methane steam reforming in asymmetric pd- and

ELSEVIER Applied Catalysis A: General 119 ( 1994) 305-325

Methane steam reforming in asymmetric Pd- and Pd-Ag/porous SS membrane reactors

Jun Shu, Bernard P.A. Grandjean, Serge Kaliaguine * Department of Chemical Engineering and CERPIC, Lava1 University, Ste-Foy, Quebec GlK 7P4, Canada

Received 31 May 1994; revised 2 August 1994; accepted 2 August 1994

Abstract

This work is devoted to applying electrolessly deposited Pd- and Pd-Ag/porous stainless steel composite membranes in methane steam reforming. The methane conversion is significantly enhanced by the partial removal of hydrogen from the reaction location as a result of diffusion through the Pd- based membranes. For example, at a total pressure of 136 kPa, a temperature of 5OO”C, a molar steam- to-methane ratio of 3, and in the presence of a commercial Ni/A1203 catalyst together with continuous pumping on the permeation side, a methane conversion twice as high as that in a non-membrane reactor was reached by using a Pd/SS membrane. These effects were examined under a variety of experimental conditions. A computer model of the membrane reactor was also developed to predict the effects of membrane separation on methane conversion.

Keywords: Hydrogen permeation; Kinetic pcrtneation model; Membrane reactor; Methane stcatu reforming; Palladium film: Palladium-silver film

1. Introduction

Methane steam reforming is one of the most important chemical processes for the production of hydrogen or synthesis gas [ 11. This technology has been proved more economic than other processes such as coal vaporization, hydrocarbon partial oxidation, water electrolysis to produce hydrogen [ 21 for the reduction of iron ores, for the use in fuel cells and for hydrocracking, etc. In recent years, the abundant availability of natural gas and the increasing demand of hydrogen have led to more and more interest to develop this process further [ 3,4].

Methane steam reforming involves two reversible reactions: the reforming ( 1) and water-gas shift reaction (2) :

* Corresponding author. E-mail: [email protected], tel. ( + 1-418) 6562708, fax. ( + 1-418) 6567763.

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306 J. Shu et al. /Applied Catalysis A: General 119 (1994) 305-325

CH,+H,O+CO+3H, A@!!,= -206kJ/mol (1)

CO + HZ0 + COZ + H2 AH& = + 41 k.I/mol (2)

CI-I, + 2Hz0 + COZ + 4Hz A&,, = - 165 kJ/mol (3)

The reforming reaction ( 1) is endothermic and involves an increase in the number of gaseous molecules so that it is thermodynamically favored by high temperature and low pressure. In the conventional technology, this reaction is conducted on supported nickel catalysts in multitubular reactors operated at temperatures up to 85O”C, pressures ranging from 1.6 to 4.1 MPa and steam-to-methane ratios between 2 and 4. The elevated pressure is used to improve the energy efficiency of the overall process [ 51. The endothermic reaction heat is supplied by burning fuel in a radiant furnace. Methane conversion is usually around 78%, limited by the thermodynamic equilibrium.

The development of palladium-based membrane separation processes [ 6-91 has opened up a new possibility to enhance the methane steam reforming conversion. With tbe continuous withdrawal of hydrogen by a palladium-based membrane from the reaction location, the equilibrium of the reforming reaction ( 1) could be significantly shifted toward the right hand side, resulting in an increase of the methane conversion. In the meantime, high purity hydrogen could be obtained directly from the permeation side of the membrane reactor. Oertel et al. [lo] performed a calculation for a modified methane steam reforming process with integrated hydrogen separation. They demonstrated that almost 90% of methane conversion could be achieved below 850°C by using a palladium membrane with a thickness of 50 pm. They also fitted a Pd-membrane in a reformer to separate the produced hydrogen and obtained a methane conversion as high as 96% [ 10,111. The behavior of a palladium membrane reactor used in methane steam reforming was simulated by Prokopiev et al. [ 121. A computer simulation on methane steam reforming in a fluidized bed membrane reactor was also performed by Adris et al. [ 131. Near complete methane conversion was expected.

However, the high cost of palladium and the relatively low hydrogen permeation capacity do not allow massive palladium sheets to be commercially used on a large scale. Certain efforts have been reported in the literature either in the selection of new membrane materials or in the improvement of membrane morphology. Besides the exploration using catalytically modified microporous alumina membranes [ 14,151, an alternative choice is to develop composite thin palladium membranes. Kikuchi and coworkers [ 16-251 developed thin Pd and Pd-Ag membranes supported on either porous glass or porous A1203 ceramic substrates. These membranes showed much higher hydrogen permeability than commercial palladium based sheets. Using these membranes combined with a low-temperature reforming nickel catalyst, these authors [20-241 showed that methane steam reforming can be promoted far beyond the equilibrium. The problem with these substrates lies in their relatively poor mechanical and/or thermal properties. The present work deals

.I. Shu et al. /Applied Catalysis A: General 119 (1994) 305-325 307

with the preparation of thin Pd-based membranes supported on a porous stainless steel (SS) substrate and their uses in performing methane steam reforming in a membrane reactor. In order to further understand the experimental results, a computer model will be described to give some general predictions of the membrane reactor behavior.

