- Email: [email protected]
Mercaptan removal from natural gas by the efficient cyclic adsorption process; a simulation study
Accepted Manuscript Mercaptan Removal from Natural Gas by the Efficient Cyclic Adsorption Process; a Simulation Study Zahra Tohidi, Shohreh Fatemi, Omid Taheri Qazvini PII:
S1875-5100(15)30037-8
DOI:
10.1016/j.jngse.2015.07.010
Reference:
JNGSE 868
To appear in:
Journal of Natural Gas Science and Engineering
Received Date: 26 May 2015 Revised Date:
7 July 2015
Accepted Date: 8 July 2015
Please cite this article as: Tohidi, Z., Fatemi, S., Qazvini, O.T., Mercaptan Removal from Natural Gas by the Efficient Cyclic Adsorption Process; a Simulation Study, Journal of Natural Gas Science & Engineering (2015), doi: 10.1016/j.jngse.2015.07.010. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
AC C
EP
TE D
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
ACCEPTED MANUSCRIPT
Mercaptan Removal from Natural Gas by the Efficient Cyclic Adsorption Process; a Simulation Study Zahra Tohidi, Shohreh Fatemi*1and Omid Taheri Qazvini
RI PT
School of Chemical Engineering, Faculty of Engineering, University of Tehran, Tehran, Iran Abstract
A more efficient and economical cyclic adsorption process was proposed for mercaptan removal
SC
from natural gas (NG) to reduce mercaptan content to less than 10 ppm and meet the environmental rules. Continuous sulfur removal is studied for the NG feed stream, with pressure of 6.8MPa, flow rate of 2850 Nm3/hr and molar composition of 95.98% methane, 0.00182%
M AN U
water vapor, 1% carbon dioxide, 0.0134% mercaptan and 3% heavier hydrocarbons (C3+). The proposed process of Pressure Vacuum Swing Adsorption (PVSA) was designed and simulated as a more efficient alternative process against the current Industrial Pressure-Temperature Swing Adsorption (PTSA). In this work, an improved PVSA process was simulated with sequences of bed pressurization, adsorption, equalization, blow down, bed regeneration by vacuum and purge by product, in each process cycle. Vacuum condition of 10 KPa with the
TE D
molar purge/feed ratio of 0.06 and temperature of 350K was required for appropriate bed regeneration from adsorbed mercaptan to approach to the continuous cyclic steady condition. Comparison between PVSA and PTSA, at the same feed characteristics, same packed columns
EP
and adsorption operating conditions, revealed that the PVSA process, with less cycle time than PTSA, could achieve the same product purity with 94.8%
productivity, whereas PTSA has the recovery of 74.04% and productivity of 2.79
AC C
recovery and 3.90.
.
. At the
same time, operating with PVSA, instead of PTSA process, would reduce the operating cost from 88 to 70
$
.
Keywords: Cyclic Adsorption Processes, Mercaptan Removal Unit, Natural Gas, PVSA Process, PTSA Process, Economic Study
* Corresponding author, email: [email protected], school of chemical engineering, Faculty of Engineering, University of Tehran, Enghelab Avenue.
ACCEPTED MANUSCRIPT
1. Introduction The progress of the international energy demand shows an average of 1.7% annual growth in the 2005 to 2020. This growth concerns all energy sources and natural gas demands which
RI PT
would be accounted for the highest growth rate in 2020 [1]. Increasing concerns of the harmful effects of natural gas contaminants on environment has led to the introduction of a number of natural gas treatments. Out coming NG from the well contains methane with impurities such as water vapor, carbon dioxide, nitrogen, hydrogen sulfide, light mercaptans, ethane and heavier
SC
hydrocarbons. Some part of the natural gas impurities must be removed before commercial use. Today, according to the recent environmental legislations, the sulfur emission in the atmosphere should be considerably reduced to less than 20 ppm [2]. Mercaptans, or more correctly thiols,
M AN U
are organic compounds in which the -SH groups are present in the molecular structure of the hydrocarbons. The mercaptans must be removed mainly for three reasons: (a) they have acidic property and can cause serious corrosion problems, (b) they have offensive odor and they are very inappropriate to be burnt, (c) most of them are highly toxic and affect the subsequent catalytic reactions [3]. In the conventional processes, the acid gases such as hydrogen sulfide and carbon dioxide are mostly removed in an amine-wash unit. Since light mercaptans are not
TE D
as acidic as hydrogen sulfide, they cannot be removed properly by amine washing and an additional step is required to reduce the sulphur concentration to an appropriate level. Gas purification by the adsorption process could be as an alternative process in progress for sulfur removal from NG [2]. Modeling and simulation of the gas adsorption processes have been
EP
investigated previously by some researchers. Some limited works have been done for investigating mercaptan removal by adsorption using different adsorbents [2, 5-12]. Shirani et
AC C
al. [13] also simulated mercaptan adsorption on 13X adsorbent in presence of water vapor at isothermal condition and presented the breakthrough curves from the packed bed. In another work, Taheri and Fatemi [14] simulated a PTSA (Pressure-temperature swing adsorption) unit for mercaptan removal in South Pars region. However, the high cycle time taking for cooling down and the high economic costs are the main disadvantages of the mentioned process. It believes that, by replacing the current PTSA process with PVSA (Pressure-vacuum swing adsorption), the efficiency of the process would be improved. In addition, in PTSA process, the temperature shock induced to the solid adsorbents by temperature variation in adsorptiondesorption cycles would reduce the life time of the adsorbents, whereas PVSA doesn’t require
ACCEPTED MANUSCRIPT
energy induction to the adsorbents. It should be noted that, there is no evidence in the literature about mercaptan removal by PVSA from NG in presence of other impurities such as CO2, water vapor and heavier hydrocarbons. In this work, a systematic simulation of mercaptan removal from NG in presence of other impurities is carried out by the PVSA process in face of the
RI PT
industrial PTSA process simulated by Taheri and Fatemi [14] with the same bed configurations and design. This comparison is carried out on the base of different regeneration conditions and the results are presented in terms of concentration and temperature profiles, and performance parameters such as purity, productivity and recovery as well as a brief economical study. In this
SC
research, industrial PTSA refers to the real mercaptan removal unit (MRU) in South Pars region of Iran that is working currently, and the process design parameters and the outlet data are taken
2. Process Description and Design 2.1.Description of process
M AN U
from there[14].
A systematic adsorption model has been presented and implemented to simulate the PVSA process proposed for purification of natural gas (NG) from light mercaptans in presence of other
TE D
impurities such as water vapor, CO2 and heavier hydrocarbons (C3+). The industrial mercaptan removal unit (MRU) is currently working in South Pars region of Iran and this process is designed in the mode of PTSA. This unit consists of six insulated two-layer packed bed columns in which the regeneration step is carried out by increasing temperature of the bed by
EP
the hot gas NG. The pressure of the upcoming feed is 6.8 MPa and the adsorption process has been initially designed to work at this high pressure. The properties of the inlet NG is specified
AC C
at Table 1. In addition, the adsorption bed dimensions in industrial scale and the type and amount of the adsorbents are presented in this table. Table 1.
Each column consists of two layers: 13X zeolite and activated alumina. The 13X zeolite is used for major adsorption of mercaptan, however because of high affinity of 13X to water vapor, water plays competitive role in mercaptan adsorption. Therefore, a pre-layer of Activated alumina is required to adsorb water vapor before entering the 13X layer. Adsorption of
ACCEPTED MANUSCRIPT
mercaptan and C3+ is ignorable compared to H2O and CO2 adsorption by activated alumina. [15, 17]. Physical properties of the adsorbents are reported in Appendix A1. In this work, the process of PVSA is designed on the base of existing PTSA, but different
RI PT
according to the regeneration step and operating conditions. A simple PVSA with no equalization and an improved PVSA with pressure equalization steps are planned to study the impact of pressure equalization step on purity, recovery and productivity of the process. Also, the results of proposed PVSA processes would be compared with the industrial PTSA process.
SC
These results would be comparable with the established industrial PTSA that is currently working with no pressure equalization. The current PTSA process consists from regeneration
M AN U
step including two heating stages; 1st stage working at 480 K and the 2nd one working at 590 K. The proposed simple PVSA process is designed to include 10 steps as following: I.
Adsorption step; during 18 hours, the NG feed is introduced from top of the bed with a flow rate of 2.2 Kmol/s at pressure of 6.8 MPa and temperature of 302 K. 1st depressurizing step; the bed pressure is decreased for 6.8 to 4 MPa during 10
II.
III.
TE D
minutes. The gas content of the bed is released to the atmosphere, from the top. 2nd depressurizing step; the bed pressure is decreased from 4 to 1.5 MPa during 10 minutes by releasing from the top of the bed to atmosphere. Blow down step; the rest of the bed pressure is reduced down to the atmospheric
EP
IV.
pressure. 10 min time is considered for this step. Evacuation step; the bed pressure is evacuated by a vacuum pump to reach to 0.01
AC C
V.
MPa during 10 minutes.
VI.
1st purge step; the bed is purged counter-currently with 2% of the product stream at the vacuum pressure of 0.01 MPa during 10 minutes.
VII.
2nd purge step; the bed is purged counter-currently with 6% of the product stream in atmospheric pressure and temperature of 350 K during 15 hours and 40 minutes.
ACCEPTED MANUSCRIPT
VIII.
1st pressurizing step; the bed pressure is increased up to 1.5 MPa through feed introducing from top of the bed during 10 minutes.
IX.
2nd pressurizing step; the bed pressure is increased from 1.5 to 4 MPa during 10
RI PT
minutes. 3rd pressurizing step; the bed pressure is increased from 4 to 6.8 MPa during 10
X.
minutes.
Rest; the bed is put on the rest for one hour to be prepared for the next cycle.
SC
XI.
The improved PVSA process was designed as the alternative of simple PVSA with the same
M AN U
conditions but considering pressure equalization steps between the columns. Pressure equalization step provides inter-connection between two columns in which one of the columns is working at depressurizing mode and the other one is at repressurizing mode. According to the above modes, step III is inter-connected to step VII, and the step II is inter-connected to step IX. The improved PVSA can be used to reduce natural gas consumption and to enhance process efficiency. In Figures 1(a), 1(b) and 1(c) the steps’ arrangement are exhibited for industrial
TE D
PTSA, simple PVSA and improved PVSA, respectively. Figure 1.
Figure 2 presents the operating pressure of the manipulated steps in the PTSA and proposed
EP
PVSA processes. In order to manipulate and control the steps of PVSA process with equalization, the adsorption time was set to 17.5 hour and the time of desorption steps (evacuation and purge) was set to 16.5 hour. This is done because pressurizing steps should be
AC C
at the first half time of the cycle in order to equalize with the depressurizing steps at the second half time of the cycle. In this situation, we could save more natural gas and enhance the process performance. Figure 3 shows the time-table of the process with equalization steps. Figure 2. Figure 3. 2.2.Model and assumptions
ACCEPTED MANUSCRIPT
A mathematical model with mass, energy and momentum balances that represents the dynamic behavior of the multicomponent adsorption bed was developed based on the following assumptions: Radial gradients for mass, velocity and temperature are neglected in the bed.
•
The bed fluid is assumed dispersed plug flow.
•
Porosity along the bed is constant with uniform packing.
•
Mass transfer to the particles is described by linear driving force (LDF) model with solid
SC
RI PT
•
phase concentration.
Diffusion and adsorption into the particles are assumed as lump kinetic transfer model.
•
There are no concentration and temperature gradients inside the particles.
•
Pressure drop along the bed is predicted by Ergun equation.
•
Gas behavior follows Peng Robinson equation of state throughout the process.
•
Adsorption equilibrium is described by Extended Langmuir Isotherm model for
TE D
M AN U
•
multicomponent adsorption, in which any isotherm parameter is temperature dependent. The flow velocity varies along the bed and is calculated by the total mass balance equation. •
EP
•
The column operates in non-isothermal and non-adiabatic conditions in which the heat is
•
AC C
exchanged with surrounding.
Temperature of fluid and solid phase is not identical and heat can be transferred between solid and fluid.
•
The solid temperature is homogeneous and a function of time during adsorption and desorption process.
•
In heat transfer to the environment, the wall resistance follows thin wall assumption.
ACCEPTED MANUSCRIPT
•
Heat capacities of the adsorbents and adsorbed phase are assumed constant.
•
Heat of adsorption is assumed constant and derived from the isotherm equation for each component.
RI PT
According to the assumptions, mass and energy balances of the gas and solid phases are shown in Table 2. Table 2.
are reported in Table 3.
M AN U
Table 3.
SC
The isotherm parameters of the components were found from literature [5, 15, 20, 21] and they
The correlations for physical properties, mass transfer and energy transfer coefficient are reported in Appendix A2 and A3, respectively. 2.3.Numerical Calculations
TE D
To obtain temperature and concentration profiles inside and at the exit of the bed versus time at each cycle, model equations should be solved, simultaneously. Because of complexity, these nonlinear equations cannot be solved in analytical form and numerical solutions should be used. The numerical algorithm implemented was finite difference scheme with the first order Upwind
EP
Differencing Scheme (UDS1) method to convert spatial derivatives of the PDEs to the discrete form and convert the PDEs to the set of ODEs. Then by using implicit Euler method, all the
AC C
ODEs were solved simultaneously for the next time step [26, 28]. The bed was discretized by 20 nodes for each packed bed of activated alumina and 13X zeolites. The sensitivity of the simulation results to the node numbers was evaluated and it was found that the results are independent of the node numbers for node numbers greater than 20. The time step for implicit Euler method was also chosen to be 100 sec. The dependence of the results on the time step was also examined and it was found that the results were independent of the time step for time step values smaller than 100 sec.
