Hydrogen Recovery from Methanol Steam Reforming in a Dense

2420

Ind. Eng. Chem. Res. 2004, 43, 2420-2432

Hydrogen Recovery from Methanol Steam Reforming in a Dense Membrane Reactor: Simulation Study Fausto Gallucci, Luca Paturzo, and Angelo Basile* Institute on Membrane Technology, ITM-CNR, c/o University of Calabria, Via P. Bucci, cubo 17/C, I-87030 Rende (CS), Italy

The methanol steam-reforming reaction to produce synthesis gas has been studied theoretically. A mathematical model has been formulated for a traditional reactor and then applied to a palladium membrane reactor in which the hydrogen production is increased by removing the hydrogen produced from the reaction mixture through a highly selective (100%) palladiumbased membrane. In agreement with the literature, the results show that it is possible to obtain both higher conversions of methanol and higher hydrogen selectivities compared to those obtained in a traditional reactor operating at the same experimental conditions. The theoretical analysis provides a set of parameters that permit one to maximize the (pure) hydrogen production and/ or methanol conversion if adopted in an experimental device. 1. Introduction The increasing levels of CO2 and other greenhouse gases in the atmosphere, 20-25% of which is due to automobiles, have led vehicle manufacturers to study different solutions for reducing these dangerous emissions. In this effort, electric drive trains with batteries and fuel-cell drive systems are of major interest as possible efficient solutions to the problem. In particular, integrated fuel-cell systems are emerging, in which pure hydrogen feeding a polymer exchange membrane fuel cell (PEMFC) is produced by hydrocarbon (gasoline, methane, methanol, etc.) reformation and subsequent hydrogen purification. The use of methanol as an onboard hydrogen source is advantageous when considering distribution infrastructures, safety aspects, and vehicle driving ranges. A fuel-cell drive system based on methanol as the fuel consists of a methanol steam reformer and a gas-cleaning unit that reduces the CO content of the hydrogen-rich gas and feeds the fuel cell. The reformer is equipped with a catalytic burner that provides the process heat for the reformer itself and converts all burnable gases in the flue gas into water and carbon dioxide. Recent studies of the methanol steam-reforming reaction have considered the application of a membrane reactor that is highly selective toward the hydrogen produced: such an approach might allow for the replacement of the traditional reformer and the following gas-cleaning unit.1 In the membrane reactor device, methanol conversion via steam reforming and hydrogen purification are combined, so that the pure hydrogen outlet stream could be directly employed by the fuel cell. Nevertheless, if steam were used as the carrier gas for hydrogen removal from the membrane reactor, the wet H2 stream could be suitable and would not need further processing before the fuel cell. Methanol has some advantages as a fuel and raw material for several chemical production processes. For example, it is more easily transported than methane or other gas fuels, it has a high energy density, and it does not require desulfurization. Furthermore, methanol fuel reacts at moderate temperature2 (200-400 °C), in such * To whom correspondence should be addressed. Tel.: (+39) 0984 492013. Fax: (+39) 0984 402103. E-mail: a.basile@itm. cnr.it.

reactions as partial oxidation (200-220 °C), steam reforming (200-300 °C), and methanol decomposition (up to 400 °C). In particular, the steam reforming of methanol is an endothermic reaction and is considered an important method for hydrogen generation in terms of feasibility for various types of on-site energy systems. Although CO2 is generated as a byproduct in both the methanol steam-reforming (MSR) reaction and partial oxidation of methanol, the amount of CO2 produced in these processes is about 50% lower than the amount of CO2 released from internal combustion engines using gasoline as a fuel. In addition, PEMFC-powered vehicles using H2 fuel do not emit any noxious gas such as NOx, SOx, or hydrocarbons,3 whereas such emissions accompany the use of traditional reformers in the conversion of methanol (or other fuels) into hydrogen. The purpose of this study is to investigate the possibility of increasing the methanol conversion in a tubular membrane reactor with respect to that obtained using a traditional catalytic system. Previous work in the literature has considered this type of reaction. A paper by Christiansen4 makes the earliest reference to the reaction between steam and methanol over a reduced Cu-based catalyst. At the end of 1970s, Santacesaria and Carra`5 published a study in which a mathematical expression for the reaction rate of methanol was derived through an empirical method; this kinetic expression is in Langmuir-Hinshelwood form, even if though was not derived from an explicit reaction mechanism. At the beginning of 1980s, Amphlett et al.6-9 performed research regarding thermodynamic and kinetic aspects of the methanol steam-reforming reaction over Cu/ZnO/Al2O3 catalysts. Other authors10-15 started kinetic studies of methanol steam reforming, to develop a reasonable kinetic model; among them, Du¨mplemann10 developed a reaction mechanism including the direct formation of CO2 by means of a reaction between methanol and steam. Scientific works on traditional reactors are mainly focused on optimization of the catalyst to reduce the CO content in the gaseous mixture exiting from the reactor. This aspect is very important in avoiding poisoning of the anodic catalyst in view of utilization of the hydrogen-containing gas in the outlet stream as the feed of a PEMFC. However, it should be said that, although the best catalyst is used in such

10.1021/ie0304863 CCC: $27.50 © 2004 American Chemical Society Published on Web 04/08/2004

Ind. Eng. Chem. Res., Vol. 43, No. 10, 2004 2421 Table 1. Methanol Steam-Reforming Reaction: Main Operating Conditions Found in the Literature authors Sekizawa et al.2 Amphlett et al.7 Jiang et al.11,12 Peppley et al.14,15 Wieland et al.18 Asprey et al.24 Segal et al.27 Idem and Bakhshi28 Velu et al.29 Tsai and Yoshimura30 Shen et al.31 Iwasa et al.32 Lindstro¨m and Pettersson33 Agrell et al.34 Liu et al.35 Breen and Ross36 Wild and Verhaak37 Takeda et al.38 Emonts et al.16 Buxbaum17 Wieland et al.18 Lin et al.20 Zaspalis et al.39 a

catalyst type

pressure temperature H2O/ (atm) (°C) CH3OH 1 1 1 1-16 25 3.33 1 1 1 1 1 1 1 1 1 1 1.3-1.5 1 21 10.2 25 6-15 0.04

150-280 150-240 170-190 160-260 280-310 202-252 200-400 170-250 180-290 300 140-290 160-450 180-250 160-310 260-310 140-400 180-270 160-280 260-280 265 280-310 350 20-500