2. Experimental

2. I. Membrane preparation

A porous seamless 3 16L stainless steel tube having an outside diameter of 0.95 cm and a nominal particle retention size of 0.2 pm was purchased from Mott Metallurgical. The tube was cut to a length of 3.6 cm, and welded to a dense stainless steel tube with the same diameter. The effective membrane area for hydrogen separation was about 10.7 cm*. This tube was then cleaned in an ultrasonic bath of carbon tetrachloride for one hour. Asymmetric Pd or Pd-Ag membranes were formed on the outer side of the porous tube by the electroless plating technique described previously [ 261. Composition and structural information about the two membranes are listed in Table 1. The impervious Pd/SS membrane was used as deposited. The deposited Pd-Ag membrane was thermally treated over their Tam- man temperature (550°C) for five hours in the presence of hydrogen. This treatment can lead to the formation of an impervious Pd-Ag alloy film [ 271.

2.2. Reaction testing

Methane steam reforming was carried out in the reaction system illustrated in Fig. 1. The detailed design of the membrane reactor used is shown in Fig. 2. The outer stainless steel tube had an I.D. of 1.7 cm and a length of 5.1 cm. The inner tube was the membrane part just described. This reactor assembly could be easily configured for the reaction test with or without membrane by a simple replacement of the inner tube. The non-membrane reactor configuration could be made by using a dense stainless steel inner tube. The shell side of the reactor served as the reaction chamber and the tube side as the permeation chamber.

Most of the parts were arranged in a modified autoclave reaction apparatus ( BTRS-JrTM , Autoclave Engineers Group). The assembly had two heating systems

Table 1

List of Pd-based membranes used in methane steam reforming

Membrane

Pdlporous SS

Pd-Aglporous SS

Composition Geometric dimension O.D. X h (cm) Estimated film thickness (pm)

Pd 0.95 x 3.6 19.8

Pd-5.1 %Ag 0.95 x 3.6 10.3


4 LC-600

nl\rir

Y

q-

HZ0

vent Plorrc*r

l-2(off) Puq~ l-S(on) Reaction

Fig. 1. Experimental apparatus for methane steam reforming.

with separate temperature controllers: a furnace and a surrounding reactor cabinet. The reactor was fixed in the middle of the furnace to maintain a temperature distribution as uniform as possible. The temperature was controlled by a Eurotherm

Fig. 2. Tubular membrane reactor design.

J. Shu et al. /Applied Catalysis A: General 119 (1994) 305-325 309

847 digital controller. The preheating cabinet allows mixers, switch valves and transfer lines to be maintained at 200°C without cold spots.

The dry reactant introduced to the reaction side was a mixture of methane with 5% helium. The inlet stream to the permeation side was pure helium used as a purge gas. Another helium stream was also employed to sweep the reaction side during start up and shut down of the reaction system. These gaseous streams were separately controlled by an MKS type 1259C mass flow controller combined with an MKS type 247C 4 channel readout. Before operation, the flow controller was calibrated with He, HZ, and CH4 + S%He gases using a digital soap bubble flow- meter (Wheaton Scientific). The reactant stream and the purge gas stream passed separately through the reaction side and the permeation side in a cocurrent flow mode (from bottom to top).

Deionized water was injected into the methane-5% helium mixture stream using an LC-600 Shimadzu Liquid Chromatograph injector. This module was a solvent delivery unit with a microvolume double-plunger reciprocating pump, which provided stable delivery with little flow pulsation compared with a conventional solvent delivery system. Its accuracy of flow-rate was within + 2%.

To enhance the reforming conversion, a commercial 12%Ni/A1203 catalyst (C 1 l-9-02) provided by United Catalysts was used. Before packing, catalyst pellets were crushed to 18-30 mesh size. About 11 g of catalyst were packed inside the reaction chamber of the reactor with glass fiber filling up small empty volumes of both the inlet and the outlet. The catalyst completely covered the Pd-based membrane surface. The permeation side contained no catalyst.

