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THE UNIVERSITY OF ALBERTA 


RELEASE FORM 


NAME OF AUTHOR BYONG KWON CHO 


TITLE OF THESIS STUDIES OF CATALYTIC PROCESSING 


FOR MODIFIED CLAUS PLANTS 


DEGREE FOR WHICH THESIS WAS PRESENTED Master of Science 


YEAR THIS DEGREE WAS GRANTED LoS 


Permission is hereby granted to the UNIVERSITY 
OF ALBERTA LIBRARY to reproduce single copies of 
this thesis and to lend or sell such copies for 
private, scholarly or scientific purposes only. 

The author reserves other publication rights, 
and neither the thesis nor extensive extracts from 
it may be printed or otherwise reproduced without 


the author's written permission. 


Edmonton, Alberta, Canada. T6G 2G6 


DATED fy 2h 1975 
















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THE UNIVERSITY OF ALBERTA 


STUDIES OF CATALYTIC PROCESSING 


FOR MODIFIED CLAUS PLANTS 


by 
(6 ror KWON CHO 
A THESIS 


SUBMITTED TO THE FACULTY OF GRADUATE STUDIES AND RESEARCH 
IN PARTIAL FULFILMENT OF THE REQUIREMENTS FOR THE DEGREE 


OF MASTER OF SCIENCE 


IN 


CHEMICAL ENGINEERING 


DEPARTMENT OF CHEMICAL ENGINEERING 


EDMONTON, ALBERTA 


Passe Lo 7s 










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THE UNIVERSITY OF ALBERTA 


FACULTY OF GRADUATE STUDIES AND RESEARCH 


The undersigned certify that they have read, and 
recommend to the Faculty of Graduate Studies and Research, 
for acceptance, a thesis entitled STUDIES OF CATALYTIC 
PROCESSING FOR MODIFIED CLAUS PLANTS submitted by 
BYONG KWON CHO in partial fulfilment of the requirements 


for the degree of Master of Science. 





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ABSTRACT 


The purposes of this work were: 

1) to develop and compare equilibrium calculations 
and a reactor modeling procedure for predicting the perfor- 
mance of a Claus process catalytic converter; 

2) to evaluate under laboratory conditions a newly 
developed catalyst for the Claus process; and 

3) to test for the maximum obtainable conversion 
level in a laboratory fixed-bed catalytic converter. 

Before undertaking experimental measurement of 
reaction conversions, preliminary investigations were made 
to improve the accuracy of the gas chromatographic analysis 
and to ensure that no additional reaction other than that 
in the catalyst bed occurred. 

To evaluate the bifunctional activity of a newly 


developed catalyst for two major reactions in a Clause unit; 


2 H.S + SO. =—™ 2 H.O + Sn (1) 


z 4 ty 


Sn (2) 


3 |W 


2.005 0-+ SO, caine mee CO, + 


an experimental study of four different types of catalysts 
was conducted. These catalysts included: pure y-alumina 
(Kaiser S-201), 5.4% Cu-on-alumina, 12.08% Cu-on-alumina and 
16.07% Cu-on-alumina. Kinetic studies using a 1 inch dia- 
meter integral fixed-bed reactor of 316 stainless steel were 


conducted at space velocities ranging from 25,000 to 


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150,000 fe + A total of 83 kinetic runs were completed at 
556 + 2.8 °K, 932 + 9 psia, and nearly constant feed compos- 
ition, 2.9 + 0.12 mole percent of HS (Oe tue) eanG «hwo ee TU NLe 
mole percent of SO., with the balance No: The Cu-on-alumina 
catalyst proved to be a good bifunctional catalyst for Claus 
Operational conditions. Its catalytic activity was depend- 
ent upon the copper content: for reaction (1), a copper 
content of about 5% by weight maximized the catalytic ac- 
tivity; while for reaction (2), the catalytic activity de- 
creased as the copper content increased. The exact reason 
for this effect of copper content on the bifunctional cat- 
alytic activities is not clear at this point. It is postu- 
lated that the catalytic activity change may be attributed 
to sulfided copper for reaction (1) combined with elimin- 
ation of Lewis-acid as well as basic sites on the y-alumina 
surface by copper. 

A total of 10 kinetic runs were carried out at very 
low space velocities, 4 and 100 hr t, to obtain the maximum 
obtainable conversion level of the Claus reaction in a fixed- 
bed integral reactor. The resulting experimental conver- 
Sions were found to be greater than the predicted thermo- 
dynamic equilibrium conversions under the same reaction 
temperature. Some possible explanations for this discrep- 


ancy are examined. 
Adiabatic reaction paths were calculated for the 
front-end burner section of a Claus unit. Their inter- 


section with the equilibrium conversion curves was used 


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to predict the conversion attained. Then the catalytic 
converter performance was computed using a reaction rate 
expression developed by Liu (66) for a pure y-alumina (Alon) 
and a one-dimensional two-phase catalytic reactor model under 
adiabatic conditions. In the Claus catalytic converter, it 
was predicted that external mass transfer resistances were 
insignificant at space velocities above 1000 hr. The 
predicted temperature and conversion profiles along the 
reactor bed revealed that the feed temperature and the space 
velocity should be above 500°K and below 2000 ake respec- 


tively, to ensure significant reaction rates at the start 


of the reactor bed. 


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ACKNOWLEDGEMENTS 


The author expresses sincere appreciation to many 
individuals who guided and assisted him in his research. 
In particular, he wishes to thank his thesis advisor, 
Professor I.G. Dalla Lana, for his encouragement, advice, 
and criticism during the course of this work. 

The author is grateful to Mr. Jerry Moser for his 
advice in testing the gas chromatographic system and for 
his preparation of the new catalyst. The author wishes 
to thank Mr. Don Sutherland of the instrument shop for 
calibration of thermocouples and for assistance in repairs 
to the gas chromatographic system. The author is indebted 
also to the staff of the machine shop in the Department of 
Chemical Engineering. Special acknowledgement is given to 
Mr. Thomas Turner whose assistance was invaluable through- 
out this work, and to Mrs. E. Sherwin for her careful 
typing. 

The author is sincerely grateful for financial sup- 
port provided by the Department of Chemical Engineering 
and the Canadian Natural Gas Processing Association 
Research Fund. 

Finally, the author is deeply indebted to his wife, 
Soo, for her patience, understanding, and support during 


his research. 


vii 
















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CHAPTER I 


CHAPTER II 


TABLE OF CONTENTS 


INTRODUCTION 


1.1 Problems Related to the Claus 
Process 


1.2 Previous Works 


1.3 Topics to be Investigated in This 
Study 


1.3.1 Evaluation of Bifunctional 
Catalysts 


1.3.2 Prediction of Claus Process 
Performance 


LITERATURE SURVEY 
2.1 Claus Process 


2.2 Sulfur Recovery Processes Related 
to Claus Units 


2.3 Simultaneous Reactions in Claus Units 


leh Ft be HS Air Reaction 


eh ae H,S - SO. Reaction 
220 35 -COS.= SO, Reaction 
2.3+4 COBSS H50 Reaction 


2.3.5 Sulfur Species Association- 
Dissociation Reaction 


2.4 Catalysis by Some Transition Elements 
2.5 Performance of a Claus Unit 


2.5.1 Importance of Combustion 
Chamber or Burner Design 


2.5.2 Performance of a Catalytic 
Converter 


2.5.3 Claus Reactor Design 


viii 


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CHAPTER III 


CHAPTER IV 


CHAPTER V 


TABLE OF CONTENTS (continued) 


Page 


PREDICTION OF CLAUS UNIT PERFORMANCE 


3.1 Performance of the Front-end Burner 63 


3.2 Performance of the Claus Catalytic 
Converter 83 


3.3 Results of Reactor Modeling 92 


DESCRIPTION OF EXPERIMENTAL SYSTEM 


4.1 Reactant Feeding System 103 
4.2 Feed-Product Analysis System 106 
4.2.1 Separation of the Component 

in GC Column 106 
4.2.2 Selection-Sampling Valve 
Mode rig be 
4.2.3 Attenuator Setting Eid 
4.3 Reaction System is 
4.3.1 Feed Pre-heater gb 
4.3.2 Reactor ft We 
4.3.3 Sulfur Condenser and Water 
Condenser 116 
3.4 sulrur, Trap 118 


4.4 Process Measuring and Control System 120 


EXPERIMENTAL PROCEDURE AND RESULTS. 


5.1 General Procedure 122 
5.L.1 Startup of the System L2e 
nc We Fae Catalyst Pretreatment and 

Kinetic Run A er 
5.1.3 Shutdown Procedure 127 
5.1.4 Materials 4 OME | 


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CHAPTER VI 


NOMENCLATURE 


TABLE OF CONTENTS (continued) 


5.2 Data Reduction Procedure 


5.3 Experimental Results and 
Discussions 


5.3.1 Preliminary Investigations 


5.3.2 Comparison of Catalyst 
Activities 


5.3.3 Performance Test on the 
Bifunctional Activity 


5.3.4 Maximum Obtainable Conver- 
sion Level 


CONCLUSION AND RECOMMENDATIONS 
6.1 Performance of Equipment 


Hee eerrearcuion of a Craus Unit 
Performance 


6.3 Evaluation of a Bifunctional 
Catalyst 


6.4 Maximum Obtainable Conversion 
in the Claus Reaction 


6.5 Reversible Reaction in the 
Claus Reaction 


BIBLIOGRAPHY 


APPENDIX A 


CALIBRATION OF GAS CHROMATOGRAPHIC 
SYSTEM 


CALIBRATION OF PROCESS MEASURING SYSTEM 


DERIVATION OF EQUATIONS FOR ADIABATIC 
REACTION PATHS IN THE FRONT-END BURNER 


ESTIMATION OF EFFECTIVE DIFFUSIVITY 


SIMPLIFICATION OF MODELING EQUATIONS 


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TABLE OF CONTENTS (continued) 


ASYMPTOTIC SOLUTION FOR EFFECTIVENESS 
FACTOR 


DERIVATION OF COLLOCATION EQUATION 
FOR EFFECTIVENESS FACTOR 


NUMERICAL SOLUTION FOR EFFECTIVENESS 
FACTOR 


SAMPLE CALCULATION OF DATA REDUCTION 


EXPERIMENTAL DATA FILE 


xi 


Page 


260 


266 


274 
287 


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Table 


LIST OF TABLES 


Heat of Dissociation of Sulfur Species 


Equilibrium Constants for Reactions 
(U.drandr({ 14.2) 


dadX/dT in the Front-end Burner 


Comparison between Different Column 
Arrangements 


Purities of Gases 
Standard Catalyst Properties 


Comparison of Reproducibility 


Calibration of Chromatograph Attenuator 
Calibration of Attenuator #1 
Calibration of Attenuator #2 
Calibration for SO,-N, Mixture 


GC Calibration Using Attenuation 
Scheme II 


GC Calibration Using Attenuation 
Scheme III 


DP-Cell Calibration (Feed Pressure 
at 30 psia) 


DP-Cell Calibration (Feed Pressure 
at 25 psia) 


DP-Cell Calibration (Feed Pressure 
at 20 psia) 


Feed Absolute Pressure Transducer 
Calibration 


Reactor Gauge Pressure Transducer 
Calibration #1 (at 480 Degree kK) 


xii 


78 


109 
128 
129 


141 


L935 
196 
L97 


200 


204 


207 


216 


P< 35 J 


218 


220 


222 


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“i 


LIST OF TABLES (continued) 


Reactor Gauge Pressure Transducer 
Calibration #1 (at 500 Degree k) 


Reactor Gauge Pressure Transducer 
Calibration #1 ( at 560 Degree kK) 


Reactor Gauge Pressure Transducer 
Calibration #2 (at 560 Degree kK) 


Reactor Gauge Pressure Transducer 
Calibration #3 (at 560 Degree k) 


Thermocouple Calibration for Reactor 
Inlet Temperature 


Thermocouple Calibration for Reactor 
Outlet Temperature 


Calibration of Water Feed Pump 
Density Correction of Feed Water 
Raw Data 


Processed Data 


XiLb 


Page 


222 


223 


224 


225 


227 


228 
229 
230 
320 


324 


sges 
EN 
ess 
BES 
ass 
Vos 
8S 


ofS 
O<e 




















(Denntgineg) Satgar gar 3O Peis 


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(2 oonped 002 36) S peveveritis ‘i 


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(x Casey 032 35) a foltsxdi lsd = 


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esiusssisgme? Joint ey 


ToIDSSF 10h noisszdils elquosomiont traf 
. la aaa aa taltco0 > 


qa beet xosaW 20 noissidifed SL. 7 
totew post 3a noigoers09 item Me elt | 

{7 aa woh Le 
“ped boaeeoost St 


Figure 


10 


a i 


iz 


as 


14 


15 


16 


LIST OF FIGURES 


Chemical Equilibrium between Sulfur 
Species at One Atmosphere 


Fraction of the Equivalent Moles of 
S, Associated to S82 of S 

2 6 8 
Effect of the Inert Content on Equili- 
brium Conversion of the Claus Reaction 


Equilibrium Conversion and Adiabatic 
Reaction Path in the Front-end Burner 


Illustration of a Reaction Path in the 
Waste-heat Boiler and the Catalytic 
Converter 


Equilibrium Conversion of co. in the 
Front-end Burner 


Front-end Reaction Path 

Effectiveness Factor as a Function of 
Thiele Modulus for Claus Reaction on 
Alon Catalyst 


Temperature and Conversion Profiles along 
the Bed Depth 


Effect of External Transport Resistance 


Check on the Existence of Multiple Solutions 
due to the External Transport Resistances 


Effect of the Feed Temperature on X-T Plot 


Effect of the Feed Temperature on Tempera- 
ture Profiles 


Effect of the Feed Temperature on Conversion 
Profiles 


Effect of the Space Velocity on Conversion 
Profiles 


Schematic Diagram of the Reactant Feeding 
System 


xiv 


66 


72 


73 


7D 


79 


80 


89 


93 


95 


96 
98 


99 


100 


101 


104 















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Figure 
17 


18 


Lg 
20 
21 


22 


oo 


24 
25 
26 
27 
28 
29 


30 


A.l 
A.2 
A.3 


A.4 
A.5 


LIST OF FIGURES (continued) 


Schematic Diagram of the Analysis System 


Typical Chromatogram for Separation of 
No, CO4, HS, COS, SO and HO 


ay 2 
Schematic Diagram of Reaction System 
Sulfur Trap Performance Test 


First Calibration of Gas Chromatograph 


Performance Comparison between Different 
Integrating Systems 


Second Calibration of Gas Chromatograph 


Simultaneous Conversions Using Different 
Catalysts 


Effect of Copper Content on the Conversion 


Level 


Comparison of Individual Reaction Rate 
on 5.4% Cu-on-alumina Catalyst. 


Effect of Water Content in the Feed on 
Hydrolysis of COS on y-alumina 


Test of Bifunctional Characteristics of 
Cu-on-alumina Catalyst 


Conversion Level as a Function of Space 
Time 


Comparison between Predicted and Experi- 


mental Conversions as a Function of Reactor 


Outlet Temperature 


GC Calibration Apparatus 
Attenuation Scheme I 
Attenuation Scheme ITI 


Attenuation Scheme III 


Homogeneous Reacton Effects in the Mixing 


Chamber During the Calibration Period 


XV 


110 
114 
134 


ite 


140 


142 


146 


148 


156 


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161 


164 


165 


191 
194 
194 


194 


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* 





CHAPTER I 


INTRODUCTION 
Problems Related to the Claus Process: 


During past decades, extensive studies were 
conducted to effectively recover the sulfur from 
natural gas using a variety of chemical processes. 
Recently these sulfur recovery processes have become 
more and more important due not only to the economic 
value of the product sulfur but also to more strict 
air pollution control policies. 

A major source of sulfur in Canada has been 
from hydrogen sulfide in natural gas, particularly in 
Alberta, and in acid-gas streams in petroleum refin- 
eries. Among many different kinds of sulfur recovery 
processes, the modified Claus process has been almost 
exclusively applied to the natural gas plant to con- 
vert hydrogen sulfide in the sour gas stream to high 
grade pure sulfur. 


The major reactions involved in the Claus plant 








are: 
Hoot 0 > H.O + SO (2 2) 

2 oe 2 2 ; 
2H.S + SO > 2H.O + 3 S eee 

2 2 <— 2 ats ae : 


WOT TOUDOATUI 
: : ia. tale 
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= bes . 
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| fon imer Oo yYtetvav 6 patey asp Aside . iy 
af 
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simorove eft oF yino ton spb tnstdioams ston bas srTom 
: Ae 
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‘ ~ = —" yee y 
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4 é - 
" ~ - ra > Sos 3 : ~ nd 
need esi shsasD ai t#ifya to somuoe LOcsm " ee 
<a “ 
at vixsivorsxsq .26e [exten ait sbitive nopoxbys mow: 


-“—1LiSsit mus 


YISVoODSy 


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“te NS 
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tt 
4 


at beviovit stolsseox 4otém ots 


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2 


which make the overall reaction of 


3 3 
3H,S + 5 0, > 3H,0 + “sy Sy Cir) 





Reaction (1.1) shows free flame combustion of 
hydrogen sulfide by air in the front-end burner to 
provide sulfur dioxide which acts as a reactant for 
reaction (1.2) in the catalytic converter. Because 
of high temperatures in the burner, reaction (1.2) 
also proceeds to the extent of roughly 50 - 70% con- 
version of HS via homogeneous kinetics. 

Though the major Claus reaction is reaction 
(1.2) other side reactions usually take place together 
with the major one. 

The most significant and important side reactim 


from the kinetic point of view is 


3 
Met OSirt SO ghee | A200 gt ek (1.4) 


Reaction (1.4) may occur due to the carbonyl 
sulfide impurity which is introduced by the reaction 
between hydrogen sulfide and carbon dioxide in the 
burner, or from COS present in the acid gas (33). 

The Claus reaction is much more complicated 
than it seems to be in its simplistic reaction formula. 
This reaction was found to be catalyzed by various 
Substances such as the glass surface, iron, water 


vapor, and even liquid sulfur, and so on (97,72,53). 


1% 


a” : x 


= Wie 


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To understand the true reaction kinetics of this 
reaction on the y-alumina surface has been an extreme- 
ly difficult subject to study due to the difficulties 
of eliminating the various catalytic effects caused by 
substances other than alumina. Among these side - 
catalysing effects, the additional catalytic effect 
due to the product sulfur in its liquid form was the 
most difficult problem to eliminate or solve. To 
eliminate side catalysing effect due to the condensed 
liquid sulfur others have tried using sulfur condenser 
(i2,osdy PLECtrLcG precipitator- (81); or ice bath: (37, 
39). But the elimination efficiencies have never been 
critically éheenea? 

Another problem related to the Claus process is 
the uncertainty on the equilibrium distribution of 
sulfur molecular species during the reaction period, 
Since above 1000°K all the sulfur species may exist in 
the form of So, while at room temperature as S. C75)-% 

Therefore, the Claus process may become exo- 
thermic or endothermic according to the operating 
temperature level of the reactor (24). The problems 
related to the equilibrium distribution between sulfur 
species with different molecular formula were attacked 
by many chemists but with controversial results. 


One of the most important problems which 































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— | | u 
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remains to be solved in the Claus plant is the effect 
of minor reactions on the overall process performances. 
In earlier studies the reaction (1.4) was found to 
affect the catalytic activity of y-alumina due to the 
poisoning effect of the product co. (65). Subsequent- 
ly, the poisoning effect was observed only under dry 
gaseous conditions, i.e. in the presence of H,0, 
reaction (1.4) is by-passed in favor of the more 


rapid reaction 


COS sen On HLS + CO (15:5) 


2 2 2 


Here the need for some type of improved bifunc- 
tional catalysts arises to improve the overall Claus 
process performance since COS impurity is almost in- 
evitably present in practical operational conditions 
(33). Actually, the loss of sulfur due to unconverted 
COS is almost 30 percent of the total sulfur loss (15). 
Until recently this problem was not solved satisfact- 
Orily because of the lack of the experimental data for 
both of the reactions on some bifunctional catalysts. 

From a practical point of view, it is surpris- 
ing that few data are available to apply to the Claus 
plant design for a catalytic fixed-bed reactor. The 
most important data for application to the plant de- 


sign is the intrinsic rate expression on the specified 


b 


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catalyst surface, physical parameters of the reaction 
systems and catalysts, and most of all, the effective- 
ness factor of catalyst pellets. 

Recently the intrinsic rate data have become 
available owing to the painstaking works by some 
researchers (65,66,72) for some type of catalysts. 
However, the availability of other information is lim- 
ited to the extent of primitive estimates of the 
physical parameters or of the effectiveness of the cat 
alyst pellets for the developed intrinsic rate expres- 
sion. No published data may be easily applied to the 
practical situation. These kinds of studies are 
really in need for improving, developing and predict- 
ing more efficient sulfur recovery in natural gas pro- 


cessing plants. 


Previous Works: 


After the Claus process to recover sulfur from 


the H,S containing gaseous stream had been invented in 


2 
1883 by the British chemist, C.F. Claus, this process 
has been widely adopted in the natural gas plants as a 
major scheme to recover the elemental sulfur since the 
first commercial application by H. Bahr in 1936. 


In 1953, Gamson and Elkins (35) carried out 


pioneering kinetic measurements of this Claus reaction 








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— , aieh : y 
a 5 -OEOL ah dst « a a ah iss 
soa ad tia Ba: : 


=. 










using Porocel catalyst in an integral fixed-bed 
reactor over the temperature range of 230 to 300°C 
at low space velocity ranging from 240 to 1920 hr -. 
In the late 1960's, an extensive research pro- 
gram was initiated at the University of Alberta to 
investigate intensively the Claus process as this 
process has become more and more widely applied in 
almost all natural gas plants in Alberta. The first 
investigation was carried out by Cormode (21) using 
Porocel catalyst to obtain kinetic data in a recycle 
differential flow reactor. McGregor (72) carried out 
a series of 80 experimental runs in a similar but im- 
proved differential recycle reactor, performed at four 
different temperature levels between 481 and 560°K 
951 er) and H,0 to 


obtain the intrinsic rate expression with activation 


and at varying partial pressure of H 


energy for the forward reaction of the reversible 
Claus reaction. 

McGregor (72) found that water vapor has a 
retarding effect on the forward rate of the Claus re- 
action at higher concentration and an accelerating 
effect at lower concentration, presumably by the 
hydrogen-bonding due to the water vapor on y-alumina 
OH-site. The rate expression obtained by Dalla Lana 
et al. (25) using McGregor's experimental data (72) 


was 


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lL. 





P P ? 
-r = 1.56 x 1072 lad ar ats be ex 7280) 
SO . I + 0.00423 P P RT 
2 H,0 g 
Eee ne ee (1.6) 


Liu (65) investigated the effect of NaOH-doping 
on the activity of y-alumina for the Claus reaction, 
and found that a 2% doped y-alumina showed highest 
activity for this reaction. Liu (65) also observed 
and explained (19) the co, poisoning effect in the 
reaction (1.4). 

Karren (53) carried out some 50 experimental 
runs uSing the same recycle differential bed reactor as 
used in McGregor's work, but with the sulfur condenser 
capacity increased. During his study Karren (53) found 
that the liquid sulfur acted as an active catalyst for 
the Claus reaction and the elimination of condensed 
liguid sulfur in the product stream was critically 
important to get accurate reaction rate expression on 
the catalyst surface. 

Most recently Liu (66) investigated the Claus 
reaction using an infrared spectroscopic reactor cell 
to get an intrinsic rate expression without external 


and internal transport resistances and obtained 


rye 


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a * a = ‘ ee iv ae Le “i ra 
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rT o 3 a 2 “ , w ) , 


1. 


2 
Pus Pso 





para, tor tee inhdefee eevee See aligns [27350 
2 ee Ce Pu? g 
Where .CGS (1.7) 


for the Claus reaction over Alon catalyst. The form 
of the rate expression obtained by Liu (66) was almost 
the same as that obtained by McGregor (72) except the 
retarding effect of water vapor was increased compared 
to McGregor's result. It is worth noting that the 
remarkable similarity held even though a pure y- 
alumina catalyst was used by Liu and a commercial baux- 


ite by McGregor. 


Topics to be Investigated in This Study: 


1.3.1 Evaluation of Bifunctional Catalysts: 


Some work has been reported (22,23,35,36,37,65) 
to develop and/or improve active catalysts for the 
reaction between HS and SO.- In parallel with the 


reaction between H.S and S05, more attention has been 


2 
paid recently to the reaction between COS and sO, 

(36,37,65,90). However, each of these studies was con- 
fined to one particular chemical reaction process with 


the view of investigating catalytic activities or a 


reaction rate expression for each separate reaction. 


~ weeny 


JNA AS 


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- 


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In practice, one approach adopted by industry has 
been to use a different catalyst more active for the 


COS-SO., reaction, but requiring a higher operating 
temperature than alumina, in plants where COS problems 
are more severe. More exotic catalysts such as Co-Mo 
supported on alumina have been used for this purpose. 
This problem could be solved more simply by 
developing a bifunctional catalyst which actively cat- 
alyses both the H,S-SO., and the COS-SO., 


reactions simultaneously. For this purpose one such 


(or COS-H.,0) 


bifunctional catalyst has been developed at the Univ- 
ersity of Alberta in 1973 to improve the Claus plant 
operational condition (23). 

The present ately has been designed to invest- 
igate the performance of this newly developed bifunc- 
tional catalyst for reaction (1.2) and (1.4) in Claus 
plants. In this study it is intended to compare its 
catalytic activity with that of standard alumina cat- 
alyst under the same range of reaction conditions 
encountered in industrial reactors but over a short- 
term reaction period. By this investigation a basis 
for understanding the bifunctional activity of the 
proposed catalyst may be provided. The catalysts 
tested included a standard y-alumina (Kaiser S-201), 
5.4% Cu-alumina, 12.08% Cu-alumina, and 16.07% Cu- 


alumina. 

























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1.3.2 Prediction of Claus Process Performance 


Leswhe be PLeg.crion.oL Thermodynamic Equilibrium 


In a conventional modified Claus process, 
elemental sulfur is formed through a two-step reaction 
procedure as was described in reaction equations (1.1) 
and (1.2). Many investigators (9,35,72,74,84) tried 
to compute and predict equilibrium conversion for 
reaction (1.2) but paid little attention to the sig- 
nificant endothermic effect at temperatures above 
850°K and to the effect of the inert gas content on 
the conversion level. These facts occur due to the 
different distribution among the sulfur species be- 

2" S¢ and So at different system temperatures. 
Regarding the sulfur species distribution over the 


tween S 


various temperature range, much work has been done 
during past decades to try to get definite 
conclusion still leaving the work in a controversial 
subject (75,88,89). 

In this work the sulfur species in equilib- 
rium were assumed to exist in the forms of So, S. and 


S That is 


8° 


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f 
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om 


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Furthermore each of the above three species was 
assumed to be in equilibrium with the others at the 
given temperature and pressure in state of the mini- 
mum total free energy. Equilibrium conversion calcul- 
ation was performed using the free energy minimization 
method originally programmed by McGregor (72). 
Finally, the computed equilibrium conversion 
level was compared to the maximum obtainable conver- 
Sion which could be obtained experimentally, i.e. at 


the lowest possible space velocity. 


1.3.2.2 Prediction of Reaction Path in the Front-End 


Burner 


The prediction of adiabatic reaction path in 
the front-end burner by the consecutive reactions (1.1) 
and (1.2) should be very useful to evaluate and com- 
pare actual process performance in individual stages 
with the corresponding idealized performance limits 
(24). To provide such a useful and convenient way of 
predicting the reaction path along the front-end 
burner before a catalytic converter, a graphical 


approach was developed in this study in attempt to 


present calculated values for "once-through" processes. 


This graphical approach may be readily extended to a 


number of problems occuring in closed systems: 


a2. 





















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12 


(1) the prediction of equilibrium conversions of 
H,S to SO. and Ss. as a function of temperature for 
acid-gas compositions ranging from 100% HS content, 


providing the diluting gas (N, or co.) remains chem- 


ically inert; 


(2) the estimation of non-equilibrium reaction 
paths in the front-end burner , from a specified inlet 
temperature and gas composition to a final flame temp- 
erature at which chemical equilibrium is assumed to be 


reached; 


(3) the use of these reaction paths to facilitate 
comparison of actual burner performance with the equil- 


ibrium prediction; and, 


(4) the estimation of sequential cooler and cata- 

lytic converter paths after the waste-heat boiler. 
This predicted reaction path diagram may be 

compared to the actual process operating conditions 
when the latter are plotted on the equilibrium conver- 
Sion -temperature diagram. Comparison of actual re- 
action paths and conversion per stage with predicted 
values should provide a useful criterion for determin- 


ing detrimental or improved process conditions. 
























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1.3.2.3 Prediction of Reaction Path in the Catalytic 

Converter 

In a "“once-through" sulfur recovery process, 
additional sulfur is formed by alternately cooling the 
gas stream and then passing it through a bed of cat- 
alysts. Elemental sulfur produced may be removed by 
cooling between stages. The graphical procedure to 
predict the overall process performance then requires 
calculating adiabatic conversion - temperature paths 
for the catalytic stages as well as for the front-end 
burner section. The heat release in the beds of cat- 
alysts is usually substantial, necessitating cooling 
between catalytic stages. The large diameter of 15 to 
20 feet and comparatively shallow depth of 3 to 4 feet 
of catalyst beds for the Claus process ensure that the 
assumption of adiabatic operation is realistic. 

In this study the catalyst bed was modeled in 
the two-phase mode originally proposed by Liu and 
Amundson (62,63) to check the effect of external trans- 
port limitations using pre-estimated internal transport 
effects. The internal transport limitations were 
checked by the Paterson's asymptotic method (82) and 
by Van Den Bosch's interior collocation method (99), 
and finally a standard computation was performed by 
applying the method of Weisz and Hicks (106) in comput- 


ing the effectiveness factor in the Claus reaction 


13 


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- d 
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sag » ebod sit ar sensisi tseot sr /MOLIOSE tenrud ; 
J _ - ‘ 


miloon paktsetieasoen ,fsaitnhstedue yilsver ex asaeis. 


~y = 
™ “7 ” = ~ 
ot CL to s1999Mmbib isi siT ».eopsiae cial -63 BO cooused: pie 


ra 


sot b of € to ddasb wollede yisvissxsqmoo: bas tos? as 


sis dgadd exvens esasoo%tq aged oft tot ebsd geyts i569 io 


7 


. »ideiises ek dotfereqo oftsdsihs ‘To Sc 


ni belebom -asw bed teaevisaiso sit yburte eid at ie 
bas uid vd Beeogord eitenivtso bom. saad: 


~pasit isnrotxys to jtoetts at Avens os (8a 83) noe 






tirogensid Isatetat botemisze-exg pata “endlsad aoe a 







StSw anndihae ies tf Proqetat: iepredmi. str 


nf ate Oy 4 


ine ($8) bos rm | oldosanyes: 2 hia ad vd 


. : = rs odin 2 


hm 


-~ (ee), 





s 
mut, 


14 


system. The computations were all performed on the 


basis of the rate equation (1.7) obtained by Liu (66) 
on the Alon catalyst. 

Using the computed results the converter 
design or catalyst performance may be easily obtained 
Since the effectiveness factor is readily accessible 
on the Thiele modulus - effectiveness factor diagram 


for this specific catalyst and reaction condition. 


. ‘ ln tan im _— = 
muotiea {is stew 2ecoisejuqnoo eff .me7aye ime 


: j Ss . 
Ye 
i Z i : < 
: 4 abs . wih soe bn | A hf cl 2 i 
LL£B7C0 ({ i) AaQd thupe Stil Bus sf 2 LS 
, . " — “a a 
wrevias sa. mm LA be i 34 
=< 
nae = 
i 2. 
3F , me | lve» , ra | o> nz pi 4 La } ; — 
= : 


i. a 


bec ico vitase od. yan ebosmtotieq Jayistso 20. 





CHAPTER II 


LITERATURE SURVEY 


Claus Process: 


Sulfur recovery units of the Claus type with a 
thermal and one or more catalytic stages have been 
applied for more than twenty years in oil refineries 
and natural gas treating plants in those instances 
where large quantities of hydrogen sulfide are produc- 
ed. 

In the modern modified form of the Claus process 
a number of stages, usually three or four today, of 
catalytic reaction are used to increase yield of the 
elemental sulfur since the high exit temperature in a 
Single-stage converter could limit sulfur yields to the 
equilibrium values. The sulfur recovery efficiency has 
been closely approached to the theoretically possible 
yield ranging 94 to 96 percent recovery (26). The 
remaining 4 to 6 percent of the sulfur present in the 
Claus unit off-gas, is converted into sulfur dioxide in 
a catalytic reactor or a thermal incinerator, and then 
discharged into the atmosphere. 

The overall Claus process consists of two stages, 
in which one-third of hydrogen sulfide is burned com- 


pletely to form sulfur dioxide in the first stage 


We 





















- -22gD079 assis’ I.8 


— —— tere + 


r i > ; ? ee See ee | > ’ 
ayvsel5 sat to eFialy vrevoss: rutine 


nesd “ ssente sttvists>. s1om' x6 sao bane Leeredns 


iy avon t oR Setiqae 


S9Y YOMows ster 


4 
~ 
4 


tw 


eoonrsteni seods ni estasiGg erEesssxs. esp isisisea bas 


~subotq ors sbiitise nsporbyd Yo setstita ABup eyisi sisiw . 


’ { SJ AN 4 . ~ ¥ 


.be~ 


Ssesnoirq euntld edt to myo? bettibom orebom sds al 


ileves ,aspete Yo rsdn B. 
| : i 
add lo bleiy seseront, o¢ bees sexs noitosea otsyissso 
‘ i : ‘ - > Pi 7) 
: M . z - ; ~ ~ — a i x ~ 
5 oi stutsisqmet tixe dpid eda somte xwiige Tedasmefe 8 - 


’ 


adit of ebleiy uwilIwe timifL Biwos 3987 rsvcoo. spste~sipnate — 


si yonsioltte yrevoves tl ve off .eenlev mukudilivpe- 
* ve 4 : , _ 
esldteeaeoq yilsaoisetosads ane oF. Reneanes eee ased 


« 





. (88) yasvonst tneored ae Se. Je “emtpast as 
q ‘ 


oft th dnsasta 20% ine sf to snso%0q, 3 a3 & erintenes 


ai ebixolb tr3ie ota beg'reva0p , enQ-At0 ska evs2 ig 


iS 


— a 
neds bus ,rotstenkons Semsets 8104 sodonion ousesase ide ‘ 
; res ey 1 ‘tet Re ae « 
s 4 -oredd it Ts F efit oat - 4 : , 

a ; pr tee. 
_Bopete ows to erakenen, hee awed LD et 


es L™ 
i- : : $ ae: oe 








according to the reaction (1.1). The product gas from 
this step is then blended with the remaining two-thirds 
of hydrogen sulfide and passed to a catalytic converter 
wherein hydrogen sulfide and sulfur dioxide formed 
elemental sulfur and water as presented in reaction 

(T <2)}. 

Hydrocarbons present as impurities fed to the 
furnace with hydrogen sulfide are converted to carbon 
dioxide or else pass through as hydrocarbons in the 
Claus reaction process. Carbon dioxide and some hydro- 
gen sulfide react in the furnace to form carbonyl sul- 
fide according to reaction (1.4). Some carbonyl sul- 
fide appears to pass through the catalytic reactors 
and is therefore lost. So desulfurization of Claus 
off-gases has recently received special attention 
(6,50,78) since stricter emission control regulations 
require increased efficiencies of the sulfur removal 
processes. However, most of Claus tail-gas treating 
processes have proven to be rather costly (6) and the 
best approach may therefore be to improve the perform- 
ance of the Claus process itself in the hope of elimin- 


ating the need for subsequent tail-gas processing. 


Sulfur Recovery Processes Related to Claus Units: 


Numerous processes have been studied and pro- 


posed for recovery of sulfur from hydrogen sulfide- 


16 













fort evap touboxrq ent (1..£) cojsdo8es eas og pal bracss& 


tdg-ows prinisme:t ofd ddiw Sobaeld asda et qese elds m4 


set¢uevnos oftyistso 's. ot Heeasq bus obitivue meporbyd 


™ 


xoih w#tIive;bis ebitipa mevosbyd niaxetw ~~ 


moijosey ak betaseetq es teteaw bas xwtive {stnemelo 


» (&.1) 


* 


esitixwami es tasasiq anodrsco1byi 
2 7 ' 
diso> o3 besrevaod e156 sbitive sovotbyti dtiw goanist ae 


a 
7 


eit at anodissotbyd es dpwowis eesg eels xo sbixolLb. 4 


byri omoe2 bas sbhixoLb nodis9 -Bago0%g noLjoss7 evel) -. 















. A oe ® 


~fge Ivnodte9 arso2 53 sosntw2 ode ai soBe1 sbitiua nep . a: 


i +e 


~ige Lynodxs> smoe .(d. £) noise neo os pakbuco 26 obit “a 


etotoseSt nisylstso orig Tipwozds | eesq oF s1se8qgs abit y ict 
avelD ho holden t203 Lames 02 deol Sxotex0d3 ak bas 
motsaesss Isiosge bevieos: yistnsos . Sins aoanp-230° 
aholteiuper foutnes noisatne testo tate sonie ($f, 02,8) 

isvomex isitve edt. tc eotonsindtie, beanetont ealupes 
paitsort asp~ihed evel to 30m (2ev8WOR | -BoREso07g 


sit bas (3) yiteaoo tedsez ees novos, wae: aea 7034 om 


a aera “a hed ow, 4 
.paie e004 nen-fied: dneuperc due 3 (ot been oft a 


ga ma 


Pa 
: 


yer: Cast. 2 y 
-nimiis to sdod peree b atesst saeco nt aE 






n 

? 7 at 
= ail . a 
; i nie hye 


i? ee) od el are ae 


‘103 a nee Oo” 


17 


containing waste gas streams (95). Most of these 
processes can be divided into three major categories 
which are related to the Claus process. 

(1) Direct catalytic conversion to sulfur. 

(2) Absorption and catalytic conversion to sulfur. 

(3) Adsorption and catalytic conversion to sulfur. 

Direct catalytic conversion scheme is applied 
for high concentration of hydrogen sulfide and is the 
most widely used process in present sulfur recovery 
plants from natural gas, and will be discussed in more 
detail in the following section. 

The absorption or adsorption scheme before con- 
version to sulfur is usually applied for gas streams 
with low concentration of hydrogen sulfide to provide 
high concentration of hydrogen sulfide in the absorp- 
tion medium or in the adsorbed state which is enough 
to initiate the catalytic reaction. In the absorption 
process, various organic and inorganic liquid solutions 
are employed as an absorbent such as potassium carbon- 
ate, sodium hydroxide, and monoethanolamine or dieth- 
anolamine (95). One typical example of such a process 
is the SCOT process (78) which was developéd to in- 
crease the efficiency of Claus conversion by absorbing 
HS from the Claus off-gas in which it is present in 


very low concentration. 


a -paidioads vd srlenerece pai Sa eae 
ie Amer pee yeth oa ae ; 
ah uc genet 


























avedt 20 320M .(2@) emsets=s Gap adesw pitiatsinod 


Iso totam seins o¢ni bohivib ed ABD eee 200 ta 


$f eng of Hosa ler exh: d>ine 


ec ~ 4 rr Pires ~ 
- ¢ "Se | ae 
fue of ndkerevadd &isyiesso soeata = §=« tf) 
ff P . _ me i 5 - ‘ , 
I 2 I 10D OSTIVLIBSSS Hits oi tquoedA gc}. i% 
¢ by 
v, 


3 rf > 3 rot tay too oisyl 5459 Hre abitqxoehé {€) ; a 


.5 


i ; omeadoe coieresvnoo oftylese: yooxrta 
, ee ‘ oe EN 
a SOL iye nmeporwy ret eo) Iaei3a = ee ‘tek d +02 
+ 0 ewes uo oe” ee @ Saenn’ tet bart «ee ff ~ re w Jaom 7 
~ ibe 2miesze tg i eseoosg epew {+= SD.L¥ &6 aa 
A 
. cotipath ed Ifiw Bas. ,.see Lstesea mort s7mss 


— o F ‘ ‘ e : : 
noisgoee pniwolfot sid at [iateb . 


oo a 


nos etotsd sreroe noksqroebs yo noljquoeds ef 


Te one . : “ee me ro 
empsisa esp tot beilqqe Yitsces af autiuve ot noOLeIsv = 
. .- 
aa > 


voug of sbitine’ negorby:i ‘to ads yertnsonod. wol dviw — 


cae 
yroeds off ai sbitive aspozbyd Te not RY Ee 


fle mons eat. doin tete. ecdtoRhs eaz- nt 0 ) my bem nois 


+1080 rf aie tole ots yingsD ort ossizint oF ~ = 


moLsragte: VAs ~ ee es 


rituloe bhuptl sia sptont Sas obaspte suotiev , 829902 


ape 


; * 2 ae 
-nodyso muteastod es move dried senda ieb a8 peyolams, STs i 


as 


"to onins loners eonom bas vobixoxbyd mu iboe evs 


3x0 & dove 30 sflqmexs Lecter? eno, = (ee) ectns fons y 


’ ie * 


-nik of bec qoleveb eau do batw ety at 





Ren (a 4 , : A 
A ‘ j iS v oe a if 


> 
oo 2r 
ae 





S 2 





The adsorption process usually involves zeolite, 
activated carbon, or alumina as an adsorbent. The 
main objectives are to use the adsorbing properties to 
increase the dilute hydrogen sulfide concentration and 
then, the ability to act as a catalyst for the reaction 
of hydrogen sulfide and sulfur dioxide. 

In the presence of free oxygen, hydrogen sul- 
fide is catalysed to elemental sulfur by activated 


carbon as indicated by the following reaction (69). 


active carbon 3 


dL. 


n 
Another different approach to hydrogen sulfide 
control is the Sulfreen process developed by Lurgi 
Apparate-Technik and S.N.P.A. (58). This process util- 
izes the catalytic ability of activated carbon for a 


high efficiency of Claus redox reaction to yield ele- 


mental sulfur. 


S 


— — 
2H,S ay SO. 2H,0 ts - Sn 


Overall removal efficiency of a Claus-plus- 
Sulfreen plant has been reported to reach 99.0% or 
more with sulfur removal efficiency of Sulfreen process 
of up to 85% (58,69). The Ram River plant in Alberta, 
started up in 1972, has a Claus unit followed by the 


Sulfreen process. 


18 























x 2 
" ‘ 
; ,.vioveii viissey eaqgoorg moisquoehs ome. > eam .' 
‘ ~ > 
adtoabs as 2n snimyile to ,nodtso Sarsves08 
ot 2 nnaidvoebs eit cau o¢ ous eevissefdo alsm 


; s oe ch 
is nolte1tteotos sbitive nspoubyd edelib eds sesatont 


noktoss¢ 4+ 402 devisiso © 6s 368 oF yii lids ei ,nons 


=" @ 
~ * ‘D> i- he at > free Te a Pau. — ; 
-obixoib tutive bas sbitise asponbyd 20 4 
L Pons <fepyxe @d12 30 sonameta eft mt 
: v we eo ~ eo 


2 . - > . 
otsvisos vd wilwe Lsdiemeto ot Beeylasso ei eh4% A 


betsoibakt es anedxzs> 


, 
e) 
, 
n 
ce 
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-s 
, 
“ 
eo 
poe 
~* 
od ‘ 
b$ 
®, 
pe 
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wat 
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oa. 


nodis> Svisos ( ~ 
= 
2 —-— +— OJ ne eer 0 « gi 
fi - * o , 
ebitive asporbyr og dosorqas tre1stib usds0nA . 
“r ‘ ‘ : . - . 4 ' = « 
. Ipind ve beqofeveb aesserq meettin? odd ai Lomsnoo .- 


a 


rou cs Sa >: Cj erdt e (82) - A - q 4 a bits: Aindoot-esexegh ae 


5 10% modzso bosev ison Bo sidings obtylstso ont eosk a 
«ifs bieiv ot anoizoses xebs7r evsid to youstai?3 5 ‘jet a 
Ao ors. = | cxuhare: issoom 





] ob } es Py Mass mas 
2= + O,.8S Ga et 20 BMS 


‘ a a fe o ; : 
; > 4 % 4, = i 


~eulg-aveld 6 20 youwiolats Oe eS (aa 
to «20,28 dosex oF ‘bairroys need asd dasitq ‘moottive 


eo 
s2an0Ig as ortive as yonibitve eer ge fy ne 
’ < Pay 4 fs rte ss Sa eee: ’ 
3 sev ted tn nt dnstq’ ve A msa +o 
a a) ere 
pity. Yo bewoltod > 
% ee P a ie ry | 


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ee fas 


| 





’ 
a" 








19 


Simultaneous reactions in Claus Units: 


Major reaction in the Claus unit in a sulfur 
recovery plant is the catalytic reaction between HS 
and SO.. However, appreciable amount of COS or CS, is 
produced in the furnace and waste-heat boilers by 
reaction between hydrocarbons and co. in air 
(33). According to Cameron and Beavon (15) one-third 
of sulfur loss in the Claus unit is due to the presence 
of COS which may not easily be converted to elemental 
sulfur. Therefore, the conversion of COS to elemental 


sulfur is a very important factor to improve the over- 


all Claus unit performance. 


2esel HS - Air Reaction 


Oxidation reaction between HS and air occurs 


in the furnace and waste-heat boilers of a Claus unit. 


2 H.S + 0. —~¥ S + 2 H.O f2oL) 


BiN 


AH = -123,924 cal/gmole of HS 


This reaction provides sensible heat by its 
large amount of heat of reaction to raise the feed 
stream temperature of Claus reactants to the catalytic 
converter. This fact necessitates a better understand- 


ing of this oxidation process than currently available 


et 
























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¥ ‘a, 
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. tod teed-etesw hoe eoér1g2 edd nt baouboxd i 
5 nl 09 bas enodisoorbyd neewted noizosex 
brirt~-eno (@L) novssd Snes roremsd of gnibroopé « (EL) a 
20f : 3 si gi tie evslD edd ak eaol tviive to 7 
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Istnomofo ot 20D Io mOkexavdos ond eiotexed? siuvtive ~ 
-19v0 oft svoigmé od rosog7 JnemeO_ME yrov « at awitye ca 
: | ~ 


“s - SOTTBeTEOT oq tins. eval {in - * 
™ , = , 
df) ; 4 ror ee NS 

< 


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ee Ee 





21u900 wie Bo6° oR asewiod. & {tosex, et ttebin av 


tins austid s to axe liod taed=ededw Nee eosaayi oft oi d 
' ry _ ~~ er 


as 


(LS) eS 2 i <—— .0 + Sal Bas e 






“hy a ; - ; i 


ati yd teed sidks atiod- eobtvow ae 


beel ord eeiss oF. nokjobo%: 


Af 1 a 


7 - 7 


- 


_ olny fete iad 10m spon 


tod 6 
‘& é 
Ae ‘ante! 


and one aspect of this is the accurate determination 
of thermodynamic equilibria between HS and air under 
Claus furnace conditions. Then, the adiabatic oxida- 
tion reaction path may be predicted based upon the 
equilibrium data. 

After Gamson and Elkins (35) made the first 
study on the reaction equilibria between hydrogen sul- 
fide and air using Kelley's data (54) for the equili- 
brium constants between the various sulfur polymers, 
Many investigators paid attention to the equilibrium 
calculation using various kinds of techniques. 

Recently McGregor (72) used White's free 
energy minimization technique (108) to compute the 
reaction equilibrium in Claus process units using the 
free energy data compiled by McBride et al. (71) to 
get results of a slightly higher equilibrium conver- 
sion level than Gamson and Elkins'. 

Bennett and Meisen (9) computed the reaction 
equilibrium for the reaction between HS and air over 
the temperature range of up to 2000°K. In their cal- 
culation they included significantly increased number 
of chemical species thought to be present at equilib- 
rium under the condition of oxygen to hydrogen sulfide 
feed ratio of 0.2 to 1.8. The computed results have 


shown that considerable amounts of SO, SH, S50, HS. 


ae, 
‘ ie : 6 
7 ~ 7 
“ Hs 
7 
eos OF. 
' ‘ ‘ ae ne pa 
1 job 4 oon oft si stds to tosqen sno lBme 
Tat < a | 
xesbhavu 2 ns @.H neewted dilinps ofmsnybormtsds to 8 ~— 
-sbixo oisedsibs e439 .aea?t .enokttibnon eosaxv? evsio 
3 uw Beasd fatotba: i vem dtsq nofsossyinois” 
Sieh muctditiups 
te +} efset (26) eaLAla Sas noetisd FotTA 
a 5 j i28Swla Sd 51 xd mee 9 of I3GHS 3 as ro ybuse 
: ot ($2) ib e'yolloxt pater sie + 5 Sbi2 
,sienmrylog twilve evolusv ond ms Wied eInS Fanos ania 
mriiditinups edt ot acitaestts bisq e10otsp.Ls teevak onal 
; Meat 
.2oupindoet to sbnia evoixsv paiey nolssinoisa 
‘ . 4 


: ee ol i$ 
est} e'ssinw 59 ($v) tops1DoM 
fi: jtugeioD oF (80L) P=y4 
oi3 piiew eatin 
os (£9) [5s 
; x 
-tevnod mutiudilics® retip 


rs 


6 
i wat 


SoBox otis besucqmos (@) coeiem Sas s3sna08 oa 


tsvo tis bas = neswited noijoser sii 
-[s9 tions ne -A°0008 ot gu to epasy 


J ¢ . 
sedmun baer 
‘eckial, 


-dilinps 3s sneeo3g ee 


of 


wat se 


bea watt aston ee ugne att ae 
he He . hike ba Aa paar rey 
x a ye “a tg 19 


2290014 ene td Fs Og “muiidt Ciups: aatsseee 
+s ehitdomM yd bsiigqmos adab as prona ‘seat : 
bid Prereate S- 0) aatdasn 2 
4i% Bas it oem nad fovet « 


sevoni xisheottimpie seven: x 
3 “siteuonts ees 
sbitive neporbys 2 se 26) HbR 























vitasoen : hie 


piasdood OEE Se vexone’ 


a? 2 


Td ce: 


ba 


F 10% moisdt 


f 


ceiine ee 


<— 


Thets : 
Ah Uy “sh ™ AN 


4 " 
+4 to 


ala A 
a 5 





| 









me 






and H, are present at equilibrium condition at higher 
temperature range of above 1500°K. They also included 
a1 Sy S3, 4 and S. in addition 


which was employed by McGregor's 


in their calculation S S S 


to S S- and S 


ai 76 8 
computation (72). For all the species included the 
results by Bennett and Meisen (9) were almost the 
same as McGregor's with their computed results falling 
in the range between Gamson and Elkins' and McGregor's. 
Bennett and Meisen (9) demonstrated that stoichio- 


metric feed ratio of HS and air is the best strategy 


2 
for sulfur yield without an additional preheater be- 
fore the furnace. 

Neumann (80) investigated the effect of CO. 
and H,O content on the equilibrium conversion of HS 
by the reaction of HS with air in the front-end 
burner of a Claus unit over the temperature range be- 
tween 800 and 2500°K. In his material balance calcul- 


ation S. was considered to be the only sulfur species 


2 
existing in the product stream. According to his 

data, the conversion of H,S was adversely affected by 
the presence of H,O but favorably by CO. in the feed 
stream. His computed results also indicated that no 
Significant amount of CS5, SO, Oo, CHy, NH, NO and 


NO, were produced in the front-end burner over the 


above temperature range. 


21 


Of 
pas OK 1a + pit ago ar 


ote gat zevo xowmid Brie 


Me a 


fso. sonsied Isireteam etd ai 2902S ane 908 noowd 


seqe «uritivge ¥iee ane) ad of ‘berebtanoy new fae 


botostis yloetevbs eew. Oph 20 nobetsvao3 ody ash a 
eo edt nd oO? Ya yidexoved, yeaa sane, a a 






















itibnos mixdilivpe 26. jaseexg ote .H ba - 
& YoAT .8°O021 svods 30 spns3 situtaiegmes . id ag 

bs ai «2 Bae 2 a ae aorteinofies riedt, at 
OM yd hevolqme esw dofdw ,2 bos a” 168 oF we 
bebuiont eefoege sit fis =F . ($8) nosis esugmos 
xew (@) asetemM bas tteqned yd adivest 
" aspheodennaie xfer? (itiw 2! ropss0om RE BINGE 
eID exee antate bes nmeeamsD deewsed epast et at 
iniots tenid Bbedsatenmomeb (2) .1seLen ba6 +tecqned 


i sft 2i xis bas 2.R to ofi¢sx Best oistom 


‘ as ° 
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ISsyvseensiad ISnols OE fis tuo! t ftw Dis Ly TOa2 j2 io3 ’ 


,soscrzet sdf s1ot 


, 


to tostie eit bestBpiseeval (09) snemvev txvok, SA 
' 4 » — 
2 J 


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brs-taori ort, ai aris dotw Bol 0-0! Lioeos ont, vt 
Spn61 stiujsisqmes sit revo: ting aunid > ‘& to renxud 


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ue 
a owe 


pnt bror0é meoxge agubexq ota st paisa, xo 


oT 


jeds beseosaek: oBls: pales : 


at: 








On the other hand catalytic oxidation of HS 
with oxygen has been an interesting subject in the 
pollution and odor control point of view. Ross and 
Jeanes (92) studied the oxidation of hydrogen sulfide 
Over cobalt molybdate on a-alumina and related cata- 
Toone using an integral reactor to reduce the hydrogen 
sulfide concentration below the odor threshold level, 
0.007 ppm. 

Steijns and Mars (96) studied the catalytic 
oxidation of hydrogen sulfide into elemental sulfur 
with molecular oxygen over the temperature range of 
130 to 200°C employing active carbon, molecular sieve 
13X and liquid sulfur as catalysts. The results 
showed that liquid sulfur acts as a catalyst for H,S 
oxidation. The kinetics and the activation energy of 
the reaction were found to be essentially equal on 
various catalysts of different chemical composition. 
They also found that the small pores less than 12 A in 
pore diameter were filled with sulfur and the catalyst 
surface area was lost as the reaction proceeds. The 
catalyst, however, was found to be almost equally 
active as the sulfur-free catalyst, which could be 
explained by the catalytic ability of liquid sulfur 


trapped in the pores. 


22 


sd 


_wrive biupit 29 i dns piseninsh 


: re “f ~ . 

} TSHEULID tiyvi yf 5 Tileze, at re 
r itasist sed ase mepyNxoO dtiw 

te 1 
y tp jrngfoq f tnoD TODO brie oo St LL0g 
: a 
a ortat t Beatbute (fe) eanset 
, 

6is7 Sri i no stabdviom jJisadoo Isvo 















, (bss o OL sipadhi ob priesr eseyl 
, biode a. ett woisd anoisayinsomoo. sbhitiae 
nig. T00.0 
ttyieseo ed3 Setbuste (3¢) Bem bab enc tose 
ott {strnomele ofai sbitige nsporbysh Fo noltsblxo — 
p= ; ; ‘ - 
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— P= - 


os me H,S - SO. Reaction 


After Claus (20) discovered the catalytic reaction 
process between HS and SO. over the iron oxide cat- 
alyst, the Claus process has become the common sulfur 


recovery scheme from HS produced in the coke oven 


2 
until the late 18th century. Starting from the signif- 
icant contribution by Gamson and Elkins (35), who 
studied this reaction in an integral bed reactor pack=- 
ed with 4/8 mesh Porocel catalysts, many researchers 
(22,35,36,65,72) devoted themselves to improving this 
famous process. Many of the research activities have 
been focused on the development of improved catalysts. 

The catalysts usually employed for this reaction 
were sulfides or oxides of aluminum and other transition 
elements. Among these iron oxide, bauxite, manganese 
oxide, alumina, glass alumino silicate, cobalt sulfide, 
molybdenum sulfide and cobalt molybdenum-alumina have 
been most frequently studied. Recently active carbon 
and liquid sulfur have also been the subject of many 
research projects (58,96). 

Claus originally used iron oxide for his inven- 
tion of this process. But today, bauxite has become 
one of the most popular catalysts used in Claus sulfur 
plants. One of the problems related to the bauxite 


catalyst is the mechanical strength of the catalyst 
pellets. 


23 


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In an early stage of this process, Gamson and 


Elkins (35) studied the kinetics of the reaction be- 
tween H,S and SO, in an integral bed reactor using 
alumina as a catalyst in their pioneering work for 
this Claus reaction kinetics. The data were obtained 
at very low space velocity of 240 to 1920 hr? over 
the temperature range of 230 to 300°C. The resulting 
conversions of HS to elemental sulfur were 92.9 to 
97.9% which were inconsistent with their thermodynamic 
analysis of this reaction because their measured con- 
versions were higher than thermodynamic equilibrium 
conversions for the given reaction temperature and 
pressure. 

Pyrex glass surface was found to be an active 
catalyst for this Claus reaction by Taylor and Wesley 
(97). The proposed reaction rate on the glass surface 
was proportional to the external surface area of 
glass with the reaction order being a first order with 
respect to SO, partial pressure and 3/2 order with 
respect to HS partial pressure. 

Hammer (45) studied the kinetics and mechanism 
of major reaction of the sulfur plant using cobald- 
molybdenum-alumina catalyst mixtures in the glass 
differential reactor. From his experimental data 


Hammer (45) concluded that Claus reaction takes place 


predominantly on the external surface of the catalyst. 


r 


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A dual site mechanism for HS and SO. on the catalyst 


2 2 


surface was proposed by Hammer (45). According to 


this mechanism adsorbed HS dissociates into he and 


SH and then reacts with adsorbed SO.. 


Deo et al. (28) studied the adsorption and 


surface reaction of HS and sO, on y-alumina using an 


infrared spectrometric technique to conclude that a 
strong hydrogen bonding exists between both HS or 


SO, and surface hydroxyl groups and a chemisorbed 


form of so, exists on y-alumina which reacts with H,S 


according to the Claus reaction. 


Most of the adsorption of HS and SO. on 


Y-alumina were found to be physical adsorption on 


Lewis-acid sites (67). But chemisorption has been gen- 


erally regarded as a precursor to catalytic reactions. 
Liu and Dalla Lana (67) found that the chemisorption 
of H,S and SO. 


small but vitally involved in the reaction mechanism. 


on basic sites of y-alumina was very 


On the other hand Khalafalla and Haas (57) proposed 
a mechanism for the chemisorption of sO, on an OH 


nucleophillic site as 


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25 


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26 


George (37) used a commercial cobalt-molybdate 
on y-alumina (Girdler G-35) as a catalyst to study 
the Claus reaction in the integral bed reactor. The 
initial rate was obtained by fitting the experimental 
data to the expression proposed by Mezaki and Kittrell 


(76) 


X = A, tanh [A, (7) We terk 


and extrapolating to zero conversion. The resulting 


rate expression for H,S-SO., reaction was 


2 2 
k "HSS 
H,S (1 + 0.1 eHLO! 


George also noted sulfidation of catalyst part- 
icles during the reaction by observing the change of 
color of the catalyst from original deep blue to 
black. 

On the effect of mass transfer resistance in 
the Claus reaction system Landau et al. (59) studied 
H,S-SO, reaction over a bauxite catalyst and concluded 
that the reaction was controlled by mass transfer 
resistance. Unfortunately, however, these authors did 
not investigate whether it was film or pore diffusion 
effect which was controlling the reaction rate. 


McGregor (72) also investigated the diffusion 


effect on the reaction rate, under the practical plant 


















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27 


operational condition, concluding that external mass 
transfer resistance had neglegible effect while inter- 
nal pore diffusion effect was the rate controlling 
factor through his diagnostic calculation. In 
McGregor's kinetic study for the reaction of H,S-SO, 
with Porocel catalyst of 28/35 mesh size it was con- 
cluded that only the external surface area of the 
catalyst was actively catalysing the reaction. In 
Spite of the diagnostic calculation results showing 
the significant effect of pore diffusion, McGregor 
failed to observe any significant effect of internal 
mass transfer resistance in the data obtained and con- 
cluded that produced sulfur, which would primarily 
be S¢ at 260°C, would be very slow in diffusing out of 
the catalyst pore and might be forming a monolayer or 
more on the internal surface. When a monolayer or 
more of sulfur forms on the capillary walls, the ef- 
fective diameter is so reduced that reactants cannot 
diffuse in and more likely, S¢ which may have been 
formed within the pore, cannot diffuse out. Sulfur 
thus presumably remains permanently inside the catalyst 
pore. 

George (37) also investigated the diffusion 
effect on the reactions HS-SO.,, COS-SO,, and found 
that there existed significant pore diffusion effect 


with H,S-SO, reaction and negligible with COS-SO,, and 


>e 


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COS-H,0 reactions. These results are to some extent 


foreseeable from the fact that H,S-SO, reaction rate 
was observed to be much faster than the other two 
reaction rates. 

According to Pearson (83) commercial cobalt- 
molybdenum catalyst was as active (84% conversion) as 
the active alumina (S-201) (83% conversion) for H,S- 
So, reaction at the reaction temperature of 275°C and 


the gas space velocity of 50,000 hr} in a micro 


reactor. 
is We Sg a 9 ia SO. Reaction 


The reduction of SO. by COS proceeds according 


2 


to the reaction 


3 
Sere t+rencesst-25n 25208 ea, on (2.4) 


z 


with heat of reaction (60) 


AH = 6760 - 2.75 T + 0.0028 17 


Lepsoe (60) investigated the kinetics and effic- 


iency of various catalysts for this reaction using a 
quartz tube reactor filled with catalysts. According 
to Lepsoe's observation this reaction proceeds slowly 
even in the absence of a catalyst at 800°C and almost 
any kind of hot surface is capable of catalysing the 


reaction at this temperature. Lepsoe (60) also 


28 






















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observed that pyrrhotite was an efficient catalyst for 
this reaction at 700°C, and at lower temperatures al- 
umina in slightly hydrated and acid-soluble forms 
(boehmite) was a remarkably efficient catalyst, and 
that lightly calcined Guiana bauxite and activated 
alumina were both satisfactory. No detailed investi- 
gations regarding the reaction mechanism have been 
made in his report except the probability of formation 
of surface compounds between sulfur dioxide and the 
catalyst. Adsorption of sulfur dioxide was manifested 


by the fact that the catalyst tenaciously retained 


sulfur dioxide once it had been exposed to concentrated 


sulfur dioxide-gas mixtures. The activity of the 
alumina catalyst was found to be reduced after it was 
exposed to high temperatures for long periods. In the 
temperature range of 300 to 600°C, the reduction of 
sO. by COS appeared to be of first order with respect 
to SO.- 

Gamson and Elkins (35) also studied this re- 
action over bauxite catalysts in an integral bed 
reactor and found that with a gas stream mole fraction 


Ot oeet oOo, eo. 155. SO and 91.75% N» yields of 902 


pm 
or better were obtainable over the temperature range 


between 250 and 350°C at a space velocity of 200 hrf, 
which was far below the practical sulfur plant opera- 


tional condition. 