ACCEPTED MANUSCRIPT
Each cycle consists of different steps in which different boundary and initial conditions are governed. The initial and boundary conditions for each step are introduced in Appendix A4. 3. Results and discussions
RI PT
To choose the best regeneration method for mercaptan desorption step, different conditions were performed and compared with each other on the base of mercaptan loading. Figure 4 shows mercaptan loading in different processes of PTSA, PVSA at 302 K and PVSA at 350 K
Fatemi [14].
M AN U
Figure 4.
SC
in the regeneration step. The PTSA results have been taken from the simulation of Taheri and
Obviously in Figure 4, PVSA with regeneration temperature of 350 K exhibits reduction of mercaptan loading to the lower level during a shorter time, whereas PTSA and PVSA without bed warming could not reach to this fast and short regeneration time. Therefore, PVSA with warming up to 350K, was considered as the regenerating operation in the cyclic process.
TE D
Some results of PVSA simulation are also reported in Appendix A5. 3.1.Cyclic steady state condition
The cyclic adsorption process as every dynamic process should pass several cycles to reach the steady state condition. Therefore at the next cycles, the extract and raffinate compositions
EP
become constant and do not change with time. A cyclic steady state (CSS) is defined for the same bed position in each cycle in which the dependent variables approach to a constant value.
AC C
In the present work, total adsorbed amount in adsorption stage between two subsequent cycles were checked to approach to the CSS as following:
∫
L 0
q i dz
L
( n −1) th cycle
− ∫ q i dz 0
<δ
(1)
( n ) th cycle
In the above equation, convergence parameter δ was considered to be 10-5 and this condition was selected as the stop criteria of the dynamic simulation.
ACCEPTED MANUSCRIPT
It was revealed that, after about twelve sequential running cycles of simple PVSA process, the steady state condition was achieved. The Figures 5, 6a and 6b show methane purity versus cycle number, mercaptan mole fraction in the outlet stream and time variation of solid average temperature, respectively. Regeneration of the adsorbent is a critical point in cyclic adsorption
RI PT
processes. Appropriate procedure is required for evacuation of the impurities from the bed. In the present case, mercaptan is the strongest adsorptive material on 13X zeolite, therefore properly bed evacuation from mercaptan at each cycle and approaching to a constant temperature is required for a continuous cyclic process. Since the adsorption of the other
SC
gaseous compounds, such as N2 is not significant, few cycles are required to approach to the cyclic steady state condition. Especially in this case, the vacuum condition helped better
M AN U
regeneration. Figure 5. Figure 6.
3.2.Validation
TE D
Average outlet mercaptan and water vapor mole fractions which are resulted from PVSA simulation at steady conditions are well validated with industrial unit data. The comparison of simulated and real values for mercaptan and water vapor mole fractions are shown in Figure 7. According to the reported values, mercaptan and water vapor relative errors are 1.5 % and 7.8
EP
%, respectively.
Figure 7.
AC C
In order to investigate the validity of the simulation results, mercaptan breakthrough curves from adsorption step in the real and simulated conditions are plotted and compared with each other, as shown in Figure 8.Mercaptan outlet from 30 hours adsorption time shows gradually increasing mercaptan at the exit section of the bed. The breakthrough results show that real plant data and simulated results are close together and there is high precision between real and model data during 18 hours adsorption which is the adsorption time of the process. Figure 8.
ACCEPTED MANUSCRIPT
3.3.Breakthrough curves The breakthrough curves of the species are shown in Figure 9 which present the mole fraction
RI PT
of each species in gas phase at different height of the bed. Feed gas enters from the top of bed and passes through the alumina pre-layer. In this layer, most of water vapor and some of CO2 content of feed is adsorbed before entering the 13X layer. As shown in Figure 9a, the bottom of the bed (13X, bottom) is approximately free from mercaptan whereas the top of 13X is rapidly
SC
reached to the feed mercaptan content. As shown in Figure 9b, the majority of water vapor has been removed in alumina pre-layer and the residue was adsorbed in 13X layer. Figure 9c and 9d
M AN U
show breakthrough curves of carbon dioxide and propane at different points of the bed, respectively. It can be seen that the fractions of carbon dioxide and heavy hydrocarbons reach to the feed content between 1 to 2 hours after passing through the column. However, the affinity of 13X towards CO2 is higher than that of propane.
Figure 9.
3.4.1. Loading curves
TE D
3.4.PVSA process in comparison with simulated PTSA process
One important factor in every cyclic adsorption process is the adsorbed amount of solid phase in
EP
adsorption and desorption stages. The loading curves show the maximum and minimum adsorbed amount and the time required for regeneration. Due to the fact that purging gas goes through the bed from bottom to top, the components accumulate at the top of the bed and this
AC C
point of the bed needs more time for desorption. Therefore the time required for mercaptan desorption at top of the bed can be an important factor for design of the cycle time which consequently affects the recovery of the process. Figure 10 shows mercaptan and water vapor loading in PVSA and PTSA processes at top of the bed [14]. According to Figure 10a, it is found that the time needed for complete mercaptan desorption at top of the bed in the PVSA process is 24 hours while in the PTSA process, it requires 30 hours to be totally desorbed. Water vapor acts as the dominant rival for mercaptan species in the adsorption by 13X. As it was seen previously in breakthrough curves, water vapor is a competitive adsorbing component
ACCEPTED MANUSCRIPT
with mercaptan and slows down bed regeneration. Therefore it is necessary to investigate water vapor competitive behavior and the impact of regeneration method on water desorption in both PVSA and PTSA process. Figure 10b indicates that loaded water vapor is evacuated from the top of 13X after 22 hours in PVSA process while in the PTSA, loaded water vapor has reached
RI PT
to 0.015 mol/kg after 25 hours (at the end of first heating step). The results reveal that it is possible to work with lower cycle time in the PVSA process rather than PTSA because lower time is required for the regeneration and this means saving recovery and productivity.
SC
Figure 10.
Table 4 reports the regeneration step time required for different cyclic processes to desorb
M AN U
mercaptan and water vapor and prepare the bed for the next adsorption cycle. The results show that using PVSA process has shortened the cyclic time compared with PTSA and it can improve the productivity of the process.
Table 4.
3.4.2. Temperature profile
TE D
Since adsorption is exothermic, every process with adsorption and desorption steps includes temperature variations. In such processes, temperature increases in adsorption stage and decreases in desorption stage. As shown in Figure 11, temperature is changed in both PVSA and PTSA process. As can be seen in Figure 11a and 11b, at the beginning of the adsorption step,
EP
the gas temperature rises to about 320 K because of adsorption of high amounts of carbon dioxide and heavy hydrocarbons, however after 2 hours, none of these components are
AC C
adsorbing. Therefore, the temperature reaches to the feed stream temperature. In Figure 11a, a sharp temperature decrease is observed after the end of adsorption time (18 hours). This is because of extreme pressure drop in blow down and evacuation step (6.8 MPa to 0.1 MPa) in the PVSA process. According to this figure, at the top of the bed, this decrement is more, since in these two steps, the gas stream flows from down to top. After that the bed temperature arises to the purge temperature of 350 K. the temperature varies from bottom to top of the bed, because the gas stream loses its temperature and becomes cooler along the bed. Figure 11.
ACCEPTED MANUSCRIPT
3.4.3. Process efficiency The performance of a cyclic adsorption process is commonly evaluated according to basic parameters of product purity, productivity and recovery defined, respectively by equations (2),
!
"##$ | ∑, % '() * ,(. %
RECOVERY =
*
*
*
6 : "##$ c89 u: |<=> dt − : $#BC#DD c89 u: |<=> dt − : BECF# c89 u: |<=> dtG AI
*
t W
*
SC
PRODUCTIVITY =
(2)
*
BECF# c89 u: |<=> dt : "##$ c89 u: |<=> dt − : $#BC#DD c89 u: |<=> dt − :
*
*
BC#DD c89 u: |<=: dt : "##$ c89 u: |<=: dt + :
M AN U
PURITY =
!
% "##$ % |'() *
RI PT
(3) and (4) as below.
(3)
(4)
The values of purity, productivity and recovery for three processes of PTSA, simple and improved PVSA are showed in Figure 12 [14]. According to the data reported in Figure 12, the purity of methane has increased by a value of 0.0299% using the PVSA processes. The incremental change of purity is not significant, whereas there are significant enhancement in the
TE D
values of productivity and recovery of the PVSA compared with PTSA. The value of productivity has been increased from 2.79[ PQR
PQR
ST.UVW
] in PTSA to 3.90[
PQR
ST.UVW
] in simple PVSA and
to a less greater value of 3.94[ST.UVW] in improved PVSA. The value of recovery has increased PVSA.
EP
from 74.04% in PTSA to 94.80% in simple PVSA and to a greater value of 98.27% in improved
AC C
The much less amount of PTSA recovery is due to the purge step in which excess amount of the natural gas is released to the atmosphere at high pressure, whereas in the PVSA process the purge step is carried out at low pressure. The impact of pressure equalization has been studied in improved PVSA on the NG recovery and this impact can be seen in comparison of the simple and improved PVSA processes. Pressure equalization has improved the recovery of the process by 3.47% from simple PVSA to improved PVSA. This enhancement is due to avoiding gas release during depressurization steps to the atmosphere. Figure. 12
ACCEPTED MANUSCRIPT
3.4.4. Economic study In order to compare three processes of simple PVSA, improved PVSA and PTSA from the operating cost point of view, some calculations have been done.
RI PT
The effective parameters of the operating cost for the PVSA process is included purge gas cost, gas stream cost for pressurization step, vacuum pump cost and purge temperature cost. Operating cost of PTSA process is influenced by the purge gas cost, gas stream cost for pressurization step and temperature change cost in heating and cooling steps.
SC
In order to calculate the operating costs, gas flow rate in each step, gas heat value (HV), vacuum pump power and efficiency, and price of the feed gas, product gas, electricity and fuel gas are
M AN U
required. In Table 5 and Figure 13, operating cost calculations for three processes are exhibited [29, 30].
Table 5.
Figure 13.
TE D
As can be seen in Figure 13, the operating cost of PTSA process is more than PVSA processes. According to Table 5, the main reason is the energy cost in heating steps (480K in 1st heating step and 590K in 2nd heating step) of PTSA process. Because of increasing the operating temperature (303K) to the high temperature and thereafter cooling down the gas consumes large
EP
amounts of energy. The difference of operating cost between simple and improved PVSA processes comes from pressure equalization steps in which the required gas for pressurizing the bed is supplied with the released gas in depressurizing steps, using improved PVSA cycle.
AC C
Therefore, a large amount of valuable gas can be saved and the operating cost can be decreased, significantly.
4. Conclusions
Process of cyclic adsorption based on PVSA was proposed as a more efficient process rather than PTSA process of South Pars Region of Iran. The most difference of the proposed PVSA process was according to the regeneration method. The improved PVSA was revealed to be more economic than the simple PVSA on the base of more energy saving.
ACCEPTED MANUSCRIPT
It was concluded that improved PVSA would successfully remove mercaptan to lower than 10 ppm with higher recovery and productivity than those in existing PTSA. At the same time, operating cost of the PVSA that is using relative vacuum for the bed regeneration is lower than the cost of PTSA that is using high gas temperature for regenerating the adsorbents. At the same
RI PT
time, cooling down the column at PTSA process takes much more time than PVSA.
Investigation of the basic performance parameters showed that the improved PVSA process has increased the recovery by 3.47% and productivity by 1% rather than the simple PVSA process.
SC
In addition, the improved PVSA process could increase the recovery and productivity by 24.23% and 44.8% respectively versus PTSA process. The economic study of the PTSA, simple PVSA and improved PVSA indicated that by using the PVSA process instead of PTSA process,
AC C
EP
TE D
M AN U
the operating cost would be decreased by 17 %.
