CALIFORNIA STATE UNIVERSITY, NORTHRIDGE
GAS PURIFICATION BY CHEMICAL SOLVENT
A graduate project submitted in partial satisfaction of the
requirements for the degree of Master of Science in
Engineering
by
¥dchael Philip Caldwell
August 1984
Protect of Michael Philip Caldwell is approved:
Ileana Costea
RajdeirBabayan
Chair
California State University, Northridge
ii
TABLE OF CONTENTS
Page
LIST OF TABLES
• vi
LIST OF FIGURES
• vii
LIST OF SYMBOLS •
• X
ABSTRACT •
xii
Chapter
1.
INTRODUCTION
2.
THE REMOVAL OF SULFUR DIOXIDE •
3.
1
. 6
Standards •
. 6
Throwaway process •
. 6
Regenerable process •
. 7
Double alkali process •
• 7
Sulfite process
• 8
Citrate process •
. 8
Dry scrubbing
• 8
THE REMOVAL OF CARBON DIOXIDE •
10
Standards •
10
General thermodynamics
10
Hydroxide solution •
12
Amine process •
12
Carbonate solution •
18
Flow arrangements
19
iii
4.
.......
........
..
THE REMOVAL OF HYDROGEN SULFIDE
Standards • • • • • • • • • •
General thermodynamics
Amine process
.
Shell process
.
a
•
•
•
•
•
•
•
•
•
•
•
•
•
•
•
................
....• .. ..
•
21
21
21
22
22
.. .
22
.
THE REMOVAL OF HYDROGEN SULFIDE AND
CARBON DIOXIDE . . . . . . . . . . . . . . . . . . . . . 24
Oxidation process
5.
Cases and standards • • • • • • • • • • • • • • • • • •
24
Inorganic salt process •
24
Carbonate solution • •
• • • 25
Organic solutions • • • • • • • • • • • • • • • • • • •
6.
ABSORPTION •••
......
Introduction •
...
....
Absorption equipment •
• • • 27
•• 27
• • 29
• 42
Chemical and physical absorption •
Diffusivity and mass transfer • • • • • • • • • • • • •
7.
..
SELECTION OF A PROCESS •
......
Carbon dioxide removal
• 59
..
66
• 66
Simultaneous removal • • •
Selective removal • • • • • • • • • •
.........
THE DESIGN OF A GAS PURIFICATION PLANT
USING MDEA FOR SELECTIVE H2S REMOVAL • • • • •
The design selection •
Assets of MDEA • • •
48
• • • 61
Hydrogen sulfide removal •
8.
26
...
Use of a computer for the design •
•• 74
• 74
.....
. . . . . . • • • 74
. . . . . . • 75
Finding or fixing boundary limits • • • • • • • • • • •
iv
66
75
Mass and energy balance around absorber • • • • • • • •
81
Heat exchangers duties • • • • • • • • • • • • • • • • • 82
Absorber diameter • • •
• • • •
=
~
e
~
•
•
•
•
•
•
Mass and energy balance around regenerator • •
REFERENCES
..........
•
83
• 85
..
Regenerator diameter •
•
• 86
• 87
APPENDIXES
A.
Flow diagram for MDEA Gas Treating Plant • • • • • • • • 89
B.
Listing of computer program • • • • • • • •
c.
Sample input and output of computer program •
v
90
. . 101
TABLES
Table
1.
2.
3.
4.
Page
Impurities Encountered in Gas
Treating
• • • • • • • •
• • • • • • • • • • • • 2
Commercial Use of Different Types
of Solvent Gas Treating Processes •
Comparative Cost of the MDEA Process
and a Total Gas Sweetening Process
(CHEMERY Case) • • • • • • • • • •
8
..........
76
Comparative Cost of the MDEA Process
and a Total Gas Sweetening Process
(Fuel Gas Treatment Case) • • • • • • • • • • • • • • • 77
vi
FIGURES
Page
Figure
1.
General Layout of an Absorption Plant . . . . . . . . . .
2.
First and Second Dissociation
Constants of Carbonic Acid •
3.
4.
5.
6.
7.
8.
Equilibrium Pressures of CO Over
Aqueous Aminoalcohol and Potassium
Carbonate Solutions at Absorber
Top Conditions (40 °C) • • • • • •
10.
11
14
Arrhenius Plot of Second-Order Reaction
Rate Constants for co - Hydroxyl Ion
2
and C0 - Amine React1ons
• • • • • • • • • • • • • • • 15
2
Flow Diagram for a Traditional MEA-C02
17
Scrubbing Unit • • • • • • • • • • • • •
Hot Carbonate C0
20
Removal Flowsheets • • • • •
2
Idealized Solvent Absorption Acid Gas
Removal Process • • • • • • • • • • • • • • • • • • • • 28
Example of a Two-Stage Tube Bundle
Co 1 'U1Jl1l
9.
..........
5
•
•
•
•
•
•
•
•
•
•
•
•
•
...........
30
Design of a Two-Stage Packed Column • • • • • • • • • • • 31
Design of a One-Stage Fluidized
Packing Column • • • • •
..........
..
11.
Design of a Jet Absorber •
12.
Three Different Plate Arrangements
13.
Design of a Hultistage Rotating
33
34
36
Disk Column • . . • . . . • . . . • • . • • . . • . . . 37
. . . . . . . . . . . . . . . 38
a Nozzle Spray Column . . . . . . . . . . . . . 39
14.
Design of a Venturi Column
15.
Design of
16.
Design of a Rotating Disk Spray
Column
........................
vii
40
17.
18.
19.
20.
21.
22.
23.
Application of a Few Absorbers as
a Function of Gas FlowRate and
Height-to-Diameter Ratio • • • • • • •
........
43
Dependence of Specific Interfacial Area
a on Bubble, Particle and Tube
Diameter d • • • • • • • • •
...........
44
Specific Interfacial Area a as a Function
of Volumetric Energy Demands for Various
Types of Absorbers • • • • • • • . •
• • • • • • • •
45
Henry Coefficient H for the Solution of
Various Gases in Water as a Function
of Temperature • • • • • • • • • •
........
47
Limiting Mass Transfer Conditions in
Absorption • • • • • • • • •
• • • • • • • • • •
52
Mass Transfer Resistance (MTR) in
Absorption Columns • •
56
...........
Selection Methodology for Acid Gas
Removal Sol vent • •
• • • • • • • • 60
24.
Range of Gas Treating Process Application
for co Removal with no H S or Other
2
2
Acidic Impurities Present • • • • • • • • • • • • • • • 62
25.
Range of Gas Treating Process Applicaton
for H S Removal with no co or Other
2
Impur12ties Present • • • • • • • • • • •
26.
27.
.......
63
Range of Gas Treating Process Application
for Simultaneous H S Plus co Removal
2
2
with no Other Impurities Present • • • •
64
Range of Gas Treating Process Application
for Selective H s Removal in the Presence
2
of C0 • • • • • • • • • • • • • • • • • • • • • • • •
2
65
28.
H s/co Solubility Ratio for Organic
2Phys1cal
2
Solvents • • • • • • • • • • • • • • • • • • • 68
29.
Arrhenius Plot for K
cs
30.
Second Dissociation Constant of H2s
31.
Computer Graphics Diagram of MDEA
Gas Treating Plant • • • • • • •
32.
.................
. . . . . . . . . . .
Equilibrium Diagrams for H s and MDEA
2
viii
.......
69
71
• 78
. 79
33.
34.
35.
Computer Graphics Equilibrium Diagram
for H S and MDEA • • • • • • • • • •
• • • • • • • • 80
2
Generalized Pressure Drop Correlation for
Dumped Packings • • • • • • • • • • • • • • • • • • • • 84
Flow Diagram for MDEA Gas Treating Plant • • • • • • • •
ix
89
SYMBOLS
A, A'
The volatile component(s)
a, a'
Concentration of A, A' in liquid
a
Value of a that would equilibrate local composition
B
A chemical base
B.
The jth non-volatile component
b.
Concentration
of B.J
.
'b.
Value of b. normalized to molarity, b.= b./m
c
Concentration
D
Diffusivity of A
f
Fugacity
H, H'
Henry's law constants for A, A' in reactive solvent
Ho
Value of H in pure solvent
I
Enhancement factor
I.1
Ionic strength
IQO
Value of I in instantaneous reaction regime
K
Thermodynamic equilibrium constant
K
g
Overall mass transfer (pressure units)
Kl
Overall mass transfer (concentration units)
k
Reaction rate constant
kG
Gas phase mass transfer coefficient
kL
Chemical mass transfer coefficient in liquid phase
L
Solution volumetric flow rate
J
J
J
/'
J
J
X
J
M
Mass flow rate
m
Solution molarity
N, N'
Rate of absorption of A, A' per unit interface area
(negative in desorption)
nA
Molar flux density
p, p'
Partial pressure of A, A'
p*, p*'
Equilibrium values of p, p'
Po
Vapor pressure of pure component
R
Gas constant
r
Rate of reaction
r
0
Rate of reaction at interface per unit area
ST
Thermodynamic selectivity
T
Temperature
tD
Diffusion time
t
Reaction time
r
u
Average velocity
v
Partial molar volume
X
Distance from interface
y
Gas phase mol fraction
y, y'
Fractional chemical saturation
y
Total saturation,
C1(
y
=Or /m
Total concentration of A, physically dissolved
plus chemically combined
ratio,~.=~ /b.
J
0
JO
«.
J
Concentration
13
Local mass transfer coefficient
Activity coefficient
Viscosity
Ratio of diffusion and reaction times
xi
' .
ABSTRACT
GAS PURIFICATION BY CHEMICAL SOLVENT
by
Michael P. Caldwell
Master of Science in Engineering
The reasons for purifying gaseous streams range from economic
considerations to environmental issues.
Economics dictates the
removal of an undesirable gas constituent from a process stream
before the stream undergoes further treatment.
Some processes demand
that the concentration of an interfering gas pollutant be as low as
parts per billion.
Environmental agencies are trying to lower the
rate of air pollution by placing restrictions on the waste gases
released to the atmosphere.
The use of chemical solvents for gas purification as opposed
to other forms of pollution control is due to the chemical solvents
ability to remove more pollutant while using less energy.
Sometimes
chemical solvents have to be used because other methods will not
remove a sufficient amount of contaminant.
With this greater ability
of contaminant removal comes the increased complexity of the process
xii
design.
The chemical reactions associated with chemical solvents
add to the unknowns which much be found if an economic process is to
be designed.
The major portion of the design is associated with the
process of absorption.
This is the actual dissolving of the gas into
the liquid stream by physical and chemical driving forces.
Mass
transfer, heat transfer, thermodynamics, and kinetics must all be
taken into account when choosing and designing absorption- equipment.
This report deals primarily with carbon dioxide and hydrogen
sulfide removal.
This can be the removal of both constituents, or
the removal of only one constituent; this is called selectivity.
The
design aspect of this report is based on one such chemical solvent to
selectively remove hydrogen sulfide from a gas stream while leaving
most of the carbon dioxide.
The chemical solvent which will do this
in the most economic way is methyldiethanolamine (MDEA).
The mass and energy balance of an MDEA gas treating plant is
done by a computer program.
The program will determine the amount of
MDEA needed for a certain gas flow as well as the tower diameters of
the absorber and regenerator.
The program also determines the duties
of the heat exchangers, condensor and reboiler.
is minimal and certain parameters are changeable.
xiii
The operator input
Chapter 1
INTRODUCTION
Gas purification is the removal of impurities from a process
gas, whether waste gas or gas about to undergo further treatment
[1:3].
The main impurities encountered are listed in table 1.
The removal
of carbon dioxide and hydrogen sulfide is referred to as acid gas
removal, and the removal of sulfur dioxide is referred to as flue gas
desulfurization.
different ways.
These gaseous impurities can be removed in two
The process gas can be brought into contact with a
medium which will absorb the impurities or chemically react with the
impurities.
Bulk removal of impurities is accomplished by absorption
into a liquid, adsorption onto a solid, or cryogenic separation.
Trace contaminants are removed by chemically changing the impurities
to other compounds or by adsorption [1:4].
Bulk removal is lowering the impurities from that of a high
level, down to that of 0.1-2.0 percent.
This is most often
accomplished by absorption, either physical or chemical.
Adsorption
is generally used in smaller plants, and cryogenic methods are the
least used [1:8].
The final or trace removal of impurities results in a
purification of the gas down to parts per million of contaminants
[1:8].
The chemical and physical solvents used for trace removal are shown
1
2
Table 1
Impurities Encountered in Gas Treating
Type of impurity
Acid gas removal
Flue gas desulfurization
Al'ld gases
Carbon dioxide
Sulfur
dioxi.Jt~
((01)
Hydrogen sulfide
(H 2 S)
Organic 'iulfur compounds
Carbonyl sulfide
(COS)
Carbon disulfide
(CS 2 )
Thiophene
Mer:::~:Jtam r RSH)a
Organk su!fides
(RS:\R, l<SR)
Oth~r
impurities
H20
HCN
NH 3
Hydrocarbons
Particulates
so 2
so]
Tar
aR is u,cd to designate an
:~lkyl
gruup.
Source: Astarita, Savage, and Basio, p. 4.
Parttculates
NOx
(S0 2 )
3
in table 2.
All of the solvent absorption methods use variations
of the flowsheet shown in figure 1.
The raw gas is contacted with
regenerated solvent and the impurities are absorbed.
The treated gas
leaves the top of the absorber and the now contaminated solvent
leaves at the bottom.
This rich solvent is reduced in pressure and
is heated by the returning clean solvent.
The rich solvent is
stripped of the impurities in the regenerator and is recycled [1:13].
These gas cleaning plants are integrated with other plant
processes.
The regenerator in the purification process may get its
low level steam from another process.
The impurities from the gas
cleaning can be used in other process, thereby reducing the cost of
the necessary removal.
The hydrogen sulfide removed may be sent to a
sulfur plant, and the carbon dioxide removed may be sent to an ammonia
plant or may be used for enhanced oil recovery [1:13].
4
Table 2
Commercial Use of Different Types of
Solvent Gas Treating Processes
Solvent
Add ga<; TC!T'OY:il
A Y. u.:ous alk ano!arnine
MEA
DEA
DGA
DiP A
Promored hot potassium carbonate
Organic pn:.n10ters
Inorgamc prorno.ters
Org:..il,.- ·;olvent-,dkanc,lami:lc:
Su!folanciDIPA
Ntun'1er o~
instalLttions
>! Ol'CJ
>740
>130
M,~OHjMEA/DEA
Aqueous ~ulution of potassium salt of amino acids
Ph:-,,i.:..'.l (organid solvents
Propy kne carbonate
Polyethylene glycol dialkyl ether
/\ -m~rhylpyrrolidone
Chi!kJ methanol
Flue g;.;s dcsulfurization
Lim<lirues~one slurries in water
Sodium sulfite
Source: Astarita, Savage, and Basio, p. 8.
--J 00
73
5
a
I' .
IJbsorption
umt
il i
,,.!'=
NA:
i!
r~gtmr:ral1on
Uf"!ll
.
;!
r,
{absorpt,~J
Figure 1
General Layout of an Absorption Plant
Source: Brauer and Varma, p. 244.
.I
Chapter 2
THE REMOVAL OF SULFUR DIOXIDE
Flue gas desulfurization is the removal of sulfur dioxide
from combustion gases or other forms of waste gas.
The sulfur
dioxide must be removed for environmental reasons since this is the
source of acid rain.
With the increase of coal usage, flue gas
desulfurization technology must be increased to insure clean air
standards [1:293].
The 1980 standards for sulfur dioxide in flue gas must
satisfy the tighter of two requirements.
The absolute emission is
limited to 1.2 lbs/10 6 BTU with sulfur dioxide being removed by
70-90 percent.
This is broken down into two parts.
If the total
r
emission is less than 0.6 lbs/10° BTU, then the sulfur dioxide needs
only be removed by 70 percent.
When the emissions are more than 0.6,
then the sulfur dioxide needs to be decreased at least 90 percent.
