United States EPA-600/R-04-107 Environmental Protection 

Agency August 2004 

Identification 

of (and Responses to) 

Potential Effects of SCR 

and Wet Scrubbers on 

Submicron Particulate 

Emissions and Plume 

Characteristics 


EPA-600/R-04/107 

August 2004 

Identification of (and Responses to)

Potential Effects of SCR and Wet

Scrubbers on Submicron Particulate

Emissions and Plume Characteristics

by

William E. Farthing, Peter M. Walsh, John P. Gooch

Joseph D. McCain, W. Scott Hinton, Robert F. Heaphy

Southern Research Institute

2000 Ninth Avenue South

Birmingham, Alabama 35205

Inter-Agency Agreement DW 64959501

EPA Project Officer: C. Andrew Miller

Office of Research and Development

National Risk Management Research Laboratory

Air Pollution Prevention and Control Division

Air Pollution Prevention and Control Division

National Risk Management Research Laboratory

Office of Research and Development

U. S. Environmental Protection Agency

Research Triangle Park, NC 27711
Abstract 

The addition of selective catalytic reduction (SCR) systems and wet flue gas desulfurization 

(FGD) scrubbers to coal-fired boilers has led to substantial reductions in emissions of 

nitrogen oxides (NOX) and sulfur dioxide (SO2). However, observations of pilot- and 

full-scale tests of these technologies reveal potential adverse side effects that may produce 

operational and particulate emissions problems. The indirect effects of SCR technology of 

immediate interest center around its catalytic enhancement of SO2 oxidation to sulfur 

trioxide (SO3) and subsequent increases of sulfuric acid aerosols, which can produce visible 

near-stack plumes and acid aerosol mists. This report summarizes the current 

state-of-the-science concerning SO3 formation processes and methods to minimize such 

formation. Corrosion of plant components such as air preheaters and other balance of plant 

problems are discussed, as are problems associated with operation of air pollution control 

equipment. Also discussed are the limitations of SO3 measurement methods and the 

resulting uncertainties in many existing SO3 emissions data. Size distributions and visible 

plume formation are discussed, including empirical predictions of conditions—such as coal 

type, sulfur content, and presence of pollution control equipment—that may lead to greater 

potential for visible plume formation. 

Four general approaches provide a high probability for successful SO3 removal: (1) alkali 

injection into the furnace, (2) humidification at the electrostatic precipitator (ESP) inlet to 

reduce the temperature to below the acid dew point, (3) alkali injection combined with 

humidification at the ESP inlet, and (4) separate wet particulate control device such as a wet 

ESP. All of these approaches have been tested in demonstration scale, but all must be 

considered new technology that will require adjustments or identifications of procedures to 

deal with specific sites. The first three have balance of plant issues associated with them; 

that is, increase risks of unforeseen negative effects on plant operation. Other concepts of 

interest, but less fully tested, include (1) alkali injection in duct leading to a wet FGD 

scrubber and (2) an electrostatically augmented mist eliminator. Various devices are being 

offered which have the purpose of contacting sorbents with flue gas for removal of SO3 as 

well as other gaseous pollutants. 

ii 
Foreword 

The U.S. Environmental Protection Agency (EPA) is charged by Congress with protecting 

the Nation’s land, air, and water resources. Under a mandate of national environmental laws, 

the Agency strives to formulate and implement actions leading to a compatible balance 

between human activities and the ability of natural systems to support and nurture life. To meet 

this mandate, EPA’s research program is providing data and technical support for solving 

environmental problems today and building a science knowledge base necessary to manage 

our ecological resources wisely, understand how pollutants affect our health, and prevent or 

reduce environmental risks in the future. 

The National Risk Management Research Laboratory (NRMRL) is the Agency’s center for 

investigation of technological and management approaches for preventing and reducing risks 

from pollution that threaten human health and the environment. The focus of the Laboratory’s 

research program is on methods and their cost-effectiveness for prevention and control of 

pollution to air, land, water, and subsurface resources; protection of water quality in public 

water systems; remediation of contaminated sites, sediments and ground water; prevention 

and control of indoor air pollution; and restoration of ecosystems. NRMRL collaborates with 

both public and private sector partners to foster technologies that reduce the cost of 

compliance and to anticipate emerging problems. NRMRL’s research provides solutions to 

environmental problems by: developing and promoting technologies that protect and improve 

the environment; advancing scientific and engineering information to support regulatory and 

policy decisions; and providing the technical support and information transfer to ensure 

implementation of environmental regulations and strategies at the national, state, and 

community levels. 

This publication has been produced as part of the Laboratory’s strategic long-term research 

plan. It is published and made available by EPA’s Office of Research and Development to 

assist the user community and to link researchers with their clients. 

Lawrence W. Reiter, Acting Director.

National Risk Management Research Laboratory

iii 
EPA Review Notice 

This report has been peer and administratively reviewed by the U.S. Environmental 

Protection Agency and approved for publication. Mention of trade names or commercial 

products does not constitute endorsement or recommendation for use. 

This document is available to the public through the National Technical Information 

Service, Springfield, Virginia 22161. 

iv 
Table of Contents 

Section Page 

List of Figures ........................................................... vi

List of Tables .......................................................... vii

Abbreviations and Acronyms ............................................... ix

1. Introduction ......................................................... 1

2. Executive Summary and Conclusions ..................................... 3

3. Conventional Experience with SO3 (Without SCR or Wet FGD) ............... 13

SO3 Formation in Boilers .............................................. 13

Ash Particle Size Components ....................................... 16

Air Preheaters .................................................... 19

Improved Electrical Resistivity and Cohesion in Electrostatic Precipitators ....... 20

Baghouses .......................................................... 21

Flue Gas Conditioning ............................................. 22

Removal of SO3 .................................................. 25

SO3 Effects on the Appearance of Stack Emissions .......................... 26

Background ...................................................... 26

Field Observation of Nucleation ...................................... 28

Survey of Scattering Coefficients for Stack Plumes ....................... 30

An Indicator of Particle Size Characteristics in Stack Plumes ............... 31

4. NOX Reduction Catalysts (SCR) ........................................ 33

Production of SO3, Relationships with Catalysts and Other SCR Parameters ...... 33

Ammonia Slip ....................................................... 36

Variation of SO3 Production and Ammonia Slip Over Catalyst Life ............. 37

Options for Controlling Ammonia and SO3 ................................ 38

Ammonium Sulfate and Bisulfate Aerosol ................................. 39

5. Sulfuric Acid Aerosol Formation and Collection in Wet FGD ................. 49

6. SO3/H2SO4 Measurement and Monitoring ................................. 55

7. Control Technology Issues ............................................. 59

8. References .......................................................... 65

v 
List of Figures 

Figure Page 

1. Processes contributing to formation of sulfur trioxide, sulfuric acid, and sulfates.

After Walsh. ........................................................ 14

2. Particle size distributions of fly ash exiting a pulverized-coal-fired utility

boiler. ............................................................. 17

3. Dew point curves for gas containing different concentrations of water vapor at

total pressure of 1 atm. ................................................ 20

4. Fabric strength retention based on Mullen Burst Strength of the fabrics in

compartments conditioned with anhydrous ammonia and hydrated lime. ......... 25

5. Specific light scattering coefficient per particle volume (left axis) and contributions

of the outlet particulate matter of Figure 2 (right axis). ....................... 27

6. Measured particle mass versus size in stack and in plume formed in plume

simulator. .......................................................... 29

7. Oxidation of SO2 to SO3 (a) for several SCR catalysts at 700 and 750 °F and

corresponding NOX reduction (b) at 700 and 620 °F from a pilot scale investigation

at a coal-fired utility boiler burning 3%-sulfur coal. ......................... 34

8. Distribution of ammonium bisulfate among ultrafine particles, fly ash, and air

preheater surface and corresponding droplet sizes versus distance along the

channel wall......................................................... 44

9. Evolution of the distribution of ammonium bisulfate between ultrafine particles

and fly ash and growth of the ultrafine particles by coagulation during 5 s at

150 °C (302 °F) ..................................................... 45

10. ESP inlet and outlet particle size distributions of fly ash with and without a

nucleated ammonium bisulfate mode expected from downstream of an air

preheater when an SCR is present. ....................................... 46

11. Growth by coagulation of nucleated sulfuric acid droplets from H2SO4 molecules. . 50

12. Scattering coefficients and average specific scattering coefficients of nucleated

sulfuric acid droplets versus time for the various H2SO4 concentrations of

Figure 11. .......................................................... 51

13. Average specific scattering coefficients of nucleated sulfuric acid droplets

versus H2SO4 concentrations for various times from Figure 12. ................ 52

vi 
List of Figures (continued) 

Figure Page 

14. Predicted ground level concentrations of particulate H2SO4 for illustrative 

conditions representing a stack effluent exhausting a wet scrubber for SO2 

control of coal-fired boiler emissions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53 

15. Concentration vs particle size for fly ash constituents, scrubber solids 

constituents, and sulfuric acid measured in the stack of a coal-fired utility 

boiler whose emissions were controlled by an ESP followed by a wet scrubber. . . 57 

16. Fine PM concentrations on a cumulative mass basis for several pollution control 

systems. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61 

vii 
List of Tables 

Table Page 

1. Illustration of Potential Effects of SO3/H2SO4 on Fine Particle Emissions and

Visible Emissions ...................................................... 9

2. Conversion of coal sulfur to sulfuric acid and inorganic sulfate. ................. 15

3. System Categorizations ................................................ 59

viii 
Abbreviations and Acronyms 

Term Definition

A/C fabric filter air-to-cloth ratio (also face velocity), fpm 

acm actual cubic meter 

Atm atmosphere 

bscat scattering coefficient 

CCM controlled condensation method for measuring SO3 

D50 particle size at which collection efficiency is 50% 

dM/dLogD derivative of particulate mass concentration by particle size with respect to 

the logarithm of particle diameter. (Also the slope of the cumulative 

particle concentration smaller than diameter curve when plotted on a 

semi-log graph.) 

dncm dry normal cubic meter ( Normal conditions are 20 °C at 1 atm.) 

dV/dLogD derivative of particulate volume concentration by particle size with respect 

to the logarithm of particle diameter. 

dγ/dLogD derivative of the scattering coefficient by particle size with respect to the 

logarithm of particle diameter. 

EPA Environmental Protection Agency 

EPRI Electric Power Research Institute 

ESP electrostatic precipitator 

FGD flue gas desulfurization 

L path length 

MW megawatt 

Op opacity 

PC pulverized coal 

PCD pollution (or particulate) control device 

PM particulate matter 

ppmv parts per million by volume 

RH relative humidity 

SCA ESP specific collection area, ft2/1000 acfm 

SCR selective catalytic reduction 

SEM scanning electron microscope 

ix 
Abbreviations and Acronyms (continued) 

Term Definition

slip the amount of ammonia that passes unreacted through an SCR catalyst bed 

SRI Southern Research Institute 

γ scattering coefficient = the total effective particle light-scattering cross 

sectional area per unit of path length 

γ* ratio of a particle’s effective scattering cross sectional area to its volume 

γ* ave average specific scattering coefficient over all particle sizes 

x 
Identification of (and Responses to) Potential

Effects Of SCR and Wet Scrubbers on Fine

Particulate Emissions and Plume Characteristics

1. Introduction 

Substantial reductions of emissions of nitrogen and sulfur oxides, which are precursors of 

ambient fine particulate matter, from coal-fired boilers have been accomplished through 

improvements in combustion practices [such as burners that reduce oxides of nitrogen 

(NOX)], application of scrubbers, and increased use of low sulfur coals. Further substantial 

reductions of these precursors of ambient fine particulate matter will be accomplished by 

widespread installation of selective catalytic reduction (SCR) systems and additional wet 

flue gas desulfurization (FGD) scrubbers. These control systems can provide low NOX and 

sulfur dioxide (SO2) emission levels while also relaxing the limitations of fuel sulfur levels. 

However, observations associated with exploratory research of SCR and FGD technologies 

and some early full-scale installations reveal some important side effects, which have 

produced operational and particulate emissions problems of a yet-to-be-determined degree 

and scope. 

The indirect effects of SCR technology on process operation that are of immediate interest 

center around its catalytic enhancement of SO2 oxidation to sulfur trioxide (SO3) which, in 

turn, reacts through multiple potential paths that can interfere with air heater performance, 

particulate collection, ash disposal, or balance-of-plant issues. [Although SO3 vapor is 

converted to sulfuric acid (H2SO4) vapor as temperature declines from 650 to 400 °F, most 

discussions refer to all forms as SO3. The present discussion utilizes both terms 

interchangeably.] Associated particulate emissions impacts of interest are substantial 

increases of sulfuric acid aerosols, which form in the wet scrubber or in the plume near the 

stack if a wet scrubber is not present. These condensation aerosols are visible and persist 

over long distances, in part, due to their particle size and, in part, due to the retention of 

substantial fractions of water. 

The low temperature of wet scrubber effluent produces less plume rise than occurs 

otherwise with hot stack gas. Furthermore, stacks for use in association with scrubbers are 

1
usually designed to operate at lower gas velocities than those for hotter stack gases in order 

to minimize entrainment of water droplets from the stack walls. This lower gas velocity 

further reduces plume rise. Less plume rise leads to higher local, ground-level 

concentrations of stack gas constituents, relative to plumes formed from hotter stack gas, 

which rise to high elevations because of high buoyancy. In other words, with lower 

temperature stack gas (around 150 °F) the degree of dilution that occurs prior to plume 

contact at the ground is greatly diminished relative to stack gas temperatures greater than 

250 °F. Persistent wet FGD scrubber plumes have been observed and referred to as plume 

blight. Physiological responses such as eye irritation and breathing difficulty have been 

reported1

 for episodes involving visible plumes containing sulfuric acid aerosols at ground 

level. 

Besides increased visible emissions associated with SO3, the increased emphasis in recent 

years on condensable emissions (such as emissions for which EPA Method 202 is utilized) 

leads to it being counted as a fine particulate emission. Therefore, it may be necessary that 

contributions of SO3 aerosols to coal and oil-fired effluent gas produced by SCR 

installations be taken into account in strategies for satisfying the ambient fine particle 

standard. 

This document summarizes, analyzes, and interprets currently available information 

concerning SO3. These results are needed for delineating its impacts on multi-pollutant 

control strategies to be applied to coal-fired utility boilers for utility and regulatory 

personnel. It is organized to first review experience with SO3 prior to and independent of 

pending changes concerning NOX and SO2 control. Parameters that effect its formation in 

furnaces, interactions with particulate control devices (PCD’s), and effects on appearance of 

stack emissions are discussed. Then SCR characteristics and operating procedures are 

discussed with emphasis on consistency across the industry. Next, principles of wet FGD 

scrubbers are reviewed. Concerns with SO3 measurement and the need and potential for 

online monitoring are discussed next followed by consideration of SO3 control. A summary 

and conclusions are given immediately below. 

2
2. Executive Summary and Conclusions 

Due to past experience with SO3, it was recognized early in the development of SCR 

technology that catalyst formulations and operating strategies for NOX control by SCR were 

needed that kept SO3 formation as low as practical. Sulfur trioxide formation can be reduced 

by using a larger volume of catalyst but this increases capital cost. The conversion rate of 

SO2 to SO3, in terms of percent of SO2, can be specified and achieved with confidence. The 

uncertainties of SCR operation primarily affect the tradeoffs between ammonia slip 

(unreacted ammonia that remains in the flue gas downstream of the SCR), NOX reduction, 

and the frequency with which one or more catalyst layers must be replaced. In principle, 

SCRs are operated at a specified minimum level of NOX reduction that satisfies a systems 

NOX compliance strategy, while increasing ammonia injection as needed to maintain the 

required degree of NOX control. As the catalyst ages, its NOX reduction performance 

degrades, and eventually catalyst must be replaced or additional catalyst must be added. In 

the latter instance, the added catalyst will result in an increase in SO3 formation. 

Of course, SO3 catalyzed in an SCR installation adds to that formed in the boiler. 

Equilibrium favors conversion of SO2 to SO3, but it is limited by kinetics to that catalyzed 

by some of the transition metals, such as iron and vanadium, in the fuel and on boiler tube 

surfaces. Alkali (alkaline and alkaline earth metals) in the coal counteracts this conversion 

so that for coals high in alkali, such as most western coals, SO3 levels exiting the boiler (i.e., 

exiting the economizer) are low. Indeed, SO3 levels generated in the boiler are so low with 

western coals (around 0.1% of the SO2) that SO3 is commonly injected upstream of 

electrostatic precipitators (ESPs) to reduce ash resistivity to levels that allow them to 

operate properly. Typically, SO3 conversion rates with low alkali coals are around 1% for 

pulverized coal (PC) boilers, although as much as 3% is not uncommon. Generally, efficient 

combustion conditions while minimizing excess air minimizes SO3 formation. Conversion 

rates with cyclone boilers burning eastern bituminous coals are about twice that of PC 

boilers. 

