Ind. Eng. Chem. Res. 2009, 48, 5415–5419
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Effect of Hydrocarbons on the Separation of Carbon Dioxide From Methane through a Polyimide Gas Separation Membrane Raza Hasan, Colin A. Scholes, Geoff W. Stevens, and Sandra E. Kentish* CooperatiVe Research Centre for Greenhouse Gas Technologies (CO2CRC), Department of Chemical and Biomolecular Engineering, UniVersity of Melbourne, VIC, 3010, Australia
The removal of carbon dioxide from natural gas by membrane gas separation is an important process in upgrading the gas for consumers. Hydrocarbons associated with natural gas have the potential to influence the CO2 separation performance of membranes. Here, a glassy polymeric membrane of poly (4,4′hexafluoroisopropylidene diphthalic anhydride-2,3,5,6-tetramethyl-1,4-phenylenediamine) (6FDA-TMPDA) was exposed to hexane and toluene at different partial pressures in a 90% CH4-10% CO2 gas mixture. The permeability of CO2 and CH4, as well as the selectivity, declined over time as both hexane and toluene competitively sorbed into the membrane. The resulting behavior fit the dual-sorption model for polymeric membrane separation and indicated Langmuir affinity constants of 0.13 ( 0.03 atm-1 for methane, 240 ( 100 atm-1 for hexane, and 980 ( 500 atm-1 for toluene (fugacity basis). The kinetics of the permeability decline were faster than that of the permeability recovery after the hydrocarbon contaminant was removed which may indicate that desorption from the Langmuir voids is the rate controlling step. Introduction Natural gas is “sweetened” to meet consumer specifications through the removal of CO2, with the end gas typically containing no more than 2 mol % CO2.1 A number of separation technologies are capable of achieving this; with polymeric membrane gas separation being an ideal choice, given the high pressures and CO2 concentration found in natural gas. However, also present are hundreds of heavier hydrocarbons, both paraffinic and aromatic,2 and while pretreatment processes remove the majority of these hydrocarbons, trace amounts remain present during membrane separation. Subsequently, hydrocarbons can influence CO2 separation through polymeric membranes with a detrimental loss in performance. It is alleged that low quantities of condensable hydrocarbons can also cause premature aging and membrane failure.3 For most glassy polymeric membranes, hydrocarbons permeate through membranes at a lower rate than CO2 because of the molecular size difference.4 Previous studies on the effect of hydrocarbons on cellulose acetate membranes show essentially constant CO2/CH4 selectivity with a 10% reduction in permeabilities when exposed to benzene, toluene, and xylene.5 Conversely, polyimides such as 6FDA-DMB and 6FDA/BPDADAM experience significant drops in selectivity upon exposure to toluene.3,6 For 6FDA-DMB, this was accompanied by an increase in permeability while a decrease in permeability was noted for 6FDA/BPDA-DAM. Similar behavior was observed for 6FDA-DMB when exposed to hexane. Alternatively, the commercial membrane material Matrimid (polyimide) separation performance is stable upon exposure to toluene.3 Here, the competitive sorption effects of hexane and toluene on CO2 permeability and selectivity against CH4 in the glassy polyimide membrane poly (4,4′-hexafluoroisopropylidene diphthalic anhydride-2,3,5,6-tetramethyl-1,4-phenylenediamine) (6FDA-TMPDA) are studied. For membrane gas separation, 6FDA-TMPDA has impressive CO2 permeability and selectivity against CH4. Hence, it has potential in natural gas sweetening * To whom correspondence should be addressed. E-mail: sandraek@ unimelb.edu.au.
