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Thermal resistance effect on methanol-steam reforming performance in micro-scale reformers
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
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Thermal resistance effect on methanol-steam reforming performance in micro-scale reformers Rei-Yu Chein a,*, Yen-Cho Chen b, J.N. Chung c a
Department of Mechanical Engineering, National Chung-Hsing University, Taichung City 402, Taiwan Department of Energy and Resource, National United University, Miaoli City 360, Taiwan c Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611-6300, USA b
article info
abstract
Article history:
We numerically investigate hydrogen production based on methanol-steam reforming
Received 13 April 2011
(MSR) using a micro-scale cylindrical packed bed reformer. The reformer wall is included in
Received in revised form
the physical model. The heat required for the reforming reaction is supplied either inter-
12 September 2011
nally using a heating rod placed along the center of the reformer or externally by a heat flux
Accepted 15 September 2011
applied at the reformer outer wall. Our results show that the thermal resistance from the
Available online 14 October 2011
heat source to the reformer environment plays an important role in the reformer performance. This thermal resistance depends on the reformer geometry, wall material and heat
Keywords:
transfer coefficients inside the catalyst bed and outside the reformer. Based on our
Methanol-steam reforming (MSR)
numerical results, it is suggested that better methanol conversion and hydrogen yield can
Packed-bed reformer
be obtained using reformer wall material with low thermal conductivity and thin thick-
Thermal resistance
ness. For both internal and external heating under the same heat rate supply, no significant
Internal and external heating
difference in reformer performance was found.
Water gas shift (WGS) reaction
A water gas shift (WGS) reaction model was included in the present numerical model. In the reformer low-temperature zone the forward WGS reaction was clearly demonstrated, resulting in a decrease in carbon monoxide (CO) selectivity. In the high temperature zone the backward WGS reaction was also clearly demonstrated in which CO selectivity increases with the increase in temperature. For both internal and external heating under the same heat rate supply, our results indicated that CO selectivity is about thirty times lower when the WGS reaction is neglected. Copyright ª 2011, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved.
1.
Introduction
The miniature fuel cell is efficient in converting chemical energy directly into electrical energy. It is a prominent alternative to batteries [1e3]. Among fuel cell designs the direct methanol fuel cell (DMFC) and polymer electrolyte fuel cell (PEMFC) have received much attention due to their low operating temperature which is suitable for portable applications.
The PEMFC has a much higher power density than the DMFC, but hydrogen storage is impractical for portable applications. A fuel reformer that converts liquid fuel into hydrogen is the possible solution. Among the fuels used for producing hydrogen, methanol has been widely adopted because of its advantages such as high hydrogen to carbon ratio, liquid form at ambient conditions, low reforming temperature (200e350 C) and low CO formation [4,5]. Methanol-steam
* Corresponding author. E-mail address: [email protected] (R.-Y. Chein). 0360-3199/$ e see front matter Copyright ª 2011, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2011.09.070
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
Nomenclature A1 B1 A2 CD CF cp CR CWGS Dij DTi dp ER ED EWGS hN K Keq kD kR kWGS Lb mcat Mi mi NG n_ i p Q qc
constant in reforming reaction rate, 1.15 106 m3 s1 kg1 constant in reforming reaction rate, 9.41 105 m3 s1 kg1 constant in decomposition reaction rate, 7.09 107 mol s1 kg1 correction factor decomposition reaction rate, 5.5 Forchheimer drag coefficient gas specific heat, J kg1 K1 correction factor for reforming reaction rate, 5.5 correction factor for WGS reaction rate, 11.2 binary molecular diffusion coefficient, m2 s1 thermal diffusion coefficient of species i, m2 s1 catalyst particle diameter, m activation energy for reforming reaction, 84,100 J mol1 activation energy for decomposition reaction, 111,200 J mol1 activation energy for WGS reaction, 70,000 J mol1 ambient heat transfer coefficient, W m2 K1 catalyst bed permeability, m2 equilibrium constant in WGS reaction rate constant of decomposition reaction, mol s1 kg1 rate constant of reforming reaction, m3 s1 kg1 rate constant of WGS reaction, mol m3 s1 Ka Pa2 reformer length, m catalyst weight, kg molecular weight of species i, g mol1 mass fraction of species i number of species in the gas mixture molar flow rate of species i pressure, Pa volumetric flow rate, m3 s1 energy source term due to the chemical reaction, J m3
reforming (MSR) for hydrogen production was reviewed recently by Holladay et al. [6,7]. The MSR reformer catalyst particles can be packed in a bed to form a packed-bed reformer [8,9] or coated onto the channel surface to form a plate reformer [10,11]. For portable power supply requirements, recent developments have focused on minimizing the size and weight of the reformer [12,13]. Due to its endothermic nature the MSR requires a heat supply from external sources. Because the chemical reactions involved in the MSR (reforming, decomposition, and water gas shift (WGS)) depend greatly on the temperature, understanding the temperature distribution and heat transfer within the reformer is essential for optimizing reformer design and operating conditions. For a packed bed reformer heated from an external source, several studies have shown that the temperature near the core of the reformer is lower than that near the wall [14e16]. The lower temperature leads to less catalyst activity and produces undesired chemical species. To overcome the adverse lower temperature in the reformer core
qi q00i q00w R Rb Ri ri rR rD rWGS SCO T tr / V Vb xi YH2
251
heat rate from heating rod, W heat flux at heating rod surface, W m2 heat flux at reformer outer wall, W m2 universal gas constant, 8.314 J mol1 K1 reformer radius, m heating rod radius, m molar generation rate of species i, mol m3 s1 reforming reaction rate, mol m3 s1 decomposition reaction rate, mol m3 s1 WGS reaction rate, mol m3 s1 CO selectivity temperature, K reformer wall thickness, m velocity, m s1 catalyst bed volume, m3 molar fraction of species i hydrogen yield
Greek symbols a temperature exponent d molar steam/CO ratio heat of decomposition reaction, J mol1 DHD DHR heat of reforming reaction, J mol1 DHWGS heat of reforming reaction, J mol1 ε catalyst layer porosity h methanol conversion l thermal conductivity, W m1 K1 m viscosity, kg m1 s1 f molar ratio of water to methanol r density, kg m3 Subscript c h m s N
reformer wall internal heating rod gas mixture catalyst ambient condition
region, Suh et al. [17] proposed an internal heating structure with heat supplied from a heating element suspended along the center of the catalyst bed. Their experimental and numerical results showed that under the same wall temperature, inlet steam-methanol flow rate and catalyst mass conditions the methanol conversion was higher for an internally heated reformer compared the externally heated reformer. Enhancing the heat transfer inside the catalyst bed can also lead to improved methanol conversion [18e20]. The heat supply to the reformer may be achieved in various ways [21e23]. Recent studies have used a catalytic combustor as the reformer heat source [24e29]. In contrast to the reformer, the catalytic combustion occurring in the combustor is an exothermic reaction. For catalytic combustion in the micro-scale dimension, it has been shown that the combustor wall material and heat loss to the environment play important roles in the combustion characteristics such as flame stability and exhaust gas temperature [30e33]. Although the reforming reaction is endothermic, the reformer wall
252
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material and external heat loss will also affect the reformer performance and need to be identified. Due to its small size experimental measurements on the detailed temperature distribution inside the reformer is difficult. For this reason, modeling and simulation have been used extensively to obtain a better understanding of the fundamental transport phenomena inside the reformer, to reduce research cost, and to shorten the design cycles [34]. Many numerical simulations have been carried out for both packed bed and plate reformers [35e39]. In most of these studies, the reformer outer wall with elevated temperature was usually treated as the heat source for the chemical reaction. In the laboratory, an elevated reformer wall temperature can be achieved by putting the reformer inside an oven. However, for the cases with heat supplied from an integrated combustor or from an electric heater placed internally or externally, a heat flux is actually applied to the reformer instead of elevated temperature [40,41]. In terms of theoretical modeling, it should be more appropriate to use wall heat flux boundary condition instead of wall temperature boundary condition [42]. In this study, we numerically examine the performance of a packed bed reactor based on the design proposed by Suh et al. [17]. In their model, cylindrical packed bed MSR reformer performance with both internal and external heating was addressed. The external heating was achieved by assigning an elevated temperature to the infinite-thin reformer wall. Extending this model, we study the reformer made by various materials and has finite wall thickness. At the reformer wall outer surface, various boundary conditions are applied to emulate the heat transfer conditions such as insulated, natural or forced convective heat loss and external heat supply. The impact of WGS reaction on the CO selectivity in MSR is also discussed.
r, v
Catalyst Bed
CH3 OH+H2 O mixture
Theoretical model and formulation
2.1.
Physical model
Fig. 1(a) shows the physical model described in the work of Suh et al. [17]. The cylindrical packed-bed reformer has the length Lb ¼ 12 mm and diameter db ¼ 2Rb ¼ 1.5 mm. A chromium alloy rod with length equal to the reformer length and diameter di ¼ 2Ri ¼ 0.33 mm was placed along the reformer centerline for use as the internal heat source. The heat flux generated by the heating rod is q00i ¼ qi =pdi Lb with qi being the heat rate generated from the electric current passing through the heating rod. CuO/ZnO/Al2O3 pellets were packed in the space between the reformer and heating rod for use as the catalyst for the reforming reaction. A methanol-steam mixture with prescribed temperature Ti, volumetric flow rate Qi and steam to methanol molar ratio f was introduced into the reformer from the inlet (z ¼ 0). At the reformer outer wall (r ¼ Rb) temperature Tw with value higher than Ti served as the external heat source. In the model shown in Fig. 1(a), the reformer wall thickness was neglected which is not realistic. In this study, the reformer wall with thickness tr was included as shown in Fig. 1(b). The reformer can be fabricated using various materials and wall thicknesses. At the outer reformer wall various
z, u
Ri
Rb
Heating rod Heat flux qi
ui , Ti Lb
b
r, v
Tw , q
w
tr Catalyst Bed
z, u
Heating rod
CH3 OH+H2 O mixture
Heat flux q i
ui , Ti Reformer wall
h , T Fig. 1 e Physical models of the cylindrical packed bed reformer. (a) model in the study of Suh et al. [17]. (b) model extended from (a) with reformer wall included.
thermal boundary conditions would be possible. As shown in Fig. 1(b), the reformer may be operated in an environment with temperature TN (¼298 K) and heat transfer coefficient hN or subject to an applied heat flux q00w serving as the external heat source.