This regulation basically requires scrubbing of low sulfur coal
[1:293].
There are two basic categories for flue gas desulfurization.
In the throwaway process, a cheap base, such as limestone, is used to
react with the sulfur dioxide to form a disposable solid.
In the
regenerable process, the base is recycled after it is cleaned.
sulfur can usually be recovered as elemental sulfur [1:295].
In the lime wet scrubbing process, sulfur dioxide can be
removed up to 90 percent.
The reactions are:
6
The
7
CaO + H20 ---> Ca(OH) 2
Ca(OH) 2 + C0 2 ---> CaC03 + H2o
CaC0 3 + C0 2 + H20 --->Ca(HC03 ) 2
Ca(HC03 ) 2 + so2 + H20 ---> CaS03 ·2H 2 0~ + 2C0 2
CaS0 3 ·2H20 + 0.502 --->CaS04 ·2H 2 o~
[10:356].
The solids are sent to a settling pond and the liquid is
recycled.
Many problems occur with this type of process such as
scaling, corrosion erosion and solid-waste disposal.
.I
Since the
scrubbing cools the waste gas, the gas must now be heated to achieve
buoyancy [10:356-7].
Magnesium oxide can be used in much the same way as lime in
the lime scrubbing method but, magnesium oxide is regenerative.
The
sulfur dioxide is scrubbed with a solution of magnesium hydroxide
which leads to the production of magnesium sulfite.
Separation and
calcination regenerates the magnesium oxide.
This system generates
no waste because the magnesium is recycled.
However, the calcination
used for this recycle requires heat above that which is needed to
raise the temperature of the waste gas [10:359-60].
The disadvantage of the throwaway system mentioned earlier
can be overcome by using the double alkali scrubbing technique.
The
flue gas is first scrubbed with a sodium oxide and this spent liquid
is sent to the second reactor where it is mixed with lime.
The lime
reacts with the sodium bisulfite to form sodium sulfite which is
recycled.
The resulting calcium sulfite/sulfate is suitable for land
fill [10:363].
Although sulfur dioxide could be absorbed by water alone, the
circulation rate and stripping steam rates would be too large to be
8
economic.
These rates can be lowered by the use of alkaline
solutions.
Two such processes use either a sulfite solution or a
citrate solution.
One of the best known methods is the Wellman-Power Gas
system.
A solution of sodium sulfite scrubs the sulfur dioxide and
forms crystal sodium bisulfite in the reaction
so2
+ Na 2 so3 + H20
--->
2NaHS03
The sodium sulfite is reclaimed by heating.
A 90 percent
sulfur dioxide stream which results is sent to a Claus unit.
This
Claus unit removes the sulfur pollutant and changes it to elemental
sulfur [10:361].
More will be said about this unit later.
The citric process is used for low sulfur loads and can
obtain up to 90 percent sulfur dioxide removal.
The sulfur dioxide
is converted into 99.5 percent pure elemental sulfur.
The reactions
are:
S0 2 (g) + H20(l) ~== HS03 + H+ (citric solution)
H+ + HS03 + H2S
--->
3S(s) + 3H20
The hydrogen sulfide must be either made available from
another process in the plant or else made from the sulfur recovered.
The utility cost for this process are very low, and the
organic acid reagent is nontoxic and biodegradable for minimal
environmental impact [10:364-5].
A newer form of sulfur dioxide removal is called dry
scrubbing.
This is much like wet scrubbing except the water in the
alkali solution is just enough to be evaporated by the flue gas.
The
subsequent particles must now be removed along with the fly ash.
This may be accomplished by using electro static precipitators or bag
houses.
Dry scrubbing is much less expensive than wet scrubbing and
requires less maintenance [10:365-6].
Chapter 3
THE REMOVAL OF CARBON DIOXIDE
Carbon dioxide must be removed in large quantities in the
manufacturing of hydrogen and ammonia as well as in the treating of
natural gas.
The carbon dioxide must be reduced to 2 percent for
pipeline gas, and reduced to 10 parts per million for the synthesis
of ammonia [1:201].
Carbon dioxide is an acid gas and when it is dissolved in an
aqueous solution, it forms carbonic acid, H2co [1:207].
3
Although
there are both physical and chemical absorption, the physical
absorption is insignificant when using alkaline solutions.
Figure 2
is a plot of the first and second dissociation constants of carbonic
acid.
Since the second dissociation constant has such a high pK, the
carbonate ion will only be formed in strongly alkaline solutions.
These solutions cannot be regenerated and are not used as chemical
solvents [1:208-9].
The formation of the bicarbonate ion is very important in the
absorption of carbon dioxide into an alkaline solution.
occur in two ways.
-~
"t"--
This may
The first mechanism has the reaction
HC03 as the kinetic step [1:209].
The rate of this
reaction is r = k[OH-]([C0 ]-[C0 ]) where [C0 ] = [H003]/[0H-]K , and
2
2
2
2
K2 is the thermodynamic constant.
The kinetic value k for dilute
solutions can be represented as log k =13.635 - 2895/T for a range of
273-213 K.
For stronger solutions where the ionic strength affects k,
10
11
5
-:;;
5u
c:
.2
t
~
.~
I.-
~
I
u
'ti
"g
Jl'"'
t-
I
!
i
I
i
iI
10-ll IL ·----~·
I
__J__J_,~~~~~-J---L--~.~~--J-~
2.5
3.0
HlOO.'T,
3.5
K
-1
Figure 2
First and Second Dissociation Constants
of Carbonic Acid
Source: Astarita, Savage, and Basio, p. 209.
12
the following is recommended:
log k/k inf dil = 0.08 Ii.
These lead
to an overall equation fork of log k = 13.635 -2895/T + 0.08 I ~.•
= ""k[C0 2 ]
second mechanism has a rate of r
The
but ./'k is so low that it is
only relevant at pH levels below 8 [1:210].
For the absorption of carbon dioxide into hydroxide
solutions, the main reaction is C0 2 + 20H- = C032-+ H 0. As long as
2
the hydroxide concentration is high enough, the bicarbonate ion
concentration is negligible [1:211].
The partial pressure of carbon
H
dioxide from this reaction is p = -Km
very low being only 6.7x10-8
y
----:z
(l-2y)
.
The H/Km term is
atm at 100 °C and 1 gmol/liter.
Therefore the hydroxide solution cannot be regenerated and can be used
only once.
However, because partial pressure is so low, the hydroxide
solution is capable of removing the final traces of carbon dioxide
which makes it an attractive process after the bulk of the carbon
dioxide has been removed [1:211].
These low partial pressures are at most 10-3
of 10-S
gmol/liter or less.
atm or ai values
This means that the Ioo from
D - m (1-2y )
OH
o
o
are on the order of 105 • Where
a.
~
D represents the diffusivity of each species. This is a fast regime
reaction such that I =fi = Jkm (1-2y )tD.
0
0
I is of the order of 10-100
so mass transfer control is liquid sided [1:212].
The reaction of carbon dioxide with an amine is believed to
be a two step mechanism.
The first reaction is slow and therefore
the rate determining step, co
reaction is RN+HCOO- + RNH
2
--->
+ RNH
--->
RN+Hcoo-.
RNH; + RNCOO-.
The fast
Therefore the
13
~1
overall reaction is first order, r =keF [00 2 ][RNH] [1:213].
Amines can be classified by the parameter P which is an
indication of the stability of the carbamate ion.
The carbamate
jP2- 4y(1-y)
P -
concentration can be written as ""b3 = -------------- , where
2
P = 1 + (1/Kc m) and ranges from 1 to infinity [1:213].
The three main reactions for amines are:
--->
CF:
C02 + 2RNH
BF:
C02 + RNH + H20
RNH 2+ + RNCOO-
--->
RNH 2+ + HC0 3-
RNCOO- + C0 2 + 2H20 --> RNH; + 2HC03
where R is an alcohol depending on the individual amine, and CF, BF,
CR:
and CR will be used to reference the equations [1:213].
For a stable carbamate ion, P = 1 and occurs in primary amine
solutions.
and at y
At y
> 0.5
< 0.5,
the carbamate formation is the main reaction
the CR reaction is dominant.
As can be seen from
figure 3, the vapor pressure of carbon dioxide is minimal at low
temperatures and low y.
This means that the amine can absorb
practically all of the carbon dioxide, to a level less than 100 parts
per million [1:216].
Since as y
< 0.5,
the CF reaction dominates, the rate is r =
1 b2b3
-
~-
K b1
A plot for kCF is given in figure 4.
equation tr
1
y
2
= - ----
K ( 1 - 2y)
[1:221].
Using this data and the
= l/kcpm0 (1-2y 0 ), the reaction time can be calculated.
These times are in the fast reaction regime so that the enhancement
factor can be calculated as I =Jtd/tr =Jkcpm0 (1-2y 0 )tD [1:221].
The general flow diagram for carbon dioxide removal with
•
14
100.0
/
/
/DEA
'/
.!!
c..
/
/
10.0
r:.'
0
"5
~
t
>
0
Ql
~
a.
Ql
1,0
E
::>
,2
::>
:r
"
N
0
u
0. i
0,01
..
~------~-----~~-----~~----~~-----~~~
0
0.2
C0
0.4
2
0.6
0.8
1.0
in solution, :nols CO.,/mol amine or
initial r:.orbonofe
Figure 3
Equilibrium Pressures of co Over Aqueous Aminoalcohol
2
and Potassium Carbonate Solutions at
Absorber Top Conditions (40 °C)
Source: Astarita, Savage, and Basio, p. 217.
0
'
15
'
'
EDA
''
''
' ' DGA
MIPA '~,
10
,,
''
'\
•.,
..
011'·'
·
''
,,
'
'''<", ·..·..·.,,
''~,
'
·..
5
·.'_,
'
''
DEA
Abbrevia~ion
., '
' '
'
'·'
'
-..:' ' ' ·..·.'
'
..~'""'
MEA··.'
', ' ' ' , ·. ••
'''
' ' ' ·..·.
'' '·.
'',,_ -, '~~·~.\.
~thv:enediamlne
EDA
o~-
' ,
~
''·,
-...l.....·---'-~--_;__,.__·,_._]
(i.,finitc di!t:tion)
MEA
DGA
MIPA
OEA
103
.. ,
hydroxyl iori
~
'
mor1oethonoi.cm i ne
~ 1 p' hydfDxyaminooihyfe rfw
·...
rr.onoiscprcponolamine
·..
dil!•horn: :<v..;(.,e
...._L
2.4
2.6
2.8
3.0
3.2
3.4
3.6
Rec iproco! temperotur~, 1000/T, !<-1
Figure 4
Arrhenius Plot of Second-order Reaction Rate Constants
for C0 -Hydroxyl Ion and co -Amine Reactions
2
2
Source: Astarita, Savage, and Basio, p. 219.
16
monoethanolamine, MEA, is shown in figure 5.
MEA is a primary amine
such that the nitrogen atom is connected to two hydrogens and one
alcohol.
The high stability of the carbamate ion in the MEA solution
causes difficulty in stripping and therefore, the reclaimer is used in
the MEA process [1:222].
When P approaches infinity, the carbamate ion becomes
unstable.
This is the case of tertiary amines where the only
reaction to be considered is the bicarbonate formation.
pressure of carbon dioxide is given as
* HKP
p = ---
y
1-y
where Kp is the amine protonation constant and Kc
The equation (1-y)/y
the hydroxide concentration is [OH-]
2
m
Kcl
dissociation constant.
1
is the first
< JmKp/Kw
= Kw/Kp *
The vapor
holds when
(1-y)/y ie.
This holds true since JmKp/Kw is of the order of 1000
for tertiary amines [1:224].
The reaction of the bicarbonate in these tertiary amines is
slow in comparison to the MEA solutions and therefore is not an
attractive process for carbon dioxide removal.
The tertiary amines,
however, because of this fact, are used to selectively remove
impurities which react faster than the carbon dioxide [1:224].
For a moderately stable carbamate ion, P is in the range of
2.
For this case when y
reactions.
A
Using db
< 0.5,
both the BF and the CF are main
A
3
as an increase of the carbamate and db
4
as an
increase in the bicarbonate, the thermodynamic progress ratio can be
found by the following [1:225]:
db3
= --- dy =
dy
1-2y
.; P 2 -4y( 1-y)
A
- dy and db
4
17
Pcrified
Sv.,thesi s Gas
I
·J
Synrne>
_Gm
I
~_
~!ution
Cooler
_t-
~---Ab~
S.,lution Hec:;t
1. ----------------·---~-c~-·~--------------------------·----------~!
~,,n~n:tor
Figure 5
Flow Diagram for a Traditional MEA-C0
2
Scrubbing Unit
Source: Astarita, Savage, and Basio, p. 222.
J
18
This is with carbon dioxide increasing ie.
becomes(~)
= (~~)
BF T
db
The ratio
1
- -
-
----- j[P 2 -4y(1-y)]/(1-2y) - 1
4
This shows that the for y
dy is positive.
[1:225].
"' is the progress of the CR
-db
3
> 0.5,
A
reaction, but that the db 4 is due to the formation of both species.
A
A
Since the BF reaction can be related as db 3 + db 4 , the ratio can be
simplified
/\
as(~~)
=
BF T
-db3
2y - 1
lh;-~-d~:
-jpZ
[1:225].
-4y(1-y)
For carbon dioxide absorbed in carbonate solution, the main
reaction is C02 + H20 + C0 23
non-volatile species [1:68].
2HC0-3 , with C0 23 and HC0 3 as the
The molarity is given as C032- + .5HC0- ,
--->
3
and the saturation as .5Hco;.
now be given as [ co23 ]
The concentrations of these species can
and [HC0- ] = 2my [1:68].
3
The equilibrium condition for the above reaction is K =
2
= m(1-y)
b2 /b 1a
with K being in the 1000 range [1:69].
gives: a
= 4m/K
* y 2 /1-y, and since p*
= Ha
Combining equations
from Henry's Law, the
equilibrium partial pressure is written as p*
* m * y2 /1-y
= 4H/K
This equation relates y and p* well but does not do very well
[1:69].
for molarity and p* •
The data is correlated by
2
p*
= 1.95
X
--T--
109 m" 4 (y _y) exp (-8160) where T is in K, p* is in atms
1
and m is in gmol/liter.
Henry's constant is correlated as log H/H0
0.125m, where H0 is H for water at the same temperature [1:229].
The hydroxide concentration is [OH-]
is the water dissociation constant.
can now be written as r
= kK 1
= K1 *
1-y/2y where K1
The rate of the main reaction
* 1-y/2y *(a-a) where
a = 2y/1-y
1/K2 where K2 is the equilibrium constant for co + OH- ~~ HC03
2
*
=
19
Now the reaction time can be written as tr = 1/kK
1
*
2y/1-y [1:230].
There are three main flow arrangements for carbon dioxide
removal.
See figure 6.
Each has its own use depending on the
amount of carbon dioxide present.
to the absorber and regenerator.
rich/lean heat exchanger.
The first method has a single flow
This method also lacks the common
This method will remove carbon dioxide
down to 5 psia partial pressure [1:241].
The second method has a split stream to the absorber but a
single stream to the regenerator.
Only 30-40 percent of the stream
is cooled by 20-70 °F before entering the absorber.
The equilibrium
back pressure in this smaller stream allows carbon dioxide removal
down to 2 psia partial pressure [1:241].
The third method has split streams to both the absorber and
the regenerator.
The major portion leaves the regenerator at the
top, and the minor stream, 15-40 percent is regenerated further and
then cooled.
Carbon dioxide removal in this method can be as low as
500 parts per million [1:241].
20
Gos Out
__ ....t._
Ab ~Qv•
Acid Gos
,,,
,.
r.r;-
·'
,,
'''
1,'
t'
~~·;
.,,
,,,
,,,
Regenerator
./I,
,_,
~::;,
,,!
.,,.
'It
'•,
.,
,,
Gas In
Rich Solution
Solution
(a)
..
C>
0
V'\
.Jt
g>
v;
Gos In
(b)
Gas Out
Add Gos
..