ACCEPTED MANUSCRIPT
NOMENCLATURE Specific area of adsorbent (1/m)
b
Heat capacity equation parameter ( J/mol/K)
B:,]
Affinity constant at infinite temperature (1/KPa)
c
Heat capacity equation parameter (K)
C]
K Z,]
Concentration of adsorption of species i (mol/m_ )
k]
K d
Pore diffusivity of species i (m gs) e
Overall mass transfer coefficient (sqr ) Effective axial thermal conductivity (W/m/K)
k ′ Vapor thermal conductivity at atmospheric pressure (W/m K)
RI PT
aZ
M
Molecular weight (g/mol)
C]a Concentration of adsorption of species i in depressurization step (mol/m_ )
Nu
Nusselt number
n Number of the adsorbed components in the mixture
C]b Concentration of adsorption of species i in evacuation step (mol/m_ )
n Number of cycle
Concentration of adsorption of species i in C]c regeneration step (mol/m_ )
n
Number of all component in gas mixture
P
Pressure (KPa)
p]
Partial pressure (KPa)
M AN U
C]Z Concentration of adsorption of species i in pressurization step (mol/m_ ) CZ
Gas heat capacity (J/mol/K)
Average heat capacity of adsorbent (J/kg/K)
d
Heat capacity equation parameter ( J/mol/K)
TE D
CZ
D d,] Effective axial dispersion coefficient of species i e (m gs)
EP
D,] Molecular diffusivity of species i in mixture e (m gs) Dh,]
e Knudsen diffusivity of species i (m gs)
e Molecular diffusivity of species i in j (m gs)
e
Heat capacity equation parameter (K)
AC C
D],i
E Characteristic energy of the adsorbent (J/mol) h] Film heat transfer coefficient between gas and solid (W/me /K) HV
Gas heat value (BTU/m3)
∆H]
Heat of adsorption of species i (J/mol)
K
k o,]
SC
C]` Concentration of adsorption of species i in adsorption step (mol/m_ )
Gas thermal conductivity (W/m/K) Film mass transfer coefficient of species i (m/s)
q] Adsorbed phase concentration of species i (mol/kg)
q∗] (mol/kg)
Equilibrium adsorbed phase concentration
q,] (mol/kg)
Maximum adsorbed phase concentration
R
Ideal gas law constant (J/mol/K)
rZ
Adsorbent radius (m)
Re
Reynolds number and
Sc
Schmidt number
Sh
Sherwood number
t
Time (s)
T
Gas temperature (K)
T
Adsorbent temperature (K)
rZ Pore radius (m)
T
T
Reduced temperature Critical temperature (K)
ACCEPTED MANUSCRIPT
z{
y:
Superficial gas velocity (m/s)
µ
_
Critical molar volume (m /Kmol) Initial mole fraction
y]
Mole fraction of species i
z
Axial coordinate in the bed (m)
Shape factor
δ
Convergence parameter
η
Pump efficiency
Subscripts
εZ
Adsorbent porosity
τ
Tortuosity factor
ρ
kg Adsorbent density ( gm_ )
1. References
F H L A R D E P
Feed High pressure Low pressure Adsorption Regeneration Depressurization Evacuation Pressurization
M AN U
Bed porosity
Gas viscosity (kg/m.s)
φ
Greek letters ε
Gas density (molgm_ )
SC
u
ρ
RI PT
Tsurr Surroundings temperature
[1]. M.Tagliabue, D.Farrusseng, S. Valencia, S.Aguado, U.Ravon, C.Rizzo, A.Corma, C.Mirodato, Natural gas treating by selective adsorption: Material science and chemical engineering interplay, Chemical Engineering Journal 155 (2009) 553–566.
TE D
[2]. J.P. Bellat · F. Benoit · G. Weber · C. Paulin ·P. Mougin · M. Thomas, Adsorption equilibria of binary ethylmercaptan/hydrocarbon mixtures on a NaX zeolite, Adsorption (2008) 14: 501–507. [3]. H.Tamai. H.Nagoya. T.Shiono, Adsorption of methyl mercaptan on surface modified activated carbon, Journal of Colloid and Interface Science 300 (2006) 814–817.
EP
[4]. A. de Angelis, Natural gas removal of hydrogen sulphide and mercaptans, Applied Catalysis B: Environmental 113– 114 (2012) 37– 42. [5]. G.Weber , F. Benoit , J-P Bellat , C. Paulin , P. Mougin ,M.Thomas, Selective adsorption of ethyl mercaptan on NaX zeolite, Microporous Mesoporous Mat. 109 (2008) 184–192.
AC C
[6]. S.Cavenati, Carlos A. Grande, Alírio E. Rodrigues, Separation of CH4/CO2/N2 mixtures by layered pressure swing adsorption for upgrade of natural gas, Chemical Engineering Science 61 (2006) 3893 – 3906. [7]. T.L.P. Dantasa, F.M.T. Lunab, I.J. Silva Jr, A.E.B. Torres, Carbon dioxide–nitrogen separation through pressure swing adsorption, Chemical Engineering Journal 172 (2011) 698– 704. [8]. V.P. Mulgundmath, R.A. Jones, F.H. Tezel, J. Thibault, Fixed bed adsorption for the removal of carbon dioxide from nitrogen: Breakthrough behavior and modeling for heat and mass transfer, Separation and Purification Technology 85 (2012) 17–27. [9]. M. Clausse, J. Bonjour1, F. Meunier, Adsorption of gas mixtures in TSA absorbers under various heat removal conditions, Chemical Engineering Science 59 (2004) 3657 – 3670.
ACCEPTED MANUSCRIPT
[10]. J. Zhang, P.Xiao, G.Li, P.A. Webley, Effect of Flue Gas Impurities on CO2 Capture Performance from Flue Gas at Coal-fired Power Stations by Vacuum Swing Adsorption, Energy Procedia 1 (2009) 1115–1122.
RI PT
[11]. J. Zhang, P.A. Webley , P.Xiao, Effect of process parameters on power requirements of vacuum swing adsorption technology for CO2 capture from flue gas, Energy Conversion and Management 49 (2008) 346–356. [12]. M.C. Campo, A.M. Ribeiro, A. Ferreira, J.C. Santos, C. Lutz, J.M. Loureiro, A.E. Rodrigues, New 13X zeolite for propylene/propane separation by vacuum swing adsorption, Separation and Purification Technology 103 (2013) 60–70.
SC
[13]. B. Shirani, T. Kaghazchi, and M.Beheshti, Water and mercaptan adsorption on 13X zeolite in natural gas purification process, Korean J. Chem. Eng., 27(1), (2010) 253-260.
M AN U
[14]. O.T. Qazvini, S. Fatemi, Modeling and Simulation Pressure–Temperature Swing Adsorption Process to Remove Mercaptan from Humid Natural Gas; A Commercial Case Study, Separation and Purification Technology (2014). [15]. D. Ferreira and R. Magalhães, Effective adsorption equilibrium isotherms and breakthroughs of water vapor and carbon dioxide on different adsorbents, Industrial & Engineering Chemistry Research, (2011) 10201–10210. [16]. R. W. Glass and R. A. Ross, Surface Studies of the Adsorption of Sulfur-Containing Gases at 423°K on Porous Adsorbents. I I. The Adsorption of Hydrogen Sulfide, Methanethiol, Et hanethiol, and Dimethyl Sulfide on y-Alumina, The Journal of Physical Chemistry, Vol. 77, No; 27, (1973).
TE D
[17]. E. Baumgarten, F. Weinstrauch, and H. Hoffkes, Adsorption isotherms of hydrocarbons on UAlumina, Journal of chromatography 138, (1977) 347–354. [18]. D. Green and R. Perry, Perry’s chemical engineers' handbook, 8th Ed. New York: McGraw-Hill, (2007).
EP
[19]. M. Mehdipour and S. Fatemi, Modeling of a PSA-TSA process for separation of CH4 from C2 products of OCM reaction, Separation Science and Technology, (2012) 1199–1212. [20]. R. Desai, M. Hussain and M. Ruthven, Adsorption of Water Vapour on Activated Alumina. I Equilibrium Behaviour, the Canadian jurnal of chemical engineering (1992) 699-706,.
AC C
[21]. D.M. Ruthven., Principles of adsorption and adsorption processes, John Wiley, New York (1974). [22]. N. Wakao, T. Funazkri, Effect of fluid dispersion coefficients on particle-to-fluid mass transfer coefficients in packed beds, Chem. Eng. Sci. 33, (1978) 1375–1384, [23]. .R. Welty, C.E. Wicks, R.E. Wilson, G. Rorrer, Fundamentals of Momentum, Heat and Mass Transfer, John Wiley and Sons, New York (2000). [24]. R. T. Yang, Gas separation by adsorption processes. Boston: World Scientific Publishing Company, (1987). [25]. R. Byron Bird, W. E. Stewart, E. N. Lightfoot, Transport Phenomena, JOHN WILEY & SONS, ING. New York Chichester Weinheim, Brisbane Singapore Toronto.
ACCEPTED MANUSCRIPT
[26]. Da Silva and A. E. Rodrigues, Adsorption equilibria and kinetics for propylene and propane over 13X and 4A zeolite pellets, Industrial & Engineering Chemistry Research 38 (1999) 2051–2057. [27]. K. Iusalaas, Numerical methods in engineering with MATLAB. Cambridge: Cambridge University Press (2005).
RI PT
[28]. G. G.Vining, Statistical methods for engineers. Pacific Grove: Duxbury Press (1997). [29]. Exclusive price list of oil, gas and energy facilities, Ministry of Petroleum, Engineering Department, 058 (2014).
[30]. M. Peters, K. Timmerhaus and R. West, Plant Design and Economics for Chemical Engineers, 5th
SC
edition (2002).
M AN U
APPENDIX A1. The physical properties of the adsorbents
The physical properties of adsorbents and the bed Characteristics are given in Table A1. Table A1.
TE D
A.2.Molecular diffusivity and physical properties prediction and correlation Molecular diffusivity of each component in the gas mixture is calculated by the equation (A.1) [18]:
1 − (. 1) ∑ = ,
EP
P, =
AC C
In this equation, is the mole fraction of the ith component and n is the number total
components in the gas mixture. For the mixture of natural gas, , can be assumed as , . , is the molecular diffusivity of binary gas mixture and is calculated by the Wilke-Lee
equation [18].
The amount of molecular diffusivity of each component through the mixture at the adsorption, evacuation and purge is brought in Table A2. Table A2.
ACCEPTED MANUSCRIPT
All physical properties are considered as methane property. Because methane is the dominate component of the gas phase. Pure gas specific heat capacity T has been calculated as the equation (A.2) [19]: e 6dZ ( ) –dZ (q ) G
+ ×
e (. 2) 6dZ ( ) dZ (q ) G
Table A3. A.3. Mass transfer and heat transfer correlations
SC
The coefficients of the above correlation are listed in Table A3.
RI PT
{
T = + ×
M AN U
The transport parameters needed in the model were calculated employing frequently used correlations. The correlations are used for the mass and energy balances in gas and solid phase are shown in Table A4.
Table A4.
TE D
A4. The initial and boundary conditions for numerical calculation As stated in numerical calculation section, to obtain the results of PVSA simulation, several non-linear equations should be solved, simultaneously. The initial and boundary conditions for the numerical solution for simple PVSA process are reported in Table A5. The initial and
Table A5.
AC C
EP
boundary condition of PTSA process is described in the paper of Taheri and Fatemi [14].
Where the subscripts F, H, L, A, R, D, E and P represent the feed, high pressure, low pressure, adsorption, regeneration, depressurization, evacuation and pressurization, respectively and L is the length of bed.
A.5. Mass transfer and heat transfer parameters for kinetic adsorption The average transfer parameters of the PVSA simulation were found based on the correlations are listed in Table A4. The parameters in different steps of the process are reported in Table A6.
ACCEPTED MANUSCRIPT
The amounts of , that are reported in the Table A6, are the overall heat transfer coefficient h
which is dependent of h] and h by equation (A.3). The amount of h is 10(.h).
h,
=
1 1 + (. 3) h] h
RI PT
1
Table A6. Figure Captions:
PTSA, b: simple PVSA, c: improved PVSA.
SC
Figure 1. Schematic of the step cycle sequences, in three processes for mercaptan removing, a:
M AN U
Figure 2. Operating bed pressure with elapsed time in (a): PTSA cycle, (b): PVSA cycle. Figure 3. The time-table of the improved PVSA process, at each column for a six- bed plant. Figure 4. Mercaptan loading on the top of 13X layer at the mentioned processes. Figure 5. Methane purity after evacuation stage, with the simulation cycle number in a simple
TE D
PVSA process with the inlet mole fraction of 95.9848% methane.
Figure 6. a: Mercaptan mole fraction at the outlet stream, b: solid temperature profile as a function of cycle number, for the simple PVSA process.
EP
Figure 7. Comparison of simulated and real values of mercaptan and water vapor mole fractions at the outlet stream.
AC C
Fig. 8. Simulated mercaptan breakthrough curve in comparison with real data. Figure 9. Breakthrough curves in the simple PVSA process, (a) Mercaptan, (b) water vapor, (c) carbon dioxide and (d) propane Figure 10. Loading values versus cycle time at PVSA and PTSA cycles at the top of 13X layer, (a) Mercaptan, (b) water vapor. Figure 11. Temperature profile along the bed at (a) simple PVSA process, (b) PTSA process. Figure 12. Basic performance parameters for PTSA, simple PVSA and improved PVSA.
ACCEPTED MANUSCRIPT
Figure 13. Operating cost for PTSA, simple PVSA and improved PVSA processes. Table Captions:
RI PT
Table1. Feed specifications and the bed Characteristics. Table2. Mathematical model of the dynamic adsorption including, total mass, mole and energy balances.
Table3. Parameters of Extended Langmuir Isotherm for Activated alumina and 13X.
SC
Table4. Comparison of the cyclic processes from the point of mercaptan and water desorption
Table5. Operating cost calculations.
M AN U
time and amount.
Table A1. Physical properties of the adsorbents.
Table A2. Average molecular diffusivity of gases in mixtureP, ( ¡e /¢).
TE D
Table A3. Heat capacity coefficients of methane.
Table A4. Mass and heat transfer correlations to solve the simulation equations. Table A5. The initial conditions for numerical calculation employed at simple PVSA
EP
simulation.
AC C
Table A6. Model parameters of the adsorption kinetics.