Significant losses of SO3, typically about 50%, occur across air preheaters, the first active 

system components encountered by the flue gas after an SCR, due to the cool metal 

surfaces. Much of this loss is the result of condensation on the air heater elements with 

evaporation into the incoming combustion air returning most of the condensate to the boiler. 

It is not clear whether the fractional loss is independent of the SO3 concentration. If the 

process is diffusion limited, as is believed to be the case, the estimate of 50% loss is 

applicable to the increase in SO3 from the SCR. Otherwise the fraction removed may be 

higher. Additional losses of SO3 occur across the PCD due to its adsorption by fly ash 

3
particles and condensation onto particles. Typical values of these removal fractions are 

approximately 50% for ESPs and 90% for baghouses. 

Sulfur trioxide reacts with water vapor in the air preheater to become H2SO4. There are plant 

operational concerns with corrosion from acid deposits on air preheater surfaces and in the 

PCD(s). In addition to corrosion of metal surfaces, bag failure rates in fabric filter 

installations have been shown to increase as the amount of H2SO4 on particles increases. The 

amount deposited in these devices can be minimized by maintaining flue gas temperatures 

higher than the H2SO4 dew point, but doing so carries a penalty in energy efficiency and 

may result in a more visible plume. 

Effects of ammonia slip on baghouse operation are also of concern. Injection of ammonia at 

levels of approximately 20 ppmv into flue gas upstream of a baghouse was attempted to 

extend bag life. It was postulated that the ammonia reaction with H2SO4 would prevent acid 

attack of the fabric. The result was that ammonia injection led to a faster loss of fabric 

strength than incurred with H2SO4 alone. The cause is believed to be formation of 

ammonium bisulfate, a sticky liquid substance of low pH, formed by reaction with H2SO4. It 

appears that the industry standard for ammonia slip will be 2 ppmv. This standard has 

evolved to limit ammonia in collected fly ash to avoid interference with its disposal or use 

as a byproduct. 

With this standard, an expected upper limit of 5 ppmv of ammonia (NH3) is appropriate. 

These expected levels are significantly less than those found to weaken fabric strength but 

still high enough to cause concern. In this same test program, injection of a small amount of 

hydrated lime into the baghouse was found to eliminate the loss in bag strength associated 

with H2SO4 or ammonium bisulfate. Alkali injection removes essentially all SO3 in 

baghouses where gas contact with particle surface is high as the flue gas passes through a 

substantial dust cake. This practice is expected to become common in operation of 

baghouses to optimize preservation of bag strength after SCR. 

Model calculations of aerosol formation for 5 ppmv of ammonia slip and 40 ppmv SO3 at 

the exit of an SCR (or air preheater inlet) were performed to predict the size and 

concentration of particles that would form by reaction of the NH3 and SO3. The results 

indicated that ammonium bisulfate formation dominates over ammonium sulfate (a less 

sticky solid), and most of it nucleates to ultrafine (less than 0.2 :m diameter) particles that 

exit the air preheater. These ammonium bisulfate particles would cause a substantial 

increase in the concentration of sub-micrometer particles entering any downstream PCD. 

4
A computer model of ESP performance2,3 was then used to estimate the effect of these added 

particles on net particulate emissions. The model was used to predict emissions from a unit 

that employed a typical cold-side ESP for particulate control by assuming the ammonium 

bisulfate aerosol component was added to a typical coal fly ash ESP inlet particle size 

distribution. The ESP model results showed that space charge would be substantially 

affected by the amount of added ultrafine aerosol predicted by the aerosol formation model, 

and the ESP current would decrease significantly for a given applied voltage prior to SCR 

installation. However, since installation of an SCR will lead to a net increase in SO3 entering 

the ESP, fly ash resistivity would be at least as low as it was prior to SCR installation and 

probably would be lower. Therefore, the applied voltage could be increased, depending on 

the power supply limitations, to reinstate the ESP current. This would result in a slight net 

improvement in collection efficiency. The opacity of the exit stream was predicted to 

decrease from 11% before SCR installation to 9% afterward for the case modeled. Instances 

of field observation of improved ESP performance with NH3 injection have been reported.4 

The levels of ammonia slip considered here, for which no degradation of ESP performance 

is indicated, are expected to be representative of most SCR installations. If NH3 and SO3 

levels present at the air preheater inlet are allowed to be substantially larger, then other 

factors should be considered. Tests performed in the field have included conditions in which 

NH3 injection led to deposition of ammonium bisulfate particles on collection plates and 

corona electrodes leading to enough particulate build-up to interfere with ESP performance. 

Thus, ESPs downstream of SCRs should be monitored for any hints of long-term 

degradation due to accumulation of ammonium bisulfate on corona and plate electrodes that 

are not removed by normal rapping. 

Removal of SO3 in wet FGD scrubbers is thought to be low because consideration of aerosol 

formation mechanisms suggests that it forms a submicrometer mist at scrubber inlets. This 

conjecture has been borne out by measurements. In a number of instances, more particles 

having diameters around 0.1 :m have been found at scrubber outlets than at the respective 

inlets when SO3 was present in the inlet gas stream.5

 The apparent cause is a rapid 

quenching of the SO3 bearing flue gas, leading to nucleation and formation of ultrafine 

particles. In addition, persistent scrubber plumes are most frequently associated with flue 

gas streams known to have SO3. In spite of the predicted low collection efficiencies of these 

ultrafine particles, attempts at direct measurement of SO3 removal using current methods 

indicate significant collection, typically in the range of 30% to 80%.6

 As discussed below, 

there are good reasons to question SO3 measurements at scrubber outlets. As a result, actual 

collection efficiencies of SO3 by scrubbers should be considered to be poorly quantified, and 

5
reported values over 50% should be considered especially suspect. Measured particle 

collection efficiencies over the relevant particle size range for low energy scrubbers (or 

low-pressure loss, 1" to 2" H2O) utilized for FGD are not available, although scrubber D50 

(size at which collection efficiency is 50%) is typically around 4 :m. Review of measured 

particle collection efficiency over the relevant particle size range for high energy scrubbers 

often indicates that some collection, albeit well below 50%, is achieved at all sizes below 

the D50. 

The inconsistent measurements of SO3 removal efficiency indicated above were obtained 

using the controlled condensation method (CCM). Since the early 1980’s, this method has 

been widely believed to produce the most accurate results for SO3 measurement at the 

concentrations encountered in flue gases from coal-fired boilers. The CCM method was 

developed primarily for measurement of vapor phase SO3. However, current needs are for 

measurement of total vapor and particulate phases. It is widely believed that much of SO3 

present in the exit gas stream of air preheaters and beyond is adsorbed to fly ash. Further, 

alternate, experimental measurements using cascade impactors for SO3 downstream of 

scrubbers have, on occasion, shown substantial acid levels in particles having diameters 

larger than ten micrometers. These results suggest substantial particle growth, driven by the 

huge capacity for water by sulfuric acid, can qualitatively be expected in moisture-saturated 

gas, such as that in exiting scrubbers if the gas has sufficient residence time for growth by 

diffusion. The presence of acid incorporated with particles larger than a micrometer could 

explain the inconsistent SO3 results by CCM downstream of wet FGDs as these may not be 

sampled correctly by the CCM method. On the other hand, the inconsistent results may be 

caused by reaction with droplets of scrubber liquor in the CCM sampling probe. 

Considering the mechanisms for producing SO3 and removing it from flue gas, it is obvious 

that the appropriate design specification for conversion rate from SO2 to SO3 in a given SCR 

is specific to parameters of the boiler and the other components of the effluent gas stream. 

For boilers using low to medium sulfur coal (less than 1.5% sulfur) the typical conversion 

rate specifications for SCRs are in the 0.5% to 1.5% range. It is expected that 0.5% to 0.75% 

will be selected for many of those boilers using high sulfur coal except, perhaps, for those 

for which explicit removal systems for SO3, discussed below, are available. Note that these 

specifications refer to levels expected after SCR catalysts have been in service to the extent 

that all design catalyst layers are installed. Actual levels of conversion are expected to be 

close to, but below, the specified levels. 

Given the regulatory impetus expected from various proposed new laws and rule-making 

6
regarding pollutant emissions (Clear Skies, Mercury Rule, and Interstate Air Quality Rule), 

the use of wet FGD scrubbers is expected to become widespread over the next 5 to 10 years. 

As a result, stack plumes will generally be at lower altitudes due to lower stack exit 

velocities and reduced plume temperature and buoyancy. Thus, under certain conditions, the 

plume will contact the ground near the stack (occasionally within 1 mile and more 

frequently within 5 miles). In that type of circumstance, characterization of the plume 

appearance in terms of opacity, as is normally used for source emission, is not reliable. 

There would be great uncertainty and variability of an appropriate path length (plume 

diameter) through which an observer would view the plume and, of course, occasionally an 

observer could be in the plume. Characterization of ambient air frequently uses the 

scattering coefficient (meters to the minus one) parameter7

, which is also a part of opacity 

considerations, to quantify the appearance (light scattering) of ambient aerosol. In the 

ambient air context this term is called “bscat.” 

There are two features about visual observations of stack emissions that are relevant: 

attenuation of light passing through the plume as measured by opacity and light scattered 

from all directions incident on the plume into the direction of the observer. When below 5% 

to 10%, opacity is difficult to measure quantitatively and becomes an insensitive monitoring 

parameter. Scattered light is still readily visible when opacity is below levels of its 

usefulness for monitoring. It is common with modern, large particulate control technology 

that opacity is well below 10% so that the plume appearance is dominated by scattered light. 

Therefore, this effort has focused on scattered light although the measure of it is the same as 

addressed in opacity concerns. Light absorption by coal-fired boiler ash is rarely significant, 

so scattering is the normal cause of opacity. 

The effects on fine particle and visible emissions by the numerous factors involved in SCR 

and wet FGD scrubbers are illustrated in Table 1. The quantities in the table are based on 

typical values currently available for the listed production and removal mechanisms. These 

results are presented assuming no specific efforts at controlling SO3 beyond current practice 

and low alkali coal. The conditions of the various effluent gas streams that were considered 

are listed in the first column. The next five columns present levels of SO3 expected to exit 

each of the components of those streams using the assumptions of typical conversion rates 

(SO2 to SO3) and removal efficiencies (as provided in the footnotes to the table). The three 

levels of coal sulfur considered—low, medium, and high—correspond to assumed values of 

800, 1500, and 3000 ppmv SO2 produced from the coal. The seventh column lists the 

corresponding fine particulate matter (PM) expected to be emitted derived from the mass 

concentration of H2SO4 per ppmv (4.03 mg/dncm per ppmv at dry normal conditions). This 

7
value does not count water that combines with H2SO4 in PM measurement methods. Since 

the rates of SO2 to SO3 conversion considered are coincidentally both one percent for boiler 

and SCR production, half of the fine PM shown represents an increment associated with 

SCR. 

In order to characterize the effects of SO3 on the appearance of emissions in this analysis, 

the scattering coefficient was resolved into two additive terms: the scattering coefficient for 

PM not associated with SO3 and that associated with SO3. Table 1addresses only the latter. 

(The appearance of the emissions will also be affected by the emissions of fly ash and NOX.) 

The scattering coefficient of an aerosol is determined by both its concentration and particle 

size distribution. For the purposes of Table 1, the SO3 contribution to the scattering 

coefficient was resolved into two factors: acid (H2SO4 and combined water) volume 

concentration and the average specific scattering coefficient. The latter was quantified 

separately based on estimated particle size and refractive index. The value of this size 

dependent component is determined by aerosol formation and growth mechanisms. These, in 

turn, depend on temperature conditions and the concentrations of H2SO4 and water. The 

mechanisms include nucleation, heterogeneous condensation onto preexisting fly ash, and 

coagulation. For hot stack effluents, consideration was given to both modeling and field data 

derived from air-dilution sampling systems to quantify the average specific scattering 

coefficient under two scenarios: (1) that in which heterogeneous condensation dominates 

(higher values given in the table notes) and (2) that in which homogeneous nucleation 

dominates. The latter case is associated with relatively low fly ash loadings presenting 

relatively little particle surface area. 

For stack effluents from wet FGD systems, the average specific scattering coefficient was 

calculated based on estimated lower limits for concentrations by particle size resulting from 

nucleation followed by coagulation. These estimates simulate nucleation of acid at the inlet 

of a wet FGD and the coagulation that follows as a function of residence time in the 

scrubber and stack. Residence time in the scrubber and stack is assumed to be 20 seconds 

for Table 1. These results are presented in more detail later in Section 5 of this report. The 

amount of water included with H2SO4 was 65% by weight of the two combined components. 

This weight percentage corresponds to ambient conditions of 20 °C at 70% relative humidity 

(RH). Acid droplets in a scrubber plume will reach those conditions in ambient air, but 

while in the scrubber and stack, where the RH is 100%, the amount of absorbed water is 

much larger even though the temperature is about 60 °C. Thus, these droplets actually grow 

by absorption as well as coagulation. A larger droplet diameter enhances coagulation 

somewhat, so it is expected that the actual size would be somewhat larger in diameter and 

8
Table 1. Illustration of Potential Effects of SO3/H2SO4 on Fine Particle Emissions 

and Visible Emissions 

Coal Type and Control Devices 

H2SO4 Concentrationa Exiting 

Effluent Stream Components 

(ppmv) 

H2SO4 

Fine PM 

H2SO4 Contribution to 

Plume Scattering 

Coefficienta 

(mS1) 

Boilrb SCRc Air 

Perhtrd PCDe Wet FGDf 

(mg/dncm) 

Emissions 

Near 

stackg 

Int. 

medh 

(×10S5) 

Disti 

(×10S5) 

Low sulfur, ESP only 8 NAk 4 2 NA 8 0.006 0.4 

Low sulfur, ESP only & low fly ashj 8 NA 4 2 NA 8 0.002 0.1 

Medium sulfur, ESP only 15 NA 8 4 NA 15 0.011 0.7 

Medium sulfur, ESP only & low fly ashj 15 NA 8 4 NA 15 0.004 0.2 

High sulfur, ESP only 30 NA 15 8 NA 31 0.022 1.4 

High sulfur, ESP only & low fly ashj 30 NA 15 8 NA 31 0.007 0.5 

Low sulfur, ESP & SCR 8 8 8 4 NA 16 0.012 0.8 

Low sulfur, ESP, SCR & low fly ashj 8 8 8 4 NA 16 0.004 0.2 

Medium sulfur, ESP & SCR 15 15 15 8 NA 31 0.022 1.4 

Medium sulfur with ESP, SCR & low fly ashj 15 15 15 8 NA 31 0.007 0.5 

High sulfur, ESP & SCR 30 30 30 15 NA 61 0.045 2.9 

High sulfur, ESP, SCR & low fly ashj 30 30 30 15 NA 61 0.014 0.9 

Low sulfur, ESP, SCR & wet FGD 8 8 8 4 3 11 H2O fog 3.1 0.3 

Medium sulfur with ESP, SCR & wet FGD 15 15 15 8 5 21 H2O fog 10.5 1.0 

High sulfur, ESP, SCR & wet FGD 30 30 30 15 11 43 H2O fog 37.7 3.8 

a For reference, the scattering coefficient of clean ambient air is 10S5

 mS1 and a value > 24×10S5 mS1 is considered an adverse 

condition7

. b Alkali typical of eastern coals with SO3 production in convection sections of boiler . 1% of SO2. c SO3 production in SCR. SO3/SO2 ratio . 1% for all sulfur levels (Note: the ratio will be about 0.75% for many boilers using high 

sulfur coals). d Air preheater H2SO4 removal . 50%. e ESP H2SO4 removal . 50% or baghouse . 90%. f Wet FGD scrubber removal . 30%. g Hot stack plume (without wet FGD); temperature of 60 °C, water vapor volume fraction of 3%, combined water . 35% by weight, 

45% by volume. Particulate volume of H2SO4 is 0.00362 cm3

/m3/ppmv. The expected range of average of the specific scattering 

coefficient is 0.2 to 2.5 m2

/cm3. Estimated here as 2.07 and 0.66 respectively for submicron fly ash loadings that are typical (>2 

mg/dncm) and low (<2 mg/dncm). Effluent gas diluted by 2.5. To convert value to opacity multiply by 2.5 × stack diameter (in 

meters). h With scrubber, temperature of 20 °C, water vapor volume fraction of 1.5%, combined water . 65% by weight, 76% by volume. 