2.2.
2.
a Tw
Mathematical model
Transport phenomena in the reformer can be described by the conservation equations of mass, momentum, energy and species leading to a set of non-linear partial differential equations. To simplify the analysis, the following assumptions are made: (1) All the species in the gas mixture are ideal gases. (2) The gas flow in the reformer is assumed to be weakly compressible, axisymmetric, steady and laminar. (3) The catalyst particles are assumed to be spherical with a diameter dp and the catalyst bed is treated as a porous medium with homogeneous porosity ε and permeability K. (4) The catalyst bed is in local thermal equilibrium with the surrounding gas mixture. Based on the above assumptions, the governing equations for the mass conservation, fluid flow, energy transport, and species transport can be written as, /
V$ðεrV Þ ¼ 0
(1a)
/ / 4 . T 1 mm / 2mm 4 . VV I V$ðrV V Þ ¼ V$ pI þ þ ðVV Þ V$V ε 3 ε2 mm / rCF / / V pffiffiffiffi jV jV (1b) K K
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2 / V$ εrcp V T ¼ V$ðle VTÞ þ qc
8 9
= NG < / X Vp VT Dij Vxi þ ðxi mi Þ V$ εrV mi rmi DTi ¼ ri : p T ; j¼1
(2)
(3)
Eq. (1) is known as the BrinkmaneDarcyeForchheimer model for the fluid flow in a porous medium with a homogeneous porosity. r is the mass-weighted density defined as, G p X xi Mi RT i¼1
N
r¼
(4)
The permeability K and Forchheimer drag coefficient CF for a packed bed with spherical particles can be written as [43], pffiffiffiffiffiffiffiffi K ¼ d2p ε3 = 150ð1 εÞ2 ; CF ¼ 1:75= 150ε3=2
(5)
In Eq. (2) for the energy transport, cp is the mass-weighted specific heat defined as,
cp ¼
NG X
mi cpi
(6)
i¼1
le is the effective thermal conductivity of the catalyst bed defined as, le ¼ εlm þ ð1 εÞls
(7)
where ls and lm are the thermal conductivities of catalyst particle and gas mixture, respectively. qc is the energy source due to chemical reaction. Eq. (3) is known as the MaxwelleStefan species transport equation. Dij and DTi are the binary molecular diffusion coefficient and thermal diffusion coefficient of the species i in the gas mixture, respectively. ri is the production rate of species i due to chemical reaction. In these equations, the gas mixture transport properties (mm, lm, Dij and DTi ) can be evaluated based on the ChapmaneEnskog theory [44]. Since there are no flows in the reformer wall and internal heating rod, their temperature distributions are simply governed by the heat conduction, Reformer wall : V$ðkc VTc Þ ¼ 0
(8)
Heating rod : V$ðkh VTh Þ ¼ 0
(9)
All governing equations were written in a cylindrical coordinate system (r, q, z). Utilizing the symmetry in the q coordinate we can recast the problem as an axisymmetric twodimensional model (r, z). The fluid velocity components in the r and z-directions for the reactant flow are u and v, respectively.
2.3.
253
Chemical reaction model
Using CuO/ZnO/Al2O3 as the catalyst, the chemical reactions taking place during the MSR are [45,46], Steam reforming : CH3 OH þ H2 O/3H2 þ CO2
(10)
Decomposition : CH3 OH/2H2 þ CO
(11)
Water gas shift ðWGSÞ: CO þ H2 O/H2 CO2
(12)
For the packed-bed reformer, Suh et al. [35] proposed semiempirical kinetic models for the steam reforming and decomposition reactions as, rR ¼ ð1 εÞrs kR CCH3 OH
(13)
rD ¼ ð1 εÞrs kD
(14)
where CCH3 OH is the molar concentration of methanol. In Eqs. (13) and (14), rs, kR and kD are catalyst density, and rate constants for reforming and decomposition reactions defined as, rs ¼
mcat mcat ¼ 2 Vb p Rb R2i Lb
(15)
kR ¼ CR ½A1 þ B1 lnfeER =RT
(16)
kD ¼ CD A2 expð ED =RTÞ
(17)
In Eq. (15), mcat is the catalyst weight and Vb is the catalyst bed volume. In Eqs. (16) and (17), A1, A2 and B1 are constants given in the study by Amphlett et al. [9]. ER and ED are the activation energies for the reforming and decomposition reactions, respectively. R and T are the universal gas constant and gas mixture temperature, respectively. CR and CD are the correction factors for the reforming and decomposition reactions accounting for the catalyst activity and effectiveness and their values are chosen as 5.5 [17]. Based on the studies by Purnama et al. [47] and Chen et al. [48], the chemical reaction model for the WGS reaction can be written as, rWGS ¼ CWGS kWGS pCO pH2 O pCO2 pH2 =Keq
(18)
EWGS kWGS ¼ 1:78 1022 1 0:154d þ 0:008d2 Ta exp RT
(19)
4577:8 Keq ¼ exp 4:33 T
(20)
where kWGS and Keq are the rate constant and the equilibrium constant of the WGS reaction, respectively, pi (i ¼ CO,
Table 1 e Heat of reaction. Reforming Decomposition Water gas shifting
DHR ¼ 4:95 104 þ ðcpCO2 þ 3cpH2 cpCH3 OH cpH2 O ÞðT 298Þ DHD ¼ 9:07 104 þ ðcpCO þ 2cpH2 cpCH3 OH ÞðT 298Þ DHWGS ¼ 4:912 104 þ ðcpCO2 þ cpH2 cpH2 O cpCO ÞðT 298Þ
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H2O, CO2, H2) is the partial pressure, a is a constant determined by the experiments, and d is the molar ratio of water to CO. Based on the study by Chen et al. [48], a and activation energy EWGS are chosen as 8.0 and 70 kJ/mol, respectively. Similar to the models for reforming and decomposition proposed by Suh et al. [35], we introduced a correction factor CWGS for the WGS reaction to account for the catalyst activity and effectiveness. Based on the experimental data reported by Suh et al. [17], CWGS is taken as 11.2 in the present study. From the reaction models described above, the production rate of each species and energy source can be written as, rCH3 OH ¼ rR rD
rCO2 ¼ rR þ rWGS
(21c)
rCO ¼ rD rWGS
(21d)
rH2 ¼ 3rR þ 2rD þ rWGS
(21e)
qc ¼ rR DHR rD DHD þ rWGS DHWGS
(21f)
where DHR, DHD and DHWGS are the heat of reaction for the reforming, decomposition and WGS reactions, respectively. Temperature-dependent heats of reaction of these reactions are listed in Table 1 [28].
Boundary conditions
Boundary conditions must be specified to complete the mathematical model. Referring to Fig. 1, the boundary conditions are specified as follows,
(22a)
(2) reformer outlet (z ¼ Lb, Ri < r < Rb) /
vV vT vmi ¼ ¼0 ¼ vz vz vz
(22b)
(3) walls of reformer and internal heating rod at inlet and outlet (z ¼ 0, z ¼ Lb, 0 < r < Ri and Rb < r < Rb þ tr) vTh vTc ¼ ¼0 vz vz
(22c)
(4) internal heating rod outer surface (0 < z < Lb, r ¼ Ri) /
V ¼ 0; lh
vTh vmi ¼ q00i ; ¼0 vr vr
(22d)
vT vTc vmi ; ¼0 ¼ lc vr vr vr
(22e)
(6) along reformer centerline (0 < z
(22f)
(7) reformer outer wall (0 < z < Lb, r ¼ Rb þ tr) vTc vTc ¼ hN ðTc TN Þ or lc ¼ q00w vr vr
(22g)
In Eq. (22a), ui ¼ Qi =ðpðR2b R2i ÞÞ and mCH3 OH;in are the inlet velocity and the mass fraction of CH3OH, respectively. At the reformer outlet, the Neumann boundary conditions are specified for the flow velocity, temperature and species concentrations as indicated in Eq. (22b). At the inlet and outlet, the walls of the reformer and internal heating rod are assumed to be insulated as described in Eq. (22c). In Eq. (22d), a constant heat flux q00i generated by the heating rod is specified while the gas flow satisfies the no-slip boundary condition and no species deposition occurs on the heating rod surface. Referring to Fig. 1(b), the conditions for no-slip gas flow, continuous heat transfer and no species deposition at the reformer inner wall are specified as indicated in Eq. (22e). The axisymmetric condition is specified along the reformer centerline as indicated in Eq. (22f). Eq. (22g) shows that various boundary conditions including an elevated wall temperature Tw, convective heat loss to the environment with heat transfer coefficient hN and temperature TN, and applied heat flux q00w for external heating can be specified at the reformer outer wall.
3.
(1) reformer inlet (z ¼ 0, Ri < r < Rb)
mH2 ¼ 0; mCO2 ¼ 0; and mCO ¼ 0
V ¼ 0; T ¼ Tc ; le
Tc ¼ Tw ; lc (21b)
u ¼ ui ; v ¼ 0;T ¼ Ti ; mCH3 OH ¼ mCH3 OH;in ; mH2 O ¼ fmCH3 OH;in ;
/
(21a)
rH2 O ¼ rR rWGS
2.4.
