OAR Box 1870 

Prepped by Ollie Stewart 

Document Number: 

14) ll-A-19 

Docket Number: 

A-90-19 

Printed 11/28/2008 7:05:04 AM Sheet 14 of 67 

P.1 
P.2 

___>_____—___— _ 

SEFA 

United States 

Environmental Protection 

Agency 

Office of Air Quality 

Planning and Standar 

Research Triangle Pai 

A-90-19 

ll-A-19 

Air 

Organic Chemical 

Manufacturing 

Volume 5: Adsorption, 

Condensation, and 

Absorption Devices 

jj) B B BIB B 

SLP I •!. lastEPA AIR DOCKET 

SA -1 IO 

/ 

RADIAN LIBRARY 

DURHAM, N.C. 
P.3 

111 

This report was furnished to the Environmental Protection Agency by IT Enviroscience, 9041 Executive Park Drive, Knoxville, Tennessee 37923, in fulfillment 

of Contract No. 68-02-2577. The contents of this report are reproduced herein 

as received from IT Enviroscience. The opinions, findings, and conclusions 

expressed are those of the authors and not necessarily those of the Environmental Protection Agency. Mention of trade names or commercial products is not 

intended to constitute endorsement or recommendation for use. Copies of this 

report are available, as supplies permit, through the Library Services Office 

(MD-35), U.S. Environmental Protection Agency, Research Triangle Park, North 

Carolina 27711, or from National Technical Information Services, 5285 Port 

Royal Road, Springfield, Virginia 22161. 

D124R 

J 
P.4 

1 

f CONTENTS 

| . Page 

J INTRODUCTION vii 

i 

! 

i Report Page 

t 

i 

1. CONTROL DEVICE EVALUATION CARBON ADSORPTION 1-i 

| 2. CONTROL DEVICE EVALUATION CONDENSATION 2-i 

r 

i 3. CONTROL DEVICE EVALUATION GAS ABSORPTION 3-i 
P.5 

1-i 

REPORT 1 

CONTROL DEVICE EVALUATION 

CARBON ADSORPTION 

H. S. Basdekis 

C. S. Parmele 

IT Enviroscience 

9041 Executive Park Drive 

Knoxville, Tennessee 37923 

Prepared for 

Emission Standards and Engineering Division 

Office of Air Quality Planning and Standards 

ENVIRONMENTAL PROTECTION AGENCY 

Research Triangle Park, North Carolina 

January 1981 

D122A 
1-iii 

CONTENTS OF REPORT 1 

P.6 

Page 

I. INTRODUCTION 

II. CARBON ADSORPTION SYSTEMS AND FACTORS INFLUENCING PERFORMANCE 

AND DESIGN 

A. Fixed-Bed Adsorption 

B. Alternative Carbon Adsorption Systems 

III. CONSIDERATIONS FOR INSTALLING NEW CARBON ADSORPTION EQUIPMENT 

A. New Plants 

B. Existing Plants 

IV. COST AND ENERGY IMPACTS OF CARBON ADSORPTION 

A. Base-Case Adsorber Design Summary 

B. Cost Basis 

C. Annual Costs 

D. Cost and Energy Effectiveness 

V. SUMMARY AND CONCLUSIONS 

VI. REFERENCES 

II II II IIIIIIIIIIIVIVIVIVIVVVI1 

1 

1 

22 

•1 

•1 

•1 

-1 

•1 

•1 

-5 

-5 

-1 

-1 

APPENDIX OF REPORT 1 

ANNUAL COST DATA A-1 
P.7 

1-V 

TABLES OF REPORT 1 

Number Page 

II-l Important Properties for Design of Adsorption System for II-6 

Selected Organics 

IV-l Factors Used for Estimating Total Installed Costs IV-3 

IV-2 Annual Cost Parameters IV-6 

IV-3 Cost Effect of Varying Carbon Life and Steam Price IV-11 

IV-4 Cost Effectiveness of Carbon Adsorption for Removal of VOC IV-12 

IV-5 Energy Effectiveness of Carbon Adsorption for VOC Removal IV-14 

IV-6 Change in Cost Effectiveness and Energy Effectiveness with IV-20 

Increasing Regeneration Steam Usage 
P.8 

1-vii 

FIGURES OF REPORT 1 

Number Page 

I-l Activated-Carbon Adsorption System 1-2 

II-l Nomograph for Determining Carbon Requirement II-4 

II-2 Bed Pressure Drop for BPL Carbon II-9 

II-3 Loading Time for Various Carbon Requirements and Bed Depths at 11-10 

100-fpm Velocity and a Carbon Density of 30 lb/ft3 

II-4 Adsorbed-Phase Profile for Trichloroethene Service Time When 11-13 

Vapor Starts to Penetrate Bed; BPL V Type Carbon 

II-5 Adsorbed-Phase Profiles for Trichloroethane at Various Stages 11-13 

of Regeneration; BPL V Type Carbon 

II-6 Desorption Efficiency Based on Steam Usage 11-14 

II-7 Effluent VOC Concentration Based on Steam Usage 11-16 

II-8 Condenser Area Based on Steam Usage 11-19 

li-9 Vacuum Regeneration System 11-21 

11-10 Fluid-Bed Carbon Adsorption System 11-24 

IV-l Installed Capital Cost of Carbon Adsorption Systems IV-4 

IV-2 Net Annual Cost vs Flow Rate for Carbon Adsorption at a Steam IV-7 

Rate of 0.3 lb of Steam/lb of Carbon 

IV-3 Net Annual Cost vs Flow Rate for Carbon Adsorption at a Steam IV-8 

Rate of 0.6 lb of Steam/lb of Carbon 

IV-4 Net Annual Cost vs Flow Rate for Carbon Adsorption at a Steam IV-9 

Rate of 1 lb of Steam/lb of Carbon 

IV-5 Net Annual Cost vs Flow Rate for Carbon Adsorption at a Steam IV-10 

Rate of 2 lb of Steam/lb of Carbon 

IV-6 Cost Effectiveness vs Feed Rate at 1.393 and 6.96 lb of IV-15 

Carbon/1000 scf 

IV-7 Cost Effectiveness vs Steam Regeneration Rates at 1000 scfm IV-16 

IV-8 Cost Effectiveness vs Steam Regeneration Rates at 20,000 scfm IV-17 

IV-9 Cost Effectiveness vs Steam Regeneration Rates at 100,000 scfm IV-18 
P.9 

II-l 

II. CARBON ADSORPTION SYSTEMS AND FACTORS INFLUENCING 

PERFORMANCE AND DESIGN 

In this section the main elements of a carbon adsorption system are discussed 

and factors influencing system design and performance are analyzed. The analysis 

is directed toward development of a design for a typical or base-case adsorption 

system. It is this base-case system and variations of it that form the basis 

for the cost estimation given in Sect. IV. 

Fixed-bed adsorption with steam regeneration is the only system examined in 

depth here. This system includes carbon beds, blowers, condenser, condensate 

decanter, piping, valves, and instrumentation. For some carbon systems, 

however, additional equipment may be required for pretreatment of the gas 

stream or for additional treatment of the VOC recovered during regeneration. 

These considerations are briefly discussed but design and cost details are not 

presented. Fluid-bed adsorption systems are also mentioned but not analyzed. 

FIXED-BED ADSORPTION 

Pretreatment 

A VOC emission stream may have to be treated before it can be sent to the 

carbon adsorber. Stream conditions or contaminants that adversely affect 

carbon adsorber performance are excessively high temperatures, high humidity, 

entrained solids, entrained liquids, and high-boiling organics. The effect 

and subsequent additional cost of these conditions or contaminants on a 

particular carbon adsorption system must be dealt with for each individual 

application. The various types of inlet-stream conditions or contaminants 

that might be encountered are discussed and the steps that might be required to 

deal with the situation are outlined. 

Temperature The adsorption capacity of the carbon and the effluent concentration 

of the adsorber are directly related to the temperature of the inlet stream to 

the adsorber. Normally the temperature of the inlet stream should be below 

100°F or else the adsorption capacity will be affected. Inlet-stream coolers 

are normally required when the waste-gas temperatures are in excess of 100°F. 

Water-cooled finned-tube types of heat exchangers are usually provided to cool 
r II-2 

P.10 

f 

L 

the inlet stream. Costs associated with cooling the inlet stream depend on the 

inlet temperature, flow rate, and temperature of the cooling water available. 

b. Humidity Relative humidities greater than 50% can have a significant effect 

on the operating capacity of carbon. The effect of relative humidity on the 

system should be dealt with for each individual application, but systems in 

which the relative humidity approaches 100% will require additional equipment 

to reduce the water content. The equipment required for humidity control is 

usually a series of coolers (to remove water) and reheaters. However, a single 

heater could be used if the temperature of the inlet stream is significantly 

low, and the increase in temperature will not greatly affect adsorption efficiency. 

Another alternative to decrease the relative humidity of a stream is to add 

dilution air to the system. This method will work if the dilution-air humidity 

is significantly less than the inlet stream humidity. Adding dilution air will 

increase the size and thus the cost of the adsorber required. The installed 

cost of humidity control by a series of heaters and coolers or a single heater 

will be $l/cfm to $2/cfm for a 5000-scfm carbon adsorption unit. 

c. Entrained Solids Entrained solids such as airborne dust, lint, and other 

general particulates may cause the carbon bed to plug over a period of time. 

The particulates that could affect adsorption performance are usually controlled 

by a cloth or fiberglass filter. Depending on the application, either a simple 

throw-away filter or a fabric filter can be used. A throw-away type of filter 

costs in the range of $100/1000 cfm. Information on fabric filters can be 

found in ref 2, and cost data can be found in ref 3. Electrostatic precipitators 

are not usually reocmmended for pretreatment of carbon adsorption inlet streams, 

because there is a possible safety problem of electric arc in the duct with 

organic vapors in the explosive range. Normally the inlet stream concentration 

to the adsorption unit is less than 25% of the LEL but upsets could cause the 

concentration to rise into the LEL range. 

d. Entrained Liquids Entrained liquids can cause operational problems for carbon 

adsorption. If the entrained liquid is water, it should be removed for the 

same reason discussed for relatively humidity. If the entrained liquid is a 

volatile organic, it will also consume adsorption capacity needlessly. A 

number of mist eliminators are available, and one should be used if this problem 
P.11 

=yII-3 

exists. For droplets greater than 5 to 10 pm in diameter a cyclone or zigzag 

baffle can be used, which, if properly designed, will result in efficiencies of 

greater than 95%. To control droplets in the 1- to 5-um range a small-mesh screen 

would be required. For droplets less than 1 um a packed filter device would be 

necessary. A packed filter device has a high pressure drop (15 to 20 in. H20) 

compared to the pressure drop (less than 1 in. H20) across the cyclone or 

zigzag baffle or the screen. For further information on mist eliminators refer 

to ref 4. The cost of a mesh-type mist eliminator is in the range of $150/1000 cfm. 

High Boilers A high percentage of high boilers in the off-gas feed stream 

will be adsorbed on the carbon and will not be removed during steam regeneration. 