Particulate volume for H2SO4 is 0.00925 cm3

/m3/ppmv. Estimated average specific scattering coefficient given by γ*ave = 

–0.0078C[H2SO4, ppmv]2

 + 0.4514C[H2SO4, ppmv], effluent gas diluted by 1000. 

Same conditions as Note “h” except flue gas diluted by 10,000. Estimated average specific scattering coefficient of Note “f “ 

(without scrubber) and of Note “g“ (with scrubber). j Low fly ash refers to low submicron loadings (< 2 mg/dncm). All other submicron loadings >> 2 mg/dncm. See Note “g.” 

i 

K NA = not applicable. 

9
have correspondingly larger average specific scattering coefficients. For example, under 

these conditions (see Section 5), it has been shown that average specific scattering 

coefficients follow a quadratic in H2SO4 concentration with coefficients which depend upon 

the residence time in the scrubber and stack, 20 s for the current example. 

The results given in Table 1 illustrate how the relevant variables, primary and secondary, 

contribute to the overall fine PM mass and appearance of PM emissions. Since the values for 

each individual variable are estimates of typical values, the frequency with which the 

indicated results occur cannot be predicted. However, to interpret these results, it is useful to 

compare them to general observation as some gauge as to the quality of the estimates. In 

addition, the circumstances for which there need not be much cause for concern may be 

indicated. Table 1 suggests that installation of SCR will increase opacity (product of near 

stack scattering coefficient and plume diameter). Considering installations without wet 

FGD, if a very efficient ESP is used for particulate control, so that fly-ash submicrometer 

loadings are low, thus promoting nucleation, the estimated opacity increase ranges from 1% 

to 8% as coal changes from low sulfur to high sulfur. With higher fly ash loadings, the 

corresponding estimated range of the increase is from 4% to 19%. This potential increase 

should be considered in choosing an SCR specification for SO2 conversion rate or in 

considering separate SO3 removal in conjunction with SCR installation. Assistance in those 

deliberations can come from pilot tests at the site using air-dilution sampling systems (also 

called plume simulation sampling systems) that mimic the processes that take place in the 

plume near the stack together with comprehensive modeling. Sites that burn high alkali 

coals or that use a baghouse for the PCD will likely experience essentially no change of 

emissions. 

Because wet FGD scrubber effluents can contact the ground relatively close to the stacks, 

installation of a scrubber (without SO3 removal) could lead to haze on the ground within 

several miles of the plant. The column labeled “Intermediate” represents an occasion when 

wet FGD plume contacts the ground at a distance on the order of one mile from the stack 

when dilution has progressed to 1,000. The results suggest that haze would result at this 

location when firing a high sulfur coal. Note the reference point concerning appearance 

given in Note “f” of Table 1. For lower sulfur coals, the impact would be marginally 

noticeable from a visual perspective. At “Distant” locations, no haze is predicted for low 

sulfur based upon the estimates of typical values presented here. In the distant plume, all 

conditions considered here lead to scattering coefficients that are near clean air conditions 

except “High Sulfur with ESP & SCR.” Consideration of the order of magnitude of the 

results suggests the potential importance of variations of variables from the typical values. 

10
For example, if most of the variations in SO3 production and removal rates were in the 

direction to increase SO3, then a very different picture emerges in which environmental 

impact could be more substantial. 

In those cases in which plume blight from SO3 is expected to be a problem, each of four 

general approaches provide a high probability for successful SO3 removal: (1) alkali 

injection into the furnace, (2) humidification at an ESP inlet to reduce the temperature to 

below the acid dew point, (3) alkali injection combined with humidification at the ESP inlet, 

and (4) a separate wet particulate control device such as a wet ESP. All of these approaches 

have been tested in demonstration scale, but all must be considered new technology that will 

require adjustments or identifications of procedures to deal with specific sites. The first 

three have balance of plant issues associated with them; that is, increase risks of unforeseen 

negative effects on plant operation. Other concepts of interest, but less fully tested, include 

(1) alkali injection in ducting leading to a wet FGD scrubber and (2) an electrostatically 

augmented mist eliminator. Various devices are also being offered which have the purpose 

of contacting sorbents with flue gas for removal of other gaseous pollutants as well as SO3. 

11
12
3. Conventional Experience with SO3 (without SCR or Wet 

FGD) 

SO3 Formation in Boilers 

The formation of sulfuric acid in a boiler occurs when a small fraction of the sulfur in the 

fuel is oxidized beyond SO2 to SO3, which subsequently reacts with water vapor to form 

H2SO4. Almost all of the sulfur is oxidized to SO2 in the combustion zone; indeed, at high 

temperature, SO2 is the equilibrium form. As the gas temperature falls between the 

combustion zone and the stack plume, shifts in the equilibria occur, illustrated in Figure 1, 

from SO2 (gas) Y SO3 (gas) Y H2SO4, (gas) Y H2SO4-H2O (liquid). Thermodynamically, 

SO2 is the most stable species in the combustion zone; SO3 is the most stable species in the 

intermediate-temperature zone between the economizer and air preheater; sulfuric acid 

vapor is the most stable species between the air preheater and the stack or wet scrubber (if 

present); and diluted liquid sulfuric acid aerosol is the most stable species in the wet 

scrubber and stack effluent. Of course, the levels of these species other than SO2 are much 

lower than that of SO2 itself, which remains the predominant fraction of the sulfur present 

due to kinetic factors that limit the key oxidation step from SO2 to SO3. Therefore, in 

addition to fuel sulfur and excess combustion air levels, the rate of formation of SO3 

depends on catalytic oxidation associated with some metals in the ash and on the boiler tube 

surfaces, especially iron and vanadium.8,9,10,11 Coals also contain alkaline metals, especially 

calcium, sodium, and magnesium, that can react with and limit the net sulfuric acid as 

illustrated in the lower right of Figure 1. 

Reported boiler oxidation rates for SO2 to SO3 run as high as 5% of the SO2 for coal-fired 

boilers8

 and 10% for residual oil-fired boilers9,12, while typical rates for low alkali coals are 

near 1%.8 When one considers utility boilers burning eastern 3%-sulfur fuels that produce 

SO2 levels greater than 2000 ppmv, quite high levels of SO3 or H2SO4 might be encountered, 

but typical levels are much lower due to several factors associated with plant operation. 

These factors include minimization of excess air, control of boiler tube fouling, fuel 

selection, and other plant operating procedures applied to limit corrosion, sulfur emissions, 

and visible emissions associated with condensed sulfuric acid. 

Walsh et al. measured SO2 oxidation by the vapor phase homogeneous reaction at the 

furnace exit of a stoker boiler to be 0.9% and suggested this observation to be near an upper 

limit set by the SO3/SO2 equilibrium at high furnace temperatures.12 Graham and Sarofim 

estimated this uncatalyzed contribution at 0.5% for a pulverized coal (PC) boiler.10 

13
Figure 1. Processes contributing to formation of sulfur trioxide, 

sulfuric acid, and sulfates. After Walsh.11 

Interaction of SO2 with catalytic compounds and SO3 with alkaline compounds involves gas 

interfacing with boiler tube surfaces, with bare metal or deposited ash, or with suspended 

(aerosolized) fly ash. Walsh, et al. investigated catalysis of SO2 to SO3 by iron in and on the 

boiler tubes considering clean tubes, tubes fouled with coal fly ash, and the combination. 

Their calculations from first principles suggest a range of conversion rates from 1% to 5% 

depending on the degree of tube fouling and amount of iron in the coal ash.12 Tube fouling 

led to higher surface temperatures of deposited ash, thus enhancing the rate of catalysis by 

Fe2O3 present in the ash, 13% of the ash mass in the instance studied. Clean tubes led to 

lower tube surface temperatures and limited catalysis to lower rates, but the amount of 

available iron was higher at the clean tube surface. The 1% (lowest) value of calculated 

conversion rate corresponded to no tube fouling, and the heavy tube fouling condition 

resulted in a 2% value. The 5% (highest) value corresponded to combined clean and fouled 

tubes. 

The modeling results by Walsh et al.12 for catalytic conversion associated with iron in 

suspended fly ash particles was 0.3%. If applied to PC boilers, this modeling result for 

suspended fly ash would be substantially higher. That conclusion follows from the 

apportioning of ash between bottom ash and flue gas typical for PC boilers, 20%/80%, 

versus that for stoker boilers, 90%/10%. 

14
Although Walsh, et al.12 conclude a total conversion rate as high as 6% for SO2 to SO3, 

measurements showed that two-thirds of it was removed in transit to and through the 

economizer of the boiler. Uptake by ash particles, by processes other than simple 

condensation, was suggested as the mechanism for SO3 removal. Graham and Sarofim10 

investigated sulfation by alkaline metals in the ash as an important mechanism that limits 

SO3 exiting PC boilers. The importance of alkaline metals was illustrated by referencing 

results from a laboratory study by Neville13 as shown in Table 2. The variation in conversion 

rate over coal type from 8.9% with Western Kentucky coal down to negligible net 

conversion for Montana lignite was explained in terms of competition between catalytic 

activity by iron oxide and subsequent sulfation by alkaline metal oxides, especially calcium, 

sodium, and magnesium in coal. These net conversion rates are considered only as upper 

limits for these coals because of the high oxygen levels utilized in this laboratory study. 

Table 2. Conversion of coal sulfur to sulfuric acid and inorganic sulfate.a 

Coal Type Coal Sulfur 

(%) 

Sulfur Converted 

to Sulfuric Acid 

(%) 

Sulfur Converted to 

Inorganic Sulfate 

(%) 

Alabama Rosa 0.85 3.4 2.8 

Illinois #6 4.1 6.2 2.17 

Montana Lignite 0.55 0. 34.3 

PSOC - 3b 0.69 4.7 1.6 

PSOC - 26b 3.95 4.1 1.4 

Western Kentucky 3.15 8.9 1.4 

a Laboratory coal study; combustion conditions: 75% oxygen, furnace temperature = 

1200 K. b From the Pennsylvania State Coal Sample Repository. 

The data showing sulfur converted to sulfuric acid in Table 2 were derived from pH data of 

distilled water after adding freshly generated coal ash that was collected at low temperature 

relative to the acid dew point. The pH of the water decreased rapidly to a minimum in about 

2 minutes and then rose substantially over a period of a few hours. The sulfuric acid 

condensed or adsorbed on the ash was calculated from this minimum pH. The subsequent 

rise in pH was caused by the dissolution into the distilled water of alkali components in the 

fly ash. In similar experiments measuring sulfuric acid on fly ash from a coal-fired boiler 

burning 3% sulfur Eastern bituminous coal, it was found that after the pH dropped to levels 

below 4, it rose in about 10 minutes to levels over 11.14 Thus, perhaps there is more than 

15
enough alkaline material in a given fly ash to neutralize formed acid, but the amount located 

on the fly ash surface structures may not be sufficient to fully neutralize it. 

Due largely to the abundance of alkaline metals in Western coals, SO3 formation is 

dramatically lower than for low alkali, Eastern bituminous coals. Hardman et al.8 identified 

three categories with SO2 to SO3 conversion rates that correlate inversely with levels of 

alkaline metals. A recent update of that initial investigation added lignite as a fourth 

category.15 The estimated values offered by these authors are 0.008 for Eastern bituminous, 

0.001 for Western bituminous, 0.001 for lignite, and 0.0005 for Western subbituminous 

coals when burned in PC boilers. In addition, this latter investigation concluded that cyclone 

boilers have conversion rates about double those of PC boilers. 

Ash Particle Size Components 

The degree of interaction of SO2 with catalytic compounds and of SO3 with alkaline 

compounds is related to a set of processes that determine the distribution of compounds by 

particle size in fly ash exiting boilers. These processes are the nucleation, condensation, and 

coagulation of vapor-phase species in the presence of other particles or droplets suspended 

in the flow. These processes are included in this discussion because they affect not only the 

vapor-phase SO3 concentration in the boiler effluent gas stream but also subsequent 

conversion to particulate sulfate components that subsequently affect if and where these 

sulfate components are removed in any given flue gas stream. In the boiler the relevant 

compounds are the metal species formed from vaporized ash constituents. 

Figure 2 illustrates typical particle size distributions for fly ash entering and leaving an ESP 

controlling emissions from a PC boiler. The ESP inlet measurements were obtained by 

combining results from a University of Washington Mark III cascade impactor operated 

in-situ (for concentrations at sizes above 0.5 :m) and a TSI Scanning Mobility Particle Sizer 

(Model 3934) with extracted, dried, and cooled sample gas (for concentrations at sizes 

below 0.7 :m). 

Metal oxide fume is formed from metal atoms and metal oxides generated in the high 

temperature reducing environment of burning char residues in pulverized coal flames 

(Quann and Sarofim,16 Senior and Flagan,17 and Helble et al.18). (Char is the remains of a 

solid carbonaceous fuel that has been incompletely combusted, such as charcoal if wood is 

incompletely burned.) Metal oxide fume is the principal source of the typical ultrafine 

particle mode (peak) centered around 0.1 :m as illustrated in Figure 2. The latter is an 

example typical of observed particle size distributions and loadings of coal fly ash exiting 

16
Figure 2. Particle size distributions of fly ash exiting a 

pulverized-coal-fired (PC) utility boiler. (The outlet size distribution 

was calculated from the measured inlet distribution using an ESP 

model applied to a typical ESP setup at typical flue gas conditions 

with 1% sulfur, Eastern-bituminous coal.) 

tangentially-fired PC boilers. These data were obtained during combustion of Eastern 

bituminous coal with 0.8% sulfur. Iron, calcium, and magnesium are among the elements 

vaporized from ash during combustion of bituminous coals (Quann and Sarofim16 and 

Helble et al.18). Iron oxide is an effective catalyst for oxidation of SO2 to SO3; therefore, the 

particle size distribution and surface area of iron-containing particles are of some 

importance in determining the conversion of SO2 to SO3 in the convection section. Finely 

divided calcium and magnesium oxides are effective sorbents for SO3 and H2SO4, so a shift 

of these elements to smaller size is likely to reduce the concentration of free acid (Graham 

and Sarofim10). 

The vaporization of inorganic species from burning char was treated in detail by Quann and 

Sarofim.16 In their mechanism, volatile metal atoms and metal suboxides are produced by 

17
reduction of the full oxides by carbon monoxide inside the char. On leaving the char, these 

vapor species encounter a more oxidizing environment and return to their normal oxidation 

states, which have extremely low vapor pressures. The products of this reaction are, 

therefore, supersaturated and homogeneously nucleate to form nanometer-size particles of 

the solid metal oxides, such as Fe3O4, CaO, and MgO. These ultrafine primary particles 

grow by coagulation in the boundary layers of char particles and in the free stream during 

the passage of combustion products through the furnace, approaching a mean size in the 0.1 

to 0.3 μm range that are characteristic of the accumulation mode. 

Combustion, and the processes described above, generate a particle size and composition 

distribution at the furnace exit consisting of three major components: (1) unburned char 

particles having sizes of the order of 100 :m, (2) ash particles having sizes on the order of 

10 :m that are formed from extraneous mineral particles in the fuel and ash released during 

burnout of the char, and (3) fine particles formed by nucleation, condensation, and 

coagulation of the most volatile inorganic elements in the ash, as described above. 

Depending on coal type and ash composition, iron, calcium, and magnesium may be present 

in both the large ash and fine particles. The char may have high internal surface area and 

absorb sulfuric acid at low temperatures. Coagulation and growth of submicrometer particles 

continues in the convection section, possibly with small additions to their mass from 

condensation of volatile trace metals as the flue gas is cooled. Superimposed on this 

complex system are the losses of vapor species, fine particles, and coarse particles by 

condensation, thermophoretic transport, and inertial impaction, respectively, to heat transfer 

surfaces in the convection section. 

