Staging Distribution of Oxygen in Circulating Fast ... - ACS Publications

Ind. Eng. Chem. Res. 2007, 46, 5493-5502

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Staging Distribution of Oxygen in Circulating Fast Fluidized-Bed Membrane Reactors for the Production of Hydrogen Mohamed E. E. Abashar* and Fahad M. Alhabdan Department of Chemical Engineering, College of Engineering, King Saud UniVersity, P.O. Box 800, Riyadh 11421, Saudi Arabia

Said S. E. H. Elnashaie EnVironmental Engineering Department, PennsylVania State UniVersity at Harrisburg, Capital College, Middletown, PennsylVania 17057-4898

A mathematical model is used to simulate coupled steam and oxidative reforming reactions of methane for heat integration and to improve the yield of hydrogen in a circulating fast fluidized-bed membrane reactor (CFFBMR). A composite membrane of a very thin layer of palladium silver alloy is used for the removal of hydrogen, shifting the thermodynamic equilibrium and achieving the production of pure hydrogen that is suitable for fuel cells from one integrated unit. A distributed oxygen feed through discrete injection points along the length of the reformer to control the heat in the reaction medium is compared with oxygen feeding through dense perovskite membranes and the conventional direct feeding of oxygen (co-feed). Equispaced and orthogonal injection points for oxygen distribution, as well as uniform and nonuniform oxygen flow distribution, are investigated. The results indicate that the staging distribution of oxygen enables one to control the hot-spot temperatures and to operate the reactor under safer operability conditions in contradistinction to the possibility of thermal runaway for the conventional co-feed mode of operation. The results also show that the nonuniform oxygen flow distribution is superior, with regard to hydrogen yield over uniform oxygen flow distribution, for both equispaced and orthogonal injection systems. This is due to the good control and high oxygen fluxes at the injection points. The equispaced injection system gives better hydrogen yield than the orthogonal injection system in a certain region of parameter space and vice versa in other regions. Among the advantages of distributed oxygen feed is that the reformer can be operated at low feed temperatures. Introduction Recently, hydrogen has come into considerable prominence as a clean fuel that produces no pollution or greenhouse gases and has the highest energy content. Methane steam reforming is the main chemical process that is used to produce the hydrogen-rich synthesis gas.1-3 The widespread availability of natural gas and the increase in the demand for hydrogen that is utilized by many processes (such as oil refining, methanol production, metallurgy, ammonia production, space transportation, etc.) have imposed a strong economic incentive to improve the hydrogen production technology. In fact, ∼50% of the demand for hydrogen is satisfied by means of methane steam reforming.4 Methane steam reforming reactions are strongly endothermic, and the number of moles increases; therefore, thermodynamic equilibrium is favored by low pressure and high temperature. The conventional steam reformers are mainly fixed-bed catalytic reactors. The large catalyst particles are normally loaded into a large number of tubes that are placed in a top-fired or sidefired furnace. The catalyst tubes operate at high temperatures, varying from 500 °C to 800 °C.1 This fixed-bed configuration suffers from high diffusion limitations and catalyst deactivation.1,5,6 Moreover, the maximum temperature is limited by the catalyst properties and the metallurgical properties of the tubes. Numerous experimental and theoretical studies have examined the different aspects of improving steam reforming of methane, such as the thermal efficiency, the nickel-based catalyst proper* To whom correspondence should be addressed. Tel.: +966-14675843. Fax: +966-1-4678770. E-mail address: [email protected].

ties, and the design and operations.7-9 Applications of palladium-based membranes in steam reforming processes have become an interesting field, because of their permselectivity to hydrogen.4,10-12 Initially, thick dense membranes were used and low hydrogen fluxes were obtained.11 An effective method to enhance the hydrogen permeation flux is to use a composite membrane in which a very thin layer of palladium or palladium alloy has been deposited on the surface of a porous metal substrate.10,12-14 Fluidized-bed membrane reactors (FBMRs) for methane steam reforming have been attracting considerable attention, because of their desirable characteristics (such as the elimination of diffusional resistances, which is due to the use of powdered catalysts; the shift of the thermodynamic equilibrium, which is due to the use of effective hydrogen-selective membranes, leading to enhancement of methane conversion; simultaneous reaction and separation of hydrogen; and excellent temperature control, allowing for a wide range of flexible operating regimes15-19). Research groups led by Grace and Lim at the University of British Columbia have shown that the endothermic heat of reaction of methane steam reforming reactions could be provided by the partial oxidation of methane with limited effect on the hydrogen production rate.20,21 In fact, the influence of oxygen addition to supply heat for endothermic reactions has been considered by many investigators.22-26 Despite the advantages of the FBMR, the reactor suffers from the bypass of the bubbles, complex dynamics and hydrodynamics, catalyst entrainment, and erosion of the tubes. The application of circulating fast fluidized bed membrane reactors (CFFBMRs) for hydrogen production has been recently

10.1021/ie061579y CCC: $37.00 © 2007 American Chemical Society Published on Web 07/18/2007

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Ind. Eng. Chem. Res., Vol. 46, No. 17, 2007

Table 1. Reaction Rate Parameters for Steam and Oxidative Reforming of Methanea Kinetics Parameters