3-5 1 1 1 0.75 1 1 1.3 1.48 1 1 1 1.3 1 1.3 1.5-2 1.5 1.2 -

Cu/ZnO/Al2O3 Cu-ZnO-Al2O3 (C18HC) Cu-ZnO-Al2O3 Cu/ZnO/Al2O3 Cu/ZnO/Al2O3 Cu/ZnO/Al2O3 Cu/Al (LDH) Cu, Al, Mn based Cu/Zn/Al(Zr)-oxide Al63Cu25Fe12 Cu/ZnO Ni/ZnO, Co/ZnO, Pt/ZnO Cu based on γ-Al2O3 Cu/ZnO/Al2O3 (G-6MR) Cu/CeO2 Cu based Cu based Cu/ZnO/Al2O3 + Ru/Al2O3 Cu-Zn/Al2O3 CuO/ZnO Cu/ZnO/Al2O3 Cu/ZnO/Al2O3 -

weight (g)

type

1-6.5 0.05-0.5 0.077 0.5-2 0.15 0.1 40 15 0.05 0.3 0.1 74 -

TR TR TR TR TR TR TR TR TR TR TR TR TR TR TR TR TR TR Pd Pd Pd75Ag25 Pd-supported ceramic, γ-Al2O3-based

a

-

reactor membrane thickness (µm) 25 20 4.5

Size ) 100 cm3.

Table 2. List of the Parameters Studied in the Theoretical Analysis and The Ranges Investigated parameter

range

reaction temperature lumen pressure time factor H2O/CH3OH feed ratio sweep gas flow rate

180-300 °C 1-10 bar 0-9 kgcat‚s/molCH3OH 1-10 (2.2 × 10-5)-2.2 mol/s

systems, other methods for reducing the CO content of the product stream must also be investigated, because catalyst optimization is not sufficient. In this scenario, the potential benefit of a hydrogen-selective membrane reactor for the recovery of pure hydrogen from the methanol steam-reforming reaction appears. To our knowledge, few scientific papers have considered the use of membrane reactors in the methanol steam-reforming reaction. Pd,16,17 Pd/V/Pd, Pd75Ag25, Pd60Cu40,18 and Pdsupported19,20 membranes have been studied in the pressure range of 1-25 atm and at temperatures between 260 and 320 °C. Table 1 summarizes the main operating conditions investigated in the studies of membrane reactors applied to methanol steam reforming found in the literature. In this paper, a palladium-silver membrane reactor (MR) is considered with particular regard to increasing the hydrogen recovery over that obtained with a traditional packed-bed reactor (TR), from a theoretical viewpoint. The few published papers16-20 dealing with MR applications in the MSR reaction mainly analyze the effects of temperature and pressure on the reaction system. The objective of this investigation is to extend the MR analysis to the effects of other parameters, such as the time factor, H2O/CH3OH feed ratio, and sweep gas flow rate, in addition to the reaction temperature and pressure. The main aim of this paper is to offer a set of operating parameters, through an engineering study, to be used in future experimental work in order maximize both methanol conversion and hydrogen production. In Table 2, a list of the operating parameters studied and their ranges is reported.

2. Comparison between TRs and MRs and Considerations A theoretical comparison between the two devices (MRs and TRs) shows quantitatively that MRs provide very high hydrogen recoveries compared to TRs at the same operating conditions, as a result of two simultaneous effects: the increase in methanol conversion (which produces an increase in the amount of hydrogen produced) and the removal of pure hydrogen. No gascleaning unit is needed for MRs. In contrast, TRs produce lower amounts of hydrogen gas (because the methanol conversion is lower than that of an MR operating at the same experimental conditions) and require hydrogen purification devices to achieve an output stream that can be exploited directly in a PEMFC. MRs are able to combine the reaction and purification processes in only one device. The reaction product (generally hydrogen) can be removed directly from the reaction system by a selective membrane. In some cases, the removal of a product can promote the reaction in the direction toward the product side according to the thermodynamic equilibrium. Therefore, the membrane reactor can carry out the reaction even at lower reaction temperature and achieve high conversion. Some equilibrium-limited reactions, such as dehydrogenation and steam reforming, have been investigated using membrane reactor to increase the conversion by shifting the equilibrium. Supported palladium membranes have usually been used as the hydrogen-permeable membrane.21 Because of their high permselectivity, highpurity hydrogen can be obtained directly from the reaction system. Therefore, no extra purification facility is needed.22 For the present investigation, a mathematical model is first used to simulate a traditional chemical reactor (TR). Figure 1 shows a scheme of such a TR, consisting of a stainless steel tube packed with a catalyst. One inlet stream (feed) and one outlet stream (products) are present, and the steam-reforming reaction takes place in the lumen of the tube. The following definitions are used to describe the performance of a TR

2422 Ind. Eng. Chem. Res., Vol. 43, No. 10, 2004

Figure 1. Scheme of the traditional reactor (TR).

Figure 2. Scheme of the membrane reactor (MR).

CH3OH conversion (%) ) NCH3OH,feed - NCH3OH,products 100 (A) NCH3OH,feed H2 selectivity (%) ) NH2,products NH2,products + NCO,products + NCO2,products

100 (B)

where Ni indicates the molar flow rate of species i in the feed stream or in the product stream. A comparison of the model results with literature experimental data23 confirmed the validity of the model. Afterward, the simulation was used to predict the behavior of an MR. Figure 2 shows a scheme of the MR, consisting of a stainless steel tube housing a Pd/Ag tubular membrane and packed with a catalyst. The Pd/Ag membrane used to achieve the permeation expression (see Scheme 1 for details) was produced by a lamination technique at ENEA Laboratories (Frascati, Italy). In particular, a palladium alloy foil was cold-rolled by a two-high laboratory mill to reduce its thickness and obtain the required characteristics of mechanical strength and surface finishing. The rolling process was operated starting from a commercial 127-µm-thick palladiumsilver foil. This metal sheet was rolled to a thickness of 50 µm in five steps. The values of the apparent activation energy and preexponential factor are 29.16 kJ/mol and 1.12 × 10-5 mol‚m/(s‚m2‚Pa0.5), respectively. Table 3 shows that these values are compatible with experimental results found in the literature. It should be noted that the Pd/Ag membrane exhibits infinite selectivity for H2 compared to other gases. Two inlet streams (feed and sweep gas inlet) and two outlet streams (permeate outlet and retentate stream) are present, and the steam-reforming reaction takes

place in the lumen of the membrane tube. A graphite gasket ensures that the lumen- and shell-side gases do not come into contact during the reaction. The sweep gas keeps the hydrogen partial pressure in the shell side as low as possible. This produces a high driving force for hydrogen permeation. Therefore, the use of the sweep gas leads to an increase in the methanol conversion compared to the case in which the sweep gas is not used. From the industrial point of view, the use of a sweep gas could correspond to keeping the shell side under vacuum. Otherwise, it is possible to use a sweep gas that is easy to remove after hydrogen extraction. In fact, it is possible to use steam as the sweep gas, thereby achieving a mixture of hydrogen and steam from which the steam can be removed by condensation, so that a pure hydrogen stream is obtained. The following definitions were used for comparing the performances of the MR and TR