29 



























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30 


Liu (19,65) conducted kinetics and mechanistic 
studies of this reaction using a recycle differential 
reactor over y-alumina in the temperature range of 
552 to 557°K. It was found that the reaction only 
proceeds over a short period of time due to catalyst 
poisoning by carbon dioxide which is one of the pro- 
ducts formed by this reaction. Their investigation 
by means of infrared spectrophotometry showed that 
carbon dioxide was irreversibly chemisorbed on the 
surface of y-alumina and thus presumably occupied the 
Sites important for the reaction, which in turn 
poisoned the catalyst. 

Liu (65) suggested through his IR study on this 
reaction mechanism that COS was not adsorbed on any 
of the surface hydroxyl groups of y-alumina but only 
physically adsorbed on y-alumina through sulfur atom 
at Lewis-acid sites (the surface aluminum ions). Liu 


(65) also suggested that SO, formed hydrogen bonding 


2 
with the surface hydroxyl group. He attributed the 

poisoning phenomena in COS-SO., reaction on y-alumina 
to the chemisorbed CO, on the surface Lewis-acid site 
as a result of the surface reaction between adsorbed 
COS and SO,, eventually making no surface Lewis-acid 


Site available for physical adsorption of COS. Sub- 


sequent studies revealed that co, poisoning did not 




























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snimiis-y nQ io ee BOG 08-809 nt snemonadig pcidostog - ° 
ofie bine-siwel sostaye enla: AO <0 ‘fiscinoe imesio eds ot 
: *% i ty 7” 
bedroeebs toows od nor soa per > o Sinser 6 en en = 
4 ’ 
_bios-eiwel SORTING. on. araal. se BAg0 ys ge 03 7 
* a ee ; weer oe naces 4 nr 


| neve: WM giye! oO -* at ee 
ang be mS 


occur when H,0 was present because of the occurrence 
of the very rapid hydrolysis of COS to HS. 

Querido (90) used Cu-on-alumina, CuO-on-alumina, 
and MoO,-on-alumina catalysts to investigate a catal- 
ytic reaction between COS and SO. at the temperature 
range of 986 to 506°F with contact times ranging from 
0.07 to 0.35 sec. in a tubular reactor. In Querido's 
experiments, CuO-on-alumina performed well only at 
temperatures above 1000°F while MoO.,-on-alumina was 
poisoned very quickly by sulfur. His results showed 
that Cu-on-alumina catalyst was most active for 
COS-SO, reaction. 

George (37) also studied on kinetics of this 
reaction using cobalt-molybdate on y-alumina catalyst 
in a integral bed reactor. The resulting rate expres- 
sion was reported to be first order with respect to 
COS and zero order with respect to SO, with an activ- 


hidotel Pats Pe In his study any 


ation energy of 18.0 
poisoning effect by CO. product in this reaction was 
not observed on cobalt-molybdate on y-alumina catalyst 
which was in contrast to the strong poisoning effect 


by CO, on pure y-alumina catalysts as reported by 


2 
Chuang etuals (19). 
George (36) also revealed that catalyst basicity 


had no significant influence on the reaction rate of 


COS-SO, reaction on the cobalt-molybdate on y-alumina 


x 


’ a1 . 
[ p< 
a Ray 
bs Fa % 
y’ ; ' 
; 
«ah oth A 
i om fe" ' 6 
t+ to eens neesuc esaw O,H setiw t89S0 
2 : 
. a Ss < ». , ig >» 
i | G09 to shevfoxyByA' bidst yxev ent tp. 
: . ~ es ‘nays So awe, 
nin yO , Boiris~s gy) been (08) obites0 a 
wnt c  perimybe-do- OOM bas -—" 
) ta 02 & 209 neswied aoftobet DE0Y 
n " tosn-t rte eof d re oe to Snnms 
= 
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| 


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od eageat ftiv tabud sexi? ed os botx9qax 'B5W ae: 
i. 
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yns .vbrte ett al: -efonp\ ~ SI to wptone nots 7 
| a | . 3S : 
esw coLicesy, aida ag soubor sod ya s203R8 stein = s 
- K 
7S . a 
teyistin soimule-y oP esehdyLom-3.sdos: ea baviosde. aye 
‘ ne ay a 
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—e Eg 
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32 


catalyst, and that co. adsorption on this catalyst 
was reversible, which is in contrast to the observa- 
tions on the pure y-alumina catalyst by Chuang et al. 
Cho) "2 


For COS-SO.-H.O reaction Pearson (83) obtained 


pow 
a conversion level of 29% with commercial cobalt- 
molybdenum catalyst and 12% with the active alumina 
catalyst at the reaction temperature of 275°C and 
the gas space velocity of 50,000 hr? in a micro 
reactor. For the low space velocity of around 


400 hr 


, however, the difference in catalyst activity 
could not be detected since the conversion level was 
100% for both catalysts with COS-SO.-H.,O reaction. 


Ausict ~<GOS, = H,0 Reaction 


COS + H,0 —» CO, + HLS (2.5) 


Buchbock (14) studied the reaction of COS in 
aqueous water to find that this reaction was catalysed 
by hydroxyl ion. The proposed mechanism of OH catal- 


ysed reaction was 
OH 


PocereH os OH > He - C = 0 (36) 
Buchbock also found that the reaction closely followed 
the unimolecular reaction mechanism at 15 ~ 40°C with 


the reaction rate constant of I1nk = - 11737/T + 45.66, 

























; 
oa 
; 
: * ” — < wd Ghee a 
} > Bin . bs. 0D 2687 Sas ,Jeyi6I59 
« 
f * i it > =< e _ ; 
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\ —— 
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- - r . ~~ lad 
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> = 


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fos ods ficgiw #81 Bas saytase 5D mune dyiom 

. x = 

oS 4o suytexaqmet nottoses eis te seylsss5 = 


i ~“xd 000,08 to Yilsolsv eosqe esp sft 


“to yiboofev sosaga wot efx tot .xO3069% ~ 
. ~~ ' iIe~ 
ivigos gevisteo ni sonstettth es yaevewod . If 0Od- = 


aew I[6vel notetevnos aft sone betoeteb - ed ton bios 


Noijuser 0,.8-.08-809 To iw eteyissso diod. 10% #00L Som 
| | s ; i. ; 


jms 


noizoBsa OH - 805 Be.f 





gee ge ii a 


, a : Ss 
¥ 4 : 


ct 209 %o noizsest ois ‘bet ute oi) tpodiione 


. 

i 

c 
"> 


aes. 
bseylstso efw notioses ekds ‘gat? bat ‘osvhedew” eve ups “7 


»« fn Ped el Bd ed 


~fsxe9 HO to meinsdoom Britiuteitie, oft sas tyxoxbyd re. 
ie gccke oe mee | aew sobasieox’ ‘beey 

HO. a EP 1 4 po f , =< 
‘ if sek Say sid aes crack 


(98) , 4. eee 
meas 


we, 


Bawol io aren ) motsanen mies 


7 ae 7, 


) Ske sw 


4 
















33 


which corresponded to an activation energy of 23240 
cal/gmole. 

Thompson et al. (98) extensively investigated 
the reaction kinetics between COS and H,O in the 
aqueous solution, in alcohol, and in the gas phase. 
Their reaction rate constant was given by 


kecaned Greta oes 


exp (-22179/R_T) 
for the reaction in aqueous solution. For the gas 
phase reaction the activation energy was found to be 
25720 cal/gmole. 

Namba and shitba (79) studied the kinetics of 


hydrolysis of COS and CS. on the alumina catalyst at 


2 
the temperature range of 220° to 330°C. The rates of 
hydrolysis in both reactions were found to be of first 
order with respect to COS or CS. with activation 
energy of 2.9 kcal/gmole and 9.6 kcal/gmole respect- 
ively. 

A reaction rate proposed by George (37) for 
COS-H.0 reaction over cobalt-molybdate on y-alumina 
catalyst was first order with respect ot COS and zero 
order with respect to H,0, and activation energy was 
12.0 kcal/gmole. According to George's data this 
hydrolysis reaction rate is five times faster than 
the rate of COS-SO, reaction but fifteen times slower 


2 


than the rate of H,S-SO, 


molybdate on y-alumina catalyst. 


reaction on the same cobalt- 


79n9 noOisevisos op od bshnoagact 103 doidw 
-elome\is> 
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f sevio aaw taatanos @sex nottasex +ttsdT 
ee Oe ; 

(TAKOViSs—) qxo "OL. %-20.4 2.2 
it 104 OLsr anOeppS aL mots “ects 20% 
2 bee £ Tens ~4Ievitosn. sd r29601 92EngG 
-Siosp\isos osvds 

: ijeeix ofd botbude (29) sdid2 bas sdmst 
& dgaevylatso.«£ 5 oat no <2? bas 207 to eievlozbyd 
to esa je EE oF OSS, Ro opnes exytsreqme edt 

1 i of, Dauort sx | 


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SE ‘bcp ft: 10 f tsuiszos Bris #2 A 93 
uy sense 
aint eteb 2 ‘ ppreed Ot, patie 
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- ’ s x 


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me: ; 


rar. i - 





- 
ps = 
-_ 
























r 


> a 
Se 


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=a 





George (36) also found that catalyst basicity 
Significantly increased the rate for COS hydrolysis 
reaction. His data showed that the hydrolysis rate 
over the cobalt-molybdate on y-alumina catalyst con- 
taining 3.9% NaOH was 25 times faster than the rate 
over the same catalyst without NaOH loading at 230°C. 
This increased rate was explained in terms of the 
presence of abstractable protons in the reactants. 

Pearson (83) found that the active alumina 
(S-201) was almost two times as active (85% conversion) 
as commercial cobalt-molybdenum catalyst (43% conver- 
sion) for the hydrolysis reaction of COS at the temp- 
erature of 275°C and the gas space velocity of 50,000 


het. 


2.3.5 Sulfur Species Association-Dissociation Reaction 


Experimental investigations on the degree of 
association between sulfur atoms existing in sulfur 
vapor can be traced back well over a century. After 
careful scrutiny of his experimental data, Preuner (88) 
concluded that four species (So, Ser S, and S,) were 
adequate to explain his experimental isotherm at 448°C. 
However, Preuner and Schupp (89) corrected Preuner's 
previous conclusion when they obtained experimental 
data covering the temperature range of 300° to 800°C 


and pressures of 7.5 to 1182 mmHg, which could be 


34 


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fitted with the assumption of three species (Sor 30, 
and So). Again, there was another correction of the 
Preuner and Schupp's data by Braune et al. (12) after 
they measured sulfur vapor density. These latter 
authors argued that Sor S¢ and S, were insufficient to 
explain their experimental results, and they revived 
S, thus returning to Preuner's initial conclusion. 

The heat of dissociation given by Preuner and 


Schupp, and Braune et al., is tabulated in Table l. 


TABLE 1. HEAT OF DISSOCIATION OF SULFUR SPECIES 


AH in cal/gmole 


Reaction 
Preuner & Schupp Braune et al. 
S. > 4 S5 99600 92180 
S¢ jpop S.5 67100 63710 
S, ai 2 S5 - 28400 
3 S. —— S¢ 30500 21700 


Berkowitz and Margquhart (11) confirmed the 


6! S4 and Se with negligible 


at a temperature of about 400°K 


presence of Sor S3, Sey S 
amount of Sg and Sio 
using a mass spectrometric technique. 


Bartlett et al. (7) studied the general kinetics 


of the conversion of S¢ to S¢ and found that commercial 





















sfooge seus to noistqmynen ods Asiw Bbotat2 
too 1zsdtons esw eyeds wnitaspA « .( bas 


wntspi ge . i 


; -f6 39 @mus1tf yd s?sb e'aqurdce? bas asavesd - 


= 
med PY : —p P ames on rs ve «vert Fee . ~e 5? “Sher : , 
ert" vtienes t0odey tefinug Sextgesom yort 
: 7 r —“~ “F = —tc , > | om re 6 
5 ASIEOM 2 ax ~~ he bys - 2 - 4c tsnig Pe DIS orl rci6 
: _ | ‘a 
: ,; a a a ee ae Se = . 
. Yons hig ,ativest Letnemtreqkne wens aisliaqxe 4 


>> Ieitini e'ysitrera ‘od weritnrutss eucdo a2 
: > ~ 


rsnueTt yd wevip acitsfosestS YS seed off 


j = FE [ Te BJ) [it if a oo .* j g% TS eng |e Ef br th sqaqurin] ¢ 
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. : . ; 


mer hai ei OL = S Nelssseak 
te de ise CLI Hoe 2 _xengoa - 


— —— t-e~y ee ee em gm 





e 
0 ; fs @ Hdaee " 
é 
OLTES COLTS 
IOS SVS z : =H 
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‘ a 
ry “Pre { ny ; > 
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t ¥ m ‘ > 
e e 


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- 2 2 oon) ye pe 
eidipiipen dtiw 2. eis e ase 


N°OGOb tuods 





alumina had a powerful catalytic effect on the conver- 
Sion of S¢ to So. Their resulted mechanism was based 
on the polymerization of the Ser followed by depolymer- 


ization of S, which could occur by thermal initiation 


8 
such as the spontaneous ring-opening of S¢- Berkowitz 
and Chupka (10) suggested a mechanism by which the 
catalytic effect takes place, involves the opening of 
the S¢ rings of rhombic sulfur with the formation of 
chains of unknown length. From the data of Preuner 
and Schupp (89), Kelley (54) developed equations for 


calculation of eqilibrium constants at any temperature 


for the following reactions. 


———y 
Sa pee (1.8) 
—_—} 


The equilibrium of the Claus reaction was theo- 
retically investigated by Gamson and Elkins (35) for 
the first time considering the equilibrium distribution 
between the sulfur molecules Sor Ser and S¢ in the 
gaseous phase using the data by Kelley (54). 

Peter and Woy (84) calculated the equilibrium 
constant of the Claus reaction from the known thermo- 
dynamic data such as the standard entropy, enthalpy, 
and heat capacity of each component involved. In their 
calculation they considered S, to be the only sulfur 


specie in the product stream, which did not seem to be 


36 


aah 







4 ‘,” 
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a 
nO tonbio oityletso Ldiwewog « bert eniomis 
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xteds at t sree nea: Lit atu 


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tadi i iups ey 3 pinanasicindtioe (wey 





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# ¢ } 





37 


realistic. 

McGregor (72) and Liu (65) using the data by 
McBride et al. (71) calculated the equilibrium compos- 
itions in Claus reaction system assuming that there 
are only three species of sulfur molecule, Sgr Se and 
So, Over temperature range of 550° to 1600°K. 

Recently Detry et al. (29) obtained heat of 
reaction and entropy data for sulfur species associa- 
tion-dissociation reactions using the electrochemical 
Knudsen-cell with a mass spectrometer over temperature 
range of 200° to 400°C. 

Rau et al. (91) derived a set of equations to 
calculate partial pressures of different sulfur species 
as a function of total pressure and temperature of the 
system by fitting their measured data of sulfur vapor 
density to equilibrium constants calculated from the 
thermodynamic properties given by Detry et al. (29) 
over the temperature range between 823 and 1273°K. In 
their calculation fugacity coefficients were included 
to correct equilibrium constants which described the 


equilibria between S., and all the other sulfur species 


2 
when pressures are high. They found that partial pres- 


6! So and Se 


completely different variation with temperature to that 


sures of S showed, at low temperatures, a 


found experimentally by Detry et al. (29). Bennett and 


Meisen (9) used the thermodynamic data for sulfur 
























= “~ P ¢ rid BaF (ST) - 1p sDoM 
ifivpo: ett betseiuots> (4%) .ls to ebizgoM. ; 


R en se ae eee j +e. 
gyons 3 + anintees meseye gorzoseas eyelid AL, NOLL ~~ 
“ “t = 
re : 2 Mee ee - oie — 
{ts : 4% eipgostont. tus ive’ sO aertoveqe $F tia Yeo 6a 
> § — 2s ele x ST cte-racdmiead ~avo > 
SU Ud Qa uct 10. Sri s 2 Tk PHASES A ify oO eG 
= is 
2 rd : a 
“a or 4 e = “ ata wal UJ 
og (24). «ce FON tieC. . JH S208 
r ~ a “> b+ a” 
iwtine sot sadeab ygotses has moisosex r 


: : ae TF . ie he oP e 
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Se 

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"oe ie | : _ = a < oe eB rere ‘< rs 

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4 b a i ’ a. # 

. y 

— 1 


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{i 


species based on Detry's measurements (29) in their 
calculation of equilibrium compositions in the Claus 
furnace at temperatues as high as 2000°K. Their re- 
sults were generally in good agreement with those of 
McGregor (72) or Gamson and Elkins (35) though the 
conversion level of H,S was lower than that of McGregor 
but higher than Gamson and Elkins. In addition to the 
minimum conversion at approximately 850°K suggested by 
Gamson and Elkins (35), Bennett and Meisen (9) suggest- 
ed that there could be maximum conversion at approxi- 
mately 1700°K. This suggestion was based upon the 
possibility of the existence of HS or SO at the equil- 
ibrium state. The physical reason of this maximum con- 
version was attributed to the fact that elemental sul- 
fur might undergo oxidation at elevated temperatures. 
For all their sophisticated arguments for including all 
the possible species involved in Claus process, Bennett 
and Meisen's results showed much larger discrepancies 
from the actual experimental data obtained in the 
present study than those predicted by McGregor. Regard- 
ing the controversial aspects on the morphology of the 
sulfur species, McGregor's simple approach was employed 
in this study to predict the equilibrium composition. 
~The equilibrium distribution of sulfur species 
between 400°K and 1100°K at the atmospheric pressure is 


shown in figure l computed by McGregor's method. 


38 


: i a? 
ari = hh ¢ 


f 
, one 
_ -~ o~, 7 
"Out > DLA 5 B19 5: ‘Ss Span ror 
Pm ¢ 
| ~—c - Yaw > + frre 
: ; f ~ . at? - L&E = se 3 cot - 


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. 


' 


MOLE FRACTION 


39 





0 200 400 600 800 
TEMPERATURE (°C) 


FIGURE 1: EQUILIBRIUM DISTRIBUTION BETWEEN SULFUR 
SPECIES AT ONE ATMOSPHERE (72) 


£ & 
As 
fo 
- sf 
* 
- a Ot Wea, 7 
> 
: a ,= we eS 
oe sf 
ff * 
: 4 7" 
A a | 
= Z & \ f 
- e 
— # % 
: ‘ f ~~ Ae 3 





4 = f On ; : 
2 S -@ i... j Pa ae es -—-— [ree tong ——— 3 0 
O08 o6e-* | OOB BE 


(2°) SRUTAMSIMST 


(S°). SSNMEBONTA GHO TA GIOage 
1 i s | 


4 


II{IUG VUSSINTAA UOLPUELATEIG MUTRBLIPGOT. +f Seon 


Catalysis By Some Transition Elements: 


High activity of transition metals in the catal- 
ysts has been known for a long time (4) and transition 
metals have been applied as effective catalysts in many 
practical processes (40,87). Nickel or cobalt sulfides 
were found to be active catalysts for the Claus reac- 
tion by Griffith et al. (41), and any readily sulfid- 
able metal was a suitable Claus catalyst according to 
Marsh and Newling (70) as cited by Pilgrim and Ingraham 
(87). 

Haas et al. (44) investigated the activity pat- 
terns in the catalytic reduction of SO, by CO on some 
transition elements on alumina in an integral flow 
reactor. In their studies the pure alumina was found 
to be not so active as the alumina that contained tran- 
Sition metal. They tested the alumina catalyst impreg- 
nated with various kinds of transition metals such as 
titanium, vanadium, chromium, manganese, iron, cobalt, 


nickel and copper to find that chromium and iron on 


alumina were the two best catalysts for SO,-CO reaction. 


The impregnated alumina tablets they used contained be- 
tween 2 and 3 percent transition metals. 
9" COS and CS, with an 


evaporated manganese surface was examined by Goodsel 


The interaction of CO 


and Blyholder (40) using infrared and mass spectral 


techniques. Their observation indicated that the C=S 


40 

















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41 


bond is more readily broken than the C=O bond on the 
Manganese surface resulting in sulfidation of mangan- 


ese. The reaction mechanism was given as 


+6 -6§ 
M, + S=C=O ——> | Sess: Cos2: 0 |——> S + CO (g) 
4 | 
M M 
n - n 


The failure of the manganese surface to adsorb CO 
after exposure to COS was assumed to occur because of 
the adsorbed sulfur. 

Haas and Khalafalla (42,43) found that the add- 
ition of a transition metal to alumina inhibited the 
COS-SO. reaction and alumina alone was the most active 


2 
catalyst for COS-SO, reaction. Khalafalla and Haas 


2 
(56) attributed the decrease in activity of a transi- 
tion metal-alumina catalyst to the partial loss of 
pellet internal surface area and porosity by sulfide 
formation and metal swelling since the molar volume 


of metal sulfide is greater than that of the metal 


alone. 


Performance of a Claus Unit: 


2.5.1 Importance of Combustion Chamber or Burner 


Design 


As environmental protection requirements became 
more strict, sulfur recovery efficiencies greater than 


99 percent are needed. Improper design of a burner may 
















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cause side reactions like carbon formation when proces- 
sing sour gases containing hydrocarbons (33). These 
Side reactions disturb the process in the catalytic 
converter and impair operational safety and optimum 
sulfur recovery. Design of a combustion chamber or a 
burner are, therefore, of utmost importance in obtain- 
ing a high overall plant efficiency. Practical oper- 
ational experience has shown that improper design of a 
combustion chamber on a burner may result in only 15 

to 20 percent conversion where about 70 percent aera 
sion should be obtainable based on the H,S concentration 
in the sour gas. This difference in performance occurs 
partly due to the incomplete mixing of feed gas and 
air, and partly due to formation of carbon soot or 
ammoniacal sulfur compounds when processing sour gases 
containing hydrocarbons or ammonia (33). Maximum 
equilibrium conversion can be attained by mixing the 
required stoichiometric amount of air with the sour 

gas (9) before entering the combustion chamber such 
that the oxygen reacts with HS just behind the burner 
mouth (33). A multiburner system distributed over the 
entire cross section of the combustion eae front 
wall has been shown to be very satisfactory for this 
purpose. 

2" H, and CO in the 
burner is dependent upon the co, and CH, content in the 


The formation of COS, CS 


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43 


sour gas. The major problem with the burner design is 
related to the fact that measurement of the exact 
composition of the gas at the burner outlet is almost 
impossible with suitable measuring methods available 

at such high temperatures (33). For this reason the 
theoretical reaction equilibrium calculation is needed 
as the next best approximation method to predict com- 
bustion chamber performance and furthermore to estimate 


the overall Claus unit performance. 


2.5.2 Performance of a Catalytic Converter 


Catalytic converters of modern Claus type sulfur 
plants are mostly designed for space velocity of 650 
to 900 SCFH of reactant gas mixture per cubic foot of 
catalyst bed volume. The catalyst is contained in 
horizontal drums of about 60 feet long and 13 feet 
in diameter with flow downward through the bed which 
is packed to a depth of 3 to 4 feet. 

The efficiency of sulfur removal in Claus units 
has recently been reported to be improved by reducing 
the reaction temperature below the dew point of sulfur 
vapor (38). However, this low temperature operation 
has a disadvantage of sulfur deposition on the catalyst 
which may necessitate catalyst regeneration. In a con- 
ventional Claus plant, the reaction between H,S and so, 


occurs exothermically below about 800°K, thus being 



















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44 


favored by decreasing the reaction temperature, but 
occurs endothermically above about 800°K, thus being 
favored by increasing the reaction temperature. The 
reaction can be pushed to completion by removal of 
product sulfur from the product mixture. The operating 
conditions for each reactor are normally selected so 
that all sulfur formed by the reaction remains in the 
vapor state, i.e., the reaction mixture is always 
above the sulfur dew point. To obtain higher conver- 
sions, several successive reaction stages are usually 
provided with intermediate condensation and removal 
of sulfur product. The removal of sulfur permits a 
reduction of temperature in successive reactors which 
makes approach to the higher equilibrium conversion: 
level attainable, while still remaining above the 
sulfur dew point. 

As the desired conversion level in Claus reac-~ 
tion increases, the possible sulfur losses due to COS, 
CS, and elemental sulfur vapor play a more important 
role. Carbonyl sulfide and carbon disulfide are form- 
ed in the burner section of the sulfur plant, but can 
be hydrolysed and eventually converted to elemental 
sulfur by applying appropriate controlling conditions 
and catalysts in the catalytic reactor system (38). 
Increasing the temperature of the first reactor of a 


two-reactor system may decrease the loss of sulfur as 














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COS and CS, in the tail gas. However, the higher temp- 


2 
erature may result in less favorable Claus reaction 
equilibrium and tend to increase the loss of sulfur as 
HS and SO, in the tail gas (38). Addition of a third 
reactor may reduce the loss of H5S and SO,, but have 
little effect on COS and CS... 

One of the practical approach to minimize the 
loss of elemental sulfur formed in the converter is to 
Operate the final condenser at low temperature and 


install an efficient mist extractor to reduce entrain- 


ment in the tail gas. 


2.5.3 Claus Reactor Design 


i) Catalytic Reactor Modeling 


Various kinds of sophisticated models have been 
recently proposed in the literature for a heterogeneous 
catalytic reaction, which may largely be divided into 
two main categories; pseudo-homogeneous or heterogen- 
eous models. 

Pseudo-homogeneous models are employed to ex- 
tend the mathematical simplicity which is used for a 
homogeneous reactor modeling. In this kind of model, 
material and energy balance equations are written sep- 
arately for the catalyst particles and the interstitial 


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concept that discrete particles exist within the 
reactor bed. All the particles in the bed are treated 
as a continuous one. 

In the heterogeneous models, a catalytic re- 
actor is treated as a series of small continuous flow 
stirred-tank reactors. 

For catalytic reactor analysis, pseudo-homogen- 
eous models are more commonly employed in one- or two- 
dimensional modes. In this work a one-dimensional 
pseudo-homogeneous model was employed since the Claus 
catalytic converter is essentially an adiabatic oper- 
ation in which radial concentration and temperature 
gradients are negligible. Even in one-dimensional 
pseudo-homogeneous models, two different approaches 
have been considered; a one-phase model and a two-phase 
model. In the one-phase model, the entire reactor is 
treated as a homogeneous empty reactor while in the 
two-phase model, it is assumed that there are two con- 
tinuous phases, solid catalyst and fluid. As mentioned 
above, this continuous two-phase model still neglects 
the particulate aspects of catalyst pellets. 

It has been customary to describe fixed-bed cat- 
alytic reactors in terms of the one-phase model (47,48). 
However, for strong exothermic reaction systems, this 
approach often yields unsatisfactory results because 


of the differences in the temperatures of the solid and 


46 

















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47 


the fluid phases. In these cases, the two-phase 
model may be applied to take the gradients between 
phases into consideration although it involves a 
highly nonlinear set of differential equations. The 
behavior of the two phase model has been an interest- 
ing subject of many publications for a long period of 
time due to its importance in practical applications 


(16,62, 102). 


ii) One-dimensional Pseudo-homogeneous model 


Liu and Amundson (62,63) developed a continuous 
two-phase model in their analysis of a packed bed re- 
actor, in which the complex behavior in the reactor is 
concentrated in two homogeneous phases; in the flowing 
fluid phase and in the fixed catalyst solid phase. 

They improved Barkelew's simple model (5), which neglec- 
ted axial and radial dispersion as well as interphase 
and intraparticle transport resistances, by introducing 
interphase resistance effects. However, Liu and 
Amundson still neglected intraparticle resistances so 
that the reaction was assumed to be controlled complet- 
ely by interphase effects. They also assumed a uniform 
velocity profile over the cross-section of the bed and 
neglected the effects of length of the reactor bed and 
temperature on the velocity profile. The major purpose 


of their model was to check the existence of multiple 























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48 


steady states in the catalytic bed depending upon 
the state of the individual catalyst particles. In 
their investigation the effects of the inlet gas temp- 
erature and the inlet partial pressure of reactants 
upon the concentration and temperature profiles were 
also examined, and the stability of the adiabatic pack- 
ed bed reactor was found to be dependent upon the 
existence of multiple steady-states for single part- 
icles; if every particle along the bed axis had only 
one steady state, then the reactor would be stable, 
and unique concentration and temperature profiles 
would result for all initial particle temperatures and 
concentrations. On the other hand if a single particle 
in the reactor has multiple steady states, then, from 
continuity of the mathematical model, adjacent partic- 
les would undoubtedly exhibit multiple steady states. 

Hlavacek et al. (49) investigated the effect of 
Peclet number for mass and heat transfer in the adia- 
batic tubular reactor using a dispersion one-phase 
model. Their results showed that higher ratio of mass 
transfer Peclet number to heat transfer Peclet number 
enlarged the domain of multiple solutions. 

The effects of axial dispersion were checked by 
Liu and Amundson (64) to find that in the multiple 
steady state the profiles of temperature and concentra- 


tion were appreciably more sensitive to axial dispersion 


° ? > _— 
m 4 3 ats i a! 4 t oy +8 5 - 
" = 
: 4 ~ 5 mio vr 
. « 
a > it tye those rt © ny TOC) 













{+ » iy ~ —~ as oeis 
Pit we — 
jasbasasb sda’ of bute? sav 107 78S" bed be 
z teta-yl oligiat Oo eseneJ2ixe 
- r 
: . & - ™ 
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° 
ia 3 p oe ! t 2 : 2 = ate eS be Pe Si. - 
~~ P A . b 
es{iitox names bas Aokttsezteeonol Supsi iy 1S 
Ps aes ; Terre 
> JE£ I ey! lina <r 3 i f 2 tO r 4 Uses 5 inaw ” 
? - ¢ - a ake ie; ae = 
sittea sipate & wet xStito eds 00 .enotssusnepae> ~ 
. . r ai - rT ~ af ; J 
owl +2 whsoete algisium esd srososex ods at 
| 54 «fehon Laaidemedsan ent do yttumtgnoa |) 
a r3 me | at ee? nes5 - ‘Sor ) fr k 23 rf TSRens Bw en3 - , ~ Ad hd 
\ - +, « Ph 
" P * . oat oolerrebnd ; arent oe - 
¢ : t3fum. tididxs betducbau biuow eel 
Z f q 
4 J 


a) 
SO 


Pras 5p tdzovriz (@6) ts a sissovelit 


“alt ct tetencit. jest bas Beem ‘aot seduun teloed a 
. fe 

sredc-ano noreiteqetb..s. piiter rodeos andudu: ofjsd 
: to citsx xanpid) todd Rowode etivess xed '. LeBom “(atg 


: ‘ . r. 
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meer b ents. s bepzaing 
ne. 






_ ’ iui. Sd 
























7 - 2 1 oe ed +s a” ‘ : 
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¥a z ba? ea - ae} bag aN 2 e as nal 4 


bier ~ 





na © 


PUES TL . Bas. aed 





49 


than in the single steady state. Eigenberger (30) ex- 
tended the model by Liu and Amundson (62) by including 
the effect of effective heat conduction term in the 
catalyst phase heat balance equation. His computations 
indicated that the temperature maximum could move to 
the front of the reactor due to the backward conduction 
of heat; compared to the case without effective heat 
conduction in the catalyst phase. 

Votruba et al. (102) used a piston flow model 
of a tubular adiabatic fixed bed reactor with external 
heat and mass transfer for description of temperature 
and concentration profiles along the reactor using the 
two-phase model developed by Liu and Anundson (62). 
They neglected axial dispersion and conduction through 
catalyst pellets as well as resistances within the 
porous catalyst structure. They only considered the 
convective mass and heat transfer in the axial direc- 
tion and external heat and mass transfer on the catal- 
yst surface as transport mechanism within the bed. 
Their computed results revealed that there could be 
multiple steady states within the reactor due to the 
external heat and mass transfer resistances over some 
range of parameter values. 

Eigenberger (31) also investigated the effect 
of different boundary conditions in the frontal surface 


of the catalyst bed using the two-phase model taking 


Ste 


4 > ° ‘4 re 7 Al a> w@ od 
» 2s Onsen ‘A522. VOERSIa -< lonie eno at nad 
¢ 
r & mn ut aly Cus fs or + 
(ca Gt my f uid va j owe mst DSeEDABS 
¢ ~~ 2 ; ‘mor :- { <) | >" Sha ye > > « ra) 259 its ans 
: ‘ 4 el - 7 r - . r < 
.MOLTEIIpPS ~St (ad teed eesdq FeyisI&6o 
2 - = 
. 
ct LL HEL SITUS Neon ong sts bots fyrts 
| 
7 : ‘ . © 2 eever > ; 
oad ori 3+ sob zotpdex ers to Snort ons 
_—— 4 4 
+259 Do s W407 betmastoo ;J&en +10 
y , oe ’ 
; iq teyistso oft ai gsottoubaoo é 
c 4 e = ~ fora Fea 
ists pee aS E \ s VF : ee AD ae 2£! 2. a 
2 at Be d Ay «aoe 
t ! set Hse oxt itedsibs tel s 10 
. i a 
— y 
atratt *-' is ~& 2. ae 
= JOLT SS My io totentstd SFesm DAS 7697 . 
« 


not as wy QO: v Baieh adT 


fuon eiad# s6Ae bs 


a4 “ 


. 
i 


" 

t2atszt teed brs. #eeam evisoevnoo _ . 
bite ged font9dne bas mols. 

; = 

ain sroqenes?: <6 gostise/ ey : 


isover 2 Luger 
Bex offs ieee, eedis? 2 vbsesre sigid ium i 


2ndnasedeex enbemie: aes 






















yn eesric-oOws = 
bot oeiper yout 
taylsse> 


,otatourme 3 dete te> 2yoToq 


besuqioo stoxT — gt. 
ae, Pa 


= 


em bn saod Eacx93x9 


panes 20 epnss “a 

’ 2% =" ¢ ee r=" ( 

Pa 2 aaa adn ~. ae < 

» 2iapuaeals 
bint tales 


=o 





wf * 





heat conduction into account. His boundary condition 


for the front end of the catalyst bed was 


eed heh orkan¢@,0f axial. (iielzadoh bso) 
e (02 7=0 R s °Z=0 2 


According to Eigenberger's results the usual assumption 
of an adiabatic termination of the front of the catalyst 
phase had to be revised in cases where a high temper- 
ature excess of the catalyst phase can occur at the 
beginning of the reactor. His reasoning was that a 
certain amount of heat would leave the front end of 
the catalyst bed by radiation if the catalyst temper- 
ature at the entrance exceeds the fluid temperature 
considerably. Eigenberger also noted that the igni- 
tion zone, defined as the region where significant 
reaction takes place on the catalyst surface for the 
first time in the bed, was fixed at the entrance of 
the reactor when the fluid velocity is low while the 
ignition ante was located in the middle of the bed 
separated from the entrance when the fluid velocity is 
high. 

Another interesting result shown by Eigenberger 
was that only one steady state existed with very small 
fluid velocities or very high fluid velocities. The 


moving velocity of the creeping profile of the 


50 


















tifnoo vneBnagod 2ift  tapooos osmt noitosbnco ter, 


. . a -_. ry 
tepvls¢so edt Yo Bre snort eds Tor 


Baw Deg 4s 
7 i" 
- mes \ , - Iny b << 
Val Fea kt h ; : F ’ 
ae - . 4 “ — 
ea i a — a 7 ee. | ee ‘ oD = : ar) ad = 
> , . oe 3 P| A rm | we - 1 
3 ; 
2 
:, of os ~~ . b iver 
; 3 y afd ediveest 2 *1eptedaspia OF PNLDZOOs 
“a 
+ - - i £ a ’ wv fo - _— al . » 
sft to teow? sid To sme LIsmemyedt COLT SBaSivs A + 
7 ~ . “ sal a ati vr 
Tr 1afiw 2oeso at Beetven sc ot Bad sesng 
: ‘ : n ¥ es anee + @ 
9 ; 270 8o seedq teytstso end te eBaeoxe SstWIs. 
a et ad ru. eS Pee 
’ 25W DMInCeseat ein -tos0Ae1 ens 10 BAe nnoiped 


ont ofd eveatl Pficow tsar to tnsoms nissieo ~ 


™ 


rosames tayistso sdv if sotietbsea yd bed tayTadso edi. 


iD 
» = 
4 
+ 


wostoanes hivl> oft ebesoxe esomeasne eds ts eints 
-tipt sittt tact Beton es 


fs ~eprednmeets .yidsrebia ado 


Jnpoliinapte oraiw. robper efit as Baniteb ,sn02 noks 


2 Se 
it +o% sostuse tay ts step ont mo evela ‘podiad aoijosex 
to sorexztine od? ts band enw bed oats ai ‘emis taxi. =. 


eit olisw wol at vi fodaw bien 2 edt new 1032891 ‘odd 


dd Yo Slpbim orig ‘ak bessnot eew. enot sotsimpk 


ww 


: eh a 
zt yttoolev Sint? ent neitw, conexiae edd Kori bessrsgee 


7 
; ” ’ . Z i 
— ; oe hig Sadat bs awe . an 
hl * Aes sy £- ?, ay e ‘an 
P4 My hy an ; Sg per ee 
P ® u 4 


a 
1s redaseid. yd awe soon eetvamanont Tons 


Hie sei aie Bae : 


alt 





a 5 


I 


conversion was significantly affected by the fluid 
velocity but almost negligibly influenced by the dif- 
ferent kind of model applied. 

The importance of axial dispersion was consid- 
ered by Karanth and Hughes (52), who recommended the 
criterion that there is no significant axial disper- 
Sion provided the bed depth be greater than 50 catalyst 


pellet diameters. 


iii) External Transport Resistances 


The concentration and temperature difference 
between the bulk fluid and the catalyst surface is 
dependent upon the mass - and heat-transfer coefficient 
between the two phases, the reaction rate constant, and 
the heat of reaction. The reaction rate constant and 
the heat of reaction have specific values corresponding 
to each specific reaction system. However, the values 
of the mass- or heat-transfer coefficients depend sole- 
ly upon the kind of fluid and the flow pattern in the 
fluid phase near the catalyst surface. 

Average transport coefficients between the bulk 
gas stream and the solid particle surface have been 
investigated by many researchers (13,34,68). Fried- 
lander's approach (34) was to get mass transfer coef- 
ficients from the boundary layer analysis around a 


Single solid spherical particle and obtained, 



















; h siosit rwspitiapia ; moraxevnoz 
‘ 7 i < 4 . ee ect 7 4 “c) Sd 
: , e/ + Te iN hase ¥ 
e. ™ ec. rf Py i 
qa Isbom bait toox82 
= <_ 
) Sw fF : ; XS >» S50nbBTTOUdaL & {T f 
; $ 
, ad a4 4 4 - 
ow . (82) ear hos dsietet yd hexe 
Ve 
sane ‘ y dees om, t *eitve, 
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re + ‘+ y~echestp od fitaqeab bad art hebivord mote 


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4 Tend - . is ? 
hw Iececeocivastpreppech ateapsiineotiiians canlineiiin 
. oa 
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Te 7G” adt+ Kee RF rc¥> St tastt ec iit cic aw ed 
TeaVvi Th. -OFis MB DEBLET AsLUG On neaew a 
™~., = 
a) a ee ‘ 7 ee es iad = yo - Se. 
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bos Jaustenoo stb1 soljosex eft .metipssx to tsed oft ©. 


ce Ne 


fousexien ssulsy sitiosqa sved mObJoae% To seen ods 


sgsisyv sfij .YevewoH ‘.messyve soltoget oittoe 2 fiose of x= 
. \ - Pn!) 


s breasb atasinitisos x4 olensid~sead 20 “23am odd 30° 
oft ai azetsed wolf? eft Bite ‘biul? to batt ods coh an yf 


oDéiazue 3 eyleteo ads xp9n oeerig biuts 


Alvd sit neawied 33 nip ite. rsagaa eQeisvé 


: ‘ 
7 won See ws” “7 - 


4 “ 
7 
r 


i 


fiesd even eosiiue ofait 164 bnew ats bas- maorte 26 


Sie eee x : 
amon ne 


- mbeixt APE eerie 





ae pteeo gotegest ee 2om 398 A aad a. 
- er Bal ore % : ia Meier i " a. it % 
: ey ‘ 7 2. ‘ faan 36 oe _ 7 1¢ ‘. ie ~ a+ 


wil, > 













- 






s B 


1/3 R L/2 


. (2.8) 


Sy = 0.89 So 


Froessling's correlation (8) used the method of 


dimensional analysis to correlate the mass transfer 


coefficient to the other important dimensionless 


groups to get 
Se: 2°* 0.6 eh 288 


h R 


. (2. 9) 
For a fixed-bed reactor, Chilton and Colburn 


(17) developed the correlation for the mass transfer 


coefficient, 


ee m Pf 3 2/3 
D G Cc 





eg (RA) (2210) 


and by applying the analogy between mass- and heat- 


transfer phenomena the correlation for the heat trans- 


fer coefficient, showed 


5 aittn 2/3 
cee et os 
p 


= f (RA) eon a 
De Acetis and Thodos (27) have summarized the 
data available up to 1960 to get the relationship be- 


tween Ip and Ra as well as between Jay and Rai 


Riga stool oS 

S 
1.10 
yee (2533) 
H Rephes as 


In this study De Acetis and Thodos' equations 


have been employed to estimate the external heat and 


mass transfer coefficients. 









¢ ey 
i t * 
4 —_ ~ a r 
- ‘<> v 4 - ue 
? 9 ae 
’ ) (8). noisélerne> 2‘ pnilesear . 
= t 26M + otsisrr0e> of eteyiene fenotanomib . 
4 Paella 
erem 16 +xOot si4o0 5ds.oc Jnstosri29809 
fen of .’2aquotp 
c.\ £ 
e \ a . ~ sy * 
St , 2 rei ) f ~ = e 
) pS? | {i 
4 i - fr ~ Aci 7 avi é rc 4 
: fit. fs / + i> rch Doan a = ~ - 
(7 + OW @ oy tO # eit : c i e ‘ey ’ (\ fr) 
, r > « 
_insiLoLtisos 
- xf “ 
4 #t 
‘nt : ; 7) +f OT ee —— | r “~ 
yf. (4) 2:= e 5 at 
-t&e IH -Ream reewted yoofens oft paiviqas yd bas 
i ete rot norseistr1os oft sasmonesrid r1siemsys . 
a . < J ~ = z 
; oe i = 
bowode ,taetoihises 29% 
; ES. = 
aX, 4 cal . 4 . 
fr ot . EN ~q ak) “. 
~ ~~ * WT... - 
; 7 i~ aa aA 
1G . * > i 
BS mia saved (TS) é@eboaT Ors erzooA oo ar - 
=a ; baa. 
r .oOr : er LY 
-ad gi + sid tap of O01 o¢ qu Sidsfteve 6305 <- 
: | ~ ie Pd alt oe C PF ayy brs '- fi 
: 2.608 of meswitsd as itew e6 .2 bras _f 
~ Oe =| a Pita: Eas * 
} 
; ¢ 5 
i i , ? 


iisups ‘sobost? 


os 


s A | . ce , > 
Sis siteoh.© 
tS. ic Fe 















: : » are ’ 
» but 







de 3 usin, beet 


38 


iv). Internal Transport Resistances . 


In order to obtain a reliable simulation of a 
fixed bed catalytic reactor performance, it is often 
necessary to consider the effect of the interaction 
between diffusion and reaction within a catalyst 
pellet as well as the effect of the external trans- 
port limitations between the fluid stream and the 
catalyst outside surface, particularly for highly 
exothermic catalytic reactions. The overall influence 
of these physical processes on the reaction rate is 
conveniently expressed by use of an effectiveness 
factor, which is defined by the ratio of the actual 
reaction rate to that which would be obtained were 
there no diffusion limitations. The determination of 
the effectiveness factor usually requires numerical 
integration of a rather complicated system of non- 
linear two-point boundary value ordinary differential 
equations. Numerous attempts have been made to simp- 
lify the equations into a more tractable form and in 
some instances semi-analytical solutions have been 
found. 

For a spherical catalyst pellet the steady 
state material and energy balance equations may be 


written as 


























- gevns tetas +sogeniEst Lestrygs at (wi 


noise fumbe afdsi fos 6 miszdo oF 19540 nt 
getio at +i ,sonmemmotie”d ros LOY a lal ped baxi2 jot 
tinsisiHs ont %0°3 fostie 444 “eUNARSS' og yiseseoos 
tJovipaso 5 nbddivw noizoset BAS aoiaytthb igewsed 
=2n6u3 rontatite iedt Jo s30t%e emt 26 A LOR 25 teileg 
eit bas ado Se bavty ods reswsed enoks sd Smit FOS 
(roid rot yisetupriwss ooptae eabLatuo savindso 


soneulin: Lisxseve oAT ene Srtylsseo “> intrerttoxs 


»} etsy folsosei shi-—n0- e98ee99034 fsolevita seort to 


2econev 2 itost?s me to sau vd chain Ci yitasinsvaos : 
feusos rid ‘Yo ofiet ois Yd ponkteb ai dpidw sodon? ‘2 
ox ow berintde ad Hivow donde Sets of ose moisonen i 
to notisoionxetsh aft Lenoisestmit nobeut ALES on oxen } 
isoixonmin estidnpet yiteues xoso87 eustiovtsoetie odd ; 
~on to meteye peseoilgnon + todyex 8190 noise poset 7 
N° 


ieisnewettee yrsn tore onisy yrsbavod, saboq2aws- 1e8nkt 
~qmLe of Spon aged ach atqusets euoxemay -enoisaliee:: 
ag bas mrot eldstu413 stom & osat ‘endisenp ong au 


need avail srottulor faptsyens~tmoe aodansent amos al 


Hac 2 dc 
De aes + dD, cs) a ao to = 0 (2718) 
1 fy 
ar 5 ar 
Reeser eae) ee oF (AH) or = 8 (2.15) 
e ar anor’ gr p 


Converting the above two equations to dimension- 
less form by introducing the relationships, 


Cc 





foe 
%=G 
Ss 
“p 
8=z 
Ss 
4 seks 
owe 
Pp 
5 2 rs (Pts) 
isla ( ) 
P Fe Cs 
neec 
= 
gives finally, 
2 Pim? 
z +2 WN - 4? omevie ps oar (2.16) 
x x 
dx r. (P. T,) 


Be ae is 
Jb c 6 4 
’ ~~ ° > oo Eo F oa — q : 
- &T.S) Q * q* 26 (3) a! 5d “ae af 
nr, Wh 5 "i os 
: : 
4 = ; a }, aS .— + nH 
4) (Hi + 53 =) at 3 “ a 







“nolanemibh o2 snot nmpe ows eveds ais paeersvoD 


LS 


.eqidenoisetor ait pods bors ns ve orxot aeol 


X 
= 
r 


4 i 4 


fae oS . 
- he be: a 


ia ay we ad | 


9 
: pl 
aera 


1225 db 
2 Caw = 


fae + = S84 26° Ene (2.17) 
ax x dx a a ee 


The boundary conditions for the above equations 
depend upon whether external transport resistances are 
included or neglected. When the latter case is assum- 


ed the boundary condition becomes the Dirichlet type 


xs aye de _ 
= Aa ax : Tac? 0 

Peas) 
x=1, 3 Shp. parag amp 


Should the external transport resistance be 
taken into account, the boundary condition is of the 
Neumann type due to the finite rates of convective 


mass and heat transfer on the surface of the catalyst 


pellet. 
% gy. oh 
Xo=oQa,; soli. thes—-omp? 
iis oe tie Ye i ane en 
Seas Y Sold S, dx’ ere N. dx 
u 


rate Sain Ke (2.40.9) 


For the Dirichlet boundary conditions, Weisz 
and Hicks (106) developed a simple transformation 
method to convert the problem into an initial value 
one which enabled a rapid evaluation of the effective- 
ness factor-Thiele modulus curve. However, they 


employed the Adams-Molton iteration technique without 


if 
(e 























igs ae oe rine 
vf .9) 6 =f peadot 3 ) a8 » 28 $4 2 
(1 a ede “OX. BD 


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1s esonsieiezez Progans7z2 tant9dxe | yodserve nogu hasqeb . 
ors 
“mBes BL 29869 192 ‘tel eda nen adh tech ele 10 bebuLfont ‘ 
sqvi steidorrka ads aamooed naksiiees yasbred ent be 
ab yb: é 
eye se aes are 
(Sf .5) te 
f.7. @ , iL 2 ee 
ed gonsteizes sxOgenst2 isaxadxe edt Biluode 
aig to ef costs sbnseS es ods ,3nuooos otdt nedet is 
svistosvaes t0 notet od bt? wis os eub saqyt nnepEen 5 
eyisiao gaz io eosiisge odd ao jetensxd json bas 2anm 
-toilisg 
6b yb . 
0 / 
os , Oe 10 =x a 
ab n we I 7 
owe > — as = = 
xB 7 ae A? f ¥ i * * > 
; \ 
(Ok, . eres &S : : apg eh ; ~~ 


seiow  Snoksti bags yaebauod joldoixka ad? x07 
coitamotensts ofqute & boqoleveb (a0) Agoin bas 
su Ley teisini ae ogni motdoxg ants sxetn0 po Prion toe 






 peebsoedte - old 20 nods es 


A | 









, j Tre Ps : 
Ee verks _.20ve ee es ee iui te. tT 1s aes 
oo Pc 7 me hd ae,, eg}. > k "2 a a +¢ . ion : ; ’ ’ arith Nei ee ll f 


Ge og 


56 


checking the accuracy at the open end of the integra- 
tion path. 

Varma and Amundson (100) applied the maximal 
and minimal principle to compute the bounds of the 
effectiveness factor for the infinite slab geometry 
with Dirichlet boundary conditions. 

For Neumann boundary conditions, McGuire and 
Lapidus (73) solved the problem using the Crank- 
Nicolson implicit method. However, the difficulties 
Still remained with the large amount of computing time 
though the accuracy was satisfactory from the engin- 
eering point-of-view. 

For the purpose of improving the conventional 
finite difference method in solving the problem related 
to diffusion with chemical reaction, an orthogonal col- 
location method has recently been developed by Villad- 
sen and Stewart (101). They employed the interior 
collocation method in which orthogonality conditions 
were applied to select the optimum collocation points. 
The accuracy was found to be comparable to that of 
conventional methods but with a much simpler computa- 
tion procedure. 

For highly active catalysts, Petersen (85) devel- 
Oped an asymtotic method to solve the non-isothermal 
Dirichlet problem. The basic idea of this method was 


based upon the assumption that when the catalyst is 


vy yi j 7 ‘ vec he tiai, His emrev 
















be pas - e = 
: : z : oti PLIL Aa J 
; + f + c yi. --0 «a : Q 
= J © O77! Vu Lu 
bric IOS LAS AJ 2 iw 
hes AvOd ansmrset 2049 
} 4 ot: qT -Si3 Ds rios e \ ) SHpLo he rq 
; c ni moe he ak! 
z G1 s Hid iV LN wv 
‘ i uC f sent sms2 lisse x 
f ? : ' a S M. 
Pe wii 4 
/ : - E is a |ri3 ipyo 3 
. — * . rn + cr 
BLY niog prires : 
$3 3 ) nity Ra j aig 304 Al 
a7 h ; ; st 
as es staat 7 
es il » 2 Ssofersii Lb Sr Lats “oy 
‘ - a P 7 iat 3%, : : > - 
» ‘é TrOLInss? soimnshdo dzeiw not Shag EB oo? ; 
+ 4 au? ? en el . Oe ow i 
. a - x ft. as . us = t i“ 
' d peaolsveb aeed ylineoot ead bonds fokispek ~ | 
DOaL IVS ft =. 9 ae 
“ ; 
. = a ’ ‘7 
i ~woLlams. % PY  «tBO 4 x EWOTS bits hak U 


to doidw ai borisem nM ots sboLloo 


ao 
_atniog notsscailon mumigao afta dosiae: oF peifags qusw 
to t6n3 of sildayvedmos 6a o7 pest Saw Ysetbose’ sar < 

wi, bi oes 
stugmoo tefqmis doum-e ddiw gud “Sbortom LEAORINAVAGD 





* ieee) pad 
~feveb (t8) tense sud vaseyistead svitos whips aot ye ; 
‘ mene te k—n0K ° i oytoe - Boul? ‘om otsoamyee ae . 
. it sa 
-portom a they ‘otaed. oe mbox te 
* av eis -_ oh: doa Per Peerage) 


highly reactive, most of the reaction would be confined 
to a thin layer near the external surface of the cat- 
alyst pellet. The conservation equations, therefore, 
May be approximated by those for a slab, whatever the 
Original pellet geometry may be, and the calculation 

of the effectiveness factor becomes relatively easy. 
But this approximation method is not applicable in 
general at low values of the Thiele modulus, usually 
below 2 (52). 

To improve Petersen's approximation method, 
Paterson and Cresswell (82) developed the effective 
reaction zone method. In their proposal, it was 
assumed that at some point within the catalyst pellet 
the concentration of a reactant drops to zero value. 

A parabolic trial function was found to be quite ade- 
quate for engineering design purposes and the computa- 
tional effort involved was very small by using the 
collocation technique. The major problem involved in 
this concept was how the appropriate optimum colloca- 
tion point can be chosen. 

Quite recently Van Den Bosch and Padmanabhan 
(99) examined the efficiencies of different inde of 
collocation methods in their ability to predict the 
effectiveness factor for the Neumann type boundary con- 


dition. They recommended the collocation point of e 


for the high reactivity model instead of = proposed 


57 


to imkog. 






pet ° 
i 


a? ab 


Colom ere ae 


‘sie fe 
und 6 HN Rh 


at 


ED 


a 


; gy 


my 
- 


‘ed nsV visnanss- atic ; di 


| 
‘ 
™ 
oh, nn 
As 
e 


A 
2H 


yaibide alia rai i 


nn smielt ‘ont. 203. xo 


to sebm ,svitosper yidesa ae 
“ y i 
ntdd #° oF 5 


6 Mose .os4 i i rel LPLLoO 


! : . 7 -~-' f. ty, 
innt. eebdoeavitesIzSs Sad. 7to 


-1980ND od, FBO ‘antod wake - 





















igjD £ Tt ,*teiisa savis : 


Li SWebaot. Bre noe tags 
-bofitem snos. neLsosed ea 


Ti 


v” 


9moe t5 tsadt Domuees 


al 


s to séitsutasonoep ame |= oy 


2 ee 
ijonv2? fstts clklodsraq-A, 7 
; iy + a 4 


t 
ri1ssnipnas 102, ogemp 


Sendeiyas +1¢ otis: ‘Isnoat ~<a 


. sypih | doss nottsno£ too 


; aS ate 


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fe 


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Paottso0ttos | 
ede eorest 


owe nee 






re V7 a 


by Villadsen and Stewart (101). It was found that the 
accuracy of the orthogonal collocation method could be 
improved by changing the optimum collocation point 


from = to #2 for the high reactivity model proposed 


V2 


by Paterson and Cresswell (82). 

The effectiveness factor in the transition 
regime between the Knudsen and the bulk molecular 
diffusion in the micropore within the catalyst pellet 
was treated by Abramov (2,3) and Abed et al. (1) with 
consideration of the Poiseuille flow in the pores due 
to the change of the number of moles of the gas along 


the pore length. 


v) Effective Diffusivity 


To obtain a poecacanie estimation of the effec- 
tiveness factor it is necessary to use accurate values 
of the effective diffusivity for the reactant concern- 
ed aswell as the physical properties of the catalyst 
pellet. Satterfield (93) collected a considerable 
amount of experimental data on effective diffusivities 
and also proposed estimation methods for the situation 
where experimental data are not available. Since ex- 
perimental data are still not sufficient for different 
reaction systems with different kinds of catalysts, it 
is frequently required to estimate an effective dif- 


fusivity from fundamental data on the reactant and the 


58 



























eit tedt baue? eawot . (£01) oxewase bas npebsif{iv yd 
ad binen bortem doisesalfloo Lencepotisie ef To yosasooR, 
triogag nobsevdfien munitqo snd eaipnatio yd bevozgm.s 


rs 


beeogoty {shor ytivitose: Apid ont 262 a ot =, pout 


. ($8) (iowees39 bas noexeted yd .- a 
noittens3d sft mi u635er suenevisoetts oft 
xsfunelom ted eft Bas aeehinx edd asowssd emipet 

jeatleq teyistso odd nidgiw sxvogowsin sds ad noLeesPEs 
djtw (1) .fe 39 BedA Sas (£,8) vortse1dA ye boesser1s enw aay 
sub astog en ni wolt of fisgqaiod sii to noid srebienoes q 

profs asp eft to seiom Io somes eis to eps ety oF 


.ftpnet srog ord 


ysivievId29 svisoe3aa (v 


-ostits oft Io doitstizes eidsnoesss 6 nistde of 
soulev sfsut0o6 Sey oF yxsaeecen at #4 rotost ssansyls 
“f190n0o InstIoset. oy rot yeivieuttib ovtioets ors to 
evistso sft To aakisemeens Isoleydq. ody es Iiswes hs -— 4 
aidstehienos s betueliies (&e) biettie336e 460 feg 
esitiviaptifh evicostts me 6386 {ng nomiteqxe 20 Javoms a 
noitsutie offs 192 ahodse noi somiszes: baeoqorg osls bas 
“xe sonia ldsiisve $on: ous stab tnanomisoaxe oxedw 
jngvettib a? saabaheRiee : Jon Beenrony tetceizeg 


bt ary. WEES fue, 






. re bee! 77. : ] 
Srae ey mi dacup > o> 


late, ee 


A 





aye! 
ae 


& 


catalyst pellet in the absence of the actual data. 

To explain the complexity of the pore geometry 
of the catalyst voids in evaluating the effective 
diffusivity, some models have been proposed. Wheeler 
(107) proposed a parallel pore model to represent the 
monodispersed pore size distribution in a catalyst 
pellet. In his model, the effective diffusivity was 


described by 

€ 
D = (=) D (2 204 
where 


my sandomly objented vy (2.21) 
(1-ay,)/Dan + 1/Dyy 


Np 
qe (2522) 
A 
a 8 R_Ty\35 
Din = 3 cael | Ry b2ine.s) 


According to Satterfield (93) the tortuosity 
factor, tT, can be approximated as 4 for non-surface 
diffusion catalyst pellets. Here the tortuosity 
factor was defined by the ratio of the actual diffus- 
ion path to the average pore length. 

Hideo Teshima (46) obtained experimental data 


on the effective diffusivity of carbon dioxide in por- 


ous chromia-alumina catalyst with average pore diameter 


° 
of 50 A and pellet porosity of 0.41 using an isotopic 


59 





















ish Isutos- ott to sonseds. oat ak fotieq sayteses 
anit eteq sft io ysixeiqmes si? pisigze. of ates: \ a 
ovisostis oft, pnigsuléve ah ebiov steyietseS ent 20: ; ee 
sise;4 .heeogorc MSc eavgd. efebor-eoe oivierti ib: 
id tnsee1gex ot ishom s10y labieten & bheeoqgo7zq: (TOL) 
tavistse> sg aAL- sonswarsaeLb os ke e109 baciscaibonom 
sew yiivieutith p¥itostis eng .febom eitt at .sellegq 


vd besdizoaeb 


se s19niw 








$. 9) yee ween! - = g ec 
ayes . gad’ (gxe-t) a5 
we io | 
(a8.'8 — + {[ =p 
A 
ey TR B ¢ ; 
(E448) Z . . . gt t ~ ax® S- 


Yiivousxes ont (€@) pisidsgasee 08 paibsoopA . | a 
sistivesnon vot § 26 bednmxorggs “Se Hike T EEN 
vt Leoudtoy -edd ozS8 saseliog: tayletso. nokevtib “s 
~eultib Lsutos sii to ove oth Re -amiNb om rosost 


.dipast etrog 


ee 





fet wsog si ca 





+ uA Oe 


ene Pree eee" ana 


s, i 


60 


exchange reaction method. His experimental values 
were in agreement with the parallel pore model when 
the tortuosity factor of 3.4 was employed. 

Since information about the surface diffusion 
effect on the catalyst pellet is still limited due to 
inadequate experimental and theoretical aspects of 
surface transport phenomena (94), the value of 4.0 
for the tortuosity factor for the Alon catalyst pellet 
was applied in this study. 