In laboratory simulations Graham and Sarofim10 investigated the portion of SO3 formation 

and sulfation of alkaline compounds associated with the fine aerosol ash component in 

boilers. Based on projections to PC boiler conditions of laboratory measurements of 

catalytic rates by 0.1-:m Fe2O3 particles, the condensed iron component (in the fine mode) 

of fly ash was estimated to contribute a 0.5% conversion rate for typical PC boiler 

conditions. This particle size mode presents the most effective catalysis per unit mass of iron 

but contains a small fraction of the total iron. The authors’ calculations addressed only the 

Fe2O3 ash fraction found in the sub-micrometer mode of the size distribution, which 

represented only 0.5% by weight of the total fly ash while typical Fe2O3 ash fractions range 

from about 5% to 20% of the total fly ash. The authors characterized typical total conversion 

rate of SO2 to SO3 as 1% to 3%, including that occurring at heat transfer surfaces and other 

particles making up the total of the fly ash. In addition to catalytic oxidation by Fe2O3 

particles, Graham and Sarofim10 measured sulfation rates by 0.015 :m CaO particles. Based 

18
on the measured rates, ranging from 10S7

 to 3×10S6 moles CaO/m2s over temperatures from 

400 to 600 °C and up to 1.5×10S3 at 1000 °C for 3000 ppmv SO2 and 5% O2, it was 

concluded that all such particles in this size mode are sulfated at typical PC boiler 

conditions. However, for the current context of impact on SO3 levels, it is not clear how 

much of the sulfation occurs in reaction with SO2 rather than SO3. 

Air Preheaters 

Substantial reduction in SO3 and/or sulfuric acid is common across air preheaters where the 

average gas temperature decreases from about 650 to 325 °F. Dew point curves are given in 

Figure 3. High losses by condensation on metal surfaces are readily understandable 

qualitatively, considering these dew point data combined with the fact that the average metal 

surface temperature varies from about 140 to 270 °F across the plane located at the last 0.3 

m of a rotary-regenerative preheater basket.19 SO3 levels downstream of air preheaters are 

typically observed to be significantly below the acid dew point levels corresponding to the 

average flue gas temperature. Several investigators suggest reduction of SO3 across air 

preheaters when burning Eastern bituminous coal by a factor around 50%.8,15,19,20 With other 

coals, for which much less SO3 is produced in the boiler, one available estimate is 10%.15 

DeVito and Oda20 reported results of parallel SO3 measurements at a preheater flue gas inlet 

and outlet and in the combustion-air side of an air preheater. They found that the 40% of the 

SO3 entering the air preheater was removed from the flue gas and attributed it to 

condensation of H2SO4 onto the air preheater surfaces.20 Measurements performed in the 

combustion air stream passing through the air preheater on its way to the furnace led to the 

conclusion that H2SO4 removed from the flue gas was revolatilized in the combustion air 

stream. 

In passing through air preheaters, flue gas temperatures decrease through the range in which 

SO3 is converted to H2SO4. Significant portions are in both the particulate and vapor phases 

as reflected by observations of ash pH like those discussed above and by the inverse 

dependence of fly ash resistivity on SO3 as characterized by investigations of ESP 

operation.21 Zhang22 performed fundamental calculations (for rotary-regenerative preheaters) 

of both H2SO4 vapor condensation on wall surfaces and transfer to suspended particulate 

matter by homogeneous nucleation, heterogeneous condensation, and adsorption by fly ash. 

All of these mechanisms were found to contribute significantly to the fate of SO3. 

The analysis of Hardman and Dismukes8

 reported above, characterizing SO3 levels, was 

based upon empirical data at the air preheater exits of 18 conventional PC boilers. 

Coefficients of variation of these field data were around 50%; from this, one can surmise 

19
Figure 3. Dew point curves for gas containing different

concentrations of water vapor at total pressure of 1 atm.23

that levels a factor of two higher or lower than the typical values used in the present analysis 

are not unusual. For example, with SO2 levels greater than 2000 ppmv (resulting from the 

use of Eastern bituminous coal), 4 to 16 ppmv of SO3 or H2SO4 could be expected at air 

preheater exits of PC boilers. Higher and lower levels cannot be ruled out because of the 

uncertainty associated with catalytic oxidation and sulfation in boilers or varying losses in 

the air preheater. Indeed, these conditions and higher variations from the norms of SO3 

values have been observed at some sites over three decades of investigation. 

Improved Electrical Resistivity and Cohesion in Electrostatic 

Precipitators 

Many electrical utility plant operators have long dealt with the presence of sulfuric acid. 

Some level of acid aerosol has been found to be beneficial because condensed or adsorbed 

acid reduces fly ash resistivity, resulting in more nearly optimum ESP performance. In fact, 

for many units burning low sulfur coal, especially those burning western coals with high 

alkali levels, SO3 is injected upstream of the ESP to improve collection efficiency by 

reducing resistivity. However, the dependence of resistivity on SO3 level is subtle. It has 

been found that, to be accurate, the empirical relationship developed for fly ash resistivity 

requires as input data the concentrations of nine common metals plus sulfate in the ash as 

well as SO3 concentration, moisture content, and temperature of the flue gas.24 

20
In addition to reducing resistivity, sulfuric acid condensed or adsorbed on the fly ash can 

also be beneficial in reducing ESP rapping reentrainment by increasing particle cohesion. 

Air preheater flue-gas exit temperatures are usually kept high enough to limit levels of 

condensed-phase acid (see Figure 3), thus both minimizing corrosion rates in, and 

downstream of, the air preheaters and maintaining the basically dry particulate properties 

necessary to insure adequate removal of collected particulate matter from ESP corona wires 

and plates by rapping. 

Essentially all of the particle phase acid is removed in the ESP. The fraction that passes 

through the ESP as vapor varies widely due to the variation of gas temperature in the ESP 

and variation of acid adsorption/reaction with fly ash composition and temperature. 

Hardman and Dismukes8 estimated 25% of the acid exiting the air preheater is removed by 

cold-side ESPs, but a recent update of that review15 resulted in an estimated 50% removal. 

Baghouses 

Performance and operating characteristics of fabric filters differ significantly between units 

burning low sulfur, high alkali Western coals and high sulfur, low alkali Eastern coals. The 

difference is such that separate and parallel research facilities were needed to concentrate on 

each type of flue gas.25 Sulfuric acid (or SO3) level is the most important variable associated 

with observed differences. 

A strong tendency toward heavier residual dust cakes is associated with high sulfur coal. 

Abundant sulfates in the residual dust cake implicate SO3 in the formation of these cohesive, 

tenacious dust cakes. Residual dust cake weights of 1 lb/ft2 (100 lb/bag) were encountered in 

some high-sulfur coal baghouses.26 This bag weight is a factor of 20 times the weight of dust 

accumulated during a single filter cleaning-cycle.27 In one of these extreme cases, the heavy 

residual dust cake exceeded the mechanical design capacity and caused failure of the bag 

suspension mechanism.28 High residual dust cake weight is the result of high particle 

cohesion and adhesion. Comparison of scanning electron micrographs (SEMs) of dust cakes 

from a Western low sulfur subbituminous coal and an Eastern high sulfur bituminous coal 

revealed the presence of cementitious structures in the latter that did not appear in the 

former. This term, cementitious structure, was ascribed to contiguous material bridging fly 

ash particles together and believed to be the cause of enhanced cohesion.29 Qualitative 

analysis by the SEM indicated that this bridging material was composed of more sulfur and 

less of other metals than individual fly ash particles. Separate analyses of dust cake samples 

and hopper samples showed up to a factor of 5 higher levels of soluble sulfate in the dust 

cakes. (That is, the residual ash that was not removed from the bags by the normal cleaning 

21
process had a much higher sulfur content than that which fell into the hopper during 

cleaning).29 These observations implicate condensed sulfuric acid as the probable cause of 

the increased cohesion. In addition to the observed correlation of residual dust cake weight 

to sulfur in the ten coals, it inversely correlated with both sodium and calcium.26 

Higher residual dust cake weights after cleaning have a positive correlation with more SO3. 

However, when normalized for differences in gas flow rate and dust cake thickness, the 

pressure drops are lower by a factor of 3 to 4 compared to normalized pressure drops in 

low-sulfur coal fabric filters. It follows that the dust cake collected in the presence of SO3 

forms a more porous structure.26 Thus, lower pressure drops are often seen in high-sulfur 

coal fabric filters even though residual dust cakes are much heavier than in low-sulfur 

installations. 

Woven fiberglass is the material of choice for bag construction. The fiberglass yarn is 

coated with a finish that is intended to protect the yarns from abrasion. The finish may also 

provide some protection from acid attack. Some yarn finishes have been offered specifically 

for resistance to acid. High dust cake weights add significant stress to bag fabric. More 

energetic means of cleaning bags, instead of the predominate method of reverse gas, have 

been investigated to reduce residual filter cake weights (specifically, sonic horns and 

shake/deflate), but these also add stress to the filter fabric.27 

Fabric strength is characterized by the Mullen Burst Strength, a fabric property that is 

measured by the method specified in ASTM D3786. The test measures the ability of a fabric 

to resist rupture as the result of pressure exerted by a hydraulically inflated diaphragm.30 

Research has correlated specific bag failures with the loss of fabric strength as measured 

with the Mullen Burst Strength test.31 The loss of fabric strength was attributed to the 

condensation of H2SO4 in the fabric/dust cake structure. It should be pointed out that, 

although the loss of fabric strength can be associated with failed bags, it is not possible to 

reliably predict bag life on the basis of the residual Mullen Burst Strength of the fabric. Bags 

that had been in service for as much as ten years have been observed to remain functional 

even though they had lost a large fraction of their original Mullen Burst Strength. The end of 

bag life is determined by more variables than incorporated in the Mullen Burst Strength test. 

Flue Gas Conditioning 

Relevant insight has been gained in three studies in which flue gas conditioning was 

attempted on order to improve baghouse performance. Sulfur trioxide was injected as a 

conditioning agent, along with anhydrous ammonia, at one site while hydrated lime, 

22
anhydrous ammonia, or a mix of the two were injected at two boilers burning high sulfur 

coal. 

SO3 Injection to Improve Dust Cake Cohesion, Porosity, and Particulate Collection 

Efficiency 

A low-sulfur lignite coal produced fly ash with very low cohesion, very high pressure drop, 

and a propensity to penetrate fabrics such that opacity exceeded 20%.32 In tests utilizing 

ammonia injection (25 ppmv), it was determined that particulate collection efficiency 

improved significantly (80% opacity reduction), while average drag experienced a moderate 

decrease (20%). In an attempt to attain still further reduction of pressure drop and emissions 

comparable to the baghouse design specification, ammonia injection was augmented with 

SO3 injection (15 ppmv SO3 with 31 ppmv ammonia) in a six month test. Pressure drop was 

reduced by 40% to 50% relative to a parallel baghouse without flue gas conditioning, and 

the emissions were effectively eliminated with combined SO3 and ammonia.33 Weights of 

conditioned bags became approximately 40% higher than bags without flue gas 

conditioning. Near the end of the trial, Mullen Burst Strength measurements were performed 

both on bags exposed to conditioned flue gas and on bags without conditioning. Both types 

of bags had been installed at the beginning of the trial. No significant difference in bag 

fabric strength was discerned.33 

After the trial with combined ammonia and SO3 conditioning ended, conditioning with 

ammonia continued using the same bags. Seven months later, large numbers of bags began 

to fail. Fabric strength measurements were then performed on another set of bags. Those that 

had not been exposed to conditioned flue gas exhibited Mullen Burst Strengths averaging 

398 psig. The bags that had been conditioned with combined ammonia and SO3 flue gas 

exhibited areas of differing strengths associated with an orange coloration on the clean side 

of bags. Areas without the coloration had average strengths of 374 psig whereas areas with 

coloration had average strengths of 147 psig. 

The orange coloring was determined to be a manifestation of the deposition of elemental 

selenium. Selenium deposition was also observed at the two other sites where ammonia 

conditioning, described below, was used. An analysis of these dust cakes and these events 

yielded the following reaction as a probable mechanism to explain the collection of 

selenium on the bag surfaces:34 

3SeO2(g) + 4NH3(g) = 3Se(s) + 2N2 + 6H2O(g) 

23
Preservation of Fabric Strength When Exposed to Flue Gas from High Sulfur Coal 

Two projects involved efforts to reduce the rate of reduction in bag strength associated with 

high-sulfur coal by injecting reactants to neutralize sulfuric acid, which had flue gas 

concentrations of about 10 ppmv at both sites. Injection of reactants eliminated SO3 from the 

downstream flue gas in both investigations. 

Compartment scale tests of ammonia injection lasting 24 months were performed at a 

full-scale baghouse controlling emissions from a boiler burning a high-sulfur, Eastern 

bituminous coal. Fabric strengths of bags from test compartments and a control 

compartment were performed after 6, 11, 13, and 25 months. Fabric strengths started at 640 

psig at the beginning of the tests. Strengths of bag fabrics from the test compartments and a 

control compartment declined at the same rates through the first 13 months. However, at the 

end of 25 months, the strength of control bags averaged 180 psig while that of the test 

compartment averaged 70 psig.28 Thus, in this instance, flue gas conditioning with ammonia 

did not reduce the rate of loss of bag strength but may have actually contributed to 

weakening of the fabric over the last 12 months of the investigation. 

The other investigation of flue gas conditioning to neutralize sulfuric acid was part of a 

multi-year pilot scale project to characterize the performance of fabric filters in service with 

high-sulfur coal. Performance comparisons were carried among three compartments, one of 

which employed ammonia conditioning, the second employed conditioning with powdered 

hydrated lime, while the third treated unconditioned flue gas as a control. The feed rates of 

the reagents were set at stoichiometric ratios ranging from 2:1 to 10:1 relative to SO3, but no 

variation of results were observed for ratios above 2:1.31 

Flue gas conditioning with ammonia produced marked differences in the performance of the 

fabric filter. Ammonia conditioning altered the filtering characteristics of the dust cake to 

reduce pressure drop. The apparent mechanism for this effect was a reaction of ammonia 

with SO3 to form an ammonium bisulfate/ammonium sulfate complex that increased the 

porosity of the dust cake as it was formed during the filtering cycle. However, the collected 

ash was more difficult to remove so that residual dust cake weight increased. In addition, the 

acidic ammonium bisulfate promoted acid attack to weaken the fabric as measured by its 

residual Mullen Burst Strength. 

Conditioning with hydrated lime had little effect on pressure drop relative to the control 

compartment even though the residual dust cake weight was much lighter in the 

compartment conditioned with hydrated lime. Actually, pressure drop relative to the control 

24
compartment was slightly higher. 

Differences in post-exposure fabric strength, depicted in Figure 4, were attributed to 

differences in flue gas conditioning. In this figure, fabric strength retention is the ratio of 

Mullen Burst Strength at a given time to its initial value (at test initiation). Superior strength 

retention in the bags conditioned with hydrated lime suggested that removal of SO3 

protected the fabric from acid attack. Fabric strength in the compartment conditioned with 

ammonia deteriorated faster than in the control compartment. 

Figure 4. Fabric strength retention based on Mullen Burst Strength 

of the fabrics in compartments conditioned with anhydrous 

ammonia and hydrated lime. 

Removal of SO3 

Without considering alkali injection or other flue gas conditioning to protect bags, Hardman 

and Dismukes8

 estimated the fraction of sulfuric acid removed from flue gas by baghouses 

to be 90%. Experiences such as those described above demonstrate that SO3 and some SO3 

mitigation methods have a significant role in the performance of fabric filters in utility 

applications. The potential for heavy bags, low pressure drop, and reduced bag life 

associated with the presence of SO3, are factors in the operation of a baghouse. Replacement 

bags are usually the most significant cost encountered in maintaining a baghouse. However, 

25
conditioning with hydrated lime may dramatically reduce loss of bag strength and extend 

bag life. When this mitigation method is implemented, baghouse removal efficiency for SO3 

is essentially 100%. 

SO3 Effects on the Appearance of Stack Emissions 

Although sulfuric acid in electrical power plant emissions has not been regulated directly, 

there has been motivation for many coal and oil-fired boilers to minimize SO3 levels to 

avoid high visibility emissions. For those boilers with hot flue gas in the stack (i.e., without 

wet scrubbers), there are sometimes lower levels of opacity measured in the stack than in the 

plume above the stack. This increase in opacity, sometimes described as a “detached 

plume,” is the result of acid vapor being converted to the particle phase resulting from flue 

gas cooling in the plume during mixing with ambient air. The purpose of the discussion to 

follow is to arrive at an estimate or range of the change in visible emissions for a given 

increase of sulfuric acid vapor. 

Background 

There are two features about visual observations of stack emissions that are relevant: 

attenuation of light passing through the plume as measured by opacity; and light scattered 

from all directions incident on the plume into the direction of the observer. When below 5% 

to 10%, opacity is difficult to measure accurately, but scattered light is still readily visible. 

Opacities well below 10% are common with modern, large particulate control technology, 

so the plume appearance is dominated by scattered light. Therefore, this discussion focuses 

on scattered light although the measure of it is the same as addressed in opacity concerns. 