0>
0
Vi
•1
:.1
··J
Goo'"
·~
(cl
Figure 6
Hot Carbonate C0
2
Removal Flowsheets
Source: Astarita, Savage, and Basio, p. 242.
Q •
Chapter 4
THE REMOVAL OF HYDROGEN SULFIDE
Hydrogen sulfide is the most common sulfur contaminant with
carbonyl sulfide, carbon disulfide and mercaptans generally occurring
at lower levels.
processes.
Hydrogen sulfide must be removed from many
It must be removed to prevent air pollution, to reduce
corrosion, to ensure good health, and to prevent catalytic poisoning.
The hydrogen sulfide must be reduced down to 4 parts per million for
pipeline gas and must be as low as 0.01 part per million to prevent
catalytic poisoning [1:245].
After the hydrogen sulfide has been removed, it must be
disposed of in a safe manner.
The most cases, the hydrogen sulfide
is sent to a Claus plant where it is turned into elemental sulfur.
These Claus plants have been used for eighty years and are still used
today [1:245].
When the gas first enters the Claus plant, it is split into
two streams.
H S + 3/2 0
2
2
One third of the stream undergoes oxidation
= so2
+ H 0 and the remaining stream is reacted with the
2
S0 2 from the first reaction 2H S + S0
2
2
= 3/N
SN + 2H 20
[1:246].
Since hydrogen sulfide is an acid, it will react with all
bases.
This is an instantaneous reaction since the reaction time of
simple proton-transfer reaction is of the order of 10This reaction can be written as H S + B
2
represents a base.
= BH+
13
sec [1:253].
+ HS - , where B
Since this the only reaction to be considered, the
21
22
concentrations can be written as [B]
= m(l-y)
where y is the fractional saturation.
is written as
y
___
Pm
p* = HK
KS1
and [BH+]
=
[HS-]= my
The equilibrium vapor pressure
2
where K is the protonation .
1-y
p
constant of the base Kp = [B][H+]/[BH+], and KSl is first
dissociation constant of hydrogen sulfide [H+][HS-]/[H S] [1:254].
2
The overall hydrogen sulfide vapor pressure can be defined as
Ks = my 2 /(1-y)p* which is inversely proportional to Kp (1:256].
The most common way to remove hydrogen sulfide from a gas
stream is the ethanolamine or Girbotol processes.
In this method,
the hydrogen sulfide and usually carbon dioxide, are absorbed in a
high pressure, low temperature column.
The stream is then sent to
the stripper where it is heated to regenerate the amine solution.
Triethanolamine was first used in the process but monoethanolamine
was found to be more suited as it could absorb more acid gas.
MEA,
however has the disadvantage of forming compounds with carbonyl
sulfide and diethanol urea.
This results in a loss of amine.
Therefore, diethanolamine is used for refinery gases and MEA is used
for natural gas treating [9:96-8].
A process developed by Shell uses a 40 percent solution of
potassium phosphate as the absorber.
This solution is more stable
than the amine solution, and is more selective towards hydrogen
sulfide [9:100].
The production of elemental sulfur from hydrogen sulfide is
the idea behind the oxidation process.
In this process, sodium or
ammonium carbonate solution is used to absorb the hydrogen sulfide.
Small amounts of ferric oxide act as a catalyst in the oxidation of
hydrogen sulfide.
Sodium thioarsenate is used in the Thylox process
with the reactions being
H2S + Na 4As 2s5o2 = Na 4As 2s6o + H20
[9:102-3].
Choosing a process for hydrogen sulfide removal depends on
the initial concentration of H2S, the degree of cleaning required,
and the presence of other contaminants.
Chapter 5
THE REMOVAL OF HYDROGEN SULFIDE AND CARBON DIOXIDE
There are two cases to consider when carbon dioxide and
hydrogen sulfide are both in the gas stream.
Case one requires that
both contaminants be significantly lowered, and the second case
requires that the hydrogen sulfide be selectively removed.
The
selective removal of hydrogen sulfide is necessary to provide high
concentrations of hydrogen sulfide to a Claus plant [1:266].
The hydrogen sulfide must be reduced to four parts per
million and the carbon dioxide is reduced anywhere from 2 percent to
100 parts per million, depending on the requirements.
The removal of
carbon dioxide to 100 parts per million is required in the
manufacture of liquid natural gas, and the less stringent requirement
of 2 percent is common for natural gas transportation in pipelines
[1:267].
One such method of simultaneous removal uses an inorganic
salt of a strong base and a weak acid.
as MX.
This salt will be indicated
The three basic reactions are:
C02 + H20 + X- = Hco; + HX
H2S + X- = HS- + HX
Hco; +
x-
=co~-+
HX
[1:284].
The equilibrium vapor pressures for the first two reactions
can be written as p* =Hm
K
X
Kc1
y(y + y')
--------1 - y - y'
24
and p*,
K y'(y +y')
= H'm
Ks1
--------1 - y - y'
leading top */p* ' = H/H' * K y/y' where y is the concentration of
cs
carbon dioxide and y' is the concentration of hydrogen sulfide
[1:285].
The hydroxyl ion concentration is written as [OH-]
(1-y-y')/y+y' assuming y andy' exist [1:286].
= Kw/KX *
This concentration is
less than the concentration in the carbonate solution ie.
m=
[ 0H- ] +2[C032- ] since Kx is larger than the second dissociation
constant for carbonic acid.
This means that the carbon dioxide
reaction mentioned above reacts at best, in the fast regime.
And the
simple proton transfer of the hydrogen sulfide reaction is
instantaneous as mentioned before [1:286].
Another method of removal involves the use of carbonate
solutions.
In this type of solution, the second dissociation
constant of hydrogen sulfide is so low that practically no S2- is
formed.
For the reaction H s + CO 22
= HS -
+ HC0 - , the
3
concentrations of carbonate and bicarbonate changes but their sum
Therefore, m(l+y)= [C0 23 ] + [HC0 3 ], where m is the
molarity given as m = [C023 ] + 0.5[HC0 3 ] + 0.5[HS ] [1:85].
stays the same.
The equilibrium vapor pressures can be written as:
p* =
Hm (2y +y')2
K
--------1-y-y'
Letting K
= ---
and
*,
p
K'
and K r =
H'm (2y + y' )y'
=
K'
---------1-y-y'
sl then
Kc2
*
*-;
p
H
=- K
P
H'
cs
[ 1:287].
This means that carbonate solutions will not remove hydrogen
sulfide unless it also removes carbon dioxide.
This is true
because if the solution absorbed only hydrogen sulfide, then p*/p*'
would increase and this would create a strong driving force for
26
desorption in the stripper, but this would not be the case unless
carbon dioxide is also absorbed [1:288].
Amine solutions are another way of removing carbon dioxide
and hydrogen sulfide from gas streams.
The stability of the
carbamate ion plays an important role as to determining the
selectivity; this was discussed in the section on carbon dioxide
removal [1:289].
The equilibrium vapor pressures can be determined using the
following equations.
+ 2RNH = RNCOO- + RNH+
[1:71]
and
RNH = RNH; + HS- [1:87], the resulting vapor pressures are
*
H y(y + y')
H' y'(y + y')
p =----------and p*' =
K'
m
K 1 + 1/Kcm
[1:291].
and since K'/K = K /K
then the ratio can be written as
sc c'
H
y
= -- K
H'
1
sc y ' 1 + K m
c
[1:290].
The selectivity is determined by the stability of the
carbamate ion as stated before and can be restated briefly.
The more
the amine behaves as a tertiary amine, the higher is its selectivity
towards hydrogen sulfide [1:292].
p
Chapter 6
ABSORPTION
Absorption is a process involving the transfer of molecules
from the gas state into the liquid state because of a concentration
gradient between the two phases.
mass transfer.
This is basically unidirectional
The soluble component in the gas phase is called the
solute and it is picked up by the absorbate of the liquid layer.
This is one of the most advanced techniques for separation of gases,
and can be used to remove pollutants from a gas mixture [3:242].
Figure 7 shows a simple layout for an absorption plant and
consists of an absorber, a regenerator, and the necessary auxiliary
equipment such as heat exchangers, pumps and holding vessels.
The
raw gas enters the absorber at the bottom and mixes with the liquid
coming down.
The pure gas exits at the top, and the now contaminated
liquid exits at the bottom.
The liquid is sent to the regenerator
where the absorbed gas component is stripped from the liquid and sent
to another unit.
absorber.
energy.
The regenerated liquid is then recycled back to the
There are usually heat exchangers in this process to save
The cooler, contaminated liquid is heated by the freshly
regenerated liquid; this lowers the temperature of the absorbing
liquid to enhance absorption.
heated to enhance stripping.
The contaminated liquid in turn is
The absorption also needs higher
pressures than the regenerator, so pumps are used to boost the
pressure as well as transport the fluids.
27
This extra pressure going
•
28
Treated
t
Go•
Out
I
I
Acid Gas lo
r - - - - - - - " 1 Sulfur Plant
F
•
••
.I
Raw Gos In--
A - Alosorber
W - Water Wa;h
llf -- HrdrocorLun F lo>h
~
- A, id Gm Flush
R - kr,genercncr
H -
l.0w -Level
(Woste) Heal
l<~i>c.iler
Figure 7
Idealized Solvent Absorption Acid Gas Removal Process
Source: Astarita, Savage, and Basio, p. 12.
29
to the regenerator can be reclaimed by using hydraulic turbines
[3:244].
Since, in an absorber, the gas must interface with the
liquid, the larger the interface area, the more efficient the
absorber will be.
The method used to generate these areas is an
important property of the absorber [3:249].
There are three groups of
absorbers depending on the method used to generate the interfacial
area:
a) generation of liquid films
b) generation of jets
c) generation of bubbles and drops
therefore the names of the absorbers are:
a) film absorbers
b) jet absorbers
c) bubble and drop absorbers [3:250].
There are three types of film absorbers:
a) tube bundle columns
b) packed columns
c) fluidized packing columns [3:250].
The important properties of a tube bundle column are:
high
gas flow rate, low pressure drop, high mass transfer rates, simple
elements, and effective liquid distribution [3:251].
See figure 8.
The packed column is packed with small cylindrical or saddle
shaped elements.
The packing material aims to ensure complete
wetting, by avoiding dead spaces and recirculation.
It is also
favorable to obtain high mass transfer and low pressure drops.
is done by using larger elements [3:252].
See figure 9.
This
30
purified
gas exit
liquid
distributor
tube
bundle
1st stage
Figure 8
Example of a Two-Stage Tube Bundle Column
Source: Brauer and Varma, p. 251.
31
~purified gas
outlet
packing''packing
packing,,_
polluted gas
inlet
absorbate exit
Figure 9
Design of a Two-Stage Packed Column
Source: Brauer and Varma, p. 252.
32
The difference between the fluidized packing column and the
packed column is that the fluidized column is operated at a much
higher gas flow rate so that the particles are fluidized.
Large
spherical particles with very low density are generally used.
These
particles move very quickly and have many collisions which prevent
fouling, therefore this makes for a self-cleaning absorber [3:254].
See figure 10.
The jet absorber introduces a turbulent jet stream of liquid
absorbent at the top of the absorber.
As the distance from the
nozzle increases, the liquid starts to form drops.
The gas is also
introduced at the top of the absorber so an intensive mixing takes
place.
This system is suitable for situations which allow for a very
slight pressure drop [3:255].
See figure 11.
The interfacial area in bubble and drop absorbers is produced
by dispersing one phase in the other.
phase can be dispersed.
Either the gas or the liquid
There are three types of gas dispersed
absorbers:
a) plate columns
b) bubble columns
c) rotating disk columns [3:256].
The plate column is a cylindrical vessel with plates spaced
about 0.5 meter apart; the number of these plates depends on the mass
transfer rate required.
The contact between gas and liquid takes
place in the pool of liquid which forms on these plates.
important designs of plates are:
a) bubble cap
b) sieve plates
The most
33
mist
s.pt~rator
absorbMt •n!~l
wpftHf fluid1 zed
packing
--gas
·
distrib:.lfion
plate
(1=
e;b;sorbat cu tiel
.i..
,_ aOsorbMI
I
_l
c:-v' - - - -
----·<;}-----~---'
Figure 10
Design of a One-Stage Fluidized Packing Column
Source: Brauer and Varma, p. 254.
abscrbent
pol(ut~gas
ml..t
purifi~
=)
gas outlet
t
turbu~nt
1
....
liq111d jet 1; '
I
r
I
gas/liquid
sep..vating-~_
vessel
absorbent
absorbat
<=
-----L---.L--{
Figure 11
Design of a Jet Absorber
Source: Brauer and Varma, p. 256.
35
c) valve plates
(see figure 12) [3:256-7].
Hundreds of bubble caps may be on one tray with an average
bubble cap diameter of 10 centimeters.
plate is on the order of 10 millimeters.
The hole size on the sieve
The valves may be thought
of as bubble caps which adjust to the gas flow rate to give the best
dispersion [3:257].
The plate columns have the advantage of variable operating
conditions but are very sensitive to foaming [3:257].
A bubble gas column is also a gas dispersing column.
The
bubbles are generated only once by means of a suitable device close
to the entrance of the raw gas.
These bubbles are then left to rise
very slowly to the top of the absorber.
This low gas flow rate makes
this absorber type impractical for air pollution control, where high
gas flows are required [3:258-9].
The rotating disk column, see figure 13, is a multi-stage
absorber.
Each stage consists of two conical sections attached by a
cylinder.
The rotating disk is in the cylindrical section and
disperses the gas in the liquid.
This type of absorber is used for
treatment of small gas flow rates, but has the highest efficiency of
all the absorbers [3:259].
The liquid dispersed absorbers are:
a) Venturi columns
(figure 14)
b) nozzle spray columns
(figure 15)
c) rotating disk spray columns
(figure 16)
[3:260].
In the Venturi column the absorbent enters at the throat of
the Venturi.
This is the section where the gas velocity is at its
greatest thus making a highly efficient absorber.
This high
36
or"3n~,.nl #II
ht1i~J
in plOt•
orra"~·,.,.,
. ~., on
ol
p/O,.
Figure 12
Three Different Plate Arrangements
Source: Brauer and Varma, pp. 258-9.
37
','
f
5epar<Jtion
t<.--1-+--'-'.,>, zone
--,~action
zone
Figure 13
Design of a Multistage Rotating Disk Column
Source: Brauer and Varma, p. 260.
38
pollut•d gas iniel
canfusor. __
liquid
diSPif"Sion-
zone
l lr
-n
0
.o:=::J--a_b_s_or.b_•_n_t_i_nl_•_t _ _ _ _~
diffusor - ___ .J-,
~
'
:' b<-I
\_~'
_::liqui~
purif1ed gas
outlet
s=r
fresh absorbent inlet
Figure 14
Design of a Venturi Column
Source: Brauer and Varma, p. 261.
39
purifi~d
gas
outl~t
!=Y?oh-:"':"<"-.i.--~mist s~parator
f~sh absorb~nt ln~t
s«:ond spray z
absorlJat~
absorbent cycie
Figure 15
Design of a Nozzle Spray Column
Source: Brauer and Varma, p. 262.
40
"'--r rotating
'
disk
dispersion elements
~-~~­
absorbate
cycle
le
fresh absorbent inlet
Figure 16
Design of a Rotating Disk Spray Column
Source: Brauer and Varma, p. 262.
41
efficiency however is offset by the energy needed [3:260].
In a nozzle spray column, a liquid mist is formed by pressure
nozzles at various levels in the column.
The gas enters at the
bottom, flows through a special gas distributor, and countercurrently
mixes with the liquid.
The pressure drop in this system is very low
and is caused by the gas distributor.
This gas distributor, however,
creates an even flow of gas which increases the efficiency of the
column [3:260].
The rotating disk sprayer uses rotating disks to throw the
liquid horizontally.
The gas however; flows in the axial direction.
This absorber has the advantage of low pressure drop and is easily
adaptable for higher flow rates or higher pollution concentrations.