because polyimides generally have high thermal and mechanical stability, as well as chemical resistance.7 Theory For a glassy membrane, the dual mode sorption model predicts that the concentration of species A within the membrane can be described by both a Henry’s Law component (CDA) and a Langmuir component (CHA): CA ) CDA + CHA
(1)
where CDA is a function of the Henry’s Law coefficient (KDA) CDA ) KDA fA
(2)
For a pure gas stream, CHA is given by the standard Langmuir adsorption relationship for gas solubility within the excess free volume: CHA )
CHA ′ bA fA (1 + bA fA)
(3)
C HA ′ is the maximum adsorption capacity, while bA is the Langmuir affinity constant. Hence, the dual-mode sorption model for the concentration of a pure gas A in a glassy membrane is written as CA ) KDA fA +
CHA ′ bA fA (1 + bA fA)
(4)
Petropoulos8 and Paul and Koros9 independently developed models where the diffusion of the gas species adsorbed in the Langmuir region is partially, or even totally immobilized. This was experimentally verified by Koros in 197610 among others.11 In this case, a parameter FA is introduced, defined either as the ratio of diffusion coefficients in the Langmuir and Henry’s Law region (DH/DD)8 or, alternatively, as the fraction of penetrant in the Langmuir region that is mobile9 that is, the mobile concentration of species A is: CMA ) CDA + FACHA
10.1021/ie801537g CCC: $40.75 Published 2009 by the American Chemical Society Published on Web 04/30/2009
(5)
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When multiple gas species are present, competition further restricts the amount adsorbed in the Langmuir free volume.12 Hence, for a ternary mixture of gases A, B, and C the mobile concentration of gas A becomes
[
CMA ) KDA fA 1 +
FACHA ′ bA KDA(1 + bA fA + bB fB + bC fC)
]
(6)
Each parameter has the same definition as in the single gas case with the subscript denoting whether it is a property of gas A or B. The concentration, and hence permeability, of all species is reduced compared to the single gas case (eq 4) and is heavily dependent on the relationship between the affinity constant, b, and fugacity. The steady state unidirectional flux of the mobile fraction of species A through the membrane can be expressed in terms of Fick’s first law where DA is the Fickian diffusion coefficient: ∂CMA (7) ∂x If component A is a plasticizing gas such as carbon dioxide, then the diffusion coefficient also varies with concentration. These effects can be modeled using a simple exponential dependence of diffusion coefficient on penetrant concentration: JsA ) -DA
13
DA ) DA(0)e(βACMA)
(8)
where DA(0) is the infinite dilution diffusion coefficient for species A and β is referred to as the plasticization potential. This yields14 JsA )
-DA(0) ∂(eβACMA) βA ∂x
(9)
If the downstream fugacity of component A (fA1) approaches zero in a permeability measurement then CMA1 also approaches zero, and assuming that the flux JsA is constant under steady state conditions, integration with this boundary condition yields JsA )
DA(0) βCMA0 (e - 1) βAδ
(10)
where CMA0 represents the concentration of solute A within the membrane at the upstream face and δ is the membrane thickness. The average permeability coefficient can be defined in terms of this steady state flux and the fugacity difference between the upstream (fA0) and downstream (fA1) faces. jA ) P
JsAδ fA0 - fA1
(11)
Again, if the downstream fugacity, fA1, approaches zero, then simultaneous solution of eqs 10 and 11 yields jA ) P
DA(0) βCMA0 (e - 1) βA fA0
(12)
This finally yields15 jA ) P
DA(0) × βA fA0
( (
(
exp βAKDA fA0 1 +
FACHA ′ bA 1 + bA fA0 + bB fB0 + bC fC0
)) )
- 1 (13)
Experimental Details 6FDA-TMPDA was synthesized by the reaction between 4,4′(hexafluoroisopropylidene) diphthalic anhydride and 2,3,5,6-
Figure 1. Mixed gas membrane testing equipment with a hydrocarbon bubbler in place: PG pressure gauge, MFI mass flow indicator, MFC mass flow controller.