This constant buildup of high boilers remaining on the carbon will greatly 

reduce the operating capacity and will require frequent replacement of the 

carbon. Also, plasticizers or resins should be prevented from entering the 

carbon bed since they may react chemically on the carbon to form a solid that 

cannot be removed from the bed during steaming. To prevent frequent replacement 

of the carbon a condenser may be necessary to reduce the high-boiler concentration. 

If the high boilers are in the form of entrained liquids, such as oil droplets, 

a mist eliminator will reduce the buildup. If the high boilers cannot be 

removed from the inlet stream to the adsorber, a carbon guard bed in the adsorber 

may result in less carbon being replaced periodically. The guard bed would be 

considered as a sacrificial device and would be replaced more often than the 

entire bed. 

Adsorption 

Defining the VOC content and the operating capacity as explained below will 

reduce the number of variables required to cover the necessary range of flows, 

concentrations, and VOC molecular weights and will simplify the presentation of 

cost and environmental impacts. 

VOC Content In this study the VOC content of the gas to be treated is defined 

in the units of lb of VOC/1000 scf. The conversion from ppm (volume) can be 

made by using the first part of Fig. II-l or by the following relationship: 
P.12 

10 II- 4 r 

o o o 

u 

•5 

_ 

z 

ui 2 

_i 

s 

10,000 - .01 

OPERATING CAPACITYt 

— lb of VOC 

100 lb of carbon 

001 .01 

i . i . • i •• ! 1 1— i i i "•'- I 

> 

_ 

a. 

i t 1000 

z 

o 

p 

< c 

Ui 

o 

z 

o 

o 

u 

o 

> 

100 

10 

0.1 1.0 

VOC CONTENT (lb of VOC/1000 »cf) 

10 

Molecular weight 

Fig. II-l. Nomograph for Determining Carbon Requirement 
P.13 

II-5 

„ ,,„_„ 4. ppm X VOC molecular weight (lb/lb-mole) 

lb of VOC/1000 scf = ** v * 1 

359 X 1000 

Operating Capacity The determination of the actual operating capacity (also 

commonly referred to as working capacity) of carbon is considered to be beyond 

the scope of this study, since it is affected by a multitude of factors. These 

factors include the properties of the carbon, the properties of the feed stream, 

such as humidity and multiple components, and the percent of containment that 

must be removed. Consideration of the theoretical aspects can be found in the 

literature.5—

12 For the purpose of this study it is assumed that for a particular 

VOC feed stream the operating capacity can be determined from operating industrial 

units, carbon manufacturers, carbon equipment vendors, or pilot-plant studies. 

Pilot tests should be continued through enough cycles to determine the residual 

constant VOC. It is important that the operating capacity used includes all 

the factors that occur in actual practice and for each specific case that the 

operating capacity is considered to be less than the saturated adsorption 

capacity that may be reported. (Some reported operating capacities are shown 

in Table II-l.) 

Carbon Requirement Carbon requirement in this study is defined as lb of 

carbon/1000 scf and is a function of the VOC content of the gas to be treated 

and the operating capacity of the carbon: 

Carbon requirement ">

 o f carbo n = VOC content (lb of VOC/1000 scf? X 100 

1000 scf operating capacity (lb of VOC/100 lb carbon) 

Figure II-l is a nomograph for determining the carbon requirement for systems 

with a variety of VOC concentrations, molecular weights, and operating capacities. 

Systems with high VOC operating capacities per lb of carbon and high concentrations 

of VOC in the process waste gas can have the same carbon requirement as systems 

with low VOC operating capacities per lb of carbon and low concentrations of 

VOC in the waste gas. 

In this study costs were determined for nine different carbon requirements, 

ranging from 0.1 to 10.0 lb of carbon/1000 scf. 
II-6 

P.14 

F 

Table II-l. Important Properties for Design of 

Adsorption System for Selected Organics 

Compound 

Benzene 

Butyl acetate 

n-Butyl alcohol 

Cydohexane 

Ethyl acetate 

Ethyl alcohol 

Heptane 

Hexane 

Isopropyl acetate 

Methyl acetate 

Methyl alcohol 

Toluene 

Trichlorotrifluoroethane 

Xylene 

Boiling 

Point 

(°F) 

176 

260 

210 

178 

171 

173 

208 

154 

199 

135 

149 

231 

118 

291 

Liquid 

Molar 

Volumea 

(cm3/mole) 

95 

152 

105 

118 

106 

61 

163 

140 

129 

83 

42 

118 

120 

140 

Lower 

Explosion 

Limita 

(ppm) 

14,000 

14,000 

17,000 

13,000 

25,000 

33,000 

12,000 

13,000 

18,000 

31,000 

60,000 

14,000 

10,000 

Reported , 

Operating Capacity 

(lb/100 lb of carbon) 

6 

8 

8 

6 

8 

8 

6 

6 

8 

7 

7 

7 

8 

10 

See ref 7. 

See ref 12. Capacities are for relatively dry compounds (or simple mixtures of compounds) 

that are particulate-free at 100°F. The effluent concentration achievable with these 

capacities is not reported. 

____ _ 
P.15 

II-7 

Plow Pa1~" (of Ga s T o B e Treated )

 I n tns study base-case costs were determined 

for six different flow rates, ranging from 300 to 100,000 scfm. 

•Y 

»n) 

Temperature The feed-gas temperature is assumed to be 100°F for the base 

case. Carbon adsorption units operate most effectively in the temperature 

range of 60 to 130°F. Most current carbon adsorption units are operated at 75 

to 100°F, with a few units being operated at 60°F and a few as high as 200°F. 

Increase or decrease of temperature within the range of 90 to 110°F will have 

little effect on adsorption. 

f Gas Velocity (Superficial) Linear velocities of 50 to 100 fpm are normally 

used in bed designs. At higher velocities the bed pressure drop becomes too 

high for standard blowers. At lower velocities the bed becomes too large and 

expensive. If inlet concentrations are low (as is expected in the SOCMI industry), 

the bed area required for the volume of carbon needed usually permits a velocity 

at the high end of this range. Therefore a superficial velocity of 100 fpm is 

used in the base case. 

nds) 

g. Bed Depth Fixed-bed carbon adsorption systems normally have bed depths of 1.5 

to 3 ft, although under certain conditions the depth may be as small as a few 

inches. The minimum bed depth of 1.5 ft is assigned for the adsorption zone. 

The maximum bed depth of 3 ft is considered to be reasonable with regard to the 

pressure developed by standard blowers. 

The difference in capital cost between a 1.5-ft-deep bed and a 3-ft-deep bed 

arises from the cost of the additional carbon and the larger-pressure-drop 

blower. The carbon bed containers are usually large enough for the extra 

1.5 ft to be added to the bed at minimal cost, and therefore there is little 

capital cost difference between the 1.5-ft bed and the 3.0-ft bed. Since the 

incremental cost of adding more carbon to a large vessel is more than offset by 

the added flexibility that a longer adsorption cycle gives, a bed depth of 3 ft 

is used in the base case. 

Pressure Drop The pressure of the feed gas will affect the design of the 

system in terms of the power requirements of the blower. A pressure drop of 

2.5 to 10 in. H20/ft of carbon bed is normally experienced with a flow velocity 
II-8 

of 50 to 100 fpm through the carbon bed. If the feed-gas pressure is high 

enough to permit a reasonable superficial velocity, then no blower is required. 

The base-case system is assumed to require a blower to develop the necessary 

pressure. 

The base-case adsorber uses 4 X 10 mesh BPL carbon, which at the base case 

velocity of 100 fpm has a pressure drop of 6.5 in. H20/ft of carbon bed. With 

the base-case bed depth of 3 ft the total pressure drop is 19.5 in. H20. (Bed 

pressure drop vs superficial velocity for two different BPL carbons is shown in 

Fig. II-2.) This differential pressure can be developed efficiently by standard 

blowers. 

i. Loading Time (per Bed) Loading time is calculated from bed depth, carbon 

density, gas velocity (superficial), and carbon requirement: 

loading time (hr) = 

bed depth (ft) X density (lb/ft3) X 1000 

velocity (fpm) X 60 (min/hr) X carbon required .(lb/1000 ft3) 

Figure II-3 shows the loading time for various carbon requirements and bed 

depths for a bed with a carbon density of 30 lb/ft3 and a superficial gas 

velocity of 100 fpm. 

j. Bed Area and Pounds of Carbon (per Bed) The bed area (per bed) is calculated 

by dividing the flow rate (scfm) by the superficial gas velocity through the 

bed (for the base-case, 100 fpm). The pounds of carbon (per bed) are most 

conveniently determined by calculating the bed volume and multiplying by the 

bed density. In the base case the bed depth is 3 ft and the bed density is 

30 lb/ft3: 

bed area (ft2) = flow rate (acfm)/gas velocity (£pm). 

carbon (lb) = bed area (ft2) X bed depth (ft) X bed density (lb/ft3). 

k. Minimum Cycle Time The minimum cycle time for the system is normally dependent 

on the time required to regenerate, dry, and cool the bed. Steam regeneration 

rates are typically 0.5 to 1 (lb of steam/min)/ft2 but can be as high as 4 (lb 

r 
P.17 

II- 9 

•o 

« 

_> 

a. 

O 

a a 

tu 

a 

(O 

(A 

Ul 

ac 

o. 

15 20 30 40 SO 

SUPERFICIAL VELOCITY (fpm) 

60 70 80 90 100 

Fig. H-2. Bed Pressure Drop for BPL Carbon 
LOADING TIME (hr) 

X 

I -

CL 

UJ 

Q 

__ 

111 

OQ 

Minimum —< 

Bed Depth 

_U—I I I I I I I , I I I I l I I i I I I I I I I 

.02 0.1 1.0 10 

CARBON REQUIREMENT (lb of carbon/1000 scf) 

2 0 

i 

»— 

o 

Fig . II-3 . Loading Time for Various Carbon Requirements and Bed Depths at 

lOO-fpm Velocity and a Carbon Density of 30 lb/ft ^ 

oo 
P.19 

11-1 1 

of steam/min)/ft2. These rates correspond to regeneration times of 30 to 

90 min at steam regeneration ratios of 0.3 or 1 lb of steam/lb of carbon. 

Cooling and drying the beds can normally be done in 15 min. The minimum possible 

time for regeneration, drying, and cooling used in this study is 1 hr. Since 

the bed is loaded for the same period of time, the total minimum cycle for one 

bed takes 2 hr. Since two beds are usually employed, each carbon bed must have 

enough adsorption capacity to allow sufficient time for the other.bed to be 

regenerated, dried, cooled, and placed back in adsorption service before breakthrough . 

]_. Bed Configuration In the base case for carbon systems handling flows below 

10,000 scfm two vertical tanks are used so that one bed can be regenerated 

while the other bed is adsorbing the VOC from the off-gas stream. For flows 

above 10,000 scfm vertical tanks would have diameters greater than 12 ft, which 

is above transportation limits; therefore horizontal tanks are required. For 

flows above 20,000 scfm three or more tanks are used in order to reduce capital 

cost by reducing the tank size and reducing carbon inventory, when three beds 

are used, one will be regenerating while the other two are adsorbing with 

parallel waste-gas flow. A sequence procedure must be established to prevent 

two tanks from being ready for regeneration at the same time. 

