  for particulate monitoring  equipment in Table  17-7.  The figures listed are
expressed as  approximate  ranges   only;  exact  figures  must be  determined by direct
quotations from specific manufacturers.
                                        17-39

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                               TABLE 17-6
    APPROXIMATE CAPITAL COSTS FOR CONTINUOUS GASEOUS STACK
                     MONITORING INSTRUMENTATION

        Instrument                       Pollutants                 Capital
          Type                          Measured                    Cost
                                                                 dollars

Electrolytic Conductivity                      SOX                  1,000-2,000
Ultraviolet Spectrophotometry            SOX, NOX, TRS             2,500-10,000
Electrochemical Transducer             SOX, NOX, CO, TRS           3,000-9,000
Coulometric Titration                      SOX, TRS               5,000-8,000
Flame lonization                            Organic                 2,000-7,000
Chemiluminescence                          NOX                  5,000-7,000
Flame Photometry                         SOX, TRS               4,000-8,000
                                TABLE 17-7
   APPROXIMATE CAPITAL COSTS FOR CONTINUOUS PARTICULATE  STACK
                     MONITORING INSTRUMENTATION

        Instrument                       Detection                  Capital
          Type                         Principle                    Cost
                                                                  dollars

Optical Bolometer                         Optical                 2,000-3,000
Optical Transmissometer                    Optical                 5,000-7,000
Electronic Sensing                        Electrical                6,000-7,000
Beta-Ray Attenuation                     Radiation               10,000-20,000
                                     17-40

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 17.7   References

  1.  Cooper, H. B. H., and Rossano, A. T., Source Testing for Air Pollution Control. Wilton,
     Connecticut. Environmental Science Service Corp., 1971.

  2.  Blosser, R. 0., Cooper, H. B. H., and Megy, J. A., Gaseous Emissions—Automatic
     Techniques—Electrolytic Titration. NCASI Atmospheric Pollution Technical Bulletin
     No. 38, National Council for Air  and Stream Improvement, New York, New York,
     December 1968.

  3.  Thoen, G. N., DeHaas, G. G., and Austin, R. R., Continuous Measurement of Sulfur
     Compounds and their Relationship to Operating Kraft Mill Black  Liquor  Furnaces.
     Tappi, 52:1485-1487, August 1969.

  4.  Thoen, G. N., DeHaas, G. G., and Austin, R. R., Instrumentation for Quantitative
     Measurement of Sulfur Compounds in Kraft Gases. Tappi, 51:246-248, June 1969.

  5.  Canfield, J., Measurement of Odors and Sulfur Compounds. (Presented at the 12th
     Conference on Methods in Air Pollution and Industrial Hygiene Studies, Los Angeles,
     California, April 7, 1971.)

 6. Altshuller, A.  P.,  and Sleva,  S.  F.,  Vapor Phase Determination  of Olefins  by
     Coulometric Method. Analytical Chemistry, 34:418-422, March 1962.

 7. Austin, R. R., Sampling and Analysis of Pulp Mill Gases for Sulfur Compounds. Tappi,
    54:977-980, June 1971.

 8.  The Barton Model 286 Sulfur Titrator. ITT Barton Instrument Co., Monterey Park,
    California, 1967.

 9. Cooper,  H. B. H.,  and Rossano, A.  T., Continuous Source Monitoring of Gaseous
    Sulfur Compounds in the Paper  Industry.  (Presented at the 12th Conference  on
    Methods in Air Pollution and Industrial Hygiene Studies, Los Angeles, California, April
    7, 1971.)

10. Miller, A. M., Brown, J., and Abrama, R., Applied Techniques of Analyses for Stack
    Emissions. (Presented at the West Coast Regional Meeting of the National Council for
    Air and Stream Improvement, Portland, Oregon, October 2, 1968.)

11. Mathis, G. V., Application of an Electrochemical Cell to NOX and S02 Monitoring.
    (Presented at  the Miami University Symposium  on Instrumentation for Continuous
    Monitoring of Air and Water Quality, Oxford, Ohio, June 21, 1973.)
                                       17-41

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12. Thoen, G. N., DeHaas, G. G., and  Baumgartel, F. A., Continuous Sulfur Dioxide
    Monitor and its Application  to Sulfite Recovery Emissions.  Tappi, 52:2304-2305,
    December 1969.

13. Saltzman, R. S., Use of Photometric  Analyses for Ultraviolet Analyzers for NOX and
    SOX.  (Presented  at  the Miami  University Symposium  on Instrumentation for
    Continuous Monitoring of Air and Water Quality, Oxford, Ohio, June 21, 1973.)