Light absorption by coal-fired boiler ash is rarely significant, so scattering is the cause of 

opacity. 

The scattering coefficient, γ, of flue gas or a plume exiting the stack represents the total 

effective particle light-scattering cross sectional area per unit of path length. When the depth 

or thickness (L) of the flue gas or plume is small enough such that γL < 0.2, the intensity of 

scattered light is directly proportion to γL. Opacity (Op) is related to the scattering 

coefficient by Op = 1 – T = 1 – e–γL

 where T is transmittance, the fraction of a light beam 

not scattered. Note that opacity is also given by γL when γL < 0.2. While light scattering 

varies dramatically with particle size, the scattering coefficient characterizes the net 

combined effects of scattering variation with size and the distribution of particle sizes for a 

particular sample. The magnitude of scattering by a particle is expressed in this discussion in 

terms of ratio of the particle’s effective cross sectional area to its volume. This ratio is called 

the specific scattering coefficient, γ*, and has units of reciprocal length. If particle size is 

26
expressed in micrometers, γ* has units of inverse-micrometers (:m2

/:m3, or :mS1). A plot of 

γ* versus particle diameter is shown in Figure 5. The differential scattering coefficient, 

dγ/dLogD, is the product of γ* and the distribution of particulate loading of the effluent with 

particle size on a volume basis, dV/dLogD. It is convenient to express dV/dLogD in units of 

cubic centimeters per cubic meter (cm3/m3) and γ* in reciprocal micrometers (:mS1)so that 

the resulting unit of the product without further conversion is in units of reciprocal meters 

(mS1), the unit normally employed for scattering coefficient, γ. Consideration of the 

differential scattering coefficient is useful to characterize the relative contributions of the 

aerosol size components to the overall scattering coefficient. Figure 5 shows dγ/dLogD for 

the example ESP outlet particulate loading and size distribution of Figure 2 (dM/dLogD, 

particle density of 2.35 g/cm3

). In this example, about half of the total γ would derive from 

particles around 1 :m, and the larger sizes, the bulk of the particulate volume (or mass), 

would contribute the other half. Note that dividing γ by V, the total particulate volume 

concentration, results in the average of γ* over size, weighting it by the normalized 

particulate volume size distribution. For the example, V is 0.026 cm3/m3, so γ/V is 0.77 

m2/cm3. 

Figure 5. Specific light scattering coefficient per particle volume 

(left axis) and contributions of the outlet particulate matter of Figure 

2 (right axis). 

27
When acid vapor is converted to particulate phase, heterogeneous condensation onto 

preexisting particles takes place as a diffusion limited process. Acid is condensed in 

proportion to particle diameter for particles larger than the gas mean free path (about 0.08 

:m) and proportional to particle area for those smaller than the mean free path.35 

Dependence on particle diameter means that the heterogeneous condensation rate across the 

particle size spectrum is weighted toward fine and ultrafine particles relative to the particle 

mass distribution. If cooling of the vapor is fast then the equilibrium vapor pressure of the 

acid can decline much faster than actual vapor pressure due to limitation by diffusion to 

available particle surfaces. Then supersaturation becomes high enough to result in 

homogeneous nucleation at the molecular level. The very large number of particles 

produced in the nucleation process rapidly grows by coagulation. The number concentration 

of particles in this mode is very much larger than the number concentration of the 

preexisting particles. For example nucleation of 10 ppmv of H2SO4 produces particle 

number concentrations on the order of 1016 per cubic meter when their size has grown to 

approximately 0.05 :m, while the total number concentration of preexisting particles is on 

the order of 5×1011 per cubic meter.36 In such cases, coagulation is characterized as 

self-preserving, dominated by self coagulation, in which the mode grows with the mean size 

increasing while the width of the mode is constant35 (1.32 geometric standard deviation37). 

Self coagulation (i.e., between particles within this mode formed by nucleation) dominates 

as opposed to coagulation with preexisting particles because of the dominance in particle 

number in this mode. Of course, particles in this nucleation mode also provide surface area 

for homogenous condensation in competition with heterogeneous condensation on 

preexisting particles. This example suggests that nucleation deposits more acid in ultrafine 

particle sizes (less than 0.2 :m) than occurs by heterogeneous condensation, depending 

upon the magnitude of the ultrafine mode of the fly ash. In view of the large variation of 

specific scattering coefficient with particle size as shown in Figure 5, it is obvious that the 

impact of sulfuric acid vapor on the appearance of stack emissions depends substantially on 

the fraction nucleated versus the fraction condensed to preexisting particles. Within the 

latter, the weighting of condensation rates as a function of particle size is important, 

resulting in relatively lower scattering if condensation is to sizes smaller than 0.2 or larger 

than 2 :m. 

Field Observation of Nucleation 

Detailed measurements of plume characteristics near a stack are impractical. Therefore, 

plume simulation hardware and procedures were developed for surrogate measurement.38,39 

A specific example that appears informative was obtained by this method during a field test 

at a coal-fired boiler.38,40 Figure 6 gives particle size distribution and mass of the aerosol in a 

28
coal-fired boiler stack, measured with cascade impactors, along with that in the “plume” 

formed in the air-dilution plume simulator measured with an electrical mobility aerosol 

analyzer. For dilution and cooling in the plume simulator, flue gas was extracted from the 

stack through a fine particle separator, maintained hot, and injected vertically upward at the 

bottom-center of a vertical, 8-inch-diameter mixing-chamber with concentric ambient air 

dilution flow. The hardware was designed to provide a residence/mixing time of 10 seconds 

with a dilution ratio of 20:1. The large increase from stack gas to plume simulator in the 

number of particles at sizes smaller than 0.1 :m (see Figure 6) is indicative of a strong 

component formed by nucleation and coagulation. A particulate mass concentration of 2 

mg/dncm in the stack for particles smaller than 2 :m became 41 mg/dncm in the “plume” 

through some combination of condensation and homogeneous nucleation. The given 

“plume” values are relative to stack conditions, before dilution, corrected to dry normal 

conditions (1 atm., 20 °C). The measured total loading is consistent with the measured 6.5 

ppmv SO3 by the controlled condensation reference method in the stack with 30% combined 

water. 

Figure 6. Measured particle mass versus size in stack and in 

“plume” formed in plume simulator. (Sample from a coal-fired boiler 

ESP exit at 300 °F with 6.5 ppmv of H2SO4 vapor was diluted so as 

to realistically simulate a near-stack plume.) 

There was no visible stack plume during the tests in which these data were obtained. 

Consideration of plume appearance or light scattering (in terms of γL) leads to consideration 

of plume concentrations, dilution, and thickness during observations. As flue gas exits the 

29
stack, it spreads and slows in velocity as well as being diluted and cooled. The spreading, 

slowing, and cooling contribute to an increase of path length, L, which tends to compensate 

the decrease in γ caused by dilution. DeFries41 investigated the degree of compensation for 

conditions of no gas to particle conversion and concluded that use of stack or undiluted 

concentrations and stack diameter to be adequate for estimating plume values of γL (less 

than 16% positive error). Note that the dilution process slows the process of particle growth 

by coagulation. The calculated scattering coefficient for the plume, including liquid-phase 

acid (acid droplet density of 1.6 gm/cm3), as depicted in Figure 6 in this example, was 0.006 

mS1. This result includes 0.0003 mS1 contributed by coarse particles totaling 22.5 mg/dncm 

(ash density of 2.35 gm/cm3

) that were separated from the flue gas passed to the plume 

simulator but were in the actual plume. For a stack diameter of 5 m, the calculated scattering 

produced an opacity of 3%, a low value consistent with no visible plume. This contribution 

to scattering coefficient by the nucleated acid in this example was 0.004 mS1

 for 6.5 ppmv 

acid, or 0.22 :mS1 for the average γ*, the specific scattering coefficient of the acid droplets 

for this case. 

Survey of Scattering Coefficients for Stack Plumes 

A comprehensive theoretical model of particulate matter formation in emission plumes has 

been developed (Damle, Ensor, and Sparks42) that provides qualitative agreement with 

observations. Unfortunately, field data for validation purposes are unavailable in which SO3, 

visible emissions, particulate size distribution, temperatures, flow rates, and moisture are 

quantified simultaneously. Even so, this model is considered useful as a survey tool. Results 

have been reported using the model to predict the effects of changes in acid vapor 

concentration. Other parameters varied were fine particle concentration, coarse particle 

concentration, water vapor, plume temperature, wind speed, and ambient humidity and 

temperature.41,43 In one series of comparisons, baseline conditions were defined that included 

100 mg/m3

 of coarse particles, 10 mg/m3 of fine particles, 10 ppmv of sulfuric acid vapor, 

and reasonable values of 6 other stack parameters, 5 other ambient parameters, and 

geometric standard deviations of the 2 preexisting particle size modes. The geometric mean 

size of these two modes were 0.15 and 4 :m on particle number basis, or 0.33 and 10 :m on 

mass basis, for the fine and coarse modes, respectively. For the baseline conditions, the 

increase in the average specific scattering coefficient, γ*ave, resulting from condensed acid 

was 2.5 :mS1, indicating heterogeneous condensation contributed significantly in the 0.5 to 2 

:m range. Other parameters were varied individually and resulted in little change in γ*ave 

except in the cases of reducing the fine particle loading to 1 mg/dncm and increasing wind 

speed to 20 m/s from 5 m/s. In both of these cases, the increase in γ*ave was reduced to 0.67 

:mS1 (from 2.5 :m S1). In the first instance, less heterogeneous condensation to the fine 

30
mode was probably the reason for a smaller increase in γ*ave. In the second instance, 

increased nucleation is suspected. Another series of calculations was performed with the 

fine mode loading reduced to 2 mg/dncm and varying acid vapor from 10 to 40 ppmv. In 

each of these instances, the increases in γ*ave were only around 1.2 :m S1 (from 2.5 :m S1 in 

the base case). Decreasing the fine particle loading to zero and varying acid vapor between 

10 and 40 ppmv produced increases in γ*ave of 0.2 :m S1 at 10 ppmv sulfuric acid, 0.3 :m S1 

at 30 ppmv, and 0.62 :m S1 at 40 ppmv. This increase in γ*ave with increasing acid vapor was 

necessarily the result of predicted nucleation. Another test case assigned 2 and 10 mg/dncm 

to the fine and coarse modes and produced a predicted increase in the scattering coefficient 

per volume of acid of 0.6 :m S1

. 

A correlation seems to exist concerning the increase in γ*ave due to condensed sulfuric acid 

vapor in stack plumes and the original concentration of fine particles in the stack. The 

apparent range of expected values of γ*ave is an order of magnitude, 0.2 to 2.5 m S1

, but 

where a specific site lies within that range is determined primarily by the loading of in-stack 

fine particles. In general, lower fine particle loadings cause lower values within that range of 

increase in the scattering coefficient, presumably due to increased nucleation. A high fine 

particle in-stack loading is 10 mg/dncm, and a low value is 2 mg/dncm. High coarse particle 

loadings limit values of the increase in scattering coefficient per unit volume in the upper 

part of the range. Coarse particle loadings are generally negligible when fine particle 

loadings are low because control strategies for fine particles tend to incidentally control 

coarse particles very well. 

An Indicator of Particle Size Characteristics in Stack Plumes 

The appearances of stack plumes are frequently reported as blue in color. This effect is the 

result of scattering of blue light by particles less than 0.3 :m being dominant over scattering 

at other wavelengths. This blue tint indicates an unusually high concentration of these 

particles relative to larger sizes. The specific scattering coefficient in Figure 5 is the spectral 

average weighted by the photopic response from 420 to 650 nm. In this particle size region 

(size much, much less than wavelength), scattering is inversely proportional to the fourth 

power of wavelength, so the scattering by each of these particles for blue light is 

significantly higher than that for the longer wavelengths. That is, this strong dependence on 

wavelength in this region below 0.3 :m causes scattering to be stronger for blue light. 

Scattering by particles of a broad size distribution greater than 0.5 :m is about the same for 

all colors of the visible spectrum. For this blue tint to be apparent, scattering by the larger 

particle size fraction must be small. This same dependence of scattering on particle size and 

wavelength causes light transmitted (that not scattered) to have a brown color; that is, 

31
depleted in blue light. There is a relevant exception to this scattering phenomenon for the 

correct interpretation of brown color that is associated with light absorption by NO2. 

Nitrogen dioxide gas concentrations are sometimes high enough to absorb blue light, 

causing the transmitted light to have a brown color. This absorption has no relation to 

particle scattering; so when blue scattered light is observed, one can be confident of ultrafine 

particles being present in significant concentrations. 

32
4. NOX Reduction Catalysts (SCR) 

Production of SO3, Relationships with Catalysts and Other SCR 

Parameters 

SCR technology has been developed with foreknowledge of the fact that catalytic oxidation 

of SO2 to SO3 accompanies NOX reduction when this technology is employed, at least with 

current vanadium-based SCR catalysts. Thus, the potential for problems related to SO3 

production has been recognized throughout the history of SCR development. Nonetheless, it 

has become the most widely used NOX control system. In Japan and Europe, early 

commercial installations in the 1980s experienced SO2 conversion rates of 1% to 2%, 

whereas current domestic installations typically have conversion rates on the order of 1% or 

less. The higher conversion rates of the early foreign installations were mainly due to the 

lower sulfur contents often present in the foreign fuels as compared with domestic fuels and 

to the relatively early nature of the catalyst design (catalysts have been optimized 

significantly to minimize SO2 conversion since these early installations). Figure 7 gives 

results of an investigation related to SCR with high sulfur coal that was performed as part of 

a DOE Clean Coal Project examining SCR applied to high sulfur coals.45 As seen in the 

upper graph (a), there was a large variation of oxidation rate from catalyst to catalyst 

produced by different manufacturers. The investigators concluded that this resulted from 

different choices by each manufacturer involving tradeoffs with other parameters such as 

catalyst volume, pressure drop, ammonia slip margin, and overall design margin. The 

project specifications required that the SO2 oxidation rate be less than 0.75% at 700 °F for 

all catalysts submitted for the test program. All catalysts satisfied that limit, and it was 

concluded that any of the manufacturers could attain lower values of oxidation rate by 

increasing catalyst volume (and capital cost) as a result of formulation changes. Catalyst 

volume could be held constant while lowering SO2 conversion, but physical changes would 

be required which would effectively increase total catalyst surface area present at the 

expense of pressure drop. These trade-offs occur because a lowering of the SO2 conversion 

rate requires that the catalyst be less active on a specific geometric surface area basis. This 

lower activity must then be offset by increased catalyst volume or by a more dense catalyst 

geometry that would produce higher pressure drops. 

Many operating factors influence SO2 conversion to SO3 in an SCR, including reactor 

temperature, flow rate, and SO2 levels. The SO3 conversion rate of an SCR is normally 

measured by simultaneous sampling from its inlet and outlet ducts using a manual 

procedure, the controlled condensation method described in a later section. The SO3 results 

are expressed in terms of percent SO2 oxidation. In most cases, SO2 conversion is linearly 

33
Figure 7. Oxidation of SO2 to SO3 (a) for several SCR catalysts at 700 and 

750 °F and corresponding NOX reduction (b) at 700 and 620 °F from a pilot 

scale investigation at a coal-fired utility boiler burning 3%-sulfur coal.45 

34 
and inversely proportional to reactor flow rate, and relatively independent of absolute SO2 

concentration. With respect to temperature, however, the change in conversion rate is more 

exponential, making temperature the most critical operating parameter affecting SO2 

conversion. Consequently, conversion may increase dramatically above the normal design 

temperature range for a particular installation. 

Figure 7(a) clearly shows the strong relationship of SO2 conversion to temperature. The 

exact nature of the relationship depends on many factors including catalyst formulation and 

the specific nature of the field application. However, SO2 conversion will always decrease 

with temperature when other factors are held constant. NOX reduction (deNOX) capability 

will generally decrease with temperature as well. This decrease is illustrated in Figure 7(b). 