tetramethyl-1,4-phenylenediamine in N-methylpyrolidone (AR grade) under Ar to give a polyamic acid, which was subsequently imidized with triethylamine and acetic anhydride to yield 6FDA-TMPDA.16 The polymer was washed with dichloromethane (AR) and methanol (AR). Dense membranes were cast from 6FDA-TMPDA 2.5 w/v % dichloromethane solution, first filtered, into casting rings on glass plates. The casting volume determined the thickness of the symmetric dense nonporous membrane. The membranes were first dried at ambient temperature for 24 h to ensure complete solvent evaporation and annealed in a vacuum oven at 80 °C for 15 h and then at 150 °C for a further 48 h. The dried membranes where stored in a desiccator before use and had a typical thickness of 55-75 µm. The pure gas CO2 permeability of the membranes was initially tested at 35 °C under a range of pressures using a constant volume variable volume apparatus as described in the work of Duthie et al.17 The permeability and selectivity of membranes under mixed gas conditions was then tested using an in-house built instrument, as shown in Figure 1. Feed gas consisting of a certified mixture of 10.1% CO2 in methane (BOC Ltd., Australia) was passed through a solvent bubbler arrangement at controlled temperature. The bubbler temperature was altered between 5, 15, and 23 °C, resulting in different partial pressures of both hexane and toluene in the feed gas, associated with the saturation pressure at the bubbler temperature. The bubbler was packed with stainless steel Ballotini to increase mass transfer area; entrainment was minimized through the use of a downstream filter (2 µm). The feed gas entered the in-house built membrane cell, where the membrane was mounted in a 47 mm circular flat sheet arrangement, within an oven maintained at 35 °C, to prevent hydrocarbon condensation on the membrane. The feed/retentate was maintained at 1025 kPa g through back pressure regulation (Extech Equipment) with a flow rate of 100 mL min-1 (mass flow controller, Aalborg). Helium sweep gas was passed across the permeate side at 20 mL min-1 to prevent concentration polarization, with the flow rate monitored by a mass flow indicator (Aalborg). The concentration of CO2 and CH4 in the permeate stream was analyzed by gas chromatography (Shimadzu Model No. GC-8A with 100/120 mesh Porapak column (Alltech)) with the carrier gas being ultra high purity helium. The permeability of both hexane and toluene are not measured, since it is the effect these have on CO2 permeability that is of interest. On the basis of similar polymeric membranes,4 6FDA-TMPDA will have a selectivity for both hexane and toluene considerably less than 1 compared to CO2, mainly attributed to the large kinetic diameters of these hydrocarbons. Therefore, the flux of these components through 6FDA-TMPDA is small, leading to negligible change in feed composition
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Figure 2. Permeability data for both pure carbon dioxide and a mixture of methane (90%) and carbon dioxide (10%). The pure gas data is also modeled using the parameters in Table 1 and eq 13. Error bars for the pure gas data are (5% based on the work of Duthie24 while those for the mixed gas point are (1 standard deviation about the mean value of twelve data points.
Figure 4. Effect of an addition of toluene to a mixture of methane (90%) and carbon dioxide (10%) at a feed pressure of 1025 kPa g: (a) CO2 permeability, (b) CH4 permeability, (c) CO2/CH4 selectivity.
Figure 3. Effect of an addition of hexane to a mixture of methane (90%) and carbon dioxide (10%) at a feed pressure of 1025 kPa g: (a) CO2 permeability, (b) CH4 permeability, (c) CO2/CH4 selectivity.