The capacity is simply increased by increasing the amount of liquid
used [3:261].
Different processes demand different absorber
characteristics.
Therefore, there is no one universal absorber.
Each absorber must have its own design dependent on the process.
quality of absorbers is described by the following three sections:
a) General process criteria:
species of pollutant
pollution concentration
gas flow rate
absorbent
physical or physical absorption
cocurrent or countercurrent flow
adaptability or rate changes
capability of dust removal
The
42
danger of fouling or plugging
danger of foaming
danger of corrosion
b) Mass transfer and fluid dynamics criteria:
size of interfacial area generated
renewal of interface
mass transfer coefficient
concentration difference between gas and liquid
minimum liquid flow rate
pressure drop or energy required for absorption
volumetric efficiency
c) Equipment criteria:
size of absorber
rationality of design
construction materials
investment cost
maintenance cost
emission proof design [3:262-3].
The chemical and physical characteristics of the pollutant
form the basis for all considerations.
The gas flow rate is important
to determine the size of the vessel; see figure 17.
Practical
operation and performance of an absorber depends on the adaptability
of the absorber to change with flow rates and pollution
concentrations.
Size must also be taken into account and a rational
design will always be the least costly design. [3:263-4].
See figures
18 and 19.
Although chemical absorption dominates in air pollution
4.)
Figure 17
Application of a Few Absorbers as a Function of
Gas Flow Rate and Height-to-Diameter Ratio
Source: Brauer and Varma, p. 264.
44
rongt! of
clo.s~ty
tu~
po r:k«<
bundt~
column
1
I;:J- 2
l
s
10 1 2
s
10-'
•l•,ln,,l_,_l,__
:L)j.___---'
-,--:.1-
1C11,_·_ _.......__ ___._.
10- 5
ur'
w-3
ro-;
to-'
10°
btlbble,portic!e, and tube diameter d [m]
Figure 18
Dependence of Specific Interfacial Area a on
Bubble, Particle and Tube Diameter d
Source: Brauer and Varma, p. 265.
45
Figure 19
Specific Interfacial Area a as a Function of Volumetric
Energy Demand for Various Types of Absorbers
Source: Brauer and Varma, p. 265.
I
'
control, the physical equilibrium conditions are still important.
Absorption in carried out under the following conditions:
a) the critical point and the dew point of the gaseous pollutant are
below the absorption conditions
b) the vaporization point of the liquid is below absorption conditions
[3:266].
The relationship between the concentration of the pollutant
in the liquid and gas phases can be represented by:
PA
= VAXAPA,s'
where pA is the partial pressure of the pollutant and is a
function of the coefficient of activity, YA, molar fraction of the
pollutant in the liquid, xA, and the saturation pressure of the
pollutant, pA ,s [3:266].
Henry's law can be stated as:
pA = HAxA, where H must be
found experimentally, and xA is defined by:
XA
= --------------- =
CA + CB + . • •
CA
c
where cA and cB [kmol/m3 ] are the molar densities of components A
and B of the liquid.
The sum of all molar fractions must be equal
n
=1
to 1 ie. [x
[3:266].
i=l
If ideal behavior of the liquid is assumed, Raoult's law
applies; therefore, YA =1, and the equilibrium constant, KA,PHY may
be defined as:
y APA,s
= -----
p
=
HA
p
I
[3:267].
Since the pollution concentration is relatively low in the
gas mixture, Henry's law applies.
on figure 20.
Some Henry coefficients are shown
High values of HA indicate poor solubility of the
47
~::~~~~~~~~;:~::=:~~--~-~-~-~~~~
6
l
..___
"'
~""
~
~
.
:t
i
:Q
6
.
::.::
0
u
~ 2
j.L-...,...,----------;-- · - - r - · - - - - l
~
.I
:
10' f.---' _
__;__----r
t-
-------'-~--,---.L
I
-+---~:--+-~.--~
;I
!
I
!
'
I
:
•
r
;
.
[E;-~: ~-.
2
JO'
270
l
j
•
.
.
290
:
.
~
.
,
.;.
HCI
___j_
[jll::i!
i_
I
:
i
I
!
~
!
·--~·
310
!
I
_ _,__ _
330
350
370
ttm~.·alu.~ T J K}
Figure 20
Henry Coefficient HA for the Solution of Various
Gases in Water as a Function of Temperature
Source: Brauer and Varma, p. 267.
48
pollutant in the liquid and low H values indicate good solubility.
Therefore, the higher the H value, the more liquid is necessary to
ensure absorption [3:267].
In a chemical absorption process, the pollutant reacts with
the liquid to produce a substance which is not a pollutant:
)(,
A+ zB .,=.= sP
"~
where s and z are stoichiometric coefficients and k
1
and k
2
are
rate constants which are determined experimentally.
increase with rising temperature and pressure [3:268].
The chemical equilibrium constant KA,CH is defined as:
CA CB
where cA, cB, and cp are concentrations of A, B, and P respectively.
For chemical absorption it is best if the reaction is product
oriented; this makes the reaction very fast and diffusioncontrolled [3:268].
So, there are some differences between chemical and physical
absorption.
Chemical absorption is improved by increasing
temperature and pressure, while physical absorption is improved by
decreasing the temperature.
In physical absorption, the pollutant
will never be able to be significantly reduced, which is desirable.
Therefore chemical absorption is most often used for air pollution
control [3:268].
From the kinetic theory of gases, the diffusivity in the gas
phase can be calculated for a single species as DG = 1/3 >-. u where A
is the mean free path and
u
is the average velocity of the molecules.
49
The mean free path is calculated from the movement of the molecules
without intermolecular forces and can be written as A= 1//2 Nv~
2
where N is the number of molecules per unit volume andu is the
The average velocity can be written as u =
diameter of a molecule.
JBRT/~ M where T is the absolute temperature, R is the universal
gas constant, and M is the molecular weight [3:63].
The diffusion of two different species can now be written as
1 NA~AuA + NB~BuB
DAB = - ---------------3
NA + NB
where NA and NB is the number of molecules of each species [3:64].
The above equation can be simplified by using the mean
molecular diameter crAB' and can be written as DAB
1
;~-N(~~)2
(uA + uB)
= __
3
27~~--}(~+ 1 )
2
3 2
N(u
AB
)
M
=
[3:64].
M
A
B
Since the normal boiling points VA and VB are proportional to
the cube of the collision diameteruAB' the above equation can be
written as DAB=
bp(VI73~;~;I73)Z ~(~ ~)
+
The constant b can be calculated from kinetic theory but more
accurate diffusivities are obtained if empirical values are used
[3: 64].
The diffusivities in the liquid phase can be calculated by
using the equation F
=-
kT/c
*
dc/dx where k is Boltzmann's constant.
This equation is based on Einstein's suggestion that an osmotic force
acts on the molecules in the direction of decreasing solute
concentration, c [3:65].
50
Stoke's law, which relates the motion of molecules, can be
used to represent the molecular resistance to motion.
written as F
= 3'ii pq- u
This can be
where cr is the diameter of the molecule, p is
the viscosity and u is the velocity of the molecule.
can now be written as u
= -kT/3'IT
cpq-
*
dc/dx.
This velocity
The rate of diffusion
NA related to the concentration and the molecular velocity as NA
uc
= -kT/3'Ti pq- *
=
dc/dx [3:65].
The diffusivity can be calculate from Fick's law, NA =
-D
*
dcA/dx, and is given as D1
= kT/3'ii p.o-
[3:66].
When a system is in equilibrium, the molal concentrations of
two species is constant and can be written as cA + cB
= constant.
Differentiating this equation with respect to x is dcA/dx
= -dcB/dx.
Where xis the direction of bulk transport and diffusion [3:66].
The amount of components A and B picked up by the liquid is
proportional to the partial pressures, p, of each component.
This
means that the total flow for a species is a sum of the bulk flow
fraction, NApA/P and the fraction absorbed.
This can be written as
Using the ideal gas law, the concentration of a species can
be written in terms of its partial pressure by cB
= pB/RT,
and since
for a two component gas, pA + pB =1, the total flow can be written as
NA = NA (1- PPB) + RT~~ . ~:B
dx
f
X
This leads to the integral NA dx
0
which after integration yields NA
[9:63-6].
51
PB2 - PBl
Defining the logarithmic mean partial pressure as pBM = - - - - - ln PB2/pBl
The pollutant transfer may be described by relating the
equilibrium curve and operating curve, see figure 21.
.
.
flow rates will be Ng and N both in [kmol/s].
1
The molar
The solute
concentration is given as y and is defined as:
y
= ------------- =
PA + PB + ...
where pA and pB are the partial pressures of the gas components, and
P is the total pressure.
Using this notation, the molar flow rate of
the pollutant into the absorber is given as Ngyb and that leaving
the absorber is given as N y
g a
NA
both in [kmol/s] [3:268-9].
The amount of pollutant transferred to the liquid is:
.
= Ng(yb
-ya)
= N1 (xb-
xa). They is plotted against x in figure
21, this is called the operating curve.
This follows from the mass
balance, and:
Nl
N
g
= ------- =
X
b
tgor where
or
is the angle between the horizontal and
- X
a
oc [3: 269].
The gas concentrations yb and ya as well as the liquid inlet
concentrations x may be given.
a
Yb -ya
~= X a
+ -.--;-N1 /Ng
Yb - Ya
= X + -----a
tgor
This leads to:
[3 :269].]
This shows that xb increases with decreasing molar liquid flow
.
rate N1 •
This lower liquid flow rate will decrease the pumping cost
52
-----·----
molar fraction in liqutd x,
Kp
Figure 21
Limiting Mass Transfer Conditions in Absorption
Source: Brauer and Varma, p. 271.
(~;', ,
0
53
but will increase the size of the column.
The equilibrium curve,
which is also shown in figure 21, shows the concentration at the
interface, Yp and
xp•
This curve must be found experimentally.
For
low concentrations, Yp and xp can be shown by Henry's law:
[3:270].
In the general case however, the local concentration is given
by:
Yp /xp
curve.
= m,
where m is the local gradient of the equilibrium
Again, this is showing that a low H is better for absorption
and that absorption is enhanced by high pressures and low temperatures
[3:270].
Using overall coefficients Kg and K1 , a relation appears such
that 1/Kg
= 1/kg +
= kg ,
small Kg
H/k 1 and 1/K1
= 1/k1 +
1/Hkg.
Therefore, if H is
and the absorption is gas film controlled.
= k1
large, then K1
If H is
and the absorption is liquid film controlled
[ 9:69].
The ratio of slopes of operating curves to equilibrium curves
is called the absorption factor¢ :
~ = ~!-~Ng = N 1 _~~~
HA /P
m
should be greater than unity and is usually between 1.3 and
1.5; increasing
~
by increasing flow rate will however, increase
pumping cost [3:270].
When the liquid enters the absorber, x for physical
a
absorption will always be non-zero because this liquid has been
recirculated.
zero.
Therefore, y
a
in the gas will also be greater than
This is why physical absorption is not enough to remove all
54
the pollutant.
Theoretically, the gas concentration can be reduced
to equilibrium concentration:
(ya )min
--->
ya,p with y a,p
--->
xaHA /P [3:271].
For such an extreme condition, the operating curve (OC) on figure 21
and the equilibrium curve (EC) coincide.
force ie.
At this point, the driving
concentration gradients, become zero.
column must be infinitely tall.
This says that the
Therefore, the inlet concentration x
a
must be greater that (x )min [3:271].
a
Another limiting factor is (yb)min=yb,p' figure 21b.
Because
the driving concentrations are so low, mass transfer does not occur.
Therefore, absorption occurs only when yb
< (yb)min = Yb,p"
The slope
of the operating curve should be equal to or greater than the slope
of the equilibrium curve with ya
¢.
For
~
. .
> 1, N1/Ng
> (ya)min.
This is a restatement of
increases when H increases.
Setting
~
to the
minimum value of 1 gives the minimum liquid flow rate for physical
absorption:
.
.
N1 ,m1n
. = NgHA/P
[3: 271].
So only minimum flow rates are required for good absorption.
~hemical
For
absorption, the stoichiometric relation between the pollutant
and the reactant must be observed.
depends on this ratio:
N "'-'yNg
1
Therefore, the minimum flow rates
[3:272].
The parameters of mass transfer include interfacial area,
mass transfer coefficient and concentration driving force.
the general equations of mass transfer is:
.
One of
-
NA = nAAP
where nA is the molar flux density and Ap is the interfacial area.
Local values for the molar flux density n are given for the gas phase
as:
55
$
[3:273]
$ g andB
1
are the local mass transfer coefficients for the gas
phase and liquid phase, T is the absolute temperature and R is the
universal gas constant.
The following can be obtained for the gas
phase:
p
nAg
= j3 g
TR (y
- Yp)
[3:273-4].
and for the liquid phase as:
nAl
=f91c(xp
- x)
The following must apply to the solute leaving the gas phase
and entering the liquid phase:
}31
c
/3 g
p
TR=
y - Yp
nAg = nAl = nA; this leads to
= tgY
Using figure 21 and the above equations, mass transfer from the gas
phase to the liquid phase can be examined.
In the direction of the molar flux, NA' of the solute,
concentrations in the gas phase changes to Yp and in the liquid layer
to x.
In the interface, the change is from Yp to xp.
in a phase shows the resistance to mass transfer.
this resistance is in both phases.
0°<
Y<
The angle
Y
This difference
For figure 22a,
is between the limits
90° and 0°< tgY <oo , OC is the operating curve, EC is the
equilibrium curve and Pl is the operating point of the absorber.
angle
The
derived above can be used to draw a line through Pl so that
it intersects EC at P2 which has the concentrations Xp and Yp in the
interface.
Therefore, with the local driving concentrations known,
the local mass flux density can be calculated.
Figures 22b and 22c
•
56
! M TR
I
II ;,. gas1:1.1B.phase IItin bethMTR
phases !in liquid phasej
-
! liquid I ga.s i liquid I gas
.Jr.
I
I
I
i
;
I
A
I
'
; l1igh solubility I
1
,
I
I
oC:./
~
/
/
.
2
~
, ;
l (y->p;
~C
I
:.·
y~A./~
/
. / '
! / Yp_,-:
~
j/:
'
!
--1Xp
i
:
I
X Xp
\
I
A
I
; low solubility
!l
,
oy
o~/
Ff_,/ f-C!
I7
/
P.,,
4!
I
X
Xp
i
l
I
1(xp-xJ-(xp-x)max':
1-G
!
"iY-J_6l,,uxi
I!rt-JT~
.
b
.'VA
Ypi
x _ soluttt
~~
I'
i
medium
solubility
: (x -xJ----D
p
...
I
I
I
!soiute:
-
,.
y
X
1 solute
~-N,
i
I
~as
y
~,:
X
-!
y
liquid
a
c
Figure 22
Mass Transfer Resistance (MTR) in Absorption Columns
Source: Brauer and Varma, p. 274.
57
show the two special cases for absorptive mass transfer.
22b, (xp- x)
--->
0.
For figure
The resistance is concentrated in the gas
phase; this occurs with a low Henry constant.
With respect to point
Pl, the concentrations differences in the gas phase reach a maximum
such that tgY
--->co. In the second case, figure 22c, the resistance
is in the liquid phase.
This is for high Henry constant such that tgY
---> 0 [3:274-5].
Calculation of the local molar flux density is done by the
following:
a) draw a diagram with EC and OC
b) plot Pl on OC
c) calculate
..81
d) calculate 8
g
e) calculate Y
f) draw a line thru Pl at angle Y
g) determine point P2 ie. the intersection at EC
[3:275].
h) calculate nA
The mass transfer conditions can be related to the size of
.
the equipment using the molar flux NA per unit volume of the absorber
vc :
.
NA
vc
=
Ap
vc
s
p
g TR
(y - yp)
It can be shown that:
Ap/Vc~l/d
bubble or a liquid drop.
substituting yields:
const P
NA
-- = ~I+m- TR (y - yp)
vc
where d is the diameter of a gas
It can also be shown that:
/3g ~ 1/dm
58
.
