Ind. Eng. Chem. Res. 2009, 48, 8381–8386
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Reactor Development for Supercritical Water Gasification of 4.9 wt% Glucose Solution at 673 K by Using Computational Fluid Dynamics Takuya Yoshida* and Yukihiko Matsumura Mechanical System Engineering, School of Engineering, Hiroshima UniVersity, 1-4-1 Kagamiyama Higashi-Hiroshima, Hiroshima 739-8527, Japan
Supercritical water gasification is suitable for gasifying biomass with high moisture content. Although biomass material can be easily decomposed in supercritical water, it can polymerize to form char products, which may result in serious problems such as plugging of the reactor. To depress char formation during gasification of biomass material in supercritical water, a combination of reactors was proposed in our previous study. In this study, we have improved the system by means of computational fluid dynamics (CFD) and verified it through experiments. With the improved reactor, 4.9 wt% glucose solution was successfully gasified at 673 K and 25.4 MPa; the carbon gasification efficiency was greater than 0.94. Introduction Biomass is now recognized as an attractive resource for energy use as well as material utilization because it is carbon neutral. Generally, raw biomass contains moisture; therefore, in some cases, the cost of drying treatment in the biomass utilization process can be significant. In the supercritical water gasification (SCWG) process, biomass with high moisture content can be easily gasified without any drying process. Therefore many researchers have studied biomass gasification process using supercritical water (SCW). Recently, Peterson et al.1 summarized the various works of this field in their review paper. In SCW, biomass is easily decomposed (i.e., gasified or liquefied); however, simultaneously, polymerizations of the biomass and its fragments can occur. The rate of polymerization varies depending on the concentrations of the reactant and the catalyst performance as well as the temperature and pressure conditions. Polymerization reactions that produce char and tar materials can cause serious problems in the SCWG process because they can plug up the reactor. In fact, polymerization and char formation become more favorable at lower temperature in the SCWG process, which makes it difficult to completely gasify biomass material at moderate temperature. Moreover the tar material formed from the fragment of biomass material can deactivate the catalyst, probably by the formation of coke on the catalyst.2,3 To achieve practical economies, the dry material content of the biomass solution or slurry to be gasified in SCW or hydrothermal processes should be greater than 15-20 wt%.1 Nevertheless, high dry material content in biomass solution or slurry easily leads to formations of tar and char material, resulting in low gasification efficiency. Therefore, one of the main targets of biomass gasification in SCW or hydrothermal processes is a complete gasification of biomass solution or slurry with high dry biomass content. Some researchers proved that complete gasification of biomass solution or slurry with more than 10 wt% of dry biomass content can be achieved by using batch reactors employing heterogeneous catalysts at 673-773 K.4-8 In most cases, the reaction time required to attain complete gasification exceeds 20 min. That implies that it requires a long reaction time to decompose the char once char has formed in the reactor, whereas, with a continuous process, we rarely find * To whom correspondence should be addressed. E-mail:
[email protected]. Tel.: +81-82-424-5762. Fax: +8182-422-7193.