Note that the figure presents NOX reduction for a single catalyst layer at a constant 

NH3/NOX ratio. A relatively small change in single-layer deNOX level is multiplied across 

layers and, therefore, results in a larger decrease in deNOX capability for the installation as a 

whole. Thus, for most catalysts, the ability to reduce NOX (assuming slip must be held 

constant) is strongly impacted by a reduction in operating temperature. In practice, the 

benefits of lowering the operating temperature for a particular installation (to lower SO2 

conversion, for instance) are offset by increased catalyst volume required to maintain 

desired deNOX capability, assuming the catalyst formulation is not changed. While it may be 

possible to minimize temperature impacts on deNOX capability in order to minimize SO3 

formation, the impacts on total plant performance are most often adverse and must be taken 

into account. Thus, no net positive effect is typically gained by lowering operating 

temperature strictly as a means of SO3 control. Optimally, the typical operating conditions 

for a facility are well known and are considered during the specification of the catalyst and 

the resulting formulation that is used. The maximum SO2 conversion rate is a specified 

parameter, along with deNOX capability, and in practice, most requirements can be met by 

the catalyst suppliers. Tight SO2 conversion specifications will result in increased catalyst 

volume, increased pressure drop across the catalysts, or both. Thus, an overly constraining 

SO2 conversion requirement would result in unnecessary increased facility capital and 

operating costs. The difficulty in specifying the allowable SO2 conversion is exacerbated by 

the poor understanding of the behavior of SO3 downstream of the boiler and upon release at 

the stack. 

The need to apply SCR to a wider variety of fuels with highly variable characteristics has 

dictated that catalyst designers meet ever widening design criteria.44 In addition, end-users 

are requiring more flexibility in SCR operating characteristics to match unit-specific 

circumstances, which contributes to the range of SO2 oxidation rates and resulting SO3 

35
levels being offered by catalyst manufacturers. SCR specialists expect that the capabilities 

of SCR catalysts will continue to expand as additional research and development is fueled 

by increased use of the technology. 

For large coal-fired utility boilers within the United States, the most common SO2 oxidation 

specification is 1%. However, facilities firing high sulfur coal or facilities known to have 

high rates of combustion-related SO2 conversion may opt for lower conversion rates, as low 

as 0.5%. Alternately, units firing very low sulfur coal, such as Powder River Basin coals, 

may opt for higher conversion rates, on the order of 1.5%, or possibly more. Cost savings 

accompany higher conversion rates. The trend in domestic plants has been a lowering of the 

specified SO2 conversion. The actual rate of SO2 conversion will vary over time as 

discussed subsequently. Thus, specifications for SO2 conversion must consider the 

long-term oxidation rates, not just the initial expected conversion. 

Ammonia Slip 

Ammonia slip refers to the small amount of ammonia that passes unreacted through the SCR 

catalyst. Ammonia slip is a function of catalyst formulation, volume, age, and specific 

operating conditions. As such, it is a specified parameter in both design and operation. Even 

a poorly operating SCR system can typically achieve required NOX reduction levels, but at 

the expense of higher ammonia injection rates and increased ammonia slip. 

Most coal-fired facilities specify ammonia slip to be held to less than 2 ppmv over the life of 

the installation and across the normal operating range for the facility while achieving the 

required NOX reduction. Slip levels as much as 5 ppmv in coal-fired facilities are often 

manageable from an operational standpoint, but concerns over fly ash contamination 

typically require that slip be held to lower levels. Usually, the SCR is located upstream of 

the particulate control device. In practice, ammonia slip is often higher than that predicted 

by the design criteria for the facility. This is due to the inability to easily measure and track 

ammonia slip, which, in turn, means that injection rates tend to be set with some margin 

over what is believed to be the minimum needed for the required NOX reduction. From this 

standpoint, actual field measured slip levels are likely to be higher than would be indicated 

by the facility design criteria discussed previously, ammonia in the presence of SO3 was also 

associated with fabric strength problems in baghouse operation. The ammonia levels of 

those investigations were an order of magnitude higher than ammonia limitations associated 

with fly ash contamination related to ash byproduct utilization. More restrictive limitation of 

ammonia slip is necessary in order to permit use of fly ash as a byproduct rather than 

disposal as solid waste. Therefore, current SCR design typically provides for lower slip than 

36
in earlier installations while maintaining the NOX reduction required by the utility system’s 

NOX compliance strategy. Equipment downstream of the air preheater may also be affected 

by ammonia byproduct formation. In this regard, SCR impacts on particulate control 

devices such as ESPs and baghouses are discussed below. 

Variation of SO3 Production and Ammonia Slip Over Catalyst Life 

The levels of SO3 and ammonia slip will vary over time for a particular SCR installation. 

This occurs primarily as a result of catalyst deactivation and catalyst management activities 

discussed below. Ammonia slip will gradually increase as the catalyst deactivates and 

ammonia injection is increased to maintain the necessary NOX reduction. A level of 

ammonia slip is specified at which catalyst addition or a partial catalyst replacement is 

necessary. It is at this point in time that the catalyst “end-of-life” has been reached, although 

a great deal of NOX reduction activity is still present. This end-of-life is consequently a 

specified condition—the point at which ammonia slip reaches some predetermined value. 

Thus, a guaranteed maximum slip value of 2 ppmv requires that the ammonia slip not 

exceed 2 ppmv at the catalyst end-of-life. Consequently, the average ammonia slip for a 

particular catalyst charge over its life will be lower than the guaranteed level, averaging 

roughly 1 ppmv for a unit with a 2 ppmv slip guarantee (assuming a linear deactivation 

rate). Most SCR reactors are constructed with space for one or more spare layers of catalyst. 

The most frequent design is a three-layer system where two layers are installed initially, and 

a third layer is installed at the end-of-life of the initial charge. This produces a step change 

in activity, carrying the reactor forward in time for several additional years of service. 

Thereafter, replacement of catalyst beds will occur, again producing step changes in activity. 

Initial catalyst lives are typically on the order of 16,000 to 24,000 operational hours, 

depending on the specific unit. Catalyst additions or partial replacements will usually extend 

the operating life by 16,000 hours or more. 

SO2 conversion to SO3 does not decline over time as would be indicated by the decline in 

NOX reduction activity for a particular charge of catalyst.46 This is somewhat 

counterintuitive and is due to the nature of the catalyst with respect to active sites as well as 

mass transfer contributions. As a result, a particular charge of catalyst will exhibit a near 

constant SO2 oxidation rate over its entire life in the reactor and at a particular operating 

condition. Only when additional catalyst is added will a step change in SO2 conversion be 

noted. Consequently, assuming that all catalyst formulations are identical, SO2 conversion 

will be a function of total installed catalyst volume only, irrespective of the age of the 

catalyst present. For example, an SCR system having a two-bed initial catalyst charge that 

has a 1% SO2 conversion rate will experience a 50% increase in SO2 conversion with the 

37
addition of the third catalyst bed, resulting in a SO2 conversion rate of 1.5%. This 

conversion rate will then be constant over the remaining life of the installation irrespective 

of catalyst replacements (assuming that all catalyst formulations are identical and the total 

catalyst volume remains constant). Thus, most facilities will experience one step-change in 

SO2 conversion, which occurs at the installation of the spare layer of catalyst, and 

conversion remains constant thereafter, assuming that three layers continue to be utilized. 

This conclusion is based on the expectation that catalyst management will proceed without 

replacing two of the layers and removing a third one. 

Options for Controlling Ammonia and SO3 

Assuming that it is well distributed, the amount of ammonia absorbed on the catalyst is 

proportional to the ammonia injection rate, and this determines the NOX reduction. 

Ammonia slip is also proportional to injection rate. Normal operation of SCR is predicated 

on the requirement that a specified NOX reduction is achieved by increasing ammonia 

injection as needed. In many systems, this is achieved by direct control of ammonia 

injection based on continuous NOX monitors. Consequently, ammonia slip is not controlled 

by limiting injection but can only be controlled through the design of the SCR facility, its 

catalyst, and its optimization to attain a desired level of slip resulting from the condition that 

the required NOX reduction is maintained. Facility design issues such as reactor inlet flue 

gas distributions (ammonia, flow, temperature, NOX) will influence the performance of the 

SCR reactor and the resulting ammonia slip. Catalyst design will of course help to minimize 

slip, with formulation and volume being related to the desired slip levels and SO3 production 

levels. Once the facility is installed, optimization procedures to tune the system will aid in 

the reduction of ammonia slip. 

When the particulate control is at the cold side of the air preheater, SO3 may enhance 

removal of ammonia through the formation of sulfates of ammonia that can be collected by 

an ESP or baghouse. When dealing with coals for which SO3 formation is very low, it may 

be desirable to design the SCR for enough SO2 to SO3 conversion to assure reaction with 

ammonia slip. The presence of reactants other than SO3 and the degree of adsorption of 

ammonia by fly ash has not been addressed. 

There may be a need to limit SO3 depending on alkali, alkaline earth, and sulfur content of 

the coal to avoid corrosion or reduce the visibility of emissions as described above. 

Concerning control of SO3, similar options are available related to the SCR facility itself, 

mainly centering on the catalyst formulation and volume. Since SO2 is evenly distributed, 

temperature is the only distribution parameter that significantly effects SO3 production from 

38
the SCR. Maintaining the design temperature for the SCR is important. 

Options for controlling SO3 elsewhere in the boiler effluent, such as injection of alkali 

downstream of the SCR or air preheater, are available. However, the costs of the auxiliary 

control schemes should be weighed against the cost of controlling SO3 via SCR design. In 

practice, for facilities that have identified a need to control SO3 production from the SCR 

(lower than the 1% rate for most SCRs), catalyst formulation changes and volume increases 

are typically preferable compared to alkali injection. This is primarily due to the 

comparative capital costs of the two options, as well as the long-term operating costs and 

attendant operating difficulties of alkali injection systems. Alkali injection requires an 

additional plant system, transport and storage of alkali material, and manpower to maintain 

the injection system versus the static system associated with the catalyst control option. 

As a comparative example of alkali injection costs versus the costs associated with a 

low-SO3 production catalyst, a 500 MW unit burning 2.6% sulfur coal is referenced (based 

on a study at EPRI’s High Sulfur Test Center).47 For this facility, it was estimated that a 

limestone injection system using purchased limestone on the open market would cost 

approximately $500,000 a year on a 15-year levelized basis. Comparatively, it is estimated 

that for a similar facility operating at 90% NOX reduction levels and using a 25% catalyst 

volume incremental increase, the additional cost for controlling SO3 production from 1% to 

0.5% would be approximately $230,000 a year on a levelized basis (based on a 24,000 hour 

life, catalyst cost of $5,500 per cubic meter, and an initial charge of 500 cubic meters for the 

1% conversion option). Catalyst-based control options are often attractive if the required 

SO3 control is within reach via catalyst modifications. Other factors will, of course, play a 

role in the decision and, consequently, the technical and economic issues are highly 

site-specific. However, to date, most facilities have opted for catalyst based SO3 control 

rather than alkali injection, assuming that proper consideration is given prior to the 

installation of the SCR facility. In situations where problems are encountered after the 

installation of the SCR system, alkali injection for control is one of the few options 

available, since a reduction in the SO3 produced by the catalyst would require a replacement 

of the entire charge of catalyst. Obviously, the desired final levels of absolute SO3 are also 

important—SO3 control using catalyst modification is, of course, very limited in magnitude 

compared to what could be achieved with alkali injection. 

Ammonium Sulfate and Bisulfate Aerosol 

It is well known that ammonia and SO3 produced in SCRs react to form ammonium sulfate 

or bisulfate and some of it is deposited on internal surfaces of the air preheater. This 

39
deposition is reflected by gradual increases in flue gas pressure drop across the air preheater 

with installation of SCR. The increased pressure drop is managed by soot blowing. 

Characteristics of the products of this reaction are of interest because of relevance to air 

preheater operation and to PCD operation downstream of air preheaters. Observations 

concerning ammonia injection upstream of a baghouse are discussed above. Previous 

observations concerning ammonia injection upstream of ESPs found that collection 

efficiency was enhanced in some instances.4 

To explore possible behavior of the ammonia/SO3 system, it is useful to make some 

approximate calculations of reaction regimes and aerosol formation. Under the conditions of 

temperature at which ammonium sulfates are observed to form, SO3 will have reacted with 

water vapor to form sulfuric acid vapor; therefore, the reactants in the cases of greatest 

interest are ammonia and sulfuric acid. Although the thermodynamically favored product of 

the reaction is ammonium sulfate, the initial product is observed to be ammonium bisulfate. 

Ammonium sulfate is formed only after longer periods of time in the presence of excess 

ammonia. The melting temperature of ammonium bisulfate is 146.9 °C (296.4 °F), below the 

temperature at which, as shown below, it typically begins to form in flue gas. Therefore, the 

initial product is liquid. The reaction forming ammonium bisulfate is then: 

NH3(gas) + H2SO4(gas) = NH4HSO4 (liquid) 

An equilibrium constant of 1 for this reaction was calculated from thermodynamic data by 

Burke and Johnson.48 Consider ammonia and SO3 concentrations at the low end of the 

ranges discussed in Section 3 (e.g., 2 ppmv of NH3 and 10 ppmv H2SO4 leaving the SCR). 

At these levels, according to Burke and Johnson,48 ammonium bisulfate formation is not 

expected until the gas has been cooled to 160 °C (320 °F). At the other end of the range of 

possible concentrations, 5 ppmv NH3 and 40 ppmv H2SO4, the onset of ammonium bisulfate 

formation is expected at 176 °C (349 °F). At a typical air preheater outlet temperature, say 

150 °C (302 °F), the equilibrium constant is equivalent to 4.7 ppmv,2

 so control of ammonia 

slip and acid formation at low levels of, for example, 0.5 ppmv NH3 and 9 ppmv H2SO4 

would be required to completely avoid ammonium bisulfate formation inside the air 

preheater. 

In an air preheater channel on the flue gas side, the temperatures of both wall and gas 

decrease in the direction of gas flow. The lowest gas temperatures at a given axial position 

are those adjacent to the wall. Formation and condensation of ammonium bisulfate, 

therefore, occurs first at the wall at the location where the metal temperature reaches the 

40
critical value according to the equilibrium constant. Starting at this axial position, a 

boundary layer forms in the gas at the wall, within which there is a gradient in ammonium 

bisulfate concentration and where the temperature is equal to, or below, the critical value for 

ammonium bisulfate nucleation and condensation. The reaction front where ammonia and 

sulfuric acid form ultrafine droplets of ammonium bisulfate is the surface where this 

boundary layer meets the free stream. The resulting droplets travel toward the wall by 

diffusion and thermophoresis, driven by the concentration and temperature gradients, 

respectively. If the critical temperature for ammonium bisulfate formation is reached far 

enough from the outlet of the air preheater, the ammonia-sulfuric acid reaction front 

eventually propagates to the center of the channel, forming ultrafine droplets over the entire 

cross section of the flow. 

Continuing in the direction of the flow, the following processes occur simultaneously as the 

gas cools: (1) formation of new nuclei and condensation of additional ammonium bisulfate 

on existing droplets, (2) growth of the droplets by coagulation, (3) deposition of droplets on 

the wall of the channel, (4) agglomeration of the droplets with fly ash particles, and (5) 

removal of ammonium bisulfate deposits from the wall by collisions of fly ash particles. As 

a result of these transformations, ammonium bisulfate can leave the air preheater in any of 

the following forms: (1) as ultrafine droplets or particles, (2) on the surface of fly ash, (3) as 

particles produced by erosion of deposits by fly ash, (4) as coarse particles produced during 

soot blowing of the air preheater, and (5) as particles reentrained into the combustion air 

when the flow direction in the air preheater channel reverses. 

Estimates of the distributions of ammonium bisulfate over ultrafine particles, fly ash, and 

the air preheater surface and its contribution to ultrafine particles entering the particulate 

control device were found by a calculation of the behavior of ammonium bisulfate during its 

formation in a cylindrical channel and its evolution during the time it would take to travel 

from the air preheater to the particulate control device. The acid and ammonia levels were 

chosen toward the upper end of the ranges of conditions expected in coal-fired boilers, as 

discussed elsewhere in this report. The most important of the processes described above 

were included in the calculation: 

1. Reaction between sulfuric acid and ammonia to form ammonium bisulfate

condensation nuclei.

2. Growth of the liquid ammonium bisulfate droplets by coagulation. 

3. Agglomeration of ammonium bisulfate droplets with fly ash. 

4. Deposition of ammonium bisulfate on air preheater surfaces by diffusion and

thermophoresis.