throughout the membrane. Similarly, the observed consistency in feed side and permeate side flow rates upstream and downstream of the membrane confirmed that back diffusion of helium was minimal. A standard experiment consisted of operating a new membrane under dry conditions at the nominated temperature and pressure for 1 h to achieve steady-state conditions. The hydrocarbon was then introduced to the feed gas by passing through the solvent bubbler at a set temperature, with the permeate composition and flow rate monitored until steady state permeability and selectivity had been achieved. Then, the feed
gas was changed back to dry conditions by bypassing the bubbler, and the change in permeability and selectivity was observed. When hexane was used as the contaminant, the original permeability and selectivity could be completely recovered by purging with dry gas overnight. Conversely, performance recovery following toluene exposure required purging for a number of hours at 70 °C. Once performance had been recovered, a second experiment was conducted in an identical manner. After two experiments at identical conditions, the used membrane was discarded. Partial pressures of hexane and toluene in the inlet gas to the membrane cell were calculated using Aspen HYSYS with the Peng-Robinson equation of state. Fugacities of all gases were calculated by software provided by Sandler et al.18 Results and Discussion Permeability data for a pure stream of CO2 is provided in Figure 2. A typical parabolic curve shape is observed with the permeability falling initially as solubility declines then rising due to plasticization at higher fugacities. The absolute values of the pure gas CO2 permeability in the present case are a little lower than in our earlier work.17 This is probably the result of minor changes to the polymer preparation, casting and drying procedures for the membrane which affect the fractional free volume (FFV) of the membrane. In turn, these changes in
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Table 1. Model Parameters for Carbon Dioxide Permeation Used in Equation 13a KD (cm3 (STP) cm-3 atm-1)
C H′ (cm3 (STP) cm-3)
3.2
58.8
a
All values are from the work of Duthie et al. to the current data set.
17
b (atm-1) 0.55
D(0) (cm2 s-1) -9
3.9× 10
with the exception of D(0) which has been adjusted from 5.1 × 10
-9
β
F
0.024
0.57
2
-1
cm s
to provide a better fit
Table 2. Langmuir Affinity Constants Predicted from Equation 13 for Methane, Hexane, and Toluene b (atm-1)
this work (fugacity basis) Scholes et al.4 (pressure basis)
CH4
hexane
toluene
0.13 ( 0.03 0.07-0.16
240 ( 100
980 ( 500
fractional free volume affect the diffusion coefficient through the well-known expression:19 D ) A exp(-B / FFV) (14) As a result, to obtain a good fit to eq 13, it is necessary to adjust the infinite dilution diffusion coefficient to the lower end of the range predicted in the work of Duthie et al.17 (5.1 ( 3.5 × 109 cm2 s-1). The adjusted value of 3.9 × 109 cm2 s-1 remains within this predicted range. The values of all other parameters (KD, C H′ , bCO2, β, and F) are retained as in the work of Duthie et al.17 (see Table 1) and components B and C are null. The resulting fit to eq 13 is shown in Figure 2. Figure 2 also shows the CO2 permeability for the mixed gas (CO2/CH4). This is clearly lower than the pure gas result, due to competitive sorption of the CH4 species which is present in much greater quantities (pCH4 ) 1014 kPa). Use of eq 13 backcalculates a Langmuir affinity constant for CH4 of bCH4 ) 0.13 ( 0.03 (component C is null). This result is consistent with the previously published literature (Table 2). The change in permeability of both CO2 and CH4 as a result of exposure to hexane, at different fugacities, can be seen in Figure 3. For all three hexane fugacities, the permeability of both CO2 and CH4 decreases as a result of competitive sorption. The decrease is more pronounced for CO2, because the solubility of this gas within the polymer is more strongly influenced by Langmuir adsorption. Indeed at the hexane concentrations used, almost all the carbon dioxide is displaced from the Langmuir voids by hexane with the permeability approaching that predicted by Henry’s Law sorption alone. The greater reduction in CO2 permeability results in a net loss of selectivity from 16 to around 10. The use of eq 13 with parameters as given in Table 1 and bCH4 ) 0.13 allows for an estimation of the b value for hexane. The resulting value of 240 ( 100 atm-1 is 2 orders of magnitude greater than that for CO2 (Table 2) and indicates that hexane has a much stronger affinity for the microvoids than CO2. A correlation between the critical temperature of a component, that is its condensability, and Langmuir constant exists.4 Hexane has a critical temperature of 507 °C,20 compared to 31 °C for CO2. The determined Langmuir constant is slightly higher than that reported for other species with similar critical temperatures.4 This probably reflects the use of fugacities in eq 13, whereas other authors have utilized partial pressures. As the fugacity value is lower than the partial pressure, the corresponding b value will be higher. The change in permeability of both CO2 and CH4 as a result of exposure to toluene follows similar trends as can be seen in Figure 4. Equation 13 can again be used to back-calculate the Langmuir constant of toluene as around 980 atm-1. This value is greater than that observed for hexane, fitting with the expected relationship between critical temperature and Langmuir constant, as toluene has a critical temperature of 592 °C.
