Volatile Organic Compound Liquid Recovery by the Dead End Gas

Article pubs.acs.org/IECR

Cite This: Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Volatile Organic Compound Liquid Recovery by the Dead End Gas Separation Membrane Process: Theory and Process Simulation Yong Ding* Air Liquide Advanced Technologies US LLC, 35A Cabot Road, Woburn, Massachusetts 01801, United States

Downloaded via DREXEL UNIV on March 19, 2019 at 14:24:34 (UTC). See https://pubs.acs.org/sharingguidelines for options on how to legitimately share published articles.

S Supporting Information *

ABSTRACT: Economically viable technologies are urgently needed for liquid VOC (volatile organic compound) recovery from refinery exhausts and flaring gases. This paper describes a potentially promising new membrane technology for such applications, i.e., separating organic liquids from gas mixtures containing condensable VOCs, such as LPG (liquid petroleum gas), and light permanent gases. The process entails utilizing a gas separation membrane operating in dead end mode to produce VOC liquids. The feed stream for the membrane is a gas mixture containing VOCs, the retentate stream is a VOC liquid rather than a gas or a gas−liquid mixture, and the permeate stream is a gas. The formation of the VOC liquids is controlled by the phase equilibrium in the feed side and dramatically enhanced by the high gas separation factors between the light permanent gases and the condensable gases and by removing the latent heat. Process simulations using the novel process for liquid propane production from methane/propane mixture and hydrogen/propane mixture were carried out, and the energy efficiency of the novel process in comparison to the cryogenic process was demonstrated.

1. INTRODUCTION Volatile organic compounds (VOCs), such as LPG (liquid petroleum gas), severely threaten human health and the ecological environment because most of them are toxic, mutagenic, and carcinogenic. The persistent increase of VOCs together with stringent regulations make the reduction of VOC emissions more imperative. Currently, there are a number of VOC treatment technologies available, such as incineration, condensation, biological degradation, absorption, adsorption, and catalytic oxidation. Among them, condensation as a nondestructive method is attractive because it offers the possibility to recover VOCs from the gas streams in a liquid state.1 However, a condensation process typically requires very low cryogenic operating temperatures in order to achieve the target.2,3 Membrane separation processes have been considered to be energy efficient processes because they do not require a thermal driving force to separate mixtures.4,5 Polymeric membranes have been widely utilized in gas separation applications, such as air dehydration, oxygen/nitrogen separation, hydrogen purification, and CO2, H2S, and higher hydrocarbon removal from natural gas and biogas.6−8 A membrane based process for VOC recovery has been deployed commercially to remove VOCs from nitrogen or air.9−11 The membrane materials utilized for VOC removal are siloxane based, such as polydimethylsiloxane and polyoctylmethylsiloxane.12 VOCs are enriched at the low pressure permeate side. The permeate gas needs to be recompressed and cooled to © XXXX American Chemical Society

produce the VOC liquids. The process has been commercially successful for certain applications containing medium amounts of VOC concentrations, but falls short for some other industrially important applications, such as separation of higher hydrocarbons from methane (for natural gas liquid recovery) and LPG separation from hydrogen mixtures.12−14 Recently, membrane condensers have been proposed and utilized to collect liquid, especially water, from gas streams.15−18 The membranes utilized for membrane condensers are porous nonselective and are utilized for both heat exchanging and gas transport. Since the membrane is nonselective toward the gases, the phase equilibrium change is due to the heat exchange rather than the gas composition change caused by the membrane gas separation. “Selective surface flow” is a separation process that uses a nanoporous membrane, especially carbon membrane, to separate gas mixtures.19−21 In selective surface flow separation, the selectivity is determined by preferential adsorption of certain components of the gas mixture on the surface of the membrane pores, as well as selective diffusion of the adsorbed molecules. The purified gas remains at the high pressure side and the condensable gases are collected at the permeate side, as is the case with the silicone based rubbery membranes. A Received: January 29, 2019 Revised: March 2, 2019 Accepted: March 3, 2019

A

DOI: 10.1021/acs.iecr.9b00586 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Article

Industrial & Engineering Chemistry Research practical issue for the selective surface flow separation is that the selectivity drops significantly once the condensable gases are removed. In this paper, we would like to propose a novel membrane process for VOC liquid production. Markedly different from conventional membrane processes using rubbery membranes, the new process collects the VOC liquids at the high pressure side and permeates the light permanent gases at the low pressure side. The novel process is a dead end membrane process utilizing molecular sieving gas separation membranes (either polymeric or inorganic). The dead end gas separation membrane process is defined as one that has only two gas streams, a feed gas stream and a permeate gas stream, i.e., a process with stage cut close to 100%. Stage cut is defined as the ratio between the permeate flow and the feed flow. The higher the stage cut, the more gas permeates through the membrane. For gas separations with product gases at the higher pressure side such as N2 and natural gases, a higher stage cut means more valuable gas is wasted. An optimized membrane process with product at the high pressure side would require the membrane to achieve the desired product purity at the lowest stage cut possible.22−24 As the stage cut approaches 100%, the permeate gas composition gradually approaches the feed gas composition, a dead end gas separation scenario. A dead end gas separation membrane process has been considered to be of no value since the feed gas composition and the permeate gas composition are identical. In contrary to the conventional belief, we have now found that a dead end gas membrane separation process is an extremely useful and energy efficient process for VOC liquid recovery from VOC gas mixtures. The process is schematically shown in Figure 1. In the process, a

to produce extremely robust and solvent resistant composite membranes, and their applications involving organic solvents have been explored. 26−28 These new membranes can potentially make the novel membrane process described in this article commercially viable. This paper will first describe the basic theory behind the novel process, and then describe process simulations for liquid propane recovery from methane/propane and hydrogen/ propane mixtures, two important industrial separations,29−32 to demonstrate the process viability. The paper further studies the effects of temperature, pressure, and feed gas composition on the process from the simulation. The energy required for the new membrane process based on the process simulation was compared with the energy required for a cryogenic process to produce the same liquid flow at the same pressure and temperatures. We hope this article can promote more research in this area and a much more energy efficient VOC liquid production process can be established.