successful cases that attained complete gasification of biomass solution at moderate temperature (673-723 K) and with high solids loading that easily form char and tar. To completely gasify biomass materials at 673-723 K through a continuous process, we proposed a combination of a pyrolysis reactor, an oxidation reactor, and a catalytic reactor in the previous study.9 In this temperature region, polymerization of biomass components and its fragments occur to some extent, resulting in the formation of tar and char products. Therefore, we designed the oxidation reactor connected after the pyrolysis reactor so that the tar and char products would be effectively decomposed. The previous reactor successfully gasified a glucose and lignin mixture solution completely at 673 K. The concentration of the solution, however, was up to 0.4 wt%. With higher concentration solution, the catalyst was deactivated probably due to coke formation on its surface in our preliminary experiment, which indicated that the tar and char products that were not effectively decomposed in the oxidation reactor flowed into the catalytic reactor. We recognized that some improvement in the design of the oxidation reactor was required for complete gasification of higher concentration glucose solution. Computational fluid dynamics (CFD) is a helpful tool to simulate and design a reactor.10-14 In addition, by studying the reactions in SCW, some researchers verified the applicability of CFD to the reactions in SCW.15,16 Sugiyama et al. conducted CFD simulations as well as experiments on supercritical water oxidation (SCWO) of a carbon particle, and they verified that the Schlieren image of the flow around the carbon particle was in good qualitative agreement with a simulated Schlieren image.15 Moussie`re et al. conducted simulations and experiments on a SCWO process and confirmed that the simulated wall temperature profiles were in good agreement with the experimental ones.16 In this study, we have conducted CFD simulations to investigate the reactant flow in the reactors. Moreover, we have developed three new oxidation reactors and have conducted experiments on SCWG and related experiments to verify the CFD simulation results and to confirm that the reactor system can gasify the glucose solution at 673 K. Experimental and Numerical Methods Experimental Section. Glucose (Kanto Kagaku, Japan) was used as the reactant and 34 wt% hydrogen peroxide solution (Kanto Kagaku, Japan) was used as the oxidant. Glucose
10.1021/ie9002188 CCC: $40.75 2009 American Chemical Society Published on Web 08/20/2009
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Figure 2. Schematic views of the oxidation reactors used in this study. (A) reactor A, (B) reactor B, (C) reactor C, (bi) bottom inlet (i.d. 4.4 mm), (mi) middle inlet (i.d. 1.0 mm), (o) outlet (i.d. 2.3 mm).
Figure 1. Schematic diagram of the gasification reactor system used in this study.
solution was prepared by dissolving 100 g of the glucose into degassed ion-exchange water, and then it was diluted to 2000 mL. The density and concentration of the resulting glucose solution were 1.02 g/mL and 4.90 wt%, respectively. The oxidant solution was prepared by diluting 250 mL of hydrogen peroxide solution (34 wt%) to 1000 mL, the density of the resulting solution was 1.03 g/mL. We employed a similar apparatus and procedure to those in the previous study9 as shown in Figure 1. The reactors were made of type 316 stainless steel. The pyrolysis reactor was made of a stainless steel tube (i.d., 1 mm; length, 24 m). Three different oxidation reactors named reactors A, B, and C were used in this study; the inner volumes of the reactors were 9.2, 12.8, and 16.9 mL, respectively. Schematic views of the reactors are shown in Figure 2. We conducted several experiments employing the combinations of the pyrolysis reactor and each oxidation reactor. In addition, we conducted a gasification experiment with a combination of the pyrolysis reactor, the oxidation reactor (reactor C), and a catalytic reactor. The catalytic reactor employed in this study comprised a series of two stainless steel tubes (i.d., 6.53 mm; length, 200 mm), which contained a total of 13.7 g of reduced nickel catalyst (Ni-5256E), which is tubular extrudate with diameter of 1.2 mm (3/64 in.) and average length of 6 mm, obtained from N.E. Chemcat (Tokyo, Japan). Prior to the experiments, the reactor was sunk into the molten salt bath that had been heated up to adequate temperature, and then degassed ion-exchange water was pressurized and fed into the reactor system by HPLC pumps until the temperature of the reactor became steady. When the catalytic reactor was employed, water was fed for approximately 2 h so that the adsorbed gas on the
catalyst (mainly hydrogen that was used for reducing the nickel) was released and became negligible. After the temperature of the reactors and the molten salt bath became stable (approximately 673 K), the glucose solution and hydrogen peroxide solution were fed instead of the water flow. The glucose solution was fed into the pyrolysis reactor. The outlet of the pyrolysis reactor was connected to the middle inlet of the oxidation reactor; thus, the effluent from the pyrolysis reactor flowed into the oxidation reactor. The hydrogen peroxide solution was fed first into the preheat line and then into the oxidation reactor from its bottom inlet. The outlet of the oxidation reactor was located at its top. When the catalytic reactor was employed, the effluent of the oxidation reactor flowed into the catalytic reactor. The effluent from the reactor system flowed through the heat