(5) reformer inner wall (0 < z
Numerical methods
All of the governing equations along with the boundary conditions were solved simultaneously using COMSOL Multiphysics (Comsol Inc., version 4.01). Weakly compressible NaviereStokes, general heat transfer and MaxwelleStefan diffusion and convection modules were applied for solving the velocity, temperature and species concentration distributions in the reformer. Because the numerical solution accuracy strongly depends on the mesh size, a refined mesh is necessary in the region where the dependent variable gradients are pronounced. Finer meshes were used to capture the subtle
Table 2 e Thermophysical property of the wall materials. Material
Chromium alloy Silica glass Steel (AISI4340) Copper
Thermal Specific Thermal Density heat capacity [kg/m3] conductivity [W/mK] [kJ/kgK] [kJ/m3K] 8400 2203 7850 8700
11.3 1.38 44.5 400
450 703 475 385
3.78 1.54 3.72 3.35
106 106 106 106
255
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
1
3000
a
b Analysis [17] experiment [17] present study
0.8
Analysis [17] experiment [17] present model
2500
2000
C O (ppm)
0.6
0.4
1500
1000 0.2
500
0 160
170
180 o
Tw ( C)
190
0 160
200
280
d
Ti=120 C Tw=180 oC
10
240
180
0.4
160
qi (W) 0 0.25 0.5 0.75
140 120
0
2
qi=0.25 W
xC O (ppm)
x M eO H
T( o C )
200
0.5 W
0.75 W
3
102
0.5
220
101 100
10-1
Ti=120 oC o Tw=180 C
10-2
4
z/R
6
8
190
104
0.6
c
180
Tw (o C)
o
260
170
0.3
10
-3
0
2
b
4
z/R b
6
8
Fig. 2 e Verification of the present numerical model by comparing the results reported in the study by Suh et al. [17]. Qi [ 900 ml/h, mcat [ 16 mg (a) methanol conversion, (b) CO concentration, (c) gas mixture temperature and methanol molar fraction distributions along the heating rod surface (r [ Ri) with various heat rates. mcat [ 16 mg, Ti [ 120 C, and Tw [ 180 C. (d) CO concentrations for the cases with and without WGS reaction based on the operation conditions used in (c).
changes in velocity, temperature and species concentration in the inlet, outlet, and fluidewall interface regions. The solution independence on the mesh size was carefully studied before reporting the final results. The numerical experiments show that the solution becomes mesh-independent when the element number exceeds approximately 4000. Hence, more than 4000 meshes were used for the results presented in this study. Important characteristics for MSR are methanol conversion, hydrogen yield and CO selectivity. Methanol conversion is defined as the ratio between converted methanol at a position along the reformer and the inlet methanol flux [32],
h¼
n_ CH3 OH;in n_ CH3 OH n_ CH3 OH;in
(23)
Hydrogen yield characterizes the reformer performance with respect to the hydrogen production. It is the ratio of the produced hydrogen to the theoretical maximum amount of hydrogen [32], n_ H3 (24) YH2 ¼ n_ CH3 OH;in CO selectivity defines the molar fraction between produced CO and the molar fraction of all carbon-containing gas phase species which are CO and carbon dioxide (CO2) in MSR. It is an
256
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
indicator for the effectiveness of carbon monoxide production and is defined as [32], SCO ¼
n_ CO n_ CO þ n_ CO2
4.
Results and discussion
4.1.
Numerical model verification
(25)
The numerical model used in this study was verified by comparing the results computed from the present numerical model with those reported by Suh et al. [17] under the same reformer geometry shown in Fig. 1(a). The inlet flow rate and catalyst mass are 900 ml/h and 16 mg, respectively. For the ratio of steam to methanol f ¼ 1.1, the inlet mass fractions of methanol and water are mCH3 OH;in ¼ 0:62 andmH2 O;in ¼ 0:38, respectively. The internal heating rod is made of chromium alloy with properties listed in Table 2. The external heating was provided by an elevated wall temperature Tw. In Fig. 2(a), the methanol conversion under various qi, Ti and Tw predicted by the present numerical model are compared with the experimental and numerical results reported by Suh et al. [17]. Good agreement between the numerical and the experimental results is obtained. In Fig. 2(b), comparison of CO concentration obtained by the present model with those reported by Suh et al. [17] is presented. It is seen that the model used in the present study under predicts the CO production. The discrepancy may be due to both WGS model accuracy and experimental uncertainty. Although the WGS reaction model used in the present study underestimates the CO concentration compared with the experimental measurement, it gives the correct variation trend. With the inclusion of WGS reaction, the CO concentration predicted by the present model is closer to the experimental data compared with the numerical results predicted by Suh et al. [17] in which the WGS reaction was neglected. For various heat rates and Ti ¼ 120 C, temperature and methanol molar fraction variations along the heating rod surface (r ¼ Ri) predicted by the present numerical model are shown in Fig. 2(c). Our model successfully reproduced the results in Fig. 8 in the study by Suh et al. [17]. To understand the WGS reaction effect, the CO concentrations for the cases with and without WGS reaction are compared in Fig. 2(d) under the same operating conditions as the results shown in Fig. 2(c). In the entrance zone at which the temperature is low, the CO concentration with WGS reaction has lower value than that without WGS reaction. This is because of the forward WGS reaction that converts CO into CO2. However as the temperature increases in the reformer downstream, the CO value with WGS reaction becomes higher than that without WGS reaction. Because the WGS reaction is a moderately exothermic reaction, the reaction tends to shift to the left side at high temperature according to the Le Chatelier’s principle. The reversed WGS reaction leads to lower CO conversion and hydrogen yield [47,49,50].
4.2.
reformer wall included in the physical domain as shown in Fig. 1(b). In order to focus on the heat transfer effect on the MSR performance, the inlet flow rate and temperature of the methanol-water mixture are fixed as Qi ¼ 900 ml/h and Ti ¼ 120 C. The catalyst mass is also fixed as 16 mg. We first examine the external heat loss effect on the MSR performance for a silica glass reformer. The silica glass property used in this study is listed in Table 2. The reformer wall thickness is taken as 0.165 mm which is the same as radius of the internal heating rod. At the outer reformer wall, various heat transfer coefficients are used to examine the heat loss effect on the MSR performance. Using the ambient heat transfer coefficient hN in the range of 0 (insulated wall) to 50 W/m2 K and
MSR performance with reformer wall
Based on the comparison described above, our numerical model can be applied to examine the MSR performance with
Fig. 3 e Typical temperature and species molar fraction distributions in the silica glass reformer. Ti [ 120 C, qi [ 0.25 W, tr [ 0.165 mm, and hN [ 5 W/m2 K.
257
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
0.6
300
a
o
Ti=120 C qi=0.25 W
b
with WGS without WGS
0.5
tr=0.165 mm 250
insulated
with WGS without WGS
o
Ti=120 C qi=0.25 W tr=0.165 mm
0.4
5
ToC
0.3
200
h =25 W/m2 K
insulated 5
0.2
150
0.1 2
h =25 W K
50 0 100
0
1
2
3
4
z/R b
5
6
7
0
8
1
1
2
3
4
z/R b
5
6
7
8
10-1 o
c
0.9
Ti=120 C qi=0.25 W tr=0.165 mm
0.8
d
with WGS without WGS
o
Ti=120 C qi=0.25 W tr=0.165 mm
10-2
0.7
insulated 5 insulated
insulated
0.6
5 25
0.5
S CO
Y H2
50
5
0.4
10-3
50 25
0.3 10-4
0.2
2
h =50 W/m K
50
0.1
h =25 W/m2 K
with WGS without wGS
0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
1
2
3
4
z/R b
5
6
7
8
Fig. 4 e Ambient heat transfer coefficient effect on the MSR performance in silica glass reformer. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
qi ¼ 0.25 W, the MSR performance are shown in Figs. 3 and 4. In Fig. 3, temperature and species molar fraction distributions for hN ¼ 5 W/m2 K are shown. As the methanol-steam mixture is introduced into the reformer, it is heated by the internal heating rod. At the same time, heat is lost to the environment from the reformer outer wall. The temperature distribution shows that there is a low-temperature zone near the reformer entrance. Temperature increases at the reformer downstream due to the heat supply from the heating rod. Because of the external heat loss, a temperature gradient also exists in the radial direction. Because the MSR depends on temperature, the methanol concentration decreases while the CO and hydrogen concentrations increase along the reformer length.