2. METHODS 2.1. Calculation of VOC Dew Points. Calculation of dew point curves was performed using the HYSYS 7.2 software package (Aspen Technology, USA) with the Soave−Redlich− Kwong (SRK) equation of state. Two sets of gas mixtures were considered: H2/propane and methane/propane. 2.2. Simulation of Gas Separation Membranes. Gas permeabilities reported for Teflon AF1600 at 20 °C were used as the basis for the simulation (Table 1).33 Teflon AF 1600 is Table 1. Teflon AF 1600 Membrane Permeabilities,33 Permeances, and Activation Energies Used for the Simulation (at 20 °C) permeability coefficient (barrer) permeance (1 μm) (GPU) activation energy (kJ/mol)

H2

CH4

C3H8

500 500 14.5

41 41 17.2

2 2 20.8

an amorphous glassy copolymer of tetrafluoroethylene (TFE, 35%) and 2,2-bis(trifluoromethyl)-4,5-difluoro-1,3-dioxole (BDD, 65%). The unit for the gas permeability is barrer (1 barrer = 10−10 cm3(STP) cm/cm2 cmHg s). The membrane permeances were calculated from the membrane permeability by dividing the membrane thickness. For the purpose of the simulation, the membrane thickness was assumed to be 1 μm. The unit for permeance is the gas permeation unit (GPU, 1 GPU = 10−10 cm3(STP)/cm2 cmHg s). Gas permeance values at different temperatures are given by the following Arrhenius relationship:

Figure 1. “Dead end” gas separation membrane process for VOC liquid recovery.

i −Ep yz zz P = P0 expjjjj z (1) k RT { where P0 is a pre-exponential factor, Ep is the activation energy of permeance (J/mol), R is the gas constant (8.314 J/mol K), and T is the temperature (K). The activation energies for Teflon AF1600 used in this paper were measured in our laboratory and are listed in Table 1. The gas transport flux through the membrane is expressed as

gas separation membrane splits a gas feed stream into two streams: one liquid stream in the retentate and one gas stream in the permeate. The formation of the liquid in the retentate as the only product at the high pressure is enabled by constantly removing the permanent gases from the gas mixtures. However, the liquid formed from the process can potentially damage conventional commercially available polymeric membranes. Recently, we have successfully developed and commercialized PEEK (polyether ether ketone) based hollow fiber membranes which are extremely robust in the presence of organic liquids.25 We combined the PEEK hollow fiber substrate with other chemically resistant polymers/oligomers

Ji = Pi(xipf − yp ) i p

(2)

where Ji is the flux (cm /(cm ·s)) for gas component i, xi and yi are the molar fractions of gas component i in the feed and 3

B

2


Article

Industrial & Engineering Chemistry Research

Figure 2. Phase diagrams for methane/propane mixtures.

Figure 3. Phase diagrams for hydrogen/propane mixtures.

implemented by the method described by Coker et al.34,35 by taking into account hollow fiber bore side pressure drops. The membrane separation model has been used extensively for commercial membrane system design. The fiber inside diameter was assumed to be 200 μm, and the outside diameter was assumed to be 450 μm. The active fiber length was 1.5 m. The outside diameter for the membrane cartridge was 0.165 m, and the fiber packing density was 45%. A cross-flow model was used for the current process simulation.

permeate streams, and pf and pp are the pressures in the feed and permeate sides (cmHg). The phase equilibrium in the feed is described as xi = K izi

(3)

where zi is the mole fraction of component i in the liquid phase. The value of Ki depends on temperature and pressure. The temperature drop due to the Joule−Thomson (JT) effect is calculated as follows: ΔT = μJT (ph − pl )

(4)

3. RESULTS AND DISCUSSION 3.1. Phase Diagrams for VOC Gas Mixtures. Figure 2 shows the phase diagrams for methane/propane gas mixtures at different propane concentrations (volume percent). It can be seen that, at the same pressure, the higher the hydrocarbon

The membrane gas separation process simulation was carried out by a customized membrane gas separation model incorporated into HYSYS for hollow fiber gas separation membranes. The customized gas separation model was C


Article


liquid is intercepted, while for the current dead end separation process, the membrane separates the noncondensable gases and condensable gases. To differentiate the current membrane process from traditional depth filtration, a new term, “Perdensation”, from “permeation + condensation”, is coined and used throughout this paper. The Perdensation process described above is conceptually shown in Figure 1 and can be viewed as a process that is opposite to a pervaporation membrane process. Pervaporation is a membrane process involving separation of a liquid mixture through a dense selective layer of an asymmetric membrane.37,38 The term “pervaporation” was first coined from “permeation and evaporation” by Kober in 1917, about 100 years ago.39 Up to 1999, more than 90 industrial pervaporation units have been installed around the world and utilized for applications such as dehydration of organics, removal of trace organics from aqueous solutions, and organic−organic separations. The mechanism for pervaporation is shown in Figure 4a. Pervaporation can be conceptually viewed as a