CH3OH conversion (%) ) NCH3OH,feed - NCH3OH,retentate 100 (C) NCH3OH,feed H2 selectivity (%) ) NH2,permeate + NH2,retentate NH2,permeate + NH2,retentate + NCO,retentate + NCO2,retentate

100 (D)

H2 recovery (%) )

NH2,permeate NH2,permeate + NH2,retentate

100 (E)

The H2 recovery is defined only for the MR. It can be considered as the hydrogen recovered throughout the shell side compared to the total amount of hydrogen produced.

Scheme 1. Details Concerning the Model Developed for Simulating the Steam-Reforming Reaction

Ind. Eng. Chem. Res., Vol. 43, No. 10, 2004 2423

2424 Ind. Eng. Chem. Res., Vol. 43, No. 10, 2004 Table 3. Apparent Activation Energy and Preexponential Factor Compared with the Literature Ea (kJ/mol)

Pe0 [10-5 mol‚m/(s‚m2‚Pa0.5)]

ref

29.16 29.73 15.70 15.50 12.48 18.45 48.50

1.12 7.71 2.19 2.54 0.38 1.02 9.33

this work Basile et al.25 Koffler et al.40 Balovnev41 Itoh and Xu42 Itoh et al.43 Tosti et al.44

This particular application of an MR in the MSR reaction is quite interesting interesting because ultrapure hydrogen can be produced at high recovery, which is particularly useful for fuel-cell applications. In fact, it is well-known that the anodic catalyst of PEM fuel cells operating at 1 atm and 80-100 °C is poisoned by CO concentrations higher than 5-10 ppm in the hydrogen stream.23,24 This theoretical analysis provides a range of parameters that permit one to obtain the maximum (pure) hydrogen production and/or methanol conversion if employed in an experimental device. 3. Theoretical Analysis of the MSR Reaction The theoretical modeling of the reaction system was developed in two different steps. In the first step, a thermodynamic analysis of the MSR reaction was carried out, on the basis of the three reactions reported below. The thermodynamic analysis applied to the MR is termed “dynamic”. The second step of this analysis was developed by using the kinetic expressions of the three reactions, indeed considering the reaction system to be controlled by kinetics. Details are explained in the following sections. 3.1. Thermodynamic Analysis and Simulation of the MSR. The first step of the present study was to investigate the thermodynamic aspects of the reaction system both without the membrane (TR) and with the membrane (MR). In other words, the maximum achievable methanol conversion and hydrogen yield were analyzed according to thermodynamics for both reactors. The thermodynamic analysis is very useful because it permits the determination of how far each reactor is from its highest limit, thus making it possible to check for improvements in the practical case of an experimental device. According to the literature, the chemical reactions involved in this process are the following

CH3OH + H2O T CO2 + 3H2 ∆H298K ) +90.70 kJ/mol (1) CO + H2O T CO2 + H2

∆H298K ) -41.19 kJ/mol (2)

CH3OH T CO + 2 H2

∆H298K ) +49.51 kJ/mol (3)

Equations 1 and 3 are both reversible and endothermic and proceed with an increase in volume, so the highest methanol conversions are obtained at high temperature and low pressures. The exothermic reaction 2 is the socalled water-gas shift reaction, which takes place simultaneously with methanol steam reforming; it proceeds without any volume change. In the thermodynamic analysis, the reaction system is considered not to be limited by chemical kinetics, and

Figure 3. CH3OH conversion vs temperature: dynamic (MR) and thermodynamic (TR) equilibrium. H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, sweep gas flow rate ) 2.2 × 10-3 mol/s (for the MR).

thus, the feed gas is assumed to be continuously at equilibrium throughout the MR as well as the TR: this means that the reaction rate is not the limiting step. All molecules in contact with the catalyst react with an infinite velocity. For the TR, the concentrations of the outlet species are a function of thermodynamic equilibrium resulting from a mass balance at the equilibrium conditions. For the MR, the fluid dynamics of the reaction system must be considered because hydrogen permeation changes the equilibrium compositions: this condition is termed “dynamic equilibrium”. The same theoretical analysis has already been developed for the partial oxidation of methane carried out in a palladium membrane reactor.25 This model was developed using all of the thermodynamic values for the reaction species; hydrogen removal was calculated using experimental hydrogen permeability data (Arrhenius plot) reported in Scheme 1. The results of the thermodynamic analysis applied to both the MR and TR operating under equilibrium conditions are reported in Figure 3. In particular, the curve representing the CH3OH conversion achieved in the MR operating at equilibrium conditions represents the dynamic equilibrium at a fixed sweep gas flow rate (2.2 × 10-3 mol/s) and at H2O/CH3OH ) 3, plumen)2 bar, and pshell ) 1 bar. The other curve represents CH3OH conversion at the equilibrium for the TR. In the range of temperatures investigated here, the CH3OH conversion increases with increasing temperature for the TR, although it maintains high values (>99%); in particular, at 190 °C, the CH3OH conversion is 99.5%, whereas at 300 °C, it reaches 100%. In comparison, the CH3OH conversion in the MR is 100% throughout the range of temperatures investigated. The best performance of the MR, due to the effect of hydrogen removal, can be evidenced by considering that the equilibrium CH3OH conversion is achieved in both the TR and MR for different lengths at a fixed feed flow rate. In particular, at the operating conditions of Figure 3, the equilibrium CH3OH conversion can be achieved using a TR that is 25 m long, whereas the equilibrium CH3OH conversion can be achieved using an MR that is 4.5 m long. The conclusion was obtained by performing a calculation using a model that takes into account the kinetic expressions and extending the reactor length until thermodynamic equilibrium is reached. An investigation at lower temperature could demonstrate the impact of the MR more clearly. However, two aspects have to be considered: (1) Hydrogen permeation de-