This shows that d strongly influences NA/Vc' and that the diameter
of the bubbles or drops should be as small as possible.
It is
fairly easy to generate small particles but it is difficult to
evenly distribute them [3:275-6].
The mass transfer coefficient depends on the type of phase
distribution:
a) particle systems that consist of
either bubbles in a continuous liquid or drops in a continuous gas
b) film systems that consist of a liquid film in a gas stream [3:276].
In film systems, the mass transfer is a steady state process,
but in the particle system the mass flux and the mass transfer
coefficient depend on time.
For this process, the mass transfer at
time= 0, independent of the mass transfer resistance:
f-3
~jD/t
t~O
This states that the mass transfer coefficient is proportional to the
square of the diffusion coefficient D, and as t --> O,f-3 -->oo.
Since for t
-->
0, mass transfer does not depend on convection,
does not depend on particle size.
This all amounts to saying that an
unsteady-state absorber will be more efficient [3:277].
Since the mass transfer resistance is reciprocal oft1, ie.
R = d/D, where d is the characteristic length of the considered phase,
reducing d will lower the resistance.
These resistances can be
summed as:
d
dl
-~ + -Dg
D1
which should be as small as possible.
So the particles of the
highly resistive phase should be made as small as possible [3:278].
Chapter 7
SELECTION OF A PROCESS
The selection of the optimum gas treating process is a
difficult problem which involves both technical and economic
reasoning.
There is no optimum gas treating process for all
applications [1:357].
The selection of the solvent used in gas purification is
based upon an economic analysis of the gas treating process and the
related recovering process.
If a Claus unit is to be used for sulfur
recovery, the acid gas to the unit must be at least 50 percent
hydrogen sulfide for the unit to operate efficiently.
There are
other processes which require less percentage hydrogen sulfide, but
these methods are very expensive.
Therefore, for economic reasons, a
high percentage of hydrogen sulfide should be sent to the Claus unit
in order to remove as much sulfur as possible [1:357].
The controlling factors in choosing an acid gas purification
process are the raw feed gas pressure, the raw feed composition, and
the required treated gas purity.
Figure 23 shows an outline for
choosing a process [1:371].
The circulation rate of the liquid absorber is dependent upon
the partial pressure of the acid gas in the raw feed.
The minimum
amount of absorber should absorb all the acid gas and still provide a
driving force for absorption at the bottom of the absorber tower
[ 1:371].
59
60
Treated
Raw Feed Gas
Gao
CompcHitior.,
Purity
Pressure, etc.
Limitations
oh
sulfur
/
Disposal
Hydrocarbon
Solct>i!it;
~~
Cono---Jsicn
I
Mau T,ansfer
~"-------1
I•
_---
Waste-heat
lntegr>Jtion
Degrcdotion
Trace Impurities
Removal
Preliminary
Select1or.
I
l
__j
I
Sele<:.tion
Figure 23
Selection Methodology for Acid Gas Removal Solvent
Source: Astarita, Savage, and Basio, p. 372.
61
The amount of acid gas in the product stream will determine
the amount of regeneration required.
In order to have a driving
force at the top of the absorber tower, the partial pressure of the
acid gas in the product must be higher than that in the regenerated
solvent [1:372].
Figures 24, 25, 26 and 27 have been developed over the
years to show which process is the optimum depending on conditions.
These graphs are based on existing process selection and may not be
the best if there are other conditions to consider [1:373].
Carbon dioxide removal with no hydrogen sulfide present
occurs in the purification of ammonia synthesis gas made by the steam
reforming of natural gas.
Carbon dioxide can be present anywhere
from 3 percent to 65 percent.
With the absorber column operating at
pressures of up to 1000 psi, the inlet carbon dioxide can have a
partial pressure of 650 psi.
The required cleaning will reduce the
carbon dioxide to 0.1 - 2.0 percent [1:373].
Using figure 24, for low carbon dioxide partial pressures,
amines are preferred.
For higher partial pressures, 75 - 100 psi, .
promoted hot carbonate solution would work.
For high partial
pressures, above 100 psi, physical solvents should be considered
[1:374].
Aqueous amine solvents will reduce carbon dioxide to 0.005
psi partial pressure, carbonate solutions to 0.2 psi and physical
solvents to 1.0 - 3.0 psi.
Therefore for a high partial pressure of
inlet carbon dioxide with a requirement to be cleaned to very low
levels, it might be best to combine a physical solvent process with a
carbonate process [1:374].
Phyoocal
Sot.,.f'lt
Plus
Amine
i
i
1.£
....
8
0
!
iis.
:;
...~
i
4r--
AQu.,..s
am.rw
i
2f-
'
I
1L0.1
I
z
4
'
'
i
6
8 1.0
2
Partoal preguno of
4
6
8 10.0
co 2 in prnduct·psi
Figure 24
Range of Gas Treating Process Application for C0 Removal
2
with no H2s or Other Acidic Impurities Present
Source: Astarita, Savage, and Basio, p. 207.
63
Physical Solvenn
(witnCiaus plant)
¥-,oo
.
-o
.£.
Aq ... ous or organic 110lvent 10lutiora of
primary or wcondory ominoaicohols (witn Claus
plant)
0.1
Partial p<e,.ure of H S in product..,si
2
Figure 25
Range of Gas Treating Process Application for H2S Removal
with no C0 2 or Other Impurities Presen~
Source: Astarita, Savage, and Basio, p. 249.
•
1,~
I
i
6
!
4
'1-.
6
0
4
~+
2
-;
..
'f
I'
I
!
l
i
!
l
I
I
!
I
:
...
~
2
0.1
i
I
I
I
l
i
I
i
I
I
!
I
I
i
i
i
I
I
2
4
a a
I
i
I
I
I
a --j"
6
I
I
i
L
!
I
!
I
I
I
I
1
!
i
I
l
!
-t-l-l
I
I
!
i I !
!
JI
_l
1.0
I
i
I
2
Partial pressure of
i
!
'
11
II
~
!
I
! I
I i
I i
l
i
..
I
i
I
II
I
i
I
i
I
I
i
I
I
••
I
II
'
..
! !
!
I
I
i
I'
i
I
I
I
l
I
!
I
iTI
I
I/'
' Iid~-~ 'i
! II
I I
I
I
t
I
!
10
4
I I I
I!
:
i
I
I
.
l
I I ii
i
I
I
hot ca:bonate or aqueous
! or Promoud
orqanic solvent solutions of P[imary or
aminoalcoho s
I
II I secondary
I -W- I
I!
I
I
l
f
II-
l
I
:I:
a
I
I
8
II
Solv.,u
I
I
100
I
Ph~ical
I
I
'i.
i
;
I
l II
i
2
1
.5
I
: i
I
I
!
I
i
·.$'.~
~ ~
I
!
I I j
I
i
.
. ~~..~~·~
ql
I I
II ~·~o~e(6
j
!;
4
s
a
H~
in product psi
I
I
!
I
I
1
",
,
10
Figure 26
Range of Gas Treating Process Application for Simultaneous H S
2
Plus C02 Removal with no Other Impurities Present
Source: Astarita, Savage, and Basio, p. 280.
65
1000
800
600
•oo
-
I
--
iI
I
i
i
PHYSICAL
SOLVENTS
200
(;)
I
80
;:
60
'
40
,..
-
_,
~
"'
::'l
...
...
..J
<(
<
"-
-·
.~
6
0.1
I
i
'
'·•
I
I
--r-:
I .
I
AQUEOUS OR ORGANIC SOLVENT SOLUTIONS
OF TERTIARY AMI~<OALCOHOLS OR WEAK
SECONDARY AMINOALCOHOLS
I
~:
~
I
I
- '_L_
' .-
~-.,:.
~"'"""
4
i
I
I!
I
j
---·-"'-··+i i i
,
I
OXiDANT
SOLUTIONS
2
.
r--r-+·
~
;.
!
.._,.----t--+-+·
i
----+----+--+-4-+----+-----rl~
i !
-
:~
I
I i
!
100
!:)
...."'"'
i '
I
in
<(
II
~--r-:-;
I
f
i
----~--~r-~-+-+----+-----+--+~1
! l
:~
!
:
:
I
--'---..;-..L.....l..
:
6
8 1.0
2
4
6
8 1C.O
i
I
I
I
J
2
4
.
1
6
8 100
H2S LEVEL IN i'ROOUCT GA.S, PPM
Figure 27
Range of Gas Treating Process Application for Selective H S
2
Removal in the Presence of C0
2
Source: Astarita, Savage, and Basio, p. 281.
66
Although in some instances, the use of a physical solvent
would appear to be optimal, hot carbonate solution is used.
This is
because the physical solvent will remove some of the hydrocarbons
from the feed gas and substitute it with carbon dioxide.
This
requires that another unit be installed to recover this absorbed
hydrocarbon gas [1:374].
Hydrogen sulfide removal with no carbon dioxide present
occurs in the purification of refinery gas, natural gas, and for the
recycle hydrogen used for oil desulfurization.
The range of hydrogen
sulfide partial pressure can range to 350 psi, and the purification
can be as low as 0.1 to 1.0 grains/ 100 standard cubic feet.
The
choice of a hydrogen sulfide removal process also depends on the
sulfur recovery process [1:374].
The most common treating problem occurs when both carbon
dioxide and hydrogen sulfide are present.
The requirements for
purification are varied and can range from maximum removal of
hydrogen sulfide and most carbon dioxide to removal of hydrogen
sulfide only [1:374].
The partial pressure of the acid gas will help determine
which process is best, but the ratio of the partial pressures will
also be of interest when selecting a process.
At high carbon dioxide
to hydrogen sulfide ratios, promoted hot carbonate solution will be
selected over amine solvents.
At low ratios, amine or other organic
solvents are preferred; this is because the carbonate solution is
difficult to regenerate with high concentrations of sulfur present
[1:375].
In some processes, it is necessary to selectively remove one
67
gas while leaving the others; the most common case being selective
hydrogen sulfide removal from a stream also containing carbon
dioxide.
I
The actual selectivity obtainable is related to the
thermodynamics and kinetics of the system.
The thermodynamic selectivity of a solvent can be defined as
p*
. I
o<'
where~
is the amount of A absorbed, and the primed
quantities refer to hydrogen sulfide.
H
2
s
content to the C0
2
When ST
> 1,
the ratio of the
content at equilibrium, is larger in the
liquid phase than in the gas phase [1:91].
For physical solvents, the selectivity equals the ratio of
H
the Henry's law constants
[1:91].
The selectivity in water is 3.05, but higher selectivities
are available in non-aqueous solvents; see figure 28.
These
physical solvents, in addition to having poor capacity, have
practically no kinetic selectivity, so that thermodynamic selectivity
is the best they can obtain.
For chemical solvents, ST
=
p*
H a y'
y'
p*' y
=
[1:92].
H' a' y
Since all chemical solvents for hydrogen sulfide and carbon dioxide
are alkaline, the following equation is possible
[1:92].
The equilibrium constant for the above equation is K as
cs
plotted in figure 29, and can be written as:
[1:92].
68
•
Solvent
N
0
~
N
:t:
I
V'>
~
,_
:::;
~
..I
0
V>
me thy I eyQOOCICI!tote
propy Iene earbcrlcte
terranydrothioFene1, 1-dioxide
dimetnvlether of
poly~thylene 'illycol
tributyl phosphate
N-methyl- t-e~ls:ctam
N-methyl pynolidone
methanol
PC
Sulfolane
Selexol
TBP
NMC
NMP
M.OHe
•
V'>
<
NMC
Me. A
TBP
SELEXCL
<.:>
a
v<(
w
•
SULFOLANE
~
•
MeOH
1:1..
.....
0
0
;::
e ePC
MCA
~
C'_l~-------:-':L:----~--·--L0
lQ
20
30
H2S SOi.~ILITY, CC/CC SOLVENT/AIM,
Figure 28
H S/C0 Solubility Ratio for Organic Physical Solvents
2
2
Source: Astarita, Savage, and Basio, p. 91.
69
j
j
·-· ·~------4--~··-----3~
3.0
3.2
3.4
Figure 29
Arrhenius Plot for K
cs
Source: Astarita, Savage, and Basio, p. 87.
~.o
70
The selectivity can now be written as
H
S
T
= H'
--
[HC03l y'
----- --
K
cs
[1:92].
[HS-] y
Since the second dissociation of hydrogen sulfide is so low,
as can be seen on figure 30, most of the combined hydrogen sulfide
is in the form of HS-.
however, can exist as
Therefore, except for
Therefore, my'
< 1,
2
The carbon dioxide
<
than HC0 - •
3
H
K
H'
cs
For aqueous solvents
the thermodynamic selectivity is always let than 3.05,
the selectivity of water.
combined C0
[HS-].
co32- or the carbamate ion other
K2co 3 solutions, my > [HC03l·
After substitution, ST
with K
cs
=
This selectivity increases as more of the
is in the bicarbonate form.
If high levels of
bicarbonate increase the selectivity, then high levels of
be avoided.
co32- should
This means that the pK of the solution should be less
than the pK for the second dissociation of carbonic acid, about 10.33
at room temperature.
With this restriction, we can see that very
strong alkaline solutions have very poor thermodynamic selectivity
[ 1:93].
If the solvent is of an organic nature, then the formation of
the carbamate ion is also possible.
The thermodynamic selectivity
will be highest when the carbamate ion is the most unstable; this
occurs with tertiary amines [1:94].
In these tertiary amines and in low pK inorganic solutions,
the chemically combined carbon dioxide will exist only as HC03.
This will be the highest thermodynamic selectivity such that
71
1o-12
5
c-
>
c
2
r
0
"'
':il
~~
~
I
10-D~
"'
r.-i=
2=t'
5
V>
N
"'
...,_
11~-1
~
5
2
10-1
1I
I
t
...
l
oi\-
!
!I
a
I
i
rl
~
l
'
:~-....-.--L-.....1..----..l.-.----l..--.-l--.-....._-___j
1)
2Q
40
60
so
100
Figure 30
Second Dissociation Constant of H s
2
Source: Astarita, Savage, and Basio, p. 93.
120
1.4l)
72
H
= -- K
H'
cs
[1:94].
Kinetic selectivity is based on the rates of reactions between
the solvent and the gases.
Solvents will have a kinetic selectivity
towards the gas which is first to react with it.
For a carbon dioxide
~nd
hydrogen sulfide gas stream being
absorbed into an amine solution, the reaction of the hydrogen sulfide
and the basic solvent will be instantaneous due to a direct proton
transfer.
The carbon dioxide however, must go through several
reactions before being absorbed in to the amine solution [8:119].
When carbon dioxide first dissolves in an aqueous solution,
it first forms carbonic acid, H co • The carbonic acid slowly
2 3
+
dissociates to form H and HC0 ions. The hydrogen then reacts with
3
the amine.
Since the formation of this hydrogen and bicarbonate is
slow, the overall absorption of C0
2
is slow [8:119].
A second co;--amine reaction is possible when labile hydrogen
atoms are present on the amine.
with primary or secondary amines.
This carbamate formation occur only
For this reaction, a carbon
dioxide molecule may react directly with two amine molecules to form
the carbamate complex.
This reaction is faster than the bicarbonate
reaction but is still slower than the direct proton transfer [8:119].
It is this difference between the amines which make tertiary
amines more selective towards hydrogen sulfide.
The tertiary amines
have no labile hydrogen atoms and therefore, carbon dioxide
absorption must take place via the slow bicarbonate route.
A way of comparing the selectivities is by using the equation:
73
~mol~~S)~~~~=-~~=:~~~~Tr~~~~~~~~~~~l~~~~~d
gas
(mol%C02 )fd gas - (mol%C02 )Treated gas/(mol%C0 2 )fd gas
The selectivity of MDEA is 3.85, for MEA it is 0.89, and for DEA, it
Sel =
is 2.27 [6:45].
Selective removal of hydrogen sulfide over carbon dioxide can
be achieved with physical solvents if the partial pressure of
hydrogen sulfide is above 60 psi.