Figure 5. Change in carbon dioxide permeability as a function of time as hexane is added through a bubbler at 23 °C and then later bypassed.
Figure 6. Change in carbon dioxide permeability as a function of time as toluene is added through a bubbler at 15 °C and then later bypassed.
Typical plots of the permeability changes as a function of time are given in Figures 5 and 6. The permeability decline following exposure to either hydrocarbon occurs over a period of around 100 min. This decline can be modeled with simple first order kinetics, as have been employed by a number of other workers:21-23,17 j A(t) - P j A0 P t ) 1 - exp j A∞ - P j A0 τ P
[
( )]
(15)
where the characteristic time, τ, is related to the rate constant (K) through the membrane of thickness, δ δ2 (16) τ The permeability decline is reasonably estimated with such first order kinetics (Figures 5 and 6) with relaxation rate constants of 34 ( 23 × 10-9 m2 s-1 for hexane addition and 22 ( 16 × 10-9 m2 s-1 for toluene. Upon reaching steady state conditions with either hydrocarbon present, the feed gas was returned to dry conditions and the permeability of both CO2 and CH4 was observed to recover. For hexane, the permeabilities approach initial values after around 300 min, implying performance is recoverable and that the polymeric structure is not permanently altered by the presence of hexane (see Figure 5). Conversely, it was not K)
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possible to readily recover the original permeability after toluene exposure (Figure 6). It was necessary to heat the membrane unit to 70 °C and continue purging with dry gas to recover this performance. The slow recovery may imply that Langmuir desorption of hydrocarbon from the free volume voids has become the rate determining step. This is consistent with the high b values for these contaminants. Conclusion The performance of this typical fluorinated polyimide changes significantly upon exposure to a mixed gas stream. In a stream of CH4/CO2, carbon dioxide permeability falls due to competitive sorption relative to the pure gas value. Addition of impurity levels of hexane or toluene causes the permeability of both gases to fall further, again through competitive sorption. The results can readily be modeled using a dual sorption model with b values consistent with those in the literature. The rate of permeability decline upon hydrocarbon exposure can be modeled using simple first-order kinetics. The recovery of membrane performance once the hydrocarbon is no longer present is much slower, indicating that Langmuir desorption may have become the rate controlling step. Acknowledgment Financial support for this project has been provided by the CRC for Greenhouse Gas Technologies (CO2CRC), and this support is gratefully acknowledged. Infrastructure support is also provided by the Particulate Fluids Processing Centre, a Special Research Centre of the Australian Research Council. Literature Cited (1) Hao, J.; Rice, P. A.; Stern, S. A. Upgrading Low-Quality Natural Gas with H2S- and CO2-Selective Polymer Membranes Part I. Process Design and Economics of Membrane Stages without Recycle Streams. J. Membr. Sci. 2002, 209, 177. (2) Legatski, T. W.; Tooke, J. W.; Grundy, L. A. Approximating Natural Gas Composition. Pet. Refiner 1953, 32, 105. (3) Vu, D. Q.; Koros, W. J.; Miller, S. J. Fouling of Carbon Molecular Sieve Hollow-Fiber Membranes by Condensable Impurities in Carbon Dioxide-Methane Separation. Ind. Eng. Chem. 2003, 42, 1064. (4) Scholes, C. A.; Kentish, S. E.; Stevens, G. W. A Review of the Effects of Minor Components in Carbon Dioxide Capture Using Polymeric Gas Separation Membranes. Sep. Purif. ReV. 2009, 38, 1. (5) Schell, W. J.; Wensley, C. G. W.; Chen, M. S. K.; Venugopal, K. G.; Miller, B. D.; Stuart, J. A. Recent Advances in Cellulosic Membranes for Gas Separation and Pervaporation. Gas Sep. Purif. 1989, 3, 162.