41
The assumed conditions were as follows: 

Initial sulfuric acid: 40 ppmv 

Initial ammonia: 5 ppmv 

Temperature gradient in the air preheater: 128 °C/m (70 °F/ft) 

Temperature leaving the air preheater: 150 °C (302 °F) 

Temperature difference, flue gas to metal: 55.6 °C (100 °F) 

Gas velocity: 14.6 m/s (48 ft/s) (constant) 

Diameter of air preheater passages: 10 mm 

Residence time between air preheater and particulate control device: 5 s 

At the initial ammonia and acid concentrations chosen for the calculation, the equilibrium 

constant of Burke and Johnson48 predicts the onset of ammonium bisulfate nucleation at 176 

°C (349 °F) and 0.205 m (8 inches) from the exit of the air preheater. Because of the strong 

interaction between ammonia and sulfuric acid, the rate of formation of ammonium bisulfate 

was estimated to be the frequency of tri-molecular collisions between ammonia, sulfuric 

acid, and any other gas molecule. The size of the critical nucleus was taken equal to the size 

of a single ammonium bisulfate molecule, consistent with the assumption that all 

tri-molecular collisions that include an ammonia and a sulfuric acid molecule result in 

formation of product. Growth of the droplets by coagulation was estimated using the 

expression of Friedlander35 for growth of a self-preserving particle size distribution. This is 

an approximation, since the distribution is not exactly self-preserving in the presence of 

continuous nucleation and condensation as the gas cools. However, the cooling rate and 

corresponding rate of production of ammonium bisulfate are slow. 

The rate of agglomeration of the droplets with fly ash was found from the standard 

description for mass transfer of nanoparticles to larger suspended particles, using the 

diffusion coefficient for the local mean size of the ammonium bisulfate droplets. Because 

the size distribution of the fly ash was broad and bimodal, it was divided into two parts 

having characteristic sizes of 0.09 and 7 :m and external surface areas of 0.1 and 3 square 

meters per cubic meter of flue gas (at 25 °C and 1 atm). Approximately twice as much 

ammonium bisulfate is collected on the smaller fly ash particles, compared to that collected 

on the larger particles. 

The Reynolds number of the flow in the air preheater channels under the conditions chosen 

is 5000. Deposition of ammonium bisulfate on the air preheater surface was treated using 

standard relations for mass transfer in turbulent flow, with the Schmidt number for the 

ultrafine particles determined using the diffusion coefficient for the local mean size of the 

42
particles. 

The assumed temperature gradient in the gas in the air preheater is shown in Figure 8a. The 

calculated distribution of ammonium bisulfate among ultrafine particles, fly ash, and air 

preheater surface is shown in Figure 8b and the corresponding droplet sizes in Figure 8c. All 

are shown as functions of distance along the 10-mm diameter air preheater channel. The 

residence time in the length of channel shown (8 inches) is 0.014 seconds. Under the 

assumption that its rate is limited by the frequency of tri-molecular collisions, the reaction 

forming ammonium bisulfate is very rapid with respect to the cooling rate; therefore, the 

incremental amount of ammonium bisulfate formed in any time or distance interval is given, 

to good approximation, simply by the change in the equilibrium constant associated with the 

drop in temperature. This process determines the rise in concentration of ultrafine droplets 

shown by the dashed curve in Figure 8b. Transport of droplets to the surfaces of fly ash 

particles and the wall of the channel are relatively slow, removing only small fractions of 

the ammonium bisulfate droplets from the gas as shown by the solid and dashed curves near 

the bottom of Figure 8b. At 150 °C (302 °F) (the temperature of flue gas leaving the air 

preheater), 4.9 ppmv of ammonium bisulfate has formed, and 4.9 ppmv of each of the 

reactants has been consumed, leaving only 0.1 ppmv of ammonia and approximately 35 

ppmv of sulfuric acid unreacted. Approximately 2.5% of the ammonium bisulfate formed is 

estimated to have deposited on the fly ash, 4% has deposited in the air preheater, and 93.5% 

remains suspended as ultrafine droplets in the flue gas. Because the amount of condensing 

material and residence time are both small, the calculated mean droplet size leaving the air 

preheater is only 0.0046 :m, as shown in Figure 8c. 

The evolution of this system over 5 s at the same temperature at which it left the air preheater (150 °C, 302 °F) and in the absence of any additional loss of particles to heat transfer 

surface or walls is shown in Figure 9. The processes occurring in the air preheater are also 

shown, but they now appear compressed into a very small fraction of the time scale (the first 

0.014 s) on the left side of the figure. Under the present set of approximations, no additional 

ammonium bisulfate forms during passage from the air preheater to the particulate control 

device because the temperature is fixed. The only processes that occur are growth of the 

ultrafine droplets and particles by coagulation and collection of the ultrafine droplets and 

particles by fly ash. At the entrance to the particulate control device, 4.1 ppmv of 

ammonium bisulfate are estimated to be present as ultrafine droplets or particles suspended 

in the flue gas, 0.6 ppmv are combined with the fly ash, and 0.2 ppmv were left behind in 

the air preheater to be reentrained as large agglomerates during soot blowing. The mean size 

of the ultrafine particles has grown to approximately 0.06 :m by this time. 

43
Figure 8. Distribution of ammonium bisulfate among ultrafine 

particles, fly ash, and air preheater surface and corresponding 

droplet sizes versus distance along the channel wall. 

44 
Figure 9. Evolution of the distribution of ammonium bisulfate 

between ultrafine particles and fly ash and growth of the ultrafine 

particles by coagulation during 5 s at 150 °C (302 °F), the time 

required for passage from the air preheater to the particulate 

control device. The processes occurring in the air preheater, shown 

in Figure 8b and Figure 8c, are compressed into the first 0.014 s at 

the left in this figure. 

So the estimated contribution of ammonium bisulfate to the particle loading approaching the 

pollution control device (ESP) is 4 ppmv. This material would be distributed over a size 

range typically characterized by a geometric standard deviation of 1.32 for a self preserving 

size distribution in the free molecule regime.37 Dismukes4

 found that, with injection of 

approximately 20 ppm ammonia, the number of particles measured with a condensation 

nuclei counter increased by factors of 3 to 4. Simultaneous measurements of ESP electrical 

parameters (voltages and currents) presented clear evidence of the effects of space charge, 

specifically a reduction of current for a given applied voltage. At three of four sites, ESP 

collection efficiency was improved. It was concluded that enhanced operation was 

associated with space charge combined with enhanced cohesion of fly ash in the collection 

45
layer on ESP collection plates. Higher cohesion reduced fly ash reentrainment during 

rapping of the plates. Fly ash resistivity was not affected by ammonia injection although it 

has been postulated that ammonia injection caused increased resistivity in some instances. 

The mechanism by which space charge enhanced ESP collection efficiency was not 

explained in the reports on the three sites mentioned above. In fact, the improvement is 

counter-intuitive as a reduction of ESP current would be expected to lead to reduced ESP 

performance. A mathematical model for ESP operation was used to evaluate the importance 

of this new aerosol mass (associated with SCR) on ESP performance to determine what 

change in performance should be expected from theory. The aerosol mode resulting from the 

above analysis was added to the inlet size distribution data of Figure 2. Then theoretical 

voltage-current (V-I) ESP operating points2,3 were calculated using both the original inlet 

size distribution and the ammonium bisulfate augmented distribution. Model projections of 

ESP collection efficiencies were then calculated using the two corresponding size 

distributions and operating currents and voltages. The results, given in Figure 10, suggest 

there to be no concern at these ammonium bisulfate loadings. The results showed a 

substantial increase in space charge that reduced current at the projected operating voltage 

Figure 10. ESP inlet and outlet particle size distributions of fly ash with 

and without a nucleated ammonium bisulfate mode expected from 

downstream of an air preheater when an SCR is present. (The ESP inlet 

curve without ammonium bisulfate is measured data, and outlet 

distributions were calculated using an ESP model for a typical ESP.) 

46
in the absence of the ammonium bisulate. However, increasing the applied voltage of the 

first ESP field from 36 to 43 kV brought current back to the level present without the 

ammonium bisulfate mode and somewhat improved collection such that predicted in-stack 

opacity decreased from 11 to 9%. The predicted outlet loading contained 0.76 mg/dncm 

more mass in the ultrafine mode, or 4.6×109 particles per cubic meter more with the 

ammonium bisulfate compared to the system without ammonium bisulfate. The calculated 

particle number loadings in this mode with ammonium bisulfate present are a factor of 38 

higher than without ammonium bisulfate, but these particles would coagulate with the 

normal fly ash present so that any concern with this potential effect can be characterized as 

increased mass in this ultrafine mode. 

The ability to increase applied voltage in the ESP analysis above is predicated on two issues. 

The first is associated with the resistivity of the fly ash layer. If the resistivity is high, the 

applied voltage is limited by sparking that originates from break down in the ash layer at 

some voltage differential across the layer. In the instances relevant here, resistivity will 

typically be lowered by installation of an SCR because of the increased net levels of SO3. 

Otherwise, there is no mechanism known by which the resistivity would be increased. 

Therefore, the original (before SCR) or higher currents can exist for the same break down 

voltage differential across the layer. The second issue is the capability of the ESP power 

supply to produce higher applied voltages. Replacement with modern equipment may be 

necessary at a few installations. 

The levels of ammonia slip considered here, for which no degradation of ESP performance 

is indicated, are expected to be representative of most SCR installations. If ammonia and 

SO3 levels present at the air preheater inlet are allowed to be substantially larger, other 

problems may be encountered. Previous field tests have included such conditions in which 

ammonia was injected upstream of an ESP. Buildup of deposits of ammonium bisulfate 

particles on collection plates and corona electrodes that were not removed by normal 

rapping interfered with ESP performance. Similar consideration applies to baghouse 

operation. Previous experience when both ammonia and SO3 concentrations were greater 

than 10 ppmv and when SO3 concentration alone was greater than 10 ppmv indicated 

problems with degradation of fabric strength. Again, significant degradation associated with 

ammonia bisulfate is not expected at ammonia levels considered above (a few ppmv). 

However, levels of SO3 greater than 10 ppmv, independent of ammonia, may be difficult to 

avoid with many SCR installations; although, as mentioned before, injection of hydrated 

lime at the baghouse inlet can be effective at preserving fabric strength. 

47
48
5. Sulfuric Acid Aerosol Formation and Collection in Wet 

FGD 

The gas-to-particle conversion processes that dominate transformation of sulfuric acid vapor 

to particulate matter in wet scrubbers are different than the analogous dominant processes 

typical for stack emissions in the absence of a wet scrubber. The rapid quench of the flue gas 

by water spray favors nucleation relative to the dilution cooling of acid vapor occurring as 

the result of mixing stack gas with ambient air. The particle size of nucleated acid in an 

FGD scrubber is governed by coagulation and absorption by acid of water vapor at 

essentially 100% relative humidity. The evolution of this nucleated aerosol can be illustrated 

directly by utilizing the self-preserving form of the size distribution associated with 

coagulation.35 The decline of particle number concentration and the growth of average 

particle size are illustrated in Figure 11 for relative H2SO4/water concentrations of 50:50 by 

mass. These are the relative concentrations for equilibrium at scrubber temperature and 50% 

relative humidity. The time scale considered was derived from the assumption that residence 

time in a scrubber is a few seconds and in the associated stack may be tens of seconds, such 

as corresponding to an 800 ft stack. These results simulate an initial nucleation event 

followed by coagulation only. Subsequent growth by diffusion of water vapor to, and 

absorption by, acid droplets that actually occurs is not taken into account. As such, these 

calculations represent lower limits of droplet diameter in the stack. The actual growth in the 

stack from the saturated water vapor also enhances coagulation somewhat, leading to still 

larger droplet sizes than calculated. After leaving the stack and mixing with ambient air, the 

acid droplets dry and become considerably smaller. The 50:50 relative H2SO4/water 

concentration on which Figure 11 is based corresponds to 20 °C temperature and 30% 

relative humidity. At higher humidities, actual droplets retain more water relative to H2SO4 

than the condition assumed for Figure 11. Figure 12 gives corresponding scattering 

coefficients and average specific scattering coefficients for the coagulating aerosols depicted 

in Figure 11. The calculated increase with time of particle size resulting from coagulation 

causes average specific scattering coefficients (γ*ave) to similarly increase with time and 

H2SO4 concentration. Figure 13 presents data of Figure 12 in terms of γ*ave versus H2SO4 

concentration for fixed coagulation times as appropriate for a given site—specifically, 

residence time in the scrubber and stack. Average specific scattering coefficient is seen to 

increase with H2SO4 concentration , nearly following quadratic functions (R2 = 0.9987). This 

means that the amount of light scattered by H2SO4 in scrubber plumes is calculated to 

increase with emission level faster than a simple linear dependence. This dependence is 

stronger for longer coagulation times. 

49
Figure 11. Growth by coagulation of nucleated sulfuric acid droplets from 

H2SO4 molecules. (This simulation is for a gas temperature of 20 °C, for 

which acid droplets are 35% H2SO4 and 65% H2O by mass.) 

50 
Figure 12. Scattering coefficients and average specific scattering 

coefficients of nucleated sulfuric acid droplets versus time for the various 

H2SO4 concentrations of Figure 11. 

51
Figure 13. Average specific scattering coefficients of nucleated 

sulfuric acid droplets versus H2SO4 concentrations for various times 

from Figure 12. 

Wet scrubbers utilized for SO2 control incorporate little energy (equivalent to about 2 inches 

of H2O pressure drop) into the relative motion of droplets and particles compared to that 

which leads to substantial collection of particles in high energy scrubbers. Energy input 

equivalent to at least 20 in. H2O is necessary to achieve a 50% collection efficiency of 1:m 

particles.49 Energy input of 2 in. H2O can achieve 50% collection of particles about 4 :m 

diameter.49 Data characterizing particulate collection efficiency of low energy scrubbers at 

particle sizes below 1 :m are not available. However, measurements on high energy 

scrubbers for sizes well below the 50% cut point show significant collection by turbulent 

and Brownian diffusion, leveling off at about 30%.49 A recent investigation found acid 

removal efficiency (i.e., acid aerosol removal) at one site to be 77% with relative standard 

deviation of 0.09 initially. Six months later, repeat measurements resulted in 31% removal 

with relative standard deviation of 0.13 with no apparent reason for the variation.50 In view 

of the previously measured data characterizing particle collection efficiency versus size and 

scrubber input energy, the 31% value appears plausible while the higher value does not. The 

higher value is believed to be associated with measurement problems of acid aerosols in 

scrubber effluent. 

The overall behavior of plumes also differs substantially between that exiting a wet scrubber 

at temperatures well below 200 °F and that exiting a stack without a scrubber at 

temperatures on the order of 300 °F. Figure 14 presents results of plume modeling for 

52
conditions typical of wet scrubber exhaust and summer days. Ground level concentrations of 

sulfuric acid are given versus downwind distance assuming 50 ppmv initial stack 

concentration of H2SO4. Mixing class A produces the most rapid vertical mixing and occurs 

only about one afternoon hour on several days of a typical summer month. Mixing classes B 

and C are more common, occurring for late morning/early afternoon hours on most summer 

days. Analogous modeling of plumes originating from flue gas in the stack at 300 °F 

predicts ground concentrations of zero over the same downwind distances because the 

buoyancy of the hot gas causes the plume to penetrate through the top of the mixing layer, 

preventing contact of the plume with the ground. 

Figure 14. Predicted ground level concentrations of particulate H2SO4 for 

illustrative conditions representing a stack effluent exhausting a wet 

scrubber for SO2 control of coal-fired boiler emissions. 

53
54
6. SO3/H2SO4 Measurement and Monitoring 

Method 8, the EPA promulgated method for measuring emissions of sulfuric acid mist,51 has 

a detection limit of about 50 mg/m3; thus, it lacks the sensitivity needed for measurements at 

electric utility plants. The controlled condensation method (CCM) developed by Cheney and 

Homolya52 is generally recognized as the most reliable method for measuring SO3/H2SO4 at 

the levels encountered at power plants.52,53 In the CCM, a sample gas stream is conveyed 

through a heated quartz-lined probe, through a quartz fiber filter, and then through a 

condenser in which the acid vapor is removed from the sample gas stream. The probe and 

filter holder are held at a temperature at or above 288 °C (550 °F), hot enough to evaporate 

concentrated H2SO4 that may be in the liquid or adsorbed phase. The condenser is 

maintained at a temperature above the moisture dew point but well below the sulfuric acid 

dew point so that all of the acid is collected in it. A second filter downstream of the 

condenser ensures that any acid aerosolized in the condenser is retained for analysis with 

that collected in the condenser itself. 