Although this will remove the
hydrogen sulfide, an amine process will also have to be used to make
the gas reach the 0.25 grains/100 standard cubic foot specifications
[1:375].
When the hydrogen sulfide is at a low partial pressure, and
its level in the product gas must be extremely low, direct oxidation
in the regenerating step appears best.
In this process, the hydrogen
sulfide reacts to be reduced chemically and then contacted with air.
The solvent is recycled and the resulting sulfur is elemental [1:375].
Kinetically selective hydrogen sulfide removal is also
possible with the use of aqueous diisopropanolamine.
Figure 27 shows
the range of partial pressures for which these processes are selected
[1:376].
Chapter 8
THE DESIGN OF A GAS PURIFICATION PLANT
USING MDEA FOR SELECTIVE H S REMOVAL
2
For the design section of this report, the feed gas selected
is gas which is obtained from the off gas of a Beavan tail gas unit.
This gas has already gone through a purification process but still
has too much hydrogen sulfide to release to the atmosphere.
Since
this gas is waste gas, most of it is made up of nitrogen, carbon
dioxide, hydrogen sulfide and water.
Since only a small fraction is
hydrogen sulfide, 3 percent, it was decided to selectively remove the
hydrogen sulfide to reduce the utility cost.
The solvent selected is
N-methyl diethanolamine, MDEA.
Gas treating with MDEA is relatively new in this country, and
has not had much attention.
It is a new process, and the data needed
to build MDEA treating plants are not readily available.
The cost of
MDEA is three times more than for MEA or DEA, and this extra cost had
no way to be offset.
In turn, it was ignored until the rising cost of
energy demanded that new processes be developed [6:49].
MDEA has many assets which make it a desirable solvent when
selective removal of hydrogen sulfide is needed.
The heats of
reaction which hinder gas absorption and increase the regenerating
cost are lower for MDEA than for DEA and MEA.
The residual acid gas
in the solvent is on the order of 4 times less for MDEA.
The
corrosion of MDEA on carbon steels is low, on the order of 0.04
74
75
mm/year.
MDEA does not breakdown to form other compounds like DEA
and MEA do.
It also does not foam; this reduces the diameter of the
absorber column.
The comparative results for a MDEA plant versus a
DEA plant can be seen in tables 3 and 4 [2:115-6].
Because MDEA purification plants have a large number of
process variables, a computer program was developed.
In this way,
MDEA plants could be designed and optimized by entering and changing
the operating parameters.
The computer program follows the basic
design routine for gas treating [4], but uses curve fits to get the
dependent values.
The program illustrates the simple layout for the
plant; see figure 31.
Computer graphics are used to extract
information from the equilibrium diagram; this requires some operator
input.
See figure 32 for the original diagram and figure 33 for the
computer graphic diagram.
The following will give details of the design, and special
notes will be added to explain how the computer interacts with the
design.
The boundary limit values are the first to be determined.
These are already set and depend on such things as plant location,
available utilities and the main plant process.
The computer program
has a set of default values for these parameters, but they can easily
be changed to fit the plant features.
Some values are fixed and cannot be changed except in the
computer program itself.
These values were chosen to comply with
engineering principles which are already proven.
These values will
be noted when used.
This design uses 12 wt. percent MDEA because that was the best
~
'
76
Table 3
Comparitive Costs of the MDEA Process and a Total
Gas Sweetening Process, (CHEMERY CASE)
SNPA-DEI
MDEA precess MOEA p•ocess
case,
Cast: 2
process
Raw 9as
• Co1'1"DOSI110n:
CO:o (% Vo!,
HzS (M~S'"JI'1" 3 j
• Pressure !M""a)
• Flow-rate (Nr-, 3 /h)
2
30
2
2
30
30
6
250.000
8
250.000
250,000
1.85
2
, .6
70
1.85
5
4(·
1
55
200
20
20
80
B
Treareo gas cor.1posir.:m
C02 ( ~t vol I
H?S tm;S 1Nr'l:')
0
L P Stearn (T'h)
E..e::tnc1ry (i(W•:)
CnE>rruca' p•ooucts
consumpt•on (Tty)
, 0
Source: Blanc, Eluge, and Lallemand, p. 116.
0.1
77
Table 4
Comparitive Costs of the MDEA Process and a Total Gas
Sweetening Process, (Fuel Gas Treatment Case)
MDEA pri.X'..ess
DE.A process
HzS
30
20
30
CH 4
50
0.8
2C
50
0.6
U.OOO
14.000
H7S
<0.05
co~
20
Raw oas
• cOmpDsitiOn (% Vel)
co,
• Pressu•e (Mpa)
• Flow rate (Nm 3 't1)
Treated
£185
compositJO'l
("'c voi)
Stea~ (T.·'h)
15
f:ie:::t'ICI!y !i<Wt":)
"0
<0.01
0.5
21
70
4
7
LP.
Chem·:::.al Products
c:on~umptior.
ty)
rr
Source: Blanc, Eluge, and Lallemand, p. 116.
78
*****
*******
H2S REMOVAL USING MDEA
ABSORBER
REGENERATOR
{.,.l ;j~!l ~ lj
I lUll. .
REFLU1~
ORUr1
1'111'1!1
1
t!l'Iii!'llli!d
l .
-:·:•
--~-~r-·=
Ll
-
.
:~Wr
~
tp~~'iln
1·---~----1
t-;.
- - - ;:.-4:::.,_
.
___ .------1.I
~...
--"'
.,.
REBOILER
Figure 31
Computer Graphics Diagram of MDEA Gas Treating Plant
Source: Generated by Michael P. Caldwel-l
79
l,ODO
li
-
0..
.II:
1.0 kmor m-3 MDEA Solution
100
10
01
001
Figure 2
--~~~l~!~l~!~·~ol--~-~~--~~
0.01
O.l
tO
Mole Ra1io in Liquid ( H 2 S/MDEA}
Figure 32
Equilibrium Diagram for H s and MDEA
2
Source: Jou, p. 7.
40
80
H2S EQUILIBRIUM WITH MDEA
10000
1000
p
p 100
H
2
s
K
10
1
p
A
.1
. 01
. 0 01
. . .· .•
1 5 8 F .. · 1 0 4: ...:-
.
.
:
:
.1
1.
..................:::::: .... :::~, ... :::................ ··········1································1···············{
.001
.01
3.2
MOLE RATIO CH2S/MOEA)
Figure 33
Computer Graphics Equilibrium Diagram for H s and MDEA
2
Source: Jou, p. 7.
by Michael P. Caldwell
Transferred to Apple Graphics
81
set of equilibrium data available [5].
is not the optimum concentration.
It should be noted that this
As was stated earlier·, this is a
relative new process and compiled data is scarce.
To perform an energy and material balance the temperature of
the fluid leaving the bottom of the reboiler is the first value to be
determined.
It is a function of the pressure and the percentage of
water in the liquid.
regenerator side.
The pressure is the summed pressure drops of the
These include pressure drops due to the reflux
drum, the condensor, the regenerator tower, the line drop, and the
control valves.
Raoult's Law can now be used to determine the vapor
pressure of the water.
The corresponding temperature can now be found
in the steam tables; the computer program has this curve fit.
Determining the amount of actual MDEA needed is the crucial
part of the design.
This will determine the amount of actual liquid
flow, which in turn, sets the amount of utilities which must be used.
This is a trial and error method which finds the MDEA needed and the
corresponding temperature at the bottom of the absorber.
Equilibrium
diagrams are used to determine when the temperature, MDEA flow, aQd
the partial pressure of the sour gas are matched.
Once this MDEA
flow is found, it is general practice to increase this value just to
be on the safe side.
The computer program has this value set at a
20 percent increase but this value can be changed while running the
program.
A heat balance around the absorber must be done to determine
the temperature at the bottom of the tower.
Then this temperature is
checked on the equilibrium diagram to determine the amount of MDEA.
The temperature now is changed with due to an updated flow rate.
82
Then this process is repeated until the system is balanced.
The heats which must be considered are the heats of reaction,
the heat of the exiting rich amine, the heat of the exiting clean
gas, the heat of the entering lean amine, and the heat of the
entering sour gas.
The heats of reaction for carbon dioxide and
hydrogen sulfide with MDEA are set in the program and can only be
changed by changing the program.
These values are set at 450 BTU/lb
for hydrogen sulfide and MDEA, and 570 BTU/lb for carbon dioxide and
MDEA.
Note that these values are much lower than for MEA and DEA;
this means that it takes less energy to strip the MDEA [2:115].
Heat capacities had to be determined for the exiting treated
gas and are dependent on the temperature.
The heat capacity for the
absorbed acid gas is set at 0.915 BTU/lb [4] and can only be changed
in the program.
The heat capacities for the amine solutions were
determined by a curve fit from know properties [7].
Assuming a value for MDEA flow, the temperature at the
bottom of the absorber can be estimated with the heat balance.
Now this MDEA flow is checked on the equilibrium diagram and a new
MDEA flow is determined.
In the computer program, the operator must
move a cursor to the corresponding temperature.
Once the values
match, the flow rate is set and the temperature of rich amine is
known.
The rich/lean heat exchanger temperatures and duty can now be
determined.
The cold side approach temperature is one of the
changeable parameters in the program with the default set at 35 °F.
This is the difference between the entering rich amine and the exiting
lean amine.
The temperature difference between the entering and
83
exiting lean amine is now used with the flow rate and curve fitted
heat capacity to determine the exchanger duty.
Once this is found,
the exiting rich amine temperature can be found; this is the
temperature of rich amine to the regenerator tower.
The duty can now be found for the trim cooler.
Using the
temperature of the cooling water and the set approach temperature,
the temperature difference can be found.
This is used along with the
curve fitted heat capacity to determine the duty.
Now that everything is known about the absorber, the diameter
of the tower can be determined.
The computer program will determine
this diameter for a packed tower with set properties.
This tower is
to be filled with #2 Hy-Pac and have a pressure drop of 0.5 inches of.
water/foot.
The diameter is determined by using the standard curves for
gas and liquid flow rates [1:402].
See figure 34.
The computer uses
a curve fit for the 0.5 inches of water/foot and a packing factor of
18.
L
The x-axis of the curve is:
(PG
GJ f~ and
1
2
G Fv·
the y-axis is p~(p~=pG) where G is the gas rate, L is the
liquid rate,
PG
is the gas density, p
1
is the liquid density, F is the
packing factor, and Vis the kinematic viscosity.
The viscosity was
determined by a curve fit obtained from known properties [7].
The duties of the regenerator side can now be determined by
doing heat balances.
The amount of steam used by the boiler is set
at 1 lb steam / 1 gal MDEA.
changing the program.
This cannot be changed except by
Using the enthalpy of the steam at the
pressure specified by the operator, the duty for the reboiler can be
84
10.0
6.0
Parame-ter of curvts is pr~ssurt
drop in incnts of water/ioot.
1.5
4.0
2.0
1.0
"""!
Co
0~1':-
u.O:
Nol:o
O.b
0.4
~
0.10
0.2
0.05
0.1
0.06
0.04
0.02
0.01
0.01
0.1
1.0
10.0
Figure 34
Generalized Pressure Drop Correlation for Dumped Packings
Source: Astarita, Savage, and Basio, p. 402.
determined.
This enthalpy is determined from a curve fit.
The heat of desorption is the same as the heat of reaction
except this time this is heat that is needed to strip the MDEA.
The heat leaving with the lean MDEA is determined using the
temperatures for the entering rich
~IDEA
and the exiting lean MDEA.
The heat capacity was found using a curve fit.
The heat leaving in
the acid gas uses the temperature of the entering rich amine and the
exiting temperature to the Claus plant.
This temperature can be
changed by the operator.
The amount the water in the acid gas is determined by finding
the partial pressure of the water at the Claus temperature, and
taking the ratio of water pressure to total pressure times the moles
of acid gas.
This is converted to pounds of flow and multiplied
times the difference in enthalpies of the exiting vapor and the
entering liquid.
The condenser duty can be found by summing the heats in
and subtracting the heats out.
Now the temperature at the top of the
tower can be found using a trial and error method.
The temperature is first estimated.
The amount of energy
used to condense the water is determined by subtracting the cooled
acid gas.
This is just the amount of acid gas times the heat
capacity times the difference in temperatures.
found with a curve fit.
This heat capacity is
This amount of energy used to condense the
water can be used to find the amount of water that is condensed.
Curve fits were used to determine the enthalpy of saturated steam at
the estimated temperature and for the enthalpy of the condensed
water.
Dividing the energy by the difference in enthalpies gives the
86
amount of water condensed and used as reflux.
Finding the partial
pressure for water for this ratio of water should be the assumed
temperature.
If it is not the same, the process repeats.
Now the information for the reboiler can be found by doing a
mass and energy balance involving the flow in the reboiler and the
utility steam.
The temperature at the top of the reboiler can be
found by Raoult's law.
From the literature, very little MDEA is
vaporized [7:10] so this overhead is assumed to be all steam. This
mass flow rate is the amount of steam generated which is also the
amount of steam condensed plus the amount of water sent to the Claus
unit.
The amount of flow to the reboiler is the amount of flow from
the top of the reboiler plus the amount of flow from the bottom of the
reboiler.
Now using the amount of heat received from the utility
steam and curve fitted values for enthalpies for the flows out of the
reboiler, the enthalpy for the flow entering the reboiler can be
found.
The temperature for this enthalpy is found using a curve fit.
The diameter of the regenerator tower can now be determined
using the same method as that for the absorber except this time, the
steam must also be counted as a gas.
Please refer to appendix A for the flow diagram of a sample
MDEA plant, figure 35.
The listing for the computer program is in
appendix B; for samples of the input and output of this program, see
appendix C.
REFERENCES
1.
Astarita, Gianni, David Savage, and Attilio Bisio. Gas Treating
With Chemical Solvents. New York: John Wiley and Sons, 1983.
2.
Blanc, C., J. Elgue, F. Lallemand. "MDEA Process Selects H2s."
Hydrocarbon Processing. 111-116. August 1981.
3.
Brauer, H., andY. B. G. Varma. Air Pollution Control Equipment.
2nd ed. Heidelberg: Springer-Verlag, 1981.
4.
Class Notes for Engineering 478B. University of Galifornia at
Northridge, Professor Babayan. Spring Semester 1983.
s.
Jou, F. Y., and others. "The Solubility of H2S and CO?. in
Aqueous Methyldiethanolamine Solutions." Paper presen'Eed at the
AIChE National Meeting, Houston, April, 1981.
6.
Pearce, R. L. "Hydrogen Sulfide Removal with Methyl
Diethanolamine." Paper presented at the 57th Annual GPA
Convention, New Orleans, ~arch 20-22, 1978.
7.
The Pennwalt Co.
Gas Sweetening".
8.
Sigmund, P. W., K. F. Butwell, and A. J. Wussler. "HS Process
Removes H S Selectively." Hydrocarbon Processing. 118-124. May
2
1981.
9.
Strauss, Werner.
Press, 1966.
10.
"Product Information: Methydiethanolamine for
Philadelphia: Organic Chemicals Division, 1980.
Industrial Gas Cleaning.
Wark, Kenneth, and Cecil Warner.
Harper and Row, 1981.
87
London: Pergamon
Air Pollution.
New York:
v .
APPENDIX A
88
I
.- -: - .,
f
rf~-~
~
.:.~,y1
~C
lFC)
.. -
. - -
-
- ·- -
-
-· - - .. - .. - -
::
~c,;
.'
CONDENSOR
----,
....
.
'*'
1
1
)-1i
--
I
--
;
- - --- - -
-- - ·
:
D-~
.
,~--t)8+-,
: i
/8 ,I
D-1 _ _ ,
--------
'
r·
I
I
,---,
I
'
I
~r- -
- - - - _ _ _
_ _ _ •
: ~,~--------
'
lI
GAS
I··-
1
:
_J
J.
I
:
~
;
I
:
·~-I
COOliNG
'-tl
IX
r-::=- ~
t><1
ReGEN.