Johnson and Stewart (51) exploited a model in 
which a porous material could be represented as a 
bundle of randomly oriented cylindrical capillaries 
with different radii. In this model the effective 


diffusivity was described as 


D, = KfD do(r) (2.24) 


where 


ses feos? 0 dv(0) 
fav (9) 


(225) 


The problem in the Johnson and Stewart model is related 


to experimental data for the geometric constant, kK, as 
defined by equation (2.21) 

Upper and lower bounds for an effective diffus- 
ivity of a binary diffusion system were calculated by 
Petty (86) but the need for experimental determination 
of the geometric constant, kK, still could not be elim- 


inated. The importance of the choice of statistics 


D8 


wabitiow (atasmiuegqxs 2h -boritem moLtonet opneiaxe 
aeftw febom ereq ielinxsq edd dpiw onemeaszys nl erow 
bevolqns caw b.E To sotost vrbeous10% sits 

nolLeetiib oostaise sds, dvodse rot temtotms eonte 
64 evb bodimil {Lite Bt solieg teyfsiss ens ne soe tie 
to adnedes Iehésetosd? Bas. [stnemiteqxs essnpebsat 
wlev fo. (he) ememoneriq srogensss sostiwe 
toifeq sayietéo wolA ofa wo’ nososT YIieous10g atid 102 
-ybute sift at beilqgs a6w 


at [ebom «6 bedioelaxe Uf2) Ftswede bas aoendot 
ce bhotnereaqet ad blued {sitet 5m eué104 s Avidw 
geixslliqs» Isoitbakiyo besneire yimoebnas1 30 edad 
esvisosita ssid L[sbom abas nl .ithsx tnsretiib ddviw 


2s bediroaeb ‘enw vtiviewtatb 


(dS.&) (x) 0b. al + * 0 
s1erw 
(fe ae 
(e)vb 8 “e0ol | 
ee (8)ybl . - 


betslax at Lebom s1swede base noandot of3 ni me idoxq art 


LL 


as «3 ‘snedanos ofxsomosp edd tot ‘sdeb Isdnomlteqxe os 
(t.8) “nol sauge ud a 

~au2? ib evisostis as xo? abauodl towol baw seqqu ~ 

‘e boas ivosss anew mopaye noiewtd Ppa ‘bo este 


o aot basa ot Hot =a 


Re us ia a Te my " lt it aa ibe A 42 res ile x ce 
.. leis Bedsits Perey ob yi rae rie ala a mye 


2 Be ‘7 , ae Bee 
a , ; is 






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61 


which best characterizes the pore size distribution 
function o(r) was emphasized by Petty (86). 

Wakao and Smith (103,104) proposed a random 
pore model for a bidispersed catalyst pellet to get 
an expression for the effective diffusivity; 

e 7 (1+3e,) 

Dy en a a ae th 


e = Duty Sidi eEgD (2.26) 


Wheeler's parallel pore model has been applied to 
estimate the effective diffusivity in the Alon catalyst 


for this study. 


vi) Effective Thermal Conductivity 


In addition to concentration gradients, a temp- 
erature gradient may also exist within individual cat- 
alyst pellets during gas-solid catalytic reactions. 
This temperature gradient may have a more significant 
effect on the reaction rate than concentration gradient 
may have, and hence on the effectiveness of the inter- 
nal surface of the pellet. 

Mischke and Smith (77) measured thermal conduc- 
tivities of pellets of alumina particles as a function 
of macropore volume fraction. They found that the 
thermal conductivity of solid alumina decreases with 
temperature. The low thermal conductivities of alumina 
itself indicated that severe temperature gradient could 


be possible in porous catalysts for highly exothermic 




























~oi toa? shh skies 2s0q ona soxitesmeusdo jaed dotriw 


, (38) vatem yo veg besdqme enw (1)? noktotut 


mobist 6 beraoqotd (AOL E02) ds Eanes brie 76%8W 


. of sol bow Sevlade > teetoge ibfd a.xoi (shom st0¢ 7 
Sivpegtilb evistostts efi4 to? roleassgxo a8 er 
4308 3 ¢ ¥ 
¥ a . aT eS aS + CG . = gd 
t4 ; a~s MoM - $ 
F hb 
+ betiees seed esd Lshom, s20q ietlcieg & a'xefeeriw 


teyletso molA ony nt yttviewits6 avitpstis et eismisas 


ybute ein sot 


yiteigoyonod Laarredt evisseits (iv 


2 a : 
:; : 1" 


~qmat s ,edagihstp noliésidneones oF soto thbs nt 


$52 [subivibat nidtiw saixe oals Yam tnetbhsete smwsste 


sityisseo bilor~-ese patauh eseliea tayis 


.anortosey 
tnavitinpte stom & ever yam jnsibese. s1vsexsqmes aidT RE 
tasibs2zY roi}sxtrsoqo> asty SIS% NMOLIOKST eds no tostis 
«ystat sft 10 sepnovisoetts ais no sored bos ,ever vem Ar 


telieq oft to sostaus idn 

feorxond bexwasem (TT) aa ime, bas aadosit 
sorsornya 6 #8 BekoLened: se kent LB: to aselieg to aoisivis 
di ont toda seid your tok 88 ‘snusfow. g10qgor38m to 


~adiw anceerne enimuts meted: ao eprvbsoutnes Kemsods 
cs ee ak nd aot -oat : eeuassemes | 


nag ste! } : men pea rs 
7 a, rR Lb ey ‘ 

f salar ae - : he» } /% te mH | a 0 
’ ’ oS : P oka f i 


* Fe al 
cats ake) BER ee 





4 y Ly + 4 
é ; oo 


reactions. But the large effects of density of pellets 
on the effective thermal conductivity suggested that 
the area of contact between particles in the pellet 

was a Significant parameter which was more important 
than the conductivity of the solid phase itself. 
According to their experimental results, the alumina 
catalyst pellet of void fraction of 0.4 filled with 

air at 120°F was found to have the thermal conductivity 


O©-O0S0G2 Btuynr. ft. of. 


62 


o 


av 


[ieqg To yitanab 29 a isotle aprel. aft tof  enoisoset 


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tolisq sdt al esfolsxseg ~sewisd seagne> Yo Ber ens 


tasti1eqmi sxom-asw tfotdw yevomarsq treolttiapre © 260. 


ifeetti scaly Diloa 44% to ysivisoubaos ea neds 
pnimois edt ,atiowes iesmomi: :aIKS tbed3 O3 pokbroo0A 
dtiw belis? 8,0 to noisoseti bige to telileq seyls365 
ifoubnoo Léniend edt sved of Bawot esw T°OSL 38 tis 


1°33, tA\wha $80.0 Fo 













CHAPTER IIL 
PREDICTION OF CLAUS UNIT PERFORMANCE 
Performance of the Front-End Burner: 


In a conventional Claus process unit, one-third 
of H,S in the feed stream is converted to SO, in the 
front-end burner and the remaining two-thirds of HS 
is eventually converted to elemental sulfur in the 
catalytic converter according to two consecutive 
reaction schemes (1.1) and (1.2) described in Chapter 
1. If an assumption is made that the acid-gas feed 
stream contains very small amounts of See es like 
hydrocarbons, other reactions than reactions (1.1) and 
(1.2) may be ignored in the calculations to predict 
the performance of the front-end burner in a Claus unit. 

The reaction stoichiometry may be written in 


terms of fractional conversions of H.S, a and 8, accord 


2 
ing to reaction schemes (1.1) and (1.2) respectively,. 
in the following way when the two reactions occur in 


sequence and not in parallel. 


3 


H,S + z 0, —_—-> sO, + HO (i 2) 
3-a 1.5(1-a) oO oF 

; 3 
2H.,S + SO. aD x S. + 2H,0 Cow) 
3-a-28 a-B Ly 56 at+2B 


63 






















ttentee int-sogsd sit 39. angeugonzes  T.€ 
(to7-en tien seenctq aust Tent rigviio> & nt 
02.0 hat rey foOD et MAS SIL hast sot As SH 29 
2.H to. 2 d+-owd. prinbeme: eri bas. teatud has-tno1t 
mst feanemals- ot bestevaoo yilsstneve ei 
VES BHrOD Owed OF pribs bet TOHTPISVNGSD sityvisiso a 
esd ni bedieveeb (4.4) Bas (2.1) gometae noitoset :: 
beet ssap-bhos edt s6d¢ sham wi Morsoqnvecs as re f 
[ 2557, £31 j etemcomes Lome Yisv aenistnos masses : 
bos (1.1) arickseaes nedd, enotdoaon 33ri30 -2nodisonbyd ; 


7s 


jotbeta. ot endian iuatiss st at Betonp. ad yam (Sui | = 
7 5s nt tenttiid boe-tnord eda To sonaateitiaeg sit ie 
it netdinw od yen yvidemoidotote ‘notdosex saT : 
fronos 4 bee 0 . SEH to ano terswiton Taniotsoest to sme 
,vievispedes: (S.-i): Bas (£05) eemorioe noij vse oz pai = 
ai tYvopo aznoftoss: Owe oft £orw yaw priwol fo orit < 
isfisseg ak tom bas eaneypes ; -~ 


~ 


H+. .02) aaa oe High 


& 


¢ ale 
ATTN | 
(S-SiL B-€: 


Te oe 


The assumption of a consecutive reaction 
scheme is quite reasonable since reaction (1.1) is a 
very fast combustion process while the reaction (2) 
is very slow in the absence of proper catalysts even 
at high temperatures. Furthermore the equilibrium 
constant is much larger in the reaction (1.1) than in 


the reaction (1.2) as shown in Table 2 (33). 


TABLE 2 


EQUILIBRIUM CONSTANTS FOR REACTIONS (1.1) and (1.2) 


Temperature, _°C Reaction (1.1) Reaction (1.2) 
700 0.5562 x 1077 0.9928 x 10° 
800 0.1388-x 1077 0.1612 x 107 
900 0.9608 x 101? 0.2394 x 107 
1000 0.1449 x 1018 0.3320 x 107 
1100 0.4018 x 101° 0.4363 x 107 
1200 0.1809 x 10° 0.5495 x 107 
1300 0.1205 x 1027 0.6689 x 107 


Since the formation of S5 by the reaction (1.2) 
may lead to other forms of sulfur molecules such as S¢ 


ox. Ss it is necessary to include reactions (1.8) and 


ee 
(1.9) along with the reaction (1.2) to completely 


describe the reaction system. 


ba 

























woicnest evituosands # TO MOLsgmAess emt 


s 2k (i.f) mokgoneax ovonte eidsneasex St Lep, es. comedians, 


(S$) notsosst sme elidy RAPOOIU moljsudmoo Jest ytew ' 


1S 2 evisdeo. teqmag To oon 22s 6dF ot woLe qaey at - 
muirtdiitups sit sxomredisi4 sudo eeemet Hpi 38 : 
7 ea 


ni seri [.P) fektones off al saprset gipuer et tistaneo 4 


(££) S*eide? aL, swore ‘ap {Sef} nofdoset od? 


¢ SHAT 


(S.f) bas. (f.0) B4O7 TUASA 09 STVATANOD Mj in alatoge 
eee ei a ee re : 


(¢.f) noksosas ({.f) sotiosea 2° _,sigtstegmet |” 


SERS ye aad. aor 


‘ 


8 * , 
Of x #S22.0 


or x Skat.0 Sor 3 eset) ee 


f 


“ 


Ob x wees.e G4g1 x gba .6 bee)” = 
Sy x oseest aE eS door 
ois theme. Sere St0k.as es 
for x wehbe. 6 Story cost a es 


“OL x 889.0 Moi gant... ae ae 


on 
7 


s. {) aoLdoset sft ya (° to gokteneot pl3 sd 
ees ove esluostom watDse Yo emo? 8m oe bes vi a 
al (8.1) encitossy sbutont oF ‘ae - ane te ag . a 
oxtageiqmen, of (Saf) solves: a ‘die! enol 4008) 


3 


65 


3 S.5 << Se Chet) 


1.5(1-v,-v,) 0.5 6 Vy 


= So pen! So (1.9) 


eo B (1-v,-v5) 0.375. 8 V5 


Here V1 and V5 are defined as "fraction of the 


equivalent moles of S. which associated to form S, or 


2 6 


So", and.can be calculated from mole fractions Xo Xe 


and Xe obtained in the equilibrium calculation by 


using the following relationships. 


3 Xe 
Pe Gea tony oer (3.1) 
X, + 3X, + 4 X, 
Che: 4 Xo 8! 
2 - 
Xo + 3X, + 4 X, 
va 5! See bs Pla nee (3 23) 
The computed results of values of Vy and V5 at differ- 


ent temperatures are shown on Figure 2 which was ob- 

tained by using the free energy minimization method. 
In this reaction, the assumption of the exis- 

tance of only three different molecular species of 


sulfur, Sor S. and S, was inevitable since the exis- 


6 8 
tence of other sulfur species was still controversial 


Pe ae 














13” ef bank¥eb ois ~V bas »v aion 
o+ £et5ide226 foittw <2 io aolom taeisvivpe - 
enoitosx? ofom got bejsivoleo ad. aso .bns »"g@ : 


- 


\ mriacdifimpe ofa ok bontstdo gt Sas 


vd noltsiIvale 7 ua A ae 


aqinsaokss fsx paiwoliot stit patey — 





y = 4 = 4 mL 


{tib te. . Has, 6 20 senlsv to edivmes beduqmoo sq? 


—“Eet gi I 
‘doitiw: S$ stapit ae oom: mee nowutsrogued ine ~~ 


“do es 


_ horisem aero imimim. yp 2sa® ax ont endes vd sey 
-eixne oft to noliqnane Say. viola pest ekets at | oO 
to setnege <= kpeby tnsxaht tb somtt eine to eonss by 


¥ abe, —7 ips S) = 








uae ond rare pil 


66 


; qUgHaSOWLY ANO iv 8s uo Is 
OL daivId0ssw s 40 SHIOW LNATWAINO’ AHL JO NOILOWIA +z gUNOTA 






es 
AVA NIN ING LAU AY YENINISIS) NA Asot" 
DON ‘- \ OOOO oe) SANS \ 0001 








zHu7T JO 


AOL TIAA 
Z ¥ © Ls f 2 AHO 3 


and no reliable data for them were available over the 
practical temperature range of the burner condition. 
In calculating the equilibrium composition of a part- 
icular gas mixture at a fixed temperature, the equil- 
ibrium distribution between Sor S¢ and S¢ was assumed. 
In tracing reaction paths which do not follow a 
sequence of equilibrium compositions, e.g. adiabatic 
reaction paths involving the reaction (1.2), the dis- 
tribution of the sy compounds may not be predicted 
because the reaction rate data are not available for 
these assocation-dissociation reactions between sulfur 
species. For this kind of situation it is often 
assumed to facilitate calculations along the reaction 
path that the various species of elemental sulfur will 
peat eani ls bices along that path. 

It was also assumed that the feed HS could be 
completely mixed with stoichiometric amount of air to 
form a uniform homogeneous reactant mixture right after 
the entrance of the burner. This condition may be 
easily achieved by modifying the burner chamber using 
some baffles or vanes in the path of the reactant 
stream. 

All reaction components were assumed to be in 
the gas phase and to exhibit perfect gas behavior over 


the range of compositions and temperatures encountered. 


A constant pressure of one atmosphere was employed 







19336 3fipit sass 


fdsilibys s190w moat TO? ets sidsifiex on bas 


conind eds to spas espteregmes Isotsosiq 
noigiaogmoo myixdilivupa eda pnizsivoiao al 


“lips oft .oxtdsrequied bexit 6 3s Stesxim e8e weafvot 


bomuers 25w.,28 Bas .2) +8 aeewted woiszudixteib muizdl 


s wollet ton of dotdw efiveq noitoset painosis al 


.enoisksoqmos muiudilinpe to sonsupes 


noftoss2x edt privioval eiséq adksoset 


betoibeta od ton. vet ebaeognos a sd to neisudtas 


fol eldsicrsvs jon sets 6465 9¢61 notfoss1 ods eSausoed 


4 snoisoss iad tsiooansb~notsanores sean 


suitiue neswss' 


‘atto- ek th notseazie to bara sitet 10% .sefoege 
cnols emoitsivoliseo estes ilios? of, beauees 


trnamels 20 Betoesge <cvokrsy oft Jade Apeg 


fiw totinoe: Is 

disq tec parole odpetél itertiem: 
ad Hfiirted Belt beat oie tadd bemtvess oats Paes 32 
of uis bo Sroms aks tsomdietase Ad iw: Beg den yladetqmas 


tnajoses euossspaned mastias 6 ogot 


em 
ad yen nokttibroo abit .tenzed sff “to: SonBtsas edt 
prtey tachi torsed sit patytibom ‘ya boveldos tien 


tosgoses oft to es att mk SoaaY to, aoitted amnoe 


v 


ai 7 ae 


ni od o3 bomvees Stew e$nonognoo noisones IfA 
tevo rolvertad ese 008m rasnclip ai scat ae ott 


poke zs 


aa: ae na Te ah 


at, 5 ‘ 


omer 





S 


ae ae ime Fi 



























68 


throughout the calculations by neglecting the effect 


of change of the mole number within the reaction 
system. Quantitatively the increase of the total 
number of moles in reaction (1.1) is 0.5 mole per 
10.14 moles of the feed mixture as illustrated in 
Figure 7. The change of the total number of moles due 
to reaction (1.2) depends upon the reaction tempera- 
ture. After reaction (1.1) has been completed, the 
temperature in the front-end burner becomes high 
enough, as shown in Figure 4, to make the average 
number of atoms in sulfur species smaller than 3. Then 
a net increase in the total number of moles can be 
predicted due to reaction (1.2). Therefore, the approx- 
imate maximum pressure decrease due to reaction (1.1) 
and (1.2) is expected to be about 5 percent of the 
feed pressure. 

The following data for heat of reaction (54) 


Mere iereiminse Pr eaCerOone: 0b.) (1. 2), 1.8) and (1.9)... 


° 


AH (1) = -123,924 cal/gmole of HS 


298 2 
AH 5 96 (2)..= 5,625 cal/gmole of HS 
AH5 95 (7) = -22,673 cal/gmole of S5 
AHj5, (8) = -24,753 cal/gmole of 8, 


The product mixtures were assumed to contain No, 


H,S, SO,, 05, Hor So) S_ and Sp: The thermodynamic 





69 


equilibrium compositions between different sulfur 
species along the reaction path were obtained using 
the free energy minimization method (72) and published 
thermodynamic data (71,105). The derivation of equa- 
tions to get adiabatic reaction paths in the front-end 
burner and the related computer programs are presented 
in Appendix C. 

Reference to the form and magnitude of the 
equilibrium constants for reaction (1.1) and (1.2) as 


tabulated in Table 1 and described in the following 


expressions 
PLavs P 
: e SO, HO 0. YH,O Et 
Pe ME 2a nine ice us (3-4) 
H,S 0. H,S 0, 
3/n 2 3/n 2 
Pp Pp y 
K = Sn ae on i oes 3/n-1 
pye diiuey E 2 3 iP =a) 
Gib wate. Yu.s Yso 
2 2 2 2 


shows that reaction (1.1) is essentially irreversible. 
Equations (3.4) and (3.5) also show that equilibrium 
compositions are relatively insensitive to pressure 
changes, and when the total pressure, 1, equals one 
atmosphere, they are independent of reaction pressure. 
The effect of the inert content on the equili- 
brium conversion can be predicted when the equilibrium 
constant is described in terms of conversion. Based 


upon the stoichiometric feed ratio of HS and SO. 


eo 






















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70 


in the presence of I-moles of inerts per a-moles of 


so the equilibrium constant for reaction (1.2) may 


Pa 


be obtained in terms of the fractional conversion of 


N (1-3/n) 
Kot Gy) © 2% s+ E-1) 7 (3/n-1) 


nates) = (3.6) 
where No is the total initial number of moles in the 
feed mixture, which is equal to (3a + I). From equa- 
tion (3.6) it may be found that the equilibrium con- 


stant for the Claus reaction is a function of X and 


N 
(2) for a constant value of n at one atmospheric 


pressure. For the value of n equal to 3, RS 2 appears 
to be constant for a fixed value of X, which means 
there is no dilution effect on the conversion level due 
to inert gases as long as the H,S-SO, feed ratio is 
maintained at the stoichiometric 2 to 1 ratio. Whenn 
is not equal to 3, the value of Bord may decrease or 
increase depending upon the values of (=) Aid jie. Ae 
nis less than 3, there is a net increase of the total 
number ot moles due to the reaction while a net decrease 
for the value of n larger than 3. When n is much less 
than 3, the equilibrium conversion level will increase 
as the inert content increases. On the other hand, 


when n is much larger than 3, the equilibrium conversion 


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[Moles of H 


level will decrease with the increasing inert content. 
For the value of n which is around 3, the effect of 
the inert content on the conversion level cannot be 
easily predicted from equation (3.6). 

The computed results plotted in Figure 3 shows 
the interesting behavior of the equilibrium conversion 
level depending upon the inert content and the average 
number of atoms in the sulfur vapor. Figure 3 really 
indicates that the conversion level of the Claus re- 
action can be improved or deteriorated by increasing 
the inert content in the feed stream depending upon 
the operating temperature. When the operating temp- 
erature is above 900°K, the conversion level increases 
with the increasing inert content, while the conversion 
level decreases with the increasing inert content in 
the temperature range below 750°K. 

Figure 4 shows "fractional conversion of HS to 
Sn plus so," as a function of the reaction temperature 
where the conversion X is defined by 


= 


> in acid-gas]-[moles of HS leaving the burner] 


[moles of H5S in acid-gas] 
ot eae) Coe) 


In Figure 4 the lines below the dotted level of 33.3% 


conversion level represent reaction paths according to 


71 





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74 


the reaction (1.1) while the lines between the equil- 
ibrium curve and the 33.3% conversion level represent 
reaction paths according to the reaction (1.2). The 
bottom curve in the family of equilibrium curves in 
the upper part of Figure 4 shows an X-T plot for 100% 
HS content in the acid-gas, i.e. for no removal of 
produced sulfur, while the other curves represent the 
same gas but after increasing levels of sulfur removal. 
Figure 4 shows that removal of 10% of produced 
equilibrium content of sulfur vapor followed by re- 
equilibration of the reaction mixture results in addi- 


tional conversion of HS to Sy according to reaction 


Z 
(1.2) from the 0% sulfur removal curve to the 10% one. 
Thus the ordinates for the family of curves represent 
the cumulative conversion attainable after a number of 
process steps during which the reaction system remains 
closed except for the removal of sulfur. 

Figure 5 shows a lower temperature portion of 
Figure 4 on an expanded scale to facilitate examina- 
tion of reaction paths for reaction (1.2) in a catal- 
ytic converter. Figure 4 and Figure 5 will be useful 
for burner and catalytic converter calculations which 
involve reactions (1.1) and (1.2) occurring sequent- 
ially but under adiabatic flame conditions. 

The equilibrium conversion of reaction (1.2) 


is temperature dependent at a fixed inert content as 
























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76 


shown in Figure 4, At low temperature ranges of below 
800°K, the equilibrium conversion decreases as the 
temperature increases, while at high temperature 
ranges of above 800°K the equilibrium conversion in- 
creases as the temperature increases. The minimum 
conversion occurs at the temperature range of about 
800°K depending upon the level of sulfur removal as 
shown on Figure 4. 

The existence of a minimum conversion point in 
the reaction (1.2) can be explained by the association- 
dissociation reaction between different sulfur species, 
which causes an exothermic reaction scheme at the low 
temperature range and an endothermic reaction scheme 
at the high temperature range. The plots in Figures 
3,4 and 5 are thus really characteristic of reactions 
(1.2), (1.8) and (1.9) rather than of the reaction 
gl opal a ao 
7 Since adiabatic temperature rises are non- 
equilibrium reaction paths in which the X-T coordinates 
are uniquely related by the condition that the enthalpy 
of the system remains constant, equilibrium consider- 
ations in the reaction mixture, except for that be- 
tween different sulfur species, are not required to get 
the X-T relationship. The magnitude of the slope ees 
at constant enthalpy of the system depends upon the 


sign and the magnitude of the heat of reaction as well 


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as upon the capacity of the reaction mixture to absorb 
the heat released, or vice-versa. 

The effect of inert gases upon the dx/dt curve 
for reaction (1.1) was computed, as shown in Table 3, 
to estimate the inhibiting effect against a temperature 
rise for cases of 10% excess amount of No, 20% and 40% 


of additional CO. contents in the acid-gas. The inert- 


ness of CO, was checked by using a feed acid-gas con- 


2 


taining HS and co. 


might contain COS and CS 


to get product distribution which 
2° The computed results in 


Figure 6 show that CO, contents in the acid-gas stream 


i 
can be treated as an inert gas in this particular study 
to predict the adiabatic reaction path in the front-end 
burner since the conversion of co. to COS and CS. is so 
small for the given temperature and pressure condition. 
The influence of other impurities present in the acid- 
gas in small amounts were neglected. 

Figure 7 illustrates the basis on which the 
adiabatic temperature rises in the front-end burner 
may be predicted. The temperature rise, (To-T)), re- 
sults from reaction (1.1) occuring to completion, 
X = 0.33, along an adiabatic reaction path; (T,-T.) is 
the slight temperature drop arising from the further 
conversion via reactions (1.2), (1.8) and (1.9) to 
form an equilibrium composition of reactants (HS and 


SO,), products (So, Ser S, and HO) and an inert (N,). 


8 


77 


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P f . ; . u 2 C cw | bh 


EE a aenemememmememmmnemneeeaa as Sa 





FRACTIONAL CONVERSION OF CO. 


0.014 


0.012 


0.010 


0.008 


0.006 


Feed composition (moles): 
H.s 30.0 
05 oak 
No 56.4 


0.004 


.002 
Sond Co 20.0 


2 
Pressure: one atm. 





0.000 
1000 1100 1200 1300 1400 1500 1600 1700 


FIGURE 6: EQUILIBRIUM CONVERSION OF CO. IN THE FRONT- 
END BURNER 


79 








— 


80 


REACTIONS COOLING (and 
(1) id): condensation of S$) 


‘ 

peer | Se Se, AN LS St Er a NS 

’ . P| . 
‘ . 


; 
<—BURNER - 
7 WASTE - HEAT 
Se eet i 









ACID 
GAS ee 3 ae 
aN CONVERTERS 
i, , t, is 
MOLES OF COMPONENTS 
HS 3 RASS 2 (1-x) 
CaS 0 O 0 
N, 5.64 5.64 5.64 5.64 
SO, O 1h 7-x 1-X 
H,O 0 1 12X 1+2X 
Saf oO O  (1-V,-V,)(.5X) 
$5.40 O° v(15Xx) ‘Gane 
Se “mad O } vh(1.5X) 


X= EQUILIBRIUM CONVERSION ATTAINABLE AT. T; 


FIGURE 7: FRONT-END REACTION PATH 





81 


Further cooling occurs within the waste heat 
boiler and is assumed to produce either a cooled gas- 
eous stream (in a once-through process) or a gas and 
liquid sulfur mixture at the vapor-liquid saturation 
temperature of sulfur. The actual location of the 
temperature, T3, may be conjectural. To predict the 
plant conversion levels use can be made of the above 
sequential reaction scheme. 

Adiabatic X-T paths have been plotted in Figure 
4 and 5 for various inlet temperatures Ti: These 
paths all terminate at X = 0.33 as described above. 
Further adiabatic temperature changes via reactions 
(1.2), (1.7) and (1.8) are shown in the region above 
X = 0.33 and below the equilibrium curve labeled as 
"0% sulfur removal". 


Each of the T,-T, paths will terminate on an 


2ru3 

re T equilibrium point. When the additional cooling 
occurs in the waste-heat boiler section, the reaction 
rate will be suppressed and the resulting change in 
state may be followed along the constant conversion 
line until the specified temperature Ty is attained. 


The extent of cooling to T, along the constant conver- 


4 
sion line will determine whether sulfur condenses or 
remains in vapor phase. 


One point to be emphasized in Figure 4 is that 


the conversion X represents total conversion based on 


cf 


7 


'_- 


im 





the initial HS feed, i.e. conversion of HS to both 


2 
SO, and S,. via two reactions (1.1) and (1.2). The 


dotted line labeled 1-2-3-4 in Figure 4 illustrates 
the reaction paths for a 100% HS acid-gas undergoing 
the process already described above. If 50% of the 
sulfur so formed is condensed in the waste-heat boiler 
and then separated from the reaction mixture, the 
curve in Figure 4 labeled "50%" would describe the 
subsequent equilibrium x" =" ‘compositions permitted 
for such a stream composition. 

Here, the equilibrium calculations and estimated 
X-T paths do not indicate whether the predicted temper- 
ature rise is sufficient to maintain a stable flame 
temperature in the burner or not. This condition must 
be specified on the basis of plant experience. 

The adiabatic reaction paths which have been 
calculated were found to be quite dramatically influ- 
enced by the equilibrium distribution between the sul- 
fur species, So Se and Sgr at some temperature ranges, 
especially between 700° and 1000°K. 

The discontinuity in X-T paths at X = 0.33, 
strictly speaking, is not realistic, but appears of 
convenience in analytic calculations, because the 
overall temperature rise during the overall reaction 
paths would become the path from Ty through T3- In 


this sense the slope of the reaction path has its 


rs 





83 


meaning ultimately to predict the terminating product 
temperature from the front-end burner just before the 


catalytic converter. 


Performance of the Claus Catalytic Converter 


Upon specifying an exit temperature and the 
corresponding composition, Ty and Xy in Figure 7,. of 
the product stream from the particular waste-heat 
boiler, further reaction may be promoted by contacting 
the stream with a bed of catalysts. 

The cooling process from T3 to Ty contributes 
not only for condensing of sulfur formed in the 
front-end burner but also for obtaining a higher equil- 
ibrium conversion level in the temperature range below 
800°K. This additional sulfur formation and increasing 
the overall conversion level by alternately cooling the 
gas stream and then passing it through a bed of catal- 
yst is the most popular scheme in a "once-through" 
sulfur recovery process. To predict the overall per- 
formance of a Claus unit a proper mathematical modeling 
of the catalytic converter is also required in addition 
to the front-end burner calculation. 

Various kinds of mathematical models for a cat- 
alytic fixed bed reactor have already been discussed in 
the literature survey. In the case of a Claus cataly- 


tic converter the catalyst bed is very large in its 


uh 


~s) 


.. 


opis 


\g 


+ 


Ys ov 


re 


at{it yeo2 


ge 


me tm 


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oL3 





diameter (15 ~ 20 feet) and the depth is much shorter 
(3 ~ 4 feet) compared to its diameter. Therefore, the 
assumption of adiabatic operation is realistic, in 
which case the one-dimensional model is applicable in 
the absence of radial concentration and temperature 
gradients. 

For the above reason, a homogeneous one-dimen- 
Sional model was adopted to simulate the Claus cat- 
alytic converter, more particularly that originally 
developed by Liu and Amundson(62), their so-called two- 
phase model. According to this model, the general 
transport and reaction processes in the catalytic bed 


may be described by the following equations: 








‘lou ee sbteys 
i — + (C. - C_.) - D. —>—= 0 (3.8) 
int az : £ s L az 
B 
aT AL h A_h 
£ #4 _B us 
Pr CoE Vint az t (Tp Tg) ‘: € (Ts T,) 
ER B 
= 0 (3.9) 
An oe (C. - Ce) +n Pp Ty (P.,/T,) = 0 f3g2.0) 
aer. 
A, h re o Te) + n AH Pp abe (P.,T,) - Ke oe, = 0 


Hangs <i e.  (3.11) 


84 


, . 
o 
— 4 — 
j 
- 
otal hs 
. a 
.? . 
7 ~~ 
Cc “4 7 
: 
: o 
Ns 
4 . 
: \ 
: 
- 
en | 


o 





’ 


85 


In the above equations, the rate expression to 
be used herein is the one proposed by Liu (66) for 
the Alon catalyst. Actually the model by Liu and 
Amundson (62) is modified by including the internal 
transport resistances through the insertion of the 
effectiveness factor which was ignored by many invest- 
igators (30,49,102). 
By assuming the following conditions: 
(1) Negligible axial dispersion 
(2) Negligible axial conduction 
(3) Uniform velocity profile across the bed 
(4) Adiabatic condition with surroundings 
(5) Constant pressure along the reactor 
(6) Stoichiometric feed ratio of HS and SO, 
(7) Equilibrium distribution of sulfur species 
between Sor S¢ and S. at the given temper- 
ature along the reactor 
(8) Uniform concentration distribution on the 
catalyst external surface 
(9) No Poiseuille flow in the pores, i.e., a 
negligible change in the number of moles of 
the reaction mixture. 
the general equations for heat and mass balances may 
be reduced to a simplified form as presented in Appen- 


dix E. The resulting simplified equations to be solved 


become, in terms of dimensionless variables, 





86 


aX 
aL + A, (X, oe X.) = 0 (3.12) 
A, (f) ee ae 
ALS (RL -2 X_ es exp (<—)= 0 
= Ah ee (urerenn Deke £2) fo 
4 Beis 
98 2916 (3.13) 


Equations (3.12) and (3.13) were solved numer- 
ically by using the Newton-Raphson method in equation 
(3.13) and the standard Runge-Kutta-Gill integration 
ened in equation (3.12) repetitively starting from 
the inlet up to the outlet of the reactor with the 
precalculated value of the effectiveness factor. 

The most important consideration in this treat- 
ment arises with the effectiveness factor of the cat- 
alyst pellets, whose effect on the actual reaction 
rate is implicitly included in the parameter A, in 
equation (3.13). The various methods to calculate 
Or estimate the effectiveness factor of the catalyst 
pellets have extensively been surveyed in the litera- 
ture survey. 

The appropriate choice of a calculation method 
for the catalyst effectiveness factor can reduce much 
of the computing time since the effectiveness factor 
should be repeatedly calculated along the reactor bed 
to obtain an accurage prediction of the reaction path 
through the bed. The first step to be taken before 


choosing any proper calculation procedure is to decide 


8 

























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87 


whether the reaction involved is in the fast reaction 
regime or in the slow one. In the fast reaction re- 
gime the concentration of a reactant rapidly drops to 
zero before it can reach the center of the spherical 
catalyst pellet while it gradually decreases until it 
can reach the center of the pellet in the slow reaction 
regime. 

To decide upon which reaction regime is applic- 
able, the concept of the Thiele modulus was used . 
For spherical catalyst pellets, the Thiele modulus is 


defined by (94) 


(3.14) 





For Thiele modulus of less than 0.5, most of the reac- 
tion occurs in the entire catalyst pellet in the slow 
reaction regime which means that the reaction rate is 
Slow enough compared to the diffusion rate for the 
catalyst effectiveness factor to become approximately 
1. On the other hand when the Thiele modulus becomes 
larger than 5 the reaction occurs in the fast reaction 
regime where the reaction rate is fast enough compared 
to the diffusion rate to make the effectiveness factor 
much smaller than unity (61). 

In this study, the calculated value of the 


Thiele modulus (Appendix F) at the reactor inlet 





7 
7 
ary 


be 


—J4 


ai 


r 


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pee s 
c 
: 
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rf 
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4 
- 
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say 





88 


conditions shows that the reaction occurs somewhere 
between the fast and the slow reaction regimes since 
the calculated Thiele modulus is about 5. This value 
of the Thiele modulus enables the asymptotic solution 
of the transport equations to be used to obtain the 
effectiveness factor with an error of less than 15%, 
as shown on Figure 8. 

To make certain whether the reaction interface, 


X defined as the point in the catalyst pore where 


I’ 
the concentration of the reactant might drop to zero, 
falls in between the external catalyst surface and 

the center, Van Den Bosch's collocation method (90) 
was modified and applied to this reaction system. The 
collocation equation was derived for this particular 
reaction system (Appendix G), which was solved using 


the false position method to get X_ with the optimum 


a 


collocation point of a as proposed by Van Den Bosch 
V2 


(99) for the high reactivity model. The calculated 


results showed that X_ was equal to zero implying that 


iE 
the positive concentration of the reactant could reach 
the center of the catalyst pellet. 

Therefore, the accurate value of the effective- 
ness factor may not be expected by employing the shell 
model on a flat slab approximation in this particular 


Situation. Judging from the above preliminary invest- 


igations, the Weisz and Hicks' conventional method (106) 





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89 


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was supposed to be the most proper method to calculate 
the effectiveness factor although it was a little slow 
and required a number of iteration. 

The detailed derivation of equations to apply 
the Weisz and Hicks' method to this study is presented 


in Appendix H with the related computer program. For 


the purpose of comparison the asymptotic solution using 


the approximate flat-slab model was also obtained as 
presented in Appendix F. The calculated values of the 
effectiveness factor by both the conventional numeric- 
al method and the asymptotic method are shown in 
Figure 8 as a function of the Thiele modulus. In 
Figure 8 it can be found that the difference between 
the two methods is significant when the Thiele modulus 
is less than 10 and negligible when it is larger than 
10. The Thiele modulus at the reactor inlet condition 
was calculated to be 4.4 and at the outlet condition 
5.4. The calculation procedure was given in Appendix 
E and H. The corresponding effectiveness factor of 
0.18 and 0.15 may be seen in Figure 8. The maximum 
percent deviation of the effectiveness factor is, 
therefore, about 8.3%. Furthermore the effectiveness 
factor is essentially a linear function between the 
Thiele modulus of 4.4 and 5.4 as may be recognized in 
Figure 8. As a result, the arithmatic average value 


of the effectiveness factor between the inlet and the 


90 


we 





91 


outlet condition was treated as an overall effective- 
ness factor for the whole length of the reactor bed. 
In further calculations to predict the characteristic 
performance data of the Claus converter the overall 
constant effectiveness factor with the value of 0.17 
has been applied. 

The effective diffusivity was calculated using 
the parallel pore model based upon Chuang's experi- 
mental data (18) for the pore size distribution of the 
Alon catalyst. The average pore size employed was 
80 A, in which range the diffusion phenomena certainly 
occurs in the Knudsen diffusion regime. The tortuosity 
factor was chosen as 4.0 in reference to data by Hideo 
Teshima (46) and the recommended value of Satterfield 
(93). The calculated value of the effective diffus- 
ivity was 0.001888 as shown in Appendix D. 

The effective thermal conductivity data were 
very limited. Fortunately, however, the catalyst 
pellets concerned in this study have pore structures 
of small enough dimensions to be operating in the 
Knudsen diffusion regime. Furthermore the pore size 
distribution range around the average value was rela- 
tively narrow according to Chuang's data (18). Oper- 
ation with this type of a catalyst would therefore be 
expected to be free of internal thermal effects within 


the catalyst pellet and justifies neglecting of these 

























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92 


effects even under the severe mass diffusion effects 
(106). Often the effect of the thermal conductivity 
becomes very important for the pore structures of 
non-Knudsen diffusion regime (106). 

The data obtained by Mischke and Smith (77) gave the 
value of the effective thermal conductivity of alumina 
catalysts with macro void fraction of 0.4 equal to 
0.082 But/hr.ft.°F at 120°F under the atmospheric 
environment. This value was used in this work to pre- 
dict the intraparticle thermal effects. The results 
of calculations indicated, as presented in Appendix F, 
that the internal thermal effect may be neglected. 

For this kind of small pore structures, the 
above analysis implies that the effective diffusivity 
rather than the effective thermal conductivity may 
play a very decisive role. Because of the absence of 
the internal thermal effect the energy balance equation 
within the catalyst pellet does not need to be consid- 
ered when computing the effectiveness factor, i.e. the 
mass balance equation is the only equation to be solv- 


ed. 


Results of Reator Modeling 


The temperature and the conversion profiles 
along the catalyst bed are plotted in Figure 9 for a 


feed temperature of 550°K and space velocity of 


oe 


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FRACTIONAL CONVERSION 


1.0 


0.6 


0.4 


0.2 


FIGURE 


9: 


X_¢ vs. depth 


Ty vs. depth 


Pressure 1 atm. 


Space Velocity = 1000 hr’ 


Feed Temp. 550° K 


Feed Composition Mole percent 


N, 69.286 
H,S 6.143 
$0, 3.071 
H,0 21.500 





0.2 0.4 0.6 0.8 1.0 
REACTOR BED DEPTH M. 


TEMPERATURE AND CONVERSION PROFILES ALONG 
THE CATALYST BED DEPTH 


JUNLVYAd SL 


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94 


1000 hr *. ‘the profiles indicate that a significant 


amount of reaction occurs at the entrance of the re- 
actor bed and almost maximum conversion may be reached 
at the depth of about 2 feet. With this prediction ba 
can be suggested that if the space velocity is in- 
creased to a higher level than 1000 hr>?, a greater 
yield may be obtained without affecting the reactor 
efficiency. 

In Figure 10, the effect of the external trans- 
port resistance is shown for the inlet section of the 
reactor since the inlet section can have the greatest 
gradient in temperature and concentration when the 
Significant reaction occurs at the inlet section. 
According to Figure 10, the effect of the external 
diffusion effect is quite negligible even in the inlet 
section while the thermal resistance is considerable. 

The negligible mass transport resistance in the 
external fluid film can provide a unique steady state 
solution of the transport equations (3.12) and (3.13) 
even with considerable thermal resistances. The pos- 
sibility of the existence of multiple solutions due to 
the external resistance was checked by solving the 


equation (3.13) for X with X,. as a parameter using 


f 
the false position method. The computed results are 
plotted in Figure 11. Obviously no multiple solutions 


may exist under the reaction conditions of this simu- 


lation and, of course, under the practical plant 


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td 
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ty 0.10 
0.08 
Pressure 1 atm. 
Space Velocity 1000 hr’ 
> 0.06 
Feed Temp. 550 K 
555 
Feed Composition Mole percent 
0.04 
N. 69.286 
H,S 6.143 
$0, 3.071 0102 
H,0 21.500 
550 


0.01 0.02 0.03 
REACTOR BED DEPTH M. 


FIGURE 10: EFFECT OF EXTERNAL TRA:ISPORT RESISTANCES 


$0, 


FRACTIONAL CONVERSION 


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A,(f£,f,) : 


+ 


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CHECK ON THE EXISTANCE OF MULTIPLE 
SOLUTIONS DUE TO EXTERNAL TRANSPORT 
RESISTANCES 


96 








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TROUWHART: JAUAITKS or ies Bes 


97 


operational condition. 

Figure 12 has been plotted to evaluate the 
effect of the feed temperature on the X-T plot with a 
Space velocity of 1000hr +. The slope of * appears 
to be almost constant for the different inlet tempera- 
tures. 

Figure 13 represents the effect of the feed 
temperature on the temperature profile along the 
reactor with a constant space velocity of 1000 hEtee 
From the slopes of the temperature profiles it can be 
predicted that the feed temperature should be above 
500°K to get high enough reaction rate right at the 
entrance of the catalyst bed. 

Figure 14 shows the effect of the feed tempera- 
ture on the conversion profile along the reactor with 
a constant space velocity of 1000 hr tl. Judging from 
the slopes of the profiles it can be confirmed that the 
fast reaction rate right at the entrance of the catal- 
yst bed can be achieved when the feed temperature is 
above 500°K. The difference in the conversion level 
appears to be negligible at the outlet of the reaction 
when the feed temperature becomes higher than 500°K. 

Figure 15 indicates that the effect of the 
space velocity on the conversion level at the fixed 


feed temperature of 550°K is significant when the 


Space velocity exceeds 2000 hr |, but negligible when 


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$rfpit etsy noidtossx dutede: seat tep ‘od ree 


E — = . 
bad Jaye 63 ‘BD aris 0 oasis 


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tojoBnex od profs SiLtorg noks aseyhon let Sd 


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1 r ‘ a 


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its 5 

ints teqmes bast ony noxrtw bovielyte . 2 a 
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i yrs 0 
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‘ Ps " 9o . 

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ws 


i? as mtd as f Level sone 






98 


LOId L-X NO SUYNLWATdWaAL GHead AHL JO LOGAAH *7ZT weno 
No ‘dUNLWUIadWnaL 


OOL 0S9 009 USS 0 OS? 


sO 
ine) 


00S“ Td 
je alee 3 
EviT. 2 
NES abe 
(qus0z0d 
:uOTITSOdUOD peadq 
ay 000T 
2AQRTOOTSA 3s0edS 
wjze [T :eansseig 


NoOS? = dWaiL ddd 





OISUGANOD 'IVNOI LOW 


x 


ay ‘Cog AO ft 


Be 


: 


rat; 
is 


iBE 


yy 
a & 


- ~ 
<3 F US) ee i 


LE 


LEED 


ik 


als Oh. wk 


ra at 


BS 












Vv 

j 

rg2 
+ SHe. 


> 
. 
a 
— 
; 
3 


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Ee 
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vv 


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+ 


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wet 
7 
=DSGe 
L6227 
B56 
al 


QD 

wil 6? 
ares 
* 

é 


228° 


ecc 


Ww 
mw 





aie 


FEED TEMPERATURE, 


700 


650 


600§ 


550 


FIGURE 13: 


FEED TEMP = 600°K 


Pressure: l atm 


Space Velocity: 
1000 hr 
Feed Composition: 
(Mole percent) 
69.286 
6.143 
Bho 2 
tp qreray500 


0.4 0.6 0.8 
REACTOR BED DEPTH, M 


EFFECT OF THE FEED TEMPERATURE ON 
TEMPERATURE PROFILES 


= J 








FRACTIONAL CONVERSION OF S05, X 5 


FIGURE 14: 


wl 


FEED TEMP 
450°K 


Pressure: 1 atm a 
Space Velocity: 1000 hr 
Feed Composition: 

(Mole percent) 


Ny Sie Nera HO 
HAS 6.143 
SO SoU sl 

CaN Wee 469, 


2 


0.4 eae 0.8 
REACTOR BED DEPTH, M 


EFFECT OF THE FEED TEMPERATURE ON 
CONVERSION PROFILES 


1 





100 


rane - 
















: o ats £ ss teres 
* ery 

A ; ti QOOL :«yoiigofeV space 

§ a 4 >- re t ~~ 
4 


tO LT LBOGMO! 2a 


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(tinea eq SOM) 





- Spe lee . on 
2 ae fae. 
a fbf 3 a YT ’ - 

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) 002. ££. OH 


5 
8 
¥, 
h 1 ® 
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HO BaOManNNG cael aati, ag) 
Med), Biba dbanil Baker towns pie 
\- th oe a ae g 
' yh ae j iv sy Bria. 
3 i x ry / 2 Bey! 












fete a ae fa 


f> ee } 
poet oe 


ry! . y = & 








By 





_ 
< 


FRACTIONAL CONVERSION OF S904, Xe 


SPACE VELOCITY 


Pressure: l atm 

Feed Temp: 550°K 

Feed Composition: 
(Mole vercent) 


No 69.286 
Ho5 Ag 
S99 Sie ted 
H90 21,5990 





Cig 0.4 0.6 0.8 
REACTOR BED DEPTH, M 


FIGURE 15: EFFECT OF THE SPACE VELOCITY ON 
CONVERSION PROFILES 


101 


GS eS 6 ee Bt 


- 


peed SoA 


a 


[2-9 or eee: area < mace cumecennne, Este 







+ EVGA A al CAL eae De 


ieee ae 


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a) 
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wi ae TY mye’ & Cy 
‘~ Y¥PLOOd BOATS ‘ 


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i+ . i ny 7 te 
(Jes ac 30) shor) 5 * sy 
: ? 
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iy F 
So hr mot +r 
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ge 4 iv r oe 
EX a eee 
- ry ’ 5 rns 
ae * VFOins x 
} > a 


8.0 ; a-s () 1 


i % urea’ g da 







VPLS 


102 


fétia below 1000 hres 


What should be noted in this modeling results 
is the fact that the reverse reaction of the Claus 
process was not taken into account due to the lack of 
knowledge on the rate of the reverse reaction. However, 
the reverse reaction rate was found to be negligible 
compared to the forward reaction rate in its magnitude 
under industrial operation conditions according to Liu 
(66) in his statistical correlation of the rate data. 
Another point which should be made here is that simul- 
taneous reactions which may possibly occur in the Claus 
reactor, for example the COS-SO,, and COS-H,0 reactions, 


have not been considered in this simulation work. 


i. 


Is 
a 
7 net 


ey 


vs a 
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+ 550A wy 


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2 = os 









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2 serie as » sotisteqo isiti Seubink: nxebag~ 
ac 
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= & em od Bisode dokaw tniog tedton4 
ideeog vem doidw enoiso694 asi 






y 9 


a eit olamsxe * 208 ~ 


CHAPTER IV 
DESCRIPTION OF EXPERIMENTAL SYSTEM 


The experimental equipment consists of three major 
parts; a reactant feeding system, a reaction system and an 
analysis system. The reactant feeding system was originally 
designed and built by McGregor (72), to which an additional 
COS feeding line was installed in this work. In the reac- 
tion system, an integral bed reactor was employed instead 
of the differential recycle reactor which had been used by 
McGregor (72), Liu (65), and Karren (53) for their kinetic 
studies on the Claus reaction. To analyse the feed and the 


product streams a gas chromatograph was used. 


4.1 Reactant Feeding System: 


The schematic diagram of the reactant feeding 
system is presented in Figure 16. The feeding system 
starts from gas cylinders, each of which has been 
equipped with its own pressure regulator. In the nit- 
rogen feed line a second pressure regulator was provid- 
ed in series with the first to improve the control of 
nitrogen pressure upstream of the flow controller 
since the nitrogen cylinder pressure tended to drop 
much faster due to its much larger flow rate compared 
to other reactants. 


Each gas stream except COS was dried with 


103 

















oe TY » is i} 
\ ie J 
bn ' : 
; « rn] y y 
| we ’ eS 
i "pslOi] 
lt ‘ 
ry ps LAHS ) ‘ 
eet 
ANT aTyaMTTadka @0 worTaa Dem 3 
: Se a 
4 _ ‘ 
- ", irr 
ls 
i r 4 tro eseie tasmo lips. Lazio beste’ ott ° a) 
- * ” ‘ . - - m ee ty ‘ ’ ; 


6 bas moteyea noitoset- s neteye cies alii 
: ? a te i+ ’ if + > et dad % > . ’ a > 


A tostose Bilt ~ -mayeya eteyleng 


F 4 : ae Y a i“ : 

é Ybs as dotiw 62 , (ST) 2opempOM Ya Se ind bne Saas D 
; ; a j ; i balteszeak 2s witht enihie® 2 209 
rLOV 7 

hsetent bevolqme ssw 1O35s897° bed Isxps eink re <i eye a 5OW 2. 


34 
| 
a 

id 


ve bewt dead Ben dotdw rotosex Sioyoss Leivaoretits ons. 


— al 


sttenin xzieds ro? (€@) nex TEA Dns uhee uid Sey x0pe er 
Sais Dee bei : oda seyisns OF .f6 of45s851 ets amis ~ eaibe 


ee . 
ty eaw dosxporsmowdd ase & aon 


é- 


4 
1 ‘ . : ‘ = uD 


stat sine, p sitet ams 









= ia Ay & : 7 He | 
. ’ ; 
oribest Jnscose1 ons to MSIpSID. ‘ot Sbm Bras | ont e is y 
"i 3 yas “Ete 
mie sev a Ott r JS t Sl; rT e o {6 ae tn8 cf bhig: i L pbshasettas naueke 
rs exe = oxi aoisiw Tt ~ diee far. sbitb yo: ase nn i Pd 


~+ bn ala nk sors! Eset St. be daad re pe ‘Ga cor a 


-hivotg aw Tote Lape! sueeea hnonge 6 seit sa 
aN . ‘ a oe eS 


lo LloxrsneD omg: svoreme oO . te wats a 
tel lors nc > wort oa Xo aia ee 


ic pow? 


PR sin, of Tr exgeeand, soba 


re a ‘possemo> aids sh ponent att be 


in a % " a i or 7 

n 4 wt Pins a, P "2-7 -* 

a Oe OT .* od, ae : ~ 
a 7 ' v2 


. 


4 
uer's 





104 


VENT 
XD FEED PRESSURE 
TRANSDUCER 
SURGE TANK 
eates) D—> TO GC 
MIXING _ THERMOCOUPLE 


VENTURI 


LD S | FLOW 
| = : CONTROLLER 
Bb sl ck 


pp fer 
| i k : ROTAMETER 


SURGE 
TANK wr) DRYER 
GAS 
CYLINDER 
cos & ~ 
N ,—o DP-CELL 
: UO FLOW 
| CONTROL 
VALVE 
HS Cw) os ie 
PRESSURE 
REGULATOR a PREHEATER 


SO, & rs 


FIGURE 16: SCHEMATIC DIAGRAM OF THE REACTANT FEEDING 
SYSTEM 


f i PG Ne Ce 
§ 
1 a 
a 
} 
ws? > 
PASY » ea ve 










~y , , 
»* aso f 5a ae i | 
Hues h 4 o { A 
oa a ten , Py ’ 
(SE G2SHAAT’ Jk : 
[} 
been amas 
"I Me) “Al eas Y* 1 7 t 
if oye ¥ . i. 
“) ’ —_ +e ~ ee el pa me are ne 
: — ee te < fs : 


ae eC da TODOMEIE ae i wtKxIM j 
{] tAurMaV 


: 

. wou ® 
ABIIOATUCD tf 

| = H 

; 

7 ' 7 

cA 





AAVYAG 


IdaDd- 4a + , 
| fost 


| . OX dogeHoc 
eo) an avaAy 







i 


-putadins mwa ABA sa 0° 


oe 7 ar z ya ne 4 it 


es 
oi 
a 
: 


105 


anhydrous calcium sulfate contained in a 500 ml stain- 
less steel cylinder to prevent the homogeneous reaction 
between H,S. and SO. in the presence of condensed 
water vapor. The COS cylinder was checked for water 
impurities and in the absence of water content a COS 
drier was unnecessary. 

Glass rotameters with stainless steel balls 
were used to control the flow rate of each reactant 
stream along with the flow controllers of diaphragm 
type. A surge tank was installed in the COS feeding 
line to improve the feed pressure control without 
using the diaphragm type controller. The rotameters 
did not give accurate readings of the flow rates, so 
they were used only as a visual indicator of whether 
the flow rate of reactant streams remained constant. 
The accurate flow rate of each reactant stream was 
obtained by the GC analysis results of the feed stream 
combined with the measured total flow rate. 

The reactants were mixed in the mixing venturi 
to form the feed mixture of desired composition, 
whose total absolute pressure was measured by a 
Foxboro 66 FR-2 electronic absolute pressure transducer 
and whose temperature was measured by an iron-constant- 
an thermocouple. 


The total flow rate of the feed mixture was 



















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# ‘ th, 
x _ 
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bre ‘gH avswied 
. 


nabroo Yo sotees1q ena mi 08 


= 


27 BW 107 ats smarD acw y~eibitd iv eon exit e sOqaV  tet8W. Tait 
“ 4 
i Hgrias si , : _ j 
193N0D Yotew to osnetds’ #O7 or Bets sed tetcegmt mY 
; ‘ Zi ral 
vrsasepsum ssw teitb 
7 - - 17 i x 
tad loode vaelniste itiw exretometot eeal> 
/ ; and =< : i a 
¢reetose: doeg to ater wold sitfo Pomago oz bone sisw 


mpsiigqsth to exelioxtaes wol? sd astw alert seagate 
w = M 


oaibest 20D sf? ai belledsadi esw danas apie & eqys 


+ponsiw Lowttoo eigesstq heet sis vce og enti a 
; ; > 


ststanaset oT .telfoudno> say? mos iigsibh estt gates. : 7 


ce ,eotex wolt elt To apmrtbset hited ovie Jon bib 


itedw te srostaorbal leveiv. s se ine beew anew “yee 
-} on Pas 

42no0> benibmet empeite siesaie St 20 ode volt ® 
' x as ee 

onpw meowte thatoser dos to Sot ant we rt ‘B82H008 


mserte Haqt odd 7 ° ed Luaes elavisus 9 ‘ode ve ba 
 esoax wold fetot bexuissom ‘ene Atiw ben 


tustnov paikint eas ai. Bextor graw ® eyansos sc Sa. 
i. <P att MF 


sols Faoqgmoan bevieeb to pee ‘howd edd 


# ve bewwesem asw 6 cet ad stat oe 
ae = ot x 


4 





106 


measured by a Foxboro 613 DL differential pressure 
cell and controlled by a Foxboro Stabiflo 6R-V4 
control valve. A surge tank with a volume capacity 
of 500 ml was used to ensure complete mixing of the 
reactants before they reach the control valve. 

A variable portion of the feed mixture on the 
upstream side of the differential pressure cell was 
continuously vented depending upon the desired flow 
rate of the feed mixture through the reactor. 
Another small portion of the feed mixture was intro- 
duced continuously to the gas chromatographic detector 


for analysis of the feed composition. 


Feed-Product Analysis System: 


The composition of both feed and product streams 
was analysed by a gas chromatograph equipped with a 
Beckman - 320 programmer, Infotronics Aerograph 471 
digital integrator, and Hewlett-Packard Model 17503 A 
SNderh, The schematic flow diagram of the analysis 


system is presented in Figure 17. 


re ee Separation of components in GC Column 


To separate all components in the gaseous re- 
actant and product mixture the GC column was arranged 
in a three-column-in-series mode. All of the possible 


components, N co HS, COS, SO. and HO were 


oe * ay r 2 
























; ' eR ee 
ra [sitnessttid ia Ele proton & me aspoenathe 


a7 OFOU ‘ont vd perekenss hae Lise 


9aq89 enmirlov s tiiw anst spage.A- svisv feasn09 as 
‘Me bf: * 


ort © mixin stelqmon sxver 6 oF beat eaw fan one 7. ca A 


fs +, 3A) 
.ovisv lotdno> od dosex add stoted atietoset 
sft no stutxim beet eft: to coLstog. sldsiasv #& 
. v5 * fea i 
asw iieo siveeeig pstinaset tee oft to obie Meotsequ 
ass y “We Pa y 
wot bertesb sit noqu parbasqeh eee Yi aragals goo 


am 4 a coxa 5 i Sckm be ? eng 209 snes 


we 


7 


. % ‘ae “+ 
, 2 : oe 4 ; 
tosgosssb GLiGABIPOTAMOTAS =6P- sit OF ‘iavoontanes omen 


nots Leoqmoes | ban? ofa to ekayisas tot 


-ovani asw e1redgnim besi sis To finan {Il — sonsonA 


i 


fi 


' a 7 
meseye eteviank. 
pines ee ee owl en, , 
\ on 7 el Ge 
emsstde t+othoxug bos Sst Mod 10 nok dbeoqmoo. we 


bes a oe ne ee) ; 
s fitiw beqqiups dqe1petsmeins 26Pp.6 Lie 


etevyinas ont to MiB Lie sift wolt ~ienapes ae eer soi S 
















ey gi aa 6D Oh, > aie —_ 
AL otugtt af Bedhowerg — Lo 
7 - z) o . OS Di ‘ 
a k e f ‘ ~ ho — en 
re ae Sf alee (eh ee ay, CS 
’ ‘ek ee ane cee os ’ 
”. a 


omwied “yo 


mel 


7 * we 
tit 


“ST spoResD seed nti cna 


+ j 
x na 


_ a 
Bie bopmsxre 28w « nis £00 90 ert om 
i : ; mr ea p ; Opes)! Bo 


7 





oy Sr M 


L Wy oF 





A: GC DETECTOR 

B: COLUMN 

C: SAMPLING COIL 

D: SELECTION VALVE 

E: SAMPLING VALVE 

F:; FEED STREAM 

G: PRODUCT STREAM 

H: UNSELECTED VENT 

I: SAMPLED VENT 

J: UNSAMPLED VENT 

K: CARRIER HELIUM 

L: REFERENCE HELIUM — 
M: REFERENCE -HELIUM VENT 


VISUAL DIGITAL 
INDICATOR INTEGRATOR 
OUTPUT 
PRINTER 


FIGURE 17: SCHEMATIC DIAGRAM OF THE ANALYSIS SYSTEM 


se CHROMATOGRAPH 
ATTENUATOR 


ATTEN #1 


aa ATTEN #2 


COMPUTER 


REMOTE 
TELETYPE 


RECORDER 


107 


Tt 


‘Te V Mba domi 


4 " * y 
: ® 
‘ a” ie 
pT 
we’ a 
' 
j 
rr. 
yh 
‘ ’ 
} 
4 
. 
i wet 
~~ a 
" : 
‘ 
rs 
» 
x 
‘ 
7 - Saal 
ifs s 
he 












FOTIA gored a. D0 4 


‘dsae ota uaa r: 
SVJAV age 2 EN 
SVGAY DUEIGMAZ 21 
| maenee o 
MASTS TOUCOAS 
aes 


“wwey CIN 
MOLISE HLA 
Ue “OWT wars va 





108 


separated using three columns in series; the first 
column was a 3-foot Chromosorb 104, the second a 
6-foot Porapak QS in 50~80 mesh, and the third a 
3-foot Porapak T. All of these columns were made of 


1/8 inch diameter SS 316 stainless steel tubings. 


According to experimental results for optimum 
column arrangements, the longer the Chromosorb 104 
column became, the better was the resolution between 
sO. and H,O peaks but the worse between COS and H,S 
peaks. In the absence of the Porapak QS column H,S 
and COS peaks completely fused into one large peak 
while the separation between SO, and HO peaks was 
excellent with almost the same elution time as ob- 


tained in the presence of that column. Porapak T 


column improved separation of H,S and COS peaks but 


2 


made the elution time of SO, and H50 peaks longer 


causing excessive tailings in both peaks. With the 
optimum three-column-in-series arrangement, 3 feet 
of Chromosorb - 6 feet of Porapak QS - 3 feet of 
Porapak T, the total elution time was about 12 minu- 
tes with good separation between each peak at the 
column temperature of 190°F. The temperature control 
of the GC system was performed by Honeywell R 7161 
temperature controller and displayed on Honeywell 


Electronik 16 multipoint temperature recorder. 


















outdo: det? paten hadersqee ik. 


1p so03-E 8 SBe vere. 


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: 10% ea ieeas ft: nixvedk>s O22 piibr0998 i aT 
somoxn>: oft 1a} 43 .pornsmaprstas ameLoo — 
ag mA Sat: 
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> ns ~ow 343. 4nd dias 0,8 Gas OB 
~ > 


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é 


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7 te "I ~~, = a 
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20D hus @.H Fo nold SIBASe povenamtt eae: 


3 va =P B95 - OA 2. (LB se as 


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sds 36 Assad dose aielens ot: STE 
~~. 









a 
¢ 5% 


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y@ =? 
« 


a 


loxsneo sivtsteqmes ort < n20et 


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 RRewyanett no. Bs hove qeib 

ee 15 ak or ha 


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was 4s. 





e 





No. 


The experimental results showing a comparison 


between different column arrangements have been sum- 


marized in Table 4. 