( )

ki ) Ai exp -

Ei RT

ki k1 (kmol kPa0.5/(kgcat h)) k2 (kmol/(kPa kgcat h)) k3 (kmol Pa0.5/(kgcat h)) k4 (kmol/(kPa2 kgcat h)) k5 (kmol/(kPa2 kgcat h)) k6 (kmol/(kPa2 kgcat h))

(

Kj ) Bj exp KCH4 (kPa-1) KH2O KH2 (kPa-1) KCO (kPa-1)

(for i ) 1, ..., 6)

Ai 9.49 × 1016 4.39 × 104 2.29 × 1016 3.96 × 109 15.08 8.71

Ei 240.10 kJ/mol 67.13 kJ/mol 243.90 kJ/mol 166.00 kJ/mol 29.00 kJ/mol 23.70 kJ/mol

Adsorption Constants

)

∆Hj RT

(for j ) CH4, H2O, H2, CO)

Kj

Bj 6.65 × 10-6 1.77 × 105 6.12 × 10-11 8.23 × 10-7

∆Hj -38.28 kJ/mol 88.68 kJ/mol -82.90 kJ/mol -70.65 kJ/mol

Reaction Equilibrium Constants

( 4400.0 K ) exp( - 4.063) T

K1 (kPa ) ) 10266.76 exp 2

26830.0 + 30.114 T

)

2

K3 (kPa2) ) K1K2 + 38.900) ( 26262.0 T 30782.0 + 42.970) K (kPa ) ) exp(T K5 (kPa2) ) exp 2

6

a

Data taken from refs 25 and 32.

co-workers.27,28

pioneered by Elnashaie and There are several advantages of circulating fluidized-bed reactors, which include the following: effective gas-solids contacting; uniform temperature distribution; the ability to use different particulate materials with high gas throughputs per unit cross-section; external recycling of particles, which allows continuous controlled operations with easy handling; and high steam-to-carbon ratio.28-31 Despite the fact that the dense perovskite membranes for oxygen permeation opened up new avenues to enhance equilibrium-limited reaction and achieving in situ isothermality, the fluxes of oxygen through these membranes are still very low and not satisfactory for commercial-scale reactors. Moreover, it was determined that the dense membrane was unstable under the conditions of methane partial oxidation reactions due to a phase transformation.24 In the present study, hydrogen is produced by the coupled steam and oxidative reforming of methane. The oxidative reforming reactions can supply the heat necessary for the endothermic reforming reactions, which can achieve in situ heat integration. Moreover, the carbon dioxide reforming eliminates this important greenhouse gas, which is harmful to the environment. The main difference between this study and the earlier studies for the production of hydrogen in CFFBMRs is related to focusing on the staging distribution of oxygen reactant along the reformer height. The advantages of oxygen staging include the following: the injection locations and distribution of oxygen can be controlled, high fluxes of oxygen may be achieved, the risk of reactor temperature run away is minimized by leveling the temperature profile, and the elimination of immersed oxygen tubes allows more free cross-sectional area for the reactions. Equispaced and orthogonal distribution of the injection points

are considered in this study, and a comparison with the use of oxygen-selective membranes also is conducted. In addition, simple dynamic programming is used to estimate the oxygen flow distribution. The cases without membranes and co-feed fast fluidized bed membrane reactors are also investigated. An external air separation system with an oxygen-permselective perovskite membrane is suggested. The distinctive features of this process are explored, and the effects of some key parameters are investigated. Reaction Kinetics Kinetic Rate Equations for Steam Reforming of Methane. The literature contains a large number of kinetic models of catalytic methane-steam reactions, and different rate expressions are available. Xu and Froment32 have developed moregeneral and realistic intrinsic rate equations for the steam reforming of methane using an integral flow reactor and a commercial nickel-based catalyst. Rate expressions for a reaction network consisting of the following three representative reactions were obtained:

CH4 + H2O h 3H2 + CO CO + H2O h H2 + CO2

[∆H°298 ) 206.0 kJ/mol] (1) [∆H°298 ) -41.0 kJ/mol]

(2)

CH4 + 2H2O h 4H2 + CO2

[∆H°298 ) 164.9 kJ/mol] (3)

CH4 + 2O2 f CO2 + 2H2O

[∆H°298 ) -802.0 kJ/mol] (4)

CH4 + H2O h 3H2 + CO

[∆H°298 ) 206.0 kJ/mol] (5)

CH4 + CO2 h 2H2 + 2CO

[∆H°298 ) 246.9 kJ/mol] (6)

Kinetic Rate Equations for Oxidative Reforming of Methane. Jin et al.25 developed reaction kinetics for the oxidation of methane, followed by steam and carbon dioxide reforming over a nickel-based catalyst, as follows: The rate expressions for reactions 1-6 are given as follows:

r1 )

(k1/PH22.5)[PCH4PH2O - (PH23PCO/K1)]

r2 )

r3 )

DEN2 (k2/PH2)[PCOPH2O - (PH2PCO2/K2)] DEN2

(k3/PH23.5)[PCH4PH2O2 - (PH24PCO2/K3)] DEN2

(7)

(8)

(9)

where

DEN ) 1 + KCH4PCH4 + KH2PH2 + KCOPCO + r4 ) k4PCH4PO2

( (

r5 ) k5PCH4PH2O 1 r6 ) k6PCH4PCO2 1 -

PH23PCO K5PCH4PH2O PH22PCO2

KH2OPH2O P H2

) )

K6PCH4PCO2

(10) (11) (12)

(13)

Ind. Eng. Chem. Res., Vol. 46, No. 17, 2007 5495

Figure 1. Schematic diagram of the circulating fast fluidized-bed membrane reactor (CFFBMR) with an external air membrane separation system.