Similar to Fig. 2(c) and (d), we use temperature, methanol conversion, hydrogen yield and CO selectivity variations along the heating rod surface r ¼ Ri to quantify the reformer performance. In Fig. 4(a), the highest temperature can be obtained for the insulated reformer outer wall case while the temperature decreases with the increase in ambient heat transfer coefficient. This is expected since a higher heat transfer coefficient implies higher heat loss to the environment. Fig. 4(a) also indicates that fully developed temperature cannot be reached when the external heat loss is present. The corresponding methanol conversion, hydrogen yield and CO selectivity are shown in Fig. 4(b), (c) and (d), respectively. Because of the decrease in reaction temperature, methanol
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300
0.6
a
with WGS without WGS
o
Ti=120 C qi=0.25 W 2 h =5 W/m K
b
o
Ti=120 C
0.5
with WGS without WGS
qi=0.25 W
250
h =5 W/m2 K
T (o C )
0.4
0.33 mm
0.3
0.165 mm
200
0.165 mm
0.330 mm
tr=0.495 mm
tr=0.495 mm
0.2
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0.1
0 100
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1
2
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z/R b
5
6
7
8
0
1
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4
z/R b
5
6
7
8
10-1
1
c
0.9
Ti=120 C qi=0.25 W h =5 W/m2 K
0.8
o
d
o
with WGS without WGS
Ti=120 C qi=0.25 W h =5 W/m2 K
10-2
0.7
0.165 mm
0.5 0.4
S CO
Y H2
0.6
0.495 mm
10-3
0.33 mm
0.330 mm
0.3
tr=0.165 mm
tr=0.495 mm 10-4
0.2
with WGS without WGS
0.1 0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
0
1
2
3
4
z/R b
5
6
7
8
Fig. 5 e Wall thickness effect on the silica glass reformer performance. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
conversion, hydrogen yield and CO selectivity decrease with the increase in ambient heat transfer coefficient. As shown in Fig. 4(b) and (c), the methanol conversion and hydrogen yield have approximately the same variation trends. In Fig. 4(d), a decrease in CO selectivity can be observed in the entrance zone due to the forward WGS reaction. As the temperature increases in the reformer downstream, CO selectivity increases due to the reversed WGS reaction. For the case of hN ¼ 50 W/m2 K, the temperature rises slightly at the entrance due to the forward WGS reaction which is an exothermic reaction. At the reformer downstream, the temperature drops significantly due to strong heat loss and weak MSR results due to insufficient heat supply. To understand the WGS reaction effect on MSR performance, results without the WGS reaction are also shown in
Fig. 4. From Fig. 4(a)e(c), the WGS reaction has no significant effect on the temperature, methanol conversion and hydrogen yield, as proven by Pepply et al. and Amphlett et al. [41e43] when gas temperature is low. However, for the CO selectivity results shown in Fig. 4(d), higher CO selectivity results for the case without WGS reaction compared with the case including the WGS in the region near the entrance zone because of the forward WGS reaction. As the temperature increases in the reformer downstream, the reverse situation occurs. That is, higher CO selectivity occurs when the WGS reaction is included. The reason for this result is due to the temperature-dependent WGS reaction discussed earlier. Due to high reaction temperature in the reformer downstream, the reversed WGS reaction is favored. Since the reversed WGS reaction becomes an endothermic reaction, the reaction
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300
0.6
a
with WGS without WGS
h =5 W/m K tr=0.165 mm
250
Ti=120 o C qi=0.25 W
0.5
2
T (o C )
b
o
Ti=120 C qi=0.25 W
silica glass
with WGS without WGS
2
h =5 W/m K tr=0.165 mm
0.4
0.3
200
silica glass steel
0.2
copper
150
0.1
copper
steel 0 100
0
1
2
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z/R b
5
6
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0
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10-1
1
c
0.9
o
Ti=120 C qi=0.25 W h =5 W/m2 K tr=0.165 mm
0.8 0.7
5
6
10-2
S CO
Y H2
silica glass
8
silica glass
tr=0.165 mm silica glass
0.6 0.5
7
o
Ti=120 C qi=0.25 W h =5 W/m2 K
d
with WGS without WGS
4
z/R b
steel copper
10-3
0.4
steel
0.3 10-4
0.2
copper
copper
0.1
with WGS without WGS
steel
0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
0
1
2
3
4
z/R b
5
6
7
8
Fig. 6 e Wall material effect on reformer performance. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
temperature is slightly decreased for the case including the WGS reaction as compared with the case excluding the WGS reaction. Consequently, the methanol conversion is slightly decreased as shown in Fig. 4(b) for the WGS reaction included case. Based on the reversed WGS reaction as indicated in Eq. (12), more CO is produced and more hydrogen is consumed. That is, the hydrogen yield is decreased as shown in Fig. 4(c) and CO selectivity is increased as shown in Fig. 4(d) in the reformer downstream. Using hN ¼ 5 W/m2 K and the same operation conditions for the results shown in Fig. 4, we examine the wall thickness effect on MSR performance in a silica glass reformer. The results are shown in Fig. 5 for wall thickness tr ¼ 0.165, 0.33, and 0.495 mm. From Fig. 5 the gas temperature decreases with
the increased reformer wall thickness. Consequently, methanol conversion, hydrogen yield and CO selectivity also decrease, as shown in Fig. 5(b)e(d). MSR performance for the case without WGS reaction has the same characteristics discussed in Fig. 4. A reformer can be made using various materials. In addition to silica glass, MSR performances in steel and copper reformers are also examined in this study. The thermophysical properties of these materials are listed in Table 2. In Fig. 6, MSR performance of reformers made of silica glass, steel, and copper are compared using the same operating conditions for the results in Fig. 4. Among these materials, copper has the best thermal conductivity and the lowest thermal resistance through the wall. As shown in
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300
1
a
0.9
qi=0.5 W
b
Ti=120 oC silica glass, tr=0.165 mm
0.8 250
0.7
0.3 W out
T out ( o C )
0.6
0.4
0.1W
200
0.5
0.3 0.2 internal heating external heating
150 0.1
0.2
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o
Ti=120 C silica glass, tr=0.165 mm 0.6
r/R b
0.7
0.8
0.9
internal heating external heating
0.1 0 0.1
1
0.2
0.3
qi (W)
0.4
0.5
100
1.5
c
d
Ti=120 oC
Ti=120 oC silica glass , tr=0.165 mm
silica glass, tr=0.165 mm 10-1
Y H 2 ,out
S C O ,out
1
10-2
0.5 10-3 internal heating with WGS external heating with WGS internal heating without WGS external heating without WGS
internal heating external heating 0 0.1
0.2
0.3
qi (W)
0.4
0.5
10
-4
0.1
0.2
0.3
qi (W)
0.4
0.5
Fig. 7 e Comparison of silica glass reformer performance under internal and external heating. (a) gas mixture temperature profiles at reformer outlet with various heat rates. (b) averaged methanol conversion at reformer outlet, (c) averaged hydrogen yield at reformer outlet, and (d) averaged CO selectivity at the reformer outlet.
Fig. 6(a) the reforming temperature decreases with the increase in material thermal conductivity. That is, the copper reformer has the lowest while the silica glass reformer has the highest reforming temperature. Note also that the temperature can reach fully developed thermal conditions for the copper reformer. This is logical since the isothermal wall solution is equivalent to the solution as the thermal conductivity approaches infinity. Because the MSR performance strongly depends on the gas mixture temperature, the corresponding methanol conversion, hydrogen yield and CO selectivity decrease with the increase in reformer wall material thermal conductivity, as shown in Fig. 6(b)e(d). Again, the MSR performance for the case without WGS reaction has the same characteristics as that discussed in Fig. 4.
4.3.
Thermal resistance concept application
The above described heat transfer effect on MSR performance can be explained further based on the simple heat conduction theory. For the internal heating case, the heat transfer from the heating rod to the environment can be approximated using one-dimension heat conduction as [42], qi ¼
TH TN Rporous þRwall þRconv
TH TN ¼ 1=2pRi Lb hp þðlog½ðRb þtr Þ=Rb =2pLb lc Þþð1=2pðRb þtr ÞLb hN Þ (26) where TH is the surface temperature of the heating rod, hp is the convective heat transfer coefficient in the porous catalyst bed
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
which depends on the gas flow velocity as well as the thermophysical properties of the catalyst bed. The correlation between the heat transfer, fluid flow and thermophysical properties in porous media can be found in the literature [51,52]. Rporous, Rwall, and Rconv are the convective thermal resistance due to flow in the porous catalyst bed, conductive thermal resistance due to wall thickness and convective thermal resistance due to the external convective heat transfer, respectively. As indicated in Eq. (26), the thermal resistance depends on the reformer geometric dimension, wall material and convective heat transfers in the catalytic bed and in the environment. In this study we do not intend to examine the Rporous effect but focus on the reformer geometry, wall material, and external heat loss mechanism effects on the MSR performance. As indicated in Eq. (26), an increase in hN results in decreases in Rconv and TH TN. For a fixed TN, TH then decreases. This is clearly shown in Fig. 4(a). Although an increase in reformer wall thickness can reduce Rconv due to the increased external heat transfer area, it also increases Rwall. The combined effect results in a decrease in temperature, as shown in Fig. 5(a). As also indicated in Eq. (26), an increase in material thermal conductivity results in decreases in Rwall and TH TN. Again, this implies that TH decreases for a fixed TN as shown in Fig. 6(a).
4.4.
MSR performance in externally heated reformers
Reformer performance using internal and external heating was also discussed by Suh et al. [17]. Their external heating was achieved by letting the reformer outer wall having an elevated temperature. Instead, we consider the external heating is provided by applying a heat flux on the reformer outer wall. In order to compare the MSR performance between the internal and externally heating, the same heat rate is applied and the same operating conditions discussed in Fig. 4 are used. For the externally heated reformer, the applied heat flux at the reformer outer wall is calculated asq00w ¼ qi =½2pðRb þ tr ÞLb . Fig. 7 shows comparisons of the silica glass reformer performance under internal and external heating cases using the temperature distribution and averaged methanol conversion, hydrogen yield and CO selectivity at the reformer outlet. Note that the results for the internal heating case are obtained with an insulated reformer outer wall while the results for the external heating case are obtained using an insulated heating rod wall. The heat rate applied is in the range of 0.1e0.5 W. As indicated in Fig. 7(a), the internal heating case has higher temperature near the heating rod surface and lower temperature near the reformer inner wall. For the external heating case, higher temperature occurs near the reformer inner wall and lower temperature is near the heating rod wall. Because the averaged temperatures are approximately the same for both internal heating and external heating cases for a given heat rate, the corresponding averaged methanol conversion, hydrogen yield, and CO selectivity are almost identical, as shown in Fig. 7(b)e(d). These results indicate that there is no significant difference in MSR performance under internal and external heating with the same applied heat rate. The CO selectivity for the case without WGS reaction is also shown in Fig. 7(d). With the inclusion of WGS reaction, the CO selectivity is approximately thirty times higher than that without the WGS reaction.
5.
261
Conclusion
We numerically investigated MSR performance in micro-scale cylindrical packed bed reformers under various heating supply mechanisms and thermal conditions at the reformer outer wall. The physical model was extended from the study reported by Suh et al. [17] with reformer wall was included. It was found that the thermal resistance from the heat source to the reformer environment plays an important role in MSR performance. The thermal resistance depends on the reformer geometry, wall material and heat transfer coefficients inside the catalyst bed and outside the reformer. Our numerical results suggested that to have high methanol conversion and hydrogen yield, a material with low thermal conductivity and thin wall thickness, such as silica glass, would be better than metals such as steel or copper. For both internal and external heating under the same heat rate, no significant difference in MSR performance was found. A water gas shift reaction model was included in the present numerical model. In the low-temperature reformer zone the forward WGS reaction was clearly demonstrated, which resulted in a decrease in CO selectivity. For the high temperature reformer zone, a reversed backward WGS reaction was also clearly demonstrated in which CO selectivity increased with the increase in temperature. When neglecting the WGS reaction it was found that CO selectivity would be higher in the low-temperature zone and lower in the high-temperature zone compared to the results including the WGS reaction. For both internal and external heating under the same heat rate, our results showed that CO selectivity would be underestimated when the WGS reaction is neglected and is about thirty times lower than the WGS reaction included case.
references
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Thermal resistance effect on methanol-steam reforming performance in micro-scale reformers Rei-Yu Chein a,*, Yen-Cho Chen b, J.N. Chung c a
Department of Mechanical Engineering, National Chung-Hsing University, Taichung City 402, Taiwan Department of Energy and Resource, National United University, Miaoli City 360, Taiwan c Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611-6300, USA b
article info
abstract
Article history:
We numerically investigate hydrogen production based on methanol-steam reforming
Received 13 April 2011
(MSR) using a micro-scale cylindrical packed bed reformer. The reformer wall is included in
Received in revised form
the physical model. The heat required for the reforming reaction is supplied either inter-
12 September 2011
nally using a heating rod placed along the center of the reformer or externally by a heat flux
Accepted 15 September 2011
applied at the reformer outer wall. Our results show that the thermal resistance from the
Available online 14 October 2011
heat source to the reformer environment plays an important role in the reformer performance. This thermal resistance depends on the reformer geometry, wall material and heat
Keywords:
transfer coefficients inside the catalyst bed and outside the reformer. Based on our
Methanol-steam reforming (MSR)
numerical results, it is suggested that better methanol conversion and hydrogen yield can
Packed-bed reformer
be obtained using reformer wall material with low thermal conductivity and thin thick-
Thermal resistance
ness. For both internal and external heating under the same heat rate supply, no significant
Internal and external heating
difference in reformer performance was found.