3. Regeneration 

a. 

Steam Requirements Regeneration of a carbon bed saturated with VOC is most 

often accomplished by the use of steam. In regenerating a bed a certain amount 

of steam is required to heat the carbon bed from its operating temperature to 

the regeneration temperature and to provide the heat of desorption. However, 

most of the steam flow is needed to provide a sufficient concentration gradient 

to promote mass transfer of the adsorbate from the carbon bed. Even with 

extensive steam regeneration some VOC will be left on the bed. The adsorbate 

left on the carbon (the heel) after regeneration accounts for most of the 

difference between the saturated adsorption capacity and the operating capacity. 

For a given compound-carbon combination the amount of heel is determined by the 

amount of steam used during the regeneration cycle. 
11-12 

Figures II-4 and II-5 illustrate the effect of regeneration on the concentration 

of the adsorbate left on the carbon. Figure II-4 is the adsorption profile of 

a single compound just before a breakthrough occurs on the bed. This is the 

usual or desired status of the system just before regeneration. Figure II-5 

shows the progression of the desorption profiles with increasingly larger 

amounts of steam passing through the bed. Profiles 1 and 2 are achieved with 

relatively small amounts of steam,- profiles 3, 4, and 5 are the results of 

increasingly larger amounts of steam.8 

Figure II-6 shows the effect that varying the steam regeneration ratio (lb of 

steam per lb of carbon) will have on the desorption efficiency. The S-shaped 

curve in Fig. II-613 indicates that a large amount of steam is required to 

achieve nearly 100% desorption efficiency (minimum heel) but that a relatively 

smaller amount is required to achieve 90% desorption efficiency (10% plus 

minimum heel left). If the designer is not limited by the effluent VOC concentration required, he has a choice of how much heel to leave on the bed. Based 

on Fig. II-6 an adsorption designer might choose between 0.25 and 0.35 lb of 

steam/lb of carbon as the preferred design condition, which will provide the 

greatest amount of operating capacity for the amount of steam used. 

The desorption efficiency curve in Fig. II-6 is a function of the compound(s) 

adsorbed on the carbon, of the operating capacity of the adsorbent, and of the 

temperature and pressure of the regeneration steam. It may also be a weak 

function of steam regeneration rate [(lb of steam/min)/ft2], but there are very 

little data in this area. A strongly adsorbed compound would have a curve 

skewed more to the right, since more steam would be required to remove it. A 

compound with a high loading capacity may require more steam per pound of 

carbon to account for the extra heat needed for the heats of desorption and the 

greater amount of material to be desorbed. Higher temperature regeneration 

steam may decrease the amount of steam required per pound of VOC removed (steam 

regeneration ratio) and may increase the amount of material desorbed. The net 

effect is that the curve will be skewed to the left and the heel remaining on 

the carbon will be less. The steam regeneration ratio may have an effect on 

the shape of the curve but will be most important when determining cycle times. 

P.20 

f 

_L -1*117111 
P.21 

30 

Saturation — 

20 

o <^ 

O OC c 

> I - — 

z — 

Q U i _ 

_ o -

00 z"—* 

o 

o 

1 0 -

11-13 

Direction of 

waste-gas flow 

* 

i i i i _J I 1 L _____ J I U J L 

10 

L (in.) 

15 

Fig. II-4. Adsorbed Phase Profile for Trichloroethylene 

at Service Time when Vapor Starts to Break 

Through the Bed; BPL V Type Carbon 

30 

Saturation 

O 

O 

> 

o 

m UJ 

z 

o I--—* 

< . 

oc c 

Zr— ~" V. 

UJ _) 

o - Z"—' 

o u 

'Relative degree of steam 

1* regeneration 

Steam 

flow 

Fig. il-5 . Adsorbed-Phase Profiles for Trichloroethylene 

at Various Stages of Regeneration; BPL V Type Carbon 
P.22 

11-14 r 

TJ 

O 

__ 

o 

CO 

(0 

Q 

100 

— E 80 -

> 

o 

z 

UJ 

E 

3 

E 

"x 

ra 

5 

u. 

UJ 

O — 

_ • 

—» m)* 

o 

r-

\ 

TJ 

O 

£> 

O 

o> 

O 

Q 

60 -

z 

o 

a 

oc 

o 

CO 

UJ 

Q 

40 -

4) 

3-

E 

3 

o 

20 -

1 2 3 

STEAM USAGE (lb of steam/lb of carbon) 

Fig. II-6 . Methylene Chloride Desorption 

Efficiency Based on Steam Usage 
P.23 

11-15 

Figure II-7 shows a generalized relationship between the amount of steam used 

for regeneration and the outlet VOC concentration during the next adsorption 

cycle. The exact relationship will depend on the type of VOC being removed and 

the operating characteristics of the system. In accordance with state-of-the-art 

carbon adsorption systems for solvent recovery, the outlet concentrations are 

usually in the range of 50 to 100 ppm .

6 This level may correspond to a steam 

regeneration ratio of 0.3 lb of steam/lb of carbon, the ratio normally required 

for solvent recovery adsorption systems. The lower achievable effluent concentration levels are in the range of 10 to 20 ppm ,

14 which would correspond to a 

higher steam ratio. However, for some compounds effluent levels are achievable 

below 1 ppm . Figure II-7 shows that reduced effluent concentration is obtained 

by increasing the steam ratio and that very low effluent concentration levels 

may be obtained with high steam ratios. Figure II-7 is not meant to correspond 

to any one particular compound or adsorption condition but is used to illustrate 

a general trend. It shows that the position of the effluent concentration 

curve for each particular compound is a function of the adsorption temperature, 

regeneration temperature, and carbon operating capacity. The effluent concentration 

curve is relatively independent of inlet VOC concentrations. When the adsorption temperature increases, the effluent outlet concentration curve baseline 

may increase. Higher regeneration temperatures may shift the effluent outlet 

concentration curve downward. A different operating capacity may shift the curve 

laterally, since different amounts of steam may be required to regenerate the 

carbon. Data in this area are not readily available but could be generated in 

pilot-plant studies. 

The achievable effluent quality values assumed for the purpose of evaluating 

adsorption systems in this study are the following: 0.3 lb of steam/lb of 

carbon will achieve an effluent quality of 70 ppm , and 1.0 lb of steam/lb of 

carobn will achieve an effluent quality of 12 ppm . Although these values may 

not be realized for all VOC compounds, they are appropriate for those compounds 

that are normally controlled by carbon adsorption. 

Blower—For the base-case carbon adsorption system a blower is used following 

regeneration to provide drying and cooling air to sweep some VOC from the 

regenerated bed. To achieve the highest possible removal efficiency, regeneration should be started before an unacceptably effluent concentration occurs and 
11-16 

pe-

> 

E 

a 

a 

< 

a 

z 

UJ 

3 

UJ 

f (capacity, type of compound, 

adsorption temperature, and 

regeneration temperature) 

J L J I L 

0 1 

STEAM USAGE (lb of steam/lb of carbon ) 

Fig. II-7. Effluent VOC Concentration Based on Steam Usage. 

(Not for Any Actual Compound) 
P.25 

11-17 

the drying and cooling air should be sent to the adsorbing bed. Routing the 

drying and cooling air through the adsorbing bed requires a larger pressure 

drop for the drying and cooling air blower and main blower, additional valves 

and controls, and a slightly larger carbon bed. When additional air enters the 

adsorbing bed, no additional carbon capacity will be required since the concentration is low; however, a slightly larger cross-sectional area will be required 

because of the increased flow rate. This practice of routing the drying and 

cooling gas may not be practical for all cases since additional water will be 

brought into the adsorbing bed, which may affect the operating capacity of the 

carbon and cause operational problems. 

The flow rate of the drying and cooling air blower is assumed to be about 30% 

of the off-gas flow rate. The pressure drop for the blower will be the total 

of the pressure drop for the regeneration bed and adsorbing bed, since the air 

is routed through both beds. For pressure-drop calculations it is assumed that 

4 X 10 mesh carbon is used, that the bed depth is 3 ft, and that the superficial 

gas velicty is 100 fpm for the total cooling air and gas flow. The drying and 

cooling time is 15 min per cycle. 

c. Condenser—All the steam required to regenerate the carbon has to be condensed 

and cooled to 100°F. Some of the heat will stay with the bed and be removed by 

the drying and cooling air, but most of it has to be removed by the condenser. 

The reduction in heat duty for the condenser is not significant; so the heat 

duty for the condenser is based on the amount of steam required to regenerate 

the bed. For the base-case design it is assumed that the steam will be used 

during 75% of the regeneration time; therefore the condenser will have to be 

sized to handle instantaneous heat duty, which is 33% more than the average 

steam consumption. 

For the base-case design, atmospheric steam at 212°F and 970 Btu/lb of heat of 

vaporization was used. The condenser was designed to condensate the steam and 

VOC, cool the condensate mixture to 100°F, and raise the cooling water from 

85°F to 120°F. 

The overall heat transfer coefficient for the condenser was assumed to be 150 

Btu/(hr)(ft2)(°F) (ref 15). The area of the condenser per pound of carbon as 
r P.26 

JL 

11-18 

shown in Fig. II-8 was based on various steam requirements and on a log-mean 

(or conservative) At of 42.4°F. 

Although a decanter to separate the organic from the water is included in the 

battery limits of the carbon adsorption system, no additional equipment is 

included to separate water-soluble organics. Some separation option must be 

considered when large quantities of organics enter the condensate since most 

local authorities will not allow such water to be discharged into municipal 

sewage treatment systems. A complex distillation system may be required to 

fractionate the mixture into separate components, or a simple system can be 

designed to merely strip out the organics. 

Other options for separation of the water-soluble compounds from the condensate 

include liquid extraction or liquid-phase carbon adsorption with solvent regeneration. 

d. Alternatives to Steam .Regeneration Hot inert-gas regeneration and vacuum 

regeneration can be used as alternatives to steam regeneration. With the first 

method, after hot gas is used to regenerate the carbon bed, the VOC-laden gas 

can be incinerated. Another method is to recirculate the inert gas through a 

preheater, into the adsorber, and through a condenser. Each pass through the 

bed with the hot inert gas will pick up additional VOC of higher molecular 

weights. Removal from the adsorber and condensation of the VOC will continue 

until an equilibrium is reached between the bed, the hot inert gas, and the 

condenser. Higher gas temperatures or lower condenser temperatures will 

improve the recovery efficiency of the system. 