The reaction rate parameters for the steam and oxidative reforming of methane are given in Table 1. Mathematical Model A schematic diagram of the circulating fast fluidizedbed membrane reactor (CFFBMR) with an external air membrane separation system to provide oxygen sequentially along the height of the reformer is shown in Figure 1. The main difference between this configuration and reactorregenerator system in the literature is that, with this system, the heat integration is achieved in one single unit (i.e., in situ heat integration instead of the circulation of hot solids from the regenerator to the cold reactor). A one-dimensional model has been developed using the following basic assumptions:28 (1) Steady-state isobaric operation is assumed. (2) Plug flow on the reaction and permeation sides, with negligible radial gradients, is assumed. (3) Negligible diffusion limitations are assumed, because of the fine catalyst particles used (the mean diameter is 186 µm). (4) Axial dispersion is assumed to be negligible, because of the high gas velocity. (5) Negligible slip velocity between the gas and solid phases is assumed. (6) Gas-phase oxidation reactions are assumed to be negligible.25 (7) Uniform permeation through the membrane is assumed. (8) The temperature profiles in the reaction side and membrane tubes are assumed to be the same. (9) Ideal gas behavior is assumed. In this study, the reforming system is a six-component system (involving CH4, H2O, H2, CO, CO2, and O2) in which some

components participate in more than one chemical reaction. In the reaction side, the component molar balances are given by

dFj dz

) FcAcL(1 - )

∑i σjiri ( Qj

(14)

where σji is the stoichiometric coefficient of component j in reaction i (negative for reactants, positive for products, zero for inerts). Qj is the permeation rate of component j between the reaction side and the membrane side. Only the hydrogen and oxygen molar balance equations have a Qj term (where j ) H2 or O2). The negative sign denotes hydrogen, and the positive sign denotes oxygen. The energy balance in the reaction side is given by

dT

FcAcL(1 - ) )

∑i (-∆Hi)ri (15)

∑ FjCpj

dz

In the hydrogen permeation side, a differential molar balance on H2 gives

dFHp 2 dz

) Q H2

(16)

and the permeation rate for H2 is given by Shu et al.10 as

QH2 ) 7.21 × 10-5

(

) (

πNH2dH2L δ H2

exp -

)

15700.0 RT

x

( PHr 2 -

xP

p H2)

(17)

and the total hydrogen yield is defined as the total amount of hydrogen produced at the exit of the reformer and the membranes per mole of methane fed:

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hydrogen yield )

(FH2 + FHp 2) - F°H2 F°CH4

Table 2. Data for Mathematical Simulations

(18)

parameter length inside diameter outside diameter catalyst diameterb densitya solid fraction of reformerc feeda feed flow rate feed composition CH4 H2O H2 CO CO2 hydrogen membrane tube outside diameterb number of tubesd flow rate of sweep gas streamd pressure of sweep gas streamd oxygen membrane tube outside diameterd thicknessd

On the oxygen permeation side, a differential molar balance on O2 gives

dFOp 2

) -QO2

dz

(19)

and the permeation rate for O2 is given by Tsai23 as -6

QO2 ) 2.642 × 10 T

( ) ( πNO2dO2L δO2

)( )

POp 2 62700.0 exp ln r RT PO2

(20)

To determine the staging distribution of oxygen along the length of the reformer, a discrete distributed flow of O2 is represented mathematically as

{

}

qm z ) zm 0 z * zm

QO2 )

(21)

in which qm is the flow of injected oxygen at the mth injection point located at height zm. Hydrodynamics of the Fast Fluidized Bed The Kunii and Levenspiel30 method is used to check the hydrodynamic regime at which the reactor is operating. Fast fluidization is only practical for very fine particles and at very high superficial gas velocities (uo > 20ut). The terminal velocity (ut) is given by

ut ) u/t

[

]

µg(Fc - Fg)g Fg2

1/3

(22)

and the dimensionless terminal velocity for any shaped catalyst particle of sphericity φ is given by

u/t )

[

]

2.335 - 1.744φ 18 + (d/c )2 xd/ c

-1

(23)

and the dimensionless catalyst particle diameter is given by

d/c ) dc

[

]

Fg(Fc - Fg)g µg2

1/3

(24)

and dimensionless gas velocity is given by

[

u* ) uo

Fg2

µg(Fc - Fg)g

]

value

reformera

1/3

(25)