Water gas shift (WGS) reaction
A water gas shift (WGS) reaction model was included in the present numerical model. In the reformer low-temperature zone the forward WGS reaction was clearly demonstrated, resulting in a decrease in carbon monoxide (CO) selectivity. In the high temperature zone the backward WGS reaction was also clearly demonstrated in which CO selectivity increases with the increase in temperature. For both internal and external heating under the same heat rate supply, our results indicated that CO selectivity is about thirty times lower when the WGS reaction is neglected. Copyright ª 2011, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved.
1.
Introduction
The miniature fuel cell is efficient in converting chemical energy directly into electrical energy. It is a prominent alternative to batteries [1e3]. Among fuel cell designs the direct methanol fuel cell (DMFC) and polymer electrolyte fuel cell (PEMFC) have received much attention due to their low operating temperature which is suitable for portable applications.
The PEMFC has a much higher power density than the DMFC, but hydrogen storage is impractical for portable applications. A fuel reformer that converts liquid fuel into hydrogen is the possible solution. Among the fuels used for producing hydrogen, methanol has been widely adopted because of its advantages such as high hydrogen to carbon ratio, liquid form at ambient conditions, low reforming temperature (200e350 C) and low CO formation [4,5]. Methanol-steam
* Corresponding author. E-mail address: [email protected] (R.-Y. Chein). 0360-3199/$ e see front matter Copyright ª 2011, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2011.09.070
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
Nomenclature A1 B1 A2 CD CF cp CR CWGS Dij DTi dp ER ED EWGS hN K Keq kD kR kWGS Lb mcat Mi mi NG n_ i p Q qc
constant in reforming reaction rate, 1.15 106 m3 s1 kg1 constant in reforming reaction rate, 9.41 105 m3 s1 kg1 constant in decomposition reaction rate, 7.09 107 mol s1 kg1 correction factor decomposition reaction rate, 5.5 Forchheimer drag coefficient gas specific heat, J kg1 K1 correction factor for reforming reaction rate, 5.5 correction factor for WGS reaction rate, 11.2 binary molecular diffusion coefficient, m2 s1 thermal diffusion coefficient of species i, m2 s1 catalyst particle diameter, m activation energy for reforming reaction, 84,100 J mol1 activation energy for decomposition reaction, 111,200 J mol1 activation energy for WGS reaction, 70,000 J mol1 ambient heat transfer coefficient, W m2 K1 catalyst bed permeability, m2 equilibrium constant in WGS reaction rate constant of decomposition reaction, mol s1 kg1 rate constant of reforming reaction, m3 s1 kg1 rate constant of WGS reaction, mol m3 s1 Ka Pa2 reformer length, m catalyst weight, kg molecular weight of species i, g mol1 mass fraction of species i number of species in the gas mixture molar flow rate of species i pressure, Pa volumetric flow rate, m3 s1 energy source term due to the chemical reaction, J m3
reforming (MSR) for hydrogen production was reviewed recently by Holladay et al. [6,7]. The MSR reformer catalyst particles can be packed in a bed to form a packed-bed reformer [8,9] or coated onto the channel surface to form a plate reformer [10,11]. For portable power supply requirements, recent developments have focused on minimizing the size and weight of the reformer [12,13]. Due to its endothermic nature the MSR requires a heat supply from external sources. Because the chemical reactions involved in the MSR (reforming, decomposition, and water gas shift (WGS)) depend greatly on the temperature, understanding the temperature distribution and heat transfer within the reformer is essential for optimizing reformer design and operating conditions. For a packed bed reformer heated from an external source, several studies have shown that the temperature near the core of the reformer is lower than that near the wall [14e16]. The lower temperature leads to less catalyst activity and produces undesired chemical species. To overcome the adverse lower temperature in the reformer core
qi q00i q00w R Rb Ri ri rR rD rWGS SCO T tr / V Vb xi YH2
251
heat rate from heating rod, W heat flux at heating rod surface, W m2 heat flux at reformer outer wall, W m2 universal gas constant, 8.314 J mol1 K1 reformer radius, m heating rod radius, m molar generation rate of species i, mol m3 s1 reforming reaction rate, mol m3 s1 decomposition reaction rate, mol m3 s1 WGS reaction rate, mol m3 s1 CO selectivity temperature, K reformer wall thickness, m velocity, m s1 catalyst bed volume, m3 molar fraction of species i hydrogen yield
Greek symbols a temperature exponent d molar steam/CO ratio heat of decomposition reaction, J mol1 DHD DHR heat of reforming reaction, J mol1 DHWGS heat of reforming reaction, J mol1 ε catalyst layer porosity h methanol conversion l thermal conductivity, W m1 K1 m viscosity, kg m1 s1 f molar ratio of water to methanol r density, kg m3 Subscript c h m s N
reformer wall internal heating rod gas mixture catalyst ambient condition
region, Suh et al. [17] proposed an internal heating structure with heat supplied from a heating element suspended along the center of the catalyst bed. Their experimental and numerical results showed that under the same wall temperature, inlet steam-methanol flow rate and catalyst mass conditions the methanol conversion was higher for an internally heated reformer compared the externally heated reformer. Enhancing the heat transfer inside the catalyst bed can also lead to improved methanol conversion [18e20]. The heat supply to the reformer may be achieved in various ways [21e23]. Recent studies have used a catalytic combustor as the reformer heat source [24e29]. In contrast to the reformer, the catalytic combustion occurring in the combustor is an exothermic reaction. For catalytic combustion in the micro-scale dimension, it has been shown that the combustor wall material and heat loss to the environment play important roles in the combustion characteristics such as flame stability and exhaust gas temperature [30e33]. Although the reforming reaction is endothermic, the reformer wall
252
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
material and external heat loss will also affect the reformer performance and need to be identified. Due to its small size experimental measurements on the detailed temperature distribution inside the reformer is difficult. For this reason, modeling and simulation have been used extensively to obtain a better understanding of the fundamental transport phenomena inside the reformer, to reduce research cost, and to shorten the design cycles [34]. Many numerical simulations have been carried out for both packed bed and plate reformers [35e39]. In most of these studies, the reformer outer wall with elevated temperature was usually treated as the heat source for the chemical reaction. In the laboratory, an elevated reformer wall temperature can be achieved by putting the reformer inside an oven. However, for the cases with heat supplied from an integrated combustor or from an electric heater placed internally or externally, a heat flux is actually applied to the reformer instead of elevated temperature [40,41]. In terms of theoretical modeling, it should be more appropriate to use wall heat flux boundary condition instead of wall temperature boundary condition [42]. In this study, we numerically examine the performance of a packed bed reactor based on the design proposed by Suh et al. [17]. In their model, cylindrical packed bed MSR reformer performance with both internal and external heating was addressed. The external heating was achieved by assigning an elevated temperature to the infinite-thin reformer wall. Extending this model, we study the reformer made by various materials and has finite wall thickness. At the reformer wall outer surface, various boundary conditions are applied to emulate the heat transfer conditions such as insulated, natural or forced convective heat loss and external heat supply. The impact of WGS reaction on the CO selectivity in MSR is also discussed.
r, v
Catalyst Bed
CH3 OH+H2 O mixture
Theoretical model and formulation
2.1.
Physical model
Fig. 1(a) shows the physical model described in the work of Suh et al. [17]. The cylindrical packed-bed reformer has the length Lb ¼ 12 mm and diameter db ¼ 2Rb ¼ 1.5 mm. A chromium alloy rod with length equal to the reformer length and diameter di ¼ 2Ri ¼ 0.33 mm was placed along the reformer centerline for use as the internal heat source. The heat flux generated by the heating rod is q00i ¼ qi =pdi Lb with qi being the heat rate generated from the electric current passing through the heating rod. CuO/ZnO/Al2O3 pellets were packed in the space between the reformer and heating rod for use as the catalyst for the reforming reaction. A methanol-steam mixture with prescribed temperature Ti, volumetric flow rate Qi and steam to methanol molar ratio f was introduced into the reformer from the inlet (z ¼ 0). At the reformer outer wall (r ¼ Rb) temperature Tw with value higher than Ti served as the external heat source. In the model shown in Fig. 1(a), the reformer wall thickness was neglected which is not realistic. In this study, the reformer wall with thickness tr was included as shown in Fig. 1(b). The reformer can be fabricated using various materials and wall thicknesses. At the outer reformer wall various
z, u
Ri
Rb
Heating rod Heat flux qi
ui , Ti Lb
b
r, v
Tw , q
w
tr Catalyst Bed
z, u
Heating rod
CH3 OH+H2 O mixture
Heat flux q i
ui , Ti Reformer wall
h , T Fig. 1 e Physical models of the cylindrical packed bed reformer. (a) model in the study of Suh et al. [17]. (b) model extended from (a) with reformer wall included.
thermal boundary conditions would be possible. As shown in Fig. 1(b), the reformer may be operated in an environment with temperature TN (¼298 K) and heat transfer coefficient hN or subject to an applied heat flux q00w serving as the external heat source.