In the second method, which involves indirect heating and a vacuum applied to 

the carbon bed, the boiling points of the VOC will be lowered and the VOC will 

be more easily removed from the bed. The vacuum regeneration cycle begins with 

the bed being heated with recirculating heated air or inert gas. Auxiliary 

heating coils within the bed are sometimes used. When the bed reaches the 

specified temperature, a vacuum pump then reduces the pressure in the bed to 

about 0.2 in. H20. The high boilers thereby removed are condensed with a 

chilled-water condenser, and the low-boiling VOC are condensed with a Freoncooled condenser (-40°C). After the desorption step is complete, the bed is 
11-19 

C 

o 

- 3 

w 

CO 

O 

o 

_3 

< 

UJ 

CC 

< 

oc 

UJ 

CO 

z 

UJ 

Q 

z 

o 

o 

0 1 2 3 . 

STEAM USAGE (lb of steam/ lb of carbon) 

Fig. II-8. Condenser Area Based on Steam Usage 
r 11-20 

restored to atmospheric pressure and is cooled by recirculating air. A flow 

diagram of a vacuum regeneration system is shown in Fig. II-9. 

The advantages of vacuum regeneration with indirect heating are that some 

high-boiling-point VOC can be recovered without the use of superheated steam. 

Also, since the desorbed VOC that are recovered are mostly free of water, 

distillation may not be needed for separation of the water-soluble organics. 

Efficiency and effluent concentration data for vacuum regeneration are limited. 

Commercial vacuum units used for solvent recovery with high solvent concentration 

have reported efficiencies of 90 to 99%, depending on the specific hydrocarbon 

adsorbed and the regeneration conditions used.16 

High boilers in the off-gas feed stream will be adsorbed on the carbon and will 

not be removed during steam regeneration. This constant buildup of high boilers 

in the heel will greatly reduce the operating capacity and require frequent 

replacement of the carbon. Disposable carbon beds preceding the regular carbon 

beds may be required to reduce the high-boiler content of the process off-gas 

and thereby reduce the carbon replacement frequency of the main beds. 

4. Achieving High Carbon Adsorption Efficiencies 

High removal efficiencies require the close control of influent gas quality. 

Performance can be impaired by particulates or a high water content in the inlet 

stream or by high temperature in the inlet stream. "Unexpected" compounds 

(those for which the adsorber was not designed) may pass through the carbon bed 

without being removed or may accumulate in the bed and reduce the operating 

capacity. 

Prevention of premature breakthrough (unacceptable effluent concentration) 

requires a favorable balance between the adsorption capacity dictated by the 

mass flow rate of VOC and the speed and effectiveness with which the adsorber 

can be desorbed, cooled, and dried. 

To avert the discharge of VOC from a hot, wet adsorber, piping must be arranged 

so that the adsorber can be dried and cooled and the drying cooling gas recycled 

to an on-line adsorber. There must be sufficient time in the total adsorption 

cycle for the bed to be dried and cooled. 

I 
11-21 

P.29 

WA*>T6 - gjA.6 1NLE1 

A\P FU-TET? ( 

hTt-Cr\ 

I 

1 

____ 

T 

FAM 

t>TeA.K 

COIL, 

VACUUM 

PUMP 

PSF^\(^ERA-SC 

CONDEM<_£^ . 

«ecEi^/£« 

3S.C0V-.EW LOCP 

Fig. II-9. Vacuum Regeneration System 
P.30 

L 

11-22 f 

5. Safety Considerations 

Spontaneous combustion of fixed beds of carbon can occur whenever the gas 

stream contains oxygen and compounds easily oxidized in the presence of carbon, 

such as ketones, aldehydes, and/or organic acids. Heat generated by adsorption 

or by oxidation of VOC in the bed is usually transported from the bed by convection. If less convection heat is removed than is generated, the bed temperature 

will rise. Higher temperatures will further increase the oxidation decomposition, 

and hot spots exceeding the autoignition temperature of the carbon may develop 

in the bed. The hot spots will develop in a shorter time and the condition 

will be aggravated if a limited amount of steam is accidentally admitted to the 

bed from a leak or from a valve being opened at the wrong time. A number of 

bed fires have occurred after the adsorber was shut down for a long period and 

then not regenerated sufficiently. When the VOC-laden stream was reintroduced, 

bed combustion occurred. Hot spots may also form rapidly under abnormal bed 

conditions, such as uneven bed depth due to carbon movement and stagnant bed 

areas.14 When ketones, aldehydes, or organic acids are involved in the operation of a bed, safety precautions are especially important. 

When the bed is adsorbing ketones, it should not be dried completely after 

regeneration, because the water present provides a heat sink to dissipate the 

heat of adsorption and oxidation. Bed fires are much more likely when the bed 

is permitted to dry. To ensure safe operation with ketones it is necessary to 

have an on-line monitoring of CO and C02 concentrations in the effluent stream. 

At shutdown the bed should be regenerated and cooled before it is left idle. 

The valves should be routinely inspected to ensure against steam leaks. Also, 

the exclusion of oxygen from the bed will help to control hot spots. This 

discussion does not imply that safe operation is impossible,- on the contrary, 

there are several examples of ketones having been successfully recovered via 

carbon adsorption. It does mean, however, that there are potential hazards 

that should be considered.14 

B. ALTERNATIVE CARBON ADSORPTION SYSTEMS 

1. Fluidized-Bed Carbon Adsorption 

Currently a fluidized-bed carbon adsorption process is conmerically available 

from Union Carbide Corporation, who licensed the process from a Japanese company. 
P.31 

11-23 

e 

,on. 

The carbon particles are designed to have a high attrition resistance so that 

they can withstand the abrasive action in the fluidized bed. The carbon particles 

a r e manufactured from shaped molten petroleum pitch. Being thermoplastic in 

nature, the material is formed into microspherical particles by its own surface 

tension. The particles are then carbonized and activated by steam. Because 

they are formed from molten material, they are almost spherical and their 

structure is homogeneous and strong.17 

Figure 11-10 is a basic flow sheet for the fluidized-bed carbon adsorption 

system. The system consists of a multistage countercurrent fluidized-bed 

adsorption section, a pressure-sealing section, a moving-bed desorption system, 

and another sealing section with carbon recycle. The regenerated carbon is 

swept by carrier gas from the bottom to the top of the column. In the adsorption section the carbon is fluidized and moves across perforated plates and 

down the column by a system of overflow weirs. The adsorption section design 

allows the carbon to contact the gas homogeneously, resulting in high mass-transfer 

efficiency. The pressure drop per stage normally ranges from 0.4 to 0.9 in. 

H20, with six to eight stages required, depending on the application. The gas 

velocity through the adsorption section is as high as 200 fpm, which is 2 to 4 

times that required for fixed beds. This high gas velocity allows the crosssectional area of the bed to be smaller, with the pressure drop across the 

entire bed being 2.4 to 7.2 in. H2 0 .1 7 

The regeneration section is a dense-phase gravity-flow bed with indirect heating. 

The VOC are removed from the bed by an inert-gas regeneration fluid being 

passed through the bed or by only direct steaming of the bed. The desorption 

temperature is normally around 250°F but can be raised to 500°F to remove 

high-boiling-point materials. By using a continuous system of regeneration no 

heat is lost due to the bed being cycled; so the heating requirements are 

reported to be less than those for fixed-bed adsorption systems.15 

For the solvent recovery systems in commercial operation the treated gas effluent 

concentration is typically around 50 ppm of VOC. Effluent concentration data 

show that outlet levels of 10 ppm have been achieved for some applications.15 
r 11-24 

ADSORPT\OU 

:secT\osa 

MEDIUM <=C£ 

RE£_.EUERAT\OW 

RE_3Ek.ePAT\OWi, 

FLUvO 

CARRIER $A*=> 

TREATED AIR TO VEUT 

<J=. <qA«b FV-OW ~" " 

cooui ki^ 

WATER 

TRAY 

<3A5 LIFT UKAE 

PREUEAT\K_G TU5E 

R€^£WePAT\OM TUQI 

RECOVERED 

SOLVENT 

Fig . 11-10. Fluidized-Bed Carbon Adsorption System 
P.33 

11-25 

2. 

Activated-Carbon Fiber Systems 

rrently a fixed-bed carbon adsorption system that uses a fibrous activated 

rbon instead of granular activated carbon is commerically available. Toyobo 

York Inc., is involved in marketing the carbon adsorption equipment in the 

ited States. Their technology was developed and is currently being used in 

japan. The fibrous carbon can be supplied in various forms, such as mats, 

corrugated paper, or honeycombs. The advantages of the fibrous carbon, as 

stated by Toyobo, over granular activated carbon are a greater apparent surface 

area, shorter adsorption paths, greater adsorption capacity, and lower pressure 

drops. The reason given for these advantages is that the fibrous carbon has a 

o 

more uniform pore distributed with an average diameter in the range of 20 A . 

The fibrous carbon can be used for such applications as recovery of solvents, 

pollution control, and gas masks. 

JZ\ 

Rotary Carbon Adsorption Systems 

Cargo Caire Engineering Corporation is currently developing a rotary carbon 

adsorption system, which is designed to function as a concentrator of low-concentration high-flow VOC emissions. The system operates by passing a high 