H
PUMP
·
l
WATER
v
I
1,
~-: ~
t-
:
'
J:C
1
: .
1
~
rc
LR-f: 1·----->~
:
i
>-
RESET
I
:
.
?(I
·
l ir~--=-,I
I
.. - -
.<--{;::l:J--~
lI k
I
- - - ·- -
r----,
-;(-.P:-f·-H-----.L.._:L{ rc.
1
i
l
COOLING
.WATER
I
REGENERATOR
LEMU RICH
Ht/lT EXCHANGER
RES[!
:- --,Il
•
FILTER
TRIM COOLER
ABSO.R~F.R
ACI D G.A3
TO CLAUS
UlilT
R Etl uX
I'
DRUM
Y ~---1I
@1
r·----~
I
' ~ : :1"
L-------(_}----------~
L_ ctrJLb
CO~DENSAT£
SHAM
:
~~
------
I
-------------------- --------'
Lf AN MDE A
PUMP
HICH MDEA
PUMP
REBOILER
Figure 35
Flow Diagram for MDEA Gas Treating Plant
Source: Drawn by Michael P. Caldwell
CXl
\0
"'c'.>
.I
APPENDIX B
1
2
GOTO 10
IF EO < 1E3 THEN
EB
=1
: RETURN
3
IF EO < 1E4 THEN
EB = lEt
RETURN
4
IF EO < 1E5
EB
= 1E2
TH~N
RETURN
5
IF EO < 1E6 THEN
EB •"' 1E3
6
IF EO
RETURN
< 1E7 THEN
EB = 1E4
RETURN
7 ·
IF E:O < 1E8 THEN
t:B "" 1E5
RETURN
8
9
IF EO < lE9 THEN
EB = lEG
RETURN
I~
EO ( 1E10 THtN
EB
~.:
1E7
RETUnN
10
REM
0::·0
HOME
INTRODUCHON
VTAB <5)
PRINT " THIS PROGRAM W:LLL HELP DE~HGN A H.2S
RHJOVAL PLANT USING MDEA"
PRINT ""
30
~0
PRINT " THE PROGRAM WILL ASK THE OPERATOR
FOR CERTAiN lNFORMATION, IT WILL ALSO
ASK T~E OPERATOR TO GRA~HICALLY SELECT
A VALUE FROM A HIGH-RESOLUTION FIGURE"
PRINT ""
PRINT "THIS WILL BE DONE USING TH£ RIGHT AND
LEFT ARROWS"
PRINT ""
PRINT "PLEASE HIT ANY KEY TO CONTINUE"
50
GET X$
50
HGR2
PRINT
D$ == CHR$ (4)
P~INT
70
8.0
D$;"BLOQD
GET X$
IA = 2.91
IB = 732.58
IC = 21.81
ID = 38.54
IE
IG
90
IH
I I
IJ
= 0.02
= 0.01
= 51.40
= 110
= 16.'3
90
PIC,A~4000"
91
100
JS "' 64.7
IW
IR
Ml:.
cu
110
120
130
140
150
160
17ill
!80
= 92
3:5
~ ... 2
74
PU = 27.7
TG
110
TEXT
HOME
VTAB (2)
P~INT "THE FOLLOWING ARE DEFAULT VALL'ES"
PRINT '"'
IN\/ERSE
PRINT "PLEASE NOTE UNITS"
NORMAL
PR!NT " "
PRINT "A) INLET GAS TEMP'':
HTAB (25>
I=RINT II;
PRINT " F'"
PRINT "BJ INLET GAS PRESS";
HTAB !25i
PRnCT IJ;
PRINT " PSIA"
PRINT "C> H2u;
WAB (25)
P~INT IA;
PRINT " MOLES/HR"
...
PRINT "0) N.-.,
<=. •
)-;TAB <25)
PRINT IB;
PRINT
MOLES/HR"
PRJ NT "El H2S";
HTAB <25>
PRINT rr.
PRINT " MOLESiHR"
PRINT "F} C02";
HTAB <25)
PRINT ID:
PRINT .. MDLES/HR"
PRINT "G> CO";
HTAB (25)
PRINT IE;
PPINT .. MOLES/HR"
PRINT "H) COS";
HTAB <25)
PRINT IG;
PRINT . MOLES/HR"
PRINT "I l H20";
HTAB <2_5)
PRINT IH;
PRINT
MOLES/HR"
PRINT "J) COOLING WATER";
HTAB <25)
PRINT IW;
PRINT . F"
PRINT "K> AVAILABLE STEAM";
HTAB <25>
PRINT IS;
PRINT
PSIA"
PRINT "Ll APP TEMP FOR HTX";
HTAB (25)
PRINT IR;
PRINT . F'"
PRINT "Ml LOADING FACTOR";
HTAB <25>
PRINT ML
..
190
·-·
2ill0
210
220
230
.
240
250
..
260
265
266
I
PRINT "Nl TEMP TO CLAUS UNIT" !
HTAB <25)
P~INT CU;
PRINT
F"
PRINT "0) Pq£S'S . i!J CLAUS UNIT";
HTAB <25>
PRINT PU;
PRINT " PSIA"
PRINT "Pl TE!IIP OF" n~EATED BAS";
II
257
258
HTAB
270
280
GET X$
HOME
VIAB
290
(25)
PRINT "TG;
PRlNT " F"
PRINT
PRINT "TYPE TKE LEiTER TO BE CHPNGED OR <CR}
I~"'"
Xfi
(5i
= ''An THEN
PRHH ''THIS IS THE INLET GAS TEi'1r'ERATURE"
PRINT
WHAT SHOULD
II\IPUT I!
!F xs = '·B'' \HEN
PRINT "T!-<IS IS THE
PRINT
PRINT "\.IHAT SHOULD
INPUT IJ
IF X$ = ·~c" THEN
PRINT "THIS !5 THE
PRINT
PRINT ''WHAT SHOULD
INPUT I.:\
IF X$ = "D" THEN
P~n;T "7HIS IS THE
PRINT
PRINT "WHAT SHOULD
INPUT IB
IF xs = "Eu THEN
PRINT "THIS IS THE
PRINT
PRINT "WHAT SHOULD
INPUT IC
IF l($ = "F'l "!HEN
PRINT TH IS IS THE
PRINT
PRINT "WHAT Si-i{)ULD
INPUT ID
IF X$ = uGu TI.!EN
PRINT "THIS !S THE
PRINT
PRINT "WHAT SHOULD
INPUT IE
IF xs = UHU THE.N
PRINT "THIS IS THE
PRINT
PRINT "WHAT SHOULD
INPUT IG
IF X$ = u I Tlo{EN
PRINT "THIS IS THE
PRINT
PRINT "WHAT SHOULD
INPUT IH
PRPH
300
310
320
330
340
11
11
350
360
370
ZT BE tF') ., ..
INLET
H
~s
PRESSURE
BE (P51Rl""
HYDROGEN CONCENTRATION"
IT BE iJ'I\OL/HR ) ?u
t.jJTROGEN CONCENTRAl"10N"
IT BE <MOLIHR
)
~~~
H2S CONCENTRATION''
IT BE (I'IQLIHR )?"
C02 CONCENTRATION"
) ?lt
H
BE <MOL/HR
co
CONCENTRATION"
IT BE <MOL/HR
)
.,~
cos CONCENTRATION"
H
BE (I'IQL/HR >?"
~•
H20 CDNCENTRAT!ON"
IT BE <I'IOLIHR
) ?•
93
380
390
400
405
4~6
407
IF XS
PRINT
PRINT
PRINT
INPUT
410
420
430
PRINT
PRINT
PRiNT
!NPUT
IF X$
PRINT
PRINT
PRINT
INPUT
I"" XS
PRI!'JT
PRINT
PRINT
INPUT
IF X$
PRINT
PRINT
PRINT
INOUT
IF X$
PRINT
PRINT
INPUT
IF X$
PRINT
PRINT
PRINT
INPUT
IF X$
Go:o
MA
450
465
470
"WHAT SHOULD IT BE <F>?"
IR
= "M" THEN
"THIS IS THE LOADING FACTOR"
"WHAT SHOULD IT BE?"
ML
= "N" THEN
"THIS IS THE TEMP TO THE CLAUS L.:NIT"
"WHAT SHOULD· IT BE?"
CIJ
= "0" THEN
"THIS IS THE PRESS TO THE CLAUS UNIT"
"WHAT SHOU'<.D IT BE?"
PU
= "P" THEN
"THIS IS THE: TEMIJ OF THE TREATED GAS"
"WHAT SHOULD IT BE""
TG
= CHR$ 113J GOTO 430
110
2.016
28.i2113
34.08
t>"D
44.01
ME
MG
MH
28.011
TZ
TM
PA
PB
PC
PD
PE
PG
TlolEbi
"WHAT SHOULD IT BE <PSIA >?"
IS
= "L" THEN
"THIS IS THE APP TEI'IP FOR THE HT )("
MB
ZP
450
"I("
"TI-'!S IS THE AVAILABLE STEAM PRESS"
MC
~.tf!
440
"WHAT SHOU..D IT BE (F)?"
IW
IE X~ "'
P~INT
408
,. "J" THEN
"THIS IS THE COOL!NG WATER TEMP"
60.075
18.015
119. 17
!A + IB + IC + ID + IE + IG +
14.7
TZ * .145
IA I
T!¥1
*
IH
!J
IB I TM * IJ
IC I TM * IJ
ID I TM * IJ
IE I TM * !J
IG I HI * IJ
PH
IH I TM * IJ
POKE 34,0
HOME
VTAB 15>
PRINT "THESE ARE THE PARTIC!L PRESSURES IKPA>
PRINT
PRINT "H2....
•1PA
PRINT "N2....
n;PB
PRINT "H2S.... ";PC
PRINT "C02....
":PD
PRINT "CO....
•;PE
PRINT "COS....
•;PG
PRINT "H20...
•;PH
REM
FIGURE TOP OF ABSORBER, THIS IS JUST
THE COOLING WATER + 16 F OR l2~ F, WHICH
EVER IS LOWER.THIS IS NOT HOWEVER THE TEMP
FRCITIIRF nF T~F TRF'QTFn RCI!';.
94
TT = IW + 16
IF TT > 125 THEN TT = 125
490
RE~
REFLUX DRUM PRESSURE IS RP
500
DP = 5
: ,XP
5
SP
3.5
480
510
LP
•5
VP
10
DP + XP + SP + LP + VP + ZP
RP
PR = Rt' I
. 98
NOW USE RAOULTS'S LAW TO DETERMINE HOW
MUCH OF THIS IS WA7ER VAPOR.IE. PARTIAL P
RESSURE EQUALS TOTAL PRESSURE TIMES MOLE P
520
REM
530
THE AMOUNT OF !I!DEA USED IS 1KMOL/W'3 SO
LUTION WHICH IS 11.917WT% "'DEA AND IS 98
MOLE" WATER.
REM
DO 'lLGORITHM FOR PRESSURE TO DEw POIN-r
FOR WATER ••• USE STEAM TABLES. .,..H!S \oi.AS D
ONE WITH A POWER CURVE FIT WITH A~ HP-41CV
. T~E RANGE WAS FROM 20.78 TO 6£.98
REM ~HE EQUATION IS y ~ AXAB WHERE RA2 IS 1
.000. A IS 115.59~7, AND B IS
.2269
BT
115. 5959 * PR /' . 2269
BO
115.5959 * RP A .2269
RC
.0025
E~CENT.
540
=50
560
555
570
REM
RD
580
590
600
6!0
620
630
640
645
650
• 001
REM
RESIDUAL ACiD GAS IN THE AMINE.FROM OP
ERATORS MANUAL, FOR H2S IT IS .0025 MOL H2
5/MOL ~D£A AND FOR C02 !i IS .001
REM WILL NOW DO HEAT BALANCE AROUND ABSORBE
Fl TO GET IDEA OF TEMPERATURE SO THAI EQUIL
!BR!UM CAN BE RUN LATER.
RE~
FIND SENSIBLE HEAT OF TREATED GAS. USE
FORMULAS FOR CP.
DT = <<<II+ TT> ! 21 ~ 459.671 I 180
CA = (13.505- 167.96 * DT
.75 + 278.44.
DT A - 1 - 134.01 * OT A - 1.5) I MA
CB = <9. 3355 - 122. !56 * DT '' - 1. 5 + 256. 3B *
DT - - 2- 196.08 * DT A - 31 I ~B
CH = <34. 19- 43.868 * DT A .25 + 19.778 * D
T A . 5 - .88407 * DTI I M~
CT = IA I ( IA + IB + IHI * CA + IB I ( IA + I
B + IHI * CB + :H I CIA + IB + IHI * C~
REM
CP AVERAGE IN BTU/LB R
DQ = .73 * lD
REM
TREATED C02
QT = <IA * MA + IB * MB + DQ * MD + IE * ME +
IG * MG + lH * MHI * CT * !TD - III
RE~
THIS IS Q TREATED GAS AND IS MIN
A-
us
660
670
580
7QI0
710
REM
THIS IS THE CP FOR ACID GAS AND IS ALSO
FOR THE SALTS CONTAINED CS IS FOR SOUR IS
ABOUT .915 BTUILB FROM CLASS NOTES.
CS = .915 * !IA * MA + <ID- DO> *MDI
REM MUST NOW GET CP FOR AMINE SOLUTION FROM
EQUATION WHICH IS CP=.916+.0004•T !N DEGR
E~S F.
ASSUME CP IS • 92 AND REFINE LATER.
HC = 450
HD = 577
REM
HEATS OF REACTION
HR = <HC * lC * MC + HD * <ID - DQI * MDl
REM
HE.ATS OF REACTION ADDES TO SYSTEM
95
p '
745
750
753
NDW GET PURE AI'IINE ROW ~T-E. F-1-RST--A
SSUME IT IS .2 MOLES ACID GAS/MOLE MDEA OR
5 !•IDEA/ACID GAS
PM = 5
IF F9 < > 1 GOTO 739
MN
<IC + RCl I X~
MO = _MN * <IC -+ '!D - DQl I IC
PM = MD J <IC + ID - DQl
PM% = PM * 100
AL =PM* IIC + ID- DQl * 119.17 I .12
REM
THIS IS THE AMINE FLOW RATE 1N L
BS.
T0 = TB
REM BEG:N FIGURING ABSORBER BASE TEMP
CP = .92
7biZI
TB = (
720
730
735
736
737
738
733
740
REM
REM
751
752
763
FIRST
01R iCS + ~,;._
763
'77~
*
QT} + CS
*
*
II + <AL
CP
*
TT))
CPl
IF ;:-a = l GOTO 1570
IF TB ) ~7 - . 1 AND TB ( T7
CP = .916 + .01Z104 * TB
T7 = TB
GOTO
7£E
T~Y
+
•
1 GOTO 766
76~
IF TB ( T0 + .5 AND TB ) T0 - .5 THEN
F8
1
q:_
AL "
~L
PM
PM
ML
*
GQTCJ 7E.0
IF F? = :
GOTO 780
PR I\, 7"
,,
PRIN~
"HIT ANY KEY TO CONTINUE";
GET
II
U
DR: t... -
7.30
PO~E
34,~
HOME
'JT,:OB
15)
'84
IF F9 = l GOTO 790
785
" t.DW BEGIN THE EQUILIBRIUM CHECK FOR
-HE H2S AND C02. ''
PR:NT
PRIN7
730
PR:~.;-
PRIN- " THE ~EMPERA-URE AT THE BOTTOM OF THE
ABSORBER HAS BEEN ESTIMATED USING ";
PRINT PM" I
DR:rc "
100;
MOLE MDEA/MOL-E ACID GAS."