2.2.
2.
a Tw
Mathematical model
Transport phenomena in the reformer can be described by the conservation equations of mass, momentum, energy and species leading to a set of non-linear partial differential equations. To simplify the analysis, the following assumptions are made: (1) All the species in the gas mixture are ideal gases. (2) The gas flow in the reformer is assumed to be weakly compressible, axisymmetric, steady and laminar. (3) The catalyst particles are assumed to be spherical with a diameter dp and the catalyst bed is treated as a porous medium with homogeneous porosity ε and permeability K. (4) The catalyst bed is in local thermal equilibrium with the surrounding gas mixture. Based on the above assumptions, the governing equations for the mass conservation, fluid flow, energy transport, and species transport can be written as, /
V$ðεrV Þ ¼ 0
(1a)
/ / 4 . T 1 mm / 2mm 4 . VV I V$ðrV V Þ ¼ V$ pI þ þ ðVV Þ V$V ε 3 ε2 mm / rCF / / V pffiffiffiffi jV jV (1b) K K
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2 / V$ εrcp V T ¼ V$ðle VTÞ þ qc
8 9
= NG < / X Vp VT Dij Vxi þ ðxi mi Þ V$ εrV mi rmi DTi ¼ ri : p T ; j¼1
(2)
(3)
Eq. (1) is known as the BrinkmaneDarcyeForchheimer model for the fluid flow in a porous medium with a homogeneous porosity. r is the mass-weighted density defined as, G p X xi Mi RT i¼1
N
r¼
(4)
The permeability K and Forchheimer drag coefficient CF for a packed bed with spherical particles can be written as [43], pffiffiffiffiffiffiffiffi K ¼ d2p ε3 = 150ð1 εÞ2 ; CF ¼ 1:75= 150ε3=2
(5)
In Eq. (2) for the energy transport, cp is the mass-weighted specific heat defined as,
cp ¼
NG X
mi cpi
(6)
i¼1
le is the effective thermal conductivity of the catalyst bed defined as, le ¼ εlm þ ð1 εÞls
(7)
where ls and lm are the thermal conductivities of catalyst particle and gas mixture, respectively. qc is the energy source due to chemical reaction. Eq. (3) is known as the MaxwelleStefan species transport equation. Dij and DTi are the binary molecular diffusion coefficient and thermal diffusion coefficient of the species i in the gas mixture, respectively. ri is the production rate of species i due to chemical reaction. In these equations, the gas mixture transport properties (mm, lm, Dij and DTi ) can be evaluated based on the ChapmaneEnskog theory [44]. Since there are no flows in the reformer wall and internal heating rod, their temperature distributions are simply governed by the heat conduction, Reformer wall : V$ðkc VTc Þ ¼ 0
(8)
Heating rod : V$ðkh VTh Þ ¼ 0
(9)
All governing equations were written in a cylindrical coordinate system (r, q, z). Utilizing the symmetry in the q coordinate we can recast the problem as an axisymmetric twodimensional model (r, z). The fluid velocity components in the r and z-directions for the reactant flow are u and v, respectively.
2.3.
253
Chemical reaction model
Using CuO/ZnO/Al2O3 as the catalyst, the chemical reactions taking place during the MSR are [45,46], Steam reforming : CH3 OH þ H2 O/3H2 þ CO2
(10)
Decomposition : CH3 OH/2H2 þ CO
(11)
Water gas shift ðWGSÞ: CO þ H2 O/H2 CO2
(12)
For the packed-bed reformer, Suh et al. [35] proposed semiempirical kinetic models for the steam reforming and decomposition reactions as, rR ¼ ð1 εÞrs kR CCH3 OH
(13)
rD ¼ ð1 εÞrs kD
(14)
where CCH3 OH is the molar concentration of methanol. In Eqs. (13) and (14), rs, kR and kD are catalyst density, and rate constants for reforming and decomposition reactions defined as, rs ¼
mcat mcat ¼ 2 Vb p Rb R2i Lb
(15)
kR ¼ CR ½A1 þ B1 lnfeER =RT
(16)
kD ¼ CD A2 expð ED =RTÞ
(17)
In Eq. (15), mcat is the catalyst weight and Vb is the catalyst bed volume. In Eqs. (16) and (17), A1, A2 and B1 are constants given in the study by Amphlett et al. [9]. ER and ED are the activation energies for the reforming and decomposition reactions, respectively. R and T are the universal gas constant and gas mixture temperature, respectively. CR and CD are the correction factors for the reforming and decomposition reactions accounting for the catalyst activity and effectiveness and their values are chosen as 5.5 [17]. Based on the studies by Purnama et al. [47] and Chen et al. [48], the chemical reaction model for the WGS reaction can be written as, rWGS ¼ CWGS kWGS pCO pH2 O pCO2 pH2 =Keq
(18)
EWGS kWGS ¼ 1:78 1022 1 0:154d þ 0:008d2 Ta exp RT
(19)
4577:8 Keq ¼ exp 4:33 T
(20)
where kWGS and Keq are the rate constant and the equilibrium constant of the WGS reaction, respectively, pi (i ¼ CO,
Table 1 e Heat of reaction. Reforming Decomposition Water gas shifting
DHR ¼ 4:95 104 þ ðcpCO2 þ 3cpH2 cpCH3 OH cpH2 O ÞðT 298Þ DHD ¼ 9:07 104 þ ðcpCO þ 2cpH2 cpCH3 OH ÞðT 298Þ DHWGS ¼ 4:912 104 þ ðcpCO2 þ cpH2 cpH2 O cpCO ÞðT 298Þ
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H2O, CO2, H2) is the partial pressure, a is a constant determined by the experiments, and d is the molar ratio of water to CO. Based on the study by Chen et al. [48], a and activation energy EWGS are chosen as 8.0 and 70 kJ/mol, respectively. Similar to the models for reforming and decomposition proposed by Suh et al. [35], we introduced a correction factor CWGS for the WGS reaction to account for the catalyst activity and effectiveness. Based on the experimental data reported by Suh et al. [17], CWGS is taken as 11.2 in the present study. From the reaction models described above, the production rate of each species and energy source can be written as, rCH3 OH ¼ rR rD
rCO2 ¼ rR þ rWGS
(21c)
rCO ¼ rD rWGS
(21d)
rH2 ¼ 3rR þ 2rD þ rWGS
(21e)
qc ¼ rR DHR rD DHD þ rWGS DHWGS
(21f)
where DHR, DHD and DHWGS are the heat of reaction for the reforming, decomposition and WGS reactions, respectively. Temperature-dependent heats of reaction of these reactions are listed in Table 1 [28].
Boundary conditions
Boundary conditions must be specified to complete the mathematical model. Referring to Fig. 1, the boundary conditions are specified as follows,
(22a)
(2) reformer outlet (z ¼ Lb, Ri < r < Rb) /
vV vT vmi ¼ ¼0 ¼ vz vz vz
(22b)
(3) walls of reformer and internal heating rod at inlet and outlet (z ¼ 0, z ¼ Lb, 0 < r < Ri and Rb < r < Rb þ tr) vTh vTc ¼ ¼0 vz vz
(22c)
(4) internal heating rod outer surface (0 < z < Lb, r ¼ Ri) /
V ¼ 0; lh
vTh vmi ¼ q00i ; ¼0 vr vr
(22d)
vT vTc vmi ; ¼0 ¼ lc vr vr vr
(22e)
(6) along reformer centerline (0 < z
(22f)
(7) reformer outer wall (0 < z < Lb, r ¼ Rb þ tr) vTc vTc ¼ hN ðTc TN Þ or lc ¼ q00w vr vr
(22g)
In Eq. (22a), ui ¼ Qi =ðpðR2b R2i ÞÞ and mCH3 OH;in are the inlet velocity and the mass fraction of CH3OH, respectively. At the reformer outlet, the Neumann boundary conditions are specified for the flow velocity, temperature and species concentrations as indicated in Eq. (22b). At the inlet and outlet, the walls of the reformer and internal heating rod are assumed to be insulated as described in Eq. (22c). In Eq. (22d), a constant heat flux q00i generated by the heating rod is specified while the gas flow satisfies the no-slip boundary condition and no species deposition occurs on the heating rod surface. Referring to Fig. 1(b), the conditions for no-slip gas flow, continuous heat transfer and no species deposition at the reformer inner wall are specified as indicated in Eq. (22e). The axisymmetric condition is specified along the reformer centerline as indicated in Eq. (22f). Eq. (22g) shows that various boundary conditions including an elevated wall temperature Tw, convective heat loss to the environment with heat transfer coefficient hN and temperature TN, and applied heat flux q00w for external heating can be specified at the reformer outer wall.
3.
(1) reformer inlet (z ¼ 0, Ri < r < Rb)
mH2 ¼ 0; mCO2 ¼ 0; and mCO ¼ 0
V ¼ 0; T ¼ Tc ; le
Tc ¼ Tw ; lc (21b)
u ¼ ui ; v ¼ 0;T ¼ Ti ; mCH3 OH ¼ mCH3 OH;in ; mH2 O ¼ fmCH3 OH;in ;
/
(21a)
rH2 O ¼ rR rWGS
2.4.