volume of low-concentration VOC stream through the carbon rotors. The carbon 

rotor slowly turns into a regeneration section, where hot gas is passed through 

the carbon rotor to desorb the organic. This lower volume regeneration gas 

then can be sent to a smaller volume incinerator or fixed-bed adsorber. 
P.34 

2-i 

REPORT 2 

CONTROL DEVICE EVALUATION 

CONDENSATION 

D. G. Erikson 

IT Enviroscience 

9041 Executive Park Drive 

Knoxville, Tennessee 

Prepared for 

Emission Standards and Engineering Division 

Office of Air Quality Planning and Standards 

Environmental Protection Agency 

Research Triangle Park, North Carolina 

December 1980 

D103R 

EH 
—«^~ ~ - • - -

W 2-iii 

i 

CONTENTS OF REPORT 2 

I. INTRODUCTION 

A. General 

II. SYSTEM DESCRIPTION 

A. General 

B. Condensation Equipment 

III. FACTORS INFLUENCING PERFORMANCE AND MODEL SYSTEMS 

A. System Efficiencies 

B. Base Case Model Condenser System 

IV. DESIGN CONSIDERATIONS 

A. General 

B. Capital Cost Parameters 

C. Operating Cost Parameters 

V. COST AND ENERGY IMPACTS OF CONDENSER SYSTEMS 

A. Capital Cost Estimates 

B. Annual Costs 

C. Cost Effectiveness and Energy Effectiveness 

VI. SUMMARY AND CONCLUSIONS 

VII. REFERENCES 

APPENDICES OF REPORT 2 

Page 

I-l 

I-l 

II-l 

II-l 

II-l 

III-l 

III-l 

III-4 

IV-l 

IV-l 

IV-l 

IV-2 

V-l 

V-l 

V-6 

V-6 

VI-1 

VII-1 

A. BREAKDOWN OF CAPITAL COSTS FOR CONDENSER SECTION AND REFRIGERATION UNIT 

B. SAMPLE CALCULATIONS 
2-v 

TABLES OF REPORT 2 

P.36 

(jumb er 

jjl-l Parameters for Condenser System Recovery Efficiencies 

V-l Capital Cost Summary for Complete Condenser Systems 

V -2 Annua l Cost Summary for 50 % VOC Removal 

V-3 Annua l Cost Summary for 80 % VOC Removal 

V -4 Annual Cost Summary for 95 % VOC Removal 

V -5 Annua l Cost Parameters 

V -6 Cost-Effectiveness Summary 

V-7 Energy-Effectiveness Summary 

III-3 

V-2 

V-7 

V-8 

V-9 

V-17 

V-18 

V-25 

m 
2-vii 

FIGURES OF REPORT 2 

dumber Page 

V-2 

V-3 

V-4 

V-5 

V-6 

V-7 

V-8 

V-9 

V-10 

V-ll 

V-12 

V-13 

V-14 

V-15 

V-16 

A-1 

A-2 

Schematic Diagram of a Shell and Tube Surface Condenser 

Schematic Diagram of a Contact Condenser 

Basic Surface Condenser System 

Vapor Pressures of Selected Compounds vs Temperature 

Installed Capital Cost vs Flow Rate for Complete Condenser 

System with a VOC Removal Efficiency of 50% 

Installed Capital Cost vs Flow Rate for Complete Condenser 

System with a VOC Removal Efficiency of 80% 

Installed Capital Cost vs Flow Rate for Complete Condenser 

System with a VOC Removal Efficiency of 95% 

Annual Cost vs Flow Rate for Complete Condenser System with 

VOC Recovery Efficiency of 50% and No VOC Recovery Credit 

Annual Cost vs Flow Rate for Complete Condenser System with 

VOC Removal Efficiency of 80% and No VOC Recovery Credit 

Annual Cost vs Flow Rate for Complete Condenser System with 

VOC Removal Efficiency of 95% and No VOC Recovery Credit 

Net Annual Cost vs Flow Rate for Complete Condenser System with 

50% VOC Removal and $0.10/lb Credit for Recovered VOC 

Net Annual Cost vs Flow Rate for Complete Condenser System with 

50% VOC removal and $0.20/lb Credit for Recovered VOC 

Net Annual Cost vs Flow Rate for Complete Condenser System with 

95% VOC Removal and $0.10/lb Credit for Recovered VOC 

Net Annual Cost vs Flow Rate for Complete Condenser System with 

95% VOC Removal and $0.20/lb Credit for Recovered VOC 

Cost Effectiveness vs Flow Rate for Condenser System with 

50% VOC Removal Efficiency and No VOC Recovery Credit 

Cost Effectiveness vs Flow Rate for Condenser System with 50% 

VOC Removal Efficiency and $0.10/lb Credit for Recovered VOC 

Cost Effectiveness vs Flow Rate for Condenser System with 50% 

VOC Removal Efficiency and $0.20/lb Credit for Recovered VOC 

Cost Effectiveness vs Flow Rate for Condenser System with 95% 

VOC Removal Efficiency and No VOC Recovery Credit 

Cost Effectiveness vs Flow Rate for Condenser System with 95% 

Removal Efficiency and $0.10/lb VOC Recovery Credit 

Cost Effectiveness vs Flow Rate for Condenser System with 95% 

Removal Efficiency and $0.20/lb VOC Recovery Credit 

Installed Capital Cost vs Condenser Area for Various Materials 

of Construction for a Complete Condenser Section 

Installed Capital Costs vs Refrigeration Capacity at Various 

Coolant Temperatures for a Complete Refrigeration Section 

II-2 

II-2 

II-4 

III-2 

V-3 

V-4 

V-5 

V-10 

V-ll 

V-12 

V-13 

V-14 

V-15 

V-16 

V-19 

V-20 

V-21 

V-22 

V-23 

V-24 

A-4 

A-5 
11-2 

P.38 

r 

COOLANT 

INLET 

VAPOR 

OUTLET 

VAPOR 

INLET 

COOLANT CONDENSED 

OUTLET VOC 

Fig . II-l . Schematic Diagram of a Shell and Tube Surface Condenser 

VAPOR INLET 

»VAPOR OUTLET 

._r 

w '''/ \ \ -

WtiW/AhtiWtWVt 

WATER 

INLET 

CONDENSATE 

OUTLET 

Fig . II-2 . Schematic Diagram of a Contact Condenser 
3-i 

lb/hr 

REPORT 3 

CONTROL DEVICE EVALUATION 

GAS ABSORPTION 

R. L. Standifer 

IT Enviroscience 

9041 Executive Park Drive 

Knoxville, Tennessee 

Prepared for 

Emission Standards and Engineering Division 

Office of Air Quality Planning and Standards 

ENVIRONMENTAL PROTECTION AGENCY 

Research Triangle Park, North Carolina 

October 1980 

< 

< 

< 

< 

< 

< 

< 

< 

< 

< 

< 

< 

< 

< 

K 

D96A 

1 
w 

3-iii 

CONTENTS OF REPORT 3 

I. INTRODUCTION I-l 

II. SYSTEM DESCRIPTION II-l 

A. Absorption Equipment II-l 

B. Solvent Requirements II-7 

C. System Flowsheets II-8 

III. SYSTEM EFFICIENCIES III-l 

A. General III-l 

B. System with Once-Through Solvent Usage III-6 

C. Systems with Solvent Recycle III-6 

IV. DESIGN CONSIDERATIONS IV-l 

A. General IV-l 

B. Capital Costs Parameters IV-l 

C. Operating Cost (Credit) Parameters IV-3 

V. COST AND ENERGY IMPACTS OF GAS ABSORPTION SYSTEMS V-l 

A. Capital Cost Estimates V-l 

B. Annual Costs V-29 

C. Cost Effectiveness and Energy Effectiveness V-31 

VI. SUMMARY AND CONCLUSIONS VI-1 

APPENDICES OF REPORT 3 

A ADDITIONAL CAPITAL AND COST SUMMARY CASES AND COST-EFFECTIVB TABLES A-1 

B SAMPLE CALCULATIONS B-l 
P.41 

Number 

V-l 

V-2 

V-3 

V-4 

A-1 

A-2 

A-3 

A-4 

A-5 

A-6 

A-7 

A-8 

A-9 

A-10 

A-ll 

A-12 

A-13 

B-l 

B-2 

B-3 

B-4 

B-5 

B-6 

3-v 

TABLES OF REPORT 3 

Capital Cost Summary for Absorber and Stripper Systems 

Annual Cost Parameters 

Cost Effectiveness Summary 

Energy Effectiveness Summary 

Capital Cost Summary for 99% VOC Removal and with Stripping 

Capital Cost Summary for 99.9% Removal and No Stripping 

Capital Cost Summary for 90% VOC Removal and No Stripping 

Capital Cost Summary for 99% VOC Removal and No Stripping 

Annual Cost Summary for 99% VOC Removal and No Stripping 

Annual Cost Summary for 99.9% VOC Removal and No Stripping 

Annual Cost Summary for 90% VOC Removal and No Stripping 

Annual Cost Summary for 99% VOC Removal, with Stripping and 

Steam Ratio of 0.1 

Annual Cost Summary for 99% VOC Removal, with Stripping and 

Steam Ratio of 0.2 

Cost Effectiveness Summary for 99% VOC Removal for Tray Column, 

with Water Discharged without Stripping and No VOC Recovered 

Cost Effectiveness Summary for 99.9% VOC Removal for Tray Column, 

with Water Discharged without Stripping and No VOC Recovered 

Cost Effectiveness Summary for 90% VOC Removal for Absorber, 

with Water Discharged without Stripping and No VOC Recovered 

Cost Effectiveness Summary for 99% VOC Removal and Steam Ratio 

of 0.1 

Flooding Constants, C. 

Constants for Use in Determining Gas Film's Height of 

Transfer Units 

Diffusion Coefficients of Gases and Vapors in Air at 25°C and 

1 atm 

Constants for Use in Determining Liquid Film's Height of 

Transfer Units 

Diffusion Coefficients in Liquids at 20°C 

Pressure-Drop Constants for Tower Packing 

Page 

VVVVAAAAAAAAAAAAAB 

B 

B 

B 

B 

B 

2 

30 

32 

33 

3 

4 

•5 

•6 

•7 

•8 

•9 

•10 

•11 

-12 

•13 

-14 

-15 

-7 

-13 

-15 

-17 

- 1 8 

- 2 1 
f 3-vii 

FIGURES OF REPORT 3 

Number Page 

II-l Schematic Diagram of a Packed Tower II-2 

II-2 Common Tower Packing Materials II-4 

II-3 Schematic Diagram of a Bubble-Cap Tray Tower . II-5 

II-4 Simple Absorption System II-9 

II-5 Absorption System with Stripping Tower (Once-Through Solvent 11-10 

Usage) 

II-6 Absorption System with Stripping Tower (Solvent Recycled to 11-12 

Absorber) 

III-l Typical Vapor-Liquid Equilibrium Curve for an Absorption System III-2 

III-2 Number of Transfer Units in an Absorption Column for Constant III-3 

mGM/LM 

II.I-3 Number of Theoretical Plates in an Absorption Column for III-4 

Constant mG„/L„ 

H M 

V-l Installed Capital Cost vs Flow Rate for Complete Absorption V-3 

System (No Stripper) with a VOC Removal Efficiency of 90.0% 

V-2 Installed Capital Cost vs Flow Rate for Complete Adsorption V-4 

System (No Stripper) with a VOC Removal Efficiency of 99.0% 

V-3 Installed Capital Cost vs Flow Rate for Complete Absorption V-5 

System (No Stripper) with a VOC Removal Efficiency of 99.9% 

V-4 Installed Capital Cost vs Flow Rate for Complete Absorption V-6 

System (No Stripper) with a Solute-Solvent System Having an 

Equilibrium Curve Slope of 2.0 

V-5 Installed Capital Cost vs Flow Rate for Complete Absorption and V-7 

Stripping System with a VOC Removal Efficiency of 99.0% 

V-6 Annual Cost vs Flow Rate for Absorber Only (No Stripper) with V-8 

5.0 wt % VOC in Waste Gas and 99% VOC Removal Efficiency 

V-7 Annual Cost vs Flow Rate for Absorber Only (No Stripper) with V-9 

0.5 wt % VOC in the Waste Gas and 99% VOC Removal Efficiency 

V-8 Annual Cost vs Flow Rate for Absorber Only (No Stripper) with V-10 

0.05 wt % VOC in the Waste Gas and a VOC Removal Efficiency of 

V-9 

V-10 

Annual Cost vs Flow Rate for Absorber Only (No Stripper) with a 

Solute-Solvent System Having an Equilibrium Curve Slope of 2.0 

and 99% VOC Removal Efficiency 

Annual Cost (Excluding BOD Surcharge) vs Flow Rate for Absorber 

Only (No Stripper) with a Solute-Solvent System Having an 

Equilibrium Curve Slope of 2.0 

V-ll 

V-12 
3-ix 

V-ll Annual Cost (Excluding BOD Surcharge) vs Flow Rate for Absorber V-13 

Only (No Stripper) with 99.9% VOC Removal Efficiency 

V-12 Annual Cost (Excluding BOD Surcharge) vs Flow Rate for Absorber V-14 

Only (No Stripper) with 99% VOC Removal Efficiency 

V-13 Annual Cost (Excluding BOD Surcharge) vs Flow Rate for . V-15 

Absorber Only (No Stripper) with 90% VOC Removal Efficiency 

V-14 Annual Cost (Excluding BOD Charge and VOC Recovery Credit) vs V-16 

Flow Rate for Absorber and Stripper with 99.0% VOC Removal 

Efficiency. 