dew point, the higher the propane concentration in the gas mixture. The dew point itself is defined as the temperature at which saturation of one component in gas mixtures occurs at a specific pressure. At the same pressure, if the temperature of the gas mixture is lowered for a saturated gas mixture, the dew point of the mixture is also lowered and the propane concentration in the saturated gas is lowered at the same time. As a result, two phases are formed by condensing the excess of propane in the original mixture in order to make the propane concentration in the saturated gas lower. On the other hand, if the temperature of the gas mixture is increased, i.e., higher hydrocarbon dew point, methane has to be partially removed from the original gas mixture to lower the methane concentration to keep the gas saturated at this new higher hydrocarbon dew point temperature. At the same pressure and temperature, if methane is continuously removed from the gas mixture, liquid propane is formed continuously in order to maintain the constant saturated propane concentration in the gas phase. Figure 3 shows the phase diagrams for hydrogen/propane gas mixtures at different propane concentrations (volume percent). Due to the extremely low liquefaction temperature for hydrogen and the low solubility of hydrogen in the liquid propane, the phase diagram does not show a phase envelope. The mixture exists as a gas phase on the right side of the curves and as a gas/liquid mixture on the left side of the curves. The general trends are the same for the hydrogen/propane mixtures. At the same pressure, the higher the hydrocarbon dew point, the higher the propane concentration in the mixture. At the same pressure, if the temperature of the gas mixture is lowered for a saturated gas mixture, i.e., lower hydrocarbon dew point, two phases are formed in order to keep the propane concentration at a lower saturated concentration. On the other hand, if the temperature of the gas mixture is increased, i.e., higher hydrocarbon dew point, hydrogen has to be partially removed from the gas mixture to lower the hydrogen concentration to keep the gas saturated at the new higher hydrocarbon dew point temperature. 3.2. Gas Separation Membrane and VOC Liquid Recovery. As shown in section 3.1, assuming the pressure is fixed, the phase equilibrium can be affected by either gas composition or temperature. To recover the VOC liquid from the gas mixtures, cryogenic cooling is conventionally utilized to affect the phase equilibrium by lowering the temperature of the gas mixtures. For example, Gupta and Verma studied the removal of VOCs by cryogenic condensation followed by adsorption.36 However, a cryogenic process is very energy intensive. Another method is to remove the less condensable components from the gas mixture. A gas separation membrane which exhibits good selectivities between the noncondensable gases and the condensable gases by allowing the noncondensable gases to permeate preferentially can be utilized for the purpose. Membrane processes have been widely accepted to be relatively energy efficient. It can be envisioned that by continuously removing the noncondensable gas from the gas mixture using the membrane, the phase equilibrium will be continuously pushed to the left to form more liquids. To the extreme, if all gas phase is removed from the feed stream through the membrane and only liquid is formed on the retentate side, the process becomes a dead end gas separation process, much like a depth filtration. The significant difference between the current process and the depth filtration is that, for the depth filtration, no gas separation takes place; only solid or

Figure 4. Comparison between pervaporation and Perdensation membrane separation processes. (a) Pervaporation process; (b) Perdensation process.

combination of two distinct processes: evaporation and gas separation.40,41 A liquid containing volatile components is fed into a membrane that has a dense separation layer. At the surface of the membrane, the volatile components along with some liquid vaporize and become a gas mixture. The gas mixture is then separated by the membrane selectively removing the volatile components. The mechanism for a Perdensation membrane separation process is conceptually shown in Figure 4b. In the Perdensation membrane separation process, a saturated gas containing condensable gases is fed into a membrane with a dense separation layer. The dense separation layer permeates the noncondensable gases faster than the condensable gases. Thus, the gas separation membrane continuously pushes the equilibrium of the feed gas into the liquid. The Perdensation membrane separation process can be conceptually viewed as a combination of two processes: gas separation and condensation. Molecular sizes for condensable gases (excluding water), such as propane and n-butane, are normally larger than those of noncondensable gases, such as methane, nitrogen, and oxygen. The separation material for the Perdensation process should exhibit gas pair selectivities based on the penetrant D


Article


formation of the liquid. It is beneficial for liquid formation to remove the heat generated from the negative JT effect from the membrane system. In the pervaporation process, the liquid covers the membrane separation area completely. However, in the Perdensation process, the membrane area should be free as much as possible from liquids so that the liquid will not interfere with the membrane gas separation. This would require the membrane cartridge design to take the liquid drainage into account, by lowering the membrane packing density and maintaining the liquid level well below the membrane cartridge, so that the liquid formed during the process will not cover the effective membrane areas. 3.4. Process Simulation and Comparison with Cryogenic Process. 3.4.1. Process Simulation Flow Diagrams. Process simulations were carried out for liquid propane production from methane/propane mixtures and hydrogen/propane mixtures. The process flow diagram (PFD) for the Perdensation process is shown in Figure 6. The

sizes. Polymeric membrane materials exhibiting size selectivities are glassy polymers.4,7 The most suitable glassy membrane materials for Perdensation separation are commercially available amorphous perfluorinated polymers, such as Teflon AF 1600 and Teflon AF 2400, copolymers of 2,2bistrifluoromethyl-4,5-difluoro-1,3-dioxole (PDD) and tetrafluoroethylene (TFE) with 65 and 87 mol % PDD, respectively (Figure 5), and Hyflon AD, copolymers of 2,2,4-trifluoro-5-

Figure 5. Chemical structures of typical amorphous perfluoropolymers.

trifluoromethoxy-1,3-dioxole and tetrafluoroethylene (Figure 5). Both Teflon AF and Hyflon AD polymer series exhibit high permeabilities for noncondensable gases and low permeabilities for condensable VOC gases.33,42−46 For example, for Teflon AF 1600, the permeability for nitrogen is 79 barrer and the nitrogen/propane selectivity is 21 at 30 °C.46 Perfluorinated amorphous polymers are readily soluble in perfluorinated solvents, and a composite membrane with a thin perfluorinated dense separation layer can be readily fabricated by a dip-coating technique.47 The membrane substrates used for the formation of the composite membranes can be polymeric or inorganic. Polymeric membrane substrates are preferred because they can be easily scaled up. However, the polymeric substrates should not be affected or damaged by the liquids formed during the Perdensation process. An ideal candidate for the substrate material is PEEK, a semicrystalline polymer not soluble in almost all conventional organic solvents.48−50 PEEK hollow fiber membrane substrates have been successfully prepared and commercialized by a melt extrusion process.28,51 Composite hollow fiber membranes with thin Teflon AF amorphous polymer as the separation layer deposited on PEEK hollow fiber substrates have been successfully commercialized. 3.3. Latent Heat and Perdensation. Condensation of the condensable gases occurs at the dew point of a gas mixture.15 The permeation of the noncondensable gases effectively increases the partial pressure of the condensable gases, thus promoting the formation of liquids by decreasing the partial pressure of the condensable component. However, latent heat is generated from the condensation,52 which causes the temperature of the gas mixture and liquid to rise. The increase of the temperature causes the vapor pressure to increase and therefore decreases the efficiency of liquid formation. Consequently, it is desirable to remove the latent heat during the Perdensation process. In the Perdensation process, gas separation takes place with a feed gas mixture at the high pressure side and a permeation gas mixture at the low pressure side, resulting in the Joule− Thomson (JT) effect. For most gases, such as CO2 and methane, the JT effect absorbs heat and reduces the temperature, and thus reduces the vapor pressure of the liquid, which helps the formation of the liquid. On the other hand, for some gases, such as hydrogen, the JT effect releases the heat and increases the temperature, and thus increases the vapor pressure of the liquid, which is detrimental to the