To ensure that the reactor is operating in the fast fluidization regime, the dimensionless catalyst particle diameter and dimensionless gas velocity are used to check the hydrodynamics regime of the reactor in the generalized map of gas-solid contacting that has been given by Kunii and Levenspiel.30 Results and Discussion Basic input data for simulations are given in Table 2. The simulation is started through the use of electrolessly deposited

a

al.17 al.28

Data based on Elnashaie and Elshishini.6 c Data based on Kunii and Levenspiel.31

2.0 m 0.0987 m 0.1154 m 186 µm 2270 kg/m3 0.2 19.228 kmol/h 20.56 mol % 71.96 mol % 5.00 mol % 0.00 mol % 2.48 mol % 0.00978 m 20 1.000 kmol/h 101.325 kPa 0.00489 m 50 µm b d

Data based on Adris et Data based on Chen et

Pd-Ag/porous stainless steel membranes for the permeation of hydrogen with a relatively thin thickness of 20 µm, as given by the experimental results of by Shu et al.10 The stainless steel substrates were known to have good mechanical and/or thermal properties. The oxygen permeation rate through dense perovskite membranes is known to be low, because it is limited by surface dissociation and diffusion rates.23 One of the highest permeation fluxes of oxygen reported in the literature was obtained by a membrane that had a thickness of 150 µm.28 In this study, the thickness of oxygen-dense perovskite tubes is chosen to be relatively thin (50 µm), to boost the oxygen permeation flux.28 Figures 2a and 2b compare the methane conversion and hydrogen yield obtained by an adiabatic CFFBMR at a feed temperature of 700 °C, a reaction side pressure of 5 bar, and a steam-to-carbon ration of 3.5 for the following cases: (a) without hydrogen removal and oxygen addition; (b) with 20 hydrogen membrane tubes only (no oxygen addition); (c) with 20 hydrogen membrane tubes and 100 oxygen membrane tubes (O/M ) 0.278); (d) without hydrogen removal and with direct oxygen feed (O/M ) 0.278); and (e) with 20 hydrogen membrane tubes and direct oxygen feed (O/M ) 0.278). For case (a), the conversion of methane is 0.2273 and the yield of hydrogen is 0.8813. The conversion is severely limited by the thermodynamic equilibrium, and the reactor performance is not satisfactory: only a small part of the reactor is utilized. The use of membrane tubes for hydrogen and oxygen clearly gives a small increase in the conversion of methane and the yield of hydrogen, as shown by cases (b) and (c), and this may be due to the low driving force and limited membrane areas used. It can be observed that the direct introduction of oxygen for cases (d) and (e) achieves substantial improvement in methane conversion and hydrogen yield. However, the heat generated by the oxidative reactions at the inlet of the reactor is much larger than the heat consumed by the endothermic reactions, leading to a hot-spot temperature at the entrance of the reactor, as shown in Figure 2c. To enhance the permeation rate of hydrogen further, a high flux composite membrane that consists of a very thin (5 µm) continuous layer of palladium-silver alloy on the outer surface


electroless plating have been successfully applied to deposit very thin palladium films (2-10 µm) on mechanically stable supports. The details of the membrane have been given in Gobina et al.13 Because the hydrogen permeation flux is inversely proportional to the membrane thickness, it is possible to have high fluxes with low cost, through the use of palladium. Also, a pressure of 20 bar has been applied, which will improve the permeation driving force substantially. These operating conditions of the composite hydrogen membranes and a high pressure of 20 bar are used throughout the remainder of this study. Figure 3 shows that substantial improvement in the performance of the CFFBMR is achieved by reducing the thickness of the hydrogen membrane tubes and increasing the pressure. The improvements in the final conversion of methane and the hydrogen yield are 16.7% and 30.2%, respectively, compared to those of a reactor working at a pressure of 5 bar and a hydrogen membrane thickness of 20 µm, respectively. The exit temperature decreases by 19.5%. It can be observed that the reformer almost approaches the maximum theoretical value for a hydrogen yield of 3.2. However, the reactor still has a hotspot region at the entrance of the reactor. Simulated profiles of the methane conversion and the hydrogen yield of the staging distribution of oxygen along the length of the CFFBMR, in comparison with direct oxygen feed (co-feed) are shown in Figures 4a and 4b, respectively. The same total amount of oxygen is fed for all cases. Respectively, 5 and 50 equispaced injection points along the CFFBMR are considered for uniform distributions of the oxygen feed. In the case of the distributed feed, ∼9% (0.1 kg-mol/h) of total oxygen is directly fed to the CFFBMR at the entrance to provide a good reaction basis for the first stage, and the remaining oxygen is distributed equally among all injection points. As shown in Figures 4a and 4b, the distributed addition of oxygen gives almost the same exit methane conversion and hydrogen yield as those obtained using the co-feed mode of operation. Figure 4c shows the corresponding temperature profiles. It can be observed that the distributed addition of oxygen releases an adequate amount of heat and the hot-spot temperature at the inlet of the reactor is totally eliminated, which implies that a commercial-scale CFFBMR could be possible. It is obvious that the use of 50 oxygen injection points gives steady smooth profiles along the length of the CFFBMR. The large number of injection points apparently has little effect on the exit methane conversion, hydrogen yield, and temperature. On the other hand, it may substantially mitigate the hot-spot conditions along the length of the reactor. The situation of using a large amount of oxygen injection points can be considered as the limiting case for an efficient high flux membrane system for oxygen permeation. Unfortunately, such a high-oxygen-flux membrane is not yet available. Figure 5 shows that the exit methane conversion, hydrogen yield, and temperature vary as a function of the O/M ratio for two feed temperatures (Tf ) 300 and 700 °C) and two types of oxygen injection distributions (namely, equispaced and orthogonal distribution). The orthogonal locations are zeros of Jacobi polynomials:33 N