creases with decreasing temperature, according to the Arrhenius-like dependence of the hydrogen permeability on the temperature [Pe ) Pe0 exp(-Ea/RT)]. (2) The lower temperature limit of the catalyst (type BASF K3110 low-temperature shift catalyst) used for carrying out the methanol steam-reforming reaction is about 180 °C. The calculated equilibrium methanol conversions at 150 °C are 94.79 and 97.68% for the TR and MR, respectively; at 100 °C, the corresponding values are 62.30 and 66.89%. Although the thermodynamic analysis (equilibrium) at T < 180 °C indicates the effect of the MR more clearly, in this temperature range (100180 °C), the catalyst is not active, and the reaction cannot take place. 3.2. Kinetic Analysis and Simulation of the MSR. The model was applied to a realistic MR as well as a TR using the kinetic expressions. In this case, the phrase “realistic reactor” is intended to mean a tubular reaction device having a fixed length (and consequently a fixed catalyst load) and in which kinetics controls the reaction system. The rate equations are taken from Peppley et al.,15 as are the kinetic and adsorption parameters (see Scheme 1 for more detail), for a particle size of catalyst in the range of 250-500 µm.

{

( )[ )]} {[ ( ) ( ) ] pCH3OH

T / r1 ) CTS1 CS1a (krKCH ) 3O1

(

pH23pCO2

krpH2OpCH3OH

1-

pH21/2

/ / 1 + KCH + 3O1

/ KHCOO1 pH21/2pCO2 + K/OH1

pH2O

pH21/2

pCH3OH pH21/2

+

+ KCO2(1)pCO2 (1 +

}

KH1a1/2pH21/2)

{

( )[ ( )]} {[ ( ) ( ) ]} { ( )[ ( )]} {[ ( ) } ( ) ]

r2 ) (CTS1)2 (k/w K/OH1) / 1 + KCH + 3O1

pCOpH2O pH21/2

pCH3OH pH2

1/2

r3 )

T CS2a

/ (kdKCH ) 3O2

/ 1 + KCH + 3O2

K/OH2

pH2O

pH21/2

pH2

pH21/2

pH21/2

1/2

kwpH2OpCO

pH2O

pCH3OH

pCH3OH

pH2pCO2

/

/ + KHCOO1 pH21/2pCO2 +

K/OH1

CTS2

1-

1-

2

+ KCO2(1)pCO2 pH22pCO

kdpCH3OH

/

/ + KHCOO2 pH21/2pCO2 +

+ KCO2(2)pCO2 (1 + KH2a1/2pH21/2)

Coke formation was not taken into account in the model, because it is not included among the available kinetic expressions of the commercial catalyst used. Mass balances for a differential reactor volume along the z axis (see Figures 1 and 2) of both the TR and MR were written for the reaction zone (lumen) and for the

Figure 4. CH3OH conversion vs time factor, model and Peppley et al.15 experimental data. H2O/CH3OH ) 1, plumen ) 1.01 bar, T ) 513 and 533 K.

permeation zone (the latter mass balance is valid only for the MR). The set of differential equations was solved using a fourth-order Runge-Kutta method with a variable step size. Scheme 1 shows all of the details concerning the model developed for simulating the steam-reforming reaction system. Moreover, the following common hypotheses were considered: plug flow, kinetic control of the reaction system, isothermicity, and constant packing of the catalyst bed (i.e., constant void fraction). The validity of the model was assessed by comparing the results achieved in this work with literature experimental data15 for the TR (Figure 4). In fact, considering the experimental results of Peppley et al.15 at 240 and 260 °C, a comparison with our model results shows good agreement. For example, at a time factor (TF) of 0.4 kgcat‚s/molCH3OH and a temperature of 260 °C, the experimental methanol conversion15 is about 4.7% versus a model prediction of about 5%. The TF results were calculated by considering the catalyst weight divided by the methanol feed rate. The maximum difference between the experimental results of Peppley et al. and our model is 2.4% at 240 °C and TF ) 4.25 kgcat‚s/molCH3OH. It should be observed that the model for both the TR and MR was developed by using the kinetic expressions of Peppley et al.15 In principle, the model is not said to fit the experimental data quite well, because the model outputs depend on the model assumptions. Unfortunately, to our knowledge, no further experimental data are available in the literature for this type of catalyst used for the methanol steam-reforming reaction. Nevertheless, the comparison reported in Figure 4 seems to be sufficient to demonstrate the model validity under the hypotheses mentioned in the paper (i.e., plug flow and kinetic control of the reaction system). Moreover, we calculated the Reynolds number (Re) for the TR at the lowest and highest operating temperatures investigated, keeping the other parameters constant (H2O/CH3OH feed ratio ) 3, plumen ) 2 bar, time factor ) 8.17 kgcat‚s/molCH3OH). The values obtained were Re ) 50 600 at 180 °C and Re ) 39 200 at 300 °C, with a quasilinear decreasing trend of Re versus temperature. These calculations also support the plug-flow hypothesis. After validation, the model was applied to the MR. The main assumption in moving from the TR to the MR is that the kinetic expressions remain the same. Al-


Figure 5. CH3OH conversion vs time factor, MR and TR model. H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, T ) 200-300 °C, sweep gas flow rate ) 2.2 × 10-3 mol/s (for the MR).

Figure 6. CH3OH conversion vs feed flow ratio (H2O/CH3OH), MR and TR model. plumen ) 2 bar, pshell ) 1 bar, T ) 260 and 300 °C, sweep gas flow rate ) 2.2 × 10-3 mol/s (for the MR), time factor ) 8.17 kg‚s/molCH3OH.

though it is possible that, depending on the fluiddynamic conditions, the kinetic expressions could change from the TR to the MR, to our knowledge, no papers have been published in the literature demonstrating that the kinetic expressions are completely different. In particular, this is true for the reaction under study in this work. This is the key assumption in our model. In general, the MR should not influence the kinetics: in fact, for example, for hydrogen production via the methane steam-reforming reaction, a recent paper26 demonstrated that a good fit was obtained by using the kinetic expressions related to a traditional reaction system in the MR model. A key comparison could be done between the model and experimental results for the MR. However, no experimental results regarding an MR operating with the same type of catalyst (BASF K3-110 low-temperature shift catalyst) are available in the literature. Moreover, this catalyst is not yet available in our laboratory, so it is not possible to perform suitable experimental tests using an MR, although we chose this catalyst because it is one of the few for which kinetic expressions are available in the literature. Finally, a comparison with an MR using another type of catalyst is not significant for the model validation: any eventual discrepancies could be attributed both to the different type of catalyst and to the use of an MR; in this case, it is not possible to distinguish between the two contributions. Using our model and following our hypotheses, we want to observe the influence of the membrane on the reactor behavior; it is likely that this contribution does not change as the type of catalyst is changed. Hereinafter, all of the simulation results are referred to the kinetic analysis.