2.3. Operating procedures

Since the original steam reforming catalyst was provided in the form of nickel oxide supported on alumina, an in-situ activation of the catalyst was necessary after packing. This was done by first heating the catalyst bed to 300°C in a helium atmosphere to prevent the Pd-based membrane from developing cracks or pinholes, followed by introducing a hydrogen stream to the reactor at atmospheric pressure. Then the temperature was increased stepwise to 400°C and 550°C each step lasted for one hour. After activation, the system was conditioned for either hydrogen permeation measurement or steam reforming.

For methane steam reforming, the activated system was first heated to 300°C in a helium atmosphere. Once the conditions were achieved, a preheated reactant mixture of steam and CI& + 5%He, in a steam-to-methane ratio of 3, was fed into the catalyst bed through an air-actuated valve. Inert helium gas at a flow-rate of 40 ml ( STP) /min (abbreviated as SCCM) passed through the permeation side to purge the permeated hydrogen. The steam reforming was conducted in the temperature range of 300 to 550°C and a typical pressure of 136 kPa. A CH, + 5%He gas flow- rate of 42 SCCM was used corresponding to a gas hourly space velocity (GHSV) of about 1067 h- ‘. After condensation of the unreacted steam by passing the


reaction product through a water-ice mixture trap, the effluent was analyzed using an on-line MS (SMD-I-100 Mass Spectrometer, VG Gas Analysis System) and a TCD (HP 5890 Series II Gas Chromatograph) .

In the present case of a dense Pd-based membrane through which only hydrogen was permeated, methane conversion was calculated using the following expression based on a carbon balance:

X cH.+= G0.r + CCO2.r

c CH4.r + G0.r + Go2.r (4)

where the index r stands for reaction side.

3. Mathematical model

In comparison with analyses for common chemical reactors, the membrane reactor analysis must consider the additional contribution of mass transfer through the membrane, that is, the permeation rate should be involved along with the reaction rate expression, To investigate the concentration profile of methane steam reforming inside Pd-based membrane reactors, a kinetic permeation model was considered. A schematic illustration of the model is shown in Fig. 3.

3. I. Basic assumptions

( 1) Isothermal and isobaric reaction conditions; (2) Steady-state operation; (3) Plug flow on both reaction and permeation sides; (4) Intrinsic kinetics for methane steam reforming (MSR) and water-gas shift

(WCS) reactions; (5) No boundary layer on membrane surfaces.

Fig. 3. Illustration of the kinetic permeation model.

J. Shu et al. /Applied Catalysis A: General 119 (1994) 305-325

3.2. Reaction rates

311

Xu and Froment [28] studied carefully the intrinsic MSR and WGS reaction kinetics on a Ni/MgAl,O, catalyst in a classical tubular reactor. The developed model was proved to be more general than previous literature models for nickel catalysts [ 291. This model was adopted in the present study. The model, based on a Langmuir-Hinshelwood reaction mechanism, contains 13 steps. The rate expres- sions for reactions ( 1) to (3) were given as:

R, = h(pCH4pH20 -9) /DEN2 Pk: 1

R2 =~(pCOpHzO -‘y) /DEN2 2

k3 4

R3 =--&pm P&O - 2

py) /DEN2 3

DEN = 1+ Kco PCO + KHZ PHI + Kcm PCH~ + ho ~~201~~2 with the partial pressures given by the relations:

pcH4= (1 -&LA/~ PHzO= (m-XCH4 -&oz) /a

PCO= (XCH4-XCOz)/a PHz = @I? + 3xCH, + Go* - YH)/(T

PC02 =&02/u (T= ( 1+m+p&+2XoHq-Yn)/Pr

(5)

(f-3)

(7)

(8)

where, P, is the pressure on the reaction side of the reactor; Xc, is the total methane converston; Xco2 the methane conversion to CO,; YH is the permeated hydrogen fraction with respect to the integrated hydrogen on the reaction side; and m the steam-to-methane ratio. The meaning of p& will be discussed below.

To estimate the model parameters, a kinetic study of methane steam reforming over C 1 l-9-02 Ni/Al,O, catalyst was performed in a Betty reactor [ 271. The study indicated that the specific model developed by Xu and Froment [ 281 fitted essentially our methane conversion results over Cl l-9-02 catalyst. Thus those model parameters obtained by Xu and Froment were also adopted in the present study, as listed in Table 2. Reaction rates for the disappearance (i.e., total conversion) of

Table 2

Kinetic input parameters [ 281

Paramete r Pre-exponential factor b Eor AH (kJ/mol )

k, (km01 MPa’.‘/kg h) k2 (kmol/kg h MPa)

ka (km01 MPa”.‘/kg h)

f&W'-')

KH2 (MPaC’)

KcH4 (ma-‘)