Figure 6. PFD for the Perdensation process. The phase separator is a virtual one to maintain the feed gas at saturation.

Perdensation membrane process is a combination of three unit operations: gas separation membrane unit, phase separator unit, and a heat exchanger unit (removing latent heat). From the HYSYS simulation point of view, a membrane separation unit does not contain the phase separation function. A virtual phase separation unit has to be placed to simulate the phase separation inside the membrane module, as shown in Figure 6. The rationale in placing the virtual phase separator in front of the membrane in the PFD (Figure 6) is to keep the feed gas at the saturated state all the time. The cryogenic process is a combination of two unit operations: heat exchange and phase separation. The membrane utilized for the simulation was a composite hollow fiber membrane prepared from Teflon AF 1600 and PEEK hollow fiber substrate. In general, the separation layer thickness can be adjusted from 30 nm up to 2 μm for composite gas separation membranes. The selection of the membrane thickness is determined by the intended application, available pressure, and pressure drop caused by the membrane device. For this feasibility study, the separation layer thickness was assumed to be 1 μm. Pinnau and Toy reported that the presence of heavy hydrocarbons can cause the plasticization of Teflon AF.43 However, unlike PTMSP and E


Article


Figure 7. HYSYS simulation PFD for the Perdensation process.

3.4.2. Liquid Propane Production from CH4−C3H8 Mixture by the Perdensation Process. First, we looked at the effects of temperature, pressure, and feed gas compositions on the liquid propane production. The results are listed in Tables S1−S3 in the Supporting Information. For temperature and pressure effect studies, the feed gas composition was fixed at 35% propane by mole and 65% methane by mole and the flow rate was fixed at 149.4 kg mol/h. The permeate pressure was fixed at 137.9 kPa. As the temperature increased from 20 to 50 °C, the liquid propane product flow rate decreased from 61.8 to 51.73 kg mol/h (the pressure was fixed at 3000 kPa). At the same time, the membrane area required increased from 604 to 1057 m2, and the C3H8 concentration in the gas stream also increased from 2.53 to 7.2 mol %. The increase of permeate gas flow and higher C3H8 concentration in the gas stream are due to the higher vapor pressure of propane and lower methane/propane selectivity at higher temperatures. As the pressure increased from 2000 to 5000 kPa, the liquid propane product flow rate increased from 43.05 to 65.45 kg mol/h (the temperature was fixed at 30 °C). At the same time, the membrane area required decreased from 3026 to 456 m2, and the C3H8 concentration in the permeate gas stream also decreased from 10.28 to 2.87 mol %. The decrease in flow and lower C3H8 concentration in the gas stream are due to the lower vapor pressure of propane at higher pressure. The lower membrane area required at higher pressure is due to the increased driving force at higher feed pressure. As the propane concentration in the feed increased from 5.0 to 35.0 mol %, the liquid propane product flow rate increased from 1.88 to 53.82 kg mol/h (the temperature was fixed at 30 °C and the feed pressure was fixed at 3000 kPa). At the same time, the membrane area required decreased from 1539 to 1122 m2, and the C3H8 concentration in the gas stream increased only slightly, from 4.03 to 4.83 mol %. The decrease in the membrane area required is a reflection of the decreased volume of methane in the feed mixture. A commercial size hollow fiber membrane cartridge typically contains a membrane area from 50 to 400 m2, depending on the cartridge diameter. Therefore, for the cases studied here, only a couple of hollow fiber membrane cartridges are required. To demonstrate the advantage of the current Perdensation process for liquid production from methane/propane mixture, a cryogenic process for propane liquid production was also simulated by HYSYS. The feed gas flow rate was fixed at 149.4

PIM related materials, Teflon AF material still maintains excellent size selectivity even though the permeability is reduced.43,53−55 In this study, membrane permeabilities reported in the literature for Teflon AF 1600 were used in the simulation to demonstrate the process feasibility.33 The HYSYS process flow diagram for the Perdensation process is shown in Figure 7. Gas stream 101 is the raw feed gas. V-100 is a virtual phase separator to separate liquid and gas. The saturated gas stream 104 is fed into the gas separation membrane OP-100 to produce a permeate gas stream 105 and a supersaturated stream 106. A heat exchanger E-100 is used to simulate the heat that is being taken out of the membrane by cooling. For the current purpose, the cooling is to maintain the feed gas temperature constant, i.e., the same as the feed gas temperature. The supersaturated gas is recycled back to the phase separator V-100 to produce a liquid stream 103. During the simulation, the membrane area is adjusted until the total flow rates for streams 105 and 103 are essentially equal to the flow rate for stream 101 (difference is less than 0.2%). Figure 8 shows the process flow diagram in HYSYS for liquid production using the cryogenic process. The feed gas

Figure 8. HYSYS simulation PFD for the cryogenic process.

stream 201 is chilled with a heat exchanger E-200 to produce a two phase stream 202, which is separated into a liquid stream 203 and a gas stream 204 by a phase separator V-200. In the current study, the temperatures of the two phase streams are adjusted until the liquid flow rate for stream 203 reaches the targeted value. F