14. Personal communication with Mr. Philip C. Stultz, Boise-Cascade Corporation, Salem,
    Oregon, June 21, 1973.

15. A Guide to the Use of Gas Chromatography in Emission Analysis. Atmospheric Quality
    Improvement  Technical Bulletin  No. 59.  National  Council  for  Air  and  Stream
    Improvement, New York, New York, February 15, 1973.

16. Adams, D. F., and Koppe, R. K., Evaluation of Gas Chromatographic Columns for
    Analyses  of Subparts per Million Concentrations  of Gaseous  Sulfur Compound.
    Environmental Science and Technology, 1:479-483, June 1967.

17. Mulik, J. D., Stevens, R. K., and Baumgardner, R., An Analytical System Designed to
    Measure Multiple Malodorous Compounds Related to Kraft Mill Activities.  (Presented
    at the TAPPI Water & Air Conference, Boston, Massachusetts, April 4, 1971.)

18. Rossano, A. T., and Cooper, H. B. H., Procedure for Calibrating a Continuous N02
    Analyzer. Journal  of the Air Pollution Control Association, 13:518-523,  November
    1963.

19. O'Keefe, A. E., Ortman, G. C.,  Primary Standards for Trace Gas Analysis.  Analytical
    Chemistry, 38:760-763, June 1966.

20. Duncan, L., and Tucker, T. W., A  Guide to  the Use of Permeation  Tubes as Primary
    Standards for  Instrument  Calibration. Atmospheric  Pollution  Technical  Bulletin
    No. 47, National Council of the Paper Industry for Air and Stream Improvement, New
    York, New York, May 1970.

21. Duckworth, S., Levaggi, D., and Lim, J., Field Dynamic Calibration ofSO^  Recording
    Instruments. Journal of the Air Pollution Control Association, 13:429-434, September
    1963.

22. Cooper, H. B. H.,  The Particulate Problem:  Continuous Particulate Monitoring in the
    Pulp  and  Paper  Industry.   (Presented  at  the  Miami  University Symposium  on
    Continuous Monitoring of Air and Water Quality, Oxford, Ohio, June 20, 1973.)
                                       17-42

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23. Determining Dust Concentrations in a  Gas Stream. Performance Test Code 27-1957.
    American Society of Mechanical Engineers, New York, New York, 1957.

24. Devorkin, H., Chass, R. L., and Fudvrich, A. P., Source Testing Manual. Air Pollution
    Control District, Los Angeles, California, December 1972.

25. Code of Federal Regulations, Part 60, Chapter I, Title 40. Standards of Performance
    for New Stationary Sources. Method 5. December 23, 1971.

26. Bosch, J. C., Pilat, M. J., and Hrutfiord, B. F., Size Distribution of Aerosols from a
    Kraft Mill Recovery Furnace. Tappi, 54:1871-1875, November 1971.

27. Walker, C. G., Manual for Counting and Sizing Particles from Kraft Recovery Furnaces.
    Atmospheric Pollution Technical  Bulletin No. 19.  National  Council of the Paper
    Industry for Air and Stream Improvement, New York, New York, July 1963.

28. Sem, G.  J.,  et al., State of the Art; 1971  Instrumentation for Measurement of
    Particulate Emissions from Combustion Sources— Volumes I & II: Particulate Mass.
    Reports  APTD  0733  and  0734,  Documents  PB 202 665  and  PB 202 666. U.S.
    Environmental  Protection  Agency, Air Pollution  Control  Office, Durham,  North
    Carolina, April 1971.

29. Leonard, J.  S., Continuous Kraft Mill Emission Monitoring. In: Blosser, R. O., and
    Cooper, H. B. H. (eds.)., Analytical Equipment and Monitoring Devices for Gases and
    Particulates. Atmospheric  Pollution Technical Bulletin No. 35, National  Council for
    the Paper Industry for Air and Stream Improvement, New. York, New York, March
    1968.

30. Tretter, V. J., Use of Continuous Monitors of Soda Loss and Malodorous Sulfur Loss in
    Process Control Tappi, 52:2324-2326, December 1969.

31. Larssen,  S., Ensor, D. S., and Pilat, M. J., Relationship of  Plume Opacity to the
    Properties of Particulate Emitted from Kraft  Recovery Furnaces.  Tappi, 55:88-92,
    January 1972.

32. Gansler,  N. R.,  The  Use of a Bolometer for Continuous Measurement of Particulate
    Losses from Kraft Recovery Furnaces. (Presented at the Annual Meeting of the Pacific
    Northwest International Section of the Air Pollution Control Association, Vancouver,
    British Columbia, November 22, 1968.)