V-15 Annual Cost vs Flow Rate for Absorber and Stripper with 99% VOC V-17 

Removal and $0.10/lb Credit for Recovered VOC. 

V-16 Annual Cost vs Flow Rate for Absorber and Stripper with 99% V-18 

'VOC Removal and Zero Credit for Recovered VOC 

V-17 Cost Effectiveness vs Flow Rate for Absorber and Stripper with V-19 

99% VOC Removal and Zero Credit for Recovered VOC 

V-18 Cost Effectiveness vs Flow Rate for Absorber and Stripper with V-20 

99% VOC Removal Efficiency and $0.10/lb Credit for Recovered VOC 

V-19 Cost Effectiveness vs Flow Rate for Absorber and Stripper with V-21 

99% VOC Removal Efficiency and $0.20/lb Credit for Recovered VOC 

V-20 Cost Effectiveness vs Flow Rate for Absorber Only (No Stripper) V-22 

with 99.9% VOC Removal Efficiency 

V-21 Cost Effectiveness vs Flow Rate for Absorber Only (No Stripper) V-23 

with 99% VOC Removal Efficiency 

V-22 Cost Effectiveness vs Flow Rate for Absorber Only (No Stripper) V-24 

with a Solute-Solvent System Having an Equilibrium Curve 

Slope of 2.0 

V-23 Cost Effectiveness vs Flow Rate for Absorber Only (No Stripper) V-25 

with 90% VOC Removal Efficiency 

B-l Number of Theoretical Plates in an Absorption Column for B-5 

Constant mGM/-M 

B-2 Correlation for Flooding Rate in Randomly Packed Towers (from B-9 

ref 4) 

B-3 Packing Factors for Raschig Rings and Saddles B-10 

B-4 Number of Transfer Units in an Absorption Column for Constant B-12 

m GM/ LM 

B-5 Tray-Spacing Constants to Estimate Bubble-Cap Tray Tower's B-26 

Superficial Vapor Velocity 

B-6 Sample Flowsheet, Material Balance, and Energy Balance B-31 
P.44 

T i-i 

I. INTRODUCTION 

Gas absorption is an operation in which one or more soluble components of a gas 

mixture are separated from the mixture by selective dissolution in a liquid, 

termed the solvent. The reverse operation, termed stripping or desorption, is 

frequently used for recovery of the absorbed components from the solvent. Gas 

absorption is widely used within the chemical industry for the separation and 

purification of gaseous process streams and is treated extensively in basic 

chemical engineering textbooks.1—

4 

Gas absorption as an emission control method is currently most widely used for 

the removal of water-soluble inorganic contaminants (e.g., sulfur dioxide, 

hydrogen sulfide, hydrogen chloride, and ammonia) from air streams, with water 

being the most common solvent or scrubbing fluid used. Water may also be used 

for the absorption of organic compounds that have relatively high water solubilities (e.g., most alcohols, organic acids, aldehydes, ketones, amines, and 

glycols). For organic compounds that have low water solubilities, other solvents 

(usually organic liquids with low vapor pressures) are used. The suitability 

of gas absorption as a VOC emission control method is generally dependent on 

the following factors: (1) availability of a suitable solvent (solvent properties 

are discussed in Sect. II-C), (2) VOC removal efficiency required, (3) recovery 

value or terminal disposal cost of the contained VOC, (4) capacity required for 

handling vapors, and (5) VOC concentration in the inlet vapor (absorption is 

usually considered when the VOC concentration is above 200—300 ppmv). Once it 

is determined that absorption is technically feasible, it can be compared (for 

efficiency and cost effectiveness) with other VOC controls (primarily carbon 

adsorption and thermal oxidation). 

XT. K. Sherwood and R. L. Pigford, Absorption and Extraction, McGraw-Hill, New 

York, 1952. 

2R. E. Treybal, Mass Transfer Operations, 2d ed., McGraw-Hill, New York, 1968. 

3R. H. Perry and C. H. Chilton, Chemical Engineers Handbook, 5th ed., McGraw-Hill, 

New York, 1973. 

4W. L. McCabe and J. C. Smith, Unit Operations of Chemical Engineering, 2nd ed., 

McGraw-Hill, New York, 1967. 
T n-i 

II. SYSTEM DESCRIPTION 

ABSORPTION EQUIPMENT 

General 

Gas absorption equipment is designed to provide thorough contact between the 

gas and the liquid solvent in order to permit interphase diffusion of the 

material. The rate of mass transfer between the two phases is largely dependent 

on the surface area exposed. Other factors governing the absorption rate, such 

as the solubility of the gas in the particular solvent and the degree of chemical 

reaction, are characteristic of the constituents involved and are more or less 

.independent of the equipment used. The types of equipment that are commonly 

used for gas-liquid contact operations include packed towers, plate or tray 

towers, spray chambers, venturi absorbers, and vessels with sparging equipment. 

The use of venturi absorbers, spray chambers, and sparging is generally limited 

to the control of particulate matter and highly soluble gases requiring very 

few transfer units. They are infrequently used for the control of VOC emissions 

in dilute concentration. The following discussion is therefore limited to 

packed and tray towers. ^ 

Packed Towers 

A schematic diagram of a typical packed tower is shown in Fig. II-l.1 The 

packing is designed to expose a large surface area. When the packing surface 

is wetted by the solvent, it presents a large area of liquid film for contact 

by the solute gas. 

Usually the flow through a packed column is countercurrent, with the liquid 

introduced at the top to trickle down through the packing while gas is introduced at the bottom to pass upward through the packing. This results in the 

highest possible efficiency, since, as the solute concentration in the gas 

stream decreases as it rises through the tower, fresher solvent is constantly 

available for contact. Consequently maximum average driving force is obtained 

for the diffusion process throughout the entire column. 

1J. A. Danielson, Air Pollution Engineering Manual, 2d ed., Air Pollution 

Control District, County of Los Angeles. 
l; 

P.46 

g^^m--..—^.—„—_ . _.- .. -

II-2 

GAS OUT 

LIQUIO DISTRIBUTOR 

PACKING 

^^KRESTRAINER 

SHELL 

RANOOM 

PtCUN G 

LIQUIO 

E-OISTRIBUTOR 

PACKING SUPPORT 

GAS IN 

LIQUID OUT 

Fig . II-l . Schematic Diagran 

of a Packed Tower (from re f 1). 
^w 

P.47 

II-3 

The packing used should be able to provide a large surface area, should be 

shaped to give a large void space when packed, should be strong enough to be 

handled and installed without excessive breakage, and should be chemically 

inert and inexpensive. Although materials such as rock, gravel, and coke are 

occasionally used, most packing consists of various manufactured shapes. The 

most common type is Raschig rings, which consist of hollow cylinders having an 

® external diameter equal to their length. Other shapes include Berl saddles, 

® ® ® 

Intalox saddles. Pall rings, Hypak, and spiral-type rings. Although slotted 

rings such as Pall rings or Hypak are generally higher in unit cost than 

Rashchig rings, their superior performance may permit the use of smaller columns 

with correspondingly lower capital costs. Figure II-21 shows several of these 

common shapes. 

3. Plate or Tray Tower 

In contrast to packed towers, where gas and solvent are in continuous contact 

throughout the packed bed, plate towers employ stepwise contact by means of a 

number of trays or plates that are arranged so that the gas is dispersed through 

a layer of liquid on each plate. Each plate is more or less a separate stage, 

and the number of plates required is dependent on the difficulty of the masstransfer operation and the degree of separation required. 

The bubble-cap plate or tray has been the most common type used in the past, and 

most general references deal primarily with it in discussions of plate towers. 

Other types of plates, including perforated trays and valve trays, are currently 

being widely used in new installations because they are less expensive and 

their performance is about equal to bubble-cap tray performance. 

A schematic section of a bubble-cap tray tower is shown in Fig. II-3.1 Each 

plate is equipped with openings (vapor risers) surmounted by bubble caps. The 

gas rises through the tower and passes through the openings in the plate and 

through slots in the periphery of the bubble caps which are submerged in liquid. 

The liquid enters at the top of the tower and then flows across each plate and 

downward from plate to plate through down spouts. The depth of liquid on the 

plate and the liquid flow patterns across the plate are controlled by various 

weir arrangements. 
IX-* 

P.48 

! 'I I 

BERL SADOLE 

RASCHIG RING 

PALL RING 

INTALOX SAODLE 

TELLERETTE 

Fig . II-2 . Common Tower Packing Materials (from re f 1). 
P.49 

r II-5 

GAS OUT 

SHELL — 

TRAY 

DOWNSPOUT -

TRAY 

SUPPORT RING 

TRAY 

ST I FFENER— ' 

VAPOR 

RISER 

FROTH-

•-•.-: < • - ^ 

^-LIQUID IN 

BUBBLE CAP 

L SIOESTREAM 

'WITHDRAWAL 

. INTERMEDIATE 

FEED 

y-yf^-y 

G AS IN 

,- LIOUIO OUT 

Fig. II-3. Schematic Diagram of a Bubble-Cap Tray Tower (from ref 1), 
1 II-6 

Comparision of Packed and Plate Tower 

While devices such as sparged vessels, spray chambers, and venturi absorbers 

have limited application for gas absorption, the choice of equipment is usually 

between a packed tower and a plate tower. Both devices have advantages and 

disadvantages for a given operation, depending on many factors, such as gas and 

liquid flow rates and degree of corrosiveness of the streams. Final selection 

will usually be based on the following comparative information: 

1. 

3. 

4. 

5. 

9. 

10. 

Packed towers are generally less expensive than plate towers when the 

materials of construction must be corrosion resistant. 

Packed towers have a smaller pressure drop than plate towers designed for 

the same throughput and thus are more suitable for vacuum operation. 

Packed towers are preferred for foamy liquids. 

The liquid holdup is usually less in a packed tower. 

Plate towers are preferable when the liquid contains suspended solids 

since they can be more easily cleaned. Packed towers tend to plug more 

readily. 

Plate towers are preferable for large diameters, to minimize channeling 

and reduce weight. Channeling may be corrected in tall, large-diameter 

packed towers by the installation of redistribution trays at given intervals. 

Plate towers are more suitable when the operation involves appreciable 

temperature variation since expansion and contraction due to temperature 

change may crush the packing. 

Plate towers are superior when the heat of solution must be removed at an 

intermediate location in the tower. The liquid downflow can be collected, 

cooled externally, and returned more simply in a plate column. 

Mass-transfer capabilities of packed towers can be predicted with less 

accuracy than plate towers, often requiring larger safety factors with 

correspondingly higher capital costs. 