Article


occurs because the removal of methane from the gas mixture allows for liquid propane to be formed at higher temperatures. 3.4.3. Liquid Propane Production from H2−C3H8 Mixture by the Perdensation Process. Separation of propane from hydrogen is another extremely important application.29,31,56 In this section, we would like to demonstrate that the Perdensation process is an energy efficient membrane process for liquid propane recovery from the hydrogen/propane mixture. The effects of temperature, pressure, and feed gas compositions on liquid propane production using the Perdensation process for the hydrogen/propane mixture are listed in Tables S5−S7 in the Supporting Information. For temperature and pressure effect studies, the feed gas composition was fixed at 40% propane by mole and the flow rate was fixed at 149.4 kg mol/h. The permeate pressure was fixed at 137.9 kPa. As the temperature increased from 20 to 50 °C (the pressure was fixed at 3000 kPa), the liquid propane product flow rate decreased from 60.83 to 58.9 kg mol/h. At the same time, the membrane area required increased from 112 to 196 m2, and the C3H8 concentration in the permeate gas stream also increased from 0.34 to 2.26 mol %. The increase of flow and higher C3H8 concentration in the permeate gas stream are due to the higher vapor pressure of propane at higher temperature and lower hydrogen/propane selectivity at higher temperatures. The membrane areas required for the hydrogen/ propane mixture are much lower than that required for the methane/propane mixture. This is due to the fact that hydrogen permeates through the membrane at a much higher rate than methane does at the same temperature and pressure differential. As the pressure increased from 2000 to 5000 kPa, the liquid propane product flow rate increased from 59.15 to 60.59 kg mol/h (the temperature was fixed at 30 °C). At the same time, the membrane area required decreased from 371 to 119 m2, and the C3H8 concentration in the permeate gas stream also decreased from 1.51 to 0.55 mol %. The decrease of permeate gas flow and lower C3H8 concentration in the gas stream are due to the lower vapor pressure of propane at higher pressure. The lower membrane area required at higher pressure is due to the increased driving force at higher feed pressures. As the propane concentration increased from 5.0 to 40.0 mol %, the liquid propane product flow rate increased from 6.99 to 60.45 kg mol/h (the temperature was fixed at 30 °C and the feed pressure was fixed at 3000 kPa). At the same time, the membrane area required decreased from 159 to 129 m2, and the C3H8 concentration in the gas stream increased only slightly, from 0.42 to 0.65 mol %. The decrease of the

kg mol/h, the same as the Perdensation process. The feed temperature was fixed at 30 °C, and the feed pressure was fixed at 3000 kPa. Table 2 lists the specific cooling duties (kJ/kg mol Table 2. Energy Consumption Difference between Cryogenic Process and Perdensation Process for Propane Liquid Production from Methane/Propane Mixture (Feed at 3000 kPa and 30 °C) energy consumption (kJ/kg mol) feed C3H8 (vol %)

liquid flow (kg mol/h)

cryogenic cooling temp (°C)

cooling process

Perdensation process

5.0 10.0 15.0 20.0 25.0 30.0 35.0

1.88 11.1 19.18 28.5 37.15 45.88 53.82

−47.36 −42.03 −35.90 −32.60 −28.60 −24.96 −20.45

272 074 55 018 35 912 28 007 23 997 21 463 19 640

31 457 13 234 9 437 10 937 10 135 10 514 10 621

liquid) required for the cryogenic process and the cooling duties required for the Perdensation process for producing the same amount of propane liquid flow. Table 2 also lists the cooling temperatures required for the cryogenic process. It can be seen that the lower the propane concentration in the feed, the lower the temperature required for the cryogenic process. The difference in cooling duty requirement between the cryogenic process and the Perdensation process is much larger at lower propane concentration in the feed. For example, at a propane feed concentration of 5.0 mol %, the cooling duty required for the cryogenic process is 8.65 times of that for the Perdensation process. On the other hand, at a propane feed concentration of 35.0%, the cooling duty required for the cryogenic process is about 1.85 times that for the Perdensation process. In all cases, the energy required for the cryogenic process is significantly higher than that for the Perdensation process. The compositions for the liquids and gases produced by the cryogenic and Perdensation processes are listed in Table S4 in the Supporting Information. The liquids produced by the cryogenic process contain more methane than that produced by the Perdensation process. The difference becomes even larger for feed gas streams containing less propane. It can also be seen that the gas streams produced by the cryogenic process contain more propane than the gas streams produced by the Perdensation process. The difference becomes larger for feed gas streams containing more propane. This favorable trend

Table 3. Energy Consumption Difference between Cryogenic Process and Perdensation Process for Propane Liquid Production from Hydrogen/Propane Mixture (Feed at 3000 kPa and 30 °C)

feed C3H8 (vol %) 5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0

cryogenic process

Perdensation process

liquid flow (kg mol/h) cooling temp (°C) energy consumption (kJ/kg mol)

liquid flow (kg mol/h) energy consumption (kJ/kg mol)

6.99 14.62 22.28 29.92 37.5 45.11 52.67 60.22

−86.20 −87.50 −90.50 −94.00 −95.50 −100.00 −100.00 −100.00

95 050 58 023 46 724 41 344 37 867 35 957 34 023 32 547 G

6.99 14.62 22.28 29.92 37.50 45.18 52.86 60.45

29 013 20 404 17 666 16 270 15 251 14 679 14 253 13 919 DOI: 10.1021/acs.iecr.9b00586 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Article


propane concentration, due to the need to recompress a large amount of permeate gas. However, the difference becomes much larger with feed gas containing more than 20% propane, due to the recompressed amount of permeate gas becoming less. The component compositions for the liquids and gases produced by the cryogenic process and the Perdensation process are listed in Table S8 in the Supporting Information. The propane concentrations in the liquids produced by both methods are about the same. The propane concentrations in the gas streams produced by the cryogenic process are slightly lower in comparison to that of the Perdensation process. The energy required for the Perdensation process is only about 40% of that required for a cryogenic process.

