Characteristics of an Air-Blown Fixed-Bed Downdraft Biomass

A total of 10 K-type thermocouple temperature sensors were installed at various .... The uncertainty of the air/fuel ratio σAFR is affected by σAR a...
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Energy & Fuels 2008, 22, 4196–4205

Characteristics of an Air-Blown Fixed-Bed Downdraft Biomass Gasifier Chih-Lun Hsi,*,† Tzong-Yuan Wang,‡ Chien-Hsiung Tsai,§ Ching-Yuan Chang,| Chiu-Hao Liu,‡ Yao-Chung Chang,‡ and Jing-T. Kuo‡ EnVironmental Protection Administration, ExecutiVe Yuan of the Republic of China, Number 83, Section 1, Jhonghua Road, 10042, Taipei, Taiwan, Republic of China, Department of Mechanical Engineering, National Taiwan UniVersity, Number 1, Section 4, RooseVelt Road, Taipei, 10617, Taiwan, Republic of China, Department of Vehicle Engineering, National Pingtung UniVersity of Science and Technology, Number 1, Shuefu Road, Neipu, Pingtung, 912, Taiwan, Republic of China, and Graduate Institute of EnVironmental Engineering, National Taiwan UniVersity, Number 71, Choushan Road, Taipei, 10617, Taiwan, Republic of China ReceiVed January 11, 2008. ReVised Manuscript ReceiVed August 5, 2008

An air-blown fixed-bed, stratified downdraft biomass gasifier was built to investigate the key operating parameters that affect the operational characteristics of a fixed bed. The main purposes of this investigation include studying the effects of operation parameters, such as air flow rate, air preheating temperature, air/fuel ratio, and fuel moisture content, on the fuel conversion rate, specific gasification rate, producer gas