exchanger and the back pressure regulator, where it was cooled down and depressurized, respectively. The temperature and the pressure in the reactor were controlled at 673 K and 25.4 MPa, respectively, in each experiment. The flow rates of the reactant solution and the oxidant solution were maintained throughout the experiment at 5.2 g/min and 1.6 g/min, respectively. Thus, the equivalence ratio, which is defined as the ratio of the oxidant fed relative to the amount required for complete combustion, was 0.27. When the gaseous and liquid products were collected, the effluent was fed into the gas-liquid separator that was vacuumed in advance, and the amount of the gaseous product was calculated on the basis of the pressure difference before and after the collection. Char products were collected immediately by suction filtration at the outlet after the back pressure regulator. The filter paper used in the filtration was GMF150 (pore size, 1.0 µm, Whatman). The gaseous product was analyzed by a GC-TCD (GC14A, Shimadzu) equipped with a SINCARBON-T column (Shinwakako). The liquid product was analyzed by a TOC (TOC-V, Shimadzu) and an HPLC equipped with an RI detector (RID10A, Shimazdu) and an SEC column (Asahipak GS-220 HQ, Shodex). The filter paper was dried in a vacuumed desiccator before and after the filtration until the weight changes became negligible, and the weights of the char products were calculated on the basis of the weight difference in the filter paper. CFD Simulation. To simulate the reactant and product flow in the oxidation reactors, we employed a commercial CFD code, FLUENT 6.3 (Ansys-Fluent Inc., USA). Although many kinds of reactions were involved in the oxidation reactors, we ignored
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all the reactions in the simulations and calculated only the flow of water and inert particles (as virtual char products). The details of the simulation are as follows. Meshing. The 3d geometry of each reactor was built using GAMBIT software. Each oxidation reactor had a plane-symmetrical shape, thus the geometry was cut in half and the half body was subjected to the simulation with a plane-symmetrical boundary condition applied to the cut surface. With turbulence, the flow could be asymmetrical. Thus we preliminarily tested applicability of symmetrical assumption by conducting CFD simulations using both of whole-reactor geometry and half-reactor geometry with symmetrical boundary condition. The results indicated that there were some differences, however negligible in terms of the particle flow. Therefore we employed half-reactor geometries with symmetrical boundary condition. In addition, the mixing effect around the reactant inlet was important in this study, the meshes were densely divided around the reactant inlet, as well as around the oxidant inlet and the outlet at the top. The resulting meshes contained 85000-92100 tetrahedral cells. Model. We employed the discrete phase model in order to simulate the SCW flow containing particles. In the reactor (i.d., 10 mm, see Figure 2), the reactant flow from the middle inlet (o.d., 1.6 mm; i.d., 1.0 mm) located at the middle of the reactor is downward, while the oxygen flow from the bottom inlet (i.d. 4.4 mm) located at the bottom is upward and the outlet (i.d. 2.3 mm) is located at the top of the reactor. Therefore the flow around the middle inlet has vortex flow or recirculation. Thus we employed the realizable k-ε model17 for turbulent flow with standard wall functions18 for near-wall treatment. An advantage of the realizable k-ε model is that it provides superior performance to the standard k-ε model for flows involving recirculation and backward facing step flow.17 The density, viscosity, and thermal conductivity of water were specified in the user defined function (UDF) in FLUENT as functions of pressure and temperature following the codes based on IAPWS-IF97.19 Because FLUENT 6.3 does not allow us to specify the specific heat in the UDF, we simply input it as a piecewise function of temperature at 25.4 MPa. Because the oxidation reactors were approximate cylindrical shapes and installed vertically, the gravity condition was set. The SIMPLE pressure-velocity coupling algorithm, the standard pressure, the first order upwind discretization scheme for momentum, turbulent kinetic energy, and dissipation energy were employed in the modeling. There was little information concerning the properties of char in the supercritical water environment, such as the density and surface character, therefore, we specified simple properties of the particles as follows. We input 1000, 1500, and 2000 kg/m3 and 1-20 µm as the density and the diameter of the inert particles, respectively. Generally, char particles obtained at the outlet of the reactor precipitate in the water at room temperature and atmospheric pressure, thus the density of the char particles can be presumed to be greater than 1000 kg/m3. However, in the supercritical water environment, char particles might expand and the density could be lower than expected. Considering all these things, we decided to assume the particle density ranges 1000-2000 kg/m3. Regarding the particle diameter, 1 µm, the lowest particle size, was simply based on the pore size of the filter paper used in the experiments. In addition, we employed a smooth spherical particle model for the inert particles, although the model could lead to an underestimation of the effect of the drag force if the actual particles had a rough surface. It must be noted that at present, it is difficult to perform accurate simulations of the char particle flow in SCW.