33. Beutner, H.  P., Measurement of  Opacity and Particulate Emissions from  Stocks.
    (Presented at the Symposium on Instrumentation for Continuous Monitoring of Air
    Water Quality, Miami University, Oxford, Ohio, June 20, 1973.)

                                       17-43

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34.  Cooper, H. B. H., Ambient and Source Odor Measurements. (Presented at the Third
     Annual  Industrial  Air  Pollution  Control  Conference,  University of  Tennessee,
     Knoxville, Tennessee, March 29, 1973.)

35.  Cooper,  H.  B. H.,  and Rossano,  A. T., Particulate Matter and Odor Control.
     Unpublished  Special  Report,  University  of  Washington,  Department  of Civil
     Engineering, Seattle, Washington, January 1971.

36.  Physiological  Effects.  In:  Air  Pollution  Abatement  Manual.  Washington, D.C.
     Manufacturing Chemists Association, 1951.

37.  Byrd, J.  F., Phelps,  A. A., Odor and Its Measurement. In:  Stern, A. C. (ed.).  Air
     Pollution, Volume II, 2nd Edition. New York. Academic Press, Inc., 1968.

38.  Prince, R. G.  H., and Ince, J. H., The Measurement of Intensity of Odors. Journal of
     Applied Chemistry, 8:314-32, May 1958.

39.  Lindvall, T., On Sensory Evaluation  of Odorous Air Pollutant Intensities. Atmospheric
     Pollution  Technical Bulletin No. 50, National Council of the  Paper Industry for Air
     and Stream Improvement, New York, New York, September 1970.

40.  Caron, A. L.,  and Adams, D.  F., Evaluation of the Use  of Humans in Measuring the
     Effectiveness  of Odor  Control  Technology at  the Source. Atmospheric Pollution
     Technical Bulletin No. 36, National Council of the Paper Industry for Air and Stream
     Improvement, New York, New York, September 1971.

41.  Standard  Method for Measurement of Odors in Atmospheres (Dilation Method). ASTM
     Standard  D 139-57. Philadelphia: American Society for Testing and Materials, 1957.
     p. 185-188.

42.  Cederlof, R.,  Edfors, M. L.,  Friberg, L., and  Lindvall, T., Determination of Odor
     Thresholds for Flue Gases from a Swedish Sulfate Cellulose Plant. Tappi, 48:405-411,
     July 1965.

43.  Huey, N. A., Broering, L. C., Jutze, G. A., and Guiber, C. W., Objective Odor Pollution
     Control Investigations. Journal of the Air Pollution Control Association, 10:441-446,
     December 1960.

44.  Walther, J. E., and Amberg, H.  R., Experience with a Mobile Laboratory in Source
     Sampling Kraft Mill Emissions. Tappi, 51:126A-129A, November 1968.

45.  Mulik, J.  D., Stevens, R. K., and Baumgardner, R. A., An Analytical System Designed
     to  Measure Multiple  Malodorous Compounds Related to Kraft Mill Activities.  In:

                                       17-44

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     Proceedings  of  the  12th  Conference on  Methods in Air Pollution and Industrial
     Hygiene Studies. University of Southern California, Los Angeles, California, April 6-8,
     1971.

46.  Megy, J. A., Design, Operation, and  Use  of a Mobile Laboratory  for  Continuous
     Monitoring of Kraft  Mill Source Gases. (Presented at the West Coast Regional Meeting
     of  the  National Council  of the Paper Industry for Air  and Stream Improvement,
     Seattle, Washington,  October 15, 1969.)

47.  Waddington, G. E.,  A Mobile Ambient and Stack Sampling System for the Pulp and
     Paper Industry—A Unified Systems Approach.  Pulp and  Paper Magazine of Canada,
     74:58-61, June 1973.
                                       17-45

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                         APPENDIX A
                   GLOSSARY OF SYMBOLS
Symbol

B&W
BOD
BOD7
BLO
BLRBAC
BP
BHV
BD
BCRC
BTU
cm
CE
ACE
LAH
cfm
cfs
cu ft, ft3
m3
m3/h
°C
°F
°K
DO
DS
ft
fpm
fps
floz
gal
gpm
gr
g
h,hr
hp
in
Definition