Figure 2. Comparison of the cases without and with hydrogen and oxygen membrane tubes and direct oxygen feed (co-feed) under adiabatic conditions: (a) methane conversion, (b) hydrogen yield, and (c) reactor temperature.

of a thermostable mesoporous support has been used for the permeation of hydrogen. Techniques such as magnetron sputtering, chemical vapor deposition, and solvated metal atom and

P(R,β) N (x) )

(-1)N-i γixi ∑ i)0

(26)

where

γi )

()

N Γ(N + i + R + β + 1)Γ(β + 1) i Γ(N + R + β + 1)Γ(i + β + 1)

(27)

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Figure 3. Effect of the hydrogen membrane thickness and pressure on (a) methane conversion, (b) hydrogen yield, and (c) reactor temperature for the case with hydrogen membrane tubes and direct oxygen feed (co-feed) under adiabatic conditions.

Figure 4. Comparison of the case with hydrogen membrane tubes and direct oxygen feed (co-feed) with hydrogen membrane tubes and distributed oxygen feed along the reformer length under adiabatic conditions: (a) methane conversion, (b) hydrogen yield, and (c) reactor temperature.


and satisfy the orthogonality property

∫01 xβ(1 - x)RPj(x)PN(x) dx ) 0

(for j ) 0, 1, ..., N - 1) (28)

The roots of the Jacobi polynomials are evaluated by the subroutine JCOBI.33 Among the elegant attractive features of the zeros of the Jacobi polynomials are as follows: (a) The roots inside the interval of orthogonality are real and distinct. (b) They provide the optimum positions of the interior collocation points for the orthogonal collocation method. (c) They are picked automatically and, thus, avoid arbitrary selection by the user. (d) They can be moved inside the interval of orthogonality by adjusting the parameters R and β. (e) They have the potential to be practically applied. Five equispaced and orthogonal injection points are used, and oxygen is distributed equally among all injection points. It is shown that the yield profiles for both feed temperatures present an optimal value for the O/M ratio, and this optimum corresponds to the maximum exit hydrogen yield almost at 100% exit methane conversion. The maximum decreases as the feed temperature decreases and shifts to higher O/M ratio values. The feed temperature apparently has a pronounced effect on the hydrogen yield and the position of the optimal O/M ratio value. In fact, the reformer performance is dependent on a complex interaction of the feed rate, reaction rate, rate of oxygen flow and its distribution, hydrogen permeation rate, and reactor hydrodynamics. We are interested in conducting further analysis on the effect of amount of oxygen distributed (nonuniform flow distribution) on the performance of the CFFBMR at a low feed temperature (Tf ) 300 °C). The low feed temperature is desirable because it is known that excessive temperatures have destructive effects on the catalyst, as well as the mechanical and chemical stability of the membrane and reactor, and it decreases the operating costs. Another positive factor for reducing the operating cost for the CFFBMR is the in situ heating provided by the oxidative reforming of methane instead of an external furnace. To keep the number of computations reasonable for this preliminary investigation, we have focused on using three injection locations for the oxygen distribution (i.e., four stages of the CFFBMR in the following analyses). The problem that we are addressing is a multivariate interconnected optimization problem. In this preliminary investigation, we considered simple oxygen distribution. The inequality constraint is 3

0e

∑ Ym e 1.0

(29)

m)1

Figure 5. Effect of the feed temperature and methods of distribution of oxygen injection points (equispaced and orthogonal) on (a) methane conversion, (b) hydrogen yield, and (c) reactor temperature.

where Ym is the fraction of injected oxygen at the mth injected point located at height zm. A simple distribution of oxygen fractions at each injection point is shown in Table 3 for cases A-G. Our objective is to determine the set of oxygen fractions that maximizes the total exit hydrogen yield from the CFFBMR. First, three equispaced injection locations with oxygen distributions according to Table 3 are tested. Figure 6 shows that the exit hydrogen yield varies as a function of O/M ratio for cases A-G. It can be observed that the hydrogen yield passes through a maximum for all cases (A-G). The nonuiform oxygen flow distribution seems to have a significant impact on the exit hydrogen yield. The results indicate that the hydrogen yield for cases G, D, and F is higher, compared to the uniform oxygen

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Figure 6. Effect of the oxygen-to-methane feed ratio (O/M) on the exit hydrogen yield for various oxygen flow distribution (cases A-G), using three equispaced injection points.

Figure 8. Effect of the oxygen-to-methane feed ratio (O/M) on the exit hydrogen yield for various oxygen flow distribution (cases A-G), using three orthogonal injection points.