AC C
EP
TE D
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
Figure 1
AC C
EP
TE D
M AN U
Figure 2
SC
RI PT
ACCEPTED MANUSCRIPT
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
Figure 3
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
Figure 4
ACCEPTED MANUSCRIPT
94.8032
94.8028
Methane purity
94.8026
RI PT
Methane purity (%)
94.803
94.8024 94.8022 94.802
94.8016 2
4
6
8 10 Cycle number
12
M AN U
0
SC
94.8018
AC C
EP
TE D
Figure 5
14
16
ACCEPTED MANUSCRIPT
(a) 4.5 cycle 1 cycle 2 3.5
cycle 3
3
cycle 4
RI PT
cycle 5 2.5
cycle 6 cycle 7
2
cycle 8
1.5
cycle 9 1
cycle 10
0.5
cycle 11
SC
Mercaptan mole fraction (ppmv)
4
cycle 12 0
2
4
6
(b)370
8
10 Time (hr)
12
14
16
310
270
20
cycle 2 cycle 3 cycle 4 cycle 5 cycle 6 cycle 7 cycle 8 cycle 9
EP
290
18
cycle 1
TE D
330
AC C
Temperature (k)
350
M AN U
0
cycle 10 cycle 11 cycle 12
250
0
5
10
15
20 Time (hr)
Figure 6
25
30
35
40
ACCEPTED MANUSCRIPT
4.5 4
3.81
3.87 (*10-6 )
3.5 real value
3 2.5
RI PT
simulated value
2 1.5 1 0.5 0
Mercaptan mole fraction
SC
0.51 0.55 (*10-7)
Water vapor mole fraction
AC C
EP
TE D
M AN U
Figure 7
ACCEPTED MANUSCRIPT
simulated result
60
real data
50 40
RI PT
Mercaptan mole fraction (ppm)
70
30 20 10
0
5
10
15
Time (hr)
20
25
AC C
EP
TE D
M AN U
Figure 8
SC
0
30
ACCEPTED MANUSCRIPT
(b)
160 140
80
10
60 40 20 0
5 0
8
10 12 14
Time (hr)
16 18
0
4
6
8
10
12
14
16
18
Time (hr)
(d) z=0 (13X bottom) z=2.3 (13X middle) z=4.7 (alumina bottom) z=5 (alumina middle)
z=0 (13X bottom) z=2.3 (13X middle) z=4.6 (13X top)
0
6000
5000
10000
M AN U
15000
2
SC
6
12000 18000 24000 30000
4
20000
2
RI PT
15
100
0 0
0.5
1 Time (hr)
1.5
2
EP
Figure 9
AC C
CO2 mole fraction (ppm)
z=0 (13X bottom) z=2.3 (13X middle) z=4.7(alumina bottom) z=5(alumina middle)
20
120
0
(c )
25 Water vapor mole fraction (ppm)
z=0 (13X bottom) z=2.3 (13X middle) z=4.65 (13X Top)
Heavy hydrocarbons C3+ (ppm)
Mercaptan mole fractio (ppm)
180
TE D
(a)
0
0.2
0.4 0.6 Time (hr)
0.8
1
AC C
EP
TE D
M AN U
Figure 10
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
M AN U
Figure 11
SC
RI PT
ACCEPTED MANUSCRIPT
ACCEPTED MANUSCRIPT
120 96.10 96.13 96.13
94.80
98.27
74.04
80
PTSA SIMPLE PVSA IMPROVED PVSA
60 40 20
RI PT
100
2.79 3.90 3.94
% Purity
% Recovery
AC C
EP
TE D
M AN U
Figure 12
Productivity (mol/kg.day)
SC
0
ACCEPTED MANUSCRIPT
88
90 80
72.94
72.705
Simple PVSA
Improved PVSA
70 60
RI PT
50 40 30 20 10 0 PTSA
AC C
EP
TE D
M AN U
Figure 13
SC
operating cost (Thousand $/year)
100
ACCEPTED MANUSCRIPT
Table 1. Parameter Flow rate (Nm3/hr)
Value 2850
Bed Bed length (m)
13X 4.65
AC 0.75
Pressure (Mpa)
6.8
Bed diameter (m)
3.7
3.7
Temperature (k)
302
Bed void fraction
0.37
0.26
CO fraction (ppmv)
Heavy hydrocarbons fraction (ppmv)
3.0 × 10
95.9848 × 10
0.82
Adsorbent weight (ton)
21.723
4.89
AC C
EP
TE D
M AN U
Methane fraction (ppmv)
1.0 × 10
0.69
RI PT
133.8
Mercaptan fraction (ppmv)
Bulk density (g/ )
SC
18.2
Water vapor fraction(ppmv)
ACCEPTED MANUSCRIPT
Table2.
Overall mass balance in gas phase Solid components’ mass balance Gas phase energy balance
' (( ) = , – (1 − )%& − *+ (() ' =− − (1 − )%& )( ) , ' = ('∗ − ' ) / / / . − (%01 − % 01 − 20 (1 − )ℎ (/ − /& ) − ℎ&455 (/ − /&455 ) = 0
*+
RI PT
Gas components’ mass balance
Energy balance of the solid
8,
AC C
EP
TE D
Extended Langmuir isotherm
(1 − )( : 37.5 (1 − ) =( = −; + 0.875% B 1 ?0 @ >?0 @A ∆7 EF, exp J L M '∗ K/ = 'CD 1 + ∑*O E exp J∆7 L M 8, F, K/
M AN U
Ergun equation
SC
/& ' %& 0& = %& )(−∆7 ) + 20 ℎ (/ − /& )
ACCEPTED MANUSCRIPT
Table 3. PCQR ) P1
, ) T0
Water vapor Carbon dioxide
0.0151 0.0011
3.22 × 10VW 1.94 × 10V
0.003580 0.016842 0.005115 0.002821
4.32 × 10VX 4.82 × 10VW 3.61 × 10V 3.36 × 10V
- ∆7 (
AC C
EP
TE D
M AN U
SC
Zeolite 13X Mercaptan Water vapor Carbon dioxide Propane
PU ) CQR
SF, (
60.014 29.805
RI PT
'C, (
Activated alumina
65.1624 62.5258 34.0031 32.5412
ACCEPTED MANUSCRIPT
Table 4. Improved PVSA
PTSA
Mercaptan time for complete desorption (hr)
6
6
12
Water vapor time for complete desorption (hr)
4
4
7
Mercaptan loading at top of the bed (Kmol/g)
0.000033
Water vapor loading at top of the bed (Kmol/g)
0.00028
RI PT
Simple PVSA
0.019
0.00028
0.015
AC C
EP
TE D
M AN U
SC
0.000033
ACCEPTED MANUSCRIPT
Table 5.
Calculations
PTSA
$ Z[p OO4C ^ a `
$ Z[ot\5C R ^ a ` $ /Z2e Z[ ^ a `
= d(be deZf ?2b ^
$ a × M?ckbl4\R 1 & ^ a `
AC C
470
235
15270
15270
15270
-
44000
44000
72000
13200
13200
88000
72940
72705
Wpump: calculated by thermodynamic correlation for isentropic work=250 KW HV: Heat Value of gas= 1/29 ηvwxv:pump efficiency= 0.7
TE D
PTSA pressurize flow rate=106000 Simple PVSA pressurize flow rate=68000 Improved PVSA pressurize flow rate=34000 purge flow rate= 1476000 PTSA consumed fuel flow rate=175000 PVSA consumed fuel flow rate=32200
EP
pricefeed= 0.2 priceproduct= 0.3 priceelectricity= 0.14 pricefuel gas= 0.41
q04C0 $ 7 =^ a × M?ckb\R\Oo5Oos ^ a × 24 × 366 h j r04C0 .q7 `
M AN U
$ Z[0451\ ^ a `
S/i = M(?nb deZf ?2b ^ a × 7g h j ` $ × M?ckb05Qm4Oo ^ a S/i
Improved PVSA
730
SC
S/i = M?b[[(?c b deZf ?2b ^ a × 7g h j ` $ Z[05\&&45]\ ^ a $ ` × M?ckbl\\m ^ a S/i
Simple PVSA
RI PT
Cost
ACCEPTED MANUSCRIPT
Table A1. Physical properties of adsorbents
Zeolite 13X
Activated Alumina
2.1
2.3
Average macropore radius (nm)
2.44
4.6
445.6
441.3
BET surface area (m /g)
Particle void fraction
0.24
1.10
Heat capacity (kJ/kg.K)
1.07
Tortuosity
1.38
0.28
1.24
3.85
1.36
AC C
EP
TE D
M AN U
SC
Particle density (g/m )
RI PT
Particle radius (mm)
ACCEPTED MANUSCRIPT
Table A2. Adsorption
2.238 × 10V
Carbon dioxide Heavy hydrocarbons Water vapor
Purge
1.6
3 × 10V
2.59 × 10V
1.1988× 10V
0.16
0.95
9.5× 10V
3.4
0.34
0.79
AC C
EP
TE D
M AN U
SC
Mercaptan
Evacuation
RI PT
Stage
7.9 × 10V
ACCEPTED MANUSCRIPT
Table A3.
(z. Ze V, . . V, ) E × 10V{ (z. Ze V, . . V, ) k × 10V (. ) × 10V (z. Ze V, . . V, ) b × 10V (. ) 2 × 10V
0.7993
2.0869
4.16
9.92
EP
TE D
M AN U
SC
RI PT
3.33
AC C
7
ACCEPTED MANUSCRIPT
The effective axial bed thermal conductivity The effective mass transfer coefficient [21]
External film resistance( Wakao and Funazkri) [22]
The axial dispersion coefficient [23]
Tortuosity Factor [11]
TE D EP AC C
= 0 + 1.5>1 − 0 A / ~, = 97?0Q5\ ( )F.{
M AN U
Knudsen diffusion coefficient [24] Chilton–Colburn analogy [25]
,/
ℎ = 2 + 1.1k Kb F. , = 20 + 0.5 k Kb C, 1 1 1 = ^ + a .0, ~, C,
SC
The macropore diffusion coefficient
. = 7 + 0.5:?Kb .1 ?0 ?0 1 ?O = + + }~ O, 3l, 15ɛ .0, 15. [ℎ C, l, = 2?0
RI PT
Table A4.
,/
( = 2.0 + 1.1:?
Kb F.
ACCEPTED MANUSCRIPT
Table A5. Adsorption stage:
RI PT
= 0:
= ,*+ + ' = ',*+ + ,
: = : = 6.8 :2 , / = / = 302 K / ( ) = 0 , = 0
( ) ( > − A , = , ,
TE D
/ (k01 : (/ − / ) , : = : , ( = ( = . K/
EP
Depressurization stage:
AC C
= = 5.4 :
M AN U
SC
= 0:
ACCEPTED MANUSCRIPT
1&o bM?b[[(?c b [bM:
= *+ + , ' = '*+ + ,
: − : / = /*+ + , : = − + : cbm\05\&& = 0:
: = 6.8 :2, : = 4.0 :2
2*m bM?b[[(?c b [bM:
RI PT
= 0:
: = 4.0 :2, : = 1.5 :2
( ) = 0 ,
35m bM?b[[(?c b [bM:
SC
/ = 0 , ( = 0 = = 5.4 :
: = 1.5 :2,
M AN U
( ) / = 0 , = 0 , P = P 8
: = 0.1 :2
cbm\05\&& = 10 c
Countercurrent evacuation stage: = 0:
TE D
= *+ + , ' = '*+ + ,
/ = /*+ + P = P = 0.01 :2 = 0:
= = 5.4:
EP
( ) / = 0 , = 0 , ( = 0
AC C
( ) / = 0 , = 0
Countercurrent regeneration stage:
ACCEPTED MANUSCRIPT
= 0:
= 0:
( ) ( / (k01 : ( − 0), (/ − / ), =− = , . K/
( = ( , P = P = 0.1 :2 = = 5.4 :
Co-current pressurization:
M AN U
SC
( ) / = 0 , = 0
TE D
= 0: = ,*+ + V, , ' = ',*+ + V, : − : / = /,*+ + V, , : = + : cb05\&& = 0:
/ ( ) = 0 , = 0 , ( = 0
EP
= = 5.4 :
AC C
( ) / = 0 , = 0
RI PT
= ,*+ + , ' = ',*+ + , / = /,*+ +
1&o M?b[[(?c b [bM:
: = 1.5 :2, : = 0.1 :2
2*m M?b[[(?c b [bM:
: = 4.0 :2, : = 1.5 :2
35m bM?b[[(?c b [bM:
: = 6.8 :2, : = 4.0 :2
cb05\&& = 10 c
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
Zeolite 13X
Activated alumina
Water vapor
0.00098
0.00014
EP
(1/s)
Mercaptan
0.00093
-
Carbon dioxide
0.00141
0.008205
Heavy hydrocarbons
0.00731
-
AC C
Step
TE D
Table A6.
adsorption , (
¡
)
Water vapor
1.7028
2.4177
Mercaptan
1.7031
2.4182
Carbon dioxide
1.7021
2.4171
Heavy hydrocarbons
1.7025
2.4172
5.848
5.941
3.1621
4.1246
£ ) .¢
¢, (
£
¤¥, ( .¢)
ACCEPTED MANUSCRIPT
0.0118
Mercaptan
0.0043
Carbon dioxide
0.00084
Heavy hydrocarbons
0.0008
¡
)
Water vapor
2.148
Mercaptan
2.186
Carbon dioxide
2.175
Heavy hydrocarbons
2.166
£
¢, (
.¢
)
£ ) .¢
¤¥, (
(1/s)
Water vapor
0.0592
Carbon dioxide Heavy hydrocarbons
-
3.493
3.528
0.0057
0.0126
3.537
3.529 0.35
0.3931
0.051 1.1369 -
TE D
)
8.843
13.217
Mercaptan
8.859
13.223
EP
¡
0.0072
Water vapor
Carbon dioxide
8.854
13.219
Heavy hydrocarbons
8.820
13.216
0.4548
0.4564
£
0.5256
0.6499
£ ) .¢
¢, (
¤¥, ( .¢)
AC C
purge
0.2135
0.0311
Mercaptan
, (
0.35
-
M AN U
Evacuation
, (
0.00142
RI PT
Water vapor
SC
(1/s)
ACCEPTED MANUSCRIPT
Research highlights:
EP
TE D
M AN U
SC
RI PT
Mercaptan removing from natural gas. Vacuum Pressure Swing Adsorption process for methane purification. Process performance parameters comparison and economic study for PVSA and PTSA processes.