TABLE 4 


COMPARISON BETWEEN DIFFERENT COLUMN ARRANGEMENTS 


Column 
Arrange- 
ment 


4A + 6B 
4A+6B+3C 


6B + 3C 


1A+6B+3C 


3A+6B+3C 


Numeric value 


Tareuosorh 104 column 


Porapak QS column 


Porapak T column 


Separation 


H,S-COS SO,-H,0 


2 Zoe 
good good 
good good 
good fused 
good fused 
good good 


Helium 
Flow Rate 
(ml/min) 


ES 


ga 


35 


6s) 


35 


the length of 


a column in feet 


Elution 
Time 
(min) 


30 
16 


excessive 
tailing 


less 
tailing 


12 


Column 
Temper- 
ature (°F) 


120 
200 


200 


200 


190 


The feature of the separation of each peak is 


illustrated in Figure 18 using the selected optimum 


column, 


column No. 5. 


109 
















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. tacnemtieexs eff oy > ah 
; 7 i 





> 6 PMiwore 74 or 
. i : 4 
reir Los drousi? ke meewied ~ 
’ is iy 
—s i 
.) oldest ok besizem 
© 
est cae 
*, = fa 
Par ’ 
s 
TKSOUARTA VoUIOD MAAN TINS HR, MOTRATIOID 
nts or 
rian | SOL dyonomomi> =A > aes 
amples 80- dager sa 8 > i 
’ * Mi + “4 : =F 
qmuios T deaqevot “J | 
— 
ial sit = aula? abtene Lam 
, 7 ’ ,4 if 
‘ a 
Yoet ak Amwlos & 
~ ~ 
“ ‘ , 
wm lo "pp Boag BS tsb at 7 roltetaqn?é . Te 
ele \ 4 . 
<—TSo VE 3/7 ; T gS $a P wold By F ? x shah 
°) side (qtmy  Ankm\To)  O,H=-0er 200-2 Hh - Tet 
i aal  N DLa Satvea s SS PI e © 
; i‘. “ ri a 
‘f ye : E . te - 
OS Ot et boos” + COR e: 
ees ; ’ fs ® 
6 m a , _? 
QOS of ec - ood -.. boog) “OS 
i . P ny : ‘. : = me ‘ei Se 
‘AOS be: ’ a be aut, ja hoog x 
ugi ike ; 1a 
a.) i sg” 


 mnisgo Bed 


a ia ‘ 
> = 






a6 


elge std p 


> gt Ampeg none to neltasaier At 







Ze boayt '  cbowg. - Oe 


ey 
ay 


> (es om <r 
Eat, 


Ny x 276 tak 


co. ee 


FIGURE 18: 


Column Arrangement: 


Helium Flow Rate 
Column Temperature 


Chart Speed 
Sample Size 
Attenuation 


Bridge Current 


HS + eee | 


eos <x ‘1 


4 


6 


8 


SO. ee 


Colum No.5 
in Table 4. 
35 ml/min 
190°F 

0.5 inch/min. 
Se Oat 

As noted 


250 mA 


H50 yaa 


TIME AFTER SAMPLE INJECTION, MINUTE 


TYPICAL CHROMATOGRAM FOR SEPARATION OF 
SO, AND H.O 


No, C05, H 


2 


5, COS, 


2 


2 


12 


110 





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4.2.2 Selection Valve and Sampling Valve Mode 


Two 6-way valves were used to select and sample 
the feed and product stream alternately. The struc- 
ture and the detailed description of the operating 
condition of both valves were given by McGregor (72) 
and Liu (65). Both valves were actuated with air 
pressure of about 20 psi initiated by a manual push 
button on the control panel; the product stream was 
selected by pushing the button and the feed stream 
was selected when the push button was in the pulled- 
out position. 

The sample injection button on the programmer 
chassis was pushed down manually for 13 seconds to 
sample the selected stream for GC analysis. The 
automatic sampling action by the cam adjustment had 
been employed by the former researchers (53,65,72) 
but was not used in this work because it was found to 
cause significant noise in the electrical circuit 
which extends from the,GC output to the recorder, the 


integrator and the computer. 


4.2.3 Attenuator Setting 


The individual attenuator in the Beckmann-320 
programmer which could be automatically actuated by 
adjusting the microswitch can also cause severe noises 


all through the electrical circuit when the cam 




























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position changed. So the timer cam was not used 


throughout this work. Instead, three separate atten- 
uators were installed; a chromatograph attenuator, 
attenuator I, and attenuator II. 

The chromatograph attenuator was set to 10.0 


which would give the maximum sensitivity while atten- 


uators I and II were set to 10.0 and 5.0, respectively. 


These attenuator settings were determined by the con- 
dition ith the maximum output of the largest peak 
(N. peak) from the GC should not exceed the maximum 
allowable input to the digital integrator at the max- 
imum sensitivity range (50 mV). 

The attenuator selection switch mode is shown 
on Figure A.4 in Appendix A. An on-off switch was 
provided in parallel with the attenuator I and II 
which were connected in series. When the switch was 
off, the attenuator I and II were connected to the 
output from the chromatograph attenuator giving the 
maximum attenuation ratio or minimum sensitivity. 
When the switch was on, the attenuator I and II were 
bypassed to obtain the maximum sensitivity or the 
minimum attenuation ratio. During the actual opera- 
tion the switch was off for the nitrogen peak, and on 
for other peaks to obtain maximum sensitivity over 
the permissible input signal range to the integrator 


and the computer. 


112 











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Reaction System: 


The reaction system was comprised of the pre- 
heater, a reactor, a sulfur condenser, a water conden- 
ser and a sulfur trap. The schematic diagram of the 


reaction system is shown in Figure 19. 


4.3.1 Feed Preheater 


The feed mixture was preheated from room temp- 
erature to the desired reaction temperature by passing 
through the preheater. The preheater was constructed 
from a solid stainless steel block. Two high-resis- 
tance heating elements were inserted in holes drilled 
within the block, and stainless steel 316 tubing was 
wrapped helically about the exterior of the solid 
block in 20 pre-machined helical groves. The entire 
assembly was insulated with a glass wool blanket. The 
power supply to the preheater was adjusted manually 
by a Variac. The temperature was measured and record- 
ed by an iron-constantan thermocouple located at the 
top of the preheater and a Honeywell Electronik 16 
Peneereears recorder. An auxiliary U-shaped heating 
element was installed between the feed preheater and 
the reactor to compensate for the heat loss from the 
line. The temperature of this element was controlled 


by an on-off temperature controller fabricated by the 


shop in the Department of Chemical Engineering. 


113 










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FEED 
MIXTURE 


PREHEATER 
WATER INJECTOR 


SRE |! 


Ss VAPORIZING 
> COIL 





REACTOR 
PRESSURE 
TRANSDUCER 







INTEGRAL 
REACTOR 







SS 316 SCREEN 


TO 
bad Gc 


FILTER 





SULFUR 
CONDENCER 





SULFUR TRAP 
(ROOM TEMPERATURE) 


WATER 
CONDENSER 
x< 
bi SULFUR ACCUMULATOR 
WATER i X SULFUR DISCHARGE VALVE 
DISCHARGE X 
VALVE 


FIGURE 19: SCHEMATIC DIAGRAM OF THE REACTION SYSTEM 


114 


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4.3.2 Reactor 


An integral fixed bed reactor was used in this 
study. The reactor assembly was fabricated using a 
stainless steel 316 tube of 1 inch diameter and 9 
inches long. A stainless steel 316 screen of 30 mesh 
Size was fixed at the depth of 7 7/8 inch from the 
top of the reactor tube, which was connected to the 
reaction system by 1 inch Swagelok fittings. The 
reactor assembly was easily removable from the system 
by disconnecting at points A and B in Figure 19 when- 
ever new catalyst was to be charged. 

The temperature of the feed stream at the reac- 
tor inlet was measured just above the top of the cat- 
alyst bed and that of the product stream at the re- 
actor outlet was measured just below the bottom of 
the bed by iron-constantan thermocouples which were 
precalibrated according to the procedure presented 
in Appendix B. 

The reactor outlet temperature was controlled 
manually by adjusting the input current to the reactor 
wall heating element by means of a Variac. The 
reactor wall heating element was made of the nichrome 
wire wound uniformly around the reactor wall. 

The pressure in the reactor during the reaction 
period was measured upstream of the reactor and con- 


trolled by a Statham electronic gauge pressure 


115 

















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116 


transducer connected to a Foxboro V4A pressure con- 
trol valve. The pressure transducer was precalib- 
rated by the procedure described in Appendix B. 

A water injection coil made of a stainless 
steel 316 tube of 1/16 inch diameter and 2 feet long 
was connected to the feed line between the preheater 
and the reactor. The water injected to the inection 
coil by a syringe pump was vaporized in the injection 
coil by heat supplied by the auxiliary heating ele- 


ment. 


4.3.3 Sulfur Condenser and Water Condenser 


The product stream from the reactor was divid- 
ed into two streams; the first portion entered the 
sulfur condenser and then the water condenser before 
being vented, while the remainder of the stream 
entered the sulfur trap to remove the product sulfur 
vapor before reaching the GC analysis system. 

The sulfur condenser originally designed by Liu 
(65) was found to be smaller than required for this 
study so that another additional one-pass condenser 
was added on the top of the original condenser to 
increase the holding time of the condenser system. 
The additional sulfur condenser was made of stainless 
steel tube of 3/4 inch in diameter and 20 inches in 


length with no interior baffles. 





















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117 


The temperature of the first condenser was 
Maintained at around 350°C at the inlet and 110°C 
at the outlet so that the product sulfur might be 
condensed into the liquid phase, which helped to 
maintain a steady flow of the gaseous product mixture 
without any serious plugging. The second condenser 
was kept at the room temperature to completely knock- 
out the sulfur vapor which would not be condensed in 
the first condenser. 

The water condenser was made of stainless steel 
pipe of 2 inch diameter and 10 inches in length, and 
packed with glass wool. The condensed water was al- 
lowed to flow by gravity into the bottom of the. con- 
denser to be accumulated before discharge through the 
vent valve attached to the bottom of the condenser. 
The product stream stripped of the sulfur and water 
vapor passed through the side opening to reach the 
reactor pressure control valve on the vent line. 

Since the product stream entered the condenser 
system merely for disposal by venting, the design and 
operational conditions were not so critical as those 
required in the previous studies (53,65,72). The only 
performance Pa itT onan for the condenser system was 
to keep the reactor pressure constant while maintain- 
f4 steady flow. Actually, however, the downstream 


part of the condenser system was gradually plugged with 







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118 


entrained sulfur dust making it necessary to clean 
out the lines every week. The pressure buildup due 
to the accumulation of the sulfur dust in the line 
was compensated by gradually increasing the opening 
of the reactor pressure control valve. 

The lines from the reactor to the sulfur con- 
denser and to the sulfur trap as well as from the 
sulfur condenser to the water condenser were kept 
heated by wrapping with nichrome wire to prevent 
sulfur vapor from condensing in the lines. The first 
sulfur condenser itself was also heated with nichrome 
wire windings. To control the input current to each 
heater, separate Variacs were used. 

A sulfur accumulator was attached to the 
bottom of the first condenser to receive the condensed 
liquid sulfur flowing downward by gravity. In adai- 
tion a vent valve was installed at the bottom of the 
sulfur accumulator to avoid the replacement of the 
sulfur accumulator by regularly removing the accumu- 
lated liquid sulfur through the vent valve. 

A detailed description of the structure of 
the first sulfur condenser was presented elsewhere 


(65). 


A...3.4.. SULEur Trap 


A sulfur trap was designed to remove the 


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product sulfur vapor before the product stream entered 
the GC analysis system. The first sulfur trap was 
made of a stainless steel 316 tube of 1/2 inch in dia- 
meter and 10 inches in length with a U-shape. No 
baffle was provided in the sulfur trap to reduce the 
dead zone around the base of the baffles which could 
cause the broader residence time distribution of the 
product stream in the trap. The second sulfur trap 
was made of 3/4 inch diameter stainless steel tube of 
7 inches in length and V-shaped, with no baffles in- 
Side. Each sulfur trap was alternately used while the 
other was undergoing cleaning of the accumulated 
sulfur. 

The temperature of the sulfur trap was kept 
just above the water vapor condensing temperature in 
the product stream, usually at the room temperature. 
This was to prevent absorption of HS and SO, in the 
condensed water with resulting chemical reaction 
which could cause incorrect GC analysis results of the 
product stream. The inlet temperature to this sulfur 
trap was measured by an iron-constantan thermocouple 
and recorded on the Honeywell 24 point electronic 
recorder. The outlet temperature from the sulfur trap 
to the GC system was also monitored by an iron- 
constantan thermocouple and recorded on the recorder. 


Just next to the outlet of the sulfur trap, a 


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120 


sponge filter made of polystyrene foam was packed 
in the line in about 1/2 inch depth to avoid any en- 
trainment of dusty sulfur particles upstream to the 


Ge 


Process Measuring and Control System: 


All temperatures were measured by stainless 
steel-shielded iron-constantan thermocouples. The 
voltage signal from each thermocouple was adjusted by 
a Acromag model 323 electronic O°C reference and then 
recorded on a Honeywell 24 point electronic recorder. 
The recorder was equipped with an integral solid 
state calibrator, a chart span selector of 5 to 
5000 mV span and a voltage range suppressor of 0 to 
1000 mV. The recorder was precalibrated using a pot- 
entiometer and occasionally checked when any malfunc- 
tioning was revealed. In almost all the experimental 
run, the 20 mV span was employed. 

The total absolute pressure of the feed stream 
was measured by a Foxboro model 66 FR-2 electronic 
absolute pressure transducer while the feed flow rate 
by a Foxboro model 613 DL electronic differential 
pressure cell. The reactor pressure was measured by 
a Statham gauge pressure transducer. The signals from 


the feed and reactor pressure transducer were recorded 


2 
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Lal 


on the Foxboro 6430 HF electronic consotrol recorder 
and controlled by the Foxboro V4A control valve. 

The feed flow rate was controlled manually by 
the Foxboro 69 PA-1 control valve in combination with 
the valve V-M on the schematic flow diagram in Figure 
oe 

The reactor inlet and outlet temperatures, the 
sulfur condenser inlet and outlet temperatures, the 
sulfur trap inlet and outlet temperatures, the pre- 
heater temperature and the auxiliary heater tempera- 
ture were controlled manually by adjusting the pe 
current through the variacs. The temperature of the 
gas chromatograph oven was automatically controlled 
by a Honeywell R 7161 temperature controller. 

The gas chromatograph and the process measur- 
ing devices were calibrated through the procedures 


described in Appendix A and B. 


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CHAPTER V 


EXPERIMENTAL PROCEDURES AND RESULTS 


The experimental work to be described herein 
deals mainly with improvements in the procedures employed 
in measuring conversions for reaction (1.2) using an 
integral fixed-bed reactor and then with evaluation of the 
performances of a number of catalysts over a variety of 
conditions. The latter experiments involve three main ob- 
jectives; a comparison of activities between the newly 
developed bifunctional catalyst and a commercial alumina 
catalyst, a determination of the influence of ratio of 
mass of activating agent to mass of y-alumina "support" 
upon the catalyst performance, and finally, an evaluation 
of maximum attainable conversions in the laboratory re- 


actor. 


5.1 General Experimental Procedures: 


5.1.1 Startup of the System: 


A known amount of a catalyst was loaded into 
the reactor, which was PReragenely vibrated to obtain 
a more uniform packing of the catalyst bed. After the 
reactor was reassembled, all of the threaded joints 


and fittings were checked for leakage using the soap 


122 













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ie . wh, ’ 7 o 


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bubble test method with positive nitrogen pressure 
of around 30 psi within the system. When the system 
was gas-tight, the auxiliary heater and the reactor 
were insulated by a box-type covering made of a 
blanket of ceramic wool. 

The nitrogen cylinder valve was then opened 
to start a flow of nitrogen through the system, and 
then all heating elements were switched on. The 
temperature of the catalyst bed was raised at the rate 
of approximately 3°C per minute until the feed temp- 
erature reached 570°K. With both the nitrogen flow 
and heating started, the flow of helium was initiated 
through carrier and reference sides in the GC system 
at the rate of 35 ml/min through each side. After 
switching on the GC systems. which included the GC oven 
and the detector filament, it took about 3 hours to 
raise the GC oven temperature to 190°F and more than 


12 hours to get a stable baseline. 


5.1.2 Experimental Measurement of Integral 


Conversions: 


All newly charged catalysts were activated by 
heating continuously with first the nitrogen flow 
through the reactor, and then with the sb ein HS mix- 


ture of about 3 mole percent in H,S concentration. 






















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; - 
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.fo0 wl ‘pimenee to: yeansid 

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a 

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a oxew 8 atmnems La oon ile ae 

+ oft ts berisx daw ted sey BI89- aft to 8% viseenene i 

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ah: 


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— a ‘oe ee : 


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wolt aspoxd tn ong. saz | t3 0 
aes 2H s Fhe ods ae : 
bs at te : AD i) Sea 
snobevn at ASS 


a aes + a 5 


124 


The feed temperature during activation and reduction 
with HS was maintained at 570°K for more than 3 
hours, respectively, which was slightly higher than 
the normal operating temperature. 

After checking the functioning of process 
measuring and controlling systems, the heaters about 
product lines leading to the GC and to the sulfur 
condenser were started. The gas chromatograph base- 
line on the recorder was readjusted by manipulating 
the zero setting on the Beckman programmer panel. 
Then, the base-line for the digital integrator was 
adjusted separately by manipulating the zero knob on 
the Infotronic digital integrator. During the gas 
chromatograph base-line adjustment, the attenuation 
ratio was set at the maximum sensitivity, i.e. at 
the minimum attenuation eatwer on both the gas chrom- 
atograph and the recorder. The base-line adjustment 
could be checked visually on the recorder reading as 
well as numerically on the integrator printout. 

Having finished the checking of the system 
behavior and readjusting the base lines.of the GC and 
the integrator as needed, the feed mixture was into- 
duced to the reaction system. To obtain desired flow 
rates and composition of the feed mixture, all the 


individual rotameters and the feed control valve were 


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125 


adjusted to the proper level. After getting the 
results of the feed GC analysis, the component flow 
rates and the total flow rate of the feed stream 
were readjusted by manipulating the rotameters and 
the feed control valve until the desired composition 
and the flow rate were obtained. The total flow rate 
and the pressure of the feed stream were checked and 
readjusted by means of the differential pressure cell 
and the feed pressure transducer as described in 
Chapter IV. 

To prevent disturbances in the analytical 
system which might occur due to the change of sample 
selection between the product and the feed stream, 
the same flow rate of the feed or the product stream 
through the GC detector was maintained by keeping 
the differential pressure constant on the mercury 
manometer which was installed in the sample vent line 
from the GC detector. Throughout this experimental 
work the differential pressure of this sample vent 
line has been kept at about 1 inch of mercury for 
both the feed and the product sample vent. 

When the reactant mixture was fed into the 
reactor, the temperature of the catalyst bed in- 
creased quickly due to the exothermic nature of the 


reaction between H,S and SsO.. As a result, to 


imge ip .epiredy odt oF om ae Sipior dotdw ee: 


ssite stonbouc silt sto best ons 0, St51 WoL) ne 
i vp 


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126 


maintain a constant temperature in the catalyst bed 


the input current to the reactor wall heating element 
had to be readjusted. 

After the reactor outlet temperature became 
stabilized, the feed and the product stream were 
alternately analysed by the gas chromatograph until 
more than three consecutive reproducible results for 
both streams were obtained. If the feed and the 
product compositions had remained unchanged, it was 
assumed that the reaction system had reached a steady 
state with a steady catalyst activity. Generally, 
when a new catalyst was introduced, a startup time 
of roughly one day was required because of the activ- 
ation period required. In making consecutive runs 
in which temperature of the reactor or space velocity 
were changed, a new steady state could be obtained 
within roughly 2 days for the reactor temperature 
change and 3 hours for the change of the space vel- 
ocity. 

The experimental data thus obtained included 
the reactor inlet and outlet temperatures, and feed 
and reactor pressure, the feed flow rate, the atmos- 
pheric pressure, the room temperature and the GC area 


results by the digital integrator printouts. 





















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5.1.3 Shutdown Procedure: 


The reactant flows with the exception of 
nitrogen were stopped when a complete set of runs 
for one type of a catalyst was obtained. The nitro- 
gen was kept flowing through the system for more 
than 5 hours to completely purge remaining reactants 
and products which might undergo additional reaction 
while standing within the closed system for prolonged 
periods. After complete purging, which could be 


checked by the H,S smell in the product vent line, 


2 
all power supplied to the system was switched off to 
cool the system to room temperature. 

Before another new fresh catalyst batch 
could be charged, the reactor assembly was disconnect- 
ed from the system. The batch of the used catalyst 
was carefully discharged from the reactor and examined 


for any changes during the reaction period before 


being stored. 


5.1.4 Materials: 
i) Feed Gases 


All gaseous reactants, HS, SO, and COS were 


obtained from Matheson Co. Nitrogen was obtained from 


Alberta Oxygen Ltd. Carbon dioxide was used for the 


che oe 


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gas chromatograph calibration and was also obtained 
from the Matheson Co. The purities specified by the 


Suppliers are listed in Table 5. 


TABLE 5 


PURITIES OF GASES 


No 99.99 % min. 
HS 59.50 4 min. 
Cos 97.50, 27min. 
SO. 99.98 % min. 
Co 99.995 5 min. 


2 


However, the oxygen contamination in the 
nitrogen cylinder was frequently detected and so, 
every nitrogen cylinder newly delivered was always 
checked for oxygen content using Molecular Sieve 5A 
column of 16 foot by 1/4 inch O.D. preconditioned in 
vacuum at 250°C for 20 hours. The thermal conductivity 
cell was kept at 220°C and the bridge current was 
150 mA. The column temperature was kept at 135°C with 
the helium flow rate of 80 ml/min. The purity of HS 


or SO. was checked from the feed GC analysis data. 


2 
The resulting feed GC data never showed peaks other 


than No, hoo Or sO. peak, which meant that no detect- 


2 


able amount of impurities existed in the HS or SO. 


cylinder. In the COS cylinder carbon dioxide was 


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always detected in very low concentration, but water 
vapor was not detected. Since carbon dioxide can be 
treated as an inert gas, which was demonstrated in 
Chapter III, the COS cylinder was used without any 


further purification of COS. 


cee Catalysts 


The standard catalyst employed in this study 
for the purpose of comparison between different cat- 
alysts was S-201 y-alumina manufactured by Kaiser 
Aluminum and Chemical Sales Inc. The properties of 
the S-201 catalyst specified by the manufacturer are 


presented in Table 6. 
TABLE 6 


STANDARD CATALYST PROPERTIES (S-201) 


Surface area: 380 m*/gm 


Chemical composition on dry basis: 


Sio. -a.0..020 5 
Fe,0, ) 020 3% 
TiO. 0.002 % 
Na,0, 0.300 
A1,0, 93.600 & 


Ignition loss 6.000 &% 


129 

















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The developed bifunctional catalyst evaluated 


in the present study were prepared by crushing Kaiser 
S-201 alumina catalyst in the size range of -3 to +8 
mesh and selecting the -12 to +24 mesh fraction 
through the use of standard sieves. This was follow- 
ed by chemical treatment with various concentrations 
3)2 yes H,0 before heat treatment in an oven 
for 20 hours at a temperature of 550°F. Heat treat- 


of Cu(NO 


ment resulted in the evolution of oxides of nitrogen 


(z097% 


Data Reduction Procedure: 


The experimental measurements of temperatures, 

pressures and flow rates were obtained in terms of 
the percentage of full scale readings on the elec- 
tronic recorder except for those of the atmospheric 
pressure and room temperature. The reactor inlet 
and outlet temperature data were converted to the 
actual temperature scale in degree centigrade using 
the calibration equation presented in Appendix B. 
The pressure data were converted to absolute pressure 
in mm Hg for the feed, and to gauge pressure in mm Hg 
for the reactor pressure according to the calibration 
equations presented in Appendix B. The feed flow 


rate data were converted to the actual flow rate in 


130 


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131 


SCFH by means of the calibration equation for the 
differential pressure cell obtained through the 
procedure described in Appendix B. 

The feed and product compositions were calcul- 
ated from the gas chromatographic peak areas using 
the calibration equation for each component as des- 
cribed in Appendix A. To calculate the molar feed 
rate of each feed component from the calculated feed 
composition and flow rate, the assumption of ideal 
gas behavior was applied for all gaseous components. 
In calculating the product stream composition, the 
nitrogen flow rate was taken to be the same in both 
the feed and the product streams since nitrogen was 
an inert gas in this reaction system. From the cal- 
culated nitrogen flow rate and the composition of 
the product stream, the molar flow rate of each of 
the reactants in the product stream was obtained. 

The fractional conversion was then obtained 
from the difference peessasatig molar flow rate of 
a reference reactant in the feed and the product 
Stream. As a reference reactant, HS was chosen for 
the H,S-SO, reaction, and COS for the COS-SO. re- 
action. 

The partial pressure of each component over 


the catalyst bed in the reactor was calculated from 


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132 


the product composition and the total pressure in 
the reactor, taking into account equilibrium dis- 
tribution of elemental sulfur species. 

A computer program for calculation of the 
material balance was written to process the exper- 
imental raw data into the calculated conversion of 
H,S and COS simultaneously. The detailed calculation 
procedure for a particular run is shown in Appendix I, 
accompanied by the listings of the corresponding 
computer program, MTBAL, which was stored in the disc 
core memory. | 

Input of the raw data to the IBM 1800 computer 
system and output from the computer in the form of 


printing of the computed results were performed on 


the remote teletype located in the laboratory. 


Experimental Results and Discussions: 


BS rb ekl Preliminary Investigations: 


i) Performance Test of the Sulfur Trap 


To check whether significant additional re- 
action due to the catalysing effect of the liquid 
sulfur occurred during the condensing within the 


trap of sulfur. vapor in the product stream, two 

















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133 


different methods were tested. 

The first method used an ice-bath. The U- 
shaped sulfur trap was immersed in a cold bath fil- 
led with ice-salt mixture to maintain a constant 
bath temperature of -20°C. When the product stream 
was introduced through the sulfur trap, both sulfur 
and water vapors were condensed into the solid 
phase new he entrance of the sulfur trap. This 
resulted in frequent plugging of the line, which 
necessitated frequent cleaning of the sulfur trap. 

In the second method, a high flow-rate method, 
the U-shaped sulfur trap was kept just above the 
water vapor condensing temperature, around 20 to 30°C 
depending upon the water content in the product 
stream. In this method, only sulfur vapor was con- 
densed in the trap. 

In these tests the feed to the reactor con- 
tained 3 percent of H5S, Lio. percent (OL SO. and the 


balance N., on the molar basis, and the temperature 


Pe 
of the product stream at the inlet to the sulfur 
trap was around 500°K. The tests were carried out 
over flow-rates of the product stream through the 
BUbLor trap eranging strom):20° to 500 mi/min. 


Figure 20 shows the comparison between the 


performances of the two methods. For the high 




















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FRACTIONAL CONVERSION OF H5S TO SO. 


100 


FIGURE 20: 


o High Flow Rate Method 


A lce Bath Method (~20°C) 





200 300 400 500 


SAMPLE FLOW RATE, ml/min 


SULFUR TRAP PERFORMANCE TEST 


134 


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flow-rate method the effect of the additional reaction 
due to the catalysing effect of the condensed liquid 
sulfur appeared to be significant at a flow rate below 
30 ml/min of the product stream as shown in curve B. 
In the ice-bath method the high conversion level of 
H,S at the low flow-rate range of below 50 ml/min, as 
shown in curve A, seemed not to be the results of any 
additional reaction but to be the result of physical 
OC ecanuiee of HS in condensed water during the con- 
densing period of water vapor. To check the effect 

of physical absorption of HS in liquid water the 
sulfur trap was kept at a temperature below the sat- 
uration temperature of water vapor. The resulting 
data showed that the H 


S and SO, peaks in the product 


2 2 


stream decreased. Sometimes HOS peaks completely dis- 


appeared from the product stream. The physical absorp- 


tion effect was much more significant for HS peaks 
than for SO, peaks under the same condition. 

In the high flow-rate method, because of the 
smaller temperature gradient between the wall of the 
sulfur trap and the product stream compared to the 
ice-bath method, the condensing rate of sulfur vapor 
might be supposed to be slower and the condensing zone 
in the trap to be longer than those in the ice-bath 
method. In spite of these phenomena, the resulting 


data shown in Figure 20 revealed that the high 


135 

















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136 


flow-rate method at room temperature gave lower con- 
version of HS than the ice-bath method. It should 
be noted here that the lower the conversion of HAS 


which was obtained during a steady-state run, the 


better the performance of the sulfur trap. 


As a result, no convincing advantage is 
apparent for using the ice-bath method when the flow- 
rate of the product stream exceeds 100 ml/min. The 
disadvantage of the ice-bath method, therefore, was 
found to be the difficulty of eliminating the H,S ab- 
sorption in condensed water while that of the high 
flow-rate method was the entrainment of sulfur mist 
downstream from the sulfur trap. 

By considering these performance test results 
and disadvantages of each method, it was decided that 
the high flow-rate method would be employed for this 
research but with a modification of the sulfur trap. 
The modification was done by packing a dust filter, 
made of styrene foam, right near the outlet of the 
sulfur trap to provide sulfur demisting. A flow-rate 
of 200 ml/min through the sulfur trap was employed 


throughout this study. 


ii) Homogeneous Reaction Test in the Preheater 


The effect of homogeneous reaction between 
HS and SO, in the preheater was checked before making 


kinetic runs. When the flow rate of the feed mixture 



















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was above 2.0 SCFH, the effect of homogeneous reaction 
in the preheater at 600°K was found to be negligible. 
However, when the preheater was contaminated with the 
product liquid sulfur, the effect of homogeneous re- 
action was quite significant. This sulfur contamin- 
ation occurred due to condensation and accumulation 
of the product sulfur in the preheater when the pre- 
heater was not completely purged out before shut-down 
of the heating system after each run. 

Considering the above preliminary investiga- 
tions the preheater was completely purged out using 


pure nitrogen for more than 5 hours after each kin- 


etic run to prévent any possible condensation or accum- 


ulation of the product sulfur due to the slow homogen- 
eous reaction in the preheater. Since the effect of 
the homogeneous reaction was found to be significant 
when the feed flow rate was below 2.0 SCFH, the feed 
flow rate above 2.0 SCFH was adopted throughout this 
study except for runs K and L for the test of the 
maximum obtainable conversion level of the Claus re- 


action. 


iii) Calibration of the Gas Chromatograph 


The gas chromatograph was calibrated according 


to the procedures presented in Appendix A. 


137 


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The first calibration, whose results were 
shown in Figure 21, was based upon the no-individual 
attenuation scheme in which only one general purpose 
attenuator (chromatograph attenuator) was used. In 
this calibration, the reliabilities of the digital 
integrator and the computer were examined by comparing 
their results with that of the disc integrator which 
could be supposed to be the most correct. In Figure 
22, the integrated results from the digital integrator 
and the disc integrator shows good consistency while 
those Gotained by the computer are very much scattered. 
This scattering may be attributed to an improper input 
signal to the computer system without any attenuation 
of the output signal from the GC detector since the 
attenuation scheme II has been employed throughout 
the first calibration procedure. The results from the 
digital integrator in the first calibration were 
applied in data reduction calculation for suns A, Ba C;, 
Dy .E and J. 

A second calibration was carried out as above 
but using different attenuation ratios for the N,-peak 
and other peaks. Since the digital integrator perform- 
ance proved to be quite reliable in the first calibra- 
tions the disc integrator was not used to check the 
accuracy of the second calibration. The attenuation 


scheme III was employed for this calibration. The | 





























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ttapasdtes yas tdodd iw msvave secre roto) ont os. tenehe 


p } 


eft. esonie totostsb I) edt Mori aia ssqgwo oda 20 re 
iqmo assed eet Tiremgn ee aokeunegss 
a 


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by 
yew noitdsrdiisos pax eds at xovsrpesnt tnthl Z 


i .A alitre tot aoidgafitolao n notsoubex wie ak bettas 
: ml bis. a 





: ‘ee 
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ait yot eortsr nokrs 


— ei ae rm 
“Miolisa trostbipesaL 1 pih™ ‘Slt | eons) i 


, ee Peer? = 
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a ¥ tage ay 
, <add goo od Been ‘don 2sw ay ast 

‘a ~% 7 T 


a 


avksegente sai. 801983 dis 


. iy NY { 
ton ec | oe 2 f 4 
4 r ri vi ae Ge fe of L& 
r ; 4 pam ¥ 4 ) 
e » 


HS (OR CoS, SO,)/N, MOLE RATIO, PERCENT 


O N5-H5S CALIBRATION 


4 N,-COS CALIBRATION 


2 


Oo N,-SO, CALIBRATION 





1 2 3 4 5 6 - 
H,S (OR COS, S0,)/No AREA RATIO, PERCENT 


FIGURE 21: FIRST CALIBRATION OF GAS CHROMATOGRAPH 


139 














oy ’ 
. ee ay 
1607 TARBTAAD a oo 
Pride =e Fae 
y Ail) pana ano 209 gif ‘ os 
ie : iki 74 
‘ AOrT RELOAD Ota * 
PS | A es 
Bate 
; mo 
~ ae Aber 
ye 5 Oy, 
ee Se 
* ‘\ 
La es 
ae 
he 
. 3 a TS 
‘TIOREE! OL TAA KBRA cM oe Pe 
% 4 ying y 


+ 


WIAROOTAMOAHD 2AD 4 


1, MOLE RATIO, PERCENT 


x 


H,S/ 


Sample: No-H,S Mixture 


Attenuation: Scheme II in 
Appendix A 


Computer calc. option: 7 


o Digital Integrator 
A Disc Integrator 


Oo Computer 





y es 6 8 10 Ba 14 16 
H,S/N5 AREA RATIO, PERCENT 


FIGURE 22: PERFORMANCE COMPARISON OF DIFFERENT 
INTEGRATING SYSTEMS 


140 


ons | fhe i UES 1 


ennai 










‘ i 
\ 
c \ 
\ ~ ( 
\ 
‘ nr +. 
Ss n 
\ \ CPs 
w . 
iat A 
‘\ ‘, 
0 & we 
\ 
\\ 
\ 
»* 
\ 
a] hh # 0, < 
ew? ee \. 6 bY 
exusx th 2Qn -) tebiqmne “ey | X. ; 3 
— 1 j oe = ‘ i 
nk IY smedo® +o iteunedgsaé - 
A xibnsqgh | 
¥ molgqo -oles vetugmnom 
roserpetol. Isecipia ‘Oo 
fossrpSdal sak. a ae oe a 
yosvGMon, 2. ; 







~ pele PL 
t eae ee 
7 pr 






- 





141 


results in Figure 23 show a very good consistency 
between the digital integrator and the computer. The 
second calibration results by the digital integrator 
was used in data reduction calculation for runs F, G, 
H} 17 *he ana i 

Reproducibility of the computer and the dig- 
ital integrator has been compared for a calibration 


run in Table 7. 
TABLE 7 


COMPARISON OF REPRODUCIBILITY 


Digital. Integrator Computer 


Peak . Ave. Area Avg. % Avg. Area Avg. % 
Deviation Deviation 

198768 °0.0039845 5319610 0.0042479 
HS 73196 0.0054839 1939811 0.0079665 
Cos 97239 0.0034246 2587350 0.0041820 
sO, 18646 0.0375630 456142 0.0126670 


In the above comparison the calculation option 
2 was employed in GCJOB definition because the cal- 
culation option 7 required a detailed elution time 
data for each peak. These were not conveniently 


available because the status of the equipment at the 


Jitltdcouborqer i * 


woataagy GMO 2 


rotsxpedal Leoiped =. 


4 pvA - sath avA 
noijsived 


eT Oe 


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<0 80 veel 7 a 

: | - t 
e9£Heeoe.0 > ener Loy + 


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te" 
- 


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bas 4 .f wR”. 


asd rovsrpetnk Is3t 


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aie ~ 


ee 


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‘il 


ae es eesre 


scion 
28 at exon 





ye 
sub isd 


inne gon atew sane 


to) ae. bofis: a5 






H,S (OR CoS, SO) /N, MOLE RATIO, PERCENT 


Attenuation: Scheme III 
Computer Calc. option: 2 
Digital Integrator: 


oO H,S-N, Calibration 


& SO,-N, 


Oo COS-N. 


Calibration 
Calibration 


Computers: 


x For-all. Calibrations 


FIGURE 23: SECOND CALIBRATION OF GAS 





CHROMATOGRAPH 


142 

















ee OS a 













. a os i 
a \ " ad Ps A 
“4 “ - i 
‘ is 
. Ls j Wr ” ae eh oe 


4 


It smadoa PRE a ee 
: S :pésgo soled xoduquOD © ot 
‘ |. ggoSaapetqT Lsebpid Ie ia 
a aolssxdé isd Pua Oo 
me > Ae aoksertdtis: gg 02 a 
res kopbret ets D F ote 2 
sao of 

enotspudiled ate xot x = 


hh. 


i] 


>, 


HOA: <YRAMOMED 240 40 worTAaEtAO 
eg ied, aie! i ca cee 


fe : la ce 


143 


time of this study was still not fixed. Detailed 
descriptions on the calculation option can be found 
elsewhere (110). 

Table 7 shows that reproducibility of the 
digital integrator is better than the computer for 


No, H,S and COS peaks but worse for the SO. peak. 


4 
The less reproducible result for the SO, peak of the 
digital integrator seemed to be due to base line 
readjustment to correct drifting during the experi- 


mental run. The SO, peak is very sensitive to the 


2 
base line adjustment due to its flat Gaussian shape 
and a long elution time. However, the need for SO, 
peak areas could be eliminated by using areas of 
other components in the sulfur balance calculation 
‘ae the reaction system. Thus, it was decided to use 
the digital integrator rather than the computer. The 
lower reproducibility in the computer results seemed 
to originate with the interfering effect of signal 
transmission noises along the electrical circuits 
leading to the computer input terminal, and also to 
the absence of detailed information on the elution 
time of each peak. This missing information prevented 
the cer from starting action at the right instants 
during the GC cycle. 

The integrated results from the digital inte- 


grator or from the computer were printed on a Victor 











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shoe i: 

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144 


digit-matic printer, or on a teletype, respectively, 


at the end of each peak. 


SS ae Comparison of Catalyst Activities 


i) Effect of Catalyst Promoter Upon Simultaneous 


Reactions 


Simultaneous conversions of both the H,S-SO. 


and the COS-SO., reactions on four catalysts were 


2 
measured in the integral-bed reactor to evaluate the 
dual activities of each catalyst. The catalysts tested 
included pure y-alumina (S-201), 5.4 % Cu-on-alumina, 
12.08 % Cu-on-alumina, and 16.07 % Cu-on-alumina. The 
weight of a batch of catalyst approximated 1.0 gram 
in the particle size range of -12 to +24 mesh. Each 
catalyst was preconditioned in the same way using the 
procedure described in the general experimental pro- 
cedure of section 5.1. Before starting to obtain kin- 
etic data, this procedure included an initial heating 
period with nitrogen flow for about 3 hours, then 
maintaining the bed with nitrogen flow at the temper- 
ature of 290°C for a further 3 hours, finally followed 
by the reduction period with combined Ny and HS flow 
for another 3 hours. 

The feed mixture was usually composed of 3 


mole percent of both H.S and COS, 1.5 mole percent of 


2 


rusite xv Hoek 2 soda x08 work?. : 


fier a i setuor a Iss iy 


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SO., and the balance, N Keeping the feed composition 


2° 
and reactor inlet and outlet temperatures constant, 
only the feed flow rate was changed to get conversion- 
Space velocity data for each catalyst. After obtain- 
ing one data point at a fixed space velocity, it took 
about 3 hours to reach a new steady state at a dif- 
ferent space velocity. 

Here it should be noted that the size of 
catalyst particles used in this study averaged around 
5 times smaller than that for catalyst pellets used 
in the field plant, making the external surface per 
unit catalyst volume also about 5 times larger than 
that of the plant catalyst pellet. To compensate for 
this increased extdfnal surface area per unit volume 
of the catalyst bed, a space velocity of 5000 hr? 
would be expected to be reasonable to simulate the 
field plant space velocity of 1000 hea However, 
upon éonsidering the external transport resistances 
existing at a space velocity of 1000 hr} as shown in 
Chapter III, a higher space velocity appears to be 
desirable in comparing the catalytic activities of 


different catalysts since the true activities might be 


disguised by the effect of the external resistances. 


Furthermore, a better discrimination between activities 


of different catalysts may be obtained at higher space 


145 


velocity region as may be seen in Figure 24. Therefore, 


4 












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FRACTIONAL CONVERSION OF H5s OR COS 


Bose i 
Z ag lee @-=-= COS Conversion 


y-alumina @® y-alumina 

5.4% Cu-on-alumina 4 5.4% Cu-on-alumina 
12.98% Cu-on-alumina @ 12.08% Cu-on-alumina 
16.07% Cu-on-alumina 


v 16.07% Cu-on-alumina 





4 6 8 10 12 14 16 18 


SPACE TIME, GM CATALYST/SCFH OF HS OR COS 


FIGURE 24: SIMULTANEOUS CONVERSIONS USING DIFFERENT 
CATALYSTS 


146 


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147 


a space velocity ranging from 25,000 hr} to, 230,000 
hr was used throughout this study, which averages 
about 17 times larger than the plant operational 
condition. 

For each experimental run, the space time, 
W/F 


(or W/F ), was calculated and plotted versus 


H,S COS 
the fractional conversion of H,S (or COS) as shown 
in Figure 24. The computer program "MTBAL" presented 
in Appendix I was used in these calculations and data 
processing was initiated through the input terminal 
(teletype) connected to the IBM 1800 computer system. 
The conversion of H5S and COS in the simult- 
aneous reaction system, HS-SO, and COS-SO., is shown 
in Figure 24 for each catalyst tested. It can be 
oe that Ae Che eer ate is most active for the 
H,S-SO, 


active for the COS-SO., reaction. Figure 25 has been 


plotted from Figure 24 at fixed space-times of 8.0 


reaction while pure y-alumina (S-201) is most 


and 14.0 (gm-catalyst/SCFH of H,S or COS) to isolate 


2 
the effect of copper content on the conversion level 


of both H.S-SO. and COS-SO. reactions. It shows that 


2 Zz 2 
the maximum catalytic activity for the H,S-SO, reaction 
appears to be at a copper content of around 5%. Figure 
25 also shows that the addition of copper markedly re- 


duces the catalytic activity for the COS-SO., reaction. 


A clear explanation for this effect of copper content 























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SWORE * OL 








cos 


NVERSION OF Es OR 


FRACTIONAL CO 


148 





LeU 
= H,S Conversion =-=-- COS Conversion 
4eSPe id 
YD Sf oe ee fis 
0.9 esr: 8 4, ST, 3S 
ST: GM CAT/SCFH of H,S Once. 
0.8 
0.7§ 
0.6 
| \ 
o.5- x \ 
\ 
Nas 
0.4 Nora 
\ a 
ee \ on pol oe 
\ 
\ 
Ore ge i hag ca chi ag 
OSE 
0 


ae PU Bg 20.0 
COPPER CONTENT, WEIGHT PERCENT 


FIGURE 25: EFFECT OF COPPER CONTENT ON THE 
CONVERSION LEVEL 








nokexevnaD, FOR aa noletevaod 2H 
bi = 72 BY ; a F ¢ 

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809 1 2.4 Bo RDA\TAD MO. Te - 
onal ete 2 7 ee 


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149 


upon the catalytic activity for the two reactions is 
not available at this stage of the study. However, 

One may conjecture that the role of copper in the 
bifunctional catalyst may be explained by the 

donation of its 4s-electron to an electrophilic Lewis- 
acid site on the alumina surface. Furthermore, since 
the energy level of 3d-orbital is only slightly lower 
than ane ea of 4s-orbital in a copper atom, a 3d-electron 
of the copper atom may easily be attracted to H,S or 


SO, molecules promoting the adsorption of HS or SO, 


2 
on the copper surface. Then copper on the Lewis-acid 
site is presumably sulfided by HS and SO, to form Cus 
(which was recognized by the color change of the cat- 
alyst particles from deep blue to black). When the 
copper surface is sulfided, the catalytic activity 
becomes stabilized and reaction conversion steadies. 
This additional catalytic effect by the sulfided cop- 
per for H,S-SO, reaction enhances the H,S conversion 
level to some extent as shown in Figure 25 ata 
copper content below 5.4%. Up to 5.4% copper content, 
most of the copper may be sited on the Lewis-acid site 
as an electron donor, but when the copper content ex- 
ceeds about 5.4%, the basic sites may also be occupied 
by copper, but now as an electron acceptor. This 


could inhibit the H.S-SO., reaction since Liu(67) has 


2 2 


shown basic sites to be necessary for the reaction to 


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150 


proceed. 

The above hypothesis may be strengthened by 
considering the effect of the copper content on the 
COS-SO., reaction conversions. As the copper content 
increases for COS-SO,, reaction, the activity of the 
alumina catalyst decreases almost linearly up to 
roughly the same copper content of 5.4%. Then the 
activity becomes almost stabilized with increasingly 
larger copper contents up to a copper content of 
approximately 12% as shown in Figure 25. When the 
copper content exceeds 12%, the activity decreases 
again with the increasing copper content. From the 
above observations it may be assumed that excess copper 
could be sited on both the remaining Lewis-acid sites 
and basic sites when the copper content exceeds 12%. 

To explain the above observations, the adsorp- 
tion and reaction mechanism of HS, SO, and COS on the 
pure y-alumina surface as well as on the Cu-on-alumina 
catalyst surface should be discussed. 

On the pure y-alumina surface, both HS and 
sO, were found to be adsorbed on the Lewis-acid sites 
in the latest resultsiofgliu's studye (66).«: “For ithe 


adsorbed H.S and SO. to react, basic sites were vitally 


2 2 
important (67). Carbonyl sulfide was also found to be 
adsorbed on the Lewis-acid sites (19). Therefore, the 


Lewis-acid sites on the pure y-alumina surface may be 


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151 


competed by HS, SO, and COS making two reactions, 


2 
H,S-SO, and COS-SO,, competitive rather than independ- 
ent. From the above reasoning the pure y-alumina 
may not be considered to be a bifunctional catalyst 
for the above two reactions. However, the individual 
reaction conversion is substantially high for both of 
the reactions as shown in Figure 24. This may be due 
to the low Petes of Lewis-acid sites by reaction 
components at the reaction temperature of about 285°C. 
A COS adsorption mechanism on the alumina sur- 
face may be proposed on the basis of the balance bond 


formalism. Carbonyl sulfide is a hybrid of three 


resonance structures (32) as shown below: 


+6 


2) See Oeics ies” ze = 


Sire es > 


(I) (IT) PEEL) 


Although resonance structure I and II predominate (32), 
resonance structure III, in-which oxygen has a residual 
negative charge, is probably more favored to the Lewis- 
acid site than the other structures because the oxygen 
atom is more electronegative than the sulfur atom. 
Since carbonyl sulfide may be considered a 
stronger base than HS or SO,, it should be chemisorbed 
on the Lewis-acid site on the alumina surface, which 
has already been demonstrated by Liu (65). Chuang et 


al. (19) proposed a mechanism in which COS was adsorbed 


ee. 


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S 


8 
AN 7 Bu 3 


sxua eft poinogset evods end nowt. ‘38 a 7 


sinitomitid ¢ sd of b steblenos od sou yom 


‘ae ? cs eat 
orf ie S| - ion engl Jonker ows: svods ots 202 


, f ~ _ 





eid? 2S exept wi owods ‘ex enoljoser eit = © 
% aed _ 4 


; ; w 
- . 2 ee o. : , ; ” 
ya eetit Sios-siwed to ep steve 59 eh edt. oF ce 
3 . v2 
- 


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ae a te 
tt no metassiosm not Iqzoabs 205° A. re aN 
! at ¢ 
ad et to ssee8d oft to boaogozg ‘od {sor 208 


Sie an 


x 
~ ») 


rwofed nworte.es ¢ &t) soustourya soaée 
3 pgs y _ hia 


= 


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y : eR ET cae b : 
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me Pivee Met a a 
* a iar oe 


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Sa * 








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+erl> 


eee * 


bay 
g 
a 


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as i ia 


152 


on the Lewis-acid site with the sulfur atom on the 
alumina surface but it is more likely that the adsorp- 
tion of COS on the Lewis-acid site is also a hybrid 

of the two mechanisms; the sulfur atom on the Lewis- 
acid site or the oxygen atom on the Lewis acid site. 
When COS adsorption occurs according to the former 
mechanism, chemisorbed COS on the Lewis-acid site 

could readily react with a chemisorbed SO, or H,0 on 

a neighboring Lewis-acid site by a Langmuir-Hinshelwood 
dual-site mechanism and leave an adsorbed sulfur atom 
on the alumina surface. On the other hand if COS 
adsorption occurs according to the Latter mechanism, 
reversible adsorption of COS may occur without reaction 
_with neighboring SO. 
From the above analysis, an adsorption and 


reaction model may be proposed as follows: 

















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On the Cu-on-alumina catalyst, the mechanism 
of COS-SO, reaction may be the same as on the pure 
Y-alumina, but most of the H,S-SO. reaction may 
occur on the copper surface rather than on the alumina 
surface. Increasing the copper content affects both 
of the reactions as may be seen in Figure 25. Blocking 
of the Lewis-acid sites with copper atoms may reduce 
the number of adsorption sites for COS causing the 
activity to decrease for the COS-SO, reaction because 
Lewis-acid sites are critically important for this 
reaction. Therefore, the inhibiting effect of copper 
on the COS conversion level may be explained by the 
blocking of the Lewis-acid sites with copper even if 


the COS-SO. reaction could be bypassed by the faster 


2 


COS-H.,O reaction. The hydrolysis reaction of COS will 


2 


be discussed in more detail in a later section. 
On the other hand, adsorption and reaction 


of H.S and SO 


2 > may be enhanced by the sulfided copper 


154 


on the Cu-on-alumina catalyst. Therefore, the reaction 


conversion increases when the copper content increases 
to some extent, beyond which copper may occupy basic 
sites to suppress the reaction rate as shown in 
Figure 25. 

Therefore, it may be concluded, for the 
Cu-on-alumina catalyst, that the H,S-SO., reaction may 


proceed largely on the sulfided copper surface rather 


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than on the Lewis-acid sites of the alumina surface, 


while the COS-SO., reaction proceeds on the Lewis-acid 


2 
Sites according to the same mechanism as on the 
y-alumina catalyst; the Cu-on-alumina acts as a bi- 
functional catalyst for two reactions, H,S-SO, and 
COS-SO., which may be confirmed by the data in 

Figure 28. 

Unfortunately, however, the reason why sul- 
fided copper could promote the rate of H,S-SO, re- 
action is not fully understood in this study. Pre- 
amably it may be due to the change of the electronic 
configuration of copper making the accessibility of 
H.S and SO, easier. 


2 2 


‘ii) Comparison of Individual Reaction Rates 


During Simultaneous Reactions 


Individual reaction rates were compared for 
one selected catalyst, 5.4% Cu-on-alumina, by varying 
the space velocity. When water was introduced with 
the feed stream, distilled water was injected through 
the vaporizing coil by means of the water injection 
pump shown in Figure 19. The water injection pump 
was precalibrated through the procedure described in 
Appendix B. All of the other experimental procedures 
were identical to those described in (i), and the 


experimental results are shown in Figure 26. 


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FRACTIONAL CONVERSION OF HS OR COS 


156 


— HS Conversion Conversion 


i) H5s tECOS tuS0 +9C0Sat080 
oO H5S + SO 





2 
H,0 

+ H50 oe0 = 8, 
SO 


2 


6 «8 10 12 14 16 18 


SPACE TIME, GM CATALYST/SCFH OF H,S OR COS 


FIGURE 26: 


COMPARISON OF INDIVIDUAL REACTION RATE 
ON 5.4% CU-ON-ALUMINA CATALYST 


anf 






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7 racy 


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1) 


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pity 
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+ 


4+ 





. 


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Wel atl ell 


157 


A slight difference in the level of the HS 
conversion may be seen between curves A and E in 
Figure 26. This small difference in the conversions 
may be explained by the effect of the difference of 
the heat of reaction between the two reaction systems. 
In the reaction system E, the total heat of reaction 
is slightly larger than that for the reaction system 
A due to the combined Bhesct of both exothermic 


reactions, H,S-SO, and COS-SO The resulting 


2° 
Slightly higher bed temperature may cause a higher 
conversion of HS. If the above reasoning correctly 
explains the observed phenomena, a bifunctional activ- 
ity of this catalyst for both H,S-SO, and COS-SO, 
reactions is quite evident. This combined heat of 
reaction effect can also be seen in the reaction 
systems C and E'. A slightly higher conversion in 

the reactions system E' can be explained by its 
combined heats of reaction being larger than that 
from the reaction system C. 

From curves B and E' it may be concluded that 
the hydrolysis reaction of COS is inhibited by the 
presence of SO., which means that SO, competes with 
H,O for adsorption sites on the alumina surface. 


That is, H.O for SO.) adsorbed on the Lewis-acid 


2 
Sites reacts with COS adsorbed on the neighboring 


Lewis-acid sites. 




















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The conversion in the reaction system D in- 
dicates that no CO.-poisoning occurs in 5.4% Cu-on- 
alumina catalyst during COS-SO. reaction period, 


which is in contrast to the findings by Liu (65) 


for pure y-alumina in the absence of water vapor. The 


reason for this contrast is not yet evident but it 


may be due to the residual amount of water vapor pre- 


sumably present on the catalyst or to an independent 
catalytic role of copper for the COS-SO,, reaction. 
The effect of water content in the feed 
stream upon the rate of COS hydrolysis has been in- 
vestigated using pure y-alumina catalyst and the 
results are shown in Figure 27. When the water con- 
tent becomes higher than 1.5 mole percent of the 
feed stream, which is one-half the stoichiometric 
amount for 3.0 mole percent of COS content in the 
feed, the hydrolysis reaction appears to be zero 
order with respect to water. On the other hand, 
when the water content remains below one-half the 
stoichiometric amount, the reaction rate increases 
as the water content increases. For stoichiometric 
amount of water content, the results in this. study 
agree with other investigations (37,79) reporting 
that the hydrolysis of COS is of zero order with 
respect to water concentration for both y-alumina 


and Co-Mo-alumina catalysts. However, no published 


158 



























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data seems to be available for low concentrations 


of water in the feed. 


5.3.3 Performance Test on the Bifunctional Activity: 


To demonstrate more clearly the bifunctional 
characteristics of the catalyst for the simultaneous 


reactions, H,S-SO., and COS-SO a 12.08% Cu-on- 


2 - A 


alumina catalyst was used. While this composition 


is not an "optimal" composition for the bifunctional 


catalyst, reference to Figure 25 shows that the result 


will still apply to the "optimal" composition of 
about 5.4% Cu. 

The amount of catalyst tested was 1.0213 gm 
with a space time of around 6.0 gm-catalyst per SCFH 


of H.S (or COS). The observed results are shown in 


2 
Figure 28. The conversion levels A and AA represent 


the conversion of HS and COS respectively when the 


2 
feed mixture contains HS, SO, and COS in 68:3:1.5 


mole percent with the balance N Then, the COS feed 


3° 
line was cut off and the conversion level Bl was 
obtained for the single reaction between HS and SO,. 


After measuring the conversion level Bl for about 


one hour, the H,S feed line was closed and then the 


COS feed line was opened to make the COS-SO, reaction 


proceed without being interrupted by H,S-SO. reaction. 


160 


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#2)+ 20° + Coe 


coe. + bo 
SSE" SHFO°S 


#2 
eo 
Cn-o 


ect J 

~ q a 

4 We id 

ms ‘ a = -} ’ » ray v) “s 
“ 5 oe Yi = .- wv ; 

2 Oo &. ~ es E 

te r ss : r —- j 

a Po hed dd ¢ . 

om . - He} mm . 

() ¥ Yo a me ’ ~@ ee 

Nee "7 7 I 

tj } - 


2 
Mons 
wht 
a 


ae | 
aaa = e 
bay 
—% ‘ 
{ 
ie 
™m 
. 5 x, 
sS 
lp, 


EIYSeED LINE’ 


i4 ee Sa TAG 


— ' 


BILAMCLIOAYT GHYEYCA 


r 

© er 

3 ae : 

F a 
5 ee 


a 


RIGAKE 58 






The resulting conversion level of Cl was obtained. 

The conversion levels B2 and C2 were then obtained 

through the same procedure as was done to obtain Bl 
and Cl. 


The conversion levels of HS in A, Bl and B2 


2 
in Figure 28 are nearly the same, which means that 
the catalyst tested essentially acts as a befunction- 
al catalyst without interaction from the other reac- 
tion. The small difference in the conversion level 
between the reaction system A and Bl (or B2) since 
the combined heat of reaction in the reaction A is 
slightly higher that that in the reaction Bl (or B2) 
which may cause a slightly higher reaction rate. The 
same analysis may be applied to the conversion level 


2 


reaction, confirming the bifunctional characteristics 


AA and Cl (or C2), for COS conversion in the COS-SO 
of this catalyst. 


5.3.4 Maximum Obtainable Conversion Level: 


Pure y-alumina (S-201) was chosen to invest- 
eave the maximum obtainable conversion level for 
comparison with the theoretical thermodynamic 
equilibrium conversion level. The feed mixture con- 
sisted of 3 mole percent of HS, 1.5 mole percent of 
SO 


and the balance, N The amount of the catalyst 


2 rs 


162 


cor Tee eee 






















.benistdo wesw ID %o level noiexevAes 
* ~ eh Peery 


bantatdo ast sxsw SD bas se elovel ¢ 


lad aisddo of adtob asw 25 sivbso07g omens eae 
ww “i i 


as ae 
cA bs If A ak Ook Bo alevel sobazevnos edt | 
tsit ensem dotdw + SMBS odd yitses ors as coset ad 
-noiscautoed-s o5 eto yitsisngzes bevees Jeylesso on 
~snes sades aie moxd noispersint ssc: setesnnts G ~# 
f evel nolatevaeo et mi sonsxe22tb. {Leme edt s es _ 
eoaie (Sa yo) fa bas A meteye molsasey eit waa i oe 
ei & nolgossa sit wi noktoser 20. sed benidmoo 
(S& 10) L@ noktosex oft ni tend send “petyta viomentn 
ad? .ats2 nottossx Yara ylsipife 6 sagen yam 
tavel ‘nolexevmoo odd os heitean ou 1 aaa 
<02-209 ant ‘mk norexsertes 20D: 0% <9 ead 1 
eoisteixzssosisiio iscoi sauté ont prtks . 


~seoevnat of nseofd Baw 1 (B88) 
got fount noketevacs ef ' 


o Lenina tnokse: 


_ TOD cadet nit f anys 


“te sheoxeq oom m Bot saet 


i D fo 
=. Bo ads 20 
Tg pa hae 








163 


charged was 35 gram packed in a bed depth of roughly 
6 inches. The feed temperature to the catalyst bed 
and the reactor outlet temperature were kept the same 


and varied from 550°K up to 700°K. 


Before examining the conversion data as a 
function of the reactor outlet temperature, the 
relationship between conversion and the space velocity 
was examined at a fixed reactor outlet temperature of 
700°K. From the results shown in Figure 29, it can 
be seen that in the region below the space velocity 
of.20 hr (or above the space time of 0.05 hour) the 
conversion has almost leveled off but is still rising 
slowly with decreasing space velocity (or increasing 
Space-time). The asymptotic experimental conversion 
level may be found to be 89.5% at 700°K in Figure 29, 
while the thermodynamic equilibrium conversion is 
TP eke, 

In additional experimental runs, the space 
velocities of 100 hr t and 4 hr + were fixed to 
examine the upper limit of the obtainable conversion 
level in the Claus reaction. The resulting experi- 
mental data are plotted in Figure 30 where they may 
be compared with the computed thermodynamic equil- 
ibrium conversions by Gamson and Elkins (35), and 


McGregor (72) as well as with the experimental data 


Lae) 


atéseront wo) ¥ yt Lsotev SOBgR: poleserseb Asin vine, 


af oupid mi MOOT, FH PE. e8 ack of bawo® ies Ysa. 
























+ to diqeb bed's ni betosq: mexe 26 ad ™ 


tjevistao eft of owIst aqtead boot 9AT: 4 shin oe 
it : vik - ‘ah 
m it tqet oyew sivtsisqned seisuo? zo+¢ oBet ett baa 


2°00) oF ay R°ORe moxt belxay ee y 


FN i. 
st6b cobereviap odd priks 5 mB xe exoted — 7 % 


; al 
nat. toldsvo dada ot to epksoni® 
— soos 


eosge sit-bos 20 buxe VaOD: - oawied qidenolveles a 


Luts reqses Seitwo rososet ‘boxi? ry ey boniasxe ae of i 


°S esupit md wore, ad! [waex odd mox a yi 
i i+ ’ Bs 


: , Ps P es ze 
olay sasde on voled..aoipes ont mi 3 oie esd 
om ! a id [- oa 
gon @9.0 to omit al eid ovods * to). ts jhe 
_— 190 
six Llige ed tud tite bolever teonts anil, nolexsva0o - 


, ss me i Toe i -e 
reviavol, Ladaemiteqxe o ivorquyss ais 7 “998q 


ei sotetevnod: murinds ftigee oinanybomeds » 43 9. 


ud Vi ® 7 a 

re pea néataacs 
b il i 
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ey 


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bre . (2) acids bas a wh 






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re: Rite ae ; 





164 


AWIL dOWdS dO NOILONNA W SW THAT NOISUYAANOOD 


UNOH ‘AWIL dOWdS 
ie 8 c.0 


Cos 


rT 
Oe s¢H 
BoC Cx 


(FUS979q STOW) NOILISOdWOO daa 


BHU OPpL : Sssgdd 
MoOOL : dWaa 


°6¢c dadNdSId 





s°H JO NOISUZANOD IWNOILOWUA 


(oU 


be 


4 


COMARve Ton 


y 


CRAET We Y LANGLIO 


RIGAHBE 3Sd* 


SbYCE LIvE* HOnE 


o°3 


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a 4 
tu 
' 

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& 


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aks. 
« 
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i pare er oe a 
ee ea 









1.0 
0.9 
x< 
” 
2 
7, 0.8 
x Equilibrium Conversion 
6 A: So, S,, Sg mixture formed 
S B+: Sg only formed 
F C+ Sg only formed 
> 
é Experimental Conversion 
0.7 Gamson & Elkins, Bauxite 
° SV= 240 hr! ov 
° 
ee Current Work 
5 
ta SV= 100 hr' © 
SV= 4 hr! Oo 
Inlet Composition (mole pct, 
0.6 HS °3.0 
SO, Abe 
N2 95.5 
0.5 


400 500 


600 


700 


Reactor Outlet Temperature, °K 


FIGURE 30: COMPARISON BETWEEN 


PREDICTED 





800 


AND EXPERI- 


MENTAL CONVERSIONS AS A FUNCTION OF 
REACTOR OUTLET TEMPERATURE 


165 










+ 


proce ag psi ' a ; ¥ 


pee an Pa ce 


bemint no ge * 
. coy 
eS hoisevag > tote ning 


atinbes oe & 
a vat oat ma = 


— 





166 


by Gamson and Elkins (35). 