membrane area required is a reflection of the decreased volume of the hydrogen in the feed mixture. For the HYSYS simulation of the cryogenic process for propane liquid production from hydrogen/propane mixtures, the feed gas flow rate was fixed at 149.4 kg mol/h, the same as the Perdensation process. The feed temperature was fixed at 30 °C, and the feed pressure was fixed at 3000 kPa. Table 3 lists the specific cooling energy required for the cryogenic process and the cooling energy required for the Perdensation process to produce the same amount of propane liquid flow. Table 3 also lists the cooling temperature required for the cryogenic process. For feed gas containing >30 mol % propane, the liquid propane flow rate cannot reach the same liquid propane flow rate as that produced by the Perdensation process, even by further reducing the temperature. A temperature of −100 °C was used for these gas streams. In contrast to the cases for the methane/propane mixture, a lower cooling temperature is required in the cryogenic process for gases containing higher propane concentration in the feed. The cooling duty requirement difference between the cryogenic and Perdensation processes is much larger at lower propane concentration, which is similar to the cases for the methane/propane mixtures. For example, at a propane feed concentration of 5.0 mol %, the cooling duty required for the cryogenic process is 3.28 times that for the Perdensation process. On the other hand, at a propane feed concentration of 35.0%, the cooling duty required for the cryogenic process is about 2.38 times that for the Perdensation process. In all cases, the energy required for the cryogenic process is significantly higher than that for the Perdensation process. Hydrogen gas is produced at the low pressure permeate side for the Perdensation process, and there is an energy loss due to the pressure drop. A fair energy consumption comparison should count the energy consumption required to recompress the permeate gas back to the original pressure if the permeate gas is not utilized at the lower pressure. The energy consumptions for the recompression of the permeate hydrogen gas were further calculated by HYSYS simulation. An adiabatic compressor with an efficiency of 75% was used to compress the permeate gases, and 90% of the thermal energy from the compressor was assumed to be recovered, such as through heat exchange or a turbo-expander.57,58 The energy consumption between the Perdensation process and the cryogenic process is shown in Table 4, taking into account energy consumption for recompression of the permeate hydrogen gas from the membrane process. The difference is smaller at low feed

4. CONCLUSIONS VOC recovery by condensation as a nondestructive method is attractive because it offers the ability to recover VOCs from gas streams in the liquid state.59 However, the condensation process often requires a very low (cryogenic) condensation temperature in order to achieve the target.1,3,60,61 A new membrane process, Perdensation, is proposed and demonstrated in this paper by process simulation, as a nondestructive process to capture the condensable gases as liquids. It provides an energy efficient operation to recover organic liquids from many different gas streams, such as LPG from hydrogen, natural gas liquids (NGLs) from natural gases, and recovery of VOCs from air. Process simulation using HYSYS has demonstrated that the energy required for liquid production using the Perdensation process is significantly lower than that for the cryogenic process. In some cases, the energy required for the Perdensation process is less than 15% of that required for the cryogenic process.

■

ASSOCIATED CONTENT

* Supporting Information S

The Supporting Information is available free of charge on the ACS Publications website at DOI: 10.1021/acs.iecr.9b00586.

■

Table 4. Energy Consumption Difference between Cryogenic Process and Perdensation Process for Propane Liquid Production from Hydrogen/Propane Mixture (Feed at 3000 kPa and 30 °C) with Permeate Gas Recompression

Effects of temperature, pressure, and feed gas composition on propane liquid production from methane/propane and hydrogen/propane mixtures; comparison of product compositions between cryogenic and Perdensation processes for propane liquid production from methane/propane and hydrogen/propane mixtures (PDF)

AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. ORCID

feed C3H8 (vol %)

energy consumption from cooling process (kJ/kg mol)

energy consumption from Perdensation process with permeate recompression (kJ/kg mol)

5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0

95 050 58 023 46 724 41 344 37 867 35 957 34 023 32 547

70 200 37 606 27 702 23 035 20 232 18 429 17 153 16 243

Yong Ding: 0000-0001-9086-8193 Notes

The author declares no competing financial interest.

■

ACKNOWLEDGMENTS This work was carried out in part as a condition for the author’s employment at Air Liquide. The author would like to thank Air Liquide Advanced Technologies US, LLC, for allowing the publication of the work, and Dr. Ramy Swaidan for proofreading the final manuscript. H