KHZ0

1.336. lOI 240.10

1.955 10’ 67.13

3.226. lOI 243.90

8.23. 1O-4 - 70.65

6.12.1OF’ - 82.90

6.65 10-S - 38.28

1.77.1@ 88.68

312 J. Shu ef al. /Applied Catalysis A: General 119 (1994) 305-325

Table 3 Influence of p&/pcti value on predicted methane conversion (conditions: 5WC, 136 kpa, H,O/CI-& = 3 and GHSV= 1067 h-l)

Required step size AL’ Predicted Xc, at the reactor exit

1 0.8 0.5 0.2 0.1 0.08 0.05 0.02 0.01

Equilibrium

0.0001 0.0001 0.0001 0.0001 0.0001 0.0001 0.00005 0.00001 o.ooooO5

0.264 0.288 0.326 0.366 0.381 0.383 0.387 0.392 0.393 0.394

methane and for the formation of carbon dioxide in methane steam reforming were thus given as:

R dx CH4

CH4 - --=R1 +Rg

R ax co2

co2 - --=R2+R3

(9)

(10)

where W is the catalyst weight, and F&, is the flow-rate of methane in the reactor entrance.

Eqs. (9) and ( 10) can be further expressed by the introduction of a dimensionless reactor length L’, dL* =dLIL:

dx,H4_Pm-h3(R +R )

dL* F& ’ 3 (11)

.o*_Pwe3 dx

dL* F&4 (RI + 4) (12)

In Xu and Froment’s kinetic model, a critical parameter is the hydrogen partial pressure. The present study used a hydrogen-free methane-steam feedstock, thus the entrance hydrogen partial pressure was zero which made the initial rate of methane conversion infinite. The problem was solved by setting a very small inlet hydrogen partial pressure introduced as an arbitrary value of &,/pcn,. An evalu- ation of the p&/p cti parameter setting upon the predicted methane conversion is shown in Table 3. Although a smallerp~,/po,, value results in a higher calculation precision, it needs much a smaller calculation step size, thus causing a longer calculation time. The computation showed that at high residence time ( W/F), the

J, Shu et al. /Applied Catalysis A: General 119 (1994) 305-325 313

predicted methane conversion was in good agreement with the equilibrium one when a ph21pCti value less than 0.1 was used. Thus the value 0.1 for p&/pcH4 was chosen in the present kinetic permeation model. This treatment is physically realistic considering the rapid conversion of methane to hydrogen and carbon dioxide, as pointed out by Xu and Froment [ 281.

3.3. Permeationjux

The hydrogen permeation rate through the palladium membrane with a thickness S has the following differential equation form:

-= (13)

with pH2,p = H p Y P / ( YH + I), where Z is the ratio of inert gas molar flow-rate with respect to methane flow-rate, Z= Fy/F’ CH4; Q. and Ep are the pre-exponential factor and the activation energy of hydrogen permeation, respectively.

This equation is de facto derived from the extended Sieve&s relation used for the a-hydride phase regime [ 271. Since methane steam reforming needs high temperature operating conditions (e.g., > 500°C)) this a-hydride approach would not lead to a significant error.

3.4. Boundary conditions

At L* =0: X,,,=X,,,=Y,=O

3.5. Numerical simulation

The set of differential Eqs. ( 11-13) was solved numerically by a 4th order Runge-Kutta method [ 301.

4. Results and discussion

4.1. Methane steam reforming in a non-membrane reactor

As a reference, methane steam reforming was first tested in a reactor without hydrogen-permeable membrane. For this purpose, a dense stainless steel tube was employed in the reactor assembly in place of the membrane. After carefully starting up, following the experimental procedures described previously, methane conversion was measured as a function of temperature. A reaction pressure of 136 kPa was set considering the favorable reaction equilibrium at low pressure and the requirement for further membrane effect test. The steam-to-methane ratio used was


0 1 2 3 4

Reaction time, II

Fig. 4. Product distribution of methane steam reforming at various temperatures in the non-membrane reactor

(conditions refer to Fig. 5).

3. Fig. 4 shows an example of the steam reforming product distribution at various reaction temperatures recorded using the mass spectrometer. The measured total methane conversion data are plotted in Fig. 5 (Curve 1) . Repeated experiments gave almost the same methane conversion results. The catalytic reaction was in a steady state over the activated catalyst. Furthermore, raising the GHSV from 1067 h- ’ to 1600 h- ’ resulted in no change in methane conversion. It can be seen that the methane conversion increases monotonically with increasing temperature in this temperature range. The main reforming product is carbon dioxide at low reaction temperature. For example, the methane selectivity to carbon dioxide is 96.6% at 400°C and 86.5% at 500°C.