In Figure 30 curve C represents the thermo- 
dynamic equilibrium conversion at different reaction 
temperatures computed on the assumption that the pro- 
duct sulfur exists only as Ser while curve B shows 
the computed results based upon the product sulfur 
in the form of S,. On the other hand, with an 


6 


equilibrium distribution between S S, and S, species 


7) aly» 8 
the equilibrium conversion curve A was predicted. 

In these experimental runs it required more 
than 24 hours to reach a steady-state conversion 
level. It was found that the smaller the space vel- 
ocity the longer the time to reach a steady state. 

In. Figure 30 the data point (1) was obtained after 

24 hours while the data point (2) was obtained after 
72 hours since the flow condition and the reactor 
temperature had been stabilized. The two data points 
show essentially the same conversion indicating that 
the steady-state condition had been reached in about 
24 hours. 

Gieetnpinectad but very interesting results 
were obtained in this experimental run as may be seen 
in Figure 30. The experimental conversion level at 
the space velocity of 100 hr! is well above the 


equilibrium conversion level for all the temperature 

















. (28). enitfa bas owned | 


fqei D> evuun OF omwpl? mi 5 i ; 
ny , nl 


tg s 


i. 
7 
7 


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mo slidw ..2 as vine eteixe wi ive foub ioe: 

g AR sia 

ivtine touborg odd soge bered & tigaet besugmon ott t a 
r-,. f =f | a) 


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itiw ,Bnsed xsrso ant es > 32-20: x08 ond ax | - 


Ad ) 

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‘ oft a) an os a. 

eni I[samemixsdxe eed? ai  — > 


oe 


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€} a : : 5 


2 
; 
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VHOD 


ste=yhseIe 6 is 29% os exbor os adit = 
bie ST. me 

-{[ev sosqe oft tsilame sit aaa bao esw roo foveal 

ee 7 3, 

.otste ybsote « dossto3 ‘mt, eds sepnol: att ws [DO ie 

' a ag ; a 


ts beoatletdo: eew’ (£) dk og “6386 baal ce ex ia 
4 


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a 


a ii 
‘ylits sare 
24 

7 LSOCGE ii herit 557 xood Bui as 1s tbmoo vee 


Jett polisobbat. acter evo ont or 








167 


range between 550° and 700°K. The overshooting of 
the experimental conversions beyond those for the 
equilibrium curve is even higher in the case of 


af 


the space velocity of 4 hr ~ than in 100 hr, These 


results are compared with those Gamson and Elkins 
(35) obtained at a space velocity of 240 HE bs 

This large discrepancy between the experimen- 
tal and theoretical equilibrium conversions cannot 
be fully explained at this stage. However, the 
following suggestions might be pertinent. 

First, possible inaccuracies in temperature 
measurement may eeeeit a misplotting of the data 
points above the equilibrium conversion curve. 
However, this possibility is very small considering 
the well-equipped experimental apparatus and measur- 
ing devices uséed in this study in contrast with the 
rather large discrepancies between the experimental 
and the equilibrium conversion level. Actually it 
may be seen in Figure 30 that an incorrect measure- 
ment of the reactor outlet temperature of 50°K higher 
than the true temperature, is required in shifting 
the equilibrium curve to the experimental curve for 
the space velocity of 100 hrt, Furthermore, for 


the space velocity of 4 hrf, a temperature deviation 


larger than 80°K is needed to shift the equilibrium 
















<25SV0 of 
me aoa 
= bros MoLey LSvnoy tognoetrogee a 


9ES9 eft ni serpin tevs et evs nwixdi tte 


oa 


mm COL @i'sisds ~ xi fo yo inolsv- eons wi: 


“y a Py cy) A 
[ntXfR San nosmeD sacs AsEW Bex peo 915 ei lueex 2 : 


13 5 


a 
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qx® ofs meewted yonsaqetoa if ephal aia ‘ 


anotatevaboe muixdi Lippe Isoty texosift bas hoa 


a 


yawou' .épate ebie #5 bait nique’ vite? on 


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I ug fe 
-f 


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sni eidiseoq Jerr © 


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stebh- ett do prisxoliqain at Jipest visio soomo ies 


OVID nolaTevaidd ah i Lip pe ort evods asin, 5 
poiirshLerem [iawe yisv aby si ttdieaog aide <re¥ PwC ie 
tuesem boas eetetsaqas Lésrem brsgxe Deaainne-th mi - ott 


sit doiw sasaendoo i's youde sind ot. -~ 


* 


isinomizeqte sift 26 sei eston: st 


svendpid A‘ d2 0 axusaxsemied tolono.* 


pristine mi beriupsr ai oe 
4 







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nolaes iveb ovata 


‘Pe 





curve to the experimental one. These large deviations 
in temperature measurement are quite improbable even 
after considering the effect of radiative heat trans- 
fer from the reactor wall to the thermocouple. 
Secondly, the higher experimental conversion 
could imply that a shift of the theoretical equili- 
brium conversion to the forward direction of the 
reaction (1.2) is occurring. This shift could result 
from a lower sulfur partial pressure on the surface 
of the catalyst pores than that in the bulk gas 
stream. This lower partial pressure of the sulfur 
vapor on the catalyst surface might be caused by 
polymerization of the product sulfur on the catalyst 
surface forming new species with longer chains like 


S S and so on. This polymerization 


Pro’ =50" F100 
phenomena on the alumina surface, while speculative, 
might be one of the characteristic properties of sul- 
fur species distribution. In such circumstances, 
additional reaction beyond the expected equilibrium 
conversion for the condition of the bulk gas stream 
could occur within the pore. 

| Thirdly, the experimental conversion level 
could be higher than the equilibrium conversion level 


if the data were taken before the reaction system 


reached a steady-state. This seems to be due to the 


168 



























ee oa en | ae! + 
exh » od Oe a ae 
+a rat be » 
yo aE a) 
; Sh et 
ib spxsi seoentT .oiTo isso tog art 0: 3 BO ei 
- ’ by 
| ni eee 5 a2 ae 
nove sldedoigmk stinp ois Jasmeiesom owt steqmet me “0 
’ cS - ani’, = Ve 
fas aaa Ma 
ansit tsad evigsiisy to tootis tid pnixebhenog: iain : 








be 
ie 
iquopomreds. eft oF Iisw  10tosex ‘ods, wont x08 Ee 
rae Pela rae fad 
[sinomitedxe sania tol ects s¥ibq0098 er aes it oh 
(soitatoans oft to Jiide s: sectt fqmt bie Pr 

it to notsoanib Bbuswre2 sold oo ‘noterevnos myixd LY 

ra = 

jineet bivyoo titsdd Skat. . pols rLI30 ak (s. ty otsones | 
— Eih< 
Quue sit no euveeekg iséousq ins tenae xewol 5 moxt | 


be ae? 


- af 2 t 


esp Aled ait ab sens. aefit } aexog ‘deyledeo ons 
viive edt to esvatetq Lsiizeq towol “pide cen! 


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teyletso ad? no 3 “sot hox gq sini to coivestiomytog 


Ef. ect: Bild YsSpaTo f, i rt a ith iY ws ie IRC . won pninivel 2 298 ous : 
iaieceall tyom; foa SLit TO on UMA: be 5 - wy ve ) 
vm x Sai ae = AJ J ” _ gor : Sal re 7* 


on 
© 


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evissiuooue oi.titw posting: Baba le, ast oi 


Me 


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aSQnesequiscrls Ape at so nsudaneaty 


wes s 
win’ xdi.Ltupe boineqxs oad buoyed: seine | 


ste ie tarde 
of eacre ie a. ee 
mseite asp alu ot te noidtbnoe s : noierse' 


* 


‘ -s070g ¢ 
lave noiaxewnod faccepin 


syel noiexevaes mu tdi thu 
4) vies a 7 
ale aokseemt.oe 8x0 od: | 
gt bod of. ae 2 ahd? ed 
bf 7 r-: ‘eae. nie hth nS 


ry 


169 


non-equilibrated adsorption and desorption of sulfur 
species in the catalyst pore (55) before the partial 
pressure of sulfur vapor could reach a saturation 
point which could result in capillary condensation. 

Another possibility, but one which seems to 
be improbable, is that the thermodynamic data used 
in the calculation of the sulfur species equilibrium 
distribution are incorrect. 

It should be emphasized here that Figure 30 
shows that the conversion level decreases as the 
temperature increases, which means that reverse re- 
action in the H,S-SO, reaction is occurring. Unfor- 
tunately, however, this reverse reaction has not been 
isolated for study to date. With the type of the 
equipment used in this report, it was not possible 


to evaluate more clearly the influence of the 


reverse reaction. 


‘gexevox ted2 sieen doiriw ane asoroak ‘gus 619 ase 


. 4 : od bee ie * oh 













te 
to moktqwereh bas nota qieebs boserdis Lupe=n 


it. gxohed (22) stog ga giagao. edd at solsegs | 
7 , ok ae a) 
geo & Coser bleos 10G BY WL ne Ro ouneess 


csshiuen yrsiliqss ms tLUe 2O% bl: SOD. ae said 
Ai ii F 


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"2s 


74 


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gos1T08Dn ons ootindiaaens 


‘ae 5 : = 


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~ 
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n says otiy dake eaeb oe bute sot paisa oat 
om gon caw ti ,oroges eins “ab beep saa 


a. 






‘ett Bo gorau s hbo Eve) Sno to sao 


De 


CHAPTER VI 


CONCLUSIONS AND RECOMMENDATIONS 


Performance of Equipment 


The Seen EAL apparatus used in this study 
encountered difficulties for the first few months, but 
after various repairs it functioned fairly well in most 
respects. Feed and reactor pressure controls, total 
feed flow-rate control, and temperature control of the 
preheater operated satisfactorily. Residual problems 
still existed in the feed composition control, GC analy- 
sis system and reactor temperature control. The feed 
composition fluctuated severely with an average devia- 
tion of +3% due to malfunctioning of the gas cylinder 
pressure regulator and the flow controller. 

| In the GC analysis system, base-line adjustment 
within the Infotronic digital integrator was difficult. 
Because the SO,-peak area waS very sensitive to these 
base-line changes, considerable error in the GC analysis 
for this component resulted. Therefore, consistency 


between conversions based upon H,S (or COS) and those 


2 
upon SO, could not be expected. 
The manual control scheme for the reactor inlet 


and out temperatures was inconvenient and inaccurate with 


a: control error of around i3°K. 


170 






















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Se. & an “ to 
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. . YY ae 


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alorinos stgeaeetg 3 989% bas: best Soaaet! 


L “ 


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eldote issbieei 4. yEasosast cline beaereq9 rosnode: 


(r 7 


tiLaoqmoo beet sds ‘ut besaixe Fier. 


ae 
as 


’ ~~ t o~ ~ r 
"Vib ts d . ‘oc. TIMOsD mou 


soot eT , oxd 10d oxusBseumed “x 1932602 bas meson ais 
~piveb & pa TSVe as dziw yie1sevs we es pedasioet? 
tebaiiyvo 2#Bp es So pakaokt soutien 9. ooh en 
. eoife (3 100° wolt: ait bak soseiuees © ° 
rremfaurbs enil-stad  eseye ekeytane, 99: of 


: wa ope 4, 


.3ivoll32b esw oss spotat fettpth eccamiuneatl 

. ree 

ia < x + a ae 

efit ot avisianoe ¥ rov 26w =e 
} te ty 


sieyisns DO eft of some sidexebiones 





Leaoco orotasest | bea 


searit iat 200 20) agli doge 


oo y : i Pa 
i gr = ig ' 
n I 


ate) 


oe zotoset oto x02 







hay ps 


To obtain more precise control of the reactor 
pressure, an absolute pressure transducer should be 
used instead of the gauge pressure transducer used in 
this study. 

After initiating some improvements, GC monitoring 
by the computer was found to be consistent with results 
from both the digital integrator and a disc integrator. 
Thus, development of an on-line data processing system 
becomes possible after devising an automatic sample 
injection scheme to get accurate elution time data for 
each peak. In addition, a new gas Pe chad both equipped 
with a temperature programmer will be very helpful to 
shorten the cycle of each analysis, which is 12 minutes 


at present. 
Prediction of a Claus Unit Performance 


The calculation of equilibrium conversions for a 
Claus unit revealed that the reaction conversion may be 
increased or decreased by changing the inert gas content 
depending upon the temperature level. When the equili- 
brium temperature is above 900°K, the conversion level 
increases with increasing inert content, while the 
conversion level decreases with increasing inert content 
in the temperature range below 750°K. For the operating 
temperature range between 750° and 900°K, the conversion 
level of the Claus reaction depends upon both the opera- 


ting temperature and the inert content. Most Claus 


pone 

“ 

ten 
# 








7 2 , ake a oa - 
risosiaom DD ,sdnemevougmi emoe patésisint 1933A" ' "Sea 


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mosere pateesvorg sisb siil-no' ts Fo: Snanqoleveb vent 

elqmse oijismosus ns paisived resis sidteeog eomosed 
ro% sisb emis notsuis siativoos top 08 omertoe notdostnt — wy 


beqaivps dqstpotsmotdo esp wens .a0i3 fbbB at seq fone” 


‘AG 















ar 


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~- 


gojunim SI at dotdw ,elaylaas fone. 20 eloys ee ; 
sd senor 98 


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i i z 2 : ae 
Sonsmro2tes 3 £a0 11D 5 2 no 29 bext vs 


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ni a 


2 Ane ims a \ 


"i 


ad ys nolezevneo notioess off sate | 





tusjnoo as~p szenl sd2 paipisrio yd 8 
~ilispe end nodW .isvel owwsesogmed ee 
Level. nolerevaoo Ey: ey vA? 008 aca a om 
atid of title 3 mas ao9 sxeat p ents an a a 
inetaoo ti9eni paiesorsad stn a ae ee a) 

Miers oo ee ae 


ert Rahs shih i a 4 i 7 a 


bd od id entacterahite BR aay e 


a no ‘ es . ‘ " 
ss) Pas fate apes mE eORL Mot oF 8 fis 
> u Yee te pee - aa ave sist ripe a ok it ‘ at a . { ; ’ dh 





> | 
Py, e = 


172 


plant furnaces or converters operate at temperatures 
outside this middle range. 

In the calculation of the adiabatic reaction 
path in the front-end burner it was assumed that the 
reaction (1.1) and (1.2) occurred consecutively and 
- reached an equilibrium condition. However, the actual 
condition in the front-end burner may not be in equili- 
brium condition. Therefore, accurate reaction kinetics 
combined with mass transfer rate in the front-end burner 
is required to understand the actual dynamic behavior 
of the reaction path. The literature suggests that the 
mixing efficiency at the entrance of the frontend burner 
should also be experimentally investigated for various 
geometrical configurations since the mass transfer rate 
depends largely upon the mixing efficiency. 

From the simulation of a Claus catalytic converter 
employing the one-dimensional two-phase model (62) and 
Liu's rate expression for reaction (1.2) for an Alon 
catalyst (66), the rate of reaction has been found to 
be sufficiently fast to proceed significantly right at 
the entrance of the catalyst bed. The maximum obtainable 
conversion may be reached within 2 feet bed-depth when 
the feed temperature is 550°K and the space velocity is 


1000 hr +, ata space velocity of 1000 hr 


and a feed 
temperature of 550°K, the external mass transfer resis- 


tance was found to be negligible compared to the 


auoiisy 1023 Bsa: peisesy nk “isdnomireqte ed. cele bis vu 


ol 1S 
si istens1? aesm ois sonte anoktsxupitnos —_— 
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internal resistance, while the external heat transfer 
resistance was found to be significant compared to the 
internal resistance. On the above results, it is re- 
commended that the space velocity should be larger than 


1000 hr 


and the catalyst particle size should be less 
than 1/8 inch in diameter, which was employed for this 
Simulation, to eliminate the effect of the external and 
internal transport resistance in investigating the 
intrinsic rate expression or in evaluating the catalytic 
activity in the laboratory reactor. In using smaller 
catalyst particles, the observed rate per unit weight of 
catalysts will increase because of the larger external 
area. Also it was observed in experiments that the per- 
formance of various catalysts could be best compared at 
lower conversion levels. For this reason, a much larger 
space velocity between 25,000 and 150,000 he voaand 
smaller catalyst particle sizing -12 to +24 mesh, were 
used throughout this study. 

To extend the applicability of the reactor simu- 
lation to Alon catalysts with different physical proper- 
ties, the effectiveness factor-Thiele modulus relation- 
ship was computed. The calculated value of the effec- 
tiveness factor, 0.17, indicated that the rate of the 
Claus reaction is so fast even at such a low concentra- 
tion of the feed reactants, 6.14 percent of H,S and 3.07 
with the balance N., and H 


percent of SO O, that most 


P z 2 


of the reaction occurred in the thin shell near the 


. ’ yi) 
| 
; we. =f 
my 42 
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4 
























} ad [ecxeoxa sdst Lkdw. yoonst ates: samses . ice R 


hig BAS 


. eqmon sosditinpie ed ot bavor 2ew sonsdeiee “a 
— inal » ¢ he r ce avi AT C34 Ae ey: ‘4 no iia < a) tear temredai oe Th 
> . ‘. , SHA i? 7S , 7 ¥ - 4 7 


ee 

: , 4 

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“ - : pe a ae 

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. “| rh 

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. ; =H 


+ b + . 2 


x 


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xo oft to. dvette ong a3 Fanimsts od, ssoksetuaie 
. ; : wn a 
nitepiteoval ai sonsteiesr +1008 ns3t, : iemxetah E 


’ 


—— ited ov thers fe a — notes ; 46% pps: 


P 
: 4. Wd J ah tre = ‘ . - ‘ 
oates of ,2otosex yrosse7 xode. ‘aid . al. whan § 
\ » ” ; ‘* Ly. 
Z jias.2eq sdar Devise eda ans seioirxba sayisdso f * 
:9txe taprat ort to sawsoed a peerodl {itw exmyinaso 
, Bi ‘E a “ee 
edd tends eddemiteqxe Ak Bon viel asw si sla ae 
; z es saclay a 
BR Ox zad { bivweo ave vi 636 Span daey 20: 904 - ae 
teprsl wm & ,fonsaer €4 le 10% eiever “ao denon wos 
im of ODO. 08: : qa ; 
bin - tod OOO. oet = 90 ‘ es noswaed: winoley & oeBqaz 
: 5 
‘ } ; 389 o an ¥ 
a1 9% eon bS4+ ot Si+ patece sloiteg daylesso rolisme 
ersew 4 , a fhe La he ‘€ ’ 
YOEDS: ehta — ox bet 
rer, ~~ e. is i 
t totosex oft 20 -y bidentiage: ‘ont nian 
; ape a! 
-xeqo1q isofayigq, 2 mexs32ib Aas iw atevieden o 


ob 


“mobo Lor eutubort oteidt=10908? 98 
rosette ent to, sulev: Davai tao: wt 
erly 30, 18% ott aed bento nbes ied 
dome 2s a9ve 


i i =iee , 
(Ne Baer. eee wires: on 


- 


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os ain : rd 
oe Pt Ete 








cam ae! 


ne “e , . ss 1 
napred: bib Ky ottaney 
Ye ‘ rig a = : é 
m +7 ‘eo , a et ba v7 ok att 


= 


174 


external surface of the catalyst particle. This result 
agrees with other observations (45, 72). 

However, for a more practical simulation of the 
Claus converter, it may be necessary to consider the 
effect of simultaneous reactions, like COS-SO, and 
COS-H.0, which possibly occur in the Claus converter. 
Here the need for knowledge of the reaction kinetics of 
such pertinent minor reactions, including the reverse 
reaction in reaction (1.2), arises. In addition, phy- 
sical properties of various kinds of catalyst particles, 
like effective diffusivity, effective thermal conducti- 
vity or pore size distributions, should be critically 
investigated to provide an improved basis for prediction 
of actual performance of a Claus converter. 
Evaluation of a Bifunctional Catalyst 

The bifunctional activity of the newly developed 
catalyst, Cu-on-alumina, was proven experimentally. The 
copper on the y-alumina surface was shown to improve the 
H,S = SO. reaction rate to some extent but with some 
deterioration of the simultaneous COS - so, reaction 
rate relative to the y-alumina used in preparing this 
catalyst. The optimum content of copper for maximum 
reaction conversion of the H,S - SO, reaction was found 
to be somewhere around 5%. Possible adsorption and 
reaction mechanisms for the two reactions, H,S - SO. and 


Cos - SO,, on both the y-alumina and the Cu-on-alumina 


~ 


es, ss PAL 3 


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i 


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‘a 
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ie 
, 


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F . 


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yifsottiio- od bLorore ano duds ak exte etog “to ytty : 
noltothaxq tot eiaed bevorgit iis ebiverq a siihsobialies 
vetrevnes adsl) 6 te en 12H 


begolsveh Bei. to eriv ven tanoksodue nett | aha 
OAT 4S issnomineges MOVOR aew eckme ts -me-wo": r} Lats: 
et avoxqmi o3 awoke cow sostxue wnkmulomy arid oo 
emoz dtiw tue taetKe ennioie os Pere noon ee 


noitosex (02 = 209, avosnad Lumis ett to 3610 
eidd paizeqer¢ at beew psi 


ie ea: 


mumixem 1oOt alc to SotecnerD = . Be I 
basol ssw notsoeer g02 - age dia a. 
% Brie cols qroebe scitenos saree toc 


Bie a i i 


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Mp : 


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Mi 


te tt 





¥ 7 >» a 
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: yi iy . 


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175 


catalyst were discussed in terms of the experimental 
data observed and the latest mechanistic model deve- 


loped by Liu (66) for the H,S - SO, reaction on 


2 2 

y-alumina. 
Summarizing, on pure y-alumina catalyst, HoS, 

COS and SO, individually appear to compete for the Lewis- 
acid sites. To account for bifunctional activity on the 
Cu-on-alumina catalyst, H,S and sO, are believed to be 
adsorbed on the sulfided copper surface more easily than 
on the Lewis-acid sites, while COS may be adsorbed on 
the Lewis-acid sites in the same mechanism as on the 
y-alumina. However, adsorbed HS and SO, on the sul- 
fided copper surface are believed to react more readily 
in the presence of neighboring basic sites. The com- 
plementary role of the basic sites for the HS - SO, 
reaction on yY-alumina, which is to stretch the H-S bond 
by the electrostatic attractive force between the basic 
oxide ion and the H-atom of H,S (66), may be similar on 
the Cu-on-alumina surface. The decreasing conversion 
of the H,S - SO, reaction beyond the copper content of 
5% may be explained by the blocking effect of copper on 
the basic sites. Therefore, the basic sites appear to 
be vitally involved in the HS - SO, reaction mechanism 
on both the Y-alumina and Cu-on-alumina catalysts. On 
the above postulations, it may be concluded that, on the 


Cu-on-alumina catalyst surface, the HS - SO. reaction 


= ¢ Eb : r 3S 
avd Fate gs 
; ae wie, Le : ‘ieee 
1 r t 1 

























Lets LOMLAGRO ais to eed: fut bezeupeib, o1eW 
-svab lebom aiteinsdosm Jeeie. ott brs ‘bovxoe 

no coltoser. 608 + SoH ons or (83) wld vd ianes: ar 
vetoes a 

Bgl seyistBo sainula~y eum nO, vent sdxsmms’ , 
-eiwed: eddon0% eteamos oF 1890q8 yiLeubiwtbnt gO, bas. 202 
eft no ytivisos isnottoautia 203) dquosoB) OR. sexie bios — o . 
ed ot bevailod ots 02 bas Soh. .deyttso sctimuts- toms 1 * 
msd3 yLiess stom enazaue ateqgeo bebiiive eds 10 bedzoabs — 
no bedxoebs ed yam 209 oLsdw saatie bibs-elwod eis 20 4 * 
eit ao es matnedoom ampe ety at potte bins-ekwed. wt © . 
-{ue edt a0 oo aos SH bedsoshe «tovewoH satay a 
ylibser siom Jongx OF psvalied exs. eoataue 49qgoo, bebe 49 
-moo saT motia skeed poisodde len ie conshera, et me, 
its Es od ons 102. eeste oissd ott: 0 te * nol 
“olesd ons neews ed eorok: ovisos3336 stdesuonsooso 
no wsliste od yam 4 (8d), Agi Re moan ed Bis i pat iY 
no kLaeusvnoo pitessx99b ont ssosisua 5 pica 4 : fn 8 5D os t= 

te tnetnoo zsggoo eng baayed. soltoeet % ee a hi 
no s9qqgoo to dost2s pattoold¢ me aott 


ot ~seqqs satke otasd ord « 


ey 


wt 


i 2 
d 


metasdoom nolsosen 08 ~ oe 
no Rae manner e gi; sibrsirs: 
Th a Mr AS 


176 


proceeds mainly on the sulfided copper surface in the 
presence of neighboring basic sites, while the COS - 
So, reaction proceeds exclusively on the Lewis-acid 
sites. 

To confirm the above explanation, an infrared 
spectroscopic study on the adsorption mechanism of 
HS, SO. and COS on Cu-on-alumina catalyst is recom- 


mended. 


' Maximum Obtainable Conversion in the Claus Reaction 


The maximum obtainable conversion level of the 
Claus react eras investigated and found to be above 
the theoretical thermodynamic equilibrium conversion 
level for the same temperature, pressure and feed com- 
position. This highly unexpected result and some 
possible explanations were discussed. No clear evidence 
is available ecepeeeent to explain this discrepancy. 
To test whether surface adsorption equilibria for 
elemental sulfur are responsible, it is recommended that 
a long-term operation (for around one week) of the 
reactor be used to provide enough time to ensure that 
such equilibrium states are reached for the adsorption- 
desorption phenomena of sulfur vapor within the catalyst 


pores. 


' Reversible Reaction in the Claus Reaction 


During the investigation of the maximum obtainable 


conversion level in the Claus reaction, it was noted on 


“7 


tnceat auyelD offs as taxewno > aids picieniinddo carded ' “4 


revaop myixrdiit.ups oimectybonrads “Isdisoxoedd © bist 
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sqexpelb, ébdd nisiqne os jnaeexg aes 


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a 
siete ee 


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- 
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vo aN/ 4 BD v i ee BA bei u ‘ 


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. | 4 bi 


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177 


figure 30 that the conversion level decreased as the 
temperature increased from 550° up to 700°K. From this 
observation, it may be concluded that the reverse 
reaction in the Claus reaction is significant even 
though its visible rate is negligible. Therefore, it 
is recommended that the reverse reaction kinetics for 


the Claus reaction should be investigated. 


et! a | ea 














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¢ a , , a a a 2 7 
’ A " 7 
a «a moe 
J ; 


i. a tigre: Pree 
a ahd 


iP « 
if “ 


Symbol 


Ayr Ag 


, eee 


Q 


NOMENCLATURE 
Meaning 
constant estimated by curve-fitting con- 


version-space velocity data 


external wall area of catalyst bed 
(cm”/bed volume) 


heat transfer area (cm*/bed volume) 
mass transfer area (cm*/bed volume) 


concentration in the fluid phase 
(gmole/m1) 


concentration within the catalyst pellet 
(gmole/m1) 


concentration on the external surface of 
the catalyst (gmole/ml1) 


heat capacity of the fluid phase 
(cal/gmole «+ °K) 


radiation constant 

molecular Mriieivity (am= /sec) 

bulk diffusivity (cm*/sec) 

effective diffusivity (cm/sec) 
Knudsen diffusivity (cm*/sec) 
diffusivity in the macropore (em?/sec) 
diffusivity in the micropore (cm*/sec) 
diameter of the catalyst pellet (cm) 


volumetric flow rate of the fluid stream 
feu tt. 7nr) 


volumetric flow rate of the fluid stream 
icin tt. /br) 


178 




















TAVTAIOMSMOM 


' painsseM 


” , : M ’ ; es 


-10> poistii+evans vd fosamizas gnstanod ° ~~ oie a Aighed 
63 6h apes sonra am Leaey 


Lt a a aid 
bed teyisisao to seis (lew Jan sxe CL ye dc ee 
ee (omy Lov bod\’ ey ee , 


(om Lov bad\! Sib) 5978 sotanst? ghet. i... Ya 
(smuiov bed\ Smo) 5915 <AARAB SS ees i oe me 


esesig Styit edt al aotistIneonos . Hf 

(fa\elomp) — 2 iM 

Jelieq seylsis> ont subsid Ew nolsstsnepnd> ‘ 
esl irk 


to eostave Learadne eit mo noisstgnsonos Lita 
(Lnr\veLomp) - ‘da¢lateo ers bas 


ean Siwv!2 od 20 yWineqen teed. Ree 
(Res olomp\ iso) > = he 
2? Sw y 


- tnstenod: acizstbsx ¥ 4 pase t 4 

(o@e\"wa) ystvieuttib vatooston™ | 

(oea\S mo} vikvieutih hi 

| (sea So) yoivieuttib ovisostts 

(pe8\* mo) ¢itvienttth, noeboan Shaan S 

(oe8\ sao) i a isa orig at vetvie ite 5 an a. > 
(se8\* m3) siacigniier ons mt ytiv 

(mo) dolLeg. Jeyletso: orth. i sé: 


ame Hius2 ong 20 e203 volt 9. 


Terns 
te . v4 
or he 
or" tae es rete t 
j : faa 





co i} 


i 
¢ s 


COS 


AH 


wn 


Sore Sets 


179 


volumentric flow rate of COS (cu.ft./hr) 
mass flux (gm/em?-sec) 

heat transfer coefficient (cal/cm? «hr: °K) 
j-factor for mass transfer 

j-factor for heat transfer 

equilibrium constant 

reaction rate constant 


effective thermal conductivity 
(cal/cmehr °K) 


thermal conductivity of the fluid phase 
(cal/cmehr+ °K) 


mass transfer coefficient (cm/sec) 
total depth of the catalyst bed (cm) 
dimensionless length 


characteristic pore length, 
L, = R/3 for a spherical pellet 


L, = R for a flat slab pellet 
molecular weight of species A 
molecular weight of species B 


average number of atoms in a sulfur 
molecule 


mole flux of component A (gmole/cm* +sec) 
mole flux of component B (gmole/cm? +sec) 
Nusselt number (= h dp/k -) 


partial pressure of H,O (mm Hg) 


partial pressure of H,S (mm Hg) 


2 


partial pressure of SO, (mm Hg) 


evi he 























(xt\.3%3.09) 20D to esst worn ‘otx300mi Lov 


(o98- Soap) xt pene 


‘ 


(4° + ads *mo\.89) tagtoriieos we2enkxd deed cm. 
‘setenes aepm x08 sosostt 
setensad ‘deed 208 sosan3-t ve 

‘Sassen outa tape oe ‘ 
or Jestenoo ote. aoltoney. 


vay bteeben learredt svttostte i 
or ane eee an ef — 


eesrig Hhivlt sett. to Side tcahaaie Lemons 
(2° «redem\IB9) | . = 


“(ont jneintttsoo xoteassd pier: aS. .: ¥ ; | 
(m5) bed sayintso siz to. dsqeb. isgos ra 
geen! seeinobanenth, | ey Leet 

” dspael. s10Gq stseizego do eS Pogue ‘i 
selieq Isolxeiqe 6 102 Ae os ae ¥ 
tolleq dale 2612 & 208 # el oe me Ne 


‘en 
be) te 


A asibedge 26 shptow xsiuseson : 

8 eoiceqe 10 iriplew sido tom. My) : 

nalts tly» ig =, oa 

whive # ak anos to sodmin spezove Nay 
(Dee Sai sini A tmemoqnoo ~onien 
teams “m3\eLomp) g Snenoqnion bly ult etom 


* 
Ag 


Prost wh " ae 


a. 
i / Pat 


- 












- 


|S 


int 


Prandtl number (= CoE Up/K,) 
radius of a catalyst pellet (cm) 
gas constant 

Reynold number (= dpG/u -) 

radius of a pore (A) 


distance from the center of the spherical 
catalyst pellet (cm) 


rate of disappearance of HS 
(gmole/sec-gm catalyst) 


rate of disappearance of SO, 
(gmole/sec-gm catalyst) 


global rate expression 


intrinsic rate expression on the external 
surface of a catalyst 


intrinsic rate expression within the 
catalyst pellet 


Schmidt number (= U-/P-°D) 

Sherwood number (= k _ap/D) 
temperature (°K) 

equilibrium temperature (°K) 
temperature of the fluid phase (°K) 


temperature of the fluid phase at the 
reactor inlet (°K) 


temperature within the catalyst pellet 
(°K) 


temperature of the external surface of the 


catalyst pellet (°K) 
interstitial velocity (em/sec) 


weight of catalyst (gm) 


180 






















ga\qu 442)? “redmun fabaaxt 
(mo) tol log site 5 to! ‘subbes ; 
: gnsdenoD ese a a 
(aoa tm) redint a bionyer hi me : 
4a) o10q 6 4p: ‘ankbss ue, 


Lsoizadqe edd to redmeo edt moud sonetetb: 
‘ (mo) ea teyiatso 


rs “to consseeqaseth to cous 


“s ne apes esis gre oes 


o8 0 eimtoniten ts to asx 


" t teyiastso a~298),9 foate) 


nobResigxe esex dedorp. to) ae 
isnretxe sft no goderexone edsa piantzgak  ~ (? 
, —. b 20 eostine ee 


edt mit iw robenalsitice ej6% Stentaaat | Me 
jolieq: ai idee er oa 


ose =) rod Feito 
(DNB x =) ‘xedmua boowred@= 
(a) oxiyeieemes ie 
(x) ow3eregmed_ mtxab fis ~ i. ie 
(x°) ound pivtd ens to. 33 


edt 36. easitg bin{t oie 20. 
1 a ge 


Jetton faytsred ‘ont aiasan 5 otis 


eit to eusiiue r far 


Greek Letters 


181 


conversion 
equilibrium conversion 
conversion in the fluid phase 


conversion on the external surface of 
catalysts 


dimensionless length 
dimensionless concentration 
mole fraction of a component A 


reactor bed depth (cm) 


dimensionless variable 
void fraction 
void fraction due to micropores 


void fraction due to macropores 


effectiveness factor 


dimensionless temperature 


dimensionless temperature difference in 
the fluid phase 


dimensionless temperature difference on 
the external surface of catalysts 


geometrical constant of catalyst pores 
viscosity of fluid phase (gm/cm+sec) 
pore orientation distribution function 
dimensionless length 


total pressure 






















4 “natexevnoa 


| a 
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to sostive Lenie3xe edd no ‘ho lguerese ‘as oan 
‘BIEY LATER 5 
dtpael exeldotememib. ‘ 
noissisasonoD evelnotenamib 
A Jnenoamen & to noitosxt Lom eo ae 
(are) Atqgob fed sodoset, } 
ay re eae 
: La \ 


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noisoex3 ae 


* 


- aetoqo1sim ot aub noitosz3 Biov. (eae 
estogoi96Mm OF sab noks on bhoy. e . ©) ee 
: be _ - ee y sar ® 
xotos? oa, ae - 


stu Fexsqmed 

ne ecinpeatiihs oe 
mo soness23tb owniexs 
esaylateo io © 

astog soyieins to danveneo’ 
(ose mo\arp) wand: ah 

: aobions? Revco t 





density of the catalyst bed (gm/m1) 
density of fluid phase (gm/ml1) 
density of the catalyst pellet (gm/m1) 


pore size distribution function 


tortuosity factor 


Thiele modulus in general definition 
Thiele modulus for a spherical pellet 
Thiele modulus for a flat slab pellet 


dimensionless concentration 


182 














hs I Mg ev 7 ) 
, P ; 
; va Me 
: Py v J \ A 
es ! “- e Bs © 
: » - > é ‘ 
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APPENDIX A 


CALIBRATION OF THE GAS CHROMATOGRAPHIC SYSTEM 


189 


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190 


A-1l. Calibration Procedure: 


The gas chromatograph was calibrated using sample 
gas mixtures the composition of which was known from the 
volumetric mixing ratio. The apparatus used for sample 
preparation is illustrated in Figure A.l, which was oOrigin- 
ally fabricated by McGregor (72). 

The sample mixing chamber was made of a 5 liter 
lucite cylinder equipped with a movable piston and a mix- 
ing fan beneath the top cover of the cylinder. The lucite 
- cylinder had been calibrated by McGregor (72) to describe 
the volume of the gas mixture in the cylinder as a function 
of the piston position. 

Before preparing the sample mixture the cylinder 
was initially purged with nitrogen for about 2 hours. Then 
the cylinder filled with pure nitrogen through the line A 
was allowed to equilibrate to the atmospheric pressure and 
temperature. Then the gas burette and the line B were 
purged with one of the calibration gasesand a volume of the 
gas was trapped in the burette by lowering the mercury 
reservoir. The pressure of the trapped gas being not equal 
to the atmosphereic pressure, the elevation of the mercury 
reservoir was manipulated to get the atmospheric pressure 
which was confirmed by the water manometer. 

Then the trapped gas was compressed to about 30 


psi by raising the elevation of the mercury reservoir and 




















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ae 
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Gas Water 
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Piston and i : 
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: H|| Tie Bar Manometer 
: 7 
H 5 


FIGURE A.1l: GC CALIBRATION APPARATUS 


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forcing it into the cylinder through the line B with a 
sufficient positive pressure. The residual volume of the 
trapped gas was brought to the atmospheric condition by 
manipulating the elevation of the mercury reservoir again. 
The difference of the gas volume between the initial and 
the residual at atmospheric condition was recorded as 

the introduced gas volume to the mixing cylinder. 

While the first calibration gas was being mixed 
with the nitrogen in the mixing cylinder by the mixing fan, 
the line B and the gas burette was purged with the nitrogen 
and then by the second calibration gas. Then the second 
calibration gas was trapped in the burette, equilibrated 
to the atmospheric condition and then forced into the 
cylinder in the same way as the first calibration gas. 

The third calibration gas was also introduced into the 
cylinder in the same way as the first one. 

After the final sample was mixed for one hour by 
the mixing fan, it was forced into the GC sampling loop 
with a positive pressure head of 1 inch mercury, which was 
measured by a mercury manometer on the vent line from the 
GC, by lifting the cylinder position. 

The sample introduced to the GC sampling loop was 
injected to the sampling coil by pushing the injection 


lamp on the GC programmer panel for 13 seconds. 























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193 


A-2. Attenuation System Design: 


The individual attenuation system originally instal- 
led on the GC programmer panel was found to be malfunction- 
ing due to the excessive electrical noise causing the 
unstable base line whenever each microswitch changed its 
electrical connection on the timer control cam. The sche- 
Matic diagram of the individual attenuation system used 
by the former investigators (53,65,72) is illustrated in 
Figure A.2. 

So the previous individual attenuating system was 
not employed throughout this study, and the first GC cal- 
ibration was done au a fixed attenuation for all peaks as 
shown on Figure A.3. With this attenuation scheme, the 
GC calibration results by the computer and che integrators 
were not consistent as described in Chapter 5. 

Finally it was found that another installation of 
attenuation system was still possible. The newly installed 
attenuation system is illustrated in Figure A.4. The atten- 
uation scheme III was calibrated by measuring the output 
Signal from each attenuator at a arbitrarily assigned in- 
put signal. The voltage signal was measured using a poten- 
tiometer. The results of the calibration of the three 
attenuators are listed in Table A.1 through A.3. The 
GC calibration results using the attenuation scheme III by 


the computer and the integrator were in a very good 



























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194 


DIGITAL 
COMPUTER N2 a RECORDER INTEGRATOR 


te) 

GC 21 aaa) 
DETECTOR cos 

AS 

‘a PO 


ATTENUATOR 


FIGURE A.2 ATTENUATION SCHEME I 


DIGITAL 
INTEGRATOR 





DISC 
INTEGRATOR 






COMPUTER 








GC 
DETECTOR 


ATTENUATOR 


FIGURE A.3 ATTENUATION SCHEME ITI 


COMPUTER 
RECORDER 


DIGITAL 
INTEGRATOR 


ATTEN #1 ATTEN #2 












GC 
DETECTOR 
ATTENUATOR 


FIGURE A.4 ATTENUATION SCHEME III 


























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TABLE A. 1 


CALIBRATION OF CHROMATOGRAPH ATTENUATOR 


ATTENUATOR SETTING 
ATTENUATION RATIO 


~< 
nou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 19.95917 


Al = -1.88738 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
1.000 18 .0 69 18.071 9.013 
229099 16.008 16.184 he te af 
3.2000 ; 14.202 14.297 0.666 
4.999 12.586 12.409 1.403 
5 000 10.735 1D. DZ 2 1.984 
6 2999 8.690 8.634 0.640 
7.000 6.759 6.747 0.174 
8.009 4.740 4.860 Zeset 
9.900 22920 2-972 1.804 

10 £999 1.073 1.085 1.095 

VARI ANCE = 0.915210 

STANDARD DEVIATION = 0.123330 

MAXIMUM PCT ERROR = 2.527461 


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TABLE Aw 2 


CALIBRATION OF ATTENUATOR #1 


< 


ATTENUATOR SETTING 
ATTENUATION RATIO 


< 
i) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 1.07781 


Al = 0.17031 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
1.000 1.210 1.248 32119 
22099 1. 398 1.418 1,296 
3.000 1.687 1.588 5.835 
4.009 1. 784 Le DY 1.425 
5.000 1.893 Levee 1.880 
6.099 2.073 2.099 1-266 
72900 22254 2-269 0.704 
8.999 26 431 2-440 0.370 
9.000 2.613 2-610 0.126 

10 909 20197 2.1780 0.603 

VARI ANCE = 0.001636 


STANDARD DEVIATION = 0.040452 


MAXIMUM PCT FRROR = 5.833227 


per i = 4 eae 





















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VABLE Ay 3 


CALIBRATION OF ATTENUATOR 


~< 
Woy 


ATTENUATOR SETTING 
ATTENUATION RATIO 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AD = lel tse 


Al = 1.760 


X MEASURED Y OBSERVED 
1.909 Cag tt el 
2-000 4.635 
32009 66471 
4.000 8.365 
5999 10.147 
6.900 11.706 
7.099 13.450 
8.000 15.244 
9.099 16.980 

10 .090 18.743 

VARIANCE = 

STANDARD DEVIATION = 

MAXIMUM PCT ERROR = 


47 


REGENERATED DATA 


0.00 8538 
0.09240 1 


4.343668 


Y CALCULATED 


Lie 735 
4.694 
6.454 
8.215 
Teo To 
el OP Be 
13.496 
15.256 
17.017 


Loos t i 


POT 


ERROR 
4.343 
i Bod ge 
0.265 
1.802 
1.699 
0.251 
0.338 
0.081 
0.219 


0.185 


197 


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agreement as described in Chapter 5. 


A-3. Homogeneous Reaction Effects In The Mixing Cylinder: 


To check the homogeneous reaction effects during 
the mixing period of the sample gas mixture in the mixing 


cylinder, the calibration results for SO, component both 


2 
in the COS-H.,S-SO,-N, mixture and in the SO,-N, mixture 
were compared as illustrated in Figure A.5. 

Figure A.5 indicates that there was no homogeneous 
reaction during the mixing period of the sample gas mixture. 
The calibration data for the SO,-N, mixture are listed in 
Table A.4. 

The first calibration results of the gas chromato- 
graph using attenuation scheme II, are listed in Table A.5 
and the second calibration results using attenuation scheme 
III are listed in Table A.6. A listing of the program 


GCCAL, used to reduce the calibration data to usable re- 


sults, is also included. 


198 


sel 












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TABLE A.4 


CALIBRATION SAMPLE NUMBER 1 


SAMPLE PREPARATION CONDITIONS 


ROOM TEMPERATUREscccccccee 29265 DEG K 


BATH TEMPERATURE secccevceee 291-7 DEG K 


ATMOSPHERIC PRESSUREseceee 697466 MM HG 


VOLUME OF NITROGEN cccecccee 461923 CC 


VGLUME OF SULFUR DIOXIDE .. 56.2 CC 


SAMPLE COMPOSITION (MOLE PERCENT) 
NEFROGENeccccecctus, 0607675 
SULEUR DIOXEOE .c.. 22324 


100X MOLAR RATIO.w-e-« 142478 


INTEGRATOR AREA RESULTS (THE LAST SET IS THE AVERAGE) 


NITROGEN SULFUR DIOXIDE 
INTEGRATED PCT OF INTEGRATED 
AREA TOTAL AREA 
0.543787E O07 98.35 @.ILOZLIOE O05 
0.543654E 07 98,33 0.920240E 05 
0.5445 17E 07 98.35 0.912500E 05 
0.544201E O7 98. 32 0.928330E 05 
0 .545233E 07 98.31 0.934960E 05 


0.544278E 07 98. 33 0.921248E 


05 


PCT OF 
TOTAL 


100X 


AREA 
RATIO 


1.6738 
1.6926 
1.6757 
1.7058 
1.7147 


1.6925 


200 





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TABLE A.w4( CONTINUED) 
CALIBRATION SAMPLF NUMBER) 2 


SAMPLE PREPARATION CONDITIONS 
ROOM TEMPERATUREscccccccee 291.5 


BATH TEMPERATURE ccccccceee 290.8 


ATMOSPHERIC PRESSUREscceee 697.28 
VOLUME OF NITROGENecccecee 4619.3 


VOLUME ‘OF SULFUR DIOXIDE .. 12.6 


SAMPLE COMPOSITION (MOLE PERCENT) 
NI TROGENes cc ee tes 99.7211 
SULFUR DIOXIDE we. 0.2788 


100X MOLAR RATIO...» 0.42796 


INTEGRATOR AREA RESULTS (THE LAST SET IS THE AVERAGE) 


NITROGEN SULFUR DIOXIDE 
INTEGRATED PCi OF INTEGRATED 
AREA TOTAL AREA 
0 .549638E 07 99.67 0.179220E 05 
0.555189E 07 99.65 O-190600E 05 
0.553106E 07 99.68 Ox iH62Z200E 05 
0.549580E 07 99.64 0.1 IZ990E U5 
0.549491F 07 99.64 Os.lI3ZTOE 05 


DEG K 


DEG K 


MM HG 
GCE 


CC 


0.551409E O07 99.66 0.186464E 05 


PCT OF 
TOTAL 


100X 


AREA 
RATIO 


0.3260 


0. 3433 


0.3185 


0.3513 


0.3516 


0.3381 


201 










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TABLE A.4( CONTINUED) 


CALIBRATION SAMPLE NUMBER) 3 


SAMPLE PREPARATION CONDITIONS 


ROOM TEMPERATUREsccceccees 293.0 DEG K 


BATH TEMPERATURE cecccceeee 292.2 DEG K 


ATMOSPHERIC PRESSUREseeeee 70261 MM HG 


VOLUME OF NITROGEN ecceceee 4619.3 CC 


VOLUME OF SULFUR DIOXIDE .. 25.5 CC 


SAMPLE COMPOSITION (MOLE PERCENT) 
NITROGENccccccccvee 9904370 
SULFUR DIOXIDE ee. 045630 


100X MOLAR RATIO... 0.5661 


INTEGRATOR AREA RESULTS (THE LAST SET IS THE AVERAGE) 


NITROGEN SULFUR DIOXIDE 
INTEGRATED PCT OF INTEGRATED 
AREA TOTAL AREA 
0 .532763E 07 99.31 0.365880E 05 
0.524177E 07 99.34 0.343750E 05 
0.55 3896E 07 99.43 0.314710E 05 
0.531029E 07 99,32. 0.359030E 05 
0.535077F O7 99.37 0 .334400E 05 


0.535388E O07 99436 0.343554E 


05 


PCT OF 
TOTAL 


100X 


AREA 
RATIO 


0.6867 
0.6557 
0.5681 
0.6761 
026249 


0.6423 


202 


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TABLE A.w4( CONTINUED) 
CALIBRATION SAMPLE NUMBER 4 


SAMPLE PREPARATION CONDITIONS 


ROOM TEMPERATURE. cccecceee 29220 DEG K 


BATH TEMPERATURE seccecccceese 291.25 DEG K 


ATMOSPHERIC PRESSURE.eeeee 706266 MM HG 


VOLUME OF NITROGENecccecee 461923 CC 


VOLUME OF SULFUR DIOXIDE .. 98.0 CC 


SAMPLE COMPOSITION (MOLE PERCENT) 
SULFUR DIOXIDE eee 261274 


100X MOLAR RATIO... 2.21737 


INTEGRATOR AREA RESULTS (THE LAST SET IS THE AVERAGE) 


NITROGEN 7 SULFUR DIOXIDE 
INTEGRATED PCT OF INTEGRATED 
AREA TOTAL AREA 
0 .539304E 07 97.35 0.146565E 06 
0 .541332E 07 97.42 0.143307E 06 
0.533457E O7 97.28 0.149082E 06 
0.537665F O7 97.30 0.149103E 06 
0.541047E 07 97.31 0.149222E 06 


0.538561E 07 O hea3 0.147455E 


06 


PCT NF 
TOTAL 


Zetl 
2269 


2-68 


100X 


AREA 
RATIO 


Zaft 
22-6473 
2.1946 
fet i Sl 
2.7580 


20.1 361 


203 









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TABLE A.5 


GC CALIBRATION FOR N2-H2S MIXTURE 
ATTENUATION SCHEME II 


AREA RATIO 


x< 
"ou 


(100XH2S/N 


2) 


MOLE RATIO ( 100XH2S/N2) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = 0.0215 


Al = 0.9468 


X MEASURED Y O08 
1.194 
1.212 
22196 
32619 


ie obs 


VARI ANCE 
STANDARD. DEVIATION 


MAXIMUM PCT ERROR 


9 


7 


REGENERATED 
SERVED 
1.084 

1. 1h 

2.060 

3.463 


5.083 


0.00 150 3 


0.038772 


5.175019 


DATA. 
Y CALCULATED 
1.067 
1.169 
2.016 
3.449 


5.100 


PG) 


FRROR 


1.542 


50195 


2136 


0.401 


0.336 


204 






















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TABLE A. 5( CONTINUED) 


GC CALIBRATION FOR N2-COS MIXTURE 


ATTENUATION SCHEME II 


x< 
wou 


AREA RATIO (100XCOS/N2) 
MOLE RATIO (100 XCOS/N2) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AD = 0.0 3809 


Al = 0.72771 


REGENERATED DATA 


X MEASURED Y OBSERVED 
1.431 1.048 
1.571 1.168 
2.761 - 2.101 
4.526 3.343 
6.784 4.956 
VARI ANCE = 0.001131 
STANDARD DEVIATION = 0.033639_ 
MAXIMUM PCT ERROR = 3.022273 


Y CALCULATED 


1.080 


1.181 


2.047 
30331 


4.975 


PCT 


ERROR 
32022 
1.150 
2-538 
0.332 


0.389 


205 


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ae 


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206 


TABLE A.w5( CONTINUED) 


GC CALIBRATION FOR N2-SO2 MIXTURE 
ATTENUATION SCHEME II 


AREA RATIO (100 XSO2/N2) 
MOLE RATIO (100XSO2/N2) 


a, 
Wott 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 0.11839 


Al = 0.8469 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
22 382 2117 22134 0.799 
3.998 3.554 3.500 beS25 
4.768 4.133 4.152 0.472 
6.452 59516 Se STE iel03 
72176 66232 6.189 0.692 

VARIANCE = (0.002294 

STANDARD DEVIATION = 0.047904 

MAXIMUM PCT ERROR = 1.525091 


a0 


AOARS 


PPT. 


hte ae 


S\.0 



















a 


’ 7 Ms “Ss = . aS nf 
38UTXIM ‘SOz=Sv 80a vournsena a 
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oe : 


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(SANSM2XOON). CITAR. 30M = Y 


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eters = 0A 


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ets) GITARIMB TSA 
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207 


TABLE A.6 


GC CALIBRATION FOR N2-H2S MIXTURE 
ATTENUATION SCHEME ITI 


AREA RATIO (100 XH2S/N2) 
MOLE RATIO (100XH2S/N2) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 0.007 33 


Al = 0.89491 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
0.742 0.638 0.672 5.338 
1.327 1,179 } ie Be) 1.351 
2.277 22138 2-045 4.357 
3.838 3, 396 3-442 1.270 

VARI ANCE = 0.093989 


STANDARD DEVIATION 


0.063159 


MAXIMUM PCT ERROR 


5.338395 


- 2Ye 


Cos cs by 4 























+ v 
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vary ; teuxzsnmond? orTan again aa = 
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resea.0 =k f “a 


a) 
i 


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208 


TABLE A.6( CONTINUED) 


GC CALIBRATION FOR N2-COS MIXTURE 
ATTENUATION SCHEME III 


AREA RATIO (100 XCOS/N2) 
MOLE RATIO (100XCOS/N2) 


< 
wou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = -0 .00149 


Al = 0.65325 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
0.670 0.431 0.436 be t5S 
1.765 1.138 Rett 1.167 
3.188 Zeit 22081 1.598 
5.2212 3-388 3.403 0.458 

VARI ANCE = 0.000528 


STANDARD DEVIATION 


0.022987 


MAXIMUM PCT ERROR 


£.598227 


80s :; Vis a 


















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98 .t . i % artes a 


Ree, 0 -_ §208.08 ‘eee 


TABLE A.6(CONTINUED) 


GC CALIBRATION FOR N2-SO2 MIXTURE . 
ATTENUATION SCHEME ITI 


AREA RATIO (100XSO2/N2) 
MOLE RATIO (100XSO02/N2) 


x< 
nou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 0.0 2466 


Al = 0.77038 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED 
0.338 06279 0.285 
0 2642 9-566 0.519 
L1eo92 © 1.247 1.328 
2.738 2-173 22134 
VARI ANCE = 0.003431 


STANDARD DEVIATION 


0.058579 


MAXIMUM PCT ERROR 


8.234605 


PCT 


ERROR 
1.979 
8.234 
62470 


1.824 


209 


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TABLE A.6( CONTINUED) 


GC CALIBRATION FOR N2-CO2 MIXTURE 
ATTENUATION SCHEME ITI 


AREA RATIO (100 XCO02/N2) 
MOLE RATIO (100XC02/N2) 


< 
nou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = Qs0SL9t 


Al = 0.86214 


REGENERATED DATA 


X MEASURED Y DBSERVED Y CALCULATED 
0.350 0.263 0.269 
0.694 0.570 0.566 
1.056 0.887 0.879 
1.697 1.347 1.353 

VARI ANCE = 0.000055 


STANDARD DEVIATION = 0.007439 


MAXIMUM PCT ERROR = 2.442953 


PGT 


ERROR 


22442 


0.625 


0.988 


0.438 


210 
































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NCOPY 
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NCROM 
IPEAK 


RTEM 
BTEM 
APRES 
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sk 
tk 
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DATA SNAM/! 
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READ( 5,1) N 
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CALCULAT 


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* 

RAM REDUCES THE PEAK AREA DATA TAKEN FROM * 
AL INTEGRATOR TO CALCULATED RESULTS USEFUL* 


FOR THE INTERNAL STANDARD PROCEDURE FOR THE CALIBR-* 


THE GAS CHROMATOGRAGH 
ATA 


NUMBER OF SETS OF DATA 

NUMBER OF COPIES OF OUTPUT DESIRED 
CALIBRATION RUN NUMBER 

NUMBER OF CHROMATOGRAM TAKEN 

PEAK NUMBER CALIBRATED 

ee e 2=H2S 

eee 3=SN2 

ee eo 4=COS 

0 ee D=CN2 

ROOM TEMPERATURE(DEG C) 

WATER BATH TEMPERATURE(DEG C) 
ATMOSPHERIC PRESSURE(MM HG) 
DISTANCE BETWEEN PISTON AND END OF 
CYLINDER(CM) 

VOLUME OF CALIBRATION GAS(CU CM) 
ATTENUATOR #1 SETTING 

ATTENUATOR #2 SETTING 

AREA OF NITROGEN PEAK 

AREA OF CALIBRATION GAS PEAK 


% 3 3¢ He OH He OH He OH He OH HH HH HH HH eH HH HH se Ht 


FE RE A RIK 2 2K 2 RI I RE AR CK RK a I KA 3 A 2 2 aK aK 2k a aK aK aK ak aK 
NAM(454) 9SMV(5)_,V(2) 5 AR(5) sAVG(5),STORE(6 


402.10422176610,5 21901.634722417.51 3 22346.34/ 

HMYOR@ SM OGEN Ts) SUL, *FIDE',s*SULFS,"UR D*, 
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yNSET 

RUN»NCROM, IPEAK 

TEM,BTEMzAPRES,V(1) 9V(2) ,ATNIZATN2 


STEMP 
STEMP 


ION OF SAMPLE COMPOSITION 


211 

































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21 


51 


212 


eee (CONT'D) 


XMN2=(11.+151.2192*V(1))*STEMP/RTEM*APRES/SPRES/SMV(1) 
V(1)=(11.+151.192*V(1)) 

XMCAL=V(2)* STEMP/BTEM*APRES/SPRES/SMV(IPEAK) 

TOT M=XMN2+ XMCAL 

XMN2=XMN2/TOTM*100. 

XMCAL=XMCAL/TOTM*100. 

RMOL=XMCAL/ XMN2*100. 

KPEAK=IPEAK~-1 


READ AND PROCESS PEAK AREA DATA 


NCAT=0 

DO 4 T=1,5 

AVG(1I)=0.0 

DO 5 ICROM=1,NCROM 
NCAT=NCAT+1 

READ (5,33) AR(1),AR(3) 


CORRECTED PEAK AREA OF NITROGEN PEAK 


AR(1)=AR(1)*(1.76047*ATN2 + 12617312)¥*(0.17031*ATNI + 


* 1.07781) 


TOTA= AR(1)+ AR(3) 
AR (2)= AR(1)/TOTA*100. 


AR (4)= AR(3)/TOTA*100. 


AR (5)= AR(3)/ AR(1)*100. 


CALCULATION OF THE AVERAGES 


DO 6 IT=1,5 

AVG(I)= AR(1I)+AVG(I) 
DO 20 J=1,5 
STORE(NCAT,J)= AR(J) 
CONTINUE 

DO 7 {=1,5 
AVG(1I)=AVG( 1) /NCROM 
NCAT=NCAT+1 

DO 21 J=1,5 
STORE(NCAT,J)=AVG (J) 
NN=NCROM+ 1 

DO 23 IC=1,NCOPY 
WRITE( 6,17) NPAGE 
IF(NRUN=-1) 51,51,52 
WRITE( 6,37) 

GO TO 53 


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213 


eee (CONT'D) 


52 WRITE( 6, 38) 

53 WRITE(6,10) NRUN,yRTEM,BTEM 
WRITE(6,9) APRESyV(1) 5 (SNAM(JyKPEAK) » J=194)9V(2) 
WRITE(6,11) XMN2y (SNAM( Jy KPEAK) 9 J=194) 9XMCALsRMOL 
WRITE( 6,12) 
WRITE(6,13) (SNAM(JyKPEAK) »J=1,4) 
WRITE(6,15) ((STORE( I,J) ,J=195) sI=1l NN) 

23 CONTINUE 

NPAGE=NPAGE+1 

CONT INUE 

FORMAT(515) 

FORMAT( 7F10.5) 

FORMAT(2F10.2) 

FORMAT( / 15X »'*CALIBRATION SAMPLE NUMBER! ,139///10Xy 
** SAMPLE PREPARATION CONDITIONS! //12X,'ROOM TEMPERATUR! 
Rete es cecweswwear ¢oteis’ DEG K's //12X, "BATH TEMPERATU! 
Mel RE earciaeinele diel ettale Ls. DEG K',/) 

9 FORMAT(/12X_9* ATMOSPHERIC PRESSURE cecece 'yFlely' MM HG! 
*y//12Xy" VOLUME OF NITROGENeccocccce 'sFlolyg!' CO%s//12X,y 
1* VOLUME OF * 54A4,_,'.e'yF5e1,' CC!) 

il FORMAT( //7,10X,_ "SAMPLE COMPOSITION (MOLE PERCENT) !/ 
MSI2Xs*NITROGBNaacccwcccce sf Beh y S//1L2K 7 4A4, Seca! p FBL4, 
1//12X,"100X MOLAR RATIOece 'sF8.4) 

12 FORMAT(//10X," INTEGRATOR AREA RESULTS (THE LAST SET ', 
T' iS THE AVERAGE )'/) 

13 FORMAT ( 19X_ *NITROGEN § 413X94A4_,9X_'100X'y//11Xy 
* © INTEGRATED! y6X,_'PCT OF',5Xy*INTEGRATED!' ,6X_,'PCT OF", 
17X,"AREA'/ 15X_y'AREA'y8Xq_!" TOTAL! 410X%_"AREA' 8Xy_' TOTAL! 
*, 6X," RATIO! /) 

15 FORMAT( 9X 9E13.693X 9F66296X9E136693X 9 F602y SXy Fl e4/) 

17 FORMAT('1'4///66X_'A='"ygI25//) 

37 FORMAT(20X," TABLE A.4' / ) 

38 FORMAT(20X,'TABLE As4(CONTINUED)! / ) 

CALL -EATT 
END 


mW 
OwWWrHN 


APPENDIX B 


CALIBRATION OF PROCESS MEASURING DEVICES 


214 








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245 
B-l. Differential Pressure Transducer: 


The electronic differential pressure transducer 
was calibrated using a pre-calibrated dry test meter 
with pure nitrogen flow at room temperature. A Foxboro 
6430 HF electronic consotrol recorder was used to record 
the pressure difference in the cell. The calibration 
data were obtained at three different feed pressure 
levels of 20, 25 and 30 psia and were fitted to the cal- 


ibration equation of the following form. 