With other conditions being equal, economic considerations generally favor 

packed towers of diameters of up to 2 ft and plate towers with diameters 

of greater than 4 ft. 
P.51 

II-7 

3- SOLVENT REQUIREMENTS 

1# Treybal2 lists some important aspects that should be considered in selecting 

absorption solvents: 

4. 

5. 

The gas solubility should be relatively high so as to enhance the rate of 

absorption and decrease the quantity of solvent required. Solvents chemically similar to the solute generally provide good solubility. 

The solvent should have relativly low volatility so as to reduce solvent 

loss. (This is particularly important in emission control applications as 

solvent losses may result in additional VOC emissions.) 

If possible, the solvents should be noncorrosive so as to. reduce construction costs of the equipment. 

The solvents should be inexpensive and readily available. 

The solvent should have relatively low viscosity for suitable 

mass-transfer rates and flooding characteristics. 

Ideally, the solvent should be nontoxic, nonflammable, and chemically 

stable and have a low freezing point. 

When applicable, water offers certain distinct advantages over other solvents. 

Because the actual cost of the solvent is very low, water can frequently be 

used on a once-through basis, with the effluent water being discharged to 

wastewater treatment either directly from the absorber or after a subsequent 

stripping operation. The discharge of absorber effluent water without stripping 

is usually practiced only when the absorbed VOC components are present in very 

low concentrations and/or have little or no recovery value. It is not considered 

to be an acceptable practice if significant secondary emissions or degradation 

in the quality of discharged water results. 

When water is stripped before being discharged to wastewater treatment, oncethrough use may result in less stringent stripping requirements than those 

involved with solvent recycle. (When the solvent is recycled, incomplete 

stripping reduces subsequent absorption efficiency.) Once-through use of water 

also usually reduces cooling requirements and the corresponding heat-exchanger 

capital costs. 

JR. E. Treybal, Mass Transfer Operations, McGraw-Hill, New York, 1955. 
P.52 

^i 

II-8 

An additional advantage in using water as a solvent compared to the use of 

organic solvents is that solvent losses do not result in additional VOC 

emissions or organic concentration in wastewater. Although organic solvents 

are selected to minimize these losses (i.e., low vapor pressure), some losses 

will result. 

SYSTEM FLOWSHEETS 

Absorption system requirements for VOC control applications and the corresponding 

equipment configurations may vary significantly, depending on specific applications and control requirements. The flow sheets discussed below illustrate 

several examples. 

Figure II-4 illustrates the simplest configuration, in which the solvent 

(usually water) is used on a once-through basis and is then either discharged 

to a waste-water treatment system or introduced as a process water stream. The 

possibility of using organic solvents on a single-pass basis may exist in those 

few situations where fresh solvent is available in large quantity as a process 

raw material or fuel. For instance, a small vent stream might be scrubbed with 

oil on a single-pass basis before being burned as fuel. With this system the 

only major equipment items required in addition to the absorption tower are, 

possibly, a blower to provide for tower pressure drop and a pump to discharge 

the effluent solvent. 

In the system shown in Fig. II-5 the solvent (usually water) is used on a 

single-pass basis but is stripped before being discharged. Additional major 

equipment items, compared to those shown in Fig. II-4, include a stripping 

tower, an additional pump, and a number of heat exchangers. Heat exchangers 

may include a feed-effluent exchanger, a steam-heated feed preheater, and a 

water-cooled reflux condenser. The effluent liquid from the absorber is usually 

preheated to near its boiling point before it is introduced into the stripping 

tower. The feed-effluent heat exchanger reduces steam requirements in the 

steam preheater and cools the stripped solvent to well below its boiling point 

before it is discharged. As is shown, the heat required for vaporization in 

the stripper may be introduced by steam being directly sparged into the base 

and the condensed steam being discharged with the stripped solvent. 
II- 9 I 

AB-bCRB-rt 

(PACA-C CR 

AR 

-CieCHA^.-E 

.AlP. 

!UL_7 

"S-FFLUS-iT 

| » 6C-V&UT 

_J 

PUMP 

6LCWE A 

Fig . II-4 . Simple Absorption System. 
AlB 

D\«_CHAR_i£ 

AB60W2ER 

C PACKED CiR 

TPA^ COLUMW) 

r <bOL.V_.r_T 

-••-'/, -

AiR lUL-EX. 3 

BLOW_R 

PUHP FEED-EFFLUENT 

HEAT EXC HM.Hl^K 

RICH _OLVEKVT 

_TEAV4 

_L_ 

6T_AM 

PREHEAT-** 

-00_\U 6 

COUD-MSEK 

/ / / 

' / / 

/ / 

fCOUDESKATC 

6TP1PPEO _0_WEWT &J 

PUKP 

_TR\PP_K 

C PACK G-P CR 

TRAY CDLUMM^ 

-TEAM 

SIRIppEp 

-OLVEkiT 

I 

r— 

O 

Fig. II-5 . Absorption System with Stripping Tower (Once-Through Solvent Usage) 
^ 

11-11 

The last flow sheet (Fig. II-6) depicts a typical absorption-stripping system 

that uses an organic-liquid solvent and recycles the stripped solvent to the 

absorber. Additional equipment items required (compared to those in Fig. II-5) 

include a steam-heated reboiler and a water-cooled solvent cooler. Because 

solvent usage and losses must be minimized, both absorber"and stripper operating 

requirements are generally more stringent than when water is used, with correspondingly higher capital and operating costs. For example, in order to attain very 

low VOC concentration in the air discharged from the absorber, the equilibrium 

concentration of residual VOC in the recycled solvent must be reduced to a very 

low level. With once-through usage of water as the solvent the initial equilibrium 

VOC concentration of the solvent is nil, and the stripping efficiency has no 

effect on absorber performance. 

Other absorption-stripping system variations (not shown) that may significantly 

affect capital and operating costs include requirements for low-temperature 

conditions (requiring refrigeration equipment and additional heat exchangers), 

vacuum stripping, and operation at elevated pressures. Corrosive conditions, 

if present, will result in more stringent construction material requirements 

and in correspondingly higher capital costs. 
•''_ • 

.' e^_<._3 c « 

TSA> ( CiLurAU) 

_=CI_!U_ 

NN^T_-< 

=;__.Y;-_ ^__V-MT 

/ / 

V/y 

A',R !M__ " 

__-V;_3 . 

•iGuvEMT 

C-CL---H 

PUM P 

-0 

r-eo-SFF--^-' 

HEAT -iCHAWS- 3. 

6TEA M 

STEA M 

COUC_UfcAT_ 

6TH'.PP_D tDUVCM T 

R\CH ___-/EMT 

CCUCtLte ? 

/ / / / 

•//y 

COOLU-lOr 

A/AT- R 

i — -

STP,iPF_R 

( PAC/.S- c a 

TRA~Y CCL_Mr_) 

PveeaL-^ 

-TEA M 

PUMP 

M 

r— 

Fig. II-6. Absorption System with Stripping Tower (Solvent Recycled to Absorber) 
m - i 

III. SYSTEM EFFICIENCIES 

GENERAL 

The VOC removal efficiencies of absorption systems are limited by the driving 

force available to transfer the VOC from the gas stream to the liquid solvent. 

The available driving force at each point within a countercurrent absorption 

tower is determined by the difference between the actual VOC concentrations in 

the gas stream and solvent at that point and the corresponding equilibrium 

concentrations. The equilibrium relationships are unique for each VOC-solvent 

system, and the equilibrium values at various concentrations are generally 

determined experimentally and displayed as an equilibrium curve. Figure III-l 

shows a typical equilibrium curve, in which x, the concentration of VOC in the 

solvent, expressed as the mole fraction, is plotted versus the corresponding 

equilibrium concentration (mole fraction) in the gas. As is illustrated by 

Fig. III-l, the equilibrium curve typically approximates a straight line 

(constant slope, m) near the origin, where VOC concentrations in the solvent 

and the gas are low, and increasingly diverges from a straight line at increasing concentrations. 

The relationships between attainable VOC removal efficiency, the equilibrium 

curve slope, and the required liquid-gas mole ratio and/or the corresponding 

absorption tower height (number of transfer units in a packed column or 

theoretical trays in a tray column) are shown by Figs. III-2 and III-3 respectively. . 

Figure III-2, which applies to packed columns, is based on the following 

equation:1 

ln 

N 

1 -

OG (mGM/LM) (III-D 

XA. P. Colburn, Trans. Am. Inst. Chem. Enqrs. 35, 211 (1939) (cited in: 

"Absorption and Extraction," T. K. Sherwood and R. L. Pigford, p 133, Chemical 

Engineering Series, McGraw-Hill, NewYork, 1952). 
P.58 

III- 2 

0.10 

u 

•mi 

~* < 

cn ^ 

<3 M l 

w o 

5 -

> n 

0 

So 

fl ^ 

* U-I fa 0 

« tn 

£

_] 

x (Mole Fraction VOC in Solvent) 

Fig. III-l. Typical Vapor-Liquid Equilibrium 

Curve for an Absorption System , 
^F 

111-3 

o 

o 

in 

4J 

•H 

c 

D 

i-l 

_ 

U-I 

01 

c 

<TJ 

u 

E-i 

U 

CU 

3 

Z 

10,000 

Fig . III-2 . Number of Transfe r Units i n an 

Absorption Column for Constant mG_/Lw 
III- 4 

P.60 

Qi 

ui 

<D 

U 

<- H 

mm 

IO 

•rt 

m-

_ 

U 

0 

_ 

A 

E+ 

0 

Fig . III- 3 

10.000 

Number of Theoretical Plates in an Absorption Column 

for Constant mGfvt/_M. 

i/: 
P.61 

III-5 

The terms used in Eq. (III-l) and Fig. III-2 are defined below: 

N = number of overall gas-phase transfer units, 

m = slope of equilibrium curve, 

G = superficial molar mass velocity of gas, lb-moles/(hr)(ft2), 

L = superficial molar mass velocity of liquid, lb-moles/(hr)(ft2), 

yx = mole fraction of VOC in inlet gas, 

y2 = mole fraction of VOC in outlet gas, 

xx = mole fraction of VOC in effluent liquid, 

x2 = mole fraction of VOC in inlet liquid. 

The analagous equation for tray columns,1 which is the basis for Fig. III-3, is 

as follows: 

ln 

N_ = k

 LM A* 2 • "^j VLM 

P ln (L„/mG_) ' (III "

2 ) 

M' M 

in which N is the number of theoretical trays. All other terms in Eq. (III-2) 

and Fig. III-3 are the same as those used in Eq. (III-l) and Fig. III-2. 

The relationships shown in Figs. III-2 and III-3 apply to the condition of 

constant mG /L^ (the equilibrium curve is a straight line). This condition is 

usually approached closely only with dilute solutions, and more complex and 

accurate calculations are usually necessary for actual design requirements. 

However, the relationships shown are useful for order-of-magnitude estimates 

and for illustrating the relationships between variables. 