For comparison, the equilibrium methane conversion was calculated using a QuickBasic program. In the calculation, equilibrium constants for MSR and WGS reactions were taken from ref. [2]. The calculated equilibrium values were also plotted in Fig. 5 (Curve 2). As can be seen, the measured methane conversion was slightly lower than the equilibrium value, and the difference increases with increasing temperature and thereby methane conversion.

A prediction from Xu and Froment’s model indicated that near equilibrium methane conversion could be achieved under the above reaction conditions. An X- ray diffraction (XRD) measurement (Fig. 6) of the reforming catalyst showed that metallic nickel, which is the active species in methane steam reforming, is present on the used catalyst after activation even though the mild reduction temperature of 550°C was used. In the original C 1 l-9-02 reforming catalyst, the nickel component was supported on a-A1203 mostly in the form of NiO. Unlike some other steam reforming catalysts, no NiA1204 spine1 was found in this original catalyst. This


40

I

30 -

20 -

250 300 350 400 450 500 550

Temperature,“C

Fig. 5. Methane conversion in a non-membrane reactor: ( 1) non-membrane; (2) calculated equilibrium.

might be the characteristics of the low-temperature reforming catalyst as the reduction of NiO is much easier than that of NiA1204. After the activation at 550°C in the hydrogen atmosphere, NiO was entirely reduced, as revealed by XRD.

4000

5 d ^

2‘ 3000 a B s

2000

loo0

20 30 40 50 60 70 2theta

Fig. 6. X-ray diffraction patterns of the steam reforming catalyst: (a) original catalyst; (b) after reduction at 550°C.


Table 4 Possible carbon deposition reactions in methane steam reforming

No. Reaction A@%? (kJ/mol)

KP at 773 K (zzFqi (units as K,)

(ZIP”‘) .lK I I

1 C&+C+2H, - 74.8 51.7 kPa 91.2 1.76 2 2co+c+c4 172.5 2.56 kPa_’ 4.5 1 1.76 3 CO+H2+C+HZ0 131.3 0.46 Wa- ’ 1.14 2.47 4 CO* + 2H2 90.1 1.87. 10d4 kPa_’ 0.29 1.57. lo3

+C+2H,O 5 cH4+2co 187.8 11.2 Wa-’ 120.4 10.8

+ 3C + 2H20 6 CH, + co* 15.3 4.35 26.7 6.14

-+ 2C + 2H20

The possibility of the catalyst deactivation caused by carbon deposition over the catalyst was examined. Table 4 lists possible carbon deposition reactions in the methane steam reforming process. Equilibrium constants of each reaction at 500°C were calculated based on thermodynamic data taken from ref. [ 3 11. Under our typical reaction conditions of XWC, 136 kPa and a steam-to-methane ratio of 3, the equilibrium criteria (nP,Yi)J& (here vj are the stoichiometric coefficients,

with sign plus for products’and minus for reactants) for each reaction exceed 1 in accordance with the calculation based on the corresponding reaction product distribution. That is to say, no carbon species would deposit over the catalyst surface. This was confirmed by the fact that no catalyst deactivation evidence was found after ca. 24 h of operation.

After a careful analysis of the experimental conditions, the difference between the measured methane conversions and the equilibrium values observed in Fig. 5 is believed to result from the temperature gradient inside the integral reactor. The reactor was made of stainless steel with a thermocouple tube welded to the reactor

body. The thermocouple measuring point was set at the exit of the reactor in order to measure the exit temperature. However, because of the integrated reactor form and the poor thermal conductivity of the catalyst, the strongly endothermic steam reforming reaction might result in a temperature difference between the catalyst bed and the reactor wall. Temperature gradients as high as 10°C do exist inside reformers [ 51. The present temperature is measured close to the reactor wall while the catalyst bed temperature tends to decrease toward the inner tube. Since the purge gas (helium) was also used in this non-membrane system, it constituted another important cooling factor. The helium stream was preheated to 200°C before entering into the tube side of the reactor. This effect would also enlarge the temperature gradient across the catalyst bed. Unfortunately, the temperature profile is hard to calculate since the heat transfer from the environment is unknown. For the sake of comparison, the measured conversion in the non-membrane system was used as a reference in the present study.