2 
ag + a,x + a5X 


Yc 
where 


Y volumetric flow rate (SCFH) 


X Square root of percent reading on the recorder 


Agra ray = calibrated parameters 


by applying the linear least-square curve fitting techni- 
que assuming that the flow behavior was in the region 
between laminar and turbulent flow. The calibration re- 


sults are listed in Table B.1, B.2 and B.3. 


B-2. Feed Absolute Pressure Transducer: 


The Foxboro 611 AH absolute pressure transducer 
was calibrated at room temperature with pure nitrogen 
flow in the feed line using a mercury manometer as a ref- 


erence. The absolute pressure signal was recorded on the 


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TABLE Be. 1 


DP-CELL CALIBRATION 


(FEED PRESSURE AT 30 PSIA) 


22 
nou 


RECORDER READING (SQRT OF PER CENT) 
FEED FLOW RATE (SCFH) 


THE COFFFICIENTS OF THE POLYNOMIAL. ARE 


AO = -) .00814 
Al = 1.05586 
A2 = 0 .0000 3 


REGENERATED DATA 


X MEASURED Y OBSERVED 

9.920 10.492 

9.099 9.560 

8.075 8.502 

7.457 7.929 

6.066 6. 368 

4.528 4.790 

3.592 So771 

2.408 2.539 
VARIANCE = 0.001113 
STANDARD DEVIATION = 0.033366 
MAXIMUM PCT ERROR = 0.775749 


Y CALCULATED 


10.469 
9.602 
8.520 
7.867 
6.398 
4.773 
3.785 


22534 


PCT 


ERROR 
0.212 
0.441 
0.216 
0.775 
0.472 
0.342 
0.371 


0.173 


216 


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TABLE Be 2 


DP-CELL CALIBRATION 


(FEED PRESSURE AT 25 PSIA) 


~< 
ou 


RECORDER READING (SQRT OF PER CENT) 
FEED FLOW RATE (SCFH) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = 0.08793 
Al = 0.94999 
A2 = 0 .00099 


REGENERATED DATA 


X MEASURED Y OBSERVED 

9.894 9.403 

9.050 8.588 

7.937 1389 

6.892 6.559 

6.253 5.978 

4.950 4.762 

32249) 3-234 

2-387 22296 
VARIANCE = 0.001578 
STANDARD DEVIATION = 0.039724 
MAXIMUM PCT ERROR = 224426989 


Y CALCULATED 


9.390 
8.604 
72565 
6.588 
7eIS9 
4.1766 
Se V55 


22349 


PCT 


ERROR 
0.137 
0.188 
0.308 
0.445 
0.191 
0.086 
22426 


2-348 


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DP-CELL CALIBRATION 


(FEED PRESSURE AT 20 PSIA) 


~< 
“ou 


FEED FLOW 


RATE. (SCFH) 


RECORDER READING (SQRT OF PER CENT) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = -0 .06161 
Al = 0.87416 
A2 = —0 .00 318 


REGENERATED DATA 


X MEASURED Y OBSERVED 

9.818 8.208 

9.311 72791 

8.438 7.098 

72964 50 91T 

5.899 4.990 

4.837 4.076 

3.479 ges i Es. 

2-998 1.778 
VARI ANCF = 0.090302 
STANDARD DEVIATION = 0.017382 
MAXIMUM PCT ERROR = 12195066 


Y CALCULATED 


8.213 
7.801 
7.087 
5.954 
4.984 
4.092 
2-941 


Le /38 


PCT 


ERROR 
0.069 
0.133 
0.145 
Veo 
0.116 
0.396 
0.892 


1.105 


218 


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219 


Foxboro 6430 HF electronic consotrol recorder. The refer- 
ence absolute pressure for the calibration was obtained 

by adding the barometric pressure to the mercury manometer 
reading. The calibration data were fitted to a straight 
line by means of a linear least-square fitting technique 


and the results are listed in Table B.4. 


B-3 Reactor Gauge Pressure Transducer: 


The reactor gauge pressure transducer (Statham 
PG-732TC-10-350) was calibrated in the same way as des- 
cribed for the feed absolute pressure transducer cali- 
bration. The data which represent the relation between 
the percent reading on the recorder and the mercury mano- 
meter reading was obtained after the reactor system 
reached steady state through which pure nitrogen was 
flowing. Since the normal operational condition of the 
reactor system was in the high temperature range of 
around 250°K, the reactor pressure transducer was cali- 
brated at elevated temperatures of 480°K, 500°K and 
560°K. A frequent malfunctioning of the reactor pressure 
transducer was found. Whenever any malfunctioning was 
detected, the transducer was checked and recalibrated 
for next runs. The results of calibration #1 were used 
for runs A, B, C, D, E and J. The calibration equation 
#2 was used for runs F and G, while the calibration equa- 


tion # 3 was used for other runs. The calibration re- 


sults are shown in Table B.5 through B.9. 











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TABLE B. 4 


FEED ABSOLUTE PRESSURE TRANSDUCER CALIBRATION 


X = RECORDER READING (PER CENT) 
Y= ABSOLUTE PRESSURE (MM-HG) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AD = 744.61367 


Al 


12.71505 


REGENERATED DATA 


X MEASURED 


2-900 
5.900 
13.809 
20 .500 
29.400 
38.900 
52.309 


65.500 


VARIANCE 


STANDARD DEVIATION 


MAXIMUM PCT ERROR 


-Y OBSERVED 


781.100 


8 20 .500 


921.900 


1004.900 


1115. 500 


1237.200 


1413. 800 


1576. 300 


5.136350 
22266351 


0.296308 


Y CALCULATED 


781.487 
819.632 
920.081 
1005.272 
1118.436 
t2s3.2e9 
1409.610 


1577.449 


PCT 


ERROR 
0.049 
0.105 
0.197 
0.037 
0.263 
0.163 
0.296 


0.072 


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REACTOR GAUGE PRESSURE TRANSDUCER CALIBRATION #1 
(AT 480 DEGREE K) 


RECORDER READING (PER CENT) 
GAUGE PRESSURE (MM-HG) 


< 
io 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AD = 8.92106 


Al = 4.41818 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
6.100 36. 300 re, Bihil Lie, BAD 
11.300 59.500 58.846 1.098 
22.150 10 6.900 106. 783 0.108 
35.150 163.300 164.220 0.563 
50 .509 230.700 232.039 0.580 
77.500 352.000 45:1).,3:310 0.190 
90 .600 409. 600 409.208 0.095 

VARI ANCE = 0.644358 


STANDARD DEVIATION = 0.802719 


MAXIMUM PCT ERROR = 12179151 






















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TABLE B. 6 


REACTOR GAUGE PRESSURE TRANSDUCER CALIBRATION #1 


(AT 509 DEGREE K) 


RECORDER 
GAUGE PR 


~< 
ou 


READING (PER CENT) 


ESSURE (MM-HG) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AD = 14. 52163 
Al = 4.33165 
REGENERATED DATA 
X MEASURED Y OBSERVED Y CALCULATED 
6.990 40. 500 40.511 
14.300 77.2100 76.464 
26.000 126. 600 127.144 
43.150 201.400 201.432 
58.800 269.100 269.222 
73.2500 332.600 332,597 
95.600 429.000 428.627 
VARI ANCE = 0.157432 
STANDARD DEVIATION = 0.396778 
MAXIMUM PCT ERROR = 0.824627 


ERROR 


0.028 


0.824 


0.430 


0.016 


0.045 


0.089 


0.086 


222 


















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TABLE Be. 7 


REACTOR GAUGE PRESSURE TRANSDUCER CALIBRATION #1 
(AT 560 DEGREE K) 


RECORDER READING (PER CENT) 
GAUGE PRESSURE (MM-HG) 


x< 
wou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 15.83784 


Al =- 4.32817 


REGENERATED DATA 


X MEASURED | Y OBSERVED Y CALCULATED PCT ERROR 
4.200 34.500 34.016 1.402 
11.500 66. 200 65.611 0.888 
17.000 89.600 89.416 0.204 
28.490 138.300 138.758 0.331 
39.700 186.700 187.666 0.517 
58.100 266. 400 267.304 0.339 
733200 333.000 332.660 0.102 
91.500 412. 600 411.865 O72 

VARI ANCE = 0.461480 


STANDARD DEVIATION = 0.679323 


MAXIMUM PCT ERROR 1.40 2393 


Y 
| 
: 

















ce FAROT IAS sanucemaar dJnivedna: ‘baa. ag 4, 
(A Rointect: 
5 ee . ay at 

Wad 93%) On OARA YBORNITA = Ri , 
(H=MM) ame2 gat FIUAD = Y a: 

a a er 

NM ; ah a see : —£ 

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et 


A) 


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—— , THOSE SA.  s 


ATA OITARIAZOIA 


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Cats finale 


TABLE B. 8 


REACTOR GAUGE PRESSURE TRANSDUCER CALIBRATION #2 


(AT 560 DEGREE K) 


~< 
wou 


RECORDER READING 
GAUGE PRESSURE (MM-HG) 


CPER “GENT 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AQ = Peeks 79 


Al = 4.05037 


X MEASURED Y OBSERVED 
15.100 63.000 
26.800 108.000 
33. 300 135.000 
39.000 160.000 
50 .000 207 .000 
61.6090 254.000 
63. 300 261.000 
80 .809 326.000 
92.200 374.000 

VARI ANCE = 

STANDARD DEVIATION = 2.283985 


MAXIMUM PCT ERROR 


REGENERATED DATA 


5.216590 


2.188647 


Y CALCULATED 


62.974 
110, 363 
136.691 
B95 he 
204.332 
Ze 3bG 
258.202 
329.053 


af 5.298 


ERROR 
0.040 
2-188 
Be252 
0.138 
1.288 
1.056 
1.071 
0.945 


0.336 


224 


- 
v2 


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B81. dae Ort. 
Saset £e9,ae! 
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aoe. : | BBO PSE” 


aee.0: ees,eve 5 | 8 
















ba 
rg 


nnad aay uraaan, s0anh = x Vig 
(arama) dels Jib “a 


32h JAI wavy 08 ‘sar Ae, = an 
‘ofc tbaby (Re 
ee 


foo @aTaqua2as ¥  OawRReER > 
hi : bd en eee 





TABLE Be. 9 


REACTOR GAUGE PRESSURE TRANSDUCER CALIBRATION #3 
(AT 560 DEGREE K) 


< 
wou 


RECORDER READING (PER CENT) 
GAUGE PRESSURE (MM-HG) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = 27.66422 


Al = 4.19851 


X MEASURED Y OBSERVED 
20 .000 112.000 
31.700 161.000 
42.800 207.000 
47.800 231.000 
54.000 252.000 
64.000 294.000 
83.000 378.000 

VARIANCE = 

STANDARD DEVIATION = 1.918356 

MAXIMUM PCT ERROR 


REGENERATED DATA 


= 1.145756 


3.6800 93 


Y CALCULATED 


111.634 
160.757 
207.360 
228.353 
254.384 
296. 369 


376.141 


ERROR 


0.326 


0.150 


0.174 


12145 


0.946 


0.805 


0.491 


A225 


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= a eee 
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(OH=MM ) wera TO0RD! « be 


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an — 


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$84 JAYMOWY SOF “BHT. AO 2vwai9191a09 3} 















$8808 = 0 
Re Mek al 0.5 
é. dese oe 


AT AQ - OSTARIMADIA 
GaTAJUDIAD ¥ ‘gavaazao v~ 
44 Sip ay ; 71s i. 3 


PESTLE = 00st 
‘ves vs ; i ) ere c ri 
yey 20dI . of abe fat wad soe 
O8é .TOS rs . 000. TOS. 
eee.asS |) OOOKFES 


ABE, SAS 


4 


ese.aeS 
i= % pat 


{1 ore. ¥ ees ‘d . 
fy ; 7 


226 


B-4. Thermocouples: 


The thermocouples for measuring the reactor inlet 
and outlet temperatures were calibrated using a platinum 
resistance thermometer as the temperature standard ina 
fluidized sand bath at the instrument shop. The cali- 


brated results are shown in Table B.10 and B.1l. 


B-5. Water Feeder: 


The syringe pump for water injection (Sage model - 
355) was calibrated by a gravimetric method. The flow 
range of the syringe was set at the scale of 1/100. The 
water sample was received in a 5 ml weighing bottle. To 
eliminate evaporation loss during the calibration period 
a capillary tube stopper was used to connect the syringe 
outlet to the weighing bottle. The weight of the sampled 
water received for a certain time interval was converted 
to the volume of water by dividing it with its density. 
The data for water density (1) was corrected for differ- 
ent temperatures as shown in Table B.12. The calibration 


results are shown in Table B.13. 


1. Perry, J.H., "Chemical Engineers Handbook", 4th ed., 
New York, 3-70, 1963. 



















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ott tt a 30) 
—- of: 

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: pe * 7 
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227 


TABLE B.10 


THERMOCOUPLE CALIBRATION FOR REACTOR INLET TEMP. 


RECORDER READING (MILLIVOLT) 
TEMPERATURE (DEGREE CENTIGRADE) 


x 
woot 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 
AO = 1.40878 


Al = Te FUUSZ 


REGENERATED DATA 


X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
1.0 34 20 -510 20.342 0.814 
20075 37.950 39.405 32834 
72970 149.350 147.351 1328 

14.211 261.510 Fels tM G Fol 0.047 
20 6254 371.610 2 a2 0.183 
26.238 481.960 481.866 0.019 

VARI ANCE = 14325184 

STANDARD DEVIATION = 1.151166 

MAXIMUM PCT ERROR = 34834459 


















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TABLE B.11 


THERMOCOUPLE CALIBRATION FOR REACTOR OUTLET TEMP. 


RECORDER READING (MILLIVOLT) 
TEMPERATURE (DEGREE CENTIGRADE) 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = 2251827 
Al = 18.2165) 
REGENERATED DATA 
X MEASURED Y OBSERVED Y CALCULATED PCT 
1.034 20 -510 21.354 
1.949 37.2950 38.022 
7-969 149.350 ) 147.685 
14.234 261.510 261.812 
20-296 371.610 3726240 
26.309 481.960 481.776 
VARI ANCE = 0.802073 


STANDARD DEVIATION = 0.895585 


MAXIMUM PCT ERROR = 4.115689 


ERROR 
4.115 
0.190 
1.114 
0.115 
0.169 


0.038 


228 


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‘ : > 
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a i 





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BRA IAT MOY JO ot 30 chelsea 


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s0as3 729 GaTANUIIAD ¥Y > Gavnezeo -¥ 
(fee | ~ Beare ‘_ oe Le. 0S : 


0°76, $$0 8€- i Yer 02? re : 

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Oitel, = 288.781 ~ > CREE 
: » 2 . : 9 Pus ue 
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Pd.0 005.816 2 [cae eee 


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229 


TABLE B.12 


CALIBRATION OF WATFR FEED PUMP 


ADJUSTMENT READING (PER CENT AT 1/100 SCALE) 
FEED RATE (CC H20/HR?) 


x< 
wou 


THE COEFFICIENTS OF THE POLYNOMIAL ARE 


AO = 0.31009 
Al = 0. 520 32 
REGENERATED DATA 
X MEASURED Y OBSERVED Y CALCULATED PCT ERROR 
10 .000 5-388 SS he gk a poo sh 
20 2099 19.849 10.716 Leses 
30 000 16.107 bi he I oe 1.165 
50 2999 25A9935 26.326 1.279 
70 .000 36.869 36a (F2 0.371 
VARIANCE = 0.049418 
STANDARD DEVIATION = 0.222303 
MAXIMUM PCT ERROR = 2.315128 


ANARS 
















quue a334 AITAM 30 moira AS 
; Bie 


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ae ~URANOSHY 1 3TAA 0334 * - Y Ne 


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aAA JAL wom 09 HT ap el 
: “ 900Ke. o = OA 4 
op Ae 


sé 0S@ 0 8 


Laws 
7 ~ ie . 
Pr. ‘aes 
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1290 GaTAWIIAD ¥  aavasee0 y a3 
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art .ot- | (088.08 
ere.e! ae : : 
ateceae 1 age ae 


SEV.8E 


TABLE B.13 


DENSITY CORRECTION OF FEED WATER 


= FEED WATER TEMP. 
Y = DENSITY OF WATER (GM/CC) 


(DEG. 


THE COFFFICIENTS OF THE POLYNOMIAL ARE 


AD = 1.00 244 


HH} 


Al 


—Ve00 22 


REGENERATED DATA 


X MEASURED Y OBSERVED 
10 .999 0.999 
15.000 6999 
20 2999 0.998 
25.000 Ong 
30 2.099 2995 
35.000 0.994 
VARIANCE = 0.000000 
STANDARD DEVIATION = 0.000346 
MAXIMUM PCT ERROR = 0.043511 


Y CALCULATED 


1.000 
OEIS9 
Oi Be a | 
0.996 
0.995 


0.994 


of ot 


ERROR 


0.043 


0.010 


0.034 


0.033 


0.007 


0.041 


230 


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ofS | <<o'r ea ee 
















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APPENDIX C 


DERIVATION OF EQUATIONS FOR ADIABATIC REACTION PATHS IN 


THE FRONT-END BURNER 


2a. 


vl 


AaMAvE OVE ~THORE SE 





LY 
- 
= 
- w rn 
3 . 
‘ wan ‘ 
y = 
‘eo 
- 4 = 
2 | 
eee 
ri : 
; 2 
' 
~~. 
ry T cA ’ 
~ a XI ua9GA " 
* 
i 

Ls ‘ bal a 

7 f ~,. 
~ » 
; aT ee as 
4 _ w vi a 

e % Gale” * 


si OUTAHATGA ween 


oe 


Y oe rs 7 


232 


C-l. Adiabatic Reaction Path in the Front-End Burner 


Section (1): 
The oxidation reaction between H.S and air to form 


2 


SO, and H50 is assumed to be the only reaction to occur 


in the burner section (1) in Figure 6. 


AH 4 gg (1) 
HS + 3/2 oP A ——» SO. + H,0 
No. of moles, initial: S La 0 0 
final: (3-X) Stl) X Pi 


ieee aol) 

where X is the number of moles of HS converted to Soh 
Since the adiabatic condition during the reaction 

period in the burner may be assumed, the total enthalpy 
of the system should remain constant before and after the 


reaction. Then the following energy balance equation can 


be written by referring to Figure 6. 


298 
3nG +1.5C + (5.64 + .2)°6. tic } dT + X*AH 1 
T 2 2 2 2 
1 
T, 
+ {(3-X) C + 1.5 (1-xX) C + (5.64 + Z) C +xXcC 
Poy S Poo Pow Poo 
298 2 2 2 2 
+xXC +¥C pees Sg Plt) ene (A.2) 


Ree ree 













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mrot of xis bas 2. B aaowsed noltosex noisebixo é 
tuDD0 of ioktones bi yl edt ad os bomuees at pee 
3 eivp ld mk a) nolye9e semmud, 





° ~ ’ ave 
(f) 5.80 pick tabi nt ; 
oH + Ve -_ <2 SXE + eon rae > : 
| ees 
o° 8 fol (OS Eaegaaes veto 
x x (Het Oe) alsnkt " me 
(AY cawekeet Oe hp AR ele ae 


+08 o7 tiaséeaied 3A 3 to eaten’ ‘to zedawa ont al 3 


ut on 
rottosex ody paineb sols tbado. oithdsibs odds 5 nko - 
vqlatitne [sted sd3 .bemees, > cat som 






t 


HA+X + Tb { om. nae 





where Y and Z represent the number of moles of co. and 
the excess nitrogen present in the feed stream as inert 
gases. 


The heat capacity of the gaseous component i may 


be described as a function of temperature (70) such as 


5 ne 
SE Pe Ce te eer bee (A.3) 


g j=1 13 


where A; = temperature coefficients for the thermodynamic 


property function 


and a oe SEOs HS 
LW? As, ECO sO, 
iL Sy 34for S¢ 
dite e for S¢ 
wa~ 5 tor S5 
tae. 6. LOY H,O 
yy ents ee oh N, 
Lo, 6. LOL H, 
Lea) SOL 0, 
i= 10: ‘for CO. 


Then equation (A.2) can be integrated to give 


(A. 4) 


Ww 


conversion in the burner section (1) = 


° 


where X = Ry * FHT/{ AH, 9g (1) - Ry * PHEAT} (A.5) 


233 


& 
jimsaybostieds sft st02 ean eioittess. oxieviome *, he 
& 


0. #efon %o tecmnin atta Jnsgexqos : nas 


: a ; 
{ 1? 4 
i eo 4 re 
aes ha a 

. « a 
a | 














a? | 


sexse Hest s it at dnseerq noporsin nies a 
‘ f 
; : ‘ y ‘ Sk 7 
IntoD Byesasp ex Wei Wiosgs 3D ‘ele anit =e 
; i a oP, 
- he a ee r 


® 
{ 
r 
He] 
y 
4 
im 7 
© 
es 
o 
-¥ 
a: 
bre. 
\& 
ra) 
:. 
yp 
m 
ow 
m 
ead 
i= 
b 
=. 


~_ 


"ide 4 foktomut | o7 


| 8m 10% os = : “fin 
~ ; . Hi 4 * 
Oe tot if 1g nt » 


, t a ks. eis 
» “ace : 7 _ ow . fn 
2; 
e 
bi 
oS i 
4 


A ae q 
evip o3 } bedexpadat © od cae 
= 

) ey : 


molsoee ‘wen . rd < 1 


234 


FHT = 
j 


i mu 


. . . ae * EX Pos eS 
’ (Ay, +1.5 Ay, + (5.64 + Z) Aj, +¥ Ajo, ) 


TE 2) Re et a RE | (A.7) 


PHEAT = 13 94 24 64 


j 


Wo Mot 


1 


The computer program BURN] was used to solve the 


above equations and also listed. 


C-2. Adiabatic Reaction Path in the Front-End Burner 


Section (2): 


In the burner section (2) in Figure 6, the reac- 


tion between HS and SO, occurs according to equation 


(A.8) following reaction (A.1) in the burner section (1). 


°o 
AH 59 (2) 3 
2 H,S + SO. ——— 2 H0 + 5 S, 
No. of moles, initial: 2 1 rE 0 
final: 2 (1-x) (1-X) (1"2xX)"" Fe5xX 
eeeeeevee (A.8) 


where X is the conversion of H,S to elemental sulfur. 
Furthermore the association-dissociation reactions 
between sulfur species are assumed to occur simultaneously 


with reaction (A.8). 


é I 
‘ie 
~ y «? 

J ‘ 
~~ 

i 
Siwia aA 
: 2 - 
ao & 
f 


 evvrae® . * ' ‘al 
\ } ft 
= t a ot at = - 
7.» <5 3 
: - po meeeet eee { , - Am. AG. A 













~~ 


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a 
sbetert ate bas. enoiseups. ' VO 
« « “Size 7 






3 eo 

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verry crt bra-$10% ; q prt at ijst ‘nolsbeok oltadst } " 3s 
Smita Diba a ow ee i 


 % 





- - 7 ne pat 
- ee — ~ “ 


By 7 . rt s) nokine? 


j (re 
i ye 
h eee 


¢) nolo08 Yesrrmd ord ar i. <a 





loin 


noitsrps of paibsoons atuno9 ° 02 bas e s= neortied do m0 


cas 


; 2 P . at te a » , 
(I) notsoee stenmid ent at (1A), soldone: pakw for, (Be . 


¢ (S) geck4 pH ¢ : 
a ) “4 inieaemieies O28 ba nh “ 
, t - © c C ¢— . = ¢ ta ¢ ; 
i ©, ' 
~ a a , 
0 ae £ ’ me va . ? ; 
Oey 
‘ty ry 
ye. k (xtet)' (x-f) (xen) aia 


or 
‘s 


suttlve isgnemel 2 09 aga te a 
rudy 
r cr | 


anoisosst holes ica bs nots isos 












ae 
a : auosnsd lume THDOO of nae = 

fea © my y Vie im .s in Ata J ~rotia € / A 
ae a $e “J 7 + a Lj ae , 
ps Fite ae Bree pee 


bss |s ict ae 


He 


a. 













* 
ays ; 
7 


° 
AH (8) 
3 298 
So ot 5; (A.9) 


(1-v)-v,) (1.5 X) v, (1.5 xX) 


oO 
AH 599 (9) 


4s ee 
i 4 
(A.10) 


(I-v)-v5) (1.5 X) Vo (1.5 X) 


Then the following energy balance equation can be 
derived by referring to Figure 6 on the basis of the adia- 


batic condition of the reaction system within the burner. 


298 


if T2200 ae & 0G G5. Ge 7, aC OG } dT 
Poy S Poco Poy O Poy Psa9 


T, 2 2 2 2 2 


° ° 


+ 2 XeAH ror v, (1-5) AH, gg (8) 4 v, (1.5%) AH, 99 (9) 


298 
a 


3 
+4 {201-x) Cc 
{ Poy g t CX) Co + (42x) C+ (1-v vy) 1.50, 


Li i 
> > > 
298 2 50, H,0 So 


, + (5.64+ + = 0 
+ v, (1.5X) Cort Vy (1. 5X) c. (5.64+2Z) oe TAC } dT 


> > > Ps 
56 Se Ny co, 


By the same procedures as in (C-1) the following results 


can be obtained. 


235 


tw 
mM 
“> 
© 












: Samh ¢ | 
(8) 25-84 <= 
(@.A) | a +h, < 


& Ns ¥ ete | ax 


a 


| Ae ae : a ; —. 
8 . 
(OLA) | + cast 
- Yee. Ye 3 2, y eet ee aaa 


ed aso not3sups eoneiad yerens palwol{o3 edt cote * : 7: 
} erie & >» aa 


~S ibs ony to eftesd odd no 3 oteplt oF pnisyotex vd £ hae 


i kn 


-xen1wd ent Pacers mozaye noisoset. ais 20 tots BBD. 


i: . 
ae D cee” Vea 
m{ 394 PS +802). | Ph Hs 
oot4 ae? > sg pe 


QAI gyngued) + DO AKSHT)' + 


oY y 
| Of : 


Total conversion in the front-end burner 


= (1 + 2x)/3 (A.12) 
where 
X = R, - FHT/(DH - R, + PHEAT) (A.13) 
Oi teat 21, ae Cee Ey YALA Poa ) 
j=1 13 23 , 7j 10j 6j j 
a Ge @late oe (A. 14) 
DH = 2 AH, 9g (2) of LESH: AH, 9, (8) FES! ONT AH, 9g (9) 
a ¢ 6b e* as (And) 
2 2 ay 
PHEAT BS {2 AD; + Ay; - Aes = 1.5 (1-v,-v,) AS - 0.5 M.° Ang 
298) - sh 
750-373 Vay? A, Coe eg ee (A.16) 


J 


A listing of the computer program BURN2, used to 


solve the above equations, is included. 


236 


*e eevee 













4y 
‘ “— F i 
bre now ars at noLarevnod issor M 
‘ } 
: cy 
£ \ f ¢ oa tr) = 4 : 
4 > [ 
f y a Ri? 
~ (TANS = A - BG)\THI + fF = ag 
: gt pat o4 io 
y ; 1 zi 
Betas A 
Bas é@ 
AY .A(S' + OVC) (# A + ASE ays. 
tol * han at 
| ete AN ' ee 
fi Pa : . ah jes a y 
ie i. : Ae 


4 ey i+ ge nso aw 


\ 
‘ 


er, 


vm vn) 2, : = 


e T wy * 
(Sree te er) “ase.0 4 


& 


t 5 F 
~ =i a. 7 
of oS bos —-_ 
oa Cie eG ld 
F ; 


Pe on 


m&xpo 24 x2 IUCMOD “eta Ro. pats 


-bobylont ek canis sine ® J 


. ik ~*~ 





OVO CII IYO CMO OOO OOM Ota ag 


237 


WE AE EC CC RE a OC OK OK OK IC IK CC a OK Ok OK OK RE OK IE OIC OI OK OK OK IK OK 2 OK OK OI OK COI OK OK oie Ok 2k i KOK Ok OK 2k OK OK ok 
* 


MAINLINE BURN1 * 


* 
THIS PROGRAM COMPUTES THE ADIABATIC REACTION PATH x 
IN THE FRONT-END BURNER SECTION (1) OF A CLAUS UNI T* 


PT TOTAL PRESSURE IN MM HG 


* 
* 

¥ 

x 

* 

* ACCORDING TO REACTION (1.1). * 
* **K 
% N = NUMBER OF COMPOUNDS IN THE REACTION * 
* MIXTURE * 
* A = TEMPERATURE COEFFICIENTS OF A THERMODYNA-* 
* MIC FUNCTION FOR HEAT CAPACITY( FOR LOWER * 
* THAN 1000 DEG K) * 
* AA = TEMPERATURE COEFFICIENTS OF A THERMODYNA-* 
* MIC FUNCTION FOR HEAT CAPACITY(FOR HIGHER* 
* THAN 1000 DEG K) * 
* ¥ = NUMBER OF MOLES OF CARBON DIOXIDE IN THE * 
* » FEED * 
* Z = NUMBER OF MOLES OF EXCESS NITROGEN IN * 
* Loe Pee * 
* TFD FEED TEMPERATURE * 
* * 
* xx 
* 3K 


So a oo kak a ak ak ak ok aki ai a ok ak aka ak ak ac oi ak a aK aa ak 2k ak 


DIMENSION A(20,%5), LO eA) eT RT DY 
10( 20,5), TPP(5) 
READ(5,1 ) N 
1 FORMAT (115) 
DO 3 I=1,N 
3 READ( 5,5) (A(I,J),J=1,5) 
5 FORMAT( 5€15.7) 
DO 4 T=1,N 
4 READ (5,5) (AA(I,J)9J=195) 
READ( 5,8) Z,Y 
8 FORMAT(2F10.5) 
WRITE( 6,10) 
10 FORMAT(1H1,25X,'ADABATIC TEMPERATURE-CONVERSION 
* CALCULATION! ) 
TR(1)=1000. 
DO 77 I=1,4 
77 TR(I+1)=1000.*TR(I) 
DO 88 NN=1,13 
XX1=0. 
TFD=300.+ 50.*NN 
WRITE(6,11) TFD 
11 FORMAT(///28X,* INLET TEMPERATURE = ',F6.21/) 
WRITE (6,101) 
101 FORMAT(30X,"EXIT TEMP.',5X, 'CONVERSION!',15Xy'DX/DT'/) 


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LvaUe SALIJATAM 
4 WOITIAIA JITABATGA JHT 23 T4MO5 
2uAI3 A 20.41) WOITOI2 AIMAUE owa-THona 3HT UL * 
AEeL) MOTTIAIA OT 
AIA BHT V 2OMUINGMOD 30 ajaMuUM. = 
JAU PRIM - 
OMAIHT A AO 2Twa12199302 JAUTARSIMST 
1 AOAYYTIDAGAD TASH ADA WOITOAVA O1 M 
(x 930 000t WAHT 
OOMSSIHT & AO 2THETIERSSOD  SAVTAAZSMST 
iM SYYTIDAGWAD THAN BOF WOTTINGA JIM. 
: (7 530 9001. WAHT 
WI 301X010 WOSAAT AO 230M FO ABSMUM 
ab 
Wil WADDATIW 2299K4 : 40 230M 30 ARAMUM 
_ @333 SHT 
~~ AUTAAIIMST 0334 
OH MM MI IAVZ2IAT JATOT 
Wrrrrrr rte tet ttt tts. 


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9 caneenenernaneanenennss 


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a . AOEMAT of 2108. By t 


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*> 
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= A 
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= - 


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81 


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64 


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92 


ac 


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39 


et 


102 


96 
94 


eee (CONT'D) 


TF(1)=TFD 
TT(1)=298. 

DO 78 IT=1,%4 
TF(I+1)=TFD*TF(I) 
TT( I+1) =298.*TT(I1) 
DO 99 NC=1,390 
TPR=TFD+50.*NC 

FHT =@, 

PHEAT=0. 

TP(1)=TPR 

DO 31 I=1,4 
TP(I+1)=TPR*¥TP(I) 
DO 74 IT=1,5 
TPP(I)=TP(I) 

DO 81 I=1,N 

DO 81 J=1,5 

C( 1, J)=A(I_J) 
IF(TPR-1000.) 64,64,55 
DO 61 J=1,5 
TP(J)=TR( J) 

L=1 

DO 22 J=1,;5 
FHT=FHT+(3e*C( lo J)+1-5¥*C(9,J5) +(5.644+2)%*C(7,/5) +Y*C( 10 


Le PPR OTROS) —- TEAS SS 


PHEAT=PHEAT +(C(1,J)+1-5*C(9,J) -C(2,J)-C( 6,5) ) ¥(TP( J) 


*—FET(C SIDS 


IF(TPR-1000.) 91,91,92 
IF(L-1) 93,93,91 
DO 95 IT=1,N 

DO 95 J=1,5 

C(I, J)=AA( I,J) 
DO 39 T=1,5 
TT(I)=TR(T) 

TFC I)=TR(I) 
TP(I)=TPP(1) 
L=L+1 

GO TO 64 


CALCULATION OF CONVERSION AT THE EXIT TEMPERATURE 


X1=1.987*FHT/(123924.9+1.987 *PHEAT) 
DXDT= (X1-XX1)/50. 

XX1=X1 

WRITE(6,102) TPRyX1_,DXDT 
FORMAT(30X,F10.195XyF1023,20X,E12.5) 
IF(X1-1.0) 96,96,88 

IF(TPR-1000.) 99,99,94 

TF(1)=TFD 

TT(1)=298. 

DO 777 I=1,4 


238 


if ‘'T¥n}3) 
’ } , ses 
















aatets iat” ores 
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| Le ITHOMTAL 4D AT “«K 
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58, 08 SCAT SAAT 
vee 1 -.Oe THA 4 
OST ASHE a: 
ie os, WARS ee 
Hale Selel SE oe a, a 
(LLVQTRAITECL +! r9F. 
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at a Sie et 0 
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a,fel 18 00. am 
> Abed stb 9199 te” 
~ 225 Oded (. OO01-R9T IAT ma 
} cg Sefer id. 00 a 
AL DAT SCL DAT 
Pole <i los : ‘ted 
ae t,t S ag . 
OLII*V+ (Le TIDE Seoavede 44,0) 98e% [4 Ge LYSE JATHISTHT A 
| ; | .) ONG AT CU IST REL 
(LIDTI¥ EM eA I9=1b eS99~" tb ger *2sFo (be tVIy+ Udit ge WSS 
ee ; . bg) te UNe ade 
See Oe (.000f= 
TOREe REE (fad) 


; “y 


TARIAMAT TRB BMT TA. wore Wi 00 v1 
; oe 


eee (CONT'D) 


TT(1+1)=298 .* TT(1) 
777 “TFC I+1)=TFO* TF (I) 
99 CONTINUE 
88 CONTINUE 

CALL EXIT 

END 


239 


(O'TUBD) oc. 


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Coro Cel G poly Gy Cy Cor Cooney wk Mier ee Cn Cob Cov) Ce ECs Cy CN Ce cece 


WE HE HE HE IE HE HE BEC FS MEIKE EC AE AE DE OK FE FE EH OE DE IS IE CE OK HE IS OE OE IK OC OIE OIE KC SE BIS IK OIE DKK AIK OK IK AK ISIC OKC OK 2S 2} 


*K *K 
* MAINLINE BURN2 os 
xk * 
* THIS PROGRAM COMPUTES THE ADIABATIC REACTION PATH * 
* IN THE FRONT-END BURNER SECTION (2) OF A CLAUS UNIT* 
* ACCORDING TO REACTION (1.22) ASSUMING EQUILIBRIUM * 
* DISTRIBUTION OF SULFUR SPECIES * 
xx x 
* N = NUMBER OF COMPOUNDS IN THE REACTION ** 
x MIXTURE x 
** A = TEMPERATURE COEFFICIENTS OF A THERMODYNA~* 
* MIC FUNCTION FOR HEAT CAPACITY( FOR LOWER * 
* THAN 1000 NEG K) * 
* AA = TEMPERATURE COEFFICIENTS OF A THERMODYNA—* 
* MIC FUNCTION FOR HEAT CAPACITY(FOR HIGHER * 
* THAN 1000 DEG K) * 
> Y = NUMBER OF MOLES OF CARBON DIOXIDE IN THE x 
* FEED x 
ok Z = NUMBER OF MOLES OF EXCESS NITROGEN IN * 
* $e FEED * 
** DHa@= HEAT OF REACTION IN H2S-SO2 REACTION * 
** ioe heats OFeREACTION IN 3582-S6 REACTION * 
* DH4 = HEAT OF REACTION IN 482-S8 REACTION me 
* TAD = "FEED - TEMPERATURE *K 
** rl = TOTAL PRESSURE IN MM HG * 
x * 
HE AS DK DK IS AE IE IE DIS DIS DI DK AIS BIE OK IE SICA IK AK DIS OE OE OK BKK IS DIK BE OK AE OE DIE IK OK DIE DIE 2 OK IS A DIK IS DIK IK OK OIE DIS SIS OE IE IS OIE DIS 2 


DIMENSION A(20,5),AA(20,5),TF(5) 49 TT(5)_TP(5),TR(5) y 
LERT ECS) sFHTH (5) 5PHTLU5).PHTH(5) 
READ(5,1 ) N 
1. JEORMSA fa) 15) 
DO 3 IT=1y,N 
3 READ(5,5) (A(TyJ)9J=1,5) 
5 FORMAT( 5E15.7) 
DO 4 I=1,N 
4 READ (5,5) (AA(I9Jé)5J=195) 
TR(1)=1000. 
DO 77 T=1s4 
1? TRUS y= L000 ee TR (CP) 
DO 88 NN=1,23 
READ(5,8) ZY yDH2,DH3,NDH4 
8 FORMAT(5F10.2) 
WRITE( 6,10) 
10 FORMAT(1H1,25X»y 'ADABATIC TEMPERATURE-CONVERSION 
* CALCULATION!) 
READ(5,54) TFD,PT 
54 FORMAT( 2F10.2) 


240 

























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A ke "». 


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: . J ra! ha ss ? i. 


* TAD WNITIAIA OITABAIGA FHT: 2471 9019 ma aonsa eso Oi * 7 

STIMU 2UAS) A AO 4S) MOTTIA2 AIMAe -—CMaaTvORA SHT WE 

+ MUTABTIIUOT OMTMUE2A 1S 6E) UNETIASA OT BULOANIIA ¥- ws 
2319992 AUT IU2 FB MOLTUATATZEO | * 4.0 i 


5 


SA 


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wo 
~ 


OYTIAIS FHF UI 20MUG9 M0390 ATA On = “. 
eh | SAUTXIM | So 
‘AUYGOMHIHT A AQ ZTHALSIARAND AMITARSSMAT = ~A- 
f4WOJ AOAVYTISAGAD TASH a AOETIVUA JIM -  s 
(4-930. OOOL WAHT © 
H<AWVYOOMRAHT A AO eTHdTaL aod JAUTAAIIMAT = AR ~) 
‘RIHDTIH ROAVYTIDAGAD TASH ‘on WT T juua 91% 
a . (» 930, 000f WAHT re 
* 3HT Vl AGIXGIO WNAHAD 30 2a30" AD AAEM =. 
° Se “CAST. ; 
* “Al V2dOATIN 2239%a9 3M 2990 A0> Aen = 
| 349 3uT ne tp 
VOLTIASA SO2-28H VE AI TIA: 23 a TAA = SHO. 
1a s 


be 
~ 4 
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7 
— re 
wu os 


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i ADT TIAIA B2-S$2e WI VOTTIAIA FAO. TAIH = aHQ 


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241 


eee (CONT'D) 


WRITE(6,11) TFD 
11 FORMAT(///28X," INITIAL TEMPERATURE = ',F6.1/) 
WRITE (6,101) 
101 FORMAT(30X,"*FINAL TEMP',5Xy_ CONVERSION! ,15X,_'DX/DT!/) 
XX2=0. 33333 
TF(1)=TFD 
TT(1)=298. 
DO 78 J=1,4 
TEC 1T+1)-=TFD* TF (1) 
78 TT(I+1)=298 .*XTT (1) 
DO 600 J=1,5 
FHTL( J) =20e*A(1lyJ) tA(293)4+(52644Z) *A(7T JS) +Y*A(10—J) 
*+A( 6, J) 
600 FHTH(J)= 26 *AA( 1,5) +AA( 2,5) 4+( 5.66442) *¥AA( 7,5) +Y*AA( 10 
*y J)+AA(65J) 
M=1." 
DO 99 NC=1,30 
TPR=TFD+ 10. *NC¥(—1.) **M 
FHT, =0¢ 
PHEAT=0. 
TP(1)=TPR 
DO 31 I=1,4 
31 TP(I+1)=TPR*TP (I) 
PS2=PT/760. 
CALL FREM(PS2, TPR _XN1_XN2) 
DH=2.-*DH2 +125* XN1 *DH3 +125%* XN2 *DH4 
DO 500 J=1,5 
PHTL( J) =2e*A( Le JIFA( 29 J) —20*A(69JS) —-1Le 5¥( 1 e-XNI-XN2) ¥*A 
1(5,J5)-0. S*¥XN1*A(45J5)-0-375*XN2*A (35 J) 
500 PHTH(J)= 26*AA(1,J)+AA( 29 J) -2e*AA(65 5) —-16.5%*(1.-XN1—XN2 
1)*AA( 5, JS) -O0. SX®XN1*AA(4 55) -06 375*XN2%AA( 3,5 J) 


FOR BOTH TFD AND TPR LOWER THAN 1000 DEG K 


IF (TFD-1000.) 335473355426 
Som 1 (1 PK=21000%) 326,336,691 
336 DO 337 J=1,%5 
FHISERTEERI LES) ATP UJ)—-TPU)) 7d 
337 PHEAT=PHEAT+PHTIL(J)*(TP(J)-TT(J) )/d 
GO TO 91 


FOR TFD HIGHER THAN 1000 DEG K AND TPR LOWER THAN 
1000 DEG K 


426 IF(TPR-1000.) 427, 427,560 
427 DO 428 J=1,5 
FHT=FHT+(FHTH( JP *(TFE(JY=TR(J)) +FHTL( J) EC TRIS) TPC) )) 
*/ J 
428 PHEAT=PHEAT+PHTL(J)*(TP(J)-TT(J))/J 
GO TO 91 


elt TiANIS Y° 
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(147 ag aT= 41 414aT - 
CPD TT 8. 8OS= (£41 TT ay q 
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eee (CONT'D) 


C FOR BOTH TFN AND TPR HIGHER THAN 1000 DEG K 


560 DO 561 J=1,5 
FHT=FHT +(FHTH(J) *(TF(J)-TP(JU))) J 
561 PHEAT=PHEAT +(PHTH(J)*(TP(J)-TR(J)) +PHTL(J)*(TR(J) 
*=-TT( JI) 75 


GO TO 91 
C FOR TFD LOWER THAN 1000 DEG K AND TPR HIGHER THAN 
C 1000 DEG K 


691 DO 692 J=1,5 
FHISPHT +tePRIM(JIe(TREISI-TPII)) +FHTLO J) S(T RUS) 
w= TR PL ee 
692 PHEAT=PHEAT +(PHTH(J)*(TP(J)-TR(J) )+PHTL(IJ)*( TRIS) 
mee TT Ca is a 


C CALCULATION OF CONVERSION AT THE EXIT TEMPERATURE 


91 X2=1.987*FHT/( DH —-1.987 *PHEAT) 
X2=0. 33333+2.*X2/3. 
DADT = Um ARe 10 * (= Le) FFM) 
XX2=X2 
WRITE(6,102) TPRyX2,NXDT 
192 FORMAT( 30XyF10.195X9F10.23912X,E12.5) 
TFC TRD+950..) §1000,999,999 
1000 M=2 
GO TO 998 
999 M=1 
930 Tho UK2=1.0) 599929758 
99 CONTINUE 
88 CONTINUE 
CALE EXT 
END 


242 





















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APPENDIX D 


ESTIMATION OF EFFECTIVE DIFFUSIVITY 


243 


“y 





t= 


x £1 i 


rad 
_ a. 
+ ie 


La 





ae 


° 
4a 4. 





AA: 


* 
te 
> 


aS 


i 


is 


~ 


a 
a 


oa w 
° 


dq aVI'TOUTIS IO MOTTAMI 





Va 
p ? M 


ear 








‘Al 











on 


oe Hey 


ry a, 
Py ae aes 


ea @ « 


>¥ vr cT y 
ait 4 


TIVIEe 


‘ 
- 


244 


D-2. Molecular Diffusivity of SO,: 


To estimate the molecular diffusivity of SO, 


through No, Wilke-Lee correlation (1) can be used. 


(0.00107 - 0.000246 | - 7 i) p3/2 ahd sages 
Pana wih. iy tis Metout a Caoenney oy. 
2 


Pap = 
m ht) DLT, 
Het hetaieie a tas as) 
where 
M, = molecular weight of SO, = 64 
Ma = molecular weight of Ny = 28 
é rn + ra 
rapR = collision diameter, A = 5 
pel for sol eae 290% A 
A 2 7, 
° 
rp for N, = oeGoL, A 
F ' : : kT 
I_ = collision integral for diffusion = — (— ) 
(2) 
€ E € / 
Seats K % Wiel 252) 151.8486 
At 
T= 550°, 
ete ees 612; ‘then Ta ~ 0.45 
SAB 
For m= 1550°K,. 7 = 1 atm, 


Shue bee cCeRs) andslee, Cov.ik 6 EC, 47, 1253 (1955),. 


(2) Perry, J.H.; "Chemical Engineers Handbook" 4th ed., 
14=20,21, 1963. 

















. ; : 
bees ed asp (1) mots sien 300 per-osl iW a pound. 
“sa 0 Die age ae nc 
ae : xi | — (a E +e | 348600.0 - - ‘youd. ne gee 
gf 
at on aie ee ; ie 
3 ; ene Mes: < 
(f.a) alt <a ee. ‘ res we 


bd = 02 to Rsn 3s shipootom> ae 


es - hf ore steiow xeiveaion = ee 





‘gy 
Ce ea - = 

° . a 4 a og 
KB oes bs = one a Je i ot. 
; rele Bac me Bi 










=> 


4 
_' 


pe 


- 


me 


- 


(Sy es nokat2ib 103 sexeoont coteisioe * Pp he ; 
“a | | Hania ay ‘3 a ae 


‘ ont 3/2 
0.00107 - 0.000246 | 55 + <4 } (550) =x + 


a (3.9855)* (0.45) 


I 


0.4215 cm/sec 
The above estimated value was used as the molec- 


ular diffusivity of SO, in the Claus reactant mixture by 


2 


assuming the Claus feed gas as a binary mixture of SO. 


and N,. 


D-2. Knudsen Diffusivity of SO,: 


From equation (2.23) 


1 
2 


: ) (2 .23) 


Daa 9>7~—x 10 Ry ( 


oes 

My 
° 

where R, = average radius of a pore = 80 A from the data 


of Chuang (18). 


Then at T = 550°K 


v 
Sci ae 


5 =o 
Porn reese Uo), CBO see) ) (=—q) 


oO 
Il 


KA 


O022 75 em”/sec. 


D-3. Effective Diffusivity: 


To estimate the combined diffusivity, equation 


(2.21) can be used, 


245 


















— ie {- nee J 
ba os ¥ ae. So BE Via 
JILL Beit Sere 
{< . U (228e -£) 
sau aad 2msh.0 = 
= Cee 
Kets aktd ‘Ge eee uy {sv pedauks ee evode ott 


* 


331 b asl, 
. ‘ aa 
ens ID eis ni 02 to yoiview i AB fR 
a a 
otxtm yisaid 5 38 nat) beet, evel) odd paimees 
é - J ikPs _— 
co hagy Ga hee 


: . ‘ta 
al Bo 


+. “. 
‘ s ~ 


vd sivtxim jassoBet 


be 


im 
ay ad = 
ae 


fpr | si 
:.08 to ytivieuttia asebur 





al > 
aa Yo * ; ; A ae ~ 
~ ; * a? *% 
(ES .&) nolssupe mort 
Ph e 


-(€8.58) 


ae 


; ‘ . 
sisbh odd mort A 08 = Sz0q 8 to ewibex opsievs * ; 
= a nobel 
“tee. Ne Uy 


x02? . 7 aire 
a 


St. OL x < 00) ( aee) dls 






2 


Sui 


246 





l-ay 
sea (2.21) 
AB KA 
where 
we 
O' Neaeng A= 0 by assuming N, component is stagnant. 
A 
Then 
1 iL af i 1 
— 2 + —— = + = ==) «46.3285 
D DAB Den 0.4215 0.02275 
or 


D 


0.02158 cm2/sec. 


Now the effective diffusivity can be estimated by 


applying the parallel pore model or equation (2.20), 


D, = D(=) (2.20) 
where 

€ = void fraction of a catalyst pellet, taken as 
0.35, 

tT = tortuosity factor, taken as 4.0. 
Then 


Dee (0.02158) (f°2>) = 0.001888 cm/sec. 


ahs , ie | ae 
I 
: 












i ‘ ' it 
2 . é Se cto, (Ah J ‘ 
(15.8) ‘ : ie - 4 a 


cBSE.OR NSC. eeeae 
ce. 
a 


muah "mS g2150.0" i re 


i+ peng) ip “Sy 


yd bss smitas ed nso voivieutiip ovisoe%e ond wou 3 


ie 5 


, (0S.S} motssups x0 Lebom a709 fet iexsg edt & 


(05.8) ee ¥ ar ‘i pei * 


QB sinha gelleg seylssso 6 s 30 otdpeat 8 haa 


2 7k See. - 

Lt +s , 
iS A eed oe 

@ meg > ’ 


.0. bas noted .x0808% sas 
i (rae AS 


Se Ty 


APPENDIX E 


SIMPLIFICATION OF MODELING EQUATIONS 


247 





E-1l. Modeling Equations: 


General modeling equations are presented in 


Chapter III. 














2 
ac Ak ac 
Ae nM am £ 
Ve —— + (C= C= 10 = 0 (3.8) 
int dZ Ep £ s L az 
h A. nD 
| nae we B s 
Pr Cot ie az eer ihe Ty) B (T., Te) . 
eecoeoeveveeeee (3.9) 
An Kn (C. - Ce) + HAD. esLg “ T,) =? 0 (S80) 
ar. 
An h i - Te) + n° AH-p.r. LE ee TS) - k, a = 0 
eae eae Coe. is) 
where 
5 
Peep wauest S03 
ry (P. TS) mole 20. xX LG 5 
J (1 + 0.006 P ) 
HO 
2 
= 0 
exp [| 22° | fk) 





gs 


248 


BBS ee oe 


| \ = tS 
(g.€) | ae aaa 














dA 3 | 
0 = (,T -42) 3 * Tet See an 


(@.£) seaenge hee ‘. - . a 


(OLE) St _ + tg es 


) 


~ ~ 


S 5, fs 
2 - 
0 pie ew (oF ya 


. Sia ae ee 
X£f.€} ; 2 & O60 86 OY! ie a ei 


ec Wie 


_ @ - - a Sey eT 
( he 300. 0 +, a - ge fen heap eh 
O<8 ae f : ure % 


249 


E-2. Simplification of Modeling Equations: 


i) Assume the axial dispersion is negligible compared 
to the convective mass transfer due to bulk motion, 
then the last term can be neglected in equation 
(3:8). 

ii) Assume the axial conduction through catalyst pellets 
is negligible compared to the convective heat trans- 
fer due to bulk motion in the fluid phase, then the 
last term in equation (3.11) can be neglected. 

iii) Assume adiabatic condition in the reactor system, 
then the last term in equation (3.9) can be neglect- 
ed. 

iv) Assume the stoichiometric feed ratio of HS and SO,, 
then the reaction rate expression can be rearranged 


by the following relations. 


E = Pp ec ter2* ¢P =e ) 
O,1 eats SO, 


where the subscript i represents the initial condi- 
tion in the feed stream. The reaction rate expres- 
sion, therefore, becomes 


x 4P Te) 





Ss i ee Bi? 
P 
-4 #P. 
= 2.56 x 10 5 
{1 + 0.006(P,, o,i* 2 Be te 2 ae ) } 
2 2 2 


exp ( RT ) 
gs 


ens ae 4 LAE ue. he 






















:3noitseupe F LLepea 
— 7 7 al , an # 
hexBqmoo atdzpldved al Hoberege tb eter Lh: 


,noisom Aind of emb. wolensid Pee ev tipevace ‘eth 
not tsupe. at betoelpsa od m9 Pam dent ‘ead 


3 , : : ; ey ; ih To 
: ' an ” i, — a ime eo) _ 
} ‘a ive 
eteileq seylsssS fgvoxd nokvoubads ‘aloe Bi y 


~ansus Jeo0 ov Ltosva0o ans of bezaqnoa ars deh ek % 
‘eit? nedd ,eandg Biolt sxitt at dobI6m Aiud o¢ sob iat & 
.betoesipen ed aso (£L,€) soltsupe nk ied deat ‘ 

meteye 1otSkOT eri ak nots Laos oisedhhbs a 


«toelpen ed 159 eve) noktsups see geal of eit. , sy 
or 


Me 


108 bos 8.8 to ottex best gms s: ss he oe # 
heprstxser ad 169 noleaeiqxe ete _natiosex © a cn) Pe 
-enoltelox eno HLoR + xd 






~ibaoD isisint ent) esneaoxqe7 at 


sannee et 51 noitoses oft + ME on 


. a cage 6 foe 
y # a 2) ess : ‘ 
nat i y : 

4 ‘e's \ ie ie a 2 i : z 


& 





250 


Now introduce dimensionless variables to simplify 


the general modeling equations with the above assumptions. 


Let 
a 
L = ? (E.2) 
ey aa eg Coes C 
afd ge 
a Des ee Tae Ben ts ey 
, e 
. Tey . Tei 
Then 
Ce = Ces (1 - X-) Cc. = Cea (1-X,) (ESS) 
Peer Leyes Tarset a (li —-@.) (E.6) 
By combining equations (3.8) and (3.9), 
An kn (C. - Ce) AH - Ay h wre ve Te) = 0 
or 
FR hg as Ge 
Ded Ce Col ee eres. te (E.7) 
From equations: (3.6), (3.7) and. (E. 7), 
dC, 
as Vint Gaze = An ee (Co - Ce) 
Yea aT 
ee = Dr G f S Vv, £ 
O2.C aT 
or os = _f “pf = = (E.8) 


Qu 
N 
> 
0 
=| 
N 


ot) Ei yy See 
. ae ae x, oe ° 
Bes Ais , 
ie 


ve ’ 
he) 


ma, 


aa “a e 

e2Ba Ts ? y oe — 

i ¢ a ‘Tae s _ 
[ a i ; 










eoldeiuey eaplaolenemih souboaseih. wom 


144.2 


svods aA ny iw’ anol +80 e pai Lobom texenoe of 





~ ¥ _ > ae 
— 78 
a ag 
0% a ia ; a” P (4% : 4) HH 4 a " 
1 | ie " 
aon : i 2! : {, 9 ‘= a ri 
» {@.£) bas sal Ey enolssyps pai 
, A ia Snub 


a . ; x “pe 


a 


ne ro 





By using equations (E.5) and (E.6), equation (E.8) 


integrated to get a X - Oo, relationship. 
X Vg ee i 9) 
Q Cg ace _f “pf fi f de. 
AH 
0 0 
_ Pe Spe Tei 
Sak. ts pl ogi 
AH 
AHO i. 
2. 
or 61. = ——- SCO™F*X 
f Pr Cf Tey £ 


ak 


can be 


(E.9) 


Equation (3.6) can be described in terms of conver- 


sion by employing equation (E.5). 


d 


rian een Se te) = 
dz 
Aa ae tGe, Vouk) Ce; (1 - X_)5 
: at = cae see (Sein eX) 
pr er CL Ep Vint Ss £ 
or ng ee A Oke eX) 
aL 1 Ss f 
LeA sk, 
where ae ee sere 


From equations (E.7), (E.5) and (E.6), 


An Kn Ce (1 - X,) - Ce (l - X 5) = 
AL h 
Rem eer ee rere: AL. Big) 
An Kn Cel 
Bealvae ine hee (X_ - x) AH 


(3.12) 


(E.10) 


(Bolt) 


a7 Ke 4 
a : Poss, . ¥ 

. ia * oa : ; 
rit 1, 


» 
mee Bee ae 


pega Re ey - 
ro toe 












ces (So oy oh} 
ed nso. (8,4) noiseupe Qa) ba (28) + 1 
sana ta is oa 8 we 


7e “ 
4 ’ Pe 


“en oe) a sy is 
| Pod 


aad 
a 


} eat Cea i 
wtagat atus ae 


> 


| A a ge See ae Le 
4X fae —- Lae aa” 
a aq. 3 See 

ot ak ‘Beadavasb ed-tee (.£) notisupS ~ 
- 


A208) ackavure ential ia 


(@.3) 


-1rsvooo to amt 


CK =) (Eee. 
2 € BA. 

. 4SL,€ | a at 
(SE68) ee 
(OL.3) ¥, 

¥ fas ' sf = ey, z 


“5 ay oem (2.9) Ba 


By introducing equation (E.9) to equation (E.11), 


Vike ih eer We An “m S£i be tee) An 
Pe ORE Tey cr s AL h Te [ s 
8. = (A, ~ Ag) X¢ + Ag X. (E22) 
AH Ces 
where A, = (Be. 13) 
Te. 
Pf pe rt 
An kn Ces AH 
aa (E.14) 
8 A, h Tes 


Now equation (3.8) can be described in terms of 
conversion only since 8 and T, can be expressed in terms 
of conversion. 

ae .. Ry Ts a ae 

= Pea (1 - Bs) (l - X.) 


= Se AN (l - 65) 4 (l - X.) 


ae 


= Pp, {li- (A, = Ag) X,°- Ag Xo} erie) 
(E.15) 
T. = TE (l - 65) 
Sed at (AJ - Ag) X_ - Ag Xo} (E.16) 
Then the reaction rate expression becomes 
oy ay ae Ag 
eerie Ts oe eee fe) (E17) 
s $58 B Cement Ci iF )? f. 
4 Seal’ ae 


zoe 












: ra Pa b \ 
vr : - et Ba ‘> Fi uae: rs 
r A i.e ‘i Bot al, 
Lseupe oF (&.&) noiteupe — a) a 


 (iL.a) mor 
oa 
yi 


aa 


> ot A | oro a 7. 
HA (:x-< hh. See oe ee ae eee i 
oe T A 2 rat POO 
a = ere ta . t tq”. a 
% oh + aX (Ga = ba) = 8 
(oi. cot a * ih | 








+ at beditoseb ed meso (e. 2) sokssupe won 


ae 


nf bseasigxs od aso fT bas | at sosie yin peeey. 


ove te 
» 


7 


(@L.4) 


i > 
F) P =e > di 
(ai.3) x 8% * ( “A> re 


- 


se 


eemooed noleee7g 


where A, = An Kn Ces 
meee saiet 10. | peers 22" 
A, = 1 + 0.006 PE o,4 eae 
Ag = 0.012 P,, 
rs I 
Goce 
f,=1-x, 
fo ie le a (APR ele Kee Ag, X. 


Finally equation (3.8) becomes, 


A, ty hg , 
Pan ee inlay Pie eee RN A ade Gg of 
2 1g s esd a \¢ 
4 mee ig 


(E.18) 


(E.19) 


(E.20) 


(E.21) 


(BE, 22) 


CH 233 


(E25) 


Go ako) 


The computer program MODEL, used to solve equations 


(3.12) and (3.13) simultaneously, is listed. 


253 





a 


ssemoned (8. é aolsoupe: ‘yite 





>= 








i 
* 
a 
£ 


x base pA. 
es.) 
i - 


, f ie a sted 
$ Kae 
1 g ‘4 = 
f q 
=F SI0, a aA 
AN f > ge 
a : a 


we i- an 


~ 
~*~ 


uy  dadoM crsooxa soiogmon 


I el ae 


as 






> 


ac 


Cee Gla rGa) CovaGed NetGy C Pak CHIC) COC Cy Kae) te) Co C) CY LY CY CGY ED CY GC 


LZ 


* 
ste 

* 

* 

xc 

* —PHASE MODEL 
* AL - 
a 

* 

x AH = 
sk 

* PT - 
* DP = 
kK, 02 = 
x EB = 
* DB = 
*  DENSP = 
* DE = 
* XL - 
*  XHS = 
* XHO = 
* XSO = 
* XS = 
*K VS = 
i OUT = 
ok 


THIS PROGRAM IS USED TO PREDICT THE PERFORMANCE OF 
A CLAUS CATALYTIC REACTOR BY A ONE-DIMENSIONAL TWO 


BE ASAE AS AE IEE IE CC EI NE KK ROIS CE OE OK FC OE AIK OK HK AE A FC IC A OK FIC IS IE OIE IS EK IC OE IK IC OK SOK OK OIC OIE OK OK OK 


MAINLINE MODEL 


HH HHH HK 


TEMPERATURE COEFFICIENTS OF A THERMODYNA-* 
MIC FUNCTION FOR HEAT CAPACITY(FOR LOWER * 
THAN 1000 DEG K) * 
TEMPERATURE COEFFICIENTS OF A THERMODYNA-* 
MIC FUNCTION FOR HEAT CAPACITY(FOR HIGHER* 
THAN 1000 DEG K) - 
TOTAL PRESSURE OF SYSTEM(MM HG) 
ULAMETER OF CATALYST PARTICLE(CM) 
INCREMENT OF INTEGRATION 
POROQCLLY* OF BED 

DENSITY OF BED 

DENS hy OE CAPALYS. PARTICLE 
EPEEGULVE. DIFFUSIVITY (CM*S*2/ SEC) 
TOTAL DEPTH DF BED 

MUSESTKACTION OF HZ2S IN THE FEED 
MALES PRACTION ‘OF H20°> INTHE FEED 
MoCeernnGi.on. OF S02 IN THE FEED 
MOLE FRACTION OF SX IN THE FEED 
SUPEREACIALS VELOCITY OF FEED 

FEED TEMPERATURE 


Hs HHH 3H He He se Ht tH He 


iv 
w 


HE DEAE BIE BE AE EE IC KE EE AE KK AK AE OK IE IC KC OKC OK FS BIE OK DE IE IE AE AE I OS FE TIS EOS OK I A IE SE IS IC OKC OK KK IE IE OK OK 


DIMENSION TT(5)yTR(5)_ ALC 795),AH( 7,5) 
COMMON AL,AH 


READ (5,12) 
FORMAT( 112 
DO 3 JI=1,N 
READ’ (5, 5) 
DO 4 J=1,N 
READ( 5, 5) 


N 
) 


CALL st) bis = Lie at) 


(AH(I,J)9J=195) 


FORMAT (5615.7) 


READ( 5,10 ) 


PT,»,DP,NDZ,EBy,DBy,DENSP,DE XL 


FORMAT( 8F10.4) 


READ( 5,11) 


XHS9XSOyXHN,XS 


FORMAT( 6F10.5) 


READ( 5,11) 
TFO=TF 


VS,TF 


AM=6. /DP*(1.-EB) 


AT=AM 
R=1.987 


254 






~ 
J 
— 

a 


4 
oe 


= 































ty a rd : 
; q ste! ‘ rad 
Pee yi 
, Mi i ’ ] \ , 
‘ a 
‘ . " ‘ee = 
se he she ae age ac oA ae ste ae aie ae fe ae ge aie a KEKE HS OE oe a 7. 
Pag e fe . > ; we 
{400M ANT IJVLAM .s ae 
y 7 *.. { Fj 


ar IWAMAOAASS BHT T7193989 OT v320 8 21 “MAR DONG - 2 rat a 
OW a |AAOT2ZHAaMIQ-3a00 A. YA ‘BOTIAS s SETYSATAD 2UAd)- A 
| | J300m Bea's 


I 3303 pas miranaal 31 a ae ro. 

are 07 vin! ramus 2IM2 4 4 Ph: 
(A 930 OOOL.WBAT 255) i 

fs 'YOOMS3HT A 3Q 2TWS1014 ami 3 AUTARAIMST 2 HA ee] 

*AAHOTH AGFVYTIIAIAD TASH #OA “nt td Hp Gof ras ‘ ves . 