Table 3. Oxygen Flow Distribution System with Three Injection Points Fraction of Oxygen Injected, Yi case

injection point 1

injection point 2

injection point 3

A B C D E F G

0 0 1/3 1/3 1/3 2/3 2/3

2/3 1/3 1/3 2/3 0 1/3 0

1/3 2/3 1/3 0 2/3 0 1/3

Figure 9. Comparison of equispaced injection points to orthogonal injection points for cases C and F of oxygen flow distribution.

Figure 7. Effect of the oxygen-to-methane feed ratio (O/M) on the exit temperature for various oxygen flow distribution (cases A-G), using three equispaced injection points.

feeding (case C). The corresponding exit temperature profiles for cases A-G is shown in Figure 7. It also seems that the nonuiform oxygen flow distribution has a considerable effect on the exit temperature. The exit temperatures for cases G, D, and F clearly are much lower, compared to the exit temperature obtained by the uniform oxygen feeding (case C).

Figure 8 shows the exit hydrogen yield versus the O/M ratio obtained by three orthogonal injection points with oxygen distribution for cases A-G. It can be observed that the hydrogen yield is sensitive to oxygen flow distribution. It is interesting to note that the same trend of the previous results is also presented here; for example, the hydrogen yield of cases D and F is much higher than that obtained by uniform oxygen feeding (case C). In Figure 9, the exit hydrogen yield obtained from oxygen distribution through three equispaced and orthogonal discrete injection points for the case of uniform oxygen feeding (case C) and the case that gives the highest maximum hydrogen yield (case F) is compared as a function of the O/M ratio. It can be observed that, for nonuniform flow of oxygen feeding (case F), the introduction of oxygen through orthogonal injection points may give better exit hydrogen yield, compared to the introduction of oxygen through equispaced injection points, whereas, for a uniform flow of oxygen feeding (case C), the introduction of oxygen through equispaced injection points gives better exit


Figure 10. Effect of the oxygen injection points distribution obtained by different orthogonal polynomials on the exit hydrogen yield for case F.

hydrogen yield, compared to the introduction of oxygen through orthogonal injection points. These interesting results indicate that the oxygen injection locations, as well as oxygen flow distribution, have an important and crucial role, in regard to reactor performance. Figure 10 shows the effect of orthogonal injection points obtained by different orthogonal polynomials on the exit hydrogen yield for case F. The polynomials tested are Jacobi, Legendre, and Chebyshev polynomials of the first and second type. The Jacobi polynomial is clearly superior to the other polynomials in this case. It is interesting that the oxygen nonuniform distribution systems obtained by all polynomials offer superior exit hydrogen yields over equispaced injection locations. Because of the fact that the reformer performance is dependent on a complex large interacting parameter space, this superiority requires further detailed study. We suspect that the nonuniform distribution of oxygen injection points may affect the thermodynamic equilibrium at the first part of the reformer where a drastic temperature decrease occurs. In the literature, it has been well recognized that nonuniform distribution offers better results than uniform distribution in many chemical engineering systems.34,35 It has been our experience that having very uniform patterns would sometimes hide other performance characteristics. Conclusions The coupling of steam and oxidative reforming reactions of methane for the production of hydrogen is investigated by a series of simulations that are conducted in a circulating fast fluidized-bed membrane reactor (CFFBMR). The effects of oxygen feeding through discrete injection points along the length of the reformer, dense perovskite membranes, and the conventional co-feed mode of operation are investigated. In the range of the parameters investigated, the simulation results reveal that very large increases in the conversion of methane and the yield of hydrogen can be realized via equilibrium shifting of steam reforming reactions, using hydrogen-selective membranes and the supplying of oxygen through injection points along the length of the CFFBMR. The results show that the effect of feeding oxygen through the membranes on the reformer performance is not satisfactory, compared to the other methods

of feeding, because of the low oxygen fluxes of the present generation of oxygen-selective membranes. However, the results of using very large numbers of oxygen injection points show that the future potential application of oxygen membranes with high fluxes may be promising as an efficient and practical tool for leveling the temperature and avoiding thermal runaway. The conventional co-feed mode of operation suffers from having a hot-spot temperature at the entrance of the reformer. This problem has been overcome by staging the distribution of oxygen. As shown, among other potential advantages of distributed oxygen feeding along the length of the CFFBMR, including well controllability of oxygen flow and operation of the CFFBMR under conditions that would not lead to thermal runaway (hot spots). The nonuniform oxygen flow distribution is shown to be superior to uniform oxygen flow distribution for both equispaced and orthogonal injection systems. It can be concluded that the method used to locate the injection points is dependent on the region of parameters, as shown by the examples of equispaced and orthogonal distribution methods considered in this study. These interesting results indicate that the oxygen injection locations, as well as oxygen flow distribution, have an important and crucial role on the reformer performance. The results also suggests that appreciable energy savings can be achieved by operating the reformer at low feed temperatures. This is an important result, because it is known that excessive temperatures have destructive effects on the catalysts, as well as the mechanical and chemical stability of the membranes and the reactor. Nomenclature AbbreViations FBMR ) fluidized-bed membrane reactor CFFBMR ) circulating fast fluidized-bed membrane reactor Symbols Ac ) free cross-sectional area of the reactor for catalyst circulation (m2) Ai ) pre-exponential factor of rate coefficient ki Bj ) pre-exponential factor of adsorption constant Kj Cpj ) specific heat of component j (kJ/(kmol K)) dj ) diameter of hydrogen or oxygen membrane tubes, where j ) H2 or O2 (m) dc ) diameter of catalyst particle (m) d/c ) dimensionless diameter of the catalyst particle Ei ) activation energy of reaction i (kJ/mol) Fj ) molar flow rate of component j (kmol/h) g ) gravitational acceleration (m/s2) ∆H ) enthalpy change of reaction or adsorption (kJ/mol) ki ) rate coefficient of reaction i K1 ) equilibrium constant of reaction 1 (kPa2) K2 ) equilibrium constant of reaction 2 K3 ) equilibrium constant of reaction 3 (kPa2) K5 ) equilibrium constant of reaction 5 (kPa2) K6 ) equilibrium constant of reaction 6 (kPa2) Kj ) adsorption constant of component j L ) reactor length (m) Nj ) number of membrane tubes of component j, where j ) H2 or O2 O/M ) oxygen-to-methane feed ratio (kmol/kmol) Pj ) partial pressure of component j (kPa) P ) total pressure (kPa) qm ) flow of injected O2 at the mth injected point located at zm (kmol/h)