0 30 40 60 80

Hz removal percent, %

100

Fig. 7. Influence of hydrogen removal on the equilibrium methane conversion (conditions: P, = 136 kPa; Pp = 101

kPa; H,O/CH.,= 3; FcH4 = 40 SCCM; FHe = 40 SCCM)

4.2. Temperature effect in the membrane reactor

The major advantage of the membrane reactor is the conversion enhancement of an equilibrium-limited reversible reaction by selective hydrogen removal. In the case of methane steam reforming, the use of a membrane reactor can lead to a significantly enhanced methane conversion at a moderate temperature. Fig. 7 shows such a thermodynamic equilibrium calculation of the methane conversion as a function of the hydrogen removal ratio. This was done by simply setting a residual hydrogen partial pressure in the equilibrium calculation for methane steam reforming.

As can be seen from Fig. 7, when the temperature is below 4OO”C, the reforming reaction could not reach a satisfactory conversion level until a high hydrogen separation efficiency, say 90%. Over 7OO”C, the hydrogen removal through the membrane can only slightly shift the reforming equilibrium due to the intrinsically favorable temperature effect. At a moderate temperature of 500 to 6OO”C, membrane separation can result in a great improvement on the methane steam reforming equilibrium. This is why moderate temperature conditions were used in the present study.

The temperature effect of membrane separation on MSR was studied under the same conditions as in the case of the non-membrane reactor. Both Pd/SS and Pd- Ag/SS membranes were examined. When methane steam reforming was performed in the palladium membrane reactor without a catalyst, no detectable conversion


136kPa, H~O/CH4=3 GHSV= 1067li’

01 I I

250 300 350 400 450 500 550

Temperature,% Fig. 8. Temperature effect in Pd-based membrane reactor: ( 1) non-membrane; (2) calculated equilibrium; (3) Pd/SS; (4) calculated from the model for Pd with 6= 20 pm; (5) Pd-Ag/SS.

occurred at XMYC, which might be due to the low specific surface area of the membrane as well as its limited intrinsic catalytic activity. Thus a catalyst was necessary to perform effectively the reforming reaction. Fig. 8 illustrates the total methane conversion against reaction temperature. In comparison with the methane conversion in the non-membrane system (Curve 1) , the membrane effect is clear in the Pd/porous SS membrane case (Curve 3). This conversion curve is also beyond the calculated equilibrium methane conversion (Curve 2). No deactivation evidence of the membrane was observed after about 47 h of reaction tests. The membrane remained still dense showing no cracks or pinholes, and permselective toward hydrogen separation. No carbon-based reactant or products were detected in the permeation side. It is believed that the close thermal expansion coefficients of palladium ( 11.8 pm/m “C at 20°C) and 3 16L stainless steel ( 14.7 pm/m “C at 20°C)) compared to the value for alumina ceramics (5.4 pm/m “C at 20°C) make the Pd/SS membrane more resistant to thermal cycling than Pd/ceramic membranes [24]. Curve 5 is the methane conversion performed in the Pd-Ag/SS membrane reactor. Evidently, methane steam reforming was greatly enhanced by using this Pd-Ag membrane. For example, at a total pressure of 136 kPa, a reaction temperature of XKY’C, and a molar steam-to-methane ratio of 3, a methane conversion of 50.9% was achieved, much higher than that in the non-membrane system (36.7%). The methane conversion was reproducible after a series of reaction tests.


When compared with the result from Pd/SS membrane, the enhancement in the case of Pd-Ag/SS should be attributed to the combination of the use of a thinner film and a possibly higher hydrogen permeability through the Pd-Ag membrane.

The temperature effect was also calculated using the kinetic permeation model for the Pd/SS membrane. Under the above mentioned reaction conditions, methane conversions at various temperatures were predicted, as shown in Fig. 8 (Curve 4). Additional input parameters include a membrane thickness of 20 pm, a purge gas flow-rate of 40 SCCM, a hydrogen permeation activation energy of 15.7 kJ/mol and a pre-exponential factor Q, of 1.56 - lo-* m3/m s Pao.5 [ 321. Comparing the calculated data with Curve 3, it can be seen that the two lines fit essentially at temperatures below 350°C. With raising reaction temperature, the measured conversions become lower than the predicted ones. The overestimated calculation is likely attributed to the temperature gradient inside the membrane reactor as well as our simplified model assumption. Both the endothermal behavior of the steam reforming and the cooling effect of the sweep gas yield a temperature decreasing toward the inner tube. A boundary layer might exist on the membrane surface toward the porous support. Besides the axial hydrogen concentration gradient, there is a radial distribution of permeated hydrogen through the porous support when the driving force is given by use of a sweep gas.