Over the past few years, potential improvements in the CCM method as it is applied to 

coal-fired utility installations have been investigated.6,50 One of the major difficulties in 

SO3/H2SO4 sampling is separation of fly ash and other extraneous particulate matter. In 

principle, this is accomplished with the heated quartz fiber filter with condensed phase 

H2SO4 being evaporated and transferred to the condenser. However, both condensed and 

vapor phase H2SO4 can react with collected particles due to the resulting intimate contact 

while passing through the filter and any deposits on it, resulting in a negative bias in the 

measurement. This artifact was evaluated in laboratory studies by passing known levels of 

H2SO4 vapor through filters prepared with representative fly ash deposits.6

 The importance 

of the reaction was found to vary with alkali content. That is, fly ash from coals normally 

associated with high SO3 levels did not produce a significant bias. 

The primary improvement implemented in recent investigations of the method has been use 

of 10-ft probes to assure that condensed or adsorbed H2SO4 in the sample gas is entirely 

volatilized in the probe before reaching the filter upstream of the condenser.50 The latter 

problem is especially relevant when downstream of a scrubber where the H2SO4 is entirely 

in the condensed phase. This sampling train was utilized in the field results discussed above 

concerning removal efficiency of H2SO4 by wet FGD. As discussed above, there remains 

concern with the sampling method due to the large variation of field measurement of H2SO4 

removal (30% to 80%) and the inability to explain high removal efficiency by mechanisms 

involved in wet FGD scrubbers. 

55
It is generally assumed that H2SO4 not in the vapor state is essentially all comprised of 

particles that are small enough that isokinetic sampling is not required in order to obtain a 

representative sample. Further, it is generally assumed that complete spatial traverses of the 

ducts (or stacks) are not required, although traversing of representative temperature zones, 

especially immediately downstream of air preheaters, is required. Data obtained in field 

measurement programs conducted by SRI over the past few years calls into question the 

assumptions regarding the lack of need for isokinetic sampling and traversing. As part of 

programs for diagnosing plume opacity problems at several coal-fired power plants, SRI has 

developed a technique for determining the size distributions of the stack particulate 

emissions in terms of the contributions of coal fly ash, scrubber generated solids, sulfuric 

acid, and combined water.54 The technique uses a sophisticated set of chemical and 

gravimetric analyses of in situ, size-fractionated samples obtained using cascade impactors 

fitted with inertial precollectors. Although the technique is still experimental and remains 

under development, some results obtained with it suggest that the assumption may not be 

valid that the sulfuric acid in the stack is entirely in the form of particles small enough that 

isokinetic sampling is not necessary. Figure 15 illustrates the data obtained during one such 

test.54 In this case, about 80% of the acid appears to have been associated with particles 

larger than 10 :m. Data obtained during a second test at the same location showed 

comparable concentrations of sulfuric acid at sizes below about 10 :m, but the peak 

between 10 and 100 :m was comparable in size to that near 0.1 :m in Figure 15, meaning it 

showed 50% of the acid was associated with relatively large particles. The CCM as it is 

conventionally applied would tend to grossly under-sample particles larger than 1 :m and 

might not be capable of volatilizing those that were sampled. Thus, the data shown here 

suggest that sulfuric acid concentration results at scrubber outlets obtained to date may have 

been biased low by significant amounts. Further, the variability in the concentrations at sizes 

greater than 10 :m may explain part of the large variability in sulfuric acid collection 

efficiencies measured across scrubbers. 

If the CCM is to be used at such locations, a traversing capability may be needed together 

with the addition of a precollector such as that used with cascade impactors. The use of 

isokinetic sampling in conjunction with the added precollector may significantly improve 

the results. The precollector catch can be analyzed for acid in the same fashion as that used 

in obtaining the data shown in Figure 15, closing the gap in the data without incurring the 

large expense incurred in analyzing the many size fractions obtained with cascade 

impactors. 

56
Figure 15. Concentration vs particle size for fly ash constituents, scrubber solids 

constituents, and sulfuric acid measured in the stack of a coal-fired utility boiler 

whose emissions were controlled by an ESP followed by a wet scrubber.54 

An accurate and reliable sulfuric acid monitor would greatly facilitate control of SO3 at 

coal-fired boilers. Several monitors have been developed over the years, some using 

colorimetric methods (e.g., the Severn Science Ltd. System), spectroscopic techniques (e.g., 

the tunable diode laser system first developed by Laser Analytics), and a conductivity cell 

system developed by the U.S. EPA.55 Physical Sciences Incorporated has announced the 

intention to develop an IR spectroscopic-based analyzer.56 Ametek Process Instruments is 

developing a prototype SO3 monitor based on their commercial Model 4600 SO2 Analyzer.57 

Ametek states that their prototype monitor uses a chemical reaction with the SO3 in a sample 

gas stream, “permitting a simple turbidity measurement proportional to the SO3 

concentration.” Instrumental H2SO4 measurement in coal-fired boilers has been limited 

while the CCM continues to be more widely used. It is believed that, as with the CCM, 

difficulties associated with separation of H2SO4 from fly ash and other entrained particulate 

matter must be solved. 

The Severn Science system has shown promise in laboratory evaluations58,59 and has been 

available commercially for many years. In two field evaluations of the Severn Science unit 

on boilers burning residual fuel oil when sampling at the air preheater inlets, measurement 

error was typically found to be less than 10% compared to controlled condensation methods. 

57
However, evaluation on a coal-fired boiler led to inconsistent results with fluctuations that 

could not be explained. 

58
7. CONTROL TECHNOLOGY ISSUES 

The degree of additional control for fine PM which may be required as a result of SCR 

installation will vary greatly depending upon the fuel, boiler design, SCR system design, 

and the pollution control system configuration. Table 3 represents an attempt to categorize 

the major types of installations that are expected to be encountered in the utility PC boiler 

population. Control technology requirements and options are listed beside each major 

system category. 

Table 3. System Categorizations 

Description of System with SCR Fine PM Control Technology Options 

Low S Western coal with fabric 

filter or SD plus fabric filter 

None required 

Low S Western coal with ESP Probably none required with high alkaline ash—otherwise 

reduce temperature ahead of ESP so H2SO4 condenses to be 

caught in ESP 

Low to medium sulfur Eastern 

coal with ESP or fabric filter 

Probably none if PCD operates <250 °F— otherwise reduce 

temperature ahead of ESP so H2SO4 condenses to be caught in 

PCD, sorbent injection. May require upgrade of ESP if it is an 

older unit with an SCA of ~300 or less. 

Eastern coal with hot-side ESP 

plus wet scrubber 

Furnace sorbent injection, sorbent downstream from SCR, or 

wet ESP after scrubber 

Eastern coal with cold-side ESP 

plus scrubber 

Sorbent injection, otherwise reduce temperature ahead of ESP 

so H2SO4 condenses to be caught in ESP. May require upgrade 

of ESP if it is an older unit with an SCA of ~300 or less. 

Alternatively, install wet ESP after scrubber 

Any coal with spray dryer and 

cold-side ESP 

Probably none but may require upgrade of ESP if it is an older 

unit with an SCA of ~300 or less. 

As indicated previously in this report, ammonium compounds resulting from SCR use are 

likely to form at temperatures above normal ESP operating temperatures. Therefore, if the 

ESP is operating with the high collection efficiency (greater than 99.5%) required for most 

particulate emission regulations, the ammonia compounds should be collected in the ESP 

with greater than 75% efficiency, as suggested by the analysis in Section 3, and are not 

likely to cause plume visibility problems. The ammonia slip, as previously noted, is 

normally limited to 2 ppmv to avoid problems with fly ash sale or disposal. 

The most difficult fine PM emission problems will occur when H2SO4 vapor or aerosol 

concentrations leaving the ESP or the wet scrubber exceed the equivalent of 2 ppmv. These 

59
concentrations have the potential to form aerosols in the plume that will have an adverse 

effect on plume visibility and appearance. Extensive work has been performed on injection 

of alkaline sorbents—especially magnesium compounds—to react with SO3 or H2SO4 vapor 

upstream of the ESP so that the vapor is converted to PM which is collected in the ESP.60 

Although many of these trials have been successful in removing the sulfuric acid vapor, long 

term side effects of the injection process are still an area of concern. The effect of the 

sorbent on the catalyst—if injected upstream of the SCR or in the furnace—is one potential 

problem area. Another potential difficulty is adverse effects on ESP performance if the 

electrical resistivity of the ash is raised above the desired value for optimum electrical 

operation.47 

Wet ESPs downstream from the wet scrubber are receiving increased attention.61 Wet ESPs 

are capable of collecting the fine PM that may result from SCR generated sulfuric acid. 

However, uncertainties remain in defining the efficiency of collection that would be 

required to maintain appropriate emission levels and in identifying a wet ESP design 

configuration that can reliably maintain the fine PM concentration below visible threshold 

limits. 

The potential plume visibility problem discussed herein is, of course, exclusively concerned 

with fine PM. Therefore, it is of interest to summarize a recent data review performed on 

PCD performance in collecting particles below 2.5 and 1.0 micrometers aerodynamic 

diameter.62 Figure 16 contains a bar chart illustrating fine PM concentrations—at both 1.0 

and 2.5 micrometers aerodynamic diameter—at system outlets for several systems treating 

flue gas from utility power boilers. Also shown are outlet fine PM concentrations from two 

pilot plant experimental programs (COHPAC II, ESFF).63,64 This graph was prepared using 

cascade impactor data obtained by sampling with traverses at each control device outlet 

sampling location. The data on the full-scale systems were collected during various field 

measurement projects dating back as far as 1975. 

Figure 16 displays cumulative mass concentrations at 1.0 and 2.5 micrometers aerodynamic 

diameter normalized to milligrams per dry normal cubic meter (mg/dnm3

 - 0% water, 20 °C) 

at 0% oxygen for the various systems (aerodynamic diameters are based on unit density 

aerosols). The first set of bars for the “E&S” system illustrate the quantity of fine PM 

measured by the impactor sampling systems when a sulfuric acid fume had formed in the 

scrubber system downstream from a high efficiency ESP.39 Ammonia injection upstream of 

the precipitator has been used at this site to minimize the quantity of sulfuric acid vapor that 

is available to condense in the scrubber. 

60
Figure 16. Fine PM concentrations on a cumulative mass basis for several 

pollution control systems. The equivalent mass concentrations for 

hypothetical SO3 emissions at 1, 5, 10, 20, and 50 ppmv are shown for 

comparison and are not directly associated with any specific control 

system. 

The second set of bars (E&S&W) are calculated results based on experiments on a 

pilot-scale wet ESP.65 Wet ESPs are designed to operate in saturated gas streams. In this 

hypothetical configuration, a wet ESP is installed downstream of a wet scrubber, and the 

particulate collection efficiency as a function of particle size is assumed to be the same as 

that measured in the pilot scale experiments. The calculated results indicate the cumulative 

mass emissions below 2.5 micrometers are reduced by 83%. This result is based on a small 

wet ESP (specific collection area of approximately 35 ft2

/1000 acfm) intended for use as a 

polishing device behind a dry ESP that does not meet current emission requirements. It is 

reasonable to expect that a wet ESP optimized for operation downstream of a scrubber could 

achieve higher collection efficiencies than those used for this hypothetical example. 

The spray-dryer plus fabric-filter combination in Figure 16 (SD&FF) achieved the lowest 

fine PM concentrations among the full-scale systems tested.5

 The full-scale fabric filter 

61
without a spray-dryer upstream (FF) used bags which had been in service for several years, 

and the fine PM emissions measured at that location may have been increased by leaks 

through the bags or in the baghouse structure.66 Both of these full-scale fabric filter 

installations, however, indicated cumulative mass efficiencies of greater than 99% for 

aerodynamic particle diameters smaller than 1 :m. 

The full-scale precipitator (E1) was a large SCA unit (560 ft2/1000 acfm) collecting ash 

from a low-sulfur western coal that produced an ash with an electrical resistivity that did not 

cause problems with electrical operation. Overall particulate mass emissions at this plant 

were very low—approximately 2.48 mg/dnm3

 at 0% O2 (0.0015 lb/million Btu).67 

The second full-scale ESP illustrated in Figure 16 (E2) is perhaps more typical of older units 

with smaller specific collection areas, in this case 241 ft2

/1000 acfm.67 Note that the PM 1.0 

and 2.5 concentration values are much higher than for E1 and higher than any of the fabric 

filters in this comparison. 

The full-scale fabric filter (FF) data are based on tests conducted at a large plant burning a 

western subbituminous coal. It is typical of the large, low-ratio, reverse-gas-cleaned 

baghouses found on many plants in the western United States. Its fine particle collection 

performance is adequate, falling between the two electrostatic precipitator examples shown 

(E1 and E2). 

Figure 16 also illustrates outlet fine PM concentrations from two pilot-scale fabric filter 

systems: 

1. An experimental electrostatic fabric filter (ESFF) without power (ESFF-OFF) and with 

power applied to the discharge electrodes (ESFF-ON),62 and 

2. A Compact Hybrid Particle Collector (COHPAC II) pilot (CPACII in Figure 16).63 

The COHPAC II pilot unit consisted of a two-field ESP operated at about 188 ft2/1000 acfm 

followed by a pulse-jet baghouse that was operated at an A/C ratio ranging from 10.4 to 11.2 

ft/min. The average tube sheet pressure drop ranged from 4.6 to 6.6 inches of water.63 

All of these pilot-scale fabric filter configurations achieved very low outlet fine PM 

concentrations-less than 1.0 mg/dnm3

 at 0% O2. The ESFF power-on configuration achieved 

the lowest values, and the value of 0.05 mg/dnm3

 for cumulative mass below 1.0 :m was at 

the lower limit of resolution for cascade impactor sampling systems under the conditions of 

these tests. 

62
Note that the bars labeled “Hypothetical H2SO4” represent the mass concentration equivalent 

of H2SO4 concentrations ranging from 1 to 50 ppmv. The graph illustrates that only 5 ppmv 

H2SO4 is equivalent to more mass in the sub-2.5 :m range than was measured at the outlet 

of any of the sources depicted, the greatest being the ESP-wet scrubber combination. It was 

apparent that condensation was occurring across the scrubber during the measurement 

program because negative collection efficiencies were observed below 1 :m aerodynamic 

diameter. 

We can derive some practical implications concerning plume appearance from the data in 

Figure 16. Let us assume that SCR introduced an increase of 10 ppmv sulfuric acid vapor 

which was reduced to 6 ppmv in passing through the air preheater and down to 4.5 ppmv 

through the ESP. This 4.5 ppmv acid is expected to nucleate in the scrubber. As discussed in 

Section 4, we expect penetration of 70% of this aerosol (3 ppmv) to the stack based on the 

low pressure drop of wet FGD scrubbers. Due to the saturated water conditions in the 

scrubber, the nucleated acid droplets will grow and coagulate until they exit the stack where 

drying proceeds slowly to a stable point at which acid content of the droplets is about 50% 

by weight,35 or 1/3 acid and 2/3 water by volume. The optical depth (γL) of the plume 

declines slowly, so a reasonable estimate of the scattering coefficient is derived from using a 

dilution factor of L/L(z) where z is the distance traveled and L(z) is the cross wind depth of 

the plume. If the stable particle size were 0.1, 0.2, or 0.3 :m, the scattering coefficient 

would be 0.006, 0.04, or 0.08 L/z mS1

, respectively. The value for 0.1 :m average particle 

size appears realistic. If the stack diameter were 10 m, the plume opacity would be 6%, a 

readily visible haze but not unusual. The average particle size is unlikely to be less than 0.1 

:m. These estimates of scattering coefficient and opacity should be considered to be linear 

with respect to sulfuric acid concentration in the stack—that is, doubling the SO3 

concentration would double the opacity. 

Additional control above that provided by current practice will likely be needed in many 

instances to handle increased levels from SCR. Further, subtleties related to particle growth 

that cannot be accounted for at present might cause formation on, or condensation to, larger 

particle sizes (0.2 or 0.3 :m) that would result in larger scattering coefficients than those 

predicted here. A control level of 80% is not an unreasonable expectation for total fine PM. 

Pilot scale wet ESP measurements have indicated more than 80% cumulative mass 

collection efficiency for particles below 0.5 :m aerodynamic diameter.5

 It is possible that a 

combination of modest amounts of sorbent injection combined with a small-footprint wet 

ESP could provide a margin of safety for adverse plume appearance events that may result 

from SCR-induced sulfuric acid concentration spikes or fuel changes at plant sites with 

63
ESP-scrubber control systems. 

In conclusion, this review indicates control technologies are available that have the potential 

to largely eliminate adverse plume appearance resulting from application of the SCR 

process. The difficulty faced by utility plant operators is that guidelines are not presently 

available to define the extent of the problem that may occur at a given site. In the absence of 

such guidelines, it is not possible to determine in advance the control strategies that are 

technically and economically feasible for a particular plant site. 

64
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