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Figure 3. Particle distribution and residence time based on CFD simulations. The colors of the particles indicate their residence time (s) in the reactor. (Particle diameter, 10 µm; density, 1000 kg/m3; calculated up to 20 s (reactor A), 26 s (reactor B), 35 s (reactor C). The sizes of the particles in the figure are magnified.)
First, the steady-state simulation of water flow without particles was performed. The linear flow rates of water from the middle inlet and the bottom inlet were set at 0.673 and 0.010 m/s, respectively, which were calculated on the basis of the experimental conditions. The temperatures of the water at the inlets were set at 673.15 K. The operating pressure at the outlet was set constant at 25.4 MPa. After the solution of the steady-state flow was obtained, the simulation of unsteady flow involving the particles was conducted using the obtained steady-state flow solution as the initial condition. The particles were introduced from the middle inlet, the flow rate of which was 25.2 × 10-6 g/min, assuming that 1% of glucose in the 5 wt% glucose solution was converted to char particles in the upstream reactor. In this paper, we use the terms “carbon gasification efficiency”, and “carbon yield of liquid product”, which are defined as follows: Carbon gasification efficiency (C in gas) ) (carbon in the gaseous product)/(carbon in the reactant). Carbon yield of liquid product (C in liquid) ) (carbon in the liquid product)/(carbon in the reactant). Results and Discussion CFD Simulation. To investigate the char particle flow in the oxidation reactors of our SCWG reactors, CFD simulations were conducted. Figure 3 shows the simulation results, with particle size of 10 µm and particle density of 1000 kg/m3. Although the terminal velocity of a particle (Table 1) in SCW at 673 K, 25.4 MPa was lower than the average fluid flow rate at the lower part of the reactor (1.8 mm/s, in i.d., 10 mm section), which was simply calculated from the oxidant flow from the bottom inlet, some part of the particles sank to the lower part of the reactor. This was because the down flow containing particles from the middle inlet changed the bulk fluid flow at the lower part of the reactor and a down flow stream appeared along the wall. This phenomenon was not observed in the simulations without particles. Table 1 summarized the results of average
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Table 1. Average Residence Time of the Particles in the Reactors and the Ratio of Precipitated Particle Based on CFD Simulationa particle diameter (µm)
density (kg/m3)
terminal velocityb (mm/s)
average residence time of escaped particlesc (s)
escaped particles ratiod (-)
particle ratio at lower part of the reactore (-)
1 5 5 5 10 10 20 10 10
1000 1000 1500 2000 1000 1500 1000 1000 1000
0.02 0.38 0.61 0.84 1.52 2.36 5.52 1.52 1.52
15.2 15.8 15.8 16.0 16.4 16.7 22.8 31.4
0.31 0.26 0.20 0.18 0.15 0.07 0.00 0.07 0.06
0.10 0.11 0.21 0.22 0.30 0.44 0.90 0.35 0.35
reactor A
reactor B reactor C
a Particle flow calculations were conducted for 20, 26, and 35 s from the beginning of the particle flow for reactor A, B, and C, respectively. Terminal velocity of a particle in water at 673.15 K and 25.4 MPa. c “Escaped particles” means the particles flew out of the reactor from the top outlet. d The ratio of the escaped particle number against the total injected particle number at 20, 26, and 35 s in reactor A, B and C, respectively. e The ratio of the number of the particles located at more than 15 mm below the middle inlet (reactant inlet) in the reactor against the total injected particle number at 20 s in each reactor. b
Table 2. Experiment Results of Char Yield, Carbon Yield of Liquid Products, And Carbon Gasification Efficiencya oxidation reactor
char (wt%)
liquid (%)b
gas (%)c
reactor A reactor B reactor C
0.19 0.04 0.03
52.7 48.6 46.3
39.8 60.5 55.7
a Reactant: 4.9 wt% glucose solution, 5.2 g/min. Oxidant: 9.2 wt% H2O2(aq), 1.6 g/min. Reactor: pyrolysis reactor + oxidation reactor, at 673 K and 25.4 MPa. b (Carbon in the liquid product)/(carbon in the reactant) × 100. c (Carbon in the gaseous product)/(carbon in the reactant) × 100.