4.3. Influence of the H,O/CH, ratio

Methane steam reforming usually proceeds in the presence of an excess of steam to prevent the carbon deposition over the catalyst surface and to enhance the steam reforming. The effect of molar steam-to-methane (H,O/CI&) ratio on methane conversion was examined by varying this ratio from 2 to 4. The results obtained at 500°C are drawn in Fig. 9. Curve 1 is the experimental data obtained from the non- membrane system. Curve 2 is the calculated equilibrium conversion, exhibiting monotonic increase with the steam-to-methane ratio. It can be seen that these points are slightly lower than the equilibrium conversion, showing the same trend as in Fig. 8. In the case of the Pd/SS membrane, an enhancement of the membrane separation efficiency was realized by continually pumping the permeation side (tube side) with a rotary vacuum pump. The pumping eliminated the radial distribution of the hydrogen concentration inside the support pores and kept the permeation side under a low pressure, resulting in a relatively high driving force of hydrogen permeation through the palladium membrane. The measured methane conversion against the steam-to-methane ratio is shown as Curve 4 in Fig. 9. It is clear that the methane conversion was greatly enhanced in this manner. Under our typical operating conditions (500°C 136 kPa and a molar steam-to-methane ratio of 3), the achieved methane conversion was 63%, almost twice as high as the one in the non-membrane system. For the Pd-Ag/SS membrane, the effect of HZ01 CH, is shown in Curve 3. In this case, helium passed through the permeation side at a flow-rate of 40 SCCM. The promotion of the reforming conversion is clearly

320 J. Shu et al. /Applied Catalysis A: General I19 (1994) 305-325

loo1

60

201 ! 1.5 2 2.5 3 3.5 4 4.5

Steam/methane ratio

SO0 %. 136kPa GHSV~1067ti’

Fig. 9. Steam-to-methane ratio effect in membrane reactor: ( 1) non-membrane; (2) calculated equilibrium; (3)

Pd-Ag/SS; (4) Pd/SS; (5) calculated from the model for Pd with 6= 20 pm.

exhibited. It should be noticed that the lower conversion obtained with the Pd-Ag/ SS membrane is due to the fact that in this experiment a smaller difference in hydrogen partial pressure in the two compartments was implemented which induced a lower hydrogen permeation flux.

It is somewhat difficult to model the steam-to-methane ratio effect in the present Pd/SS membrane case for lack of the exact pressure value on the permeation side. A calculation using the kinetic permeation model indicates that a very low pressure, as low as 1.3 Pa, on the permeation side would result in a complete methane conversion. However, it is possible for a certain amount of residual gas to exist on the permeation side considering the continual hydrogen permeation from the reaction side and a long transfer tubing between the membrane reactor and the vacuum pump. If a residual gas pressure of 30 kl?a is supposed, the calculated methane conversion is approaching the measured conversion for the Pd/SS membrane, as shown in Curve 5 of Fig. 9.

4.4. Pressure efSect in the membrane reactor

In a conventional reactor, the increase of pressure results in a decrease of the methane conversion. However, the incorporation of a hydrogen permeable membrane would change this general trend. Although a high reaction pressure


SOO°C.H,0/C&=3 GHSV=1067b"

I

6

5 . . . . . --.

100 120 140 160 180 200 220

Reaction pressure, kPa

Fig. 10. Pressure effect on methane conversion: ( 1) non-membrane; (2) calculated equilibrium; (3) Pd-Ag/SS,

FHc = 40 SCCM; (4,5,6) calculated from the model for Pd with 6= 20 pm and FHe = 40, 140, 200 SCCM,

respectively.

depletes the steam reforming conversion equilibrium, the increase of hydrogen partial pressure on the reaction side would increase the driving force for hydrogen permeation, resulting in an enhancement of the methane conversion under certain conditions.

The effect of pressure on methane steam reforming conversion is shown in Fig. 10. As expected, in the case of the non-membrane system, the methane conversion decreased with increasing total reaction pressure, in agreement with the equilibrium calculation. However, in the case of Pd-Ag/SS membrane, a decreasing variation in the methane conversion with the increase of pressure was also observed. This seems to indicate a low membrane separation efficiency.

Uemiya et al. [ 241 noticed a conversion increase with increasing reaction pressure. The sweep gas they used was an argon stream in a flow-rate of 400 SCCM, ten times higher than that in our case. An increasing conversion upon pressure was predicted by Prokopiev et al. [ 121 based on a mathematical model calculation. In their calculation, the partial pressure of permeated hydrogen was taken as zero. We also calculated the dependence of methane conversion on the total pressure under various purging conditions for a palladium membrane of 20 pm thick. It was found that the increasing tendency exists in the case where hydrogen separation is efficient. If the separation is relatively weak, for example with a mild purging effect on the permeation side, the decreasing methane conversion upon pressure is maintained


in the low pressure region. Large flow-rates of the purge gas bring in a direct increase of methane conversion with the pressure. These results are reported in Fig. 10 where three levels of helium flow-rate are examined. There is thus no contra- diction between our observation and the literature results. In our case, the use of a mild purge flow-rate (40 SCCM) and a low reaction pressure yielded a decreasing conversion dependence upon the pressure.