AC C
• • •
S1875-5100(15)30037-8
DOI:
10.1016/j.jngse.2015.07.010
Reference:
JNGSE 868
To appear in:
Journal of Natural Gas Science and Engineering
Received Date: 26 May 2015 Revised Date:
7 July 2015
Accepted Date: 8 July 2015
Please cite this article as: Tohidi, Z., Fatemi, S., Qazvini, O.T., Mercaptan Removal from Natural Gas by the Efficient Cyclic Adsorption Process; a Simulation Study, Journal of Natural Gas Science & Engineering (2015), doi: 10.1016/j.jngse.2015.07.010. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
AC C
EP
TE D
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
ACCEPTED MANUSCRIPT
Mercaptan Removal from Natural Gas by the Efficient Cyclic Adsorption Process; a Simulation Study Zahra Tohidi, Shohreh Fatemi*1and Omid Taheri Qazvini
RI PT
School of Chemical Engineering, Faculty of Engineering, University of Tehran, Tehran, Iran Abstract
A more efficient and economical cyclic adsorption process was proposed for mercaptan removal
SC
from natural gas (NG) to reduce mercaptan content to less than 10 ppm and meet the environmental rules. Continuous sulfur removal is studied for the NG feed stream, with pressure of 6.8MPa, flow rate of 2850 Nm3/hr and molar composition of 95.98% methane, 0.00182%
M AN U
water vapor, 1% carbon dioxide, 0.0134% mercaptan and 3% heavier hydrocarbons (C3+). The proposed process of Pressure Vacuum Swing Adsorption (PVSA) was designed and simulated as a more efficient alternative process against the current Industrial Pressure-Temperature Swing Adsorption (PTSA). In this work, an improved PVSA process was simulated with sequences of bed pressurization, adsorption, equalization, blow down, bed regeneration by vacuum and purge by product, in each process cycle. Vacuum condition of 10 KPa with the
TE D
molar purge/feed ratio of 0.06 and temperature of 350K was required for appropriate bed regeneration from adsorbed mercaptan to approach to the continuous cyclic steady condition. Comparison between PVSA and PTSA, at the same feed characteristics, same packed columns
EP
and adsorption operating conditions, revealed that the PVSA process, with less cycle time than PTSA, could achieve the same product purity with 94.8%
productivity, whereas PTSA has the recovery of 74.04% and productivity of 2.79
AC C
recovery and 3.90.
.
. At the
same time, operating with PVSA, instead of PTSA process, would reduce the operating cost from 88 to 70
$
.
Keywords: Cyclic Adsorption Processes, Mercaptan Removal Unit, Natural Gas, PVSA Process, PTSA Process, Economic Study
* Corresponding author, email: [email protected], school of chemical engineering, Faculty of Engineering, University of Tehran, Enghelab Avenue.
ACCEPTED MANUSCRIPT
1. Introduction The progress of the international energy demand shows an average of 1.7% annual growth in the 2005 to 2020. This growth concerns all energy sources and natural gas demands which
RI PT
would be accounted for the highest growth rate in 2020 [1]. Increasing concerns of the harmful effects of natural gas contaminants on environment has led to the introduction of a number of natural gas treatments. Out coming NG from the well contains methane with impurities such as water vapor, carbon dioxide, nitrogen, hydrogen sulfide, light mercaptans, ethane and heavier
SC
hydrocarbons. Some part of the natural gas impurities must be removed before commercial use. Today, according to the recent environmental legislations, the sulfur emission in the atmosphere should be considerably reduced to less than 20 ppm [2]. Mercaptans, or more correctly thiols,
M AN U
are organic compounds in which the -SH groups are present in the molecular structure of the hydrocarbons. The mercaptans must be removed mainly for three reasons: (a) they have acidic property and can cause serious corrosion problems, (b) they have offensive odor and they are very inappropriate to be burnt, (c) most of them are highly toxic and affect the subsequent catalytic reactions [3]. In the conventional processes, the acid gases such as hydrogen sulfide and carbon dioxide are mostly removed in an amine-wash unit. Since light mercaptans are not
TE D
as acidic as hydrogen sulfide, they cannot be removed properly by amine washing and an additional step is required to reduce the sulphur concentration to an appropriate level. Gas purification by the adsorption process could be as an alternative process in progress for sulfur removal from NG [2]. Modeling and simulation of the gas adsorption processes have been
EP
investigated previously by some researchers. Some limited works have been done for investigating mercaptan removal by adsorption using different adsorbents [2, 5-12]. Shirani et
AC C
al. [13] also simulated mercaptan adsorption on 13X adsorbent in presence of water vapor at isothermal condition and presented the breakthrough curves from the packed bed. In another work, Taheri and Fatemi [14] simulated a PTSA (Pressure-temperature swing adsorption) unit for mercaptan removal in South Pars region. However, the high cycle time taking for cooling down and the high economic costs are the main disadvantages of the mentioned process. It believes that, by replacing the current PTSA process with PVSA (Pressure-vacuum swing adsorption), the efficiency of the process would be improved. In addition, in PTSA process, the temperature shock induced to the solid adsorbents by temperature variation in adsorptiondesorption cycles would reduce the life time of the adsorbents, whereas PVSA doesn’t require
ACCEPTED MANUSCRIPT
energy induction to the adsorbents. It should be noted that, there is no evidence in the literature about mercaptan removal by PVSA from NG in presence of other impurities such as CO2, water vapor and heavier hydrocarbons. In this work, a systematic simulation of mercaptan removal from NG in presence of other impurities is carried out by the PVSA process in face of the
RI PT
industrial PTSA process simulated by Taheri and Fatemi [14] with the same bed configurations and design. This comparison is carried out on the base of different regeneration conditions and the results are presented in terms of concentration and temperature profiles, and performance parameters such as purity, productivity and recovery as well as a brief economical study. In this
SC
research, industrial PTSA refers to the real mercaptan removal unit (MRU) in South Pars region of Iran that is working currently, and the process design parameters and the outlet data are taken
2. Process Description and Design 2.1.Description of process
M AN U
from there[14].
A systematic adsorption model has been presented and implemented to simulate the PVSA process proposed for purification of natural gas (NG) from light mercaptans in presence of other
TE D
impurities such as water vapor, CO2 and heavier hydrocarbons (C3+). The industrial mercaptan removal unit (MRU) is currently working in South Pars region of Iran and this process is designed in the mode of PTSA. This unit consists of six insulated two-layer packed bed columns in which the regeneration step is carried out by increasing temperature of the bed by
EP
the hot gas NG. The pressure of the upcoming feed is 6.8 MPa and the adsorption process has been initially designed to work at this high pressure. The properties of the inlet NG is specified
AC C
at Table 1. In addition, the adsorption bed dimensions in industrial scale and the type and amount of the adsorbents are presented in this table. Table 1.
Each column consists of two layers: 13X zeolite and activated alumina. The 13X zeolite is used for major adsorption of mercaptan, however because of high affinity of 13X to water vapor, water plays competitive role in mercaptan adsorption. Therefore, a pre-layer of Activated alumina is required to adsorb water vapor before entering the 13X layer. Adsorption of
ACCEPTED MANUSCRIPT
mercaptan and C3+ is ignorable compared to H2O and CO2 adsorption by activated alumina. [15, 17]. Physical properties of the adsorbents are reported in Appendix A1. In this work, the process of PVSA is designed on the base of existing PTSA, but different
RI PT
according to the regeneration step and operating conditions. A simple PVSA with no equalization and an improved PVSA with pressure equalization steps are planned to study the impact of pressure equalization step on purity, recovery and productivity of the process. Also, the results of proposed PVSA processes would be compared with the industrial PTSA process.
SC
These results would be comparable with the established industrial PTSA that is currently working with no pressure equalization. The current PTSA process consists from regeneration
M AN U
step including two heating stages; 1st stage working at 480 K and the 2nd one working at 590 K. The proposed simple PVSA process is designed to include 10 steps as following: I.
Adsorption step; during 18 hours, the NG feed is introduced from top of the bed with a flow rate of 2.2 Kmol/s at pressure of 6.8 MPa and temperature of 302 K. 1st depressurizing step; the bed pressure is decreased for 6.8 to 4 MPa during 10
II.
III.
TE D
minutes. The gas content of the bed is released to the atmosphere, from the top. 2nd depressurizing step; the bed pressure is decreased from 4 to 1.5 MPa during 10 minutes by releasing from the top of the bed to atmosphere. Blow down step; the rest of the bed pressure is reduced down to the atmospheric
EP
IV.
pressure. 10 min time is considered for this step. Evacuation step; the bed pressure is evacuated by a vacuum pump to reach to 0.01
AC C
V.
MPa during 10 minutes.
VI.
1st purge step; the bed is purged counter-currently with 2% of the product stream at the vacuum pressure of 0.01 MPa during 10 minutes.
VII.
2nd purge step; the bed is purged counter-currently with 6% of the product stream in atmospheric pressure and temperature of 350 K during 15 hours and 40 minutes.
ACCEPTED MANUSCRIPT
VIII.
1st pressurizing step; the bed pressure is increased up to 1.5 MPa through feed introducing from top of the bed during 10 minutes.
IX.
2nd pressurizing step; the bed pressure is increased from 1.5 to 4 MPa during 10
RI PT
minutes. 3rd pressurizing step; the bed pressure is increased from 4 to 6.8 MPa during 10
X.
minutes.
Rest; the bed is put on the rest for one hour to be prepared for the next cycle.
SC
XI.
The improved PVSA process was designed as the alternative of simple PVSA with the same
M AN U
conditions but considering pressure equalization steps between the columns. Pressure equalization step provides inter-connection between two columns in which one of the columns is working at depressurizing mode and the other one is at repressurizing mode. According to the above modes, step III is inter-connected to step VII, and the step II is inter-connected to step IX. The improved PVSA can be used to reduce natural gas consumption and to enhance process efficiency. In Figures 1(a), 1(b) and 1(c) the steps’ arrangement are exhibited for industrial
TE D
PTSA, simple PVSA and improved PVSA, respectively. Figure 1.
Figure 2 presents the operating pressure of the manipulated steps in the PTSA and proposed
EP
PVSA processes. In order to manipulate and control the steps of PVSA process with equalization, the adsorption time was set to 17.5 hour and the time of desorption steps (evacuation and purge) was set to 16.5 hour. This is done because pressurizing steps should be
AC C
at the first half time of the cycle in order to equalize with the depressurizing steps at the second half time of the cycle. In this situation, we could save more natural gas and enhance the process performance. Figure 3 shows the time-table of the process with equalization steps. Figure 2. Figure 3. 2.2.Model and assumptions
ACCEPTED MANUSCRIPT
A mathematical model with mass, energy and momentum balances that represents the dynamic behavior of the multicomponent adsorption bed was developed based on the following assumptions: Radial gradients for mass, velocity and temperature are neglected in the bed.
•
The bed fluid is assumed dispersed plug flow.
•
Porosity along the bed is constant with uniform packing.
•
Mass transfer to the particles is described by linear driving force (LDF) model with solid
SC
RI PT
•
phase concentration.
Diffusion and adsorption into the particles are assumed as lump kinetic transfer model.
•
There are no concentration and temperature gradients inside the particles.
•
Pressure drop along the bed is predicted by Ergun equation.
•
Gas behavior follows Peng Robinson equation of state throughout the process.
•
Adsorption equilibrium is described by Extended Langmuir Isotherm model for
TE D
M AN U
•
multicomponent adsorption, in which any isotherm parameter is temperature dependent. The flow velocity varies along the bed and is calculated by the total mass balance equation. •
EP
•
The column operates in non-isothermal and non-adiabatic conditions in which the heat is
•
AC C
exchanged with surrounding.
Temperature of fluid and solid phase is not identical and heat can be transferred between solid and fluid.
•
The solid temperature is homogeneous and a function of time during adsorption and desorption process.
•
In heat transfer to the environment, the wall resistance follows thin wall assumption.
ACCEPTED MANUSCRIPT
•
Heat capacities of the adsorbents and adsorbed phase are assumed constant.
•
Heat of adsorption is assumed constant and derived from the isotherm equation for each component.
RI PT
According to the assumptions, mass and energy balances of the gas and solid phases are shown in Table 2. Table 2.
are reported in Table 3.
M AN U
Table 3.
SC
The isotherm parameters of the components were found from literature [5, 15, 20, 21] and they
The correlations for physical properties, mass transfer and energy transfer coefficient are reported in Appendix A2 and A3, respectively. 2.3.Numerical Calculations
TE D
To obtain temperature and concentration profiles inside and at the exit of the bed versus time at each cycle, model equations should be solved, simultaneously. Because of complexity, these nonlinear equations cannot be solved in analytical form and numerical solutions should be used. The numerical algorithm implemented was finite difference scheme with the first order Upwind
EP
Differencing Scheme (UDS1) method to convert spatial derivatives of the PDEs to the discrete form and convert the PDEs to the set of ODEs. Then by using implicit Euler method, all the
AC C
ODEs were solved simultaneously for the next time step [26, 28]. The bed was discretized by 20 nodes for each packed bed of activated alumina and 13X zeolites. The sensitivity of the simulation results to the node numbers was evaluated and it was found that the results are independent of the node numbers for node numbers greater than 20. The time step for implicit Euler method was also chosen to be 100 sec. The dependence of the results on the time step was also examined and it was found that the results were independent of the time step for time step values smaller than 100 sec.