Article


■

(23) Huang, Y.; Merkel, T. C.; Baker, R. W. Pressure ratio and its impact on membrane gas separation processes. J. Membr. Sci. 2014, 463, 33−40. (24) Brunetti, A.; Scura, F.; Barbieri, G.; Drioli, E. Membrane technologies for CO2 separation. J. Membr. Sci. 2010, 359, 115−125. (25) Ding, Y.; Bikson, B. Preparaion and characterization of semicrystalline poly(ether ether ketone) hollow fiber membranes. J. Membr. Sci. 2010, 357, 192−198. (26) Ding, Y.; Bikson, B. Composite perfluorohydrocarbon membranes, their preparation and use. U.S. Patent 9,610,547, 2017. (27) Li, S.; Rocha, D. J.; Zhou, S. J.; Meyer, H. S.; Bikson, B.; Ding, Y. Post-combustion CO2 capture using super-hydrophobic, polyether ether ketone, hollow fiber membrane contactors. J. Membr. Sci. 2013, 430, 79−86. (28) Li, S.; Pyrzynski, T. J.; Klinghoffer, N. B.; Tamale, T.; Zhong, Y.; Aderhold, J. L.; Zhou, S. J.; Meyer, H.; Ding, Y.; Bikson, B. Scaleup of PEEK hollow fiber membrane contactor for post-combustion CO2 capture. J. Membr. Sci. 2017, 527, 92−101. (29) Pinnau, I.; He, Z. Pure- and mixed-gas permeation properties of polydimethylsiloxnae for hydrocarbon/methane and hydrocarbon/ hydrogen separation. J. Membr. Sci. 2004, 244, 227−233. (30) Yang, Y.; Burke, N.; Ali, S.; Huang, S.; Lim, S.; Zhu, Y. Experimental studies of hydrocarbon separation on zeolites, activated carbons and MOFs for applications in natural gas processing. RSC Adv. 2017, 7, 12629−12638. (31) Lin, H.; Wagner, E. V.; Freeman, B. D.; Toy, L. G.; Gupta, R. P. Plasticization-enhanced hydrogen purification using polymeric membranes. Science 2006, 311, 639−642. (32) Takht Ravanchi, M.; Kaghazchi, T.; Kargari, A. Application of membrane separation processes in petrochemical industry: a review. Desalination 2009, 235, 199−244. (33) Alentiev, A. Y.; Yampolskii, Y. P.; Shantarovich, V. P.; Nemser, S. M.; Plate, N. A. High transport parameters and free volume of perfluorodioole copolymers. J. Membr. Sci. 1997, 126, 123−132. (34) Coker, D. T.; Freeman, B. D.; Fleming, G. K. Modeling Multicomponent Gas Separation Using Hollow-Fiber Membrane Cohtactors. AIChE J. 1998, 44, 1289−1302. (35) Coker, D. T.; Allen, T.; Freeman, B. D.; Fleming, G. K. Nonisothermal model for gas separation hollow fiber membranes. AIChE J. 1999, 45, 1451−1468. (36) Gupta, V. K.; Verma, N. Removal of volatile organic compounds by cryogenic condensation followed by adsorption. Chem. Eng. Sci. 2002, 57, 2679−2696. (37) Roy, S.; Singha, N. R. Polymeric nanocomposite membranes for next generation pervaporation process: Strategies, challenges and future prospects. Membranes 2017, 7, 53−116. (38) Ong, Y. K.; Shi, G. M.; Le, N. L.; Tang, Y. P.; Zuo, J.; Nunes, S. P.; Chung, T.-S. Recent membrane development for pervaporation process. Prog. Polym. Sci. 2016, 57, 1−31. (39) Kober, P. A. Pervaporation, perstillation and percrystallization. J. Am. Chem. Soc. 1917, 39, 944−948. (40) Wijmans, J. G.; Baker, R. W. A simple predictive treatment of the permeation process in pervaporation. J. Membr. Sci. 1993, 79, 101−113. (41) Baker, R. W.; Wijmans, J. G.; Huang, Y. Permeability, permeance and selectivity: A preferred way of reporting pervaporation performance data. J. Membr. Sci. 2010, 348, 346−352. (42) Alentiev, A. Y.; Shantarovich, V. P.; Merkel, T. C.; Bondar, V. I.; Freeman, B. D.; Yampolskii, Y. P. Gas and vapor sorption, permeation, and diffusion in glassy amorphous Teflon AF1600. Macromolecules 2002, 35, 9513−9522. (43) Pinnau, I.; Toy, L. G. Gas and vapor transport properties of amorphous perfluorinated copolymer membranes based on 2,2bistrifluoromethyl-4,5-difluoro-1,3-dioxole/tetratfluoroethylene. J. Membr. Sci. 1996, 109, 125−133. (44) Arcella, V.; Colaianna, P.; Maccone, P.; Sanguineti, A.; Gordano, A.; Clarizia, G.; Drioli, E. A study on a perfluoropolymer purification and its application to membrane formation. J. Membr. Sci. 1999, 163, 203−209.

REFERENCES

(1) Belaissaoui, B.; Le Moullec, Y.; Favre, E. Energy efficiency of a hybrid membrane/condensation process for VOC (volatile organic compounds) recovery from air: A generic approach. Energy 2016, 95, 291−302. (2) He, T.; Ju, Y. Design and optimization of a novel mixed refrigerant cycle integrated with NGL recovery process for small scale LNG plant. Ind. Eng. Chem. Res. 2014, 53, 5545−5553. (3) Parthasarathy, G.; El-Halwagi, M. M. Optimum mass integration strategies for condensation and allocation of multicomponent VOCs. Chem. Eng. Sci. 2000, 55, 881−895. (4) Sanders, D. F.; Smith, Z. P.; Guo, R.; Robeson, L. M.; McGrath, J. E.; Paul, D. R.; Freeman, B. D. Energy-efficient polymeric gas separation membranes for a sustainable future: A review. Polymer 2013, 54, 4729−4761. (5) Park, H. B.; Kamcev, J.; Robeson, L. M.; Elimelech, M.; Freeman, B. D. Maximizing the right stuff: The trade-off between membrane permeability and selectivity. Science 2017, 356, eaab0530. (6) Zhang, Y.; Sunarso, J.; Liu, S.; Wang, R. Current status and development of membranes for CO2/CH4 separation: A review. Int. J. Greenhouse Gas Control 2013, 12, 84−107. (7) Baker, R. W.; Low, B. T. Gas separation membrane materials: A perspective. Macromolecules 2014, 47, 6999−7013. (8) Chen, X. Y.; Vinh-Thang, H.; Ramirez, A. A.; Rodrigue, D.; Kaliaguine, S. Membrane gas separation technologies for biogas upgrading. RSC Adv. 2015, 5, 24399−24448. (9) Baker, R. W. Membrane Technology and Applications, 2nd ed.; McGraw-Hill: New York, 2000. (10) Poddar, T. K.; Sirkar, K. K. A hybrid of vapor permeation and membrane-based absortion-stripping for VOC removal and recovery from gaseous emissions. J. Membr. Sci. 1997, 132, 229−233. (11) Gales, L.; Mendes, A.; Costa, C. Removal of acetone, ethyl acetate and ethanol vapors from air using a hollow fiber PDMS membrane module. J. Membr. Sci. 2002, 197, 211−222. (12) Schultz, J.; Peinemann, K. V. Membranes for separation of higher hydrocarbons from methane. J. Membr. Sci. 1996, 110, 37−45. (13) Grinevich, Y.; Starannikova, L.; Yampol’skii, Y.; Gringol’ts, M.; Finkel’shtein, E. Membrane separation of gaseous C1-C4 alkanes. Pet. Chem. 2011, 51, 585−594. (14) Elsherif, M.; Manan, Z. A.; Kamsah, M. Z. State-of-the-art of hydrogen management in refinery and industrial process plants. J. Nat. Gas Sci. Eng. 2015, 24, 346−356. (15) Crabtree, E. W.; El-Halwagi, M. M.; Dunn, R. F. Synthesis of hybrid gas permeation membrane/condensation systems for pollution prevention. J. Air Waste Manage. Assoc. 1998, 48, 616−626. (16) Brunetti, A.; Santoro, S.; Macedonio, F.; Figoli, A.; Drioli, E.; Barbieri, G. Waste gaseous streams: From environmental issue to source of water by using membrane condensers. Clean: Soil, Air, Water 2014, 42, 1145−1153. (17) Hu, H. W.; Tang, G. H.; Niu, D. Wettability modified nanoporous ceramic membrane for simultaneous residual heat and condensate recovery. Sci. Rep. 2016, 6, 27274. (18) Yan, S.; Zhao, S.; Wardhaugh, L.; Feron, P. H. M. Innovative use of membrane contactor as condenser for heat recovery in carbon capture. Environ. Sci. Technol. 2015, 49, 2532−2540. (19) Rao, M. B.; Sircar, S. Nanoporous carbon membranes for separation of gas mixtures by selective surface flow. J. Membr. Sci. 1993, 85, 253−264. (20) Arruebo, M.; Coronas, J.; Menendez, M.; Santamaria, J. Separation of hydrocarbons from natural gas using silicalite membranes. Sep. Purif. Technol. 2001, 25, 275−286. (21) Neubauer, K.; Dragomirova, R.; Stohr, M.; Mothes, R.; Lubenau, U.; Paschek, D.; Wohlrab, S. Combination of membrane separation and gas condensation for advanced natural gas conditioning. J. Membr. Sci. 2014, 453, 100−107. (22) Alshehri, A. K.; Lai, Z. Attainability and minimum energy of multiple-stage cascade membrane systems. J. Membr. Sci. 2015, 495, 284−293. I