4.5. Purge effect

Purging the permeation side is important in maintaining a high hydrogen partial pressure difference between the two sides of the membrane and moving out hydrogen product from the reaction system. As discussed above, the increase of the purge gas flow-rate is expected to enhance the methane conversion owing to the decrease of the hydrogen partial pressure on the permeation side. This effect was examined in the Pd/SS membrane system, as shown in Fig. 11 (Curve 1). The methane conversion was found slightly increased with raising the purge gas flow-rate.

The computer model analysis for the palladium membrane predicts a stronger sweeping flow-rate effect on the methane conversion (Curve 2 in Fig. 11) than the one determined in our experiment. The effect reported by Uemiya et al. [24] is also lower than the calculated purge effect for the same reaction under similar

ioor

80 -

20 -

O- 20

-

soo”c, H*O/CH,=3 136kPa. GHSV=1067li’

60 100 140 180

He flow rate, SCCM

Fig. 11. Purge flow-rate effect on methane conversion: (1) Pd/SS; (2) calculated from the model for Pd with S=2Opm.


conditions. This may imply that there is a great potential to improve the membrane separation efficiency. In our experiment, the higher the flow-rate of the purge gas, the more the measured methane conversion deviated from the predicted value. This was quite possibly caused by the cooling effect of the purge gas stream as well as the oversimplified boundary layer condition. In future work, the improvement of the reactor setup to eliminate this cooling effect or measuring the temperature profile in the reactor will be useful to completely understand the membrane effect- iveness.

In general, increasing the purge gas flow-rate enhances the hydrogen separation efficiency. However, the economy in a large scale industrial application may not permit the use of large amounts of inert gas. Instead, in the case of methane steam reforming, steam may be used as the purge gas. This would also make the subsequent separation of hydrogen from the purge stream much easier since steam can be easily condensed from the mixture. Another choice to increase the hydrogen permeation driving force is the vacuum operation on the permeation side, as already demonstrated in this study. It seems that operation under rough vacuum ranging from 10 1 kPa to ca. 100 Pa is industrially acceptable from economic considerations [ 331.

5. Conclusions

Electrolessly prepared Pd/SS and Pd-Ag/SS composite membranes were used to perform methane steam reforming in the presence of a Ni/A1203 catalyst. The methane conversion was significantly enhanced as a result of diffusion through the Pd-based membrane by the partial hydrogen removal from the reaction location. For example, at a total pressure of 136 kPa, a temperature of 500°C and a molar steam-to-methane ratio of 3 together with continuous pumping on the permeation side, a methane conversion twice as high as the conversion in the non-membrane system was reached by using a Pd/SS membrane. No deactivation evidence of this membrane was observed after about 47 h of operation. These effects were examined under a variety of experimental conditions. A computer model of the reactor was developed and utilized in predicting the effects of membrane separation on methane conversion.

Nomenclature

c molar fraction (%) E activation energy (J/mol) EP activation energy of hydrogen permeation (J/mol) F flow-rate ( m3( STP) /s) GHSV gas hourly space velocity (h-i) h height (m)

324

AH

k K L L* m

O.D.

P P

Qo

i? T

W X Y

J. Shu et al. /Applied Catalysis A: General I1 9 (1994) 305-325

heat of reaction (J/mol) rate constant equilibrium constant reactor length (m) dimensionless reactor length molar steam/methane ratio outer diameter (m) partial pressure (Pa) total pressure (Pa) pre-exponential factor ( m3/m s Pa”.5) reactor radius (m) reaction rate (kmol/kg-cat h) ; or gas constant (8.314 J/mol K) temperature (K) catalyst weight (kg) conversion (%) permeated hydrogen fraction with respect to the produced hydrogen (%)

6. I. Greek letters

s membrane thickness (m)

P catalyst density ( kg/m3)

6.2. Subscripts

r reaction side

P permeation side i inner 0 outer

vj stoichiometric coefficient

6.3. Superscript

0 initial

Acknowledgements

This work was supported by the Natural Sciences and Engineering Research Council (NSERC) of Canada. The authors thank CANMET-EDRL at Varennes for providing facilities to perform methane steam reforming, and United Catalysts, Inc. (Louisville, KY) for supplying Ni/A1203 steam reforming catalysts. They


also acknowledge Drs. Gilles Jean, Jean Paquette and Michel Pokier for valuable experimental suggestions.

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methane steam reforming in asymmetric pd- and

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