particle residence time and the ratio of the particle at the lower part of the reactor based on CFD simulations with several particle diameters and particle densities. The ratios of the particles at the lower part of the reactor were calculated on the basis of the particle number located more than 15 mm below the middle inlet, where the flow was not significantly affected by the strong down flow from the middle inlet. The results revealed that even though the terminal velocity was lower than the average fluid flow rate at the lower part of the reactor (the particle size less than 5 µm), some part of the particles sank to the bottom of the reactor due to the down-flow stream containing particles. With slightly higher terminal velocity than the average fluid flow rate (particle size 10 µm and particle density 1500 kg/m3), 44% of particles were precipitated; however, some parts of the particles were trapped in upflow stream and escaped from the reactor. Moreover the average residence times of the escaped particles of 1-10 µm diameter particles in reactor A were in the same range (15.2-16.7 s) due to strong upflow at the upper part of the reactor. In the case where the particle diameter was larger than 20 µm, most of the particles sank and precipitated at the bottom of the reactor. The configuration of the reactor was designed in such a manner that the heavy polymerized material (i.e., char particle) could sink to the lower part of the reactor, where the decomposition of heavy polymerized material could be enhanced due to the oxidation.9 However, the simulation with reactor A revealed that some parts of the particles precipitated, while large parts of particles smaller than 5 µm were expelled from the top outlet of the reactor. Probably, using a reactor with a larger diameter, precipitation of smaller particles can be enhanced. However, because of the size limitations of our experimental setup, we could not employ an oxidation reactor with a larger diameter. Instead, we employed longer reactors, and thus we conducted the simulations with longer reactors as shown in Table 1 and Figure 1. Moreover, the lower parts of other two reactors
were shortened, and the upper parts were extended so that the residence time of the particles could be extended effectively. The average residence time of the escaped particles (10 µm, 1000 kg/m3) from reactor B was 22.8 s, which was certainly longer than in reactor A (16.4 s). Simulation results of further extended reactor (reactor C) indicated that the residence time would be lengthened to as long as 31.4 s. It was also indicated that some of the particles precipitated to the bottom of the reactor. Interestingly, values of the particle ratio at the lower part of the reactor in reactors B and C were slightly larger than those in reactor A, even when the lower part of the reactor was shortened. We need further investigation to understand this difference in the particle ratios at the lower part of the reactor among the three reactors. Considering all these results, it is expected that char products of larger diameter than 20 µm in a real reactor will precipitate in each reactor efficiently, while some parts of the char particle of small diameter less than 10 µm will escape from the top outlet. The difference results regarding the ratio at the lower part of the reactor could imply that reactors B and C have less char product expelled from the reactors. Moreover, it can be expected that reactors B and C have less char product expelled than reactor A, because precipitated char products will react with oxidant at the bottom and decompose to smaller molecular weight materials. Longer residence time would enhance the decomposition and gasification of char and high molecular weight materials, which will lead to less high molecular weight products and larger gas yield. Experimental Verification of the Oxidation Reactors. In the previous section, we estimated the char particle flow in the oxidation reactors by means of CFD simulations. To verify the obtained results, we conducted experiments and investigated the performance of each oxidation reactor. First, we employed the pyrolysis reactor followed by each oxidation reactor, and we conducted experiments using glucose solution. Table 2 provides the results of the weight yields of char trapped on the filter papers together with the carbon yields of liquid product and the carbon gasification efficiencies. The char yields of reactors B and C were significantly lower than that of reactor A. The char products obtained from both reactors B and C were negligible in quantity, and there was little difference between their char yields. However, the char products captured from reactor B contained some visible black particles, which were not obtained from reactor C. Because the main purpose of the oxidation reactor was to decompose the char product and to depress the char forma-