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Qj ) permeation rate of component j, where j ) H2 or O2 (kmol/ h) ri ) rate of reaction i (kmol/(kgcat h)) R ) gas constant (kJ/(mol K)) S/M ) steam-to-methane feed ratio (kmol/kmol) T ) temperature (K) Tf ) feed temperature (K) u* ) dimensionless gas velocity uo ) superficial gas velocity (m/s) ut ) terminal velocity (m/s) u/t ) dimensionless terminal velocity Ym ) fraction of injected O2 at the mth injected point located at height zm z ) dimensionless reactor length Greek Letters δj ) thickness of hydrogen or oxygen membranes, where j ) H2 or O2 (µm) ) void fraction µg ) viscosity of gas (kg/(m s)) Fc ) catalyst density(kg/m3) Fg ) gas density (kg/m3) σji ) stoichiometric coefficient of component j in reaction i (negative for reactants, positive for products, zero for inerts) φ ) sphericity of solid particles Superscripts p ) permeation side r ) reaction side Literature Cited (1) Abashar, M. E. E. Modeling and Simulation of Industrial Natural Gas Steam Reformers and Methanators Using the Dusty Gas Model. M.Sc. Thesis, University of Salford, U.K., 1990. (2) Elnashaie, S. S. E. H; Abashar, M. E. E. Steam Reforming and Methanation Effectiveness Factors Using the Dusty Gas Model Under Industrial Conditions. Chem. Eng. Process. 1993, 32, 177. (3) Abashar, M. E. E; Alhumaizi, K. I.; Adris, A. M. Investigation of Methane-Steam Reforming in Fluidized Bed Membrane Reactors. Trans. Inst. Chem. Eng., Part A 2002, 80, 251. (4) Barbieri, G.; Di Maio, F. P. Simulation of the Methane Steam Reforming Process in a Catalytic Pd-Membrane Reactor. Ind. Eng. Chem. Res. 1997, 36, 2121. (5) De Deken, J. C.; Devos, E. F.; Froment, G. F. Steam Reforming of Natural Gas: Intrinsic Kinetic, Diffusional Influences, and Reactor Design. In Chemical Reaction Engineering; Wei, J., Georgakis, C., Eds.; ACS Symposium Series 196; American Chemical Society: Washington, DC, 1982; p 181. (6) Elnashaie, S. S. E. H.; Elshishini, S. S. Modeling, Simulation and Optimization of Industrial Fixed Bed Catalytic Reactors; Gordon and Breach Science Publishers: London, 1993. (7) Kaihu, H.; Hughes, R. The Kinetics of Methane Steam Reforming Over a Ni/-Al2O Catalyst. Chem. Eng. J. 2001, 82, 311. (8) Venkataraman, K.; Wanat, E. C.; Schmidt, L. D. Steam Reforming of Methane and Water-Gas Shift in Catalytic Wall Reactors. AIChE J. 2003, 49, 1277. (9) Levent, M.; Donald, J. G.; El-bousiffi, M. A. Production of Hydrogen-Rich Gases from Steam Reforming of Methane in an Automatic Catalytic Microreactor. Int. J. Hydrogen Energy 2003, 28 (9), 945. (10) Shu, J.; Bernard, P. A. G.; Kaliaguine, S. Methane Steam Reforming in Asymmetric Pd-and Pd-Ag/Porous SS Membrane Reactors. Appl. Catal., A 1994, 119, 305.