;::JR;P\1-
739
8ill0
-8%
=
TB
+
.5
PRB:i "USE THE ARROWS TO MOVE THE CURSOR :o
-HE TEMPERATURE THAT WAS ESTIMATED, THIS
-~MPERATURE IS
";
F;_As>-
81IZI
"THE CURSOR IS DA!"PED AND WILL BECOME
MORE RESOLVED AS YOU CENTER IN ON THE
EXACT POINT"
;JR:;:~,;-:-
PR~NT
PRINT " wHEN TI-'E CORRECT TEMPERATURE HAS BEE
N
CHOSEN, HIT THE <ESC>"
820
CR!'\i-
INV£RSE
PRINT "NOW, '"'IT ANY KEY TO CONTINUE";
NORI"H'-.
830
840
GET a
PRIN'i
HGR2
I
96
85tll
aet21
870
880
890
'300
910
'320
930
'3"'0
'350
960
'370
D~
CHR$ (4)
:::
PRINT DS; "BLOAD EH2S,AS400~"
YC ::: '33.857 - 9.4304 * LOG <PC I ML>
HCOLOR= 3
X = 45
HPt.OT X,YC
NC = 12
NC = NC - 4
F1
F2
IF
0
=0
<NC
= 1 ) THEN GOTO '350
NC = 1
GET K'5
1-iCOLOR= 0
HPLOT X,YC
IF (K$ = CHRS (21>) GOTO 10..,0
lF (K$ = CHRS U3)) GOTO 1.0S0
rr. (K$ = CHR$ (c:7 > > GOTO 1140
'380
'3'30
1000
GOTO '350
1010
X
!~20
-~
1030
12••0
X --- 277
Fl = 1
.. r-
--
X + NC
277>
(
{X
I HEN GOTO 1040
HCOLOR= 3
r1:.:1LQT X, YC
11?150
GOTQ 1120
1060
F2 = 1
Y. = X - NC
1 070
~080
IF
::.0'30
11il'l0
X = 44
HCOLOR=: 3
HPLOT X,YC
<X ) 44) GOTO 1100
1110
GOTO 11--20
1120
ll30
IF tF1 = F2l THSN GOTO 910
C->OTO '350
11£;0
HCOLOR= 3
116G
HPLOT X,YC TO X,158
HP~OT X,YC TO 44,YC
TEXT
HOME
VTAB (5)
XC= .0002192 * EXP <.03449
MN ~ <!C + RCi I XC
REM
MOLES MDSA NEEDED
MD = <ID + RD) I MN
1540
F9
1150
1160
1165
=
*
X>
1
Go;:J 736
1570
1580
1590
16~0
REM WILL NOW DO THE RICH/LEAN HEAT EXCHANGE
R DUTIES
TE
TB + IR
CP = .916 + .~004 * <BT + TE) I 2
AA = AL + !C * MC + <ID - DQ} * ~D
RL = AA * CP * \BT - TEl
CP = .91G + .0004 * <TE +TTl I 2
TC = AL * CP * <~E - TTl
PQi-".E 34, 0
HOME
VT~-'
1510
PRH-JT
(5)
"
NOW
BEGIN THE ABSORBER SIZING.
THIS WILL BE A PACKED COLUMN USING •2
METAL HY-DAK.
THIS PACKING HAS A
PACKING FACTOR OF 18."
97
1620
PRINT ""
PRINT " THE GENERAL! ZED PRESSU~E DROP FOR A
PACKED COLUMN WILL BE DETE~INED USING
SEVERAL CURVE ~="ITS. THE DU>COfllE OF
THIS .wlLL BE UiE
~.R
DlPillETER-"
PRINT ""
PRINT "HIT ANY KEY TO
CONTINUE.
u
f
GET X.
1630
1635
1640
!650
1660
167121
1680
REM ~!GURE EVERYTHING TO GET X AXIS COOROIN
ATE.
LB = CIA* MA + IB * MB + IC * ~ • ID *MD +
IE * ME + IG * MG + IH * MHJ
FM = LB I !IA + IB + lC + ID + 1E + !G + lHl
REM
DF =
II
REM
PL =
REM
XA =
RE~
MW OF GAS
<FM I 379.5) * !IJ I 14.696> * (520 I (
+ 460l)
DENSITY OF FEED GAS
63.227
DENSITY OF LIQUID
AL / U3 * <DF / PL>
5
NOW DO CORREL~TIDN FOR PRESSJRE DROPS @
A
•
0.5
1690
1700
1710
1720
1730
1740
1741
1745
YA = .4331 - .3681 * LQG CXAI
REM DQ CORRELATION FOR VISCOSITY IN CENTIST
OkES
VS = 6-:iE.. 522
GS =
REM
~R =
REM
DM =
REM
01-
:=
*
C<TB + TT>
I
2l
" -
L 3987
YA / \ 1 8 * VS ·'- . 1 l * DF * ( PL - Dl=)
THIS !S G ~ 2
LB I
*
<3600
GS- .5)
AREA FOR NON-FOAMING OF
<AR
*
4 /
-~
3.14J_6J
WHIC~
MDEA IS.
.5
THIS IS DIAMETER
* 100
DM
HOM~
V''<B (5>
PRINT "TOWER DIAMETER !S
";01:
PRINT ,, FT 11
P~INT
PRINT "WILL NOW DO REGENERATOR BALANCE"
!='<!NT
1900
1810
182~
1830
:840
1850
PRINT " HIT ANY KEY TO CONTINUE"
GET X$
REM NOW BEGIN REGENERATOR SECTION
RM = AL + CIC * MC + (!D
DO> .. 1'10)
REM RICH I'!DEA
TR = RL I <RM * (. '316 + .0004 * <BT + TR) I
2> + < IC
MC + < !D - DQl * 1"0' * • 915)
IF TR ( Tl *
+ • 1 AND T~ > T1 OOTO !850
Tl = TR
-
..
GOTO 1820
TR = TR + TB
RE~
THIS IS THE TEMPERATURE TO
TH~
TOP OF T
HE REGENERATOR
1860
HF
= 1l74.5776 *
IS~-
.0610
1890
AVAILABLE STEAM
QR = RM I 8.4 * HF
ClL = AL * (.916 + .0004 * <BT + TR) I 2) * (
BT - TRl
QA = <IC * MC + <ID - DQl *MD> * .915 • <TR
1900
REM
REM
1870
1880
- CUl
THE FOLLOWING IS A TEMP TO PRESSURE CUR
THIS IS FOR T~£ ~ANSE 6
0-100 DEGREES F AND IS ONLY GOOD IN TM!S R
ANGE.THESE ARE COMMON TEMPS ~QR CLAUS UNIT
S HOWEVER'
Vl = .0331 * EXP <CU * .03411
REM CU 15 TEMP TO C~AUS UNIT
V2=Vl I !PU-Vll * !IC+ liD-OOlJ
VE F:T USING EXP.
191~
1920
P~M
THTS rq MOl FS
~
H2n VAPOP
98
1925
1'326
1927
1'?30
31.724 + .9977 * cu
.4330 * TR
=-
H2
Hl
= 106!.72 +
REM HV AND HG
QW = V2 * MH *
FOR H20
CHl - H2)
1945
= QR - HR TS = 207
1350
INITIAL GUESS
P.EM
FG = <IC * MC + !1D- DQ)
19'+0
QC
QL
QA - QW
~
CU>
1953
H3
REM
1954
H4
REM
1360
1970
1980
1390
2000
2010
UW
PO
=
HV AT TOP OF
I
=
H5
REM
2100
TS
CU
*
*
CXP + PU + 34J
TS
4. !838
A
IF PP < PQ + .1 AND PP r PQ- .1 GOTO 2050
IF PP { PQ THEN
TS = TS - 1.
GOTO 1950
IF PP > PQ THEN
TS + 1.
C>OTO 1350
2060
.25 » ITS -
H4)) I MH
IUW + V2 + IC + ID - DQ)
34J I
9
*
MD>
STRIPPE~
=- 31.724 + .3977
HF Ai CLAUS UNIT
= ((QC- FG> I tH3-
= 'UW + V2)
<<XP + ~U> *
PQ = 2.6348E-
TS
2050
*
*
1069.822 + .3804
~
= 954.98-
.667!
*
IS
H=-C AVAILABLE STEAM
SN =
Q~
REM
LE~
H5
I
STEAM NEEDED PER HOUR
H5 = 1073.78 + .362
REM
HG AT
*
BO
TOP OF REB'JILER
*
21!0
H7 = - 34.234 + !.0112
2120
REM f.JF AT BCT..-Ot!f OF REBOILER
TH = IIUW + V2) • M4 * H6 + RL
BT
*
H7- QR)
I
<UW + V2 + AL)
REM
=
ENTHALPY
33.32~5 +
TE~P I~TU
2130
Bl
3500
REM
POKE 34,0
HG E
VT~B
RE~ANING
.9918
*
TH
REBOILER
<5>
3510
PRINT " NOW BEGIN THE REGENERATO~ SIZING.
THIS WILL BE A PACkED COLUMN USING
+02
METAL HV-O~K. THIS PACK.ING HAS n
3620
PRINT
PRINT " THE GENERALIZED PRESSURE DROP FOR A
PACKED COLUMN WI~L BE DETERMINED USING
SEVERAL CURVE F!TS. THE OUTCOME OF
THIS WILL BE THE TOWER DIAMETER. "
PRINT ""
PRINT "HIT ANY KEY TO CONTINUE.";
GET X$
REM
FIGURE EVERYTHING TO GET X AXIS COORD!
NATE.
JL = <IC * MC + <ID- DQ) * MD + (UW + V2> *
MH>
PACKING FACTOR OF !8."
3530
3540
365~
JF
REM
3550
JD
3670
JP
=
=
JL I
( IC + ID •· DGI + UW + V2l
MW OF GAS
<JF I 379.5) • <PU I 14.696)
< <TS + BT> I 2> + 460> l
REM
REM
DENS!TY OF FEED GAS
= 63.227
DENSITY OF LTQUID
*
<520 I
(
3&80
JX = <AL •
.5
l&90
~1'1
<UW + V2l
*
MHl I JL
~00-GGAA£-LATION
*
<JD I
JP)
FOR PR.ESS:JRE DROPS
@ 0.5
3700
37!0
3720
3730
3740
3750
3760
3770
JY = .4331 - .3681 *LOG <JX)
REM
DO CORRELATION FOR VISCOSITY IN CENTIS
TOKES
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JG ;_ JY I (18 * JV A • 1) * JD * (JO - JDl
~EM
THIS IS G A 2
:~ = JL I
13600 * JG ~ .5l
REM
AREA FOR NON-FOAMING OF WHICH MDEA IS
.
4 I 3. 14l&l
JM = (JA
THIS IS DIAMETER
REM
Q')(. = JM
* 100
02 = 0':4 / 100
HOME
VTAB
,.. .5
(5)
;02;
PR:NT "TOWER DIAMETER IS
PRINT " FT"
iJRINT
PRINT " HIT ANY KEY TO CONTINUE"
GET X$
HOME
VTAB (2)
PRINT "THESE ARE THE PROGRAM I.'ALUES"
PRINT ""
PRlNT "ABSORBER TOP";
HTAB 125>
TT')(. = TT + .5
PRINT TT')(.;
IJ
5000
~010
PR1'J..,. " F"
5020
PRINT "ABSORBER BOTTOM";
HTAB 1251
TB')(. = "7"B + .5
PRINT TBY.;
~RIN!"
5030
II
F"
PRINT "ABSORBER DIAMETER";
HTAB '25>
C')(.
OA
= OM *
= O'X I
100
100
PRINT OA;
P"<:NT " FT"
5040
5050
5060
5070
PRINT "AMINE SOLUTION FLOW":
HTAB <25l
EO = AL
GOSUB 2
AL1- = AL I EB
PRINT AL1- * EB;
PRINT " LBSIHR"
PRINT "PERCENT OF C02 PASSED";
HTAB (25>
PRINT "73 7-"
P~ItH "MDEA CONCENTRATION";
HTAB (251
MP = 1 I PM
01- = MP * 1000
DB = 0')(. I 1000
PRINT OB;
PRINT " ACID/MDEA"
PRINT "TRIM COOLER DUTY";
HTAB <25)
EO = TC
GOSUB 2
O'X = TC I EB
DC = 0')(. I 100
PRINT DC;
PRINT " MM BTUIHR"
A
100
S080
PRINT "RICH/LEAN DUTY";
HTAB <25>
EO ~ RL
0~ = RL I EB
OD ~ 0" I 100
PRINT OD;
5090
PRINT " f'IM BTU/f-iR"
PFUNl "TEMP LEAN OUT";
HTAB <25l
TE% = TE + .5
PRINT TE~;
PRINT " F"
5112'0
PRINT
"1"EI"'P RICH
<25>
TR,;"' TR + .5
OUT";
HTAB
5110
::i~INT TR";
PRINT " F"
;:JRINT "C:ONDENSOR DUTY";
HTAB <25)
EO ::: QC
0~ = QC I EB
G!O: = 0':4 / 100
~"RJ.NT
OE;
i="•RINT "
5120
~~M
BTUIHR"
PRINT "REBOILER DUTY";
i-TAB
i.c;5)
EO
QR
C:i. = OR / EB
OF
=
0" I 100
PJ:tlNT OF;
PRINT " r1M BTU/HR"
5122
PRINT "REBO!LER TEMP AT TOP";
HTi~B
<.::5l
0" "' BO + • 5
;::RINT O:X.;
5124
PRINT " F"
PRINT "REBOILER TEMP IN";
HTAB
0"
5130
C25i
=
Bl
·~<-
•
5
PRINT O:X.;
PRINT " F"
PRINT "STEAM NEEDED";
HTAB (25>
EO '-~ SN
GOSUB 2
= SN
SN"
I
PR!NT SN"
5140
EB
* EB;
PRINT " LBS/HR"
PRINT "REGEN TOP";
HTAB (25>
= TS
TS;(
+ .5
PRINT TS~;
PRINT " F"
5145
PRINT "REGEN BOTTOM»;
HTAB (25)
= BT
BT')(.
5150
+
.5
PRINT BT~;
PRINT " F"
PRINT "REGENERATOR DIAMETER";
HTA8 <25)
0~
= .JM
0"'
OG :::
•· 100
I 100
PRINT OG;
PRINT " I=T"
APPENDIX C
THE FOLLOWING ARE DEFAULT VALUES
PLEASE NOTE UNITS
A) INLET GAS lEMP
B> INLET GAS PRESS
C) H2
D) N2
E:.) H2S
F) C02
G) CO
H> COS
I> H20
J) COOLING WATER
K) AVAILABLE STEAM
L) APP TEMP FOR HTX
M> LOADING FACTOR
N) TE~P TO CLAUS UNIT
0) PRESS TO CLAUS UNIT
P) TEMP OF TREATED GAS
~YPE
110 F
16.'3 PSlA
2.'31 MOLES/HR
732. 58 MOLES/H!~
2i. 81 MOLES/HR
38.54 MOLES/HR
.02 MOLES/HR
. 01 MOLES/HR
6i.4 MOLES/HR
92 F
64.7 PSIA
35 F
1. 2
74 F
27.7 PSIA
11~ F
THE LETTER TO BE CHANGED OR
101
<CR>
102
THESE ARE THE PROGRAM VALUES
ABSORBER TOP
ABSORBER BOTTOM
ABSORBER DIAMETER
AMINE SOLUTION FLOW
PERCENT OF C02 PASSED
MDEA CONCENTRATION
TRIM COOLER DUTY
R ICH/LEAr.J DUTY
T£MP LEAN OUT
TEtf1P RICH OUT
Cm·mENSOR DUTY
Rt:BOILEI-~ DUTY
REBOILER TEMP AT TOP
i~~~E10 IL.ER TEMP ll\l
s·:t-::AM NEEDED
Ri:-~GEN
TOP
HEf:;EN ·;;mTTOM
R~GENERATOR
DIA~ETER
108 F
117 F
4.02 FT
143~tlllll
LBS/HR
73 -;<.
.222 ACID/MDEA
G. lb MM BTU/HI~
16. ::. MM B1 U/1-H~
152 F
231 F
10.17 MM BTU/HR
15.72 MM BTU.tHR
265 F
17200
215 F
266 F
LBS/Hi~
2. 88 Fl
~
.