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Table 3. Comparison of the Gas Compositions between the Experiment Results (exptl) and the Equilibrium Calculations (equil)a H2
CO
CH4
CO2
this work
exptl equilb
0.265 0.261
0.003 3.6 × 10-4
0.171 0.205
0.562 0.534
previous workc
exptl equil
0.536 0.586
n.d.e 9.3 × 10-5
0.025 0.002
0.439 0.411
a The values are in molar ratio. b Equilibrium gas composition calculated by STANJAN.13 c The gasification experiment in the previous work9 was conducted using 0.4 wt% glucose and lignin mixture solution at 673 K and 25.7 MPa. e CO concentration was less than the detection limit.
Figure 4. SEC chromatograms of liquid effluents obtained from the combinations of the pyrolysis reactor and each oxidation reactor. The experimental conditions were the same as those described in Table 2.
Figure 5. Gaseous products obtained from the combinations of the pyrolysis reactor and each oxidation reactor. The experimental conditions were the same as those described in Table 2.
tion, the results showed that the reactor performance was improved by the extension of the reactor length. Regarding the carbon yield of liquid products, the yields decreased by extending the reactor length. Figure 4 shows the SEC chromatogram of the liquid effluent, in which the peak of the polymerized materials appears before that of the glucose (approximately 20 min). For reactor A, the chromatogram indicates a large peak area attributed to the polymerized material, while those peaks in the chromatograms of reactors B and C were depressed. In particular, only a trace amount of polymerized material remained in the effluent from reactor C. Therefore, it can be stated that not only were the carbonaceous materials in the liquid effluent depressed, but also the polymerized materials were decomposed to lower molecular weight materials in the oxidation reactors by extending the residence time. The carbon yield of the gaseous product was increased by replacing reactor A with reactor B. Figure 5 shows the
compositions of the gas products. The H2 and CO yields of reactor B doubled from those of reactor A, while the yield of CO2 increased by only 13%. With regard to the gaseous products from reactor C, there was only a little difference between them and those from reactor B, but the yield of CO2 decreased. There is no appropriate explanation for the CO2 decrement, and we might face some problems in CO2 analysis at the moment. Except for the CO2 yields, the results of the gaseous product indicated the enhancement of glucose decomposition and gasification. All the results of the experiments with the pyrolysis reactor and the following oxidation reactor indicate that the reactor has significantly improved from the original design (reactor A), which was expected from the CFD simulation results described in the previous section. Thus it implies that CFD simulation can be a useful tool to estimate the flow of SCW containing particles. The configuration of the reactor dimension (reactor C) almost reached the maximum limit of our apparatus; therefore, further improvement in the reactor dimension will be one of the aims of our future studies. Complete Gasification of 4.9 wt% Glucose Solution at 673 K, 25.4 MPa. We have conducted a gasification experiment with a combination of the pyrolysis reactor, the oxidation reactor (reactor C), and the catalytic reactor. The gasification was operated for 10 consecutive hours. The gaseous product yield and the carbon gasification efficiency and carbon yield of liquid product are given in Figure 6. The effluent at 30 min contained less gaseous and liquid products than the expected amount. This phenomenon can be attributed to the catalyst, because gas and liquid products were obtained steadily from the beginning in the experiment without the catalytic reactor described in the previous section. It is likely that the light-molecular-weight products obtained from the oxidation reactor were adsorbed on the catalyst until the conditions in the catalytic reactor, such as the state of the catalyst surface and the reactant concentration in the bulk region, became stable.