(11) Dittmeyer, R.; Hollein, V.; Daub, K. Membrane Reactors for Hydrogen and Dehydrogenation Processes Based on Supported Palladium. J. Mol. Catal. A: Chem. 2001, 173, 135. (12) Hughes, R. Composite Palladium Membranes for Catalytic Membrane Reactors. Membr. Technol. 2001, 131, 9. (13) Gobina, E. N.; Oklany, J. S.; Hughes, R. Elimination of Ammonia from Coal Gasification Streams by Using a Catalytic Membrane Reactor. Ind. Eng. Chem. Res. 1995, 34, 3777. (14) Liang, W.; Hughes, R. The Effect of Diffusion Direction on the Permeation Rate of Hydrogen in Palladium Composite Membranes. Chem. Eng. J. 2005, 112, 81. (15) Elnashaie, S. S. E. H.; Adris, A. A Fluidized Bed Steam Reformer for Methane. In Proceedings of the Fourth International Fluidization Conference, Banff, Canada; AIChE Publication 319; American Institute of Chemical Engineers (AIChE): New York, 1989. (16) Adris, A. M.; Elnashaie, S. S. E. H; Hughes, R. A Fluidized Bed Membrane Reactor for the Steam Reforming of Methane. Can. J. Chem. Eng. 1991, 69, 1061. (17) Adris, A. M.; Lim, C. J.; Grace, J. R. The Fluidized-Bed Membrane Reactor for Steam Methane Reforming: Model Verification and Parametric Study. Chem. Eng. Sci. 1997, 52, 1609. (18) Abashar, M. E. E. Stagewise Fluidized Bed Membrane Reactors for Methane-Steam Reforming. Afinidad 2003, 60 (504), 184. (19) Abashar, M. E. E. Coupling of Steam and Dry Reforming of Methane in Catalytic Fluidized Bed Membrane Reactors. Int. J. Hydrogen Energy 2004, 29, 799. (20) Roy, S.; Pruden, B. B.; Adris, A. M.; Lim, C. J.; Grace, J. R. Fluidized-Bed Steam Methane Reforming with Oxygen Input. Chem. Eng. Sci. 1999, 54, 2095. (21) Grace, J. R.; Li, X.; Lim, C. J. Equilibrium Modeling of Catalytic Steam Reforming of Methane in Membrane Reactor with Oxygen Addition. Catal. Today 2001, 64, 141. (22) Itoh, N.; Govind, R. Combined Oxidation and Dehydrogenation in Palladium Membrane Reactor. Ind. Eng. Chem. Res. 1989, 28, 1554. (23) Tsai, C. Y.; Dixon, A. G.; Moser, W. R.; Ma, Y. H. Dense Perovskite Membrane Reactors for Partial Oxidation of Methane to Syngas. AIChE J. 1997, 43, 2741. (24) Lu, Y.; Dixon, A. G.; Moser, W. R.; Ma, Y. H.; Balachandran, U. Oxygen-Permeable Dense Membrane Reactor for the Oxidative Coupling of Methane. Membr. Sci. J. 2000, 170, 27. (25) Jin, W.; Gu, X.; Li, S.; Huang, P.; Xu, N.; Shi, J. Experimental and Simulation Study on a Catalyst Packed Tubular Dense Membrane Reactor for Partial Oxidation of Methane to Syngas. Chem. Eng. Sci. 2000, 55, 2617. (26) Roy, S. Fluidized Bed Steam Methane Reforming with High-flux Membranes and Oxygen Input, Ph.D. Thesis, University of Calgary, Alberta, Canada, 1998. (27) Prasad, P.; Elnashaie, S. S. E. H. Coupled Steam and Oxidative Reforming for Hydrogen Production in a Novel Membrane Circulating Fluidized-Bed Reformer. Ind. Eng. Chem. Res. 2003, 42, 4715. (28) Chen, Z.; Yan, Y.; Elnashaie, S. S. E. H. Novel Circulating Fast Fluidized-Bed Membrane Reformer for Efficient Production of Hydrogen from Steam Reforming of Methane. Chem. Eng. Sci. 2003, 58, 4335. (29) Brereton, C. M. H.; Grace, J. R. Microstructural Aspects of the Behavior of Circulating Fluidized Beds. Chem. Eng. Sci. 1993, 48, 2565. (30) Kunii, D.; Levenspiel, O. Fluidization Engineering, 2nd Edition; Butterworth-Heinemann: Stoneham, MA, 1991. (31) Kunii, D.; Levenspiel, O. Circulating Fluidized Beds. Chem. Eng. Sci. 1997, 52, 2471. (32) Xu, J.; Froment, G. F. Methane Steam Reforming, Methanation and Water-Gas Shift. I. Intrinsic Kinetics. AIChE J. 1989, 35 (1), 88. (33) Villadsen, J.; Michelsen, M. L. Solution of Differential Equation Models by Polynomial Approximation; Prentice Hall: Englewood Cliffs, NJ, 1978. (34) Morbidelli, M.; Servida, A.; Varma, A. Optimal Catalyst Activity Profiles in Pellets. 1. The Case of Negligible External Mass Transfer Resistance. Ind. Eng. Chem. Fundam. 1982, 21, 278. (35) Huss, R. S. Collocation Methods for Flexible Distillation Design, Ph.D. Thesis, Carnegie Mellon University, Pittsburgh, PA, 1995.

ReceiVed for reView December 8, 2006 ReVised manuscript receiVed May 21, 2007 Accepted June 7, 2007 IE061579Y

Staging Distribution of Oxygen in Circulating Fast ... - ACS Publications

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