fixed temperature, the methanol conversion is higher in the MR than in the TR, and increasing TF increases the difference between the MR and TR. For example, at TF ) 8.17 kgcat‚s/molCH3OH and T ) 260 °C, the CH3OH conversion is 33% for the TR, but it is double that (68%) in the MR. The better performance of the MR compared to the TR in terms of CH3OH conversion is due first to the presence of the membrane, which removes the hydrogen produced during the reaction. The improvement in MR performance with increasing TF is due to the increase in the membrane surface area of MR. In other words, when TF increases, the catalyst weight increases as well (for a constant feed flow rate as well as a constant packing of the catalyst); the reactor length increases, as does the tubular palladium membrane length. In fact, considering the constant packing of the catalyst bed (i.e., the void fraction does not change with reactor length), an increase of TF produces an increase in the catalyst weight and consequent increases in the membrane reactor length, membrane length, and membrane surface area. This results in increasing hydrogen removal with increasing TF, so the curve for the MR exhibits a higher slope than that for the TR. It should be observed that, at each value of TF, the CH3OH conversion for the MR operating at 260 °C is close to the CH3OH conversion for the TR operating at 300 °C. This demonstrates that the use of the membrane permits one to achieve the same CH3OH conversion as in the TR but at lower temperature, with consequent energy savings. The effect of temperature on the TR is clearly demonstrated in Figure 5. At low temperature, the methanol conversion is extremely low. In any case, the methanol conversion in the MR exceeds that in the TR at each temperature investigated. The choice to stop the modeling at TF ) 8.17 kgcat‚s/molCH3OH is due to the very long time required for the computer calculations. Moreover, a value of TF ) 8.17 kgcat‚s/molCH3OH is easy to adopt in an experimental laboratory-scale device, to be investigated in the future. Figure 6 shows the CH3OH conversion versus the H2O/CH3OH feed ratio at plumen ) 2 bar, pshell ) 1 bar, T ) 260 and 300 °C, and NN2,sweep ) 2.2 × 10-3 mol/s. Considering reaction 1 (see section 3), an increase in the H2O/CH3OH feed ratio produces an increase in the methanol conversion; with regard to the other two reactions, the effect of the H2O/CH3OH feed ratio is not

4. Results and Discussion Figure 5 reports the CH3OH conversion versus TF at different temperatures in the range of 200-300 °C for H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, NN2,sweep ) 2.2 × 10-3 mol/s. As expected, increasing TF increases the CH3OH conversion for both the TR and the MR. In fact, considering a constant feed flow rate, increasing TF means that the catalyst weight increases, so the residence time of the gaseous reactant over the catalyst surface increases as well, giving a higher reactant conversion. For the whole range of TFs considered, at a


Figure 7. CH3OH conversion vs temperature, MR and TR model. H2O/CH3OH ) 3, plumen ) 2 and 10 bar, pshell ) 1 bar, sweep gas flow rate ) 2.2 × 10-3 mol/s (for the MR), time factor ) 8.17 kg‚ s/molCH3OH.

Figure 8. Hydrogen selectivity and hydrogen recovery in the MR vs temperature, MR and TR model. H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, sweep gas flow rate ) 2.2 × 10-3 mol/s (for the MR), time factor ) 8.17 kg‚s/molCH3OH.

immediate, because of the presence of common chemical species in the reactions. Concerning the TR, increasing the feed ratio from 1 to 6 always increases the CH3OH conversion. For example, at H2O/CH3OH ) 1 and T ) 260 °C, the CH3OH conversion is 34.5%, whereas at H2O/CH3OH ) 6 and the same temperature, it reaches 40.3%. In contrast, for the MR, the CH3OH conversion shows a minimum at H2O/CH3OH ) 1.25. In fact, the CH3OH conversion is 73.1% at H2O/CH3OH ) 1, 72.8% at H2O/CH3OH ) 1.25, and 76.4% at H2O/CH3OH ) 6. In each case, the CH3OH conversion in the MR is again higher than that in the TR, although the difference in the methanol conversions at H2O/CH3OH ) 6 is lower than that at H2O/CH3OH ) 1. At 300 °C, the same trend is observed, but the difference between the MR and TR is lower than it is at 260 °C. With this evidence, the H2O/CH3OH feed ratio parameter does not produce a relevant variation in the reactors’ behavior in terms of CH3OH conversion in the range investigated. Figure 7 shows a plot of the CH3OH conversion versus temperature for both the TR and MR, at H2O/CH3OH ) 3, plumen ) 2 and 10 bar, pshell ) 1 bar, and NN2,sweep ) 2.2 × 10-3 mol/s. Increasing the temperature increases the CH3OH conversion for both reactors, according to the endothermicity of the overall reaction system (see section 3). For the whole range of temperatures considered, the CH3OH conversion is higher in the MR than in the TR. For example, at T ) 247 °C and plumen ) 2 bar, the CH3OH conversion is 27.6% for the TR and 59.9% for the MR. In the middle region of the figure, the difference between the two curves increases with increasing temperature; this phenomenon is more evident at high lumen pressures. This is due to the temperature dependence of hydrogen permeation through the membrane. In fact, the hydrogen permeability increases with increasing temperature according to the Arrhenius expression (see Scheme 1), so a benefit in terms of methanol conversion is obtained. At temperatures higher than 300 °C, 100% CH3OH conversion is achieved in the MR at both lumen pressures considered, compared to 80.5 and 70.5% for the TR operated at plumen ) 2 and 10 bar, respectively. At the same operating conditions as reported in Figure 7, Figure 8 reports the H2 selectivity versus temperature for both the TR and MR. The H2 selectivity increases with increasing temperature. The investigation starts from 180 °C, that is, lower than the starting