(5) reformer inner wall (0 < z
Numerical methods
All of the governing equations along with the boundary conditions were solved simultaneously using COMSOL Multiphysics (Comsol Inc., version 4.01). Weakly compressible NaviereStokes, general heat transfer and MaxwelleStefan diffusion and convection modules were applied for solving the velocity, temperature and species concentration distributions in the reformer. Because the numerical solution accuracy strongly depends on the mesh size, a refined mesh is necessary in the region where the dependent variable gradients are pronounced. Finer meshes were used to capture the subtle
Table 2 e Thermophysical property of the wall materials. Material
Chromium alloy Silica glass Steel (AISI4340) Copper
Thermal Specific Thermal Density heat capacity [kg/m3] conductivity [W/mK] [kJ/kgK] [kJ/m3K] 8400 2203 7850 8700
11.3 1.38 44.5 400
450 703 475 385
3.78 1.54 3.72 3.35
106 106 106 106
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1
3000
a
b Analysis [17] experiment [17] present study
0.8
Analysis [17] experiment [17] present model
2500
2000
C O (ppm)
0.6
0.4
1500
1000 0.2
500
0 160
170
180 o
Tw ( C)
190
0 160
200
280
d
Ti=120 C Tw=180 oC
10
240
180
0.4
160
qi (W) 0 0.25 0.5 0.75
140 120
0
2
qi=0.25 W
xC O (ppm)
x M eO H
T( o C )
200
0.5 W
0.75 W
3
102
0.5
220
101 100
10-1
Ti=120 oC o Tw=180 C
10-2
4
z/R
6
8
190
104
0.6
c
180
Tw (o C)
o
260
170
0.3
10
-3
0
2
b
4
z/R b
6
8
Fig. 2 e Verification of the present numerical model by comparing the results reported in the study by Suh et al. [17]. Qi [ 900 ml/h, mcat [ 16 mg (a) methanol conversion, (b) CO concentration, (c) gas mixture temperature and methanol molar fraction distributions along the heating rod surface (r [ Ri) with various heat rates. mcat [ 16 mg, Ti [ 120 C, and Tw [ 180 C. (d) CO concentrations for the cases with and without WGS reaction based on the operation conditions used in (c).
changes in velocity, temperature and species concentration in the inlet, outlet, and fluidewall interface regions. The solution independence on the mesh size was carefully studied before reporting the final results. The numerical experiments show that the solution becomes mesh-independent when the element number exceeds approximately 4000. Hence, more than 4000 meshes were used for the results presented in this study. Important characteristics for MSR are methanol conversion, hydrogen yield and CO selectivity. Methanol conversion is defined as the ratio between converted methanol at a position along the reformer and the inlet methanol flux [32],
h¼
n_ CH3 OH;in n_ CH3 OH n_ CH3 OH;in
(23)
Hydrogen yield characterizes the reformer performance with respect to the hydrogen production. It is the ratio of the produced hydrogen to the theoretical maximum amount of hydrogen [32], n_ H3 (24) YH2 ¼ n_ CH3 OH;in CO selectivity defines the molar fraction between produced CO and the molar fraction of all carbon-containing gas phase species which are CO and carbon dioxide (CO2) in MSR. It is an
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indicator for the effectiveness of carbon monoxide production and is defined as [32], SCO ¼
n_ CO n_ CO þ n_ CO2
4.
Results and discussion
4.1.
Numerical model verification
(25)
The numerical model used in this study was verified by comparing the results computed from the present numerical model with those reported by Suh et al. [17] under the same reformer geometry shown in Fig. 1(a). The inlet flow rate and catalyst mass are 900 ml/h and 16 mg, respectively. For the ratio of steam to methanol f ¼ 1.1, the inlet mass fractions of methanol and water are mCH3 OH;in ¼ 0:62 andmH2 O;in ¼ 0:38, respectively. The internal heating rod is made of chromium alloy with properties listed in Table 2. The external heating was provided by an elevated wall temperature Tw. In Fig. 2(a), the methanol conversion under various qi, Ti and Tw predicted by the present numerical model are compared with the experimental and numerical results reported by Suh et al. [17]. Good agreement between the numerical and the experimental results is obtained. In Fig. 2(b), comparison of CO concentration obtained by the present model with those reported by Suh et al. [17] is presented. It is seen that the model used in the present study under predicts the CO production. The discrepancy may be due to both WGS model accuracy and experimental uncertainty. Although the WGS reaction model used in the present study underestimates the CO concentration compared with the experimental measurement, it gives the correct variation trend. With the inclusion of WGS reaction, the CO concentration predicted by the present model is closer to the experimental data compared with the numerical results predicted by Suh et al. [17] in which the WGS reaction was neglected. For various heat rates and Ti ¼ 120 C, temperature and methanol molar fraction variations along the heating rod surface (r ¼ Ri) predicted by the present numerical model are shown in Fig. 2(c). Our model successfully reproduced the results in Fig. 8 in the study by Suh et al. [17]. To understand the WGS reaction effect, the CO concentrations for the cases with and without WGS reaction are compared in Fig. 2(d) under the same operating conditions as the results shown in Fig. 2(c). In the entrance zone at which the temperature is low, the CO concentration with WGS reaction has lower value than that without WGS reaction. This is because of the forward WGS reaction that converts CO into CO2. However as the temperature increases in the reformer downstream, the CO value with WGS reaction becomes higher than that without WGS reaction. Because the WGS reaction is a moderately exothermic reaction, the reaction tends to shift to the left side at high temperature according to the Le Chatelier’s principle. The reversed WGS reaction leads to lower CO conversion and hydrogen yield [47,49,50].
4.2.
reformer wall included in the physical domain as shown in Fig. 1(b). In order to focus on the heat transfer effect on the MSR performance, the inlet flow rate and temperature of the methanol-water mixture are fixed as Qi ¼ 900 ml/h and Ti ¼ 120 C. The catalyst mass is also fixed as 16 mg. We first examine the external heat loss effect on the MSR performance for a silica glass reformer. The silica glass property used in this study is listed in Table 2. The reformer wall thickness is taken as 0.165 mm which is the same as radius of the internal heating rod. At the outer reformer wall, various heat transfer coefficients are used to examine the heat loss effect on the MSR performance. Using the ambient heat transfer coefficient hN in the range of 0 (insulated wall) to 50 W/m2 K and
MSR performance with reformer wall
Based on the comparison described above, our numerical model can be applied to examine the MSR performance with
Fig. 3 e Typical temperature and species molar fraction distributions in the silica glass reformer. Ti [ 120 C, qi [ 0.25 W, tr [ 0.165 mm, and hN [ 5 W/m2 K.
257
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
0.6
300
a
o
Ti=120 C qi=0.25 W
b
with WGS without WGS
0.5
tr=0.165 mm 250
insulated
with WGS without WGS
o
Ti=120 C qi=0.25 W tr=0.165 mm
0.4
5
ToC
0.3
200
h =25 W/m2 K
insulated 5
0.2
150
0.1 2
h =25 W K
50 0 100
0
1
2
3
4
z/R b
5
6
7
0
8
1
1
2
3
4
z/R b
5
6
7
8
10-1 o
c
0.9
Ti=120 C qi=0.25 W tr=0.165 mm
0.8
d
with WGS without WGS
o
Ti=120 C qi=0.25 W tr=0.165 mm
10-2
0.7
insulated 5 insulated
insulated
0.6
5 25
0.5
S CO
Y H2
50
5
0.4
10-3
50 25
0.3 10-4
0.2
2
h =50 W/m K
50
0.1
h =25 W/m2 K
with WGS without wGS
0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
1
2
3
4
z/R b
5
6
7
8
Fig. 4 e Ambient heat transfer coefficient effect on the MSR performance in silica glass reformer. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
qi ¼ 0.25 W, the MSR performance are shown in Figs. 3 and 4. In Fig. 3, temperature and species molar fraction distributions for hN ¼ 5 W/m2 K are shown. As the methanol-steam mixture is introduced into the reformer, it is heated by the internal heating rod. At the same time, heat is lost to the environment from the reformer outer wall. The temperature distribution shows that there is a low-temperature zone near the reformer entrance. Temperature increases at the reformer downstream due to the heat supply from the heating rod. Because of the external heat loss, a temperature gradient also exists in the radial direction. Because the MSR depends on temperature, the methanol concentration decreases while the CO and hydrogen concentrations increase along the reformer length.
Similar to Fig. 2(c) and (d), we use temperature, methanol conversion, hydrogen yield and CO selectivity variations along the heating rod surface r ¼ Ri to quantify the reformer performance. In Fig. 4(a), the highest temperature can be obtained for the insulated reformer outer wall case while the temperature decreases with the increase in ambient heat transfer coefficient. This is expected since a higher heat transfer coefficient implies higher heat loss to the environment. Fig. 4(a) also indicates that fully developed temperature cannot be reached when the external heat loss is present. The corresponding methanol conversion, hydrogen yield and CO selectivity are shown in Fig. 4(b), (c) and (d), respectively. Because of the decrease in reaction temperature, methanol
258
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
300
0.6
a
with WGS without WGS
o
Ti=120 C qi=0.25 W 2 h =5 W/m K
b
o
Ti=120 C
0.5
with WGS without WGS
qi=0.25 W
250
h =5 W/m2 K
T (o C )
0.4
0.33 mm
0.3
0.165 mm
200
0.165 mm
0.330 mm
tr=0.495 mm
tr=0.495 mm
0.2
150
0.1
0 100
0
1
2
3
4
z/R b
5
6
7
8
0
1
2
3
4
z/R b
5
6
7
8
10-1
1
c
0.9
Ti=120 C qi=0.25 W h =5 W/m2 K
0.8
o
d
o
with WGS without WGS
Ti=120 C qi=0.25 W h =5 W/m2 K
10-2
0.7
0.165 mm
0.5 0.4
S CO
Y H2
0.6
0.495 mm
10-3
0.33 mm
0.330 mm
0.3
tr=0.165 mm
tr=0.495 mm 10-4
0.2
with WGS without WGS
0.1 0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
0
1
2
3
4
z/R b
5
6
7
8
Fig. 5 e Wall thickness effect on the silica glass reformer performance. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
conversion, hydrogen yield and CO selectivity decrease with the increase in ambient heat transfer coefficient. As shown in Fig. 4(b) and (c), the methanol conversion and hydrogen yield have approximately the same variation trends. In Fig. 4(d), a decrease in CO selectivity can be observed in the entrance zone due to the forward WGS reaction. As the temperature increases in the reformer downstream, CO selectivity increases due to the reversed WGS reaction. For the case of hN ¼ 50 W/m2 K, the temperature rises slightly at the entrance due to the forward WGS reaction which is an exothermic reaction. At the reformer downstream, the temperature drops significantly due to strong heat loss and weak MSR results due to insufficient heat supply. To understand the WGS reaction effect on MSR performance, results without the WGS reaction are also shown in
Fig. 4. From Fig. 4(a)e(c), the WGS reaction has no significant effect on the temperature, methanol conversion and hydrogen yield, as proven by Pepply et al. and Amphlett et al. [41e43] when gas temperature is low. However, for the CO selectivity results shown in Fig. 4(d), higher CO selectivity results for the case without WGS reaction compared with the case including the WGS in the region near the entrance zone because of the forward WGS reaction. As the temperature increases in the reformer downstream, the reverse situation occurs. That is, higher CO selectivity occurs when the WGS reaction is included. The reason for this result is due to the temperature-dependent WGS reaction discussed earlier. Due to high reaction temperature in the reformer downstream, the reversed WGS reaction is favored. Since the reversed WGS reaction becomes an endothermic reaction, the reaction
259
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
300
0.6
a
with WGS without WGS
h =5 W/m K tr=0.165 mm
250
Ti=120 o C qi=0.25 W
0.5
2
T (o C )
b
o
Ti=120 C qi=0.25 W
silica glass
with WGS without WGS
2
h =5 W/m K tr=0.165 mm
0.4
0.3
200
silica glass steel
0.2
copper
150
0.1
copper
steel 0 100
0
1
2
3
4
z/R b
5
6
7
8
0
1
2
3
10-1
1
c
0.9
o
Ti=120 C qi=0.25 W h =5 W/m2 K tr=0.165 mm
0.8 0.7
5
6
10-2
S CO
Y H2
silica glass
8
silica glass
tr=0.165 mm silica glass
0.6 0.5
7
o
Ti=120 C qi=0.25 W h =5 W/m2 K
d
with WGS without WGS
4
z/R b
steel copper
10-3
0.4
steel
0.3 10-4
0.2
copper
copper
0.1
with WGS without WGS
steel
0 0
1
2
3
4
z/R b
5
6
7
8
10
-5
0
1
2
3
4
z/R b
5
6
7
8
Fig. 6 e Wall material effect on reformer performance. Distributions of (a) gas mixture temperature, (b) methanol conversion, (c) hydrogen yield, and (d) CO selectivity along the heating rod surface (r [ Ri).