Babcock & Wilcox
biochemical oxygen demand
biochemical oxygen demand as determined by 7-day test
black liquor oxidation
Black Liquor Recovery Boiler Advisory Committee
boiling point
bomb heat values
bone dry
British Columbia Research Council
British thermal unit
centimeter
Combustion Engineering
CE system of contact evaporation using combustion air
CE system of evaporation using Laminaire air heaters
cubic feet per minute
cubic feet per second
cubic foot
cubic meter
cubic meters per hour
degree Celsius
degree Fahrenheit
degree Kelvin
dissolved oxygen
dry solids
foot
feet per minute
feet per second
fluid ounce
gallon
gallons per minute
grain
gram
hour
horsepower
inch
                             A-l

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Symbol
Definition
ID
SI
kcal
kg
kj
km
kW
LK
MJ
MPa
t
/xm
mm
mg
mV
min
ng
nm
NCASI
NSSC
NCG
R
NOX
SOX
ppb
ppm
Ib
pcf
psia
psig
psi
RF

ton
SSL
std
scf
scm
sdcf
SBLO
induced draft
International System of Units ("metric system")
kilocalorie
kilogram
kilojoule
kilometer
kilowatt
lime kiln
megajoules
megapascal (= 106 kg • m/sec2 per m2)
metric ton
micrometer
millimeter
milligram
millivolt
minute
nanogram
nanometer
National Council for Air and Stream Improvement
neutral sulfite semi-chemical (process)
noncondensable gases
organic  chemical radical
oxides of nitrogen
oxides of sulfur
parts per billion
parts per million
pound
pounds per cubic foot
pounds per square inch, absolute pressure
pounds per square inch, gauge pressure
pounds per square inch
recovery furnace
second
short ton (2000 Ib)
spent sulfite liquor
standard
standard cubic foot
standard cubic meter
standard dry cubic foot
strong black liquor oxidation
                               A-2

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Symbol         Definition

TRS            total reduced sulfur
w.g.            water gauge
W              watt
WBLO          weak black liquor oxidation
yd              yard
yr              year
                               A-3

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                   APPENDIX B
              CHEMICAL  FORMULAS
Acetic Acid
Acetone
Acrolein
Ammonia
Ammonium bisulfite
Ammonium sulfate
Ammonium sulfite
Arsenic
Beryllium
Boric acid
Bromine, molecular bromine
Cadmium
Cadmium sulfate
Calcium bisulfite
Calcium Oxide (lime)
Calcium sulfite
Carbon dioxide
Carbon monoxide
Carbonic Acid
Carbonyl hydrogen sulfide
Carbonyl sulfide
Chlorine dioxide
Chlorine, molecular chlorine
Citric Acid
Dimethyl disulfide
Dimethyl Sulfide
Ethanol, Ethyl alcohol
Ferric chloride
Ferrous sulfide
Fluorine
Formaldehyde
Hydrobromic acid
Hydrochloric acid
Hydrogen, molecular hydrogen
Hydrogen peroxide
Hydrogen sulfide
Lead
CH3COOH
CH3COCH3
CH2CHCHO
NH3
NH4HS03
(NH4)2 SO4
(NH4)2 S03
As
Be
H3B03
Br, Br2
Cd
CdS04
Ca(HS03)2
CaO
CAS03
CO2
CO
H2C03
COSH
COS
C102
Cl, C12
C6H807
CH3SSCH3
CH3SCH3
CH3CH2OH
Fe2Cl6
FeS
Fl
HCHO
HBr
HC1
H,H2
H202
H2S
Pb
                       B-l

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Magnesium bisulfite
Magnesium hydroxide
Magnesium oxide, magnesia
Mercury
Methane
Methanol, methyl alcohol
Methyl mercaptan, methanethiol
Nitric oxide
Nitrogen dioxide
Nitrogen, molecular nitrogen
Nitrogen oxides
Oxygen, molecular oxygen, ozone
Phosphorus
Potassium acid phthalate
Selenium
Silver nitrate
Sodium bicarbonate
Sodium hydrogen sulfide
Sodium bisulfite
Sodium carbonate
Sodium chloride
Sodium chlorite
Sodium hydroxide
Sodium mercaptide
Sodium sulfate
Sodium sulfite
Sodium thio sulfate
Sulfur
Sulfur dioxide
Sulfur trioxide
Sulfuric acid
Vanadium
Vanadium pentoxide
Water
Zinc
Mg(HS03)2
Mg(OH)2
MgO
Hg
CH4
CH3OH
CH3SH
NO
N02
N,N2
NOX
0,02,03
p
rvHCg 0404
Se
AgN03
NaHC03
NaHS
NaHS03
Na2C03
NaCl
NaC102
NaOH
CH3CH2SNa
Na2S04
Na2S03
Na2S203
S
SO2
SO3
H2S04
V
V20S
H2O
Zn
                         B-2
                   U S GOVERNMENT PRINTING OFFICE 1976-660-859

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