ACCEPTED MANUSCRIPT
Each cycle consists of different steps in which different boundary and initial conditions are governed. The initial and boundary conditions for each step are introduced in Appendix A4. 3. Results and discussions
RI PT
To choose the best regeneration method for mercaptan desorption step, different conditions were performed and compared with each other on the base of mercaptan loading. Figure 4 shows mercaptan loading in different processes of PTSA, PVSA at 302 K and PVSA at 350 K
Fatemi [14].
M AN U
Figure 4.
SC
in the regeneration step. The PTSA results have been taken from the simulation of Taheri and
Obviously in Figure 4, PVSA with regeneration temperature of 350 K exhibits reduction of mercaptan loading to the lower level during a shorter time, whereas PTSA and PVSA without bed warming could not reach to this fast and short regeneration time. Therefore, PVSA with warming up to 350K, was considered as the regenerating operation in the cyclic process.
TE D
Some results of PVSA simulation are also reported in Appendix A5. 3.1.Cyclic steady state condition
The cyclic adsorption process as every dynamic process should pass several cycles to reach the steady state condition. Therefore at the next cycles, the extract and raffinate compositions
EP
become constant and do not change with time. A cyclic steady state (CSS) is defined for the same bed position in each cycle in which the dependent variables approach to a constant value.
AC C
In the present work, total adsorbed amount in adsorption stage between two subsequent cycles were checked to approach to the CSS as following:
∫
L 0
q i dz
L
( n −1) th cycle
− ∫ q i dz 0
<δ
(1)
( n ) th cycle
In the above equation, convergence parameter δ was considered to be 10-5 and this condition was selected as the stop criteria of the dynamic simulation.
ACCEPTED MANUSCRIPT
It was revealed that, after about twelve sequential running cycles of simple PVSA process, the steady state condition was achieved. The Figures 5, 6a and 6b show methane purity versus cycle number, mercaptan mole fraction in the outlet stream and time variation of solid average temperature, respectively. Regeneration of the adsorbent is a critical point in cyclic adsorption
RI PT
processes. Appropriate procedure is required for evacuation of the impurities from the bed. In the present case, mercaptan is the strongest adsorptive material on 13X zeolite, therefore properly bed evacuation from mercaptan at each cycle and approaching to a constant temperature is required for a continuous cyclic process. Since the adsorption of the other
SC
gaseous compounds, such as N2 is not significant, few cycles are required to approach to the cyclic steady state condition. Especially in this case, the vacuum condition helped better
M AN U
regeneration. Figure 5. Figure 6.
3.2.Validation
TE D
Average outlet mercaptan and water vapor mole fractions which are resulted from PVSA simulation at steady conditions are well validated with industrial unit data. The comparison of simulated and real values for mercaptan and water vapor mole fractions are shown in Figure 7. According to the reported values, mercaptan and water vapor relative errors are 1.5 % and 7.8
EP
%, respectively.
Figure 7.
AC C
In order to investigate the validity of the simulation results, mercaptan breakthrough curves from adsorption step in the real and simulated conditions are plotted and compared with each other, as shown in Figure 8.Mercaptan outlet from 30 hours adsorption time shows gradually increasing mercaptan at the exit section of the bed. The breakthrough results show that real plant data and simulated results are close together and there is high precision between real and model data during 18 hours adsorption which is the adsorption time of the process. Figure 8.
ACCEPTED MANUSCRIPT
3.3.Breakthrough curves The breakthrough curves of the species are shown in Figure 9 which present the mole fraction
RI PT
of each species in gas phase at different height of the bed. Feed gas enters from the top of bed and passes through the alumina pre-layer. In this layer, most of water vapor and some of CO2 content of feed is adsorbed before entering the 13X layer. As shown in Figure 9a, the bottom of the bed (13X, bottom) is approximately free from mercaptan whereas the top of 13X is rapidly
SC
reached to the feed mercaptan content. As shown in Figure 9b, the majority of water vapor has been removed in alumina pre-layer and the residue was adsorbed in 13X layer. Figure 9c and 9d
M AN U
show breakthrough curves of carbon dioxide and propane at different points of the bed, respectively. It can be seen that the fractions of carbon dioxide and heavy hydrocarbons reach to the feed content between 1 to 2 hours after passing through the column. However, the affinity of 13X towards CO2 is higher than that of propane.
Figure 9.
3.4.1. Loading curves
TE D
3.4.PVSA process in comparison with simulated PTSA process
One important factor in every cyclic adsorption process is the adsorbed amount of solid phase in
EP
adsorption and desorption stages. The loading curves show the maximum and minimum adsorbed amount and the time required for regeneration. Due to the fact that purging gas goes through the bed from bottom to top, the components accumulate at the top of the bed and this
AC C
point of the bed needs more time for desorption. Therefore the time required for mercaptan desorption at top of the bed can be an important factor for design of the cycle time which consequently affects the recovery of the process. Figure 10 shows mercaptan and water vapor loading in PVSA and PTSA processes at top of the bed [14]. According to Figure 10a, it is found that the time needed for complete mercaptan desorption at top of the bed in the PVSA process is 24 hours while in the PTSA process, it requires 30 hours to be totally desorbed. Water vapor acts as the dominant rival for mercaptan species in the adsorption by 13X. As it was seen previously in breakthrough curves, water vapor is a competitive adsorbing component
ACCEPTED MANUSCRIPT
with mercaptan and slows down bed regeneration. Therefore it is necessary to investigate water vapor competitive behavior and the impact of regeneration method on water desorption in both PVSA and PTSA process. Figure 10b indicates that loaded water vapor is evacuated from the top of 13X after 22 hours in PVSA process while in the PTSA, loaded water vapor has reached
RI PT
to 0.015 mol/kg after 25 hours (at the end of first heating step). The results reveal that it is possible to work with lower cycle time in the PVSA process rather than PTSA because lower time is required for the regeneration and this means saving recovery and productivity.
SC
Figure 10.
Table 4 reports the regeneration step time required for different cyclic processes to desorb
M AN U
mercaptan and water vapor and prepare the bed for the next adsorption cycle. The results show that using PVSA process has shortened the cyclic time compared with PTSA and it can improve the productivity of the process.
Table 4.
3.4.2. Temperature profile
TE D
Since adsorption is exothermic, every process with adsorption and desorption steps includes temperature variations. In such processes, temperature increases in adsorption stage and decreases in desorption stage. As shown in Figure 11, temperature is changed in both PVSA and PTSA process. As can be seen in Figure 11a and 11b, at the beginning of the adsorption step,
EP
the gas temperature rises to about 320 K because of adsorption of high amounts of carbon dioxide and heavy hydrocarbons, however after 2 hours, none of these components are
AC C
adsorbing. Therefore, the temperature reaches to the feed stream temperature. In Figure 11a, a sharp temperature decrease is observed after the end of adsorption time (18 hours). This is because of extreme pressure drop in blow down and evacuation step (6.8 MPa to 0.1 MPa) in the PVSA process. According to this figure, at the top of the bed, this decrement is more, since in these two steps, the gas stream flows from down to top. After that the bed temperature arises to the purge temperature of 350 K. the temperature varies from bottom to top of the bed, because the gas stream loses its temperature and becomes cooler along the bed. Figure 11.
ACCEPTED MANUSCRIPT
3.4.3. Process efficiency The performance of a cyclic adsorption process is commonly evaluated according to basic parameters of product purity, productivity and recovery defined, respectively by equations (2),
!
"##$ | ∑, % '() * ,(. %
RECOVERY =
*
*
*
6 : "##$ c89 u: |<=> dt − : $#BC#DD c89 u: |<=> dt − : BECF# c89 u: |<=> dtG AI
*
t W
*
SC
PRODUCTIVITY =
(2)
*
BECF# c89 u: |<=> dt : "##$ c89 u: |<=> dt − : $#BC#DD c89 u: |<=> dt − :
*
*
BC#DD c89 u: |<=: dt : "##$ c89 u: |<=: dt + :
M AN U
PURITY =
!
% "##$ % |'() *
RI PT
(3) and (4) as below.
(3)
(4)
The values of purity, productivity and recovery for three processes of PTSA, simple and improved PVSA are showed in Figure 12 [14]. According to the data reported in Figure 12, the purity of methane has increased by a value of 0.0299% using the PVSA processes. The incremental change of purity is not significant, whereas there are significant enhancement in the
TE D
values of productivity and recovery of the PVSA compared with PTSA. The value of productivity has been increased from 2.79[ PQR
PQR
ST.UVW
] in PTSA to 3.90[
PQR
ST.UVW
] in simple PVSA and
to a less greater value of 3.94[ST.UVW] in improved PVSA. The value of recovery has increased PVSA.
EP
from 74.04% in PTSA to 94.80% in simple PVSA and to a greater value of 98.27% in improved
AC C
The much less amount of PTSA recovery is due to the purge step in which excess amount of the natural gas is released to the atmosphere at high pressure, whereas in the PVSA process the purge step is carried out at low pressure. The impact of pressure equalization has been studied in improved PVSA on the NG recovery and this impact can be seen in comparison of the simple and improved PVSA processes. Pressure equalization has improved the recovery of the process by 3.47% from simple PVSA to improved PVSA. This enhancement is due to avoiding gas release during depressurization steps to the atmosphere. Figure. 12
ACCEPTED MANUSCRIPT
3.4.4. Economic study In order to compare three processes of simple PVSA, improved PVSA and PTSA from the operating cost point of view, some calculations have been done.
RI PT
The effective parameters of the operating cost for the PVSA process is included purge gas cost, gas stream cost for pressurization step, vacuum pump cost and purge temperature cost. Operating cost of PTSA process is influenced by the purge gas cost, gas stream cost for pressurization step and temperature change cost in heating and cooling steps.
SC
In order to calculate the operating costs, gas flow rate in each step, gas heat value (HV), vacuum pump power and efficiency, and price of the feed gas, product gas, electricity and fuel gas are
M AN U
required. In Table 5 and Figure 13, operating cost calculations for three processes are exhibited [29, 30].
Table 5.
Figure 13.
TE D
As can be seen in Figure 13, the operating cost of PTSA process is more than PVSA processes. According to Table 5, the main reason is the energy cost in heating steps (480K in 1st heating step and 590K in 2nd heating step) of PTSA process. Because of increasing the operating temperature (303K) to the high temperature and thereafter cooling down the gas consumes large
EP
amounts of energy. The difference of operating cost between simple and improved PVSA processes comes from pressure equalization steps in which the required gas for pressurizing the bed is supplied with the released gas in depressurizing steps, using improved PVSA cycle.
AC C
Therefore, a large amount of valuable gas can be saved and the operating cost can be decreased, significantly.
4. Conclusions
Process of cyclic adsorption based on PVSA was proposed as a more efficient process rather than PTSA process of South Pars Region of Iran. The most difference of the proposed PVSA process was according to the regeneration method. The improved PVSA was revealed to be more economic than the simple PVSA on the base of more energy saving.
ACCEPTED MANUSCRIPT
It was concluded that improved PVSA would successfully remove mercaptan to lower than 10 ppm with higher recovery and productivity than those in existing PTSA. At the same time, operating cost of the PVSA that is using relative vacuum for the bed regeneration is lower than the cost of PTSA that is using high gas temperature for regenerating the adsorbents. At the same
RI PT
time, cooling down the column at PTSA process takes much more time than PVSA.
Investigation of the basic performance parameters showed that the improved PVSA process has increased the recovery by 3.47% and productivity by 1% rather than the simple PVSA process.
SC
In addition, the improved PVSA process could increase the recovery and productivity by 24.23% and 44.8% respectively versus PTSA process. The economic study of the PTSA, simple PVSA and improved PVSA indicated that by using the PVSA process instead of PTSA process,
AC C
EP
TE D
M AN U
the operating cost would be decreased by 17 %.
ACCEPTED MANUSCRIPT
NOMENCLATURE Specific area of adsorbent (1/m)
b
Heat capacity equation parameter ( J/mol/K)
B:,]
Affinity constant at infinite temperature (1/KPa)
c
Heat capacity equation parameter (K)
C]
K Z,]
Concentration of adsorption of species i (mol/m_ )
k]
K d
Pore diffusivity of species i (m gs) e
Overall mass transfer coefficient (sqr ) Effective axial thermal conductivity (W/m/K)
k ′ Vapor thermal conductivity at atmospheric pressure (W/m K)
RI PT
aZ
M
Molecular weight (g/mol)
C]a Concentration of adsorption of species i in depressurization step (mol/m_ )
Nu
Nusselt number
n Number of the adsorbed components in the mixture
C]b Concentration of adsorption of species i in evacuation step (mol/m_ )
n Number of cycle
Concentration of adsorption of species i in C]c regeneration step (mol/m_ )
n
Number of all component in gas mixture
P
Pressure (KPa)
p]
Partial pressure (KPa)
M AN U
C]Z Concentration of adsorption of species i in pressurization step (mol/m_ ) CZ
Gas heat capacity (J/mol/K)
Average heat capacity of adsorbent (J/kg/K)
d
Heat capacity equation parameter ( J/mol/K)
TE D
CZ
D d,] Effective axial dispersion coefficient of species i e (m gs)
EP
D,] Molecular diffusivity of species i in mixture e (m gs) Dh,]
e Knudsen diffusivity of species i (m gs)
e Molecular diffusivity of species i in j (m gs)
e
Heat capacity equation parameter (K)
AC C
D],i
E Characteristic energy of the adsorbent (J/mol) h] Film heat transfer coefficient between gas and solid (W/me /K) HV
Gas heat value (BTU/m3)
∆H]
Heat of adsorption of species i (J/mol)
K
k o,]
SC
C]` Concentration of adsorption of species i in adsorption step (mol/m_ )
Gas thermal conductivity (W/m/K) Film mass transfer coefficient of species i (m/s)
q] Adsorbed phase concentration of species i (mol/kg)
q∗] (mol/kg)
Equilibrium adsorbed phase concentration
q,] (mol/kg)
Maximum adsorbed phase concentration
R
Ideal gas law constant (J/mol/K)
rZ
Adsorbent radius (m)
Re
Reynolds number and
Sc
Schmidt number
Sh
Sherwood number
t
Time (s)
T
Gas temperature (K)
T
Adsorbent temperature (K)
rZ Pore radius (m)
T
T
Reduced temperature Critical temperature (K)
ACCEPTED MANUSCRIPT
z{
y:
Superficial gas velocity (m/s)
µ
_
Critical molar volume (m /Kmol) Initial mole fraction
y]
Mole fraction of species i
z
Axial coordinate in the bed (m)
Shape factor
δ
Convergence parameter
η
Pump efficiency
Subscripts
εZ
Adsorbent porosity
τ
Tortuosity factor
ρ
kg Adsorbent density ( gm_ )
1. References
F H L A R D E P
Feed High pressure Low pressure Adsorption Regeneration Depressurization Evacuation Pressurization
M AN U
Bed porosity
Gas viscosity (kg/m.s)
φ
Greek letters ε
Gas density (molgm_ )
SC
u
ρ
RI PT
Tsurr Surroundings temperature
[1]. M.Tagliabue, D.Farrusseng, S. Valencia, S.Aguado, U.Ravon, C.Rizzo, A.Corma, C.Mirodato, Natural gas treating by selective adsorption: Material science and chemical engineering interplay, Chemical Engineering Journal 155 (2009) 553–566.