¥ (x 3a0 COOL WAHT 

% (oH MM)MAT2¥Y2 90 J3Av2eTRS JATOT | 7a ety 


4a 


(MI)SSDITAAT T2YJATAD 30 AITSMAEG = qai* & 
VYOTTAAOTTVT AQ THAMBAQWT & SO eae 
o38 WW YT 190809 = as * 


age ae ¥TF2naIG = a 
2¥ JS TAD AQ) VFI 2va0 = - aenaG  * 
4yiz(aaid 3V1T239A2 > 3 oa 
a3a 40 HT930 SATOT dk , 
30 WOTTIARA 3JOM = 72h 


auDTT aN 
37 7 


rh ~ 


7 3HT AL 28H 40 aes 
osaa. set AL OSH 30 MOTTIARA AJOM = . OF en 
Q354 aHXT Al $O2 a0 WOTTIART 308 3 3 
: 3393 3HT Wl X2) 30 VOLTOARS SION = t 
* 9333. 40 YTIOOISV JAI D1 IARWE.2 eee 
3) AUTA AIIMST 'Q339 = wana ak Ei * 


un 


: ' ble heya: 
tee de a eR ewe a ep eh ae a ssnennanengennes adie 
eal 


a a Here ey ee 
4x He ak Oe NK oe ‘? 
7h a) -_ \ 


[2 et HAG (By v 1A stat CIT a 
5 ee a wt +P 
ae 





Bon 
i 


iy 


+ ‘gee 
(Btn by ¥ me & 
ABs Fated La 


- 

























’ a in ‘ - 
J%,30 q2nad, 14083450) 
- ’ iv > he ty - 
& ty ee oe i 
¥ vs F Pk a 
| » 


a : 


ie 


20 


27 FORMAT(1H19///914Xy_ "LENGTH", 8X5 !XFPy8X_ 'XC89 7Xy "THF! gy 
17Xy "THC! 8X ye" TE 8X9" TC8 yg Xo *XN1" 45 7Xy XN2Z! 9 8Xy DH //) 


et. 


gl 


44 


88 


as 


eee (CONT'D) 


RR=82.06*760. 
VINT=VS/EB 


WMOL= 28.*(1.-XHS—XSO-XHO-XS) +34.*XHS +64.¥*XSO +18.* 


1XHO +64.*XS 
PFO=PT*XSO 
PHO=PT*XHO 
CFO=PFO/RR/TFO 
PS=PT*XS 

A3=1. +0.006* (PHO+2.*PFO) 
PF=PFN 

PS2=PT/760. 
TO=298. 
TR(1)=298. 

DO 20 J=1,4 

TR( J+ 1) =298.*TR(J) 
WRITE( 6,27) 


XF=0. 

DO 1000 IITI=1,1001 
TT(1)=TF 

DO 77 J=194 

TT( J#1)=TF*XTT(J) 
DENST=PT*WMOL/RR/TF 
VISCO=0.00025+0.000034*(TF-473.)/100. 
CPE= ALT Ts i} 

DO 21 L=1,4 

CPF=CPF+AL( 7,L+1)¥*TT(L) 
CPF=R*CPF/28. 

XK=0. 2976/3600. 

D=0 .010996%*TF**1.5/PT/0.44 
G=DENST*VS 
SC=VISCO/DENST/D 
PR=VISCO*XCPF/ XK 
RE=G*NP/VISCO 
RRE=RE**0.41-1.5 
¥ID=0.725/RRE 

XJH=1.10/RRE 
XKM=XJD*¥G/DENST*SC**(-2./3-) 
H=XJH*CPF*¥G*PR**(—-2./3.) 
IF(III-1) 88,88,44 


CALL FREM (PS2y,TFeXN1y,XN2 ) 

CALL HTORN (TF yXN1_9XN2_ Ry TTyTRyDH) 
GO TO 55 

XN1=0.135 

XN2=0.865 


DH=11250.-1.5* (XN1*22673~6 +XN2*24753.) 
C1=AM* XKM*C FO 

C2=NB*PFO**1.5*0.17 *  2+56E-04 
C3=1.+0.006* (PHO+2.*PFO) 


255 





















; A. a in 
+087 son sseans R 
| r aaeveraly % se 
wae ZHKw SEs X-OHX-N12X- 24k EIR 8S ep . ris 
. exe, feb 


ale re ae: , “2 K8T4 jens 
eight Vk Poe oneereaone 


ot as\nad at : 


- + 2XST9 224 fe > 
(Q49# .S#0H4) #300,0+ sfeEA 
Pe Pea ; HAs ead 


Renton eee © fact 4 


| ae aineuns ae 


ov ot ele os 0 
~ (LVAT# BOS thie is: 
ce) Oe yatt 
t XB e MHT DMA OL WN eLHL IT. ed 


tAHT! KT eh DK eh hea 
ENN HHO X80 FSU ext ie yersaapee er Kt VAT fy XBq ‘DHT! « vi 
O=AX 
HOO aH EL 0001 90 
f “ATs rT 


sores rvs yi feb: ¥ 
eT wa renatetss , 
> oo ATA RANUOMHATS Tava 


.OO1\ (ETe=AT) | #350000 40485000 06 ae 
e 7 i - 


25 


1009 
1001 


256 


eee (CONTIN) 


C4=0.012*PFO 
C5=-7350./R/TFO 

A1l=XL*AM*XKM/VINT/EB 

A2=DH* CFO/DENST/CPF/TFO 

A3=C1*DH/AT/H/TFO 

CALL NEWTN( XF 4A2,A39C1,C29C39C4_C5 XC) 
Z=D2*(ITI-1) 

THF=A2®XF 

THC= ( A2-A3)*XF+A3*XC 

TF=TFOX( 1.-THF) 

TC=TFO*( 1.-THC) 

WRITE(6,25) ZyXFyXCyTHFyTHCyTF 9 TC yXN1_¢XN2q DH 
FORMAT( 10X,9F10.5, F10.2 ) 


CALCULATE BULK FLUID CONDITION 


CALL RKGS (A1,DZ_yXCyXF) 

CHK=1.0-XF 

Bret TABSOCHR) —1.0E-07) 1001,1001,1000 
CONTINUE 

CALL: & XE. 

END 
























ir 


TANI eee 45 4 ; ee aan 
9994S 10.0809 © | 
| Bp wren Gd ae. 
aa \TVLVAMAX #MAKSKSEA e a 
Nat \A9I\T2 MIQ\OIIFHISSA: ake 
7 vai “DAT A\H\TANHOAL DEA 7 
LDN BICPDs ‘SDefDeEAs Se AX UTHAM JJAD ne tae 
= (L-hL Ee R083. 
3X*S AAMT 
F ORR EAHA EA a a 
na ; Ps ‘* (ANT =.5) 8O0ATSAT 
hy) | A DH ToL ROAT=OT 
HO «SUX, LAKe OTs 4% HT FAT DKe IK eD pf een 
( SeOL7 ¢@s OFAC. xOL, LMAOS 





wort amar onda awe STADIA. > 6 


UAK GOR Sona) 2048 | oon 

sn 4a 'AX=0 ef=HD > 

OO08 1 00L—¢f00L (O+30.1- OD Z8AE IAT 
Pal) eo = SUMETMOD 608 

| ee 0 


Cy CCG Cree eyo) eG 


HE HE AE HK DE HE HE DK AE EA FE CDI EAE DC AC KK DIC DC HE DHE DHE HE HIE AE IE AEC IE IC TE AE CIE AIC IKE FS ICO DIC AKC DK AK OK KK EAE OIE 2K IS 2K 


SUBROUTINE HTORN 


“ACTION OF THE CLAUS REACTION CONSIDERING THE EFFE 
=Gi OF COUIEERIUM DISTRIBUTION OF SULFUR SPECIES 


ok 
x 
A 
His SUBROUTINE CALCULATES THE STANDARD’ HEAT OF RE * 
AT A DIFFERENT TEMPERATURE > 


HHH HHH HK 


o 


ak y 
7 


HE AK IE EE AE HE AE IE TE IE OK OE I IK I A EAS OK IS IE HE IS AS OIE IS BIE IK OS BK OK OK OK IK BK AK KAS OK OS OK OK ISIE OK > 


+ 
4 
4% 
+ 


3 


SUBROUTINE HTORN (TyXN1_9XN2yq Ry TTs TReDH) 
BIMENSIUON TINS Ts TRUS), AH Ts5) 5 ALC7Ts 5) 
COMMON’ AL,AH 
N=5 
HR=11250.—-XN1*226/73.%*1¢5-XN2*24753.%*1.25 
IF( T~-1000.) 203920, 30 
290 DO 50 J=1,N 
50 HR=HR+(2.*AL (695) 4+1.5%*( 1e-XNI-XN2)¥*AL(59J5)4065* XNI*AL 
1(49J) +0.375% XN2¥*AL (35 J3)-20*AL(1,JS)-AL(2,J5)) *(TT(JU)- 
2TR(J)) /J*R 
GO TO 100 
30 DO 40 J=1,N 
40 HR=HR+(2.*AH(6,))4+125*%(12-—-XNI—XN2) *AH(5,J5) +0. 5*XN1*AH 
1(4,J) +0.375*XN2*AH(35J5)—-26*AH(1,J)-AH(25J)) *(TT(J)- 
2TR(J)) /J*R 
100 DH=HR 
RETURN 
END 


257 


% 
: A 
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| 34? 
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3H CAACMAT2 FHT: 237A JUD J® ‘t. qv FTuNRBv2 21ut <— 7h 
j ree \A 1209 WOITIAIA 2A. BHT FO een * Be 
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7 AAUTRAAIMSE TE TWAAIIAIG A diay 
whe is “ge : 7 9 

ae Ae aE a I te sg meee ie ae Me waseme enews mice Dn 
ciiadies an 

a 
7 


an SUX | PUK eo TH yinotH surtuosaus 


(HOgAT TT 
ea B48) Hb (DAT Tu dl = 
HA ‘eA MO ae 
} 7 “A A Bat + 
a.m Ger asnGudedeltett erase tik 088s a 
| aren 06 0S, 0S iene i= 


20) Oo Mei of a0 


{Lee )JAKISMK-PMK—, £8 faa Se He cen 
({U aS) SA= 0b gf PIARe Gb ED LAE SUN - 
aS ae ip a OP 
P| amet br 
a 


(UU SHA (by LYMAM S- (eltnesienete 


CYORMOAOGAM OA A DP 


258 


BE IE EE A AE IC OK IE IK IE KC SI EOE IC OK OE OIC IS AK IE OK OK OE SK OK OIE IK IS AE IC OIE OC OIC IK Ig IC 2k OK AK aK 2K IS aK 2K OK 2K aK OK 


% s 
* SUBROUTINE NEWTN xe 
*K x 
* THIS SUBROUTINE IS TO SOLVE MODELLING EQUATION * 
* (3.13) AND GET THE CONVERSION ON THE CATALYST SUR- * 
* FACE AT A GIVEN CONVERSION IN THE BULK FLUID USING * 
* NEWTON-RAPHSON ITERATION TECHNIQUE * 


“ 
w 


sk 
HK RE HK OIC I ik Se i OK I IC KOK aK DIK OK IK I IK SI SIS DK AC OK I OC 2K IK OK OS KOI SK IE OK OI OK OK IK OK OK OK OIC OK OK OS OI KOK OI ok OK OK 


SUBROUTINE NEWTN(XFyA29A39C1_C29C39C4yC5_yXMM) 
DIMENSION X(500) 

F1l(X)=1.-X 

F2(X)=1.—-XF*(A2-A3)-A3*X 

F3(X)=F1(X)*F2(X) 

F4(X)=C3-C4*F3(X) 

F5(X)=EXP(C5/F2(X)) 

ROAMH=CLIE(XF—A) +C2*FS(X)**1,.5*F5(X) /(CF4(X) ) *¥*2, 
ORDA] -Cl +C2*F5(X) ¥( 1 S*(FS1X) ) ¥*0.5%*(-F2(X)-ABS*FI(X 
1}) SCPS(X)VRF2. —(F3(X) FF] .5*2.% (C4*F2(X) +A3*C4*F1(X) 
2) S(F4(X))#*3. +AZB¥*¥CS*(F1L(X) )®*L SSI F4(X) RR2./(F2(X)) 
3¥*0 2.5 ) 

G(X)=X-F(X)/DF(X) 


INET TAL GUESS OFX 
X(1)=XF+0.002 
CALCULATE CONVERSION AT CATALYST SURFACE 


DO 99 1=1,499 
X(I+1)=G(X(1)) 
BSPSADSUE MIE LIRKR CL ve ACTH L)) 
IF(ESP-0.0001) 100,100,99 
99 CONTINUE 
190 IF(I-499) 88,55,55 
55 WRITE(6,77) 
77 FORMAT(/10X,'NO CONVERGE !/) 
88 XMM=X(I+1) 
RETURN 
END 




















, uteis 
“Ss 
‘ vig 
A ane 
“ = 7 
. _ a 
4 RY me wit te setae aie a ak we ae oe she hk aye aie ate i he oe tekek eM Heap phan 
j : ’ 7 mo 
“TWA AMI THORS! hele 


, \" 4 ms > aoe 

ITAUOSA DMUIJIS00N av jo2 oT 2k IAT Rave. 2 LHT * 

* -MU2 T2YJATAD 3HT WO VOLERaVVND SRT TS OMA fete Ee) * 

* SUI2U OIUIA AJUS FHT. WAL MOleRaVvuRD WBVIa- A TA 394A * 
IWOTUHSST WOITARSYT E. VOQZNRA RHO TH 

psi ak a ale ok we see sik ak Me ae as ea oR oe ae me nee dy ae ate. ak aie tea eal 3 ae a ke bola THREARAREE: 


(MMX eS DeP DED eSIe LO 4A SA ax} MTWan. auiruosene: ie 
. (002)X woravaMia — 

X= fat XoLa 1 

KeEAREEAM ~SA)#AX= fE(XPSA | 7 

| - (KPSPROX DAI KIER oe 

~ | AKVETESIMEDEAK IBA 

ACK) SFINGD 19K3 500989 hn 


SS#eC(X)OT)\ (RSIS, LROKIEIESIE (XARA DRAKA & a 
KX) LAeEA-( KSA) we, Om (XPE TFET SIXES Oe i=, sien 
(KD LASBISEAF TAI SIERO SES THR CUXV EAD “SHE CXIOTIN 16 
((XISTINS* (UX) OAIN|s LHL OX) £3) 


H2OWEAS “eile & 
‘ 2. a 


peice 


oe oo $0 at 


MS 10 223ua JALTIOT | 
| Genel x 
“ADAAAU2 T2NJATAD TA wntzRayMo2 “gtasuasaa a 
Be rere ee { e 
speciftl ti ag x. 


Aes Caen \ (Cnet +EK) CHAS 
Toa ea a2 00k oe Yar A2 


ie civ 7) 


« ae a 


‘ te 
4 ws 
— aa Vii! 








oh 


flan se 1 BR SA a 8 (CRA: 
ae h: Shae £ \ 


’ io) ‘out, xe \ ¥7 w 
aay (1308904 Taian FN a 
Sai an a mip A s beat a4 5 t 

‘ - . £41 ¢ 


> ia e -~ 


aa : 
' 3 
7 7 49 7 
iy. a oe 
ie iy 


CO Gitar auc) Caco) bo 


FE AE AE TE IE IKE SS HE IE IC A ONS IE AE OE AE OE FE HE OE BE IE IK KE IC IC OIE IS KC IS FIC 2 OIC A IC BIE IC IE OK IE OK OK OIC OKC IK OK OIE 2K OK 2 


SUBROUTINE RKGS 


THIS SUBROUTINE IS TO INTEGRATE THE MODELLING EQU- 
RIIGN Cae 142) "(RUM CI nGeeINLe?: 10 THE OUTLET OF THE 
REACTOR USING RUNGE-KUTTA-GILL FOURTH ORDER INTE- 
GRATION PROCEDURE 


we HH EH HH 


HE EAE AE EE OK OE TK EE IE OIC I TIS A OY I IEE DK DIK OK IE BS BIAS FC OS IS KC BIS IS 2 IC DIK AK FSIS OK BIC OK IS OIC OK OK FC BIC OK 3 


SUBROUTINE RKGS(A1,0Z,XCyY) 

F(Y)=A1*(XC-Y) 

D1=DZ*F(Y) 

N2=DZ*F(Y+D1/2. ) 

NZ=DZ*F(Y+0.207107*D1 +0.292893%*D2 ) 
D4=DZ*F(Y-0.707107*D2+1.707107*D3 ) 

Y=Y + (D1+D4)/6. +0.0976310*D2 +0.569035%*D3 
RETURN 

END 


ee ee 4 


Hu HH 


st 


ay 


< 


> 


259 


ett ok Ke ak ee 5 Aenean een ne mee hme eee q 


DAP. OF - OL ESTO. o+ ee + Y= 


a 
El 


* 

: et ah) 

2 xH AATTY naaue é ig Ae { : : " 
yar i 

‘7 aTAAOITHI OT Se anftuosaue sat 
OF rAd FHT MOAY 4S2eE ANITA * 
ated rTUN-SOML aa wtrau. ant Dasa * sae 1 

9 380039089 VOITAAD * 

y Se > ’ * 
er er pare eae a Ue me eee sie 


a 


(Yn DX SO+LAb2OAA “gu rTyoReU2 
(Y¥=D)*TASCYIR 
ae es. Poder te 
7 Se | aS \E0-4Y )A#SO8 
f SAE CBS ,OS 6 O+ fdaTOLTOSad+¥ )RSM= 
ATED TY. [4608TOL YO TeOeY IAS aan 


: rc 
: Foe} | 


way ss % 


‘ 


eee 
aw 


pout 























APPENDIX F 


ASYMPTOTIC SOLUTION FOR EFFECTIVENESS FACTOR 


260 


' ’ 
4 
. 
a 
i 
4 
> 
> 
"7 1 
4 Le 
= ‘ eo ham ats 
ee hore 
~ y 
: 
i J 
: j 
- 
; Py e . 
* 7 ee « 
j ty o * 
- wi edit 
ae 
é . 
. * 
: <- 
oo 
vf XIGUAITA 5 
* 
~~ 
t 
é af 
‘= * 
- " 
eters, ' 


™~ 


aviToo 


. 4 
' “ 7) 


2 
= 
, 
’ 
> oF é 
os, 
Pad 
St 
‘ a) 





F-1. Check the Validity of the Isothermal Catalyst Pellet 
Assumption: 
The validity of the assumption of an isothermal 
catalyst pellet may be checked by the value of the heat of 
reaction parameter, 8, at the inlet condition of the 


catalyst bed. 

DA 
e -s 
Keen 
e s 


RD 
II 


(— AH) 





0.001888 649 2)10." 


(26000) ( ye ae EL 
sI9 x 107 956 


He 


cal om* /sec ) (mole/em*, 


 eecpeg gear prarsoe) ad 
mole SO, can ay oe °K 


4 


174852 x 108 
= 0 


Therefore the whole pellet can be treated ina 
isothermal condition, and there is no need for an energy 


balance equation to calculate the effectiveness factor. 


F-2. Modeling Equation for the Effectiveness Factor: 


A general material balance equation within a 


catalyst pellet can be written as 


p 
roe o = —+ 5 r (P T,) ate (F.1) 
Bs (= 


261 












'ae5 oft Io yatbitev, oft 


maritort ng tq morsqnl 


av ors yd bexoeds ad \ ya solleq cod 


+ "Sei rozes%9q aoisoset 


+ %o notsibaoo setnt edt 76 48 
| | = hate 
bed sayied 59° 
a 


2 a rele) 
— <4 (BA #) = § d 
e 2 ts L Pe 
7 4 ; os a 
: ‘ +4 we. | 
: 





of x T.9, (< : “988100. 0, . (9008S) & : 7 
"OE 2 abit 


— ~ 
— j ‘ ' : oe 
\ f -” “Lat . : 
»\8io0n oSée mo “Lpo 
pmo \efom) (___ 228 -=—-) (a5 Slom ) 4 
2 992 °° igus a 4 
F * ‘ -- 
>= “. b A® oi a 
~ > _ 
oi x edd SS bs ‘ 
1 pe ~ i co 7 - ' * - nr 
" Je et . a) 
a ? La ce * 
a ~ - et a 
f 


T ed heited . tefleq tod ‘odd ‘asote 


i3 °. 


ey 


‘ypisne as 102 bs en on ai etedt bas “\sokaibnos Ie 


; easnevisoeit 9 od statueiso wh nots pe 


.tos0s2 ab 


y 


:tosos Tt _eeonevi3ge33a eds 20% 


it ar . ae 


~] A 


s mids iw noiseepe sons isd Lsiae: 





ee ie a 


262 


0 for a flat slab. pellet 


where a 


al for a cylindrical pellet 


se) 
II 


a= 2 for a spherical pellet 


To check the validity of the asymptotic solution, 
the concept of Thiele-modulus, ¢#, can be employed for this 


particular reaction condition. 





GRP aS, Ie heed 
Se ts ps s s 
* a4 pec (3.14) 
ea rs 
(oeee 160} phe 
where fo) (Bae) oS 
s SiS a 
Plt eee Eris o, i + 2 Pea - 2 P,)] 
7350 
EXP (- 5) 
g Ss 
(ech 105 e207) 6545)>°> 


(oe oedde. 0 Ole s 20.6545) - 


-7350 


EXP (3-997 x 556) 


Pee G Nea ati (g mole SO,/ 
sec-gm catalyst) 


ah 
eae ne L (923175, /- 053 x 7.6047. x10 ees aeas 
= 0.001888 x 5.957 x 10 


Since the estimated value of Thiele modulus is 


larger than 5.0, the catalyst pellet may be approximated by 
















| it ; ie 
i 1 5 ‘; 
a ae 
ee al q 


: 3 : nes, 
Jolleq dela Jef? 6 303 < 0 mm ener 


t ’e to re 
é ;* - < 2 yy. dae 
teileq Isotubnilyo 6 102  seuimgteh oe 
ye 
- \ 4; 


tol iog {soixedgqe & sot Ss 


a ' 
‘ : 


,noftuloe oito sqmyes ods te yiiwttev sits toed ‘or * 
etdt rot beyolaqmse ed As° . eu iubom-etstan pare tr le 
; P ; et (as “Ye a] 

A023 tbao8. nois2se% reivoisieg 


ct < _ 





c o% ot Se - “+ es 
4: ( OL x Be s8Ri9 35 vate 


[( ¢s = qt s g00 oe 
iF | :.32 * tyo,a)?° a eg ra . 











cel ap. a : Rank 
Bei _. ie hed. OS) cr OL x.d¢.8)/ _ , i 
*(znda.08 x $10.0 s Mae. ees 

: . Sf pee - 


see eeD x8 a gi 


ge ie Re). ae OL x te 


(tavisitso mp-292 


male | 

ES8€.c = <= {2 : Xeon, Tx ta, oa. 
OL x vee.2 x “988100.0 
ths. fT 4 

ie 









a flat slab model. The modeling equation can be obtained 


from equation (F.1) by letting a equal to zero. 








2 
3 Pas p 
ts - De TD ee T,) = Q (Pea) 
In a dimensionless form, 
ar 38 ae Epis 2 5 (F.3) 
az? £ Yr, ea Tet ; 
fe Pp 
where Y= ae ee sf (F.4) 
s s 
‘=, 
Eg = R (B55) 
Op 1g (P ee ) 
62 s R2 s s s (F.6) 
7 De C. 
Furthermore, 
ie Pp 5 
(Pe ) : 
ry Bey T,) 
(2956 sho ta), Bes ae 
EXP e ee 
ee 
{1 + ES Ata + 2P ei - oes gp 
(2-56. x 10uenpe 
s EXP (22350) 
2 Reel 
{l + Nea hea + 2Pe i ~ 2P.)} gs 
‘ eo heidi 2 (F.7) 
2 3 2 
(C, ~ Co Y) 


263 


Lebo sta 
ksted vd ut. 4) cotanuve went a 


vin rose 
So eG | Es 


2 iol of x aes 





oa im | 
a 







‘ aa ok i Ns 
98 me 


S >. 


where Cc) = 1 + 0.006 Pano, 2 + 2 Peyi? 
Co = 0.012 
C405 €,7P = 


(F.3) becomes, 


2 1.5 
os - 02 (c, - ¢,)? -———~ = 0 (F.8) 
dé (oe. C4!) 
3 2 
with boundary conditions 
Y= 0 at —& = 0 
eh te 
ae 0 at e = 0 (F.9) 
: ip pep A at £  -&).1 


F-3. Asymptotic Solution for the Effectiveness Factor: 


An asymptotic solution of equation (F.1) can be 


obtained from the solution of equation (F.8). To solve 
equation (F.8), let 
oot 

y de 

Then equation (F.8) becomes, 
i 
og i‘ of Ser ies c,)* —+—__, us 
(C, - Cy Y) 


After multiplying both sides of equation (F.10) by d¥ and 
rearranging, equation (F.10) can be integrated such as 


y (€=1) Y=] 1.5 
{ E 
y (€=0) 


_ 42 he 2 y 
ay= OF (C, Co) \on Reoeaat ie ay (FS LL) 


264 









r. ¥ Td , ~ 
‘ 7 : N i: ~ ’ \~ ‘ : 
7 a A 


ue , 



















‘te 054 eget bien ome eres. 
| * ie Fea 7 Ss 
onl SEE ON ‘st. — a 
F an 9 = hs 7 5 ae ; 
~ P ps Se Ee ” ae £ 
.2émooed. (€.3) voltae WA 
re ae , i ae a 
(3.7%) 9 =. anaes tis 3 =. Re a 
MEM he K Sy 


enot#ibaco casbaied 4 


(@,%) 


ean oe 
ed ma9 (L. 4) sidnied to noLsuloe. ohtoo geen 98 ‘ims 


evioe of . (8. m an se) noisulos aor | 


(OL .%).. 


bas ¥b yd tots sai 20. 


Pe ae 


aS q 88 cave ey ods 


The integration on the right hand side of equation (F.11) 


can be done analytically to obtain a value of 50.008. 


2 2 2 
Yiea1) = 2 2% (C3 - Cy)* (50.008) 
2 2 
= 2 2 (0.1094 - 0.012)* (50.008) 
= 0.94885 07 
£ 
or TireN) = 0.9741 o, 


The effectiveness factor can be calculated according 


PGeLte ceflini tion. 


le Actual diffusion rate into the pores 


Reaction rate within the catalyst pellet 


Here actual diffusion rate 


reaction rate = re 
Therefore 
ie wl. ¥ 
a e2e> (e=1) 
3 Pate ee a ee eG | 
ii to) iran eee | 


re 


dc ( 
e dr a ts. 
PoP 


) 





(D 


Cs. ay 63% 


D_ (=) (==) -_, ( 
6° R dé" €=1 p Pp 


5 (09741 0.) 
5 £ 
oF 


223 


eto) em 0.1858 


Then 


265 


ae, Ah at a 


a - . : ae 
cos . “ ih UJ ¥. ? eh ; as é 
; — aay ! on wae ‘s ae 

i Po is alia ads et MPL 

11.3%) aoitsype to ebia Ouse Paprt ort ao not a: 4 4 7 

et 


. aa ¢ 7 
ri" 800,02 20 opisv 6 nisido oF yilsoistytenas ated abs a iy 














, 
| 
> ‘ 


an 
<< 
> 


7 
' " a 4s 


. A Cc i 
(BOD ,02) [4.9 —-p9) 40 Soe at 


¢c a 
Bers ca 
en 
“ve : : ee SS 
00.02) “($10.0 — de0r.o7 Go f= : 
; ot * y 
C r , = a 
A ee egage.0 eric 
; : 7 
: “Sas as 
a, a 
) a f " 0 = , es “ 
hs aad rete. (La3)* 
- Cee Fe . 
: 85 od > yotost gaenevisgert 19 ost 
»' a e . J 2 : 


i ee i mapa! a3: 


’ e . 


ge10g off oct etst noieutiib LausoA ey 


Na 
emecenggretignemntacs ipa > een a 7 3 fas oh . 
a? ss BROS aaa 4 t ay 
JHLit= | 3Byis be eAQ aid iS iw SB3BE fic . F a eS me 
’ -_" ; 
; = es ‘\ 
: * 





_a) = ost mokeut3 tb . 
- . a 17 


APPENDIX G 


DERIVATION OF COLLOCATION EQUATION 


FOR INTERNAL RESISTANCE OF CATALYST PELLETS 


266 


. Ll} 


>" 
A 


pes ~ + 
vii } tTASOY 


(IATAD. FO SOMAT 


1109 40 ) wornaivzniia - 





rai 
ne : <P 
# 
r 
Cs ” ns 
“y' 
i mye j 
“ihe a 
= | ’ 
< : e 
] 
i j 7 
on 
« ‘ ~ 
~~ ’ + 
. ’ 
« 
4 os 
- 
a s 
J 
il 2, ra 
* n _ ’ 
i 
ak o —— 
. 


AS 5 rd 


267 


The general material balance equation for an iso- 


thermal spherical catalyst pellet becomes 


d*c Pian Pp 
gizo + r az = De ae) (Pe Test = 0 (Goad) 


By introducing dimensionless variables 


2 ae 
Y= ame (G.2) 
Ss Ss 
= = 
Ste (G.3) 
equation (G.1) becomes 
Sn aed Be es eet 
az? — dé Seah y rs 
where a represents Thiele modulus defined by 
Se R Pp s Ps! T,) 
Ss 3 eG 
e “s 
But as was shown in Appendix E, 
ihe pie) yi-5 5 
ee = A - ) (G.5) 
s‘'s s (C, - Cy ) 
where Cc, = L ts0.006 (PH0,4 + 2 Pe;) 
er ig UO Ob2 
C3 = Ci/P, 
Now equation (G.5) becomes 
2 de 
2 bay 2” ay 2 2 y 
apr eiae  S 9 @ (C, - C,)° ——— 7 = 0 
deg. oad > ee ET ee 


tot notaaupe sonsisd iskeesam {sense oft 










\ a 5 F 
*,0 
. neat &- 
on oa 
ye ah 
n 


i 
a 
4G 
Ht 
alo 
he 
o.- 
dae: 
a 
ty. 


ch 
eoldsizev exsinolan emi pit touboxsat, va 
| ‘2 ens 
2 g.. z= p- & =: y i 

el ot ‘ p a 2 s 


irs ee Re Hes ; 
e i > 

Ws 3 
eemoved ‘L 2) av JsupE 





/ . ge > 
re = \ 
oe S 6 ‘. 2 
Q-:< = ae > ~o- + —_— 
Te sure) “3 8 °.- SBS a 
“ > . = 


yd bentteb eu Lubom- eterat asaede 


268 


To check the validity of the flat slab model for a 
spherical catalyst in solving equation (G.6), an approxima- 
tion method can be applied for the steeply descending 
concentration profiles within the catalyst pellet, which was 
Originally proposed by Paterson and Cresswell (76) in terms 
of the effective reaction zone defined by the region beyond 
which the reactant concentration becomes essentially zero 
where no reaction can occur as a consequence. To apply this 
reaction zone concept the reaction interphase er is defined 
as 


oe 


x = Laaeet (G.7) 


In equation (G.7) x varies from zero to one for the 
value of € between er and one. Now equation (G.6) can be 
described in terms of the reaction interphase by using 
equation (G.7). 


ie 
1 d fn 


eke E,)? dx* oT 


= 


2 ( ui ay 
1 - é,)* 1 - er ax 


9 92 (c, - C4)? i, = 0 (G.8) 
dea ce 
Boundary conditions are 
y=] he J ee rad A (G.9) 
weetoy Gobo intaeelk =o (G.10) 


It should be noted here that the boundary condition 
(G.9) is the Dirichlet type rather than the Neumann type in 


this problem. That is because the approximate value of the 


Bas oe te yaa 












s 10% isbom dele soir ert Ro yibiiev, amd ‘ 
-smixorggs as (9.2): nobssupe enivior nee sextedao J AC 

Le 
on ismeoeal viqoose ods x03 bollags od ms bostson 


aeaw tiotdw ,telleq seyls2e> ons - abdd iw ‘est ktorg noktsxdas 

4, Pik ae a 
amxet ot (3%) Lloweser bas doaret64 yal bezogoxq ct | 
bnoyed rnolpss esis yd bealieb Senos no iso8ex evitoerie ‘oem 


~~ 


O2eS8 yilsttneste Bemozed nobséxsasoato,, tasdosex ont a 


ra 


~ vo 
eids yilaqs of . somenpeertoe 5 2S xtops ris noktosex on os, 

’ ya? 
fo IO8 7 
at4ae 











benitseb el te sesdgquedni noktosex oni 4q9pn03 ue 


F 
4 


sh ud at. OP a 
(Tm) eee | 3-4 a 
L ah - Pa 


———~ 


efit 102 sto oF Ofe8 mort eeitav, ee i notsaups ar ne 
od ms9 (8.0) motssups wort yen. bas ioe nes 


pateu yd seach noisoses Pe ae 


ae 
(8.2) OQ = 
(€.9) : ; ie, oa ere 
(OL-D) , ee Oi" ge. di: 


ngistbaoo: yusbaved. ‘edd sass vet be 


a Aetna ms a dex save 


| zr fevicond : coos 


269 


Sherwood number in this particular case is 


pare 
Se Le ee ee 
sh Ba Fed o0le 166 


So the ratio of the convective mass transfer effect 
in the external film to the diffusive mass transfer effect 
within the pore is around 166. Therefore the external 
resistance is 166 times less than the internal resistance. 
Here the assumption that the bulk gas concentration can be 
rigorously used as the catalyst surface concentration to 
calculate the effectiveness factor. Actually the bulk gas 
phase concentration is different from the catalyst surface 
concentration as was indicated in the study on the effect of 
the external resistances. However, this difference has 
negligible effects on the calculated value of the effective- 
ness factor on that assumption since the major resistance 


exists in the inner part of the catalyst. 


By employing the Paterson and Cresswell's collocation 
technique (76) improved by Van Den Bosch and Padmanabhan (90), 
the concentration profile in the reaction zone can be appro- 
ximated by a parabola. 

2 


Yo) ae (G.11) 
Ss 


As expected in the internal collocation technique, 
equation (G.11) already satisfies the boundary conditions 
(G.9) and (G.10). From the boundary condition (G.9), Le 


becomes one in equation (G.11) to make 


ee a ww € 
r y , 
7 ( “ Of 4 
+ c Ww ~. Vv L | m 7 
0 r = _ oGawwaree §84E de 
~ Q f 60 en «€ 



















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} 


a 


[ud edi yilsaaDA .2oO79872 PY 


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sf aonsistiib eftis.xesveworR . oot 1m niger, * feern 
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ide 
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+ wi 3 “ - | 
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| | hte 
ai OS 


: } ‘ou 

moitsoolion e'ifswaa x19 brs noexes84 odd peck jotqma a 
PAs wes tes 7 

(0@) oasddsnembsd ons. dowod aed ns: vd ovens | av) 


re 


-o1gas ed msD> Smos no btosst" ott al Late 


te 


ia ma ; 


eupliadoes notyss0les Asaxert 


Nal 


; pais tong yasbawod edt eoite abe 
tte 


- 
+ 











pA as 


aie ¥, 








: 





ays nots 


aa a 
ob! Sir tap 


re x (G.12) 


for any collocation point i. By substituting equation 


(G.12) into equation (G.8) the collocation equation can be 


obtained. 
a . 4 #51 
fe eee er! 

3 

2 ed 
SEOh oes (Ome Cs) = 0 (G.13) 

3 (oe ee ee 
3 Py eee 


The optimum collocation point can be taken as 


xX. = recommended by Van Den Bosch and Padmanabhan (90) 


1 
» 2 


for the high reactivity model. 


For the value of x; = et and calculated value of 
¥ 
eo er can be obtained from equation (G.13) to see the 


applicability of the high reactivity model using the false 
position iteration method. If the high reactivity model is 
really applicable, then the effictiveness factor can be 
calculated as 


& 
(ok d 
De (qe) (ae) 


E=1 6 1 dv 


7 =—_—__ = — (==) (G.14) 
dik<o 2 dé’ =] 
Deke ey T,) Deyo 3 o 
But 
ay net gay dx i 2 
Ea = Cpa ier ey ned = traes (G5) 
By combing equations (G.14) and (G.15) 
2 af 
n= —-——s (>=) (G.16) 
Bepoe otis Cet 


Ss 


270 


os 


sivigosex dpid edd 21 ,Box {yom au 


: “\) ee" 
<< Wen 
5 7 7 
: on DT 
1 oe ’ 
ae 










ri 4 : i ta 
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notssoolion sit (8.0) mos wala oss (91.9) 


' ° do 
* a <o05 £ 7 ape 
5) -' > : 5 
x - Zs 
> s 
“ —- — — rt ee + 7 
ij + aay 4 1 = 


Weert 
x ew 2 22) a : a ; 
aso tnteg nottspolioo mumiz 10 oat ee - 

~ = "eon ~ 
bas doeos aed nav. ais babfonmones = = x 
- a A fibge 


' , a 


.Lebom ys hivisoset cet edt IC 
ee, ae i A 
as Bite} 
soles bab =.= eRe sul sv arn; 308) 
: S f : ie Yo ; p * ame ‘ . 
(€£.5) mo ee Ye) mo xt beaissdo a ye Ne 













tite ae: 4 


fabom ytivisoset doit eft 2 to watt 
ee 





ss one ae 
yELlags 


aeonevissitis odd ovine a ae 


< 
Te 


Then the concentration profiles become using 


equations (G.11) and (G.6), 


Y= qaniess = (RE +.B) 7 (G.17) 
- & . 


A listing of the computer program COLLO, used to 


calculate the reaction interface ere is included. 


271 









; 
aol 
, 1 


¥ | Et 
onieuw emoged aelitorq Pca eat: on? enwhh 


i 

(8.0) bias (ff. 2) li 

| : *e ¢ yah aan 

(Tia) ee Se ee “ =T * Y <s 


= sj : ites 
= .* 


ot beev , O10 mexpoxg xotuqmod sniy to pa repita A 


ext 


VO AO Ml OQAGMH MO 


10 


12 


90 
60 


1090 
20 


Tas 
88 


99 


BE AE AE EE IK IE AE OK OE SEI EC BIE I AK OK IE IE OK OIC OIC IC IE IK SK OK OK OE OIE OIE FS FE AE OK OIE DIS OIE SICK OK OIE AK OK OK OK OK OKO 2K 
* : 

* MAINLINE COLLO 

* 

* THIS PROGRAM COMPUTES THE EFFECTIVENESS FACTOR OF 
* A ALON CATALYST FOR THE CLAUS REACTION UNDER CON- 
* DITIONS WHERE THE REACTION INTERFACE IS GREATER 

* THAN ZERO USING A INTERNAL COLLOCATION METHOD. 

x 

** PT’ = TOTAL PRESSURE 

* XHO = MOLE FRACTION OF H20 

* XSO = MOLE FRACTION OF SO2 

x 

HE AE DE AK AE AS IK IEA OK IE OK OI RCI ICI III IG IOI I aI aK ak af ak ak ak aka ak ak akc afc ak akc aka ask 


READ( 5,10) PT 
FORMAT(1F10.5) 
READ(5,12) XHO,XSO 
FORMAT(2F10.5) 
PHO=PT*XHO 
PFO=PT*XSO 


ASSUME NEGLIGIBLE CONCENTRATION DIFFERENCE IN 
EXTERNAL FILM 


PC=PFO 

C1=1.+0.006* (PHO+2.*PFO) 

C2=0.012 

C3=C1/PC : 

DO 100 J=1,16 

H=4.0 +J 

CALL FALS2 (C2,C39H,Z1) 

RECZRM T7%sf1,90 

WRITE( 6,60) 
FORMAT(1H1,///19Xy "THIELE 97Xy "REACTION ', 2X, © 
1* EFFECTIVENESS! 4/18X_y "MODULUS ' 56Xy "INTERFACE! ,9X,y 


2"*FACTOR® /) 


EFACT=2./H¥®*2./(1--Z1)/3.~ 
WRITE(6,20) HyZIy,EFACT 
FORMAT(//10X,3F15.4) 
SULTUS9 

WRITE( 6,88) 


FORMAT(1LH15///10Xy "THE REACTION INTERFACE IS NOT', 
1" GREATER THAN ZERO.! /) 


CALL EXIT 
END 


% 
x 
HHHH sz H HH HH HHH HK 


ott 


+ 
3 


272 

















: a 7 


a) 


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-OOHT 3M Nd ern JAMASTAT A OMT2U OAS 


Het & & Oe 
ie 


x 3 ae i‘ F ; * fe ie 2 
* , 3022404 JATOT = 19 bie 3 
e ASH AO MOTTSASA FjnM = “tee “a 
% go2 20 WOT TIARA 3.0 = e os au. 

a 

Me Mek ee is eek oe <AEASEORERAG SERS SERERAAESAESES OE ATEEES ORR of ‘Bs 


Te (0Ly2 10038 
x Dstt eee ‘7 


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| 3 . (260175 TAM < 
‘paw oer aone 
| aang tie 4 DeKsTaeoa8. 


i=; 
wi aqmanaaaia WoITAATHS3H02 398191 Ja3M sauce a 
ag: Re . FY: 


a Ere ee 
ie 


(TSaHe£9eS9)- wtp i ia a 
ar orate te 1) y —. 


e%S s'MOTTOARRE aKVe ‘353 tKTs aha 
XO ‘I DAIRSTUL? tXOe ‘ener. eXettem 


oh a es Mt 2 


ean ar 1 * 
rads 4” iy nea ds aT rte ] . 
‘ ayer et. ‘ ti es aes Pe ey Ee 
Loan ae v cin s i Le ae go ue wt * Ps 
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ie prad Bary phe yer et 





ie IM my 181 ey ae Fe ae, om ae VE 


AEE AC IC AE CC IC a I IE 2 ae ak Re RC 2 A 2 IC aC OC OK IE SC OK KC A OK IC AEE OIC IE A OC A IC AKC IE KK OK 2K 2K OK aK Ik OK ik 3K 


SUBROUTINE FALS2 


FROM THE COLLOCATION EQUATION USING THE FALSE- 
POSITION ITERATION METHOD. 


et HH He H He 


x 
* 
* 
* THIS SUBROUTINE COMPUTES THE REACTION INTERFACE 
* 
¥ 
x 
* 


BEE I EI CE AE ISIE IK CIE IK OK IK RE AE AE OE IE OK BIC A IC IS OE IC IEE AS EK OE OK OE OK IK FE OK OK OK OKC OK KK AE OK OK OK OK 2K 


SUBROUTINE FALS2 (C2,C39HyXMM) 
DIMENSION XM(500) 

F(X)=2e/( 1Le—-X) *¥¥ 2. +40/(10-X )¥XI/(X+(12.-X) *XI) — (H¥* (C3 
1-C2)*X1**1.5/ (C3-C2*X1**2.))**2, *9, 
X1=0.707107 

XL=0.0 

XR=1.0 

XM ( 1)=0. 

DO 99 T=1,499 

AMC I+]) =(XL*¥F(XR)=-XR¥FCUXL)) 7 CRIXR)-F(XL)) 
IF(FC(XM( I+1)) *F(XL)) 22933544 
XR=XM(I+1) 

GO TO 70 

XR=XM(I+1) 

XL=XM(I+1) 

ESP=ABS((XM(I+1)—-XM(1)) /XM(I+4+1)) 
IF(ESP-0.001) 100,100,99 

CONTINUE 

LEO1 499)" 666,555 59 

WRITE( 6,77) 

FORMAT(/10X,'NO CONVERGE'/) 
XMM=XM(I+1) 

RETURN 

END 


va 

























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APPENDIX H 


NUMERICAL SOLUTION FOR EFFECTIVENESS FACTOR 


BY WEISZ AND HICKS' METHOD 


274 


S is 
4 ié war | 
= _ 
i ’ 
A I 
’ = 









t ql : 
} » 
‘ 
Le ae - 
, ? rs 
J 
a: el 
» l~ 
. y 4 c 
i 
' J P , ey . ' 
\ - - 
‘ _ ~oe 
» is « 
yy A ai 
; c “— f ~ - 
- - 2 y { 1 
i J '@ 4 “ 
i “T¢ 
Z ¢ ” W, I 
- _ 2 2 : u fr 
4 XI CUAaAITA ie ea 
= : nat 
ft Pe 
’ 
~ : = z 
i © 
. ; : 
“7 “ 
"> 
i > wi & Lae, y 
- J f ' 
) : et ae : 8 r i hl ry a ’ 
tens , a Teeny ; 1 ATF oe 9 aa 
AOTIAY LAANAIVITOATIIA APY WOLTUIO® TADTAAMUMT A 
of. se an 4 y te . re 
i - ’ j ne i . oy Teg: = ¢' Fi ; 
: meres i. ’ T LZ a a é « a 7 ; 
i : COHTSM ‘eASisn GWA searaw Ya A Ton oft 
; ~ a 8 or , ha t% 


275 


H-l. Numerical Solution for the Effectiveness Factor: 


Use can be made of the equation (G.6) 


2 5 
os i: = oF 29 9? (c4-C,)* —_—, = 0 (G.6) 
dé (C4-C5Y) 


with the boundary conditions of 


y= 1 at = 1 
al at —E = 0 


Equation (G.6) can be transformed by introducing a new 


variable © 
A 
ao= 5 (H.1) 
Then equation (G.6) becomes 
2 fest 
f+ lS - 9 a? 0% (c,-c,)* —~—, = 0 (H.2) 
dx (C,-C,¥) 


Equation (H.2) can be numerically solved by the open-end 
method with a known initial value instead of treating it as 
a two-point boundary value problem, which was originally 
proposed by Weisz and Hicks (97). 

The computing procedures are 

(1) Choose an arbitrary value of a ae 

(2) Assume an initial value of Y=0 at x=0. 

(3) Integrate the equation up to Y=l1, and get the 

value of x at that point. 
(4) Solve for a using a = & = ~ at Y=1 from the 


boundary conditions. 

















st eagnev! Ibs 2 ie 
(2.0) soltespe sit 
, i 
ae 
r & S { *) § a) ¥B < — b 
ete aaa eheneteenael a 7 h ’ ‘ 3 . 
: ur s S$ € pall Pa) r © 3 : “3b = 
i‘ w~* c i ; | 
S~ Ri ee Pak 
& é 4 ty 5 4 
to enolstibsaod yrshaL oe ae a 
i 7 ~~ ; 7 4 | 
| [ = 3.35 oe 4 


4 
S 
'?) 
P| 
+ 
| 
a) 
ss 
AU] 
2 
By 
re 
at 
o 
po 
Br. 
© 
~ 
fv] 
@ 
— 
ro 
Cc 
be S 
re 
ie 





e 4 
i . 
‘ a. fare 
~ Por 
: [ eH) : “2 
eemosed (2. ay ever 
ny e 
aie cg) vag 
% < ‘ : a) 
sc f ps ss eishieahs ar ey - 7 fe ry ; ®&.; v a Vee = 
(hk. 0 r - Fi Pe) > 5 
Diet Saccialees'. Malini, ae , mb ® 
s d ; om h- a) 4 fr 
re * | .) 
bas-asgo sit yd bevloe wits so btemusin ed. nee os oe 


ye 
es ti pttittserst to bases ani eulev teistak wond & Ati 


yvitsate tic esw do idw me Ldoxa, ouley “yrahauoe wi 
5 n Wwe 

(8 e) 229 it H be siseines 

»~ 


218 es rubeo03g sara mI. 0 | 2 et : 
Eee ‘eh 









a . om santo aaee os 


geen a. 
— 5. 8 


hgh we eae eae 


ao i= il : 


Piast “ ots fmond & 


= 






‘, ’ 


- Pre - ie a ’ pac j 
: ae Pet fee nar * 


276 


(5) Solve for ®. using Conditven (1). 
(6) Solve for effectiveness factor by the relation 
PS eat (at) 
‘ a ou 1 


3 9 x=— 
s a 


The condition (6) can be derived from the definition 


of the effectiveness factor; 


ae day 
— Deke Dede eel 6 oe De. ' gray, 
BAP Via) qPp EAP Sy TS "R2P, dé° €=1 
since 
byte: 
es x 1 ay my cay 
* and (Fe) e=1 iy (adx) x=2 


———_- 
r{P. eT) .R Pp (30.)° 


Equation (H.2) can be transformed into two first 


order differential equations by letting 


o = y (H. 3) 
Then 
LV5 
oe eee 252 eee 
pt 9 ca De (CoC) oe tS (H. 4) 
(C,-C,¥) 


Equation (H.3) and (H.4) can be solved simultaneously 
using the Runge-Kutta-Gill integration procedure. To deter- 


mine the value of the indeterminate form of ey on the left 
















- ‘ f net, — 
| 


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(<=) re > eee 

aay ® £ ' > s 

eS . z 5 a } 

“t ' y i J hae 

moit HevixedS sd aso’ {9) noi? ibaoo ont ne 
: 5 44. 

,190j0681 Be gnsvijostie | 






} rs 
. 
) g 
3 2 
} - we ey SE 
> ate - 
y \ ao +h : 
: 2 q 
pe, 

- 2 ‘ 

(=< = 4 her pms 

4 4 =3 .* 

} f 
>. 
Tr 
—ay 
1 
+ Fs e° 
727 Owy OFFAL Demtotersis 
=< Orit 
c y\ ‘ 
(ft o ct) 
4 
i 


| 2 vdy y ‘ Fi 5- fe 
(Db. H) Bie a Bat +e 
f Phi sex: 5). ” 


rast {.. te 
; 


“yleupenss tum te eA <ta 
rete At be : snk | 





thO<g 
hoes -s 
a » 
ee hie -” 
fr oe tet 





pS 


277 


hand side in equation (H.4) use can be made of the 


L'hospital's theorem at x=0. 


lim 2y _ ady 
x>0 ee: (Be) 


Now equation (H.4) becomes, at x=0 


Vas 
Ye 3/h66") (co, -Gey*) eee (H.6) 
ax s EM 2 


For other points than x=0, equation (H.4) still 
applies. 

The computer program CHOWH was used to calculate the 
effectiveness factor by applying Weisz and Hicks' method to 


the Claus reaction system and also listed. 


H-2. Difference in Thiele Modulus Between At the Inlet and 


At the Outlet Condition: 


The ratio of the Thiele modulus at the reactor bed 


outlet condition to that at inlet condition can be written 


as 
® : ed Pd a a ee A 
Fits. re) ores si (H.7) 
o. rs (P ae. Cs 


by using the definition of the Thiele modulus, equation 
(3.14), and assuming constant pellet density and effective 
diffusivity along the bed. In equation (H.7) subscripts 

o and i represent the outlet and the inlet condition 


respectively. 


to ebsm Sd nso ean (but) Aobsneys ake ab! 


A LT TOA 





















? pe 


v5 s' . 
\ hee: = x Ȣ 
xD x. , J oe 7 - 
ot 4 ae es 
han er rea eet f “A cri 
=% 36 «4Boemoosed {e.h) 10.3 eeHEP WR 7. 
‘ , \y t : * & 
J + jw red ° © of. vb a 
gmemenmnacten, fen e ‘e. = ; 
: Tok ie Y ad ait x - 
> su — om _ ny Me best - 7 
‘+ t c 2 TD r Ae ds 4 7 
a St wy ee r P 
a | ; ee: : "ee ’ 
, ~ “ 7 wore 
iteopo ,0@=x asAdy esatog FT yahto ‘a0 
« : _ 
: d a" - wi 
€ 


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“ . 3 a 


ie 


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: nots k prio) t0L2u0 ods re) . 
acti bom ole. bag. ends Yo. Ofte: ot Vita 
deh Oil eae 


olzibnaoo teint 48 ys (© wokath o> Je! 





aa 


Paleo 
ee tadaity 
7 ty 


> & 


~~ 


278 


If the denominator change in the reaction rate 
expression may be ignored equation (H.7) can be described 
in terms of concentration and temperature at the inlet and 


outlet of the catalyst bed. 








me NF | 
se = aed (H. 8) 
i so 
Assuming 98 percent conversion and Le = 550°K, 
1 = 680°K, equation (H.8) gives 
Se AS OS tee os i oar a Rae a 
5 50 550 a987e 42550 680 
= 0.839 


Therefore, from the calculated value of o. in Appendix F 


qa ty (8 ae o. -e0. casos lo. oons) = 4.4 





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re 


bas telni eft 3s stutsteqmes bas ‘no isextneoaoo io an ‘ 


















, in * a 4 vo AS hh 2 
a bad te b aneaatas et to telsuo — 
SS aay ia 
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L! a 2 O28 , ,oe2 | ; “AN 


asae ™ 
“ie Pe 7 . 


ne oi vie 


Miss (aia ee rae oe 5 
gevip (8. HH) ‘ten 7082 = 3 


< 
~ 
=) 
| 
be 
re | 
Oo 
~ 
a 
D 
cee 
iw 
© 
ie 
oe 
ip 
QO 
4 - 
‘p 
1 - 
oo 
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re 
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| we 








oh 


I xibnasqgA at ,% 20 srlsyv boss {vole os3. moxd 4 





Cr Coreen Ca €aeGy Crt) CIC) Ca Gr G) Co Ch Ct? 


10 


bs 


18 


30 


BE AK FC IC KC RC IE OK FE I CIE KAR I ISIC OK OK KC OIE OE I IE AK OE IS OIC IC IS OK KK OK aK aK OK 2K OK ak aK Og aK 2K oie 2K 2K ac 


* 
s MAINLINE CHOWH 

* 

* THIS°PROGRAM COMPUTES THE EFFECTIVENESS FACTOR OF 
* A ALON CATALYST IN A CLAUS REACTOR USING THE WEISZ 
* AND HICKS METHOD 

* 

* RF = TOTAL PRESSURE 

*K XHO = MOLLE FRACTION OF H20 

* XSO = MOLE FRACTION OF S02 

** PRMT(1) = STARTING POINT OF INTEGRATION 

* PRMT(2) = END POING OF INTEGRATION 

x PRMT(3) = INCREMENT OF INTEGRATION 

* PRMT(4) = ERROR BOUND OF INTEGRATION 

* NDIM = NUMBER OF DIFFERENTIAL EQUATIONS 

* 


AE AE AE AE TK A OE BE EC ES AE OE AE AE AK OK AE AE IE OK KOE IE IC AE IC OE IE IS OK SK IS IE OK NS OK OE IE OE OIE OIC KK 2K OIE IE AIK 


EXTERNAL FCT,QUTP 

DIMENSTON Y¥(2)y,DERY(2) »PRMT(5),AUX (8,2) 
COMMON C2,C3,AH 

READ(5,10) PT 

FORMAT(1F10.5) 

READ(5,12) XHO,XSO 

FORMAT(2F10.5) 

PHO=PT*XHO 

PFO=PT*XSO 


ASSUME NEGLIGIBLE CONCENTRATION DIFFERENCE IN 
EXTERNAL FILM 


PC=PFO 
C1=1.+0.006% (PHO+2.*PFO) 

C2=0.012 

C3=C1/PC 

WRITE( 6,18) 
FORMAT(1H1_9///24X,"THIELE',6X,"EFFECTIVENESS!/ 24X, 
1* MODULUS! ,8X,"' FACTOR! // ) 

DO 100 KK=1,41 

AH=KK 

PRMT(1)=0. 

PRMT(2)=100. 

PRMT(3)=0.01 

PRMT(4)=0.00001 

NDIM=2 

IF(KK=20) 30,40,40 

Y(1)=1.0-0.05*KK 

GO TO 90 


279 


evs 


>, >» 7 é > . a ae 7 - s ; Ve ; 
a be 4 = = a a MA Tas 7 , 4 
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280 


eee (CONT'D) 


40 IF(KK-24) 50,60,60 
50 Y(1)= 0.05-0.01*(KK-19) 
GO TO 90 
60 IF(KK-33) 70,80,80 
70 Y(1)=0.01-0.001*(KK-23) 
GO TO 90 
89 Y(1)=0.001-0.0001*(KK-32) 
90 Y(2)=0. 
DERY(1)=0.5 
DERY(2)=0.5 
CALL RUNGE(PRMT,Y,DERY yNDIM,IHLFyFCT,sOUTP,AUX) 
100 CONTINUE 
CALL EXIT 
END 















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BEAK IC AE COE KC EI CE IC OE KK OK OK OK IC KO OE KE SE OK OK BIE IE AI OE OK IE OK IK IK IE OK OE OK OIC OK OK OK OK OK OOK OK 


SUBROUTINE FCT 


H+ tH 


THIS SUBROUTINE GIVES THE ORDINARY DIFFERENTIAL 
EQUATIONS TO BE SOLVED TO THE SUBROUTINE RUNGE. 


* 


* 
BE BEE AE EK EK AE IK EAS IKK OK IK AE OE EK OC OIC KS DIS AK AE IC IE OK BE OK OIC AE AC OE OK OIE OIC IS IK OK IE OE HK OIC OK AEDK OK 


% HH Ht 3 3 OK 


+ 
3 


SUBROUTINE FCT (X+sY,DERY) 

DIMENSION Y(2),DERY(2) 

COMMON C2,C3,AH 

DERY(1)=Y(2) 

IF(X) 20,720,30 
30 DERY FA} S=2 ee V1 20 7At CAR (C3 -C2)/(0C3=C2*Y01))) )**20%9, * 

LYUD)F*1E5 

GO TO 40 
20 SDERWR eStart (063-C2)) 71 C3-C2*YC1)) )**2. *Y¥ (1) **1.5*3. 
40 RETURN 

END 


281 





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x 

* SUBROUTINE RUNGE 

a 

* THIS SUBROUTINE SOLVES THE SYSTEMS OF ORDINARY 


KUTTA-GILL METH WITH THE TEST OF ACCURACY 


t 3% 


HE AEE HE IE IEE OE A IE SEA IE IS IC IE IE IE IE AE OIE OK IE OS IE IE IC IE OK IE IS OC OC IS 2K IC OIC IC AC IE IE IE BIE OK BIE AS OK OK IRA OIE 2K OS 


SUBROUTINE RUNGE (PRMT,Y,DERYyNDIM,THLFyFCT,OUTP,AUX) 


DIMENSTON Y(2) sDERY (2), AUX(892),A(4),B(4) »C( 4), PRMT(5) 
DO 1 IT=1,NDIM 
1 AUX(8,1)=0.06666667*DERY (1) 
X=PRMT( 1) 
XEND=PRMT( 2) 
H=PRMT( 3) 
PRMT(5)=0. 
CALL FCT(X,Y,DERY) 


ERROR TEST 
IF(H*(XEND-X)) 38937 2 
PREPARATIONS FOR RUNGE-KUTTA METHOD 


2 A(1)=0.5 
A(2)=0. 2928932 
A(3)=1.707107 
A(4)=0.1666667 
B(1)=2. 
B(2)=1. 
B(3)=1. 
B(4)=2. 
C(1)=0.5 
C(2)=0. 2928932 
C(3)=1.707107 
C(4)=0.5 


PREPARATION OF FIRST RUNGE-KUTTA STEP 


DO 3 T=1,NDIM 
AUX(1,1)=Y( 1) 
AUX(2,1)=DERY(1) 
AUX(3,1)=0. 