temperature of Figure 7. Throughout the temperature range considered, the H2 selectivity for the MR is higher than that for the TR, except for T ) 180 °C: at this temperature, the H2 selectivity is the same for both the TR and MR, and its value is 60.7%. This is probably due to the negligible hydrogen permeation through the membrane at this temperature. In other words, the exponential temperature dependence of the hydrogen permeability does not produce significant hydrogen removal from the reaction side, so the MR practically works similarly to the TR. The difference between the MR and TR increases with the temperature, because the H2 permeation exhibits an Arrhenius-like temperature dependence through the palladium membrane. This result is due to kinetic control of the reaction system, which can be verified also by considering the Figure 7, where the methanol conversion increases with increasing temperature, even though the thermodynamic predictions are 100% methanol conversion at each temperature investigated. It should be noted that the H2 selectivity was calculated by considering all of the H2 produced with respect to the byproducts (eq D, section 2): this means that, for the TR, only one outlet stream was considered, whereas for the MR, both the outlet streams (permeate + retentate) were considered for the hydrogen calculation (byproducts are present in the retentate stream only). Considering the three reactions reported in section 3, an increase in the CH3OH conversion means that reactions 1 and 3 are shifted toward products, so the total amount of byproducts (CO and CO2) increases, although the hydrogen production increases as well. However, using the MR, it is possible to recover part of the total amount of the hydrogen produced in the shellside stream. To show this, in the same figure, the percentage of H2 recovered in the shell side of the MR is also reported: with increasing temperature, the amount of H2 recovered increases. In particular, at T ) 255 °C, the percentage of H2 recovered is 95%; this means that 95% of the hydrogen produced is recovered in the shell side. Considering Figure 8 with particular regard to the H2 recovered in the shell side, defined as the amount of H2 recovered compared to the amount of H2 produced in the lumen side, three cases can, in principle, be distinguished: (1) inert gas is used as the


Figure 10. CH3OH conversion vs lumen-side pressure, MR and TR model. H2O/CH3OH ) 3, pshell ) 1 bar, T ) 260 and 300 °C, sweep gas flow rate ) 2.2 × 10-3 mol/s (for MR) time factor ) 8.17 kg‚s/molCH3OH.

Figure 9. Effects of the reaction pressure and H2O/CH3OH feed ratio on the H2 recovery in the MR. T ) 300 °C, sweep gas flow rate ) 2.2 × 10-3 mol/s, time factor ) 8.17 kg‚s/molCH3OH.

sweep gas, (2) steam is used as the sweep gas, or (3) vacuum is realized to promote H2 removal. In case 1, the H2 needs to be separated from the inert gas. Adopting the operating conditions of Figure 8, the maximum H2 recovery is about 95% at 255 °C; the corresponding flow rate of the H2-containing stream is 0.0039 mol/s, and the H2 percentage is 43.5%. In case 2, the same considerations as in case 1 can be repeated, but now, the H2 recovery is very simple, because steam can be removed via condensation. This condition is interesting from an industrial viewpoint. In case 3, the vacuum could be tuned to realize the same driving force as obtained in the previous cases. In this case, the outlet stream of the shell side is 100% H2, with the same flow rate of 0.0039 mol/s. This condition should also be interesting from an industrial viewpoint. In each case, the maximum H2 recovery corresponds to the final value. This value could be changed by adjusting some of the other operating parameters, for example, the reaction pressure and the H2O/CH3OH feed ratio. Figure 9 shows the effects of these two parameters on the H2 recovery in the MR. The H2 recovery decreases with increasing reaction (lumen) pressure at each value of the H2O/CH3OH feed ratio investigated. For example, at H2O/CH3OH ) 10, the H2 recovery is 76.8% at plumen ) 10 bar and 20.82% at plumen ) 2 bar. Moreover, the H2 recovery also decreases with increasing H2O/CH3OH feed ratio for a fixed value of plumen. For example, at plumen ) 4 bar, the H2 recovery is 79.24% at H2O/ CH3OH ) 3 and 44.85% at H2O/CH3OH ) 10. The maximum hydrogen recovery is 95.82% at plumen ) 10 bar and H2O/CH3OH ) 1. Considering Figures 7 and 8, it is clear that, at high temperatures (>227 °C), higher methanol conversion, higher hydrogen selectivity, and much higher hydrogen recovery can be achieved using the MR compared to the TR. Finally, it should be observed that the H2 selectivity predicted by our model for the TR is lower than the experimental H2 selectivity achieved in TRs studied by other authors.26 In fact, in previous work, the authors generally found a H2 selectivity of about 70% for their

experimental conditions, whereas our model predicts an average H2 selectivity of about 62%. The main reason for this discrepancy could be either the difference in the catalyst type (on which the kinetic expressions depend) or the difference in the catalyst amount (and the corresponding difference in reactor length). Figure 10 reports the CH3OH conversion versus the reaction-side (lumen) pressure, for a constant shell-side pressure (1 bar, MR), at T ) 260 and 300 °C, H2O/CH3OH ) 3, and NN2,sweep ) 2.2 × 10-3 mol/s. Conceptually, for a TR, an increase in the operating pressure results in a decrease in the equilibrium methanol conversion, as reactions 1 and 3 proceed with a volume increase (see section 3) whereas reaction 2 does not involve a volume change. With regard to the TR, the CH3OH conversion shows a maximum for plumen ) 4.6 bar, where it reaches 38.5% at 260 °C. In contrast, at plumen ) 10 bar, the CH3OH conversion is 37.4%, so a decreasing trend is observed in the range of lumen pressure between 4.6 and 10 bar at 260 °C. Probably, at lower values of lumen pressure, the reaction performance is affected by the adsorption velocity of the reactant gases on the catalyst surface, and this phenomenon is favored by increasing pressure. This could justify the increasing trend up to 4.6 bar. At 300 °C, the maximum CH3OH conversion for the TR shifts toward plumen ) 1.9 bar, and the increasing lumen pressure decreases the CH3OH conversion more rapidly than in the case of 260 °C. When the reaction system is not at equilibrium, the effect of pressure is not immediately predictable from a thermodynamic viewpoint. In fact, for example, for a TR, Peppley et al.15 found a positive effect of pressure on CH3OH conversion at low values of the time factor both experimentally and theoretically. In fact, at time factor lower than 10 kgcat‚ s/molCH3OH the increasing pressure led to an increase in CH3OH conversion, whereas the opposite effect was observed for time factor values higher than 10 kgcat‚s/ molCH3OH. In our case (Figure 10), however, increasing the reaction-side pressure in the MR causes the CH3OH conversion to increase. In this case, it is possible to mention two opposite effects: (a) The increase in the lumen pressure negatively affects the CH3OH conversion, as already mentioned for the TR. (b) An increase in the CH3OH conversion occurs because of the increase in the driving force for hydrogen permeation. Through-


Figure 11. Lumen pressure corresponding to the maximum CH3OH conversion of Figure 9, as a function of temperature and time factor. TR, H2O/CH3OH ) 3, plumen ) 2 bar.