Equation (III-l) or Fig. III-2 may be used to estimate the number of transfer 

units required in a packed tower. In order to estimate the total packed height 

required, the equivalent height of a transfer unit must also be determined. 

Similarly, in order to estimate the actual number of trays required in a tray 

tower from the number of theoretical trays predicted by Eq. (III-2) or Fig. III-3, 

the tray efficiency must be estimated. Sample calculations that illustrate the 

estimation of transfer-unit height and tray efficiency are given in Appendix B. 

Examples that illustrate the use of Fig. III-3 to estimate VOC removal efficiency 

are given in Sects. III-B and III-C. 
P.62 

III-6 

SYSTEM WITH ONCE-THROUGH SOLVENT USAGE 

As is discussed in Sect. II, for systems in which the solvent is used on a 

once-through basis (Figs. II-4 and II-5), which is frequently the case when the 

solvent is water, the initial VOC concentration of the solvent is usually nil 

and VOC stripping efficiency has no effect on the VOC removal efficiency of the 

absorber. Referring to Figs. III-2 and III-3, with once-through solvent usage 

x2 becomes 0 and the ordinate (yx - mx2)/(y2 " mx2) becomes simply yi/y2. The 

VOC removal efficiency in percent is then expressed as (1 - y2/Yi) X 100. As 

an example, if a plate absorption column containing 16 theoretical plates is 

operated with a gas/liquid mole ratio (G /L^) of 0.35 and if the equilibrium 

curve slope, m, for the VOC solvent system is approximately 2.0 over the operating range, the value of mG_/L„ becomes 0.7 (i.e., 2.0 X 0.35). Then, referrinq 

M M

 3 

to Fig. III-2, the corresponding value of (yj - mx2)/(y2 - mx2) is read as 

approximately 1000 and the corresponding VOC removal efficiency is estimated as 

(1 - 1/1000) X 100, or 99.9%. 

SYSTEMS WITH SOLVENT RECYCLE 

Systems that utilize organic liquids as solvents usually include the stripping 

and recycle of the solvent to the absorber (Fig. II-6). In this case the VOC 

removal efficiency of the absorber is dependent on the solvent stripping efficiency. 

Based on the same criteria used in the example in Sect. III-B (i.e., m = 2.0, 

Gw/Iv. = 0.35, and 16 theoretical plates), on a concentration of 1.0 mole % VOC 

in the inlet gas, and on a solvent stripping efficiency of 98.0%, the value of 

(yx - mx2)/(y2 - mx2) is again determined as 1000 from Fig. III-2; however, in 

this case the VOC removal efficiency is less because the mole fraction of VOC 

in the inlet solvent, x2, is greater than 0. The value of x2 must be determined 

to solve for y2 in the expression 

Yi mxc 

y2 - mx2 

= 1000. (III-3 

The value of x2 may be determined from a VOC material balance across the absorb* 

expressed as i 
P.63 

III-7 

(Yi ' Y2) % = (*_ " x2) ^ (III-4) 

Although y2 also appears in Eq. (III-4), with a relatively high VOC removal 

efficiency the effect of neglecting the VOC in the exhaust air is negligible 

(as shown later in the example) and Eq. (III-4) may be approximated as 

Xi - x. 

Yi 

_M 

LM 

(III-5) 

Substituting the numerical values in Eq. (III-5) 

xx - 0.02 Xi 

0.01 

Solving for xx and x2 

= 0.35, (III-6) 

xx = 0.00357, and x2 = 0.02 xx = 0.0000714. 

Substituting numerical values in Eq. (III-3): 

0.01 - (2)(0.0000714) = 10Q0 < 

y2 - (2)(0.0000714) 

Solving for y2 

y2 = 0.000153. 

The corresponding VOC removal efficiency is determined as 

3) 

E =

 Y^—^ X 100 = °-01 -.°;?°0153 X 100 = 98.5%. 

Yi 0.01 

To verify that neglecting the value of y2 in the material-balance equation [(III-4)] 

has only a slight effect on the calculated values of xx and x2, the calculated 

value of y2 is substituted in Eq. (III-4): 

M 

r" <yi • y-) = «i - »2 r 

0.35 (0.01 - 0.000153) = xx - 0.02 xt , 
P.64 

18

 li! 

III-8 

then xx = 0.00352, and x2 = 0.0000703. 

This example illustrates the effect of the stripping efficiency on the VOC 

removal efficiency. 

i 

• ti 
IV-l 

IV. DESIGN CONSIDERATIONS 

AGENERAL 

The actual design of an optimum absorption or absorption-stripping system for a 

given emission control application is relatively complex. "An optimum design 

requires the selection of a combination of equipment and operating conditions 

that will satisfy emission control requirements at minimum overall cost. Often 

a change that reduces the cost of one element must be balanced against the 

effect it has on other costs. Although detailed design procedures are beyond 

the scope of this report, some of the most significant design considerations 

and the relationships between them are discussed briefly in this section.1 For 

simplicity, these factors are discussed specifically as they apply to absorption-stripping systems using plate or tray towers and designed to control VOC 

emissions at relatively low concentrations in air, with water used as the 

solvent on a once-through basis (flowsheet, Fig. II-5). 

B. CAPITAL COSTS PARAMETERS 

1. Absorber 

The diameter of the absorber is primarily a function of the total gas flow 

rate, and the height is determined primarily by the number of trays required. 

The number of trays is, in turn, determined by the solvent/gas mole ratio used, 

the slope of the equilibrium curve for each specific VOC-solvent system, and 

the VOC removal efficiency required. Thus the combination of the tower height 

and the solvent rate selected must be balanced for optimum operation. The 

relationships between the number of trays, the solvent-gas ratio, the operating 

curve slope, and the VOC removal efficiency are shown by Fig. III-2. 

2. Stripping Tower 

As with the absorber, the height of the stripper is determined primarily by the 

number of trays required. The number of trays is, in turn, determined by the 

relative volatilities of the solvent and the contained VOC (equilibrium curve 

slope), the reflux ratio used, and the degree of separation (or stripping 

efficiency) required. A relationship of variables for the stripper similar to 

1For a detailed study of wet scrubber systems see: S. Calvert e_t al., Wet 

Scrubber System Study. Vol. 1. Scrubber Handbook, PB—213 016, NTIS (July 1972) 
ammmmmmm '

<—J* —m - c nm 

P.66 

IV-2 

that described for the absorber exists (see Fig. III-2); however, because more 

centrated gas mixtures occur, the deviation of the actual equilibrium curve from 

a straight line generally becomes much greater and the assumption of constant 

mG_/L_ introduces greater error. As with the absorber, the combination of M M 

stripping tower height and reflux ratio must be balanced for optimum cost. 

The required tower diameter is primarily a function of the vapor flow rate. 

The vapor flow rate is, in turn, determined by the reflux ratio used and by the 

solvent flow rate from the absorber. Therefore the required diameter becomes 

dependent on the height selected, and, as the liquid rate is dependent on the 

absorber design parameters, the absorber and stripper parameters become interdependent and must be balanced for optimum cost. 

3. Blower 

Blower capital costs are primarily dependent on horsepower requirements. The 

required horsepower is determined by the air flow rate and by the pressure 

increase (AP) that must be developed to overcome the pressure drop through the 

absorber and inlet duct. The absorber pressure drop is roughly proportional to 

the number of trays in the absorber tower, and therefore the effect of changes 

on blower capital costs and energy requirements must be considered in selecting 

the optimum number of trays in the absorber. (The pressure drop is usually 

significantly less in packed towers than in tray towers and may be a significant 

factor in the type of tower selected). 

4. Heat Exchangers 

a. General The capital cost of heat exchangers is primarily a function of the 

heat-transfer surface area required. For a specific application in which the 

heat-transfer coefficients and the required temperatures remain relatively 

constant, the required areas will be approximately proportional to the required 

heat-transfer rates and to the corresponding flow rates of fluids that are 

heated, cooled, vaporized, or condensed. 

b. Feed-Effluent Heat Exchanger The required areas and the corresponding capital 

costs are primarily dependent on the solvent flow rates. 

c. Steam Preheater The capital cost primarily depends on the solvent flow rate. 
IV-3 

J con. 

"om 

Stripper Condenser The capital cost primarily depends on the vapor rate in 

the stripper, which in turn is determined by the solvent flow rate and the 

reflux ratio selected. 

le 

Other Capital Requirements 

Most of the other capital items to be considered, including piping and instrumentation, are either relatively minor, remain relatively constant, or are independent of other design considerations. 

C. OPERATING COST (CREDIT) PARAMETERS 

e 

to 

s 

ng 

ant 

2. 

General 

Certain operating cost components are interdependent with the equipment design 

parameters and must be considered in arriving at an optimum design for an 

absorption-stripping system. Following is a brief discussion of primary operating cost components. 

Solvent Cost 

For the flow sheet shown in Fig. II-5 in which the solvent (water) is used on a 

once-through basis, the cost of using process water must be considered in the 

design of the absorber tower. The combination of the number of trays, or the 

height of the absorber tower, and the corresponding liquid rate required must 

be balanced for optimum cost. 

ed 

:al 

Electrical Power 

Electrical power is used primarily for operation of blowers and pumps. Blower 

power requirements are primarily determined by the quantity of gas passing 

through the absorber and by the absorber-tower pressure drop. Since pressure 

drop is primarily dependent on the type of tower selected (plate or packed) and 

on the number of plates or height of packing required, electrical power usage 

must be considered in arriving at the optimum absorber tower design. 

Power requirements for pump operation are generally relatively small compared 

to requirements for blower operation. The power required for pump operation 

will be roughly proportional to the liquid rate through the system and will 

therefore vary with the liquid/gas ratio selected. 
IV-4 

Process Steam 

Steam is used for heating the stripper feed (feed preheater) and for vaporization in the stripper tower, by direct injection or by means of a reboiler. 

Preheater steam requirements are determined primarily by the liquid flow rate 

to the stripper. 

The quantity of steam for vaporization is determined primarily by the liquid 

flow rate and by the reflux ratio selected in the design of the stripping 

tower, and thus vaporization steam usage is influenced by both absorber and 

stripper design. 

Cooling Water 

The primary use of cooling water is in the stripper condenser. The quantity 

used will depend on the liquid flow rate and on the stripping-tower reflux 

ratio. 

Wastewater Treatment 

Wastewater treatment costs are usually dependent on both the quantity of wastewater discharged and the VOC concentration (usually measured as BOD or COD). For 

a once-through absorber-stripper system in which the solvent (water) is discharged to wastewater treatment, treatment costs will be dependent on the 

liquid flow rate to the absorber and on the VOC removal efficiency obtained by 

the stripper. Treatment costs will therefore be influenced by the design of 

both the absorber and the stripper. 

Product Recovery Credits 

Product recovery credits will be determined by the volume of gas treated and 

the corresponding VOC concentration, by the unit value of the specific organic 

compounds (either as process materials or as fuel), and by the removal efficiencies of the absorber and stripping towers. As removal efficiencies are 

influenced by specific design parameters (e.g., tower height, reflux ratio) 

they must be considered in the design of an optimum system.