TE D
[2]. J.P. Bellat · F. Benoit · G. Weber · C. Paulin ·P. Mougin · M. Thomas, Adsorption equilibria of binary ethylmercaptan/hydrocarbon mixtures on a NaX zeolite, Adsorption (2008) 14: 501–507. [3]. H.Tamai. H.Nagoya. T.Shiono, Adsorption of methyl mercaptan on surface modified activated carbon, Journal of Colloid and Interface Science 300 (2006) 814–817.
EP
[4]. A. de Angelis, Natural gas removal of hydrogen sulphide and mercaptans, Applied Catalysis B: Environmental 113– 114 (2012) 37– 42. [5]. G.Weber , F. Benoit , J-P Bellat , C. Paulin , P. Mougin ,M.Thomas, Selective adsorption of ethyl mercaptan on NaX zeolite, Microporous Mesoporous Mat. 109 (2008) 184–192.
AC C
[6]. S.Cavenati, Carlos A. Grande, Alírio E. Rodrigues, Separation of CH4/CO2/N2 mixtures by layered pressure swing adsorption for upgrade of natural gas, Chemical Engineering Science 61 (2006) 3893 – 3906. [7]. T.L.P. Dantasa, F.M.T. Lunab, I.J. Silva Jr, A.E.B. Torres, Carbon dioxide–nitrogen separation through pressure swing adsorption, Chemical Engineering Journal 172 (2011) 698– 704. [8]. V.P. Mulgundmath, R.A. Jones, F.H. Tezel, J. Thibault, Fixed bed adsorption for the removal of carbon dioxide from nitrogen: Breakthrough behavior and modeling for heat and mass transfer, Separation and Purification Technology 85 (2012) 17–27. [9]. M. Clausse, J. Bonjour1, F. Meunier, Adsorption of gas mixtures in TSA absorbers under various heat removal conditions, Chemical Engineering Science 59 (2004) 3657 – 3670.
ACCEPTED MANUSCRIPT
[10]. J. Zhang, P.Xiao, G.Li, P.A. Webley, Effect of Flue Gas Impurities on CO2 Capture Performance from Flue Gas at Coal-fired Power Stations by Vacuum Swing Adsorption, Energy Procedia 1 (2009) 1115–1122.
RI PT
[11]. J. Zhang, P.A. Webley , P.Xiao, Effect of process parameters on power requirements of vacuum swing adsorption technology for CO2 capture from flue gas, Energy Conversion and Management 49 (2008) 346–356. [12]. M.C. Campo, A.M. Ribeiro, A. Ferreira, J.C. Santos, C. Lutz, J.M. Loureiro, A.E. Rodrigues, New 13X zeolite for propylene/propane separation by vacuum swing adsorption, Separation and Purification Technology 103 (2013) 60–70.
SC
[13]. B. Shirani, T. Kaghazchi, and M.Beheshti, Water and mercaptan adsorption on 13X zeolite in natural gas purification process, Korean J. Chem. Eng., 27(1), (2010) 253-260.
M AN U
[14]. O.T. Qazvini, S. Fatemi, Modeling and Simulation Pressure–Temperature Swing Adsorption Process to Remove Mercaptan from Humid Natural Gas; A Commercial Case Study, Separation and Purification Technology (2014). [15]. D. Ferreira and R. Magalhães, Effective adsorption equilibrium isotherms and breakthroughs of water vapor and carbon dioxide on different adsorbents, Industrial & Engineering Chemistry Research, (2011) 10201–10210. [16]. R. W. Glass and R. A. Ross, Surface Studies of the Adsorption of Sulfur-Containing Gases at 423°K on Porous Adsorbents. I I. The Adsorption of Hydrogen Sulfide, Methanethiol, Et hanethiol, and Dimethyl Sulfide on y-Alumina, The Journal of Physical Chemistry, Vol. 77, No; 27, (1973).
TE D
[17]. E. Baumgarten, F. Weinstrauch, and H. Hoffkes, Adsorption isotherms of hydrocarbons on UAlumina, Journal of chromatography 138, (1977) 347–354. [18]. D. Green and R. Perry, Perry’s chemical engineers' handbook, 8th Ed. New York: McGraw-Hill, (2007).
EP
[19]. M. Mehdipour and S. Fatemi, Modeling of a PSA-TSA process for separation of CH4 from C2 products of OCM reaction, Separation Science and Technology, (2012) 1199–1212. [20]. R. Desai, M. Hussain and M. Ruthven, Adsorption of Water Vapour on Activated Alumina. I Equilibrium Behaviour, the Canadian jurnal of chemical engineering (1992) 699-706,.
AC C
[21]. D.M. Ruthven., Principles of adsorption and adsorption processes, John Wiley, New York (1974). [22]. N. Wakao, T. Funazkri, Effect of fluid dispersion coefficients on particle-to-fluid mass transfer coefficients in packed beds, Chem. Eng. Sci. 33, (1978) 1375–1384, [23]. .R. Welty, C.E. Wicks, R.E. Wilson, G. Rorrer, Fundamentals of Momentum, Heat and Mass Transfer, John Wiley and Sons, New York (2000). [24]. R. T. Yang, Gas separation by adsorption processes. Boston: World Scientific Publishing Company, (1987). [25]. R. Byron Bird, W. E. Stewart, E. N. Lightfoot, Transport Phenomena, JOHN WILEY & SONS, ING. New York Chichester Weinheim, Brisbane Singapore Toronto.
ACCEPTED MANUSCRIPT
[26]. Da Silva and A. E. Rodrigues, Adsorption equilibria and kinetics for propylene and propane over 13X and 4A zeolite pellets, Industrial & Engineering Chemistry Research 38 (1999) 2051–2057. [27]. K. Iusalaas, Numerical methods in engineering with MATLAB. Cambridge: Cambridge University Press (2005).
RI PT
[28]. G. G.Vining, Statistical methods for engineers. Pacific Grove: Duxbury Press (1997). [29]. Exclusive price list of oil, gas and energy facilities, Ministry of Petroleum, Engineering Department, 058 (2014).
[30]. M. Peters, K. Timmerhaus and R. West, Plant Design and Economics for Chemical Engineers, 5th
SC
edition (2002).
M AN U
APPENDIX A1. The physical properties of the adsorbents
The physical properties of adsorbents and the bed Characteristics are given in Table A1. Table A1.
TE D
A.2.Molecular diffusivity and physical properties prediction and correlation Molecular diffusivity of each component in the gas mixture is calculated by the equation (A.1) [18]:
1 − (. 1) ∑ = ,
EP
P, =
AC C
In this equation, is the mole fraction of the ith component and n is the number total
components in the gas mixture. For the mixture of natural gas, , can be assumed as , . , is the molecular diffusivity of binary gas mixture and is calculated by the Wilke-Lee
equation [18].
The amount of molecular diffusivity of each component through the mixture at the adsorption, evacuation and purge is brought in Table A2. Table A2.
ACCEPTED MANUSCRIPT
All physical properties are considered as methane property. Because methane is the dominate component of the gas phase. Pure gas specific heat capacity T has been calculated as the equation (A.2) [19]: e 6dZ ( ) –dZ (q ) G
+ ×
e (. 2) 6dZ ( ) dZ (q ) G
Table A3. A.3. Mass transfer and heat transfer correlations
SC
The coefficients of the above correlation are listed in Table A3.
RI PT
{
T = + ×
M AN U
The transport parameters needed in the model were calculated employing frequently used correlations. The correlations are used for the mass and energy balances in gas and solid phase are shown in Table A4.
Table A4.
TE D
A4. The initial and boundary conditions for numerical calculation As stated in numerical calculation section, to obtain the results of PVSA simulation, several non-linear equations should be solved, simultaneously. The initial and boundary conditions for the numerical solution for simple PVSA process are reported in Table A5. The initial and
Table A5.
AC C
EP
boundary condition of PTSA process is described in the paper of Taheri and Fatemi [14].
Where the subscripts F, H, L, A, R, D, E and P represent the feed, high pressure, low pressure, adsorption, regeneration, depressurization, evacuation and pressurization, respectively and L is the length of bed.
A.5. Mass transfer and heat transfer parameters for kinetic adsorption The average transfer parameters of the PVSA simulation were found based on the correlations are listed in Table A4. The parameters in different steps of the process are reported in Table A6.
ACCEPTED MANUSCRIPT
The amounts of , that are reported in the Table A6, are the overall heat transfer coefficient h
which is dependent of h] and h by equation (A.3). The amount of h is 10(.h).
h,
=
1 1 + (. 3) h] h
RI PT
1
Table A6. Figure Captions:
PTSA, b: simple PVSA, c: improved PVSA.
SC
Figure 1. Schematic of the step cycle sequences, in three processes for mercaptan removing, a:
M AN U
Figure 2. Operating bed pressure with elapsed time in (a): PTSA cycle, (b): PVSA cycle. Figure 3. The time-table of the improved PVSA process, at each column for a six- bed plant. Figure 4. Mercaptan loading on the top of 13X layer at the mentioned processes. Figure 5. Methane purity after evacuation stage, with the simulation cycle number in a simple
TE D
PVSA process with the inlet mole fraction of 95.9848% methane.
Figure 6. a: Mercaptan mole fraction at the outlet stream, b: solid temperature profile as a function of cycle number, for the simple PVSA process.
EP
Figure 7. Comparison of simulated and real values of mercaptan and water vapor mole fractions at the outlet stream.
AC C
Fig. 8. Simulated mercaptan breakthrough curve in comparison with real data. Figure 9. Breakthrough curves in the simple PVSA process, (a) Mercaptan, (b) water vapor, (c) carbon dioxide and (d) propane Figure 10. Loading values versus cycle time at PVSA and PTSA cycles at the top of 13X layer, (a) Mercaptan, (b) water vapor. Figure 11. Temperature profile along the bed at (a) simple PVSA process, (b) PTSA process. Figure 12. Basic performance parameters for PTSA, simple PVSA and improved PVSA.
ACCEPTED MANUSCRIPT
Figure 13. Operating cost for PTSA, simple PVSA and improved PVSA processes. Table Captions:
RI PT
Table1. Feed specifications and the bed Characteristics. Table2. Mathematical model of the dynamic adsorption including, total mass, mole and energy balances.
Table3. Parameters of Extended Langmuir Isotherm for Activated alumina and 13X.
SC
Table4. Comparison of the cyclic processes from the point of mercaptan and water desorption
Table5. Operating cost calculations.
M AN U
time and amount.
Table A1. Physical properties of the adsorbents.
Table A2. Average molecular diffusivity of gases in mixtureP, ( ¡e /¢).
TE D
Table A3. Heat capacity coefficients of methane.
Table A4. Mass and heat transfer correlations to solve the simulation equations. Table A5. The initial conditions for numerical calculation employed at simple PVSA
EP
simulation.
AC C
Table A6. Model parameters of the adsorption kinetics.