Figure 12. CH3OH conversion vs temperature for the TR and MR operating at different sweep gas flow rates. The H2 partial pressure in the shell side versus temperature is also reported. H2O/ CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, time factor ) 8.17 kg‚ s/molCH3OH.

out the range of lumen pressures investigated, the second effect prevails. In fact, at 260 °C, for example, the CH3OH conversion increases from 67.6% at plumen ) 1 bar to 86.4% at plumen ) 10 bar. Considering Figure 10, it should be noted that the curve of the TR operating at 300 °C crosses the curve of the MR operating at 260 °C at plumen ) 3.5 bar. This shows that, in the TR operated at 300 °C and plumen values higher than 3.5 bar, the CH3OH conversion in the MR operated at 260 °C is higher than that in the TR. The difference between the two curves increases with increasing the lumen pressure, so that, in the MR operating at 260 °C and plumen ) 10 bar, the CH3OH conversion is 86.4%, whereas in the TR operating at 300 °C, the CH3OH conversion is only 70.5%. For these conditions, the MR provides a higher CH3OH conversion than the TR but at a lower temperature. In the regime of kinetic effects, the conversion can increase with increasing pressure. Therefore, an analysis with a lower TF seems to be necessary to observe the kinetic effects caused by the reaction pressure without the influence of thermodynamic equilibrium. For this purpose, considering the TR, Figure 11 shows a 3D plot indicating the dependence of the lumen pressure (corresponding to the maximum of the methanol conversion) on the temperature and the TF. Figure 11 demonstrates that, at lower TF values, the lumen pressure corresponding to a maximum in methanol conversion is higher, but at a higher temperature. For example, at 280 °C and TF ) 1 kgcat‚s/molCH3OH, the value of plumen corresponding to the maximum methanol conversion (15.4%) is 5.05 bar; at the same temperature and TF ) 8.17 kgcat‚s/molCH3OH, the value of plumen corresponding to the maximum methanol conversion (61%) is 3.25 bar; and at 200 °C and TF ) 8.17 kgcat‚s/ molCH3OH, the value of plumen corresponding to the maximum methanol conversion (33.5%) is about 1 bar. Figure 12 shows the CH3OH conversion versus temperature for the TR and MR operating at different sweep gas flow rates (2.2 × 10-5, 2.2 × 10-3, and 2.2 mol/s), at H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, and TF ) 8.17 kgcat‚s/molCH3OH. The CH3OH conversion of the MR is always higher than that of the TR operating

under the same conditions. Moreover, the CH3OH conversion in the MR increases as the sweep gas flow rate increases, because of the higher hydrogen removal from the reaction system, which produces a higher equilibrium shift toward the product. A CH3OH conversion of 100% is achieved in the MR operating at about 280 °C and 2.2 mol/s of sweep gas, compared to the about 60% CH3OH conversion obtained using the TR at the same temperature and sweep gas flow rate. Significant improvements in terms of the CH3OH conversion are achieved when the sweep gas flow rate is increased, depending on the entire set of operating parameters of the reaction. In principle, a limit in the sweep gas flow rate exists, such that increasing this parameter over the limiting value no longer increases the methanol conversion. This can be explained considering that, because of the high values of sweep gas flow rates used, the hydrogen removal rate (through the shell side of the MR) might surpass the hydrogen production rate (in the lumen side of the MR). Thus, the hydrogen production would become the limiting parameter. The operating sweep flow rate should not be greater than this limit; otherwise, the increase in the sweep gas flow rate is not useful and can be considered a wasted effort. In our case, changing the sweep gas flow rate from 2.2 × 10-5 to 2.2 mol/s resulted in improvements in the CH3OH conversion; the most significant improvement was observed at 260 °C by changing this parameter from 2.2 × 10-3 mol/s (about 60% CH3OH conversion) to 2.2 mol/s (more than 90% CH3OH conversion). It should be emphasized that the development of the modeling work in this paper mainly focused on methanol conversion and hydrogen selectivity: which implies an analysis of only the outlet streams of the TR and MR. To briefly investigate the interiors of the reactors, we extended the theoretical analysis to include what happens along both the TR and MR. Results are presented in Figures 13 and 14, showing the flow rates of each reaction species (CH3OH, H2O, CO, CO2, H2) along the two reactors. Both figures are represented with two x axes, the reactor length (z), and the time factor: these two parameters (flow rate and time factor) are correlated, because of the hypothesis of constant void


Figure 13. Flow rate of each reaction species (CH3OH, H2O, CO, CO2, H2) along the TR. H2O/CH3OH ) 3, plumen ) 2 bar, T ) 300 °C, z ) reactor length.

Figure 14. Flow rate of each reaction species (CH3OH, H2O, CO, CO2, H2, lumen, H2, shell) along the MR. H2O/CH3OH ) 3, plumen ) 2 bar, pshell ) 1 bar, T ) 300 °C, sweep gas flow rate ) 2.2 × 10-3 mol/s, z ) reactor length.

fraction in the catalyst bed. In particular, considering the TR (Figure 13), the flow rates of the reactants are observed to decrease, with a resulting increasing trend for the products. The flow rate of H2 at TF ) 8.17 kgcat‚ s/molCH3OH (i.e., at the exit of a 20-cm-long TR) is about 0.003 mol/s at the operating conditions of this figure. With regard to the MR (Figure 14), a decreasing trend is again observed for the reactants, and an increasing trend is found for the products, with the exception of CO, which shows a maximum of about 0.001 mol/s at z ) 8 cm (TF ) 3.4 kgcat‚s/molCH3OH). The flow rate of H2 at the outlet of the MR is about 0.003 mol/s at the lumen exit and 0.0027 mol/s at the shell exit, with a resulting H2 production of 0.0057 mol/s in the MR, versus 0.003 mol/s in the TR. Therefore, the H2 production in the MR is about twice that of the TR, at the operating conditions of these figures. 5. Conclusions Results related to the modeling of the methanol steam-reforming reaction in both a TR and an MR