Article

Industrial & Engineering Chemistry Research (45) Shantarovich, V. P.; Kevdina, I. B.; Yampolskii, Y. P.; Alentiev, A. Y. Positron annihilation lifetime study of high and low free volume glassy polymers: Effects of free volume sizes on the permeability and permselectivity. Macromolecules 2000, 33, 7453−7466. (46) Bikson, B.; Ding, Y.; Lerous, J. K.; Nelson, J. K. Gas separation using membranes formed from blends of perfluorinated polymers. U.S. Patent 6,723,152, 2004. (47) Nelson, J. K.; Bikson, B.; Macheras, J. T. Method of preparing composite gas separation membrane from perfluoropolymers. U.S. Patent 6,540,813, 2003. (48) da Silva Burgal, J.; Peeva, L.; Livingston, A. Negligible ageing in poly(ether ether ketone) membranes widens application range for solvent processing. J. Membr. Sci. 2017, 525, 48−56. (49) Werner, P.; Verdejo, R.; Wollecke, F.; Altstadt, V.; Sandler, J. K. W.; Shaffer, M. S. P. Carbon nanofibers allow foaming of semicrystalline poly(ether ether ketone). Adv. Mater. 2005, 17, 2864−2869. (50) da Silva Burgal, J.; Peeva, L.; Marchetti, P.; Livingston, A. Controlling molecular weight cut-off of PEEK nanofiltraion membranes using a drying method. J. Membr. Sci. 2015, 493, 524− 538. (51) Ding, Y.; Bikson, B. Preparation and characterization of semicrystalline poly(ether ether ketone) hollow fiber membranes. J. Membr. Sci. 2010, 357, 192−198. (52) Lin, C. X.; Wang, D.; Bao, A. Numerical modeling and simulation of condensation heat transfer of a flue gas in a bundle of transport membrane tubes. Int. J. Heat Mass Transfer 2013, 60, 41− 50. (53) Nemser, S. M. Cyclic membrane separation process. U.S. Patent 6,887,300, 2005. (54) Pinnau, I.; He, Z.; DaCosta, A. R.; Amo, K. D.; Daniels, R. Gas separation using C3+ hydrocarbon-resistant membranes. U.S. Patent 6,361,582, 2002. (55) Scholes, C. A.; Stevens, G. W.; Kentish, S. E. Impact of heavy hydrocarbons on natural gas sweetening using perfluorinated polymeric membranes. Ind. Eng. Chem. Res. 2016, 55, 7696−7703. (56) Ockwig, N. W.; Nenoff, T. M. Membranes for hydrogen separation. Chem. Rev. 2007, 107, 4078−4110. (57) Zhai, H.; Rubin, E. S. Techno-economic assessment of polymer membrane systems for postcombustion carbon capture at coal-fired power plants. Environ. Sci. Technol. 2013, 47, 3006−3014. (58) Yang, D.; Wang, Z.; Wang, J.; Wang, S. Potential of two-stage membrane system with recycle stream for CO2 capture from postcombustion gas. Energy Fuels 2009, 23, 4755−4762. (59) Ojala, S.; Koivikko, N.; Ainassaari, K.; Brahmi, R.; Keiski, R. L.; Laitinen, T.; Mouammine, A.; Seelam, P. K.; Laassiri, S. Utilization of volatile organic compounds as an alternative for destrutive abatement. Catalysts 2015, 5, 1092−1151. (60) Rebollar-Perez, G.; Carretier, E.; Lesage, N.; Moulin, P. Volatile organic compound (VOC) removal by vapor permeation at low VOC concentrations: Laboratory scale results and modeling for scale up. Membranes 2011, 1, 80−90. (61) Kimmerle, K.; Bell, C. M.; Gudernatsch, W.; Chmiel, H. Solvent recovery from air. J. Membr. Sci. 1988, 36, 477−488.

J


Volatile Organic Compound Liquid Recovery by the Dead End Gas

Recommend Documents