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(6) White, L. S.; Blinka, T. A.; Kloczewski, H. A.; Wang, I. Properties of a Polyimide Gas Separation Membrane in Natural Gas Streams. J. Membr. Sci. 1995, 103, 73. (7) Langsam, M. Polyimides for Gas Separation. Plastics Eng. 1996, 36, 697. (8) Petropoulos, J. H. Quantitative Analysis of Gaseous Diffusion in Glassy Polymers. J. Polym. Sci., Polym. Phys. Ed. 1970, 8 (10), 1797. (9) Paul, D. R.; Koros, W. J. Effect of Partially Immobilizing Sorption on Permeability and the Diffusion Time Lag. J. Polym. Sci., Polym. Phys. Ed. 1976, 14 (4), 675. (10) Koros, W. J.; Paul, D. R.; Rocha, A. A. Carbon Dioxide Sorption and Transport in Polycarbonate. J. Polym. Sci., Polym. Phys. Ed. 1976, 14 (4), 687. (11) Kesting, R. E.; Fritzche, A. K., Polymeric Gas Separation; Wiley: New York, 1993. (12) Koros, W. J.; Chern, R. T.; Stannett, V.; Hopfenberg, H. B. J. Polym. Sci. Polym. Phys. Ed. 1981, 19, 1513. (13) Rogers, C. E. Solubility & Diffusivity. In Physics and Chemistry of the Organic Solid State; Fox, D., Labes, M. M. , Weissberger, A., Eds.; John Wiley & Sons: New York, 1965; p 975. (14) Stern, S. A.; Saxena, V. Concentration-Dependent Transport of Gases and Vapors in Glassy Polymers. J. Membr. Sci. 1980, 7 (1), 47. (15) Stern, S. A.; Saxena, V. Concentration-Dependent Transport of Gases and Vapors in Glassy Polymers. J. Membr. Sci. 1980, 7 (1980), 47. (16) Powell, C. E.; Duthie, X. J.; Kentish, S. E.; Qiao, G. G.; Stevens, G. W. Reversible diamine cross-linking of polyamide membranes. J. Membr. Sci. 2007, 291, 199. (17) Duthie, X.; Kentish, S.; Powell, C.; Nagai, K.; Qiao, G.; Stevens, G. Operating Temperature Effects on the Plasticization of Polyimide Gas Separation Membranes. J. Membr. Sci. 2007, 294 (1-2), 40. (18) Sandler, S. I. S. Chemical, Biochemical, and Engineering Thermodynamics; John Wiley & Sons: Hoboken, NJ, 2006. (19) Fujita, H. Diffusion in Polymer-Diluent Systems. Fortschr. Hochpolym. Forsch. 1961, 3 (1), 1. (20) Lange, N. A. Lange′s Handbook of Chemistry; McGraw-Hill: New York, 1992. (21) Berens, A. R.; Hopfenberg, H. B. Diffusion and Relaxation in Glassy Polymer Powders: 2. Separation of Diffusion and Relaxation Parameters. Polymer 1978, 19, 489. (22) Wind, J. D.; Sirard, S. M.; Paul, D. R.; Green, P. F.; Johnston, K. P.; Koros, W. J. Relaxation Dynamics of CO2 Diffusion, Sorption, and Polymer Swelling for Plasticized Polyimide Membranes. Macromolecules 2003, 36 (17), 6442. (23) Wessling, M.; Huisman, I.; Boomgaard, T. v. d.; Smolders, C. A. Dilation Kinetics of Glassy, Aromatic Polyimides Induced by Carbon Dioxide Sorption. J. Polym. Sci.: Part B, Polym. Phys. 1995, 33, 1371. (24) Duthie, X. A Study of Polyimide Gas Separation Membranes: Carbon Dioxide-Induced Plasticization and the Influence of Temperature. Ph.D. Thesis, University of Melbourne, 2007.
ReceiVed for reView October 11, 2008 ReVised manuscript receiVed January 20, 2009 Accepted March 12, 2009 IE801537G