confirm the potentiality of using a membrane reactor to increase the reactant conversion, H2 selectivity, and H2 production over those obtained with a traditional system. By using an MR, it is possible to easily separate H2 from CO and all of the other reaction species, thereby generating a pure H2 stream that can be directly utilized in a polymer electrolyte membrane fuel cell. Improvements in both the CH3OH conversion and the H2 recovery were observed upon changing several reaction parameters, including the time factor, H2O/CH3OH feed ratio, reaction temperature and pressure, and sweep gas flow rate. Each parameter exerts an influence on the reactor behavior, both in the TR and in the MR. In particular, the following effects were observed: (1) The H2O/CH3OH feed ratio does not produce significant effects on the CH3OH conversion when the reaction (lumen pressure) is low (around 2 bar). Howvere, it is a key parameter at high reaction pressure: in this case, the H2O/CH3OH feed ratio causes a decrease of the CH3OH conversion. For example, for the MR at plumen ) 4 bar, the H2 recovery is 79.24% at H2O/ CH3OH ) 3 and 44.85% at H2O/CH3OH ) 10. (2) The effect of the reaction pressure on the behavior of the MR depends on the operating temperature. For example, at 300 °C, the CH3OH conversion in the MR is about 100% for plumen in the range 1-10 bar, whereas at 260 °C, the CH3OH conversion in the MR increases from about 70% at plumen ) 1 bar to 85% at plumen ) 10 bar. Concerning the TR, a value of the lumen pressure corresponding to a maximum in the methanol conversion was found; this value changes with the temperature and the TF. For example, at 280 °C and TF ) 1 kgcat‚ s/molCH3OH, the plumen value corresponding to the maximum methanol conversion (15.4%) is 5.05 bar; at the same temperature and TF ) 8.17 kgcat‚s/molCH3OH, the plumen value corresponding to the maximum methanol conversion (61%) is 3.25 bar; and at 200 °C and TF ) 8.17 kgcat‚s/molCH3OH, the plumen value corresponding to the maximum methanol conversion (33.5%) is about 1 bar. (3) In principle, an increase in the sweep gas flow rate produces an increase in the CH3OH conversion up to a plateau. In this work, changing the sweep gas flow rate from 2.2 × 10-3 or 2.2 × 10-5 to 2.2 mol/s yielded improvements in the CH3OH conversion. The most significant improvement was observed when this parameter was changed from 2.2 × 10-3 mol/s (about 60% CH3OH conversion at 260 °C) to 2.2 mol/s (more than 90% CH3OH conversion at 260 °C). The most interesting application concerns the use of steam as the sweep gas, and in each case, the use of a sweep gas leads to an increase in the methanol conversion compared to the case in which a sweep gas is not used. From an industrial point of view, the use of a sweep gas could correspond to keeping the shell side under vacuum. Otherwise, it is possible to use steam as the sweep gas, which is easy to remove after hydrogen extraction. In this way, it is possible to achieve a hydrogen/steam mixture from which the steam can be removed by condensation, so that a pure hydrogen stream can be obtained. (4) The main tuning parameters are the operating temperature and the time factor: high CH3OH conversions were obtained at high time factors. The best CH3OH conversion was reached at high temperature: 100% CH3OH conversion at 280 °C for the MR with a sweep


flow rate of 2.2 mol/s and a time factor of 8.17 kgcat‚s/ molCH3OH, compared to a corresponding value of 60% for the TR. (5) The maximum hydrogen recovery obtained in the MR was 95.82% at plumen ) 10 bar and H2O/CH3OH ) 1. (6) Using the MR, the H2 production can be about twice that obtained using the TR: 0.0057 mol/s of H2 produced in a 20-cm-long MR (TF ) 8.17 kgcat‚s/ molCH3OH) versus 0.003 mol/s of H2 produced in a 20cm-long TR, at T ) 300 °C, H2O/CH3OH ) 3, and plumen ) 2 bar. Nomenclature A ) cross section of the reactor (lumen) (m2) C ) total surface concentration (mol/m2) D ) internal diameter of the reactor (lumen) (m) E ) activation energy (kJ/mol) Ea ) apparent activation energy for hydrogen permeability () 29.16 kJ/mol) ∆H ) heat of adsorption of a species or heat of formation of the species (kJ/mol) J ) flux [mol/(s‚m2)] k ) rate constant [m2/(s‚mol)] K ) equilibrium constant or adsorption coefficient m ) H2O/CH3OH feed ratio MR ) membrane reactor N ) molar flow rate (mol/s) p ) partial pressure (bar) Pe0 ) preexponential factor of hydrogen permeability [) 1.12 × 10-5 mol‚m/(s‚m2‚Pa0.5)] R ) universal constant for ideal gas [J/(mol‚K)] rj ) rate of reaction j per unit weight of catalyst [mol/(s‚ kgcat)] ∆S ) entropy change for adsorption or reaction [kJ/(mol‚ K)] T ) absolute temperature (K) TR ) traditional reactor V ) reactor volume (m3) W ) mass of catalyst (kg) z ) axial coordinate (m) Greek Letter δ ) membrane thickness () 50 × 10-6 m) Superscripts 0 ) feed conditions ∞ ) corresponds to 1/T ) 0 (or 1/T ) ∞) in the intercept of the Arrhenius expression * ) composite parameter, whose expression and value are in Scheme 1 lumen ) lumen side of the membrane reactor shell ) shell side of the membrane reactor T ) total concentration of active sites Subscripts 298K ) absolute temperature of 25 °C ()298 K) CH3O1 ) adsorbed species CH3O on active site S1 CH3O2 ) adsorbed species CH3O on active site S2 CO2(1) ) adsorbed species CO2 on active site S1 CO2(2) ) adsorbed species CO2 on active site S2 d ) decomposition (reaction 3) H1a ) adsorbed species H on active site S1a H2a ) adsorbed species H on active site S2a HCOO1 ) adsorbed species HCOO on active site S1 HCOO2 ) adsorbed species HCOO on active site S2 i ) chemical species i (i ) H2, CO, CO2, H2O, CH3OH) j ) reaction j (j ) 1-3) OH1 ) adsorbed species OH on active site S1

OH2 ) adsorbed species OH on active site S2 r ) direct reforming (reaction 1) S1, S1a, S2, S2a ) active sites of the catalyst w ) water-gas shift (reaction 2)

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Received for review June 12, 2003 Revised manuscript received January 7, 2004 Accepted February 25, 2004 IE0304863

Hydrogen Recovery from Methanol Steam Reforming in a Dense

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