3 AUX(6,1)=0. 


282 


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13 
14 


15 


eee (CONT'D) 


IREC=0 
H=H+H 
IHLF=-1 
ISTEP=0 
IEND=0 


START OF A RUNGE-KUTTA STEP 


ITF((X+H-XEND)*H) 79695 
H=XEND-X 
ITEND=1 


RECORDING OF INITIAL VALUES OF THIS STEP 


CALL OUTP(XyY,DERY,IRECyNDIM,PRMT) 
IF(PRMT(5)) 40,78,40 

ITEST=0 

ESTEP=ISTERS) 


START OF INNERMOST RUNGE-KUTTA LOOP 


J=1 

AJ=A( J) 

BJ=B( J) 

CJ=C( J) 

DO 11 IT=1,NDIM 
R1=H*DERY(1) 
R2=AJ*(R1-BJ*AUX (6,1)) 
Y(T)=Y(1)+R2 
R2=R2+R2+R2 
AUX(6,1)=AUX(6,1)+R2-CJU*R1 
DE J=4-0 = iba 15.15 


J=J+1 

IF(J-3) 13914,13 

X=X+0. 5H 

CALL FCT(XysY,DERY) 

Gi Tal 10 
END OF INNERMOST RUNGE-KUTTA LOOP 
TEST OF ACCURACY 

IF(ITEST) 16,16,20 


BN GASE ETEST=0 THERE IS NO POSSIBILITY FOR TESTING 
OF ACCURACY 


283 




















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24 


25 
26 
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28 


29 


eee (CONT'D) 


DO 17 T=1l,NDIM 
AUX (4,1) =Y(1) 
ITEST=1 
ISTEP=ISTEP+ISTEP-2 
IHLF=THLF+ 1 

X=X—-H 

H=0 .5%*H 

DO 19 T=1,NDIM 
Y(1)=AUX(1,1) 
DERY(1T)=AUX(2,1) 
AUX(6,1)=AUX(3,1) 
GOKTO 9 


INTGASE/PTEST=1-TESTING OF ACCURACY IS POSSIBLE 


IMON=ISTEP/2 
TEUISTER= Ph 0D-1M0D)- «21523,21 
CALL FCT(XyY,DERY) 

DO 22 IT=1,NDIM 

AUX(5,1)=Y(1) 
AUX(7,1)=DERY (1) 

CT 8 IM SE Ia 


GOMPUTATION OF TEST VALUE DELT 


DELT=0. 

DO 24 T=1,NDIM 
DELT=DELT+AUX(8,I1)*ABS (AUX (4,1)-Y(1)) 
IF(DELT-PRMT(4)) 28,28,25 


FRROR IS TOO GREAT 


IF( IHLF-10) 26436936 
DO 27 T=1,NDIM 
AUX(4,1)=AUX(5,1) 
ISTEP=ISTEP+ ISTEP—4 
X=X-H 

I END=0 

GO TO 18 


RESULT VALUES ARE GNOD 


CALL FCT(XyY,DERY) 

DO 29 T=1l,yNDIM 

AUX(1,1)=Y(1) 

AUX(2,1)=DERY(1) 

AUX(3,1)=AUX(6,1) 

Y(I)=AUX(5,1) 

DERY(1)=AUX(7,1) 

CALL OUTP(XsY,DERY,IHLFyNDIM,PRMT) 


284 




























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vaasa oe 13 


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29 


eee (CONT'D) 


IF(PRMT(5)) 40,530,440 
NO 31 T=1,NDIM 
Y(IT)=AUX(1,1) 
DERY( 1) =AUX( 2,1) 
IREC=IHLF 

IF(IEND) 32,32,39 


INCREMENT GETS DOUBLED 


IHL F=IHLF-1 

ISTEP=ISTEP/2 

H=H+H 

LFCIHEF?Y 4.33533 
IMOND=ISTEP/2 

IF( ISTEP-—IMOD-IMOD) 493454 
IF(DELT-0.02*PRMT (4) ) 3593554 
IHLF=IHLF-1 

ISTEP=ISTEP/2 

H=H+H 

GO TO 4 


RETURNS TO CALLING PROGRAM 


THLF=11 

CALL FCT(XyY,DERY) 

GOoTO 39 

THLF=12 © 

GE010 39 

IHLF=13 

CALL OUTP(X,Y »,DERY,ITHLFyNDIM,PRMT ) 
RETURN 

END 


285 


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BEA A EE ICI IE EK aK IIS KC EE KC aK aKa KC aK KC 2 aR aK IS AS aK IC REC ICC KK RC I KC aK aK AR KK SOK 2 OK KC 2 2K 
SUBROUTINE OUTP 


Ke 

x 

* THIS SUBROUTINE SPECIFIES THE OUTPUT OF THE 
* SUBROUTINE RUNGE 
Ne 
ok 
x 
sk 
* 


ErAG hh 
H 


EFFECTIVENESS FACTOR 
THIELE MODULUS 


ete weHH HHH H HK 


BE IK EK HERE SIC I AE I EK IE IE A IKE IE IE OIC AKC OE OC OK OK OK I IE OI IE OE OKC IC IE OIE OK FE DIK OK OK OK OK A OK AS SIR IK OK 2K OK OE AK 2 


SUBROUTINE OUTP (X,Y ,DERYyIHLF»NDIM,PRMT ) 
DIMENSION Y(2),DERY(2),PRMT(5) 
COMMON C2,C3,AH 

WRITE(6,40) Xy¥(1)9¥(2) THLE 
FORMAT ( LOX» SELO.S sp ke CF 
IF(1.-Y(1)) 200,100,50 

itt beet A er eee OYE ey YA 1) -Y¥ 1D) 
X= XX#(X=—-XX)¥*(1.0-Y1) / (Y¥(1)-Y1) 
H= AH* X 

EFACT=X¥*Y(2)/3./H**2. 

WRITE (6980) HyEFACT 

FORMAT ( 4/20X2F10.5/ ) 
PRMT(5)=1. 

XX=0. 

Y1=0. 

Y2=0. 

GO TO 60 

XX=X 

Y1=Y(1) 

Y2=Y(2) 

RETURN 

END 


286 





















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Nc: 


APPENDIX I 


SAMPLE CALCULATION OF DATA REDUCTION 


287 


A294 


: 


\ 


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bd 
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7 : 
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288 


I-l. Input Data: 


Imput data were entered on request after keyboard 
queuing the material balance program, MTBAL, on the remote 
teletype. A typical input request sequence is listed in 
Table I.1. All entered input data were preceded by a ">" 
and each input request precedes the input data. A typical 


output of the processed data is also listed in Table 1.2. 


I-2. Reduction of Input Data: 


I-2.1 Temperature: (70) 
areTreactorsinter.=—..40878 + 18.31152 x (M.V.) 


B) Treactor*outlet = 2351827 f 18.21650 *x"(M-vV*) 


I-2.2 Pressure (mm Hg) 


a) P coed Cabs*) "= 744761367°+ 12.71505 x (PCT) 
b) PesteCr ae roduges == loon 379 =F"4705037 x (PCT) 
for runsF and G. 
Ben ate (gauge) = 27.66422 + 4.19851 x (PCT) 
fOrtvruns ua; 0;, & anda™L 
Cee. (gauge) = 15.83784 + 4.32817 x (PCT) 


Tor stuns a, bs Cy, U0, b.and J. 


I-2.3 Feed Flow Rate (SCFH) 


a) een 0.08793 + 0.94999 YPCT - 0.00099 x (PCT) 
2 
b) composition correction for the feed mixture 
4 
xX FDCOM. x MW. 
z=1 * - 
Pmix ~ 4 
py 


FDCOM, x Ves 


i=l 


T 
- 

Vv 
Pog 
ae 

{Pc |} 
(TO) 


-(T99). x 


Ss tos x 


+ tO 
iil et 
bebsveig 
4 
3 
aT as: f 4 
x &cLIE.S 
oda! ; 
\ 
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\ » 

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[és @ i or + 
5 ww . T i 

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2 
= 2 H.S 
= 3 cos 
= 4 SO. 
Px. = 28/V; 
FF = FF ( ) 
MIX No Purx 


4 Raa ae 


a) 


b) 


c) 


2) 


ATN 1 Leip ol. + 


ATN 2 Letps be 7 


II 


corrected area of 


i) For runs A, B, 
corrected area 
For funs oF 7-G, 


corrected area 


calculate average 


289 





MW; Ay, & (ml1/gmole) 
28 22402.10 
34 221/76. 10 
60 224.7% 35 
64 2190163 


Gas Chromatograph Data 


0.17031 x (ATTEN #1 Setting) 
1.76047 x (ATTEN #2 Setting) 
No + peak 

Creb/sBpaand J 

= measured area 

I, K¥and 5B 

= measured area x (ATN 1) x 


(ATN 2) 


area for each component 


d) calculate area ratios, *y0/4N,' Acos/*n,, and 


(1) 


Braker, W., 


Ag ard. 
SO, *N, 


and Mossman, 


A.L., "Matheson Gas Data Book", 


5th ed., Matheson Gas Products, East Rutherford, N.J., 


NG at 


ay ht) wr 2 4) 400 Zia 
Vv P 
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5 - te 
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OJ ~— ’ 
0a 200. i! een 
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te + v2 5 ; 
raore a) oz 4 b - i ‘ 
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- 4 » 
V\8S = a2 
< aie 
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us Pere 1 Hasxpo ames nad 260 B.S 


‘= 


(pntstee [4 WETTA) x LEOVE 0 + 19vT0. f =f wEs (s ba ey 
(onistea S# WATTA). x ehoat.t/4 ‘greti.t = § WA 


ray - : 


g43 x 
; P an a) 7 
L Oas 2 a 49 »&, A. aos 10% (. ~ - 
‘ 5 ce « 
a a eee 0) 
59is6 beivesem = 69is bodes: ! ze Ms 
eee ts ty<> = 
I bas % ,I ,H «Od a nee x0F Gt 
, Y i a ent 
‘ \ ‘A 
(f UTA) = sexs Hbexzvesom.= & 5918 haved 


ae ee 
a es va of. 


‘e 





oe by 7 


eae 


‘alle 


290 


e) calculate molar ratios from GC calibration 
equations: 


LirOor Lee fh, fo; Gp.) and J 


Mi o/My = 0.01259 + 0.94687 =x (A, o/ Ay ) x 100 
2 P 2 2 

“cos/Mn,, = O,Gsc09 + OL, 72771 x cos! Ane! x 100 

oa jai = 0.11839 + 0.84600 x es0n/ Nat x 100 


J erOrurounessr, Gye, LL, Ko and G& 


My o/ My = 0.00733. 740689481 x (A, o/ Ay ) x 100 
2 2 2 2 

“cos/Mn, = 0.00149 + 0.65325 x becOssaNa! x 100 

BS0,/' No = 0.02466 + 0.77038 x ROne aN x 100 


£) calculate mole fractions 


FDCOM 


1 OF PRCOM, = 100/(100 + Mus! MN oP 


1 2 


My) 


“cos. 5 ceo, , 


FDCOM. = FDCOM, x alan eons 


FDCOM 


3 = FDCOM, x Moog /My,)/100 


FDCOM 


4 FDCOM, x OI aed 


I-2.5 Water Feeder 


a) ml/HR = 0.31009 + 0.52032 x (dial setting) 


I 


b) gm/HR (m1/HR)/(1.00244 - 0.00022 x Ty oO) 
2 









7 Lia ies dl A on Fr an 
Pa . ‘ a ‘ + 
oes q ‘ ae : _ 7 > ie 
: ; a, rr ee 


a Le 


noiserdifgo DD mou noite? xsiom ssnteten 8 


oe  tenoisaupe eee 














tL bas 3 .o 2 (a. em: 108 ( ere 


OOL x.¢ .A\. oA)-x. Teabe.o + esto. o = git\eca ‘ae eae 
/f o “7 , mal Sloe 


-— 


ion + 208£0.0-— Naa i Chee 


OOL x C LAN aon A) * 
‘ 


OOL x ( \ oe x o0ang..0 + erat, 0 = a as. * ‘+ B 
$ Re 1 i PF anus 108! as ea 
ooL x ( ‘aon? x £8)08,0 4 €€T00.9 =! zene ‘ ¥ 
O0I x Cu \200 x @$623,0°+ 2100. 0 mel | Pret pi Be 
OOL.x ( AN oe) * scort.0 + 2080.0 & * et 
7 “> 


enotdosxt atom whens a aie 


ean 
0 0 = “7 4 
+ sf 2. es £) KOOL Mooaa esas 
bie diary “be anaes DBE 
on \con™ of a00! ca | & y 


raat car a” ROD . 
oor\( choc) x = sito = as 
oor a oe et: | 


é _ a es 
~ * , P me : 
: : ‘ a ‘s 


(enidsen tsb) x senso He 



















* Pace ny 7 OG a | 
: x $5000 on DPN VE 
nal se 6 Pa: aie ihe e 1 eS pce a ator a 
ro hit — nis - C mA o} 4 at oo ‘ 
te ee ee 7 


: —_ a: 2a ee ons 
} m a ns . 7 = 3 " . 
ay ae 


I-2.6 


Compounent Flow Rates 


a) 


b) 


c) 


average molar volume (£t? /mole) 


AVGMV = 
i. 


Il Mos 


(V. x FDCOM, )/28317.016 
ii i 


where 28317.016 is a conversion factor from 
ft? torml 2 


component flow rate (gmole/HR) 


in’ = FDCOM, x (FF,,,/AVGMV) 
FFy 8 = FDCOM, x (FF,.y/AVGMV) 
FFugg = FDCOM, x (FFy7y/AVGMV) 
FF 50, = FDCOM, x (FF,.,/AVGMV) 
FFs 0 = (gm H,0/HR)/18.0588 


component flow rate (SCFH) 


ny 
Il 


H.s FDCOM., x PPIX 


Hy 
II 


cos FDCOM, x PP MIX 


ty 
II 


PDCOM, xX “FE 


4 MIX 


duct Flow Rates (gmole/HR) 


=P, x (PRCOM,/PRCOM, ) 


= Pp x (PRCOM/PRCOM, ) 


II 
tg 


x (PRCOM ,/PRCOM, ) 


291 










* oe weet 
ry Ba nhs > » a 
ae 8 ay eAt * 
ate i at 7) iy 
; ADist 4 7 . bad ce 
asseA wold ahold .S- : i 

(efom\~+2) omulov 1rslom epsiteve (56 


7 ¥ Ve 
, | . _ 1% ica 


— 


y 
@ 
A 
é 
. 


* 
> 





<7 ‘e x “si 


, 


n - _—. \ oc? ee ” a wMn ae 
CAV a) i £ CG $ \ { © Deh: sm « Vv) Pe = oMi VA | i. — ’ 
L . Pic © az a + 
~*~ é i e 7 
~ . a oN i 
‘ ~ ‘ , - <raie regen Sood ‘ } 
2 rosost noietavnos s et dLOTV{iess sienw- 
i » © , ¢ 


: : Ps er - 
‘ ~ 4 eo | « 
i ; : ,~km oF 32 pra 


| “ ~ 


iL=-2.8 


I-20 


ose age PA 


CO 


Psn 


ll 


FF 


at 


COS COS 


hae 


x (FF Far 
H5s H,S 


FFoos 


- Pugs) /2 


292 


where n is the average number of atoms in a sulfur 


molecule. 


Space Velocity (SCFH/gm cat.) and Space Time 
(gm _cat/SCFH) 


SVurx 


OVE S 


2 


SVcos 


SV 
SO. 


where 


= FF 


M 


= F 


Ho 


= Eco 


Fso 


rx/We ST tx 

_/WC Ea 
2 

_/WC aia 

/WC ST 

2 805 


= 1/SV 


1/Sv 
2 


1/SV 


1/SVoo 


MIX 


HS 


COs 


2 


WC is the weight of catalyst in gram. 


Fractional Conversions 


H5SCM 


CoSCN 


SO.CN 


Correction of Feed Composition for Water Injection 


FDCOM; 


(FF 


(FF 


(FF 


- P. .)/FF 
H.S HS HS 


) /FF 


cos ~ Pcos cos 


- P_..)/FF 
SO, SO, sO. 


Partial Pressures in the Reactor (mm Hg) 


eo aes 
1 


P; > ae 


reactor 












<<a 
i ‘ : 
, ‘ an 1 
y 
d+ ae wh 
20972 ~ 200 <9 
y 7 
q Ty 4 gq + . oth) & ¢.f{ = at : 
200 205 ~ &..H SB: 
1K ¢ § wy 
Se 
: Pa 7 x 
~ —_ — - ‘+. ® - B siw 2 
TO ‘Sons mi Spe I SVG suis et a St ’ = _ 
: . = j ri at, _ ia 
Ry,‘ A oe seins fom ’ 
eae bas .4(.36D. me\H We) yas :poLev x osge 8. 


Bi 3 ec orig fae amp) 


OWN yin? 


: . " . ‘ ~~ : 
ovr =< 7 . Ti > Time gale ve a | 
ais ae ie nee 


me te 


aonVeMt* geot® « i eos iS aod 


| 2 se Ad aa a aes 
r f o De y , a 
an.Ve\a * neo = 02 
’ 4 
gn wt Ge ¢ 
> > in 
r * a 

mesrp «ai JaYLSIBO 290 

cick 
8, 


- - ta? - 
stn seteW tot 4s 


———— 


I=-3. 


I-31 


where i=l for 
i= 2 for 
i= 3 ror, 
i= 4 for 
i=5 fox 
i= 6 for 
i=7 for 


Gas Chromatograph Data 


a) 


b) 


attenuator settings 
ATTEN #1 = 10.0 
ATTEN #2 = 5.0 
feed analysis 

Bo 
Measured area 


200960 


201201 


Corrected area 5583325 
5574808 
5581494 


Avg. corrected 5579876 
area 


201267 


on 


Sample Calculation for Run F-6: 


H,S 


200308 


202253 


203878 
200308 


203233 


202473 


203878 


COs 


219590 


278458 


276263 


279190 
278458 


276263 


eis 20 


SO. 


Bivees 


114565 


122204 


A eS 
114565 


122204 


Gal Ba des A 


433 


-02 


ee 


» CESVIL 


Goda ll 


 pOssse 


EssvilL 


eecatl 


boSsst 


peers orerys “ev bsOs | are 


ar. it 


209 





PeLers 
 p2aarts 
Easars 
oerers 
“Beners, 


eases 


ots 



















Pas 


sretos 


BOLDOS 


GESEOS - 
avagas 
a \ 
B0g008” 
_ REREOS, 


i 


GSES882 2 


’ Sa Pe 


ee a 


eonisoed rossunes3s oe 
i Se os + 14 metre 7 ch 





a i 
= 0 +. ~* St METTR. ue 
rae 
neers eae nook 4d ve 
* Aly aie 
eae gk os =i 
Tastos see bowensn 
| 2S Sth 
o8eo0s. es Sar sg 


5 ; at 


sa ie tah eras! 


aosates “te i. x fe 
ben oes | 


_ 






c) 


d) 


e) 


product analysis 
N HS 


Be. 


Measured area 199037 133398 
199785 131616 


200025. 132677 


Corrected area 5521462 133398 

Baer LSL61L6 

Soeanue70. 132677 

Avg. corrected 553 7515 72132564 
area 


area ratios 


Feed 
Ay o/Ay 0.036286 
2 2 
*cos/4n. 0.049817 
aa oN 0.021147 


mole ratios (percent) 


Feed 
Myos/™N, 3.254238 
Mcos/™n, 3.252806 
Moo /My 1/653483 


2 é 


COs 


244301 
244920 


243946 


244301 
244920 


243946 


244389 


Product 


SO 





62466 
63089 


62988 


62466 
63089 


62988 


62848 


0.023939 


0.044133 


0.011349 


Product 


2.149416 


2.881498 


0.898964 


———s 


on): 


eBC 
3 “ 
8b8Sa 


— > 
mAh 
7" 

: S 


soubor? 


+ 


eer. 
ourbor? 


ee gee 


at 


baes eB . 


5 


bBebL.S 


ep1es.S 


eidl - 28 eel “> aN 
“—~, is a | 
oc c be no i te 
; SEL @sooos " - 
Bectel gael [See SATB badser109 
; se i <f 
ach te P ve wel ry Ve 
OLGLCL Fisséce a S30) 
| — re 
4 br a mae 
rcace £ otaabee or 
7 wr. ot oorc ; i we: 
/ tee 
















saasel. @Levedd be¥petto> .pvA © 
; ted ope 

~ : . Ye 
20i3s1 5918 (be 

baet | in g ee i Dia ed 4 


Ne 


gasetv.0 Ne g 
TL8@b0.0 RNa 


RBLLS0.0 


1 ee 


£f) mole fractions 


Feed Product 
No 0.924549 0.944021 
HS 0.030087 0.020291 
COS 0.030074 0.027202 
SO. 0.015290 0.008486 


Temperatures {$C} 


a) = 1.40878 + 18.31152 (15.40) 


mie TS tay inlet 


283.41 


b) Mieco woes + 1O.21050 (15. 40) 


Teactor outlet 
233.05 


Pressure (mm Hg) 
a) Peoed (abs...) on 1 44,60136.,+442.21505 (40) 


Te Oowed 


b) P 


reactor (gauge) Passi 2 ee SO D037- (55) 


224.58 
c) Atmospheric pressure = 705.6 


Feed Flow Rate (SCFH) 


a) FFy =IUeusesereunsag99 yO) = 0.00099: (90) 
2 
= 9.052247 
b) Putx = (.092455x28 + .03009x34 + .03007x60 + 


.01529x64) /(.92455%22402.1 + 
“VIUCGUReet Ont iot .S007K2Z2417.51 + 
-01529x21901.63) 


= 0.001326 


292 


~~ + ehh 2.) Pen 
F . raat 




























soubor | boot | 4 
LSOBRE.0 * ebaase 0 yt in : ee i ny | im 
resoso.0 TEOOED.D fi ahs oat Sa 
sogrso.6 "ato0eO.0 ry - 209 ine ig Ages 
senso0.d peseio.8:” tr"? Tape tel Te fe 

EY ay eel s.-t 14 
(Ob.2L) S@LLei8L + BYBOR, a =. ot 


sbcat scanatee 


£b,€85 = tS elites ats 
(OB. 2£) oaats. Gf. + TSBLe,S = toltue xo2080x" poe 


20.688 m eo el 


5 : 
’ ; « 
: * 
i 7 
i ‘ “ . oie ‘es 
| : Hit) oxy £.£ 
ain. s ‘ . . 
on 5 2 
= 


rr 7 


(05) epett. Si + veers. hat = iG autai aa 
if we: = ies. 
»GS.a ease SL a oie ae a 
“ > Cee oP 9 Oe 


(22) TED20.b + ete rB, L= (seuiioh xogoBes! (a ee 
; iy 
82. pss = kt es bain 


ih 20% = Sesh, sneak (6 


‘4 3 bes hy 
+ Od&xTOOEO, + BExeOOEO.. + ass 


+f. ont 


a, 2 ' fe. TESSxTO0eO. 


4 
ri" 
CD ¢ 


rf Pie ae Ads 


a” a5) 
» 






e+ 
7 ” > 


4 oo “ep 4 


‘wh aie 4 a 


296 


p. == 28/22402114= 0.00125 
No 


FF wrx = 9.05225 x (.00125/.001326)°> = 8.789 


I-3.5 Component Flow Rates 
a) AVGMV =(.92455x22402.1 + .03009x22176.1 + 
.03007x22417.51 + .01529x21901.63) 
/ 28317.016 
= 0.790624 (£t?/gmole) 


b) component flow rates (gmole/HR) 


FFy = 0.92455 x 8.789/0.790624 = 10.2778 
y 
Ee s = 0703009 =x 8.789/0.790624 = 0.3345 
2 
FF aos =.0,03007 x 8.789/0.790624 = 0.3343 
FF oo = 0.01529 x 8.789/0.790624 = 0.1700 
2 
c) component flow rates (SCFH) 
"HS = 0.03009 x 8.789 = 0.26446 
Fuos =-0 703007 x 8.789 .= 0.26429 
F =O eOLoee Xo. foo =)0 513438 
SO, 
I-3.6 Product Flow Rates (gmole/HR) 
Py = 10.2778 
2 
Pa s = 1032778 = ~(0.020291/0.944021) = 0.22091 
2 
Poos = 10.2778 x (0.027202/0.944021) = 0.29616 
Pso = 10.2778 x (0.008486/0.944021) = 0.09239 










ester wolt 3 Snesiogmod C4 cr 
a ree oe Lage 


= wy oO es oe ~ oe ree, 
“f avy iccxe®obed .--4 e SOBSSxERESe Le VMOVA (s 


\ i a 
- ‘ - ‘ 
ce a 
" * 14> scnaor O24 
(wa ’ 32) o> eT 5 i ie? ee 
t ’ ek * aoa 
‘bore a i 7 F : a Ww é 
‘SH\elomp).zets1 wolt “Sa@qogaon (dt 


VOeT. O\NCSTA.B x ZAbSOLO = a 


s > é 


2bEeE.O = BSIOCT.O\CSY.B x QDOED.O = ) ae 


‘vopE 50.0 





oe 


297 


P = Oe (OTS SES = Os 2209 fe 0 Se EL36 
HO 
°co, = "0.3343 -°0.29616 = 0.0381 
Pon SLs ee ers a4 = eeeo9 - OF3343 <"0 729616) 
1 et 
= 0.0313 


Space Velocity (SCFH/gm cat) and Space Time 


(gm cat/SCFH) 


SVutx = -§.189/170Z13 = 8.6057 

SV = 0.26446/1.0213 = 0.2589 
H,S 

SVags = 0.26429/1.0213 = 0.2588 

SV = 0.13438/1.0213 = 0.1316 
SO. 

ST rx = 1/8.6057 = 0.1162 

Pelgs = 1/0.2589 = 3.8625 

STaos = 1/0.2588 = 3.8640 

‘Lee = 1/0.1316 = 7.5988 


Fractional Conversions 


H,SCN Set Usoto we User OIL 7 0.3345 = 0.3396 
COSCN = (0.3343 - 0.29616) /0.3343 = 0.1140 
SO5CN = (055700 — 0.09239)70.1700 = 0.4565 


Partial Pressures in the Reactor 


| MN 


P; = 11.07026 


i=l 





ves ° ; ied Gy Sera.) 


ra 





















Ri 


d€1L.0 = “(eoss 0 =- anf .0). + 0 ore i 


- Z 
Is£0.0 « ar9es.0 + ERED, 0" = a a 
(aLICS.0 ~ EEE, 0 + eOSs, ® - avet dl x 25 bs eaten °; 
: ho Ac aes Pee ae | 
} Ete0.0 = 
omit 9 ; 


(HD : } ay , cn, } 


e825,0 = ELSO.F\ObRBS.O * , Ve 
a f ry s ¢ 7 
8822.0 = ELS0.t\eSHaE 0 =) aos? tae 


 BLeEto “Sesto. INBEREL. Oe 


| ‘Sarr, 0 = ‘T2098. ML 


meen 


dsde/e = wes. wa _ 


aan 


‘ 


0h8.£ = sees. vig Fe 


Ae diy acd - a 
. asee ater, OL = ome 
1 


a a _ 
| pao pa: no taints 
“fA 


(Q@LE.O © -ZREE. Vues 0 es et 


ae or 


2920.0 = ones, estos oot 
the a seas 


pind hy os rete ey 
| mer 9 a9 We 


“ee be Te 

> "es oy ¢ 
as, vine r a : ae Ws ; 
n 1) i Be bg e . | 


err 2 j 
Me eee, a 


PP = (102778 x 


phe 
PP, = (0.22091 x 
2 
PPigg = (0.29616 x 
PP. = (0.09239 x 
2 
PRs egr=p(Galla6 ox 
2 
PP gg coe (0.0381 x 
2 
PP. = (0.0313 x 


The processed data 
the following units: 

Weight of catalyst 

temperature 

pressure 

space velocity 

space time 

feed rate 


conversion 


930.18) /11.07026 = 863.59 
osc nep7l..07026 = 18.56 
930.18)/11.07026 = 24.88 
930.18) /11.07026 = Lis 
930013) /11.0/026 = vier 
930,18) /11.07026 = a. a0 
B30 ¢Le@psLL.O7026 = £200 


in 


Table I.2 were based upon 


gram 


degrees Kelvin 


mm Hg absolute 


SCFH/gm catalyst 
gm catalyst/SCFH 
SCFH 


fractional conversion 


298 


¢ C08 mw @gf 


n 
re 
fe 
| 
f™ 
{Oo 
* 
bre) 
Ps | 
“1 
> 

. 
oS 
RA 
— 
fh 

s 
cad 
4 














Bf = OSOTO.LI\ (BL. 0F@ x. L00SS.0), = 


zz.@ © 9SOVO, LIX 4 (8£.00 oe x | VEEILO); @. 48 


a) <a 


le oe ay 
; Oe aoe Pye eS re 
| - QSOTO.LL\(BL.06€ x $8£0,0) = ga 5 


= _ bi 0 
~~ | ve 
: ~ - 5 he ¥ a; =. 
E3.$ = BSOTO.LL\ (BL. O£@ x ELE0.0) =. nee eit) 
3 - 4 ; rah 4 
ii 9 Te eee 
~ e “S 
a * 
é a 3 a 


1oqu Dbeesad eitew $.T sidgt nt ‘sa8b boeaeoorg oat f 
“ ig : 


2 fn A sad Laat ‘paiwolio? 2, 9f 
r . ’ ; ger, ee lie) 
mssp sevFodno to adotow' : 


ae: Po / H i‘ 
aivion Bsexpsh $ os a ay a 
. a F ‘ + oe a ae J nt : *, _ 






eee 1d Sls, ea 
ejuloeds pHi mt <5 9, 948 2010 
To) oe 
> : , ie iy Be 
seylseteo me mA) 3/6 ytio 





HaDe\; 


CALEVCY Cm G Gl 6) Ge CYS) OC) C1 Co Gt CO) 


BE AE AE HE IE HE I OE IE IE IS EI IC OE OK BE KE I OKC IC OIE IK OIE OE OK OK IE IE IE OIE IE IK OIC OIC OIC OE OK IK OK OC IK IK OIE KK SCD 3K IK 2K >} 
* 

* MAINLINE MTBAL 

bs 4 

* THIS PROGRAM COMPUTES THE MATERIAL BALANCE IN THE 

* INTEGRAL BED REACTOR FOR H2S-SO2-COS-H20 REACTION 

2esvS TEMCVINPULS TS REQUEST EDSON THE TELETYPE INA 

* FREE FORMAT INPUT STYLE. THE RAW DATA AND THE 

* PROCESSED DATA MAY BE ENTERED IN USER DEFINED DISK 
* FILES BY THE PROGRAM 

** PRESSURE eeeecceeeeMM HG 

* SPACEVVEUGCITY «..<.SCEH OF A FEED GAS/GM=CAT 

** SPAGE TI MEsiccsece eGM-CAT/SCFH OF A FEED GAS 

* COMPOSITIONseeeee eMOLE PERCENT 

x VOLUME ecccccvcvceee STANDARD CURIC FEET 

* CATALYST WEIGHT. eeGRAM 


* 


HE BE AK HS AE AE IC OE BE IS IC OIE IE IC IC OK AS IKE AIS CIE IC TIE OIE OS BIE AC IE IIS FE IE BIE OS IIS AS IC KC OIC IE AIS OK EE IS OIE IE ISIC DIS OIC OK 


DEFINE FILE 100 (1359119,UyNEXT) »200(13494,U,NEXT) 


a 


HoH eH eH He He HH He HF 


DIMENSION PRS(3)_TEMP(2) »PRESS(3),FDCOM( 10) »PRCOM(10) 


tk 
7 9 


LBAL( 2, 7) 9V(4)5TC(293)PC( 2,2) sDPC(3)¢PCR(4y6) »HC(2) 


1FCR( 4, 6) 

DATA V/22402 610, 22176.10,22417.251,21901.63/ 
DATA TC/1.40878,18.31152, 2.51827,18.21650/ 
DATA PC/744.613674912671505, 1.81379,4.05037/ 
DATA HC/0.31009,0.52032/ 

DATA DPC/0.08793 ,0.94999,-0.00099/ 


CALL GETTY(LUNR) 


LUNW=LUNR 
ISTRT=1 
IEND=0 


WRITE(LUNW,121) 


121 FORMAT( PILE STORAGE FLAG(I-STORE,2-NO) */'FUNCTION FL* 
1,*AG(1-—PROC NEW DATA OR 2—-PROC FILE DATA)*/*PRINT FLA! 


1,*'G(1-TTY,2—NONE ) *) 
CALL FFINP(LUNR,3,0,ITSTOR,O,TFUNC,0,IPR, IEROR) 
IF(TEROR) 998,432,999 
32 GO TO( 88,33), 1TFUNC 


READ DATA FROM DISK FILE 


33 WRITE(LUNW,122) 
122 FORMAT(*HOW MANY POINTS,AT WHAT RUN NO. IN FILE TO 


299 

























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pe? 
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ae i Pie 

Te TEtrtirC iT Tetris ett uk cK RSS corraeeee he i 
; ie, ; > 

: ; s Meh J 

{,aTM aut JMTAM Le ~ 

e a. I ‘yy 


75 £ one 


HT wt 4D Af JATASTAM AHT 25% aqnos mARonaa 20HT 
OLTQA OSH=-2no~She-28H AOA AGTOASR Oae PAROS TH aed _ 
Al YT3jaT AMF “uO GIT23V0sA VT TUAAT oMaTe¥e - <4 

ur ci ‘ATA WAS 3HT 23 ri TUQM! AMADA 3399 a 
12] Vilqad: ae v TRARTHA,, YAM -ATAG 032239089 * 
at ol “WARDORS SBT YR erg e 
z ae Satins aP 
oH MMs essen ae e+ SAUZ 2AM aie 


2A0 3 A an WAD2 ee es YT II0U8V. ADAM _ 
| A AN HADSNTAD“MDe eanecss Sel? guage. 
rMAIHSS 310M 4s ove b MOLT L209NOD ; 


te 


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oo 


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q. MAM ee Tl tO aw biel 


be rt ese ic ie ae ee ed nenenaeunen easy Stkee tes ee 


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eoe (CONT'D) 


* START!) 


CALL FFINP(LUNR,2,0yNUM,O,ITSTRTgIERDR) 
IF(ITEROR) 998,436,999 

ISTRT=ISTRT=-1 

TEND=ISTRT+NUM 

ISTRT=ISTRT+1 

NFIL=ISTRT 


READ (100'*NFIL) NFILyNFDCRyNPRCRy RUN gWC,TEMP, ((FCR(I y 


1J),PCR( I,J) s1=194)9J=196) sPRSyHPCTsHITMP,ATMP 


WRITE(LUNR,123) ISTRT 
FORMAT(*RUN NOw*1XsA4,1X_y "FINISHED ' ) 
GO TO 35 


READ DATA FROM TTY 


WRITE(LUNW,90 ) 

FORMAT('WT OF CATALYST(GM) $ NUMERIC FILE NO.') 

CALL FFINP(LUNRy291,WCyO,NFILyIEROR) 

IF(TEROR) 998,101,999 

WRITE(LUNW,100) 

FORMAT(*RUN NOwsNO] OF FEED GeCeyNQe OF PROD. GeC.') 


CALL FFINP(LUNR,393yRUN,OysNFDCR 40 sNPRCRgIEROR ) 
IF(TEROR) 998,103,999 

WRITE(LUNW,105) 

FORMAT('TEMP—REACT. INLET,OUTLET, PRESS-FEED,REACT.»D 


ree CELA 


CALL FFINP(LUNR,5y91y,TEMP(1)91,TEMP(2),1,PRS (1) 91 9PRS 
1(2),1,PRS(3), TEROR) 

IF(IEROR) 998,104,999 

WRITE(LUNW, 106) 

FORMAT( USPECIFY H2U0 FEED PUMP FCT AND TEMP") 


CALL FFINP(LUNRy2y1yHPCTy1,HTMPyIEROR) 
IF(IEROR) 998,107,999 

WRITE(LUNW,108 ) 

FORMAT ('ATMOSPHERIC PRESSURE'®) 


CALL FFINP(LUNR,1_1,ATMP,IEROR) 

IF(ITEROR) 998,109,999 

WRITE(LUNW, 113) 

FORMAT('FEED GeC~w ARFAS-N2,H2S,COS_SN2,H20,C02! ) 


DO 116 T=1,NFDCR 
CALL FFINP(LUNRy6y1y,FCR(T,1)91lyFCR( 1,2) 1leFCR( 193) 919 
LFCR( 194) 91lyFCR(I1,5)91,FCR(1,6),TEROR) 



































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eee (CONT'D) 


IF(IEROR) 998,116,999 
116 CONTINUE 


WRITE(LUNW,117) 
117 FORMAT(*PROD. G.eC. AREAS!) 


DO 120 I=1,NPRCR 
CALL FFINP(LUNR,6y419PCR(151)91,PCR(152)51,PCR(143)91 
*» PCR ( I,4)y91l,PC 
IR(1,5)y1l,PCR(I 56) ,1TEROR) 
IF(TEROR) 998,120,999 
120 CONTINUE 


GO TO (40,35),ISTOR 
STORE RAW DATA IN PNISK FILE 


40 WRITE(1OO'NFIL) NFIL»yNFDCR»NPRCR, RUN gWCy TEMP, ((FCR( I, 

1J) yPCR( 19d) pl =194) »J=1,6) yPRSyHPCT »HTMPyATMP 
35 CALL CHROM(FDCOM,FCR,NFDCR) 

CALL CHROM( PRCOM,PCR yNPRCR) 
DO 444 J=1;,2 
TEMP OT) STeUie T+ 1C(2,1)*TEMP(I) 
TEMALILSTEMECE) + 273. 

444 CONTINUE 


H20 FEED RATE CALCULATION (CC/HR) 
HRATE=HC(1)+HC(2) *HPCT 
H20 FEED RATE IN GRAM/HR 
HRATE=HRATE /(1.00244-0.00022*HTMP ) 
TFC HPCT)®© 23923724 
23 HRATE=0. 
REACTION TEMPERATURE 
24 RTEMP=(TEMP(2)+TEMP(1))/2. 


CALCULATE ABSTRACT PRESSURE OF FEEN STREAM AND 
REACTOR 


PRESS(1)=0. 

PRESS(2)=ATMP 

DO 17 I1=1;2 

PRESS(1)=PRESS(1)4PC( 151) *¥PRS(1) ¥*( 1-1) 
17 PRESS(2)=PRESS(2)4+PC( 15 2) *PRS(2) ** (1-1) 


FEED MIXTURE FLOW RATE 























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eee (CONTIN) 


PRS(3)=PRS(3)**0.5 
PRFSS(3)=0. 
DO 19 [=1,3 


19 PRESS(3)=PRESS(3)+DPC(1I)*PRS(3) **(I-1) 


10 


FLOW RATE CORRECTION 


ROMIX= (28.*FNCOM(1)+ 34.*FDCOM(2) +60.*FDCOM(3)+64.% 
LFDCOM(4) )/ (FDCOM(1)*V(1) +FDCOM(2)*V(2) +FDCOM(3)*V 
1(3) +FNCOM( 4) *V(4)) 

PRESS(3)=PRESS(3)* (28./V(1)/ROMIX) **O0.5 


FH2S = PRESS(3)*FDCOM(2) 
FCOS = PRESS (3)*FDCOM(3) 
FSO2 = PRESS(3) *FDCOM(4) 


SPACE VELOCITY AND SPACE TIME 


SV=PRESS(3) /WC 
ST=1./SV 
SVH2S=FH2S/WC 
SVCOS=FCOS/WC 
SVSO2=FSO2/ WC 
STH2S=1./SVH2S 
STCNS=1./SVCOS 
STSO2=1./SVS02 


MATERIAL BALANCE 


AVGMV=(FDCOM(1)*V(1)+FDCOM(2)*V(2) +FDCOM(3)*V(3)+FDCO 
1M(4)*V(4)) /28317.016 


FEED COMPONENT FLOW RATE (GM-MOLE/HR) 


DO 10 J=1%94 

BAL(1,J) = FDCOM(J)*PRESS(3) /AVGMV 
WFA = WC/BAL(1,2) 

WFB=WC/BAL(1,3) 


H20 FEED RATE (GM-MOLE/HR ) 
BAL( 1,5) =HRATE/18.0588 
PRODUCT COMPOSITIONS 


L=0 

BAL(2,1) = BAL(1,1) 

BAL( 2, 2) =BAL (2,1) *PRCOM( 2) /PRCOM(1) 
BAL(2,3)=BAL(2,1)*PRCOM(3)/PRCOM(1 ) 

BAL (2,4) =BAL (2,1) *PRCOM(4)/PRCOM(1) 
BAL(2,5)=BAL(1,5) + BAL(1,2) -— BAL(?,2) 

































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eee (CONT'D) 


BAL(2,6) = BAL(1,3)-BAL(2,3) 
BAL(2,7)=1.5 * (BAL(1,2)-BAL(2,2)+BAL(1,3)-BAL( 2,3) 


AVERAGE MOLECULAR WEIGHT OF SULFUR 
INITIAL ASSUMPTION XS=8. 


PPS = PRESS(2)/760.¥*BAL(2,7)/8. 

CALL FREM( PPS,RTEMP,XS ) 

PRS1=PRESS(2) /760.*BAL(2,7)/XS 
IFUABSTCPPSHPRSPYPPRST? -0.005) 20,20, 21 
PPS=PRS1 

GO TO 22 

BAL(2,7)=BAL(2,7)/XS 


H2S/S02 RATIO 


FHVS=BAL(1,2) /BAL(1 44) 
PHVS=BAL(2,2)/BAL (294) 
FCVS= BAL(1,3)/BAL(194) 
PCVS= BAL(2,3)/BAL(2,4) 


CONVERSTON 


H2SCN=( BAL(1,2)-BAL(2,2)) / BAL(1,2) 
COSCN= (BAL(1,3)-BAL(2,3)) / BAL(1,3) 
SO2CN= (BAL(1,4)-BAL(254)) / BAL( 1,4) 


ADJUST FEED COMPOSITION WITH H20 FEED 


TOT=0. 

DUSo -J=1,5 
TOT=TOT+BAL(1,J) 

NO 11 J=1,5 
FOCOM(J)=BAL(1,J)/TOT*100. 


PARTIAL PRESSURE IN THE REACTOR 

TOT=0. 

NO 12 J=l1l,7 

TOT=TOT+BAL(2,J) 

DO 13 J=ly7 

PRCOM( J) =BAL (24S) *PRFESS(2)/TOT 
DATA OUTPUT 

GO TO (43,44), IPR 


ty eOUT PUT 


CALL OUTPT(RUN,PRESSyRTEMPyFHVSyPHVS,FCVSyPCVS,H2SCNy 


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41 WRITE(200'*NFIL) 


998 
S97 


eee (CONT'D) 


1COSCNy, SO2ZCNy FOCOM,PRCOM,BALyLUNWyXSyWCy SV_ySTySVH2S,y ST 
1H2S,SVCOS,STCOS ySVSO2,STSO2,TEMP,IPR) 
44 GO TO (41,42), ISTOR 


PILE OU ERUT. 


WC »RTEMPyPRESSyRUN, BAL,yFDCOM, PRCOM, 


LH2SCN,COSCN ySN2ZCNyWFA WEB yFHVS yPHVSyFCVS yPCVSy XS_SVy 
1SVH2S,SVCOS,SVSO02 
42 IF(ISTRT-IEND) 34,65,65 


WRITE(LUNW,997) 


FORMAT(! 
GO TO 65 


FFINP CALL ERROR 


999 WRITE(LUNW, 996) 


596 
65 


FORMAT(! 
CALL EXIT 
END 


FFINP INPUT ERROR 


(EXE CALLED)-*) 


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xK *K 
7 SUBROUTINE FREM * 
* SUBROUTINE FREM IS USED TO CALCULATE EQUILIBRIUM ¥* 
* COMPOSITIONS FOR SULFUR MOLECULES BY FREE ENERGY x 
* MINIMIZATION METHOD * 
x Bs 
MK AE OK IE IK AE aI BIE Xe fs ae aie ok aig OI OK BK IK OK SIE IC IK IK OK SIC SI SIE KK OK SK AC OK OK DIK OK DIS BK OK IK IK OIE SIS OK OK OK OEE OK 3K 2K OK OK 3x 


SUBROUTINE FREM(PRESS,T,XS) 

DIMENSTON FRE(397)9X(3) »GA(39 3) ,GB(3),C(3) 9 F( 3) 9A(3,y 3) 
1,Y¥(3),B( 3) yNG1(3) GX (3) 

DATA FRE/7.7838968E+00 56.089 2429E+00 9 2-6999349E+00 » 
12.5099 820E-02 ,1.8824865E-02 »6.2749549 E-0 3,-3.2 /148310E- 


20 5, —2- 7861233E-05 5-9. 2870775E-06,2.6157310E-08,1.961 79 


BESE iy Os DDO et OLEH U9 s—1 es 2091 2SE-—1 27-5. S40684S6E-1 29-14 


4780 2282E-12,1.0114584F+04, 1.1264370E+04,144504935E+04, 


54. 7621792E+00 »7.3202322E+00 1 .0534222F+01 / 

DATA A/8.e. 9609220396*00.0/ 

DATA Kiletlearlal 

DATA YVils¢i<sief 

DATA B/16.52*0.0/ 

M=] 

N=3 

NO 12 I=1,N 

NG1(1)=0 
FRT=FRE(1,1)*(1.-ALOG(T))-FRE(I,2)*T/2.-FRE(1,3)*T 
1**2/6.-FRE( 1,4) *T*¥*®3/12.-FRE(1 95) *T**4/20.+FRE(1,6)/T- 


2FRE( 1,7) 


C(I)=FRT+ALOG(PRESS) 

TFC I“N) 12,12,13 

C(1I)=FRT 

CONTINUE 

DO 14 ITER=1,300 

CALL FREN (Y,Cy9FyYBARyNy NG1) 

MG =M+ l 

CALL GSET (A sY,GAyGByBeF ye MyMGyN) 
CALL GAUSS(GA;,;GB,MG,GX) 

DO 18 I=1,N 

IF(NG1(1I)) 19,19,18 

X(I)=-Y( 1) *( (C(1)4+AL0G(Y(1)/YBAR))-GX(1)) 
DO 21 J=l_M 

I1G= J+l 

X(T )=X(1)+GX( IG) *A(T J) *Y (1) 
CONTINUE 

CALL NEZE (XY 5N ,NG1) 

QUIT=1. 

DO 22 I=l,N 


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IF(NG1(1)) 23,523,22 
TEST=(X( 1) -Y(I))/X(1) 
TF ( ABS( TEST) -0.5E-04) 22,22,24 
QUIT=-1. 

CONTINUE 

IF(QUIT) 25,254526 

DO 27 IT=l,N 

Y(1)=X(1) 

CONTINUE 

DO 32 I=1,N 

NG1(1)=0 

IF(X(1)) 33933534 
Y(I)=X(T) 

GO TO 32 
Y(1)=0,000001 
CONTINUE 


XS=(86*Y(1)4+6.¥*Y(2)+2-*¥(3)) / (Y(1)+¥(2)+Y(3)) 


RETURN 
END 


306 



















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SUBROUTINE OUTPT * 

x 

THIS TSUBRUUTTNErPSPECIFRIES. Fre QUTPUT /FORM, OFUTHE ** 
MAINLINE MTBAL ON THE TELETYPE ** 


rd 
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HEE AIK BK CIE EE SE OE ICIS SE OC IEE OE IK AS EK OIE IC IE OE AS AE AK BIS OE AE FE EOS IE BIE FE OE IS BIS IS OK AS IS AIS IS SSK OK OK OC OK OK 


Het tH 


SUBROUTINE OUTPT(RUNsPRESSyRTEMP, FHVS,PHVS,FCVS,PCVS, 
LH2SCN,COSCNySOZ2CNsgFDCIM gPRCOMyBALy LUNWyXSyWCySVyST, 
2SVH2Sy_ STH2SySVCOS_ySTCOS»SVSO2,STSO2,TEMP,IPR ) 

DIMENSION PRESS (3),PRCOM(7) »FDCOM(7),BAL( 2,7), TEMP( 2) 


WRITE(LUNW,1) RUN 
1 FORMAT(//// TX "RUN NUMBER", 3X 9A4,///) 


12 WRITE(LUNW,210) WC 
210 FORMAT( 7X y*WT. OF CATALYST ',1X,F 10.3) 


WRITE(LUNW,219) TEMP(1),TEMP(2),PRESS(2) 
219 FORMAT(/ 7X_,*REACTOR INLET TEMP Petiwee! “1hs 
PeREACTORSUUICED Semen tae iae f 1Xy"*REACTION PRESSURE’, 
feLlv2s/) 


WRITE(LUNW,778)SV_9STySVH2Sy_STH2Sy SVCOS,STCOSsSVS02 
* STSO2 
Pre FORMAT 8() 32%, SPACE VELOCITY";sBX," SPACE TIME',/ 1X, 
1' BASE ON TOTAL FEED GAS',54X,F10.3910Xy,F1l0039/ (Xy'BA', 
PUSE ON HS FEED RATE -5X5F10.3910XsF 10.397 (Xs! BASE", 
3" ON COS FEED RATE',5XyF10.3910XyF10.39/ 7X_'BASE ON', 
ei SO2* FEED. RATE s5X5F10.5910X%,F10.3 ) 


WRITE(LUNW,218) PRESS(3) 
218 FORMAT(// 7X_y'VOLUMETRIC FEED RATE',F9.3,/) 


WRITE(LUNW,220) H2SCN,yCOSCN,SO2CN 
220 FORMAT(/ 7X_y"CONVERSION OF H2S '43XyF?7e29/ TX! CONVER! 
1y*SION OF COS '43XsFT.29/7Xy*CONVERSION OF SO2 '43Xy 
3 ehh Ie Me, 


WRITE(LUNW, 250) | 

250 FORMAT(/ 7Xy*MOLECULAR',5Xy "FEED! ,6X,'PARTIAL PRESSUR! 
1y'E',3X,'MATERIAL BALANCE 'y/ 8Xy"SPECIE',4Xy"COMPOSI! 
2," TION',4Xy"IN REACTOR',8Xy_ FEED’, 3X, !PRODUCT! ) 


WRITE(LUNW, 789) 
789 FORMAT( 17X_' (MOLE PERCENT) 'y5X_*(MM HG) ',8Xq_!' (GM 
x=MOL) SHR)! ) 


307 


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eee (CONT'D) 


WRITE(LUNW,211) (FDCOM( J) yPRCOM(J) yBAL( 1,5) _yBAL(2,J) 
*, J=1, 5) ; 
211 FORMAT (/10X9*N2 "y7XyF6e29 9Xq Fel ye BX eF le 39 2X Fle3y/ 
1/10 Xy "H2S' 9 7X gF6e2 9X gF6el gy BX 9 Fle3dg2XeFleo3e//10Xy 
2'COS', 7X yF60249X 9 F601 9 8X 9F le 39 2X 9 Fle 3e//1LOX9!# S024, TXy 
ZFOe2y9Xe Fe le BX eFle3de2X oF leo 3d9//10Xy'H20! g7Xg F602 99Xy 


P61 8X sh 1 sa e2XeF le 3s ) 
WRITE(LUNW,911) (FDCOM(I+5) ,PRCOM(1+5),BAL(1,I1+5), 


LBAL( 2, 1+5) ,T=1,2) 
911 FORMAT (/10X9!*CO2'y7X yF6e299X yg F601 y BX yFT eo 39 2Xy Flo 39 / 


L/1OXe*SX Fy TXgF60299X gF6el BX gFleo3y2XyF 723 ) 


WRITE(LUNW,212) XS 
212 FORMAT(//, 7X,"AVERAGE NO OF SULFUR ATOMS/MOLECULE = ! 
*y FO. 2) 


Pra GIP R= 2 eS eier hs 
13 RETURN 
END 


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x * 
* SUBROUTINE CHROM * 
* SUBROUTINE CHROM IS USED TO CALCULATE THE COMPO-— * 
* SITION OF FEED AND PRODUCT STREAM FROM THE GC AREA* 
* RESULTS * 
*K * 
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BE AEA HE EE IE OKC IE OIE IE OK CII AIC IS OK OK OK A IS OK OK AIK ISK OK OE IS OK OS DIS IK AC AS OS KC OK OK OIC IK OK OK OKIE OK KC KOK 3 


SUBROUTINE CHROM(COMP,CRyN) 

DIMENSTON COMP(10) ,CR(496) sHSN2(2),SON2(2),COSN2(2) 
*y ATN( 2) 

DATA HSN2/0.00733 50.89481/,S0N2/0 .02466,0277038/ 
DATA COSN2/-0.00149,0.65325/ 

XN=N 


CORRECTED AREA OF N2 PEAK 


DO 20 I=1,N 
CR(Ig1)=CR(191)*(1.76047*ATN(2)+ 1217312) *(0.17031 
xX ATN(1)+1.07781 ) 


AVERAGE AREA OF EACH PEAK 


DO 1 J=1,6 
CR(4,J)=0. 
NO 1 T=1,N 
CR(4, J) =CR(49J') +CR(1,J) /XN 
IF(CR(4,2)) 393,95 
HNRAT=0. 
GO TO 7 
HNRAT=HSN2(1) + HSN2(2)*CR(4_2)/CR(4,1)*100. 
IF(CR(4,4)) 858,49 
SNRAT=0. 
GO TO 14 
SNRAT=SON2(1) + SON2(2)*CR(4_4)/CR( 4,1) *100. 
PF(ICR(4,3)) 15,15917 
CNRAT=0. 
GO TO 16 
CNRAT=COSN2(1) + COSN2(2)*CR(4_,3)/CR(4,1)*100. 
COMP(1)=100./ (100.+HNRAT+SNRAT+CNRAT ) 
COMP (2) =HNRAT*COMP(1)/100. 
COMP( 3) =CNRAT*COMP(1)/100. 
COMP (4) =1.-COMP(1)-COMP(2)-COMP (3) 
RETURN 
END 


309 


























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x 

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* IS SOLVED USING SUBROUTINE GAUSS. 
ok 


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SUBROUTINE GSET (AsYsGAsGByBeFy MyMGyN) 
DIMENSTON R(393)9A(393)9Y(3)9GA(353)96B( 3) »B(3),F(3) 
DO or K=1,M 

DO 1 J=1,K 

R(J,K)=0.0 

DO 2 L=feN 

R( Je KIER SoK FACT oJ) ¥ACT »K)¥Y(T) 
R(KyJ)=R(JyK) 

NO 3 IT=1,MG 

DO 3 J=1,MG 

GA(I,J)=0.0 

DO 4 IG=1,M 

DO 5 T=1,N 
GA(IG,1)=GA(IG,1)+A(1,1G)*Y (IT) 
No 9 JaliM 

JGsJ+ . 51 

GA(IG,JG)=R( 1G ,J) 

JG= IG+1 

IGG=M+1 

GA(1GG,JG)=GA(1G,1) 

DO 10 J=1,_M 

GB(J)=B(J) 

NO 10 T=1,N 
GB(J)=GB(J)+A(1T,J)*F(T) 

JGB=M1 

GB(JGB)=0.0 

DO 11 T=1,N 

GB(JGB)=GB( JGB)+F(1) 

RETURN 

END 


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* Kk 
* SUBROUTINE GAUSS * 
x xx 
* THE FUNCTION OF THIS SUBROUTINE IS TO SOLVE THE * 
* SET OF EQUATIONS A*X=B USING GAUSSIAN ELIMINATION * 
* AND BACK SUBSTITUTION ROTATING ABOUT THE ELEMENT x 
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a * 
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SUBROUTINE GAUSS (A gRyN,X) 
DIMENSTON A(3593),R(3),X(3) 
M=N-1 

DO 11 J=1l_M 

S=0. 

Ho. SrA IT=J,N 

U= ABS(A( I,J) ) 


TF(U-S) 12,12,112 


S=U 

L=] 

CONTINUE 

BFL L=J3)  LiSy it slid 

NO 14 T=J,N 

S=A(L, 1) 

A(L,I)=AlJe1) 

A(J,I)=S 

S=R(L) 

R(L)=R(J) 

R(J)=S 

Leis ADS CAs Jie b—S008 bids 115,15 
WRITE( 6,3) 

FORMAT (1H »'MATRIX SINGULAR !) 
RETURN 

MM= J+ 1 

DO 11 I=MM,N 
PPC ABS UAC so mtb 30) aie lds bi 
S=A(JyJ)/A( Ty J) 

ACI, J) 20.0 

No 16 K=MM ,N 
A(T,yK)=A(J9K)—-S¥A( 14K) 
R(I)=R(J)-S¥*R(T) 

CONTINUE 

DO 17 K=1,N 

I=N+ 1-K 

S=0.0 

TFC I“-N) 117,17-117 

MM=I+1 

DO 18 J=MM,N 


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eee (CONT'D) 


18 S=S+A(1,J)*X(J) 

17 X(T)= (ROT) -S)/A(T,1) 
RETURN 
END 








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SUBROUTINE FREN 


THIS SUBROUTINE CALCULATES THE FREE ENERGY 
CONTRIBUTION OF EACH SPECIE TO THE SYSTEM. 


Heke uUt He HK € 
HHH HH H HH 


BE A BIE OK IE OK AE OIC IK AE IC IC KK BK AS KK AE OK OIE OE OK IK IK IE OK DK OK OK OK AIS AIS OK DIS DIE AE IS IK OK IC AS AE OK FE IS COIS IS OKC OC OK 2K 


SUBROUTINE FREN (Y9CyFsYBAR»Ny NG1) 
DIMENSION Y(3) 5C(3) 9F (3) 9NG1(3) 
YBAR=0.0 

DO 1 I=1,N 

YBAR=YBAR+Y (I) 

DO 2 T=1,N 

IF(NG1(1I)) 35394 
F(IT)=Y(1I)*(C(I)+ALOG(Y(1I)/YBAR) ) 
GO TO 2 

F(1I)=0.0 

CONTINUE 

RETURN 

END 


A 


















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SUBROUTINE NEFZE 


% % % + 


THIS SUBROUTINE TESTS FOR NEGATIVE OR ZERO AMOUNTS 
CF MOL ECUDAR SPECIES: ANU TAKES THE CORRECTIVE 
ACTION AS INDICATED IN THE METHOD REVIEW. 


4% 
3% 3 3 3t 3 3 


x 
BEAK IE DK RIE 2K SIE IC DI AIC OK ICC 3 OK OK IK OE ANC EOI A OK OK DIK OK SIE 2 CIE AK DIK IE OK SIS IK SK IK IS OE DIS OIE OO OK OKIE OIE OE OK IS aN OK 


SUBROUTINE NEZE (XY _N yNG1) 
DIMENSION X(3)9Y¥(3)9NG1(3) 
TEST=1.0 

DO 1 I=1,N 

IF(NG1(1)) 29291 

IF(X(1I)) 393 91 
SLAM=-0.99*Y(T)/(X(1)-Y(I)) 
TF(SLAM-TEST)494 1 
TEST=SLAM 

CONT INUE 

DO 5 IT=1,N 

TF(NG1(1)) 797,95 
X(T)=Y¥CT)+TEST*(X(1)-YCT)) 
IF(X(1)-0.10E-10) 6 96,5 
X(1)=0.0 

NG1(I)=1 

CONTINUE 

RETURN 

END 


314 
















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315 


TABLE I.1 


>QMTBAL 
FILE STORAGE FLAGC1-STORE»s 2-NO) 
FUNCTION FLAGC(1-PROC NEW DATA OR 2-PROC FILE DATA) 


PRINT FLAGC1-TTYs 2-NONE ) 


>2 1 1 

WT OF CATALYSTCGM) $ NUMERIC FILE NO- 
>1-0213 1 

RUN NOesNOe OF FEED GeCesNO~e OF PRODe GeCe 
>F-06 3 3 


TEMP*REACTe INLETs»OUTLETs PRESS-FEED»s REACT-»D/P CELL 
>152¢4 1564 40¢ SSe 91-8 

SPECIFY H20 FEED PUMP PCT AND TEMP 

>Be Be 

ATMOSPHERIC PRESSURE 

>7956 

FEED GeCe AREAS=-N2sH2Ss2COSs SO02sH205 C02 
>201267 203878 279198 117223 @ @ 
>200969 290308 278458 114565 BB @ 
>201201 203233 276263 122204 @ @ 

PROD. GeCe AREAS 

>199037 133398 244301 62466 32932 21117 
>199785 131616 244920 63089 30115 29994 
>200025 132677 243946 62988 27860 20994 


















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TABLE I.2 
RUN NUMBER F-86 
WTe OF CATALYST 1-821 
REACTOR INLET TEMP 556+ 48 
REACTOR OUTLET TEMP 556-05 
REACTION PRESSURE 9380-18 
SPACE VELOCITY SPACE TIME 
BASE ON TOTAL FEED GAS 8-611 H-116 
BASE ON H2S FEED RATE 6-259 3°859 
BASE ON COS FEED RATE @-258 3-861 
BASE ON S02 FEED RATE 0-131 72594 
VOLUMETRIC FEED RATE 8-795 
CONVERSION OF H2S 0-33 
CONVERSION OF COS B-11 
CONVERSION OF S02 8-45 
MOLECULAR FEED PARTIAL PRESSURE MATERIAL BALANCE 
SPECIE COMPOSITION IN REACTOR FEED PRODUCT 
(MOLE PERCENT) CMM HG) CGM-MOL)/HR) 
N2 92-45 863-5 16-285 10-285 
H2S 3°08 18-5 0-334 @-221 
cos 3°80 24-8 0-334 0-296 
S02 1-52 77 0-170 8-892 
H20 0-86 95 0-800 8-113 
co02 0-80 3-2 0-880 0-838 
SX 8-AB 2-6 0-869 6-631 


AVERAGE NO OF SULFUR ATOMS/MOLECULE = 7227 


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EXPERIMENTAL DATA FILE 


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Experimental data were tabulated using the following 


symbolic representation. 


Catalyst 
l= 
2= 
3 = 


Aa 


Attenuation scheme; 


2= 
3 = 
Experimental 
A= H,S 
B = COS 
Cp COn: 
D = COS 
E = COs 
F = COS 
G = COS 
H = COS 
I = COs 
J = 


type; 


y-alumina 


5.4% Cu - 


12.08% Cu 


16.07% Cu 


(Kaiser S-201) 
alumina 


alumina 


alumina 


attenuation scheme II in Appendix A 


attenuation scheme III in Appendix A 


run; 


SO. 


BeO 


Z 


HO 


2 


reaction on catalyst type 2 
reaction on catalyst type 2 
- so, reaction on catalyst type 2 
reaction on catalyst type 2 
at SO, reaction on catalyst type 2 
- SO, reaction on catalyst type 3 
_ SO. reaction on catalyst type 4 
= SO. reaction on catalyst type l 


reaction on catalyst type 1l 


Test of bifunctional characteristics of 


catalyst type 3 


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reaction on catalyst type 1 (SV = 4 hr‘) 


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