AC C
EP
TE D
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
Figure 1
AC C
EP
TE D
M AN U
Figure 2
SC
RI PT
ACCEPTED MANUSCRIPT
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
Figure 3
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
Figure 4
ACCEPTED MANUSCRIPT
94.8032
94.8028
Methane purity
94.8026
RI PT
Methane purity (%)
94.803
94.8024 94.8022 94.802
94.8016 2
4
6
8 10 Cycle number
12
M AN U
0
SC
94.8018
AC C
EP
TE D
Figure 5
14
16
ACCEPTED MANUSCRIPT
(a) 4.5 cycle 1 cycle 2 3.5
cycle 3
3
cycle 4
RI PT
cycle 5 2.5
cycle 6 cycle 7
2
cycle 8
1.5
cycle 9 1
cycle 10
0.5
cycle 11
SC
Mercaptan mole fraction (ppmv)
4
cycle 12 0
2
4
6
(b)370
8
10 Time (hr)
12
14
16
310
270
20
cycle 2 cycle 3 cycle 4 cycle 5 cycle 6 cycle 7 cycle 8 cycle 9
EP
290
18
cycle 1
TE D
330
AC C
Temperature (k)
350
M AN U
0
cycle 10 cycle 11 cycle 12
250
0
5
10
15
20 Time (hr)
Figure 6
25
30
35
40
ACCEPTED MANUSCRIPT
4.5 4
3.81
3.87 (*10-6 )
3.5 real value
3 2.5
RI PT
simulated value
2 1.5 1 0.5 0
Mercaptan mole fraction
SC
0.51 0.55 (*10-7)
Water vapor mole fraction
AC C
EP
TE D
M AN U
Figure 7
ACCEPTED MANUSCRIPT
simulated result
60
real data
50 40
RI PT
Mercaptan mole fraction (ppm)
70
30 20 10
0
5
10
15
Time (hr)
20
25
AC C
EP
TE D
M AN U
Figure 8
SC
0
30
ACCEPTED MANUSCRIPT
(b)
160 140
80
10
60 40 20 0
5 0
8
10 12 14
Time (hr)
16 18
0
4
6
8
10
12
14
16
18
Time (hr)
(d) z=0 (13X bottom) z=2.3 (13X middle) z=4.7 (alumina bottom) z=5 (alumina middle)
z=0 (13X bottom) z=2.3 (13X middle) z=4.6 (13X top)
0
6000
5000
10000
M AN U
15000
2
SC
6
12000 18000 24000 30000
4
20000
2
RI PT
15
100
0 0
0.5
1 Time (hr)
1.5
2
EP
Figure 9
AC C
CO2 mole fraction (ppm)
z=0 (13X bottom) z=2.3 (13X middle) z=4.7(alumina bottom) z=5(alumina middle)
20
120
0
(c )
25 Water vapor mole fraction (ppm)
z=0 (13X bottom) z=2.3 (13X middle) z=4.65 (13X Top)
Heavy hydrocarbons C3+ (ppm)
Mercaptan mole fractio (ppm)
180
TE D
(a)
0
0.2
0.4 0.6 Time (hr)
0.8
1
AC C
EP
TE D
M AN U
Figure 10
SC
RI PT
ACCEPTED MANUSCRIPT
AC C
EP
TE D
M AN U
Figure 11
SC
RI PT
ACCEPTED MANUSCRIPT
ACCEPTED MANUSCRIPT
120 96.10 96.13 96.13
94.80
98.27
74.04
80
PTSA SIMPLE PVSA IMPROVED PVSA
60 40 20
RI PT
100
2.79 3.90 3.94
% Purity
% Recovery
AC C
EP
TE D
M AN U
Figure 12
Productivity (mol/kg.day)
SC
0
ACCEPTED MANUSCRIPT
88
90 80
72.94
72.705
Simple PVSA
Improved PVSA
70 60
RI PT
50 40 30 20 10 0 PTSA
AC C
EP
TE D
M AN U
Figure 13
SC
operating cost (Thousand $/year)
100
ACCEPTED MANUSCRIPT
Table 1. Parameter Flow rate (Nm3/hr)
Value 2850
Bed Bed length (m)
13X 4.65
AC 0.75
Pressure (Mpa)
6.8
Bed diameter (m)
3.7
3.7
Temperature (k)
302
Bed void fraction
0.37
0.26
CO fraction (ppmv)
Heavy hydrocarbons fraction (ppmv)
3.0 × 10
95.9848 × 10
0.82
Adsorbent weight (ton)
21.723
4.89
AC C
EP
TE D
M AN U
Methane fraction (ppmv)
1.0 × 10
0.69
RI PT
133.8
Mercaptan fraction (ppmv)
Bulk density (g/ )
SC
18.2
Water vapor fraction(ppmv)
ACCEPTED MANUSCRIPT
Table2.
Overall mass balance in gas phase Solid components’ mass balance Gas phase energy balance
' (( ) = , – (1 − )%& − *+ (() ' =− − (1 − )%& )( ) , ' = ('∗ − ' ) / / / . − (%01 − % 01 − 20 (1 − )ℎ (/ − /& ) − ℎ&455 (/ − /&455 ) = 0
*+
RI PT
Gas components’ mass balance
Energy balance of the solid
8,
AC C
EP
TE D
Extended Langmuir isotherm
(1 − )( : 37.5 (1 − ) =( = −; + 0.875% B 1 ?0 @ >?0 @A ∆7 EF, exp J L M '∗ K/ = 'CD 1 + ∑*O E exp J∆7 L M 8, F, K/
M AN U
Ergun equation
SC
/& ' %& 0& = %& )(−∆7 ) + 20 ℎ (/ − /& )
ACCEPTED MANUSCRIPT
Table 3. PCQR ) P1
, ) T0
Water vapor Carbon dioxide
0.0151 0.0011
3.22 × 10VW 1.94 × 10V
0.003580 0.016842 0.005115 0.002821
4.32 × 10VX 4.82 × 10VW 3.61 × 10V 3.36 × 10V
- ∆7 (
AC C
EP
TE D
M AN U
SC
Zeolite 13X Mercaptan Water vapor Carbon dioxide Propane
PU ) CQR
SF, (
60.014 29.805
RI PT
'C, (
Activated alumina
65.1624 62.5258 34.0031 32.5412
ACCEPTED MANUSCRIPT
Table 4. Improved PVSA
PTSA
Mercaptan time for complete desorption (hr)
6
6
12
Water vapor time for complete desorption (hr)
4
4
7
Mercaptan loading at top of the bed (Kmol/g)
0.000033
Water vapor loading at top of the bed (Kmol/g)
0.00028
RI PT
Simple PVSA
0.019
0.00028
0.015
AC C
EP
TE D
M AN U
SC
0.000033
ACCEPTED MANUSCRIPT
Table 5.
Calculations
PTSA
$ Z[p OO4C ^ a `
$ Z[ot\5C R ^ a ` $ /Z2e Z[ ^ a `
= d(be deZf ?2b ^
$ a × M?ckbl4\R 1 & ^ a `
AC C
470
235
15270
15270
15270
-
44000
44000
72000
13200
13200
88000
72940
72705
Wpump: calculated by thermodynamic correlation for isentropic work=250 KW HV: Heat Value of gas= 1/29 ηvwxv:pump efficiency= 0.7
TE D
PTSA pressurize flow rate=106000 Simple PVSA pressurize flow rate=68000 Improved PVSA pressurize flow rate=34000 purge flow rate= 1476000 PTSA consumed fuel flow rate=175000 PVSA consumed fuel flow rate=32200
EP
pricefeed= 0.2 priceproduct= 0.3 priceelectricity= 0.14 pricefuel gas= 0.41
q04C0 $ 7 =^ a × M?ckb\R\Oo5Oos ^ a × 24 × 366 h j r04C0 .q7 `
M AN U
$ Z[0451\ ^ a `
S/i = M(?nb deZf ?2b ^ a × 7g h j ` $ × M?ckb05Qm4Oo ^ a S/i
Improved PVSA
730
SC
S/i = M?b[[(?c b deZf ?2b ^ a × 7g h j ` $ Z[05\&&45]\ ^ a $ ` × M?ckbl\\m ^ a S/i
Simple PVSA
RI PT
Cost
ACCEPTED MANUSCRIPT
Table A1. Physical properties of adsorbents
Zeolite 13X
Activated Alumina
2.1
2.3
Average macropore radius (nm)
2.44
4.6
445.6
441.3
BET surface area (m /g)
Particle void fraction
0.24
1.10
Heat capacity (kJ/kg.K)
1.07
Tortuosity
1.38
0.28
1.24
3.85
1.36
AC C
EP
TE D
M AN U
SC
Particle density (g/m )
RI PT
Particle radius (mm)
ACCEPTED MANUSCRIPT
Table A2. Adsorption
2.238 × 10V
Carbon dioxide Heavy hydrocarbons Water vapor
Purge
1.6
3 × 10V
2.59 × 10V
1.1988× 10V
0.16
0.95
9.5× 10V
3.4
0.34
0.79
AC C
EP
TE D
M AN U
SC
Mercaptan
Evacuation
RI PT
Stage
7.9 × 10V
ACCEPTED MANUSCRIPT
Table A3.
(z. Ze V, . . V, ) E × 10V{ (z. Ze V, . . V, ) k × 10V (. ) × 10V (z. Ze V, . . V, ) b × 10V (. ) 2 × 10V
0.7993
2.0869
4.16
9.92
EP
TE D
M AN U
SC
RI PT
3.33
AC C
7
ACCEPTED MANUSCRIPT
The effective axial bed thermal conductivity The effective mass transfer coefficient [21]
External film resistance( Wakao and Funazkri) [22]
The axial dispersion coefficient [23]
Tortuosity Factor [11]
TE D EP AC C
= 0 + 1.5>1 − 0 A / ~, = 97?0Q5\ ( )F.{
M AN U
Knudsen diffusion coefficient [24] Chilton–Colburn analogy [25]
,/
ℎ = 2 + 1.1k Kb F. , = 20 + 0.5 k Kb C, 1 1 1 = ^ + a .0, ~, C,
SC
The macropore diffusion coefficient
. = 7 + 0.5:?Kb .1 ?0 ?0 1 ?O = + + }~ O, 3l, 15ɛ .0, 15. [ℎ C, l, = 2?0
RI PT
Table A4.
,/
( = 2.0 + 1.1:?
Kb F.
ACCEPTED MANUSCRIPT
Table A5. Adsorption stage:
RI PT
= 0:
= ,*+ + ' = ',*+ + ,
: = : = 6.8 :2 , / = / = 302 K / ( ) = 0 , = 0
( ) ( > − A , = , ,
TE D
/ (k01 : (/ − / ) , : = : , ( = ( = . K/
EP
Depressurization stage:
AC C
= = 5.4 :
M AN U
SC
= 0:
ACCEPTED MANUSCRIPT
1&o bM?b[[(?c b [bM:
= *+ + , ' = '*+ + ,
: − : / = /*+ + , : = − + : cbm\05\&& = 0:
: = 6.8 :2, : = 4.0 :2
2*m bM?b[[(?c b [bM:
RI PT
= 0:
: = 4.0 :2, : = 1.5 :2
( ) = 0 ,
35m bM?b[[(?c b [bM:
SC
/ = 0 , ( = 0 = = 5.4 :
: = 1.5 :2,
M AN U
( ) / = 0 , = 0 , P = P 8
: = 0.1 :2
cbm\05\&& = 10 c
Countercurrent evacuation stage: = 0:
TE D
= *+ + , ' = '*+ + ,
/ = /*+ + P = P = 0.01 :2 = 0:
= = 5.4:
EP
( ) / = 0 , = 0 , ( = 0
AC C
( ) / = 0 , = 0
Countercurrent regeneration stage:
ACCEPTED MANUSCRIPT
= 0:
= 0:
( ) ( / (k01 : ( − 0), (/ − / ), =− = , . K/
( = ( , P = P = 0.1 :2 = = 5.4 :
Co-current pressurization:
M AN U
SC
( ) / = 0 , = 0
TE D
= 0: = ,*+ + V, , ' = ',*+ + V, : − : / = /,*+ + V, , : = + : cb05\&& = 0:
/ ( ) = 0 , = 0 , ( = 0
EP
= = 5.4 :
AC C
( ) / = 0 , = 0
RI PT
= ,*+ + , ' = ',*+ + , / = /,*+ +
1&o M?b[[(?c b [bM:
: = 1.5 :2, : = 0.1 :2
2*m M?b[[(?c b [bM:
: = 4.0 :2, : = 1.5 :2
35m bM?b[[(?c b [bM:
: = 6.8 :2, : = 4.0 :2
cb05\&& = 10 c
M AN U
SC
RI PT
ACCEPTED MANUSCRIPT
Zeolite 13X
Activated alumina
Water vapor
0.00098
0.00014
EP
(1/s)
Mercaptan
0.00093
-
Carbon dioxide
0.00141
0.008205
Heavy hydrocarbons
0.00731
-
AC C
Step
TE D
Table A6.
adsorption , (
¡
)
Water vapor
1.7028
2.4177
Mercaptan
1.7031
2.4182
Carbon dioxide
1.7021
2.4171
Heavy hydrocarbons
1.7025
2.4172
5.848
5.941
3.1621
4.1246
£ ) .¢
¢, (
£
¤¥, ( .¢)
ACCEPTED MANUSCRIPT
0.0118
Mercaptan
0.0043
Carbon dioxide
0.00084
Heavy hydrocarbons
0.0008
¡
)
Water vapor
2.148
Mercaptan
2.186
Carbon dioxide
2.175
Heavy hydrocarbons
2.166
£
¢, (
.¢
)
£ ) .¢
¤¥, (
(1/s)
Water vapor
0.0592
Carbon dioxide Heavy hydrocarbons
-
3.493
3.528
0.0057
0.0126
3.537
3.529 0.35
0.3931
0.051 1.1369 -
TE D
)
8.843
13.217
Mercaptan
8.859
13.223
EP
¡
0.0072
Water vapor
Carbon dioxide
8.854
13.219
Heavy hydrocarbons
8.820
13.216
0.4548
0.4564
£
0.5256
0.6499
£ ) .¢
¢, (
¤¥, ( .¢)
AC C
purge
0.2135
0.0311
Mercaptan
, (
0.35
-
M AN U
Evacuation
, (
0.00142
RI PT
Water vapor
SC
(1/s)
ACCEPTED MANUSCRIPT
Research highlights:
EP
TE D
M AN U
SC
RI PT
Mercaptan removing from natural gas. Vacuum Pressure Swing Adsorption process for methane purification. Process performance parameters comparison and economic study for PVSA and PTSA processes.
AC C
• • •