Figure 6. Gas yield (a) and carbon yield and carbon gas efficiency (b) obtained in the experiment conducted for 10 h. Reactant: 4.9 wt% glucose solution, 5.2 g/min. Oxidant: 9.2 wt% H2O2aq, 1.6 g/min. Reactor: pyrolysis reactor + oxidation reactor + catalytic reactor, at 673 K and 25.4 MPa.
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Table 4. Comparison of the Conditions of the Supercritical Water Gasification Experiment between This Work and the Previous Work
this work previous work9 b
reactant concentration (wt%)
diluted concentrationb (wt%)
residence time in pyrolysis reactor (s)
WHSV total mass flow/catalyst weight [gtotal (gcatalyst h)-1]
WHSV reactant flow/catalyst weight [gglucose (gcatalyst h)-1]
4.9 0.4
3.75 0.309
37 30
30 38
1.1 0.12
Reactant concentration after mixed with the oxidant flow.
The gaseous products after 1 h mainly consisted of CO2, H2, and CH4, whereas the gaseous product in the effluent from the oxidation reactor shown in Figure 5 mainly consisted of CO, CO2, and H2. It is generally recognized that the water gas shift reaction shown in eq 1 proceeds in the supercritical water environment. The improvement in H2 yield indicates that the water-gas shift reaction was enhanced due to the catalytic reaction. CO + H2O f H2 + CO2
(1)
Moreover, the yield of CH4 was significantly improved, which is attributed to methanation on the catalytic surface, as follows: CO + 3H2 f CH4 + H2O
(2a)
CO2 + 4H2 f CH4 + 2H2O
(2b)
Compared with the results obtained in the previous work,9 in which the gaseous products mainly consisted of H2, CO2, and a small amount of CH4, the yield of CH4 in this study was significantly high. Table 3 provides the gas composition in the experiments and the equilibrium composition calculated by STANJAN.20 Each experimental result is approximately consistent with the equilibrium composition, which implies that the composition of the gaseous products after the catalytic reactor is mostly attributed to the thermochemical equilibrium. The average cool gas efficiency (LHV) and carbon gasification efficiency calculated from the average of the gaseous products after 4 h reached 0.61 and 0.94, respectively. Table 4 shows the comparison of experimental conditions between this experiment and the previous work.9 The value of WHSV of the reactant (i.e., glucose) in this experiment, which is the value of the reactant flow rate (g/ h) devided by the catalyst weight (g), was improved dramatically from the previous work. Future studies will be dedicated to achieving higher WHSV and to attaining complete gasification of more concentrated biomass materials and also achieving higher cool-gas efficiency by reducing the oxidant. Conclusions We have performed CFD simulations concerning the flow of SCW containing particles in reactors. In addition, we have conducted experiments on supercritical water gasification of 4.9 wt% glucose solution under the conditions of 673 K and 25.4 MPa. The results are concluded as follows. 1. The CFD simulation revealed that char particles are not likely to sink to the bottom of the original reactor that we proposed in the previous study. 2. By extending the upper part of the reactor, the simulation results implied that the residence time of the particles could be lengthened. 3. The results of the experiments with the combination of the pyrolysis reactor and the modified oxidation reactors indicated that the modified reactors were significantly improved from the original one.
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ReceiVed for reView February 8, 2009 ReVised manuscript receiVed July 6, 2009 Accepted July 8, 2009 IE9002188