temperature is slightly decreased for the case including the WGS reaction as compared with the case excluding the WGS reaction. Consequently, the methanol conversion is slightly decreased as shown in Fig. 4(b) for the WGS reaction included case. Based on the reversed WGS reaction as indicated in Eq. (12), more CO is produced and more hydrogen is consumed. That is, the hydrogen yield is decreased as shown in Fig. 4(c) and CO selectivity is increased as shown in Fig. 4(d) in the reformer downstream. Using hN ¼ 5 W/m2 K and the same operation conditions for the results shown in Fig. 4, we examine the wall thickness effect on MSR performance in a silica glass reformer. The results are shown in Fig. 5 for wall thickness tr ¼ 0.165, 0.33, and 0.495 mm. From Fig. 5 the gas temperature decreases with
the increased reformer wall thickness. Consequently, methanol conversion, hydrogen yield and CO selectivity also decrease, as shown in Fig. 5(b)e(d). MSR performance for the case without WGS reaction has the same characteristics discussed in Fig. 4. A reformer can be made using various materials. In addition to silica glass, MSR performances in steel and copper reformers are also examined in this study. The thermophysical properties of these materials are listed in Table 2. In Fig. 6, MSR performance of reformers made of silica glass, steel, and copper are compared using the same operating conditions for the results in Fig. 4. Among these materials, copper has the best thermal conductivity and the lowest thermal resistance through the wall. As shown in
260
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
300
1
a
0.9
qi=0.5 W
b
Ti=120 oC silica glass, tr=0.165 mm
0.8 250
0.7
0.3 W out
T out ( o C )
0.6
0.4
0.1W
200
0.5
0.3 0.2 internal heating external heating
150 0.1
0.2
0.3
0.4
0.5
o
Ti=120 C silica glass, tr=0.165 mm 0.6
r/R b
0.7
0.8
0.9
internal heating external heating
0.1 0 0.1
1
0.2
0.3
qi (W)
0.4
0.5
100
1.5
c
d
Ti=120 oC
Ti=120 oC silica glass , tr=0.165 mm
silica glass, tr=0.165 mm 10-1
Y H 2 ,out
S C O ,out
1
10-2
0.5 10-3 internal heating with WGS external heating with WGS internal heating without WGS external heating without WGS
internal heating external heating 0 0.1
0.2
0.3
qi (W)
0.4
0.5
10
-4
0.1
0.2
0.3
qi (W)
0.4
0.5
Fig. 7 e Comparison of silica glass reformer performance under internal and external heating. (a) gas mixture temperature profiles at reformer outlet with various heat rates. (b) averaged methanol conversion at reformer outlet, (c) averaged hydrogen yield at reformer outlet, and (d) averaged CO selectivity at the reformer outlet.
Fig. 6(a) the reforming temperature decreases with the increase in material thermal conductivity. That is, the copper reformer has the lowest while the silica glass reformer has the highest reforming temperature. Note also that the temperature can reach fully developed thermal conditions for the copper reformer. This is logical since the isothermal wall solution is equivalent to the solution as the thermal conductivity approaches infinity. Because the MSR performance strongly depends on the gas mixture temperature, the corresponding methanol conversion, hydrogen yield and CO selectivity decrease with the increase in reformer wall material thermal conductivity, as shown in Fig. 6(b)e(d). Again, the MSR performance for the case without WGS reaction has the same characteristics as that discussed in Fig. 4.
4.3.
Thermal resistance concept application
The above described heat transfer effect on MSR performance can be explained further based on the simple heat conduction theory. For the internal heating case, the heat transfer from the heating rod to the environment can be approximated using one-dimension heat conduction as [42], qi ¼
TH TN Rporous þRwall þRconv
TH TN ¼ 1=2pRi Lb hp þðlog½ðRb þtr Þ=Rb =2pLb lc Þþð1=2pðRb þtr ÞLb hN Þ (26) where TH is the surface temperature of the heating rod, hp is the convective heat transfer coefficient in the porous catalyst bed
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 7 ( 2 0 1 2 ) 2 5 0 e2 6 2
which depends on the gas flow velocity as well as the thermophysical properties of the catalyst bed. The correlation between the heat transfer, fluid flow and thermophysical properties in porous media can be found in the literature [51,52]. Rporous, Rwall, and Rconv are the convective thermal resistance due to flow in the porous catalyst bed, conductive thermal resistance due to wall thickness and convective thermal resistance due to the external convective heat transfer, respectively. As indicated in Eq. (26), the thermal resistance depends on the reformer geometric dimension, wall material and convective heat transfers in the catalytic bed and in the environment. In this study we do not intend to examine the Rporous effect but focus on the reformer geometry, wall material, and external heat loss mechanism effects on the MSR performance. As indicated in Eq. (26), an increase in hN results in decreases in Rconv and TH TN. For a fixed TN, TH then decreases. This is clearly shown in Fig. 4(a). Although an increase in reformer wall thickness can reduce Rconv due to the increased external heat transfer area, it also increases Rwall. The combined effect results in a decrease in temperature, as shown in Fig. 5(a). As also indicated in Eq. (26), an increase in material thermal conductivity results in decreases in Rwall and TH TN. Again, this implies that TH decreases for a fixed TN as shown in Fig. 6(a).
4.4.
MSR performance in externally heated reformers
Reformer performance using internal and external heating was also discussed by Suh et al. [17]. Their external heating was achieved by letting the reformer outer wall having an elevated temperature. Instead, we consider the external heating is provided by applying a heat flux on the reformer outer wall. In order to compare the MSR performance between the internal and externally heating, the same heat rate is applied and the same operating conditions discussed in Fig. 4 are used. For the externally heated reformer, the applied heat flux at the reformer outer wall is calculated asq00w ¼ qi =½2pðRb þ tr ÞLb . Fig. 7 shows comparisons of the silica glass reformer performance under internal and external heating cases using the temperature distribution and averaged methanol conversion, hydrogen yield and CO selectivity at the reformer outlet. Note that the results for the internal heating case are obtained with an insulated reformer outer wall while the results for the external heating case are obtained using an insulated heating rod wall. The heat rate applied is in the range of 0.1e0.5 W. As indicated in Fig. 7(a), the internal heating case has higher temperature near the heating rod surface and lower temperature near the reformer inner wall. For the external heating case, higher temperature occurs near the reformer inner wall and lower temperature is near the heating rod wall. Because the averaged temperatures are approximately the same for both internal heating and external heating cases for a given heat rate, the corresponding averaged methanol conversion, hydrogen yield, and CO selectivity are almost identical, as shown in Fig. 7(b)e(d). These results indicate that there is no significant difference in MSR performance under internal and external heating with the same applied heat rate. The CO selectivity for the case without WGS reaction is also shown in Fig. 7(d). With the inclusion of WGS reaction, the CO selectivity is approximately thirty times higher than that without the WGS reaction.
5.
261
Conclusion
We numerically investigated MSR performance in micro-scale cylindrical packed bed reformers under various heating supply mechanisms and thermal conditions at the reformer outer wall. The physical model was extended from the study reported by Suh et al. [17] with reformer wall was included. It was found that the thermal resistance from the heat source to the reformer environment plays an important role in MSR performance. The thermal resistance depends on the reformer geometry, wall material and heat transfer coefficients inside the catalyst bed and outside the reformer. Our numerical results suggested that to have high methanol conversion and hydrogen yield, a material with low thermal conductivity and thin wall thickness, such as silica glass, would be better than metals such as steel or copper. For both internal and external heating under the same heat rate, no significant difference in MSR performance was found. A water gas shift reaction model was included in the present numerical model. In the low-temperature reformer zone the forward WGS reaction was clearly demonstrated, which resulted in a decrease in CO selectivity. For the high temperature reformer zone, a reversed backward WGS reaction was also clearly demonstrated in which CO selectivity increased with the increase in temperature. When neglecting the WGS reaction it was found that CO selectivity would be higher in the low-temperature zone and lower in the high-temperature zone compared to the results including the WGS reaction. For both internal and external heating under the same heat rate, our results showed that CO selectivity would be underestimated when the WGS reaction is neglected and is about thirty times lower than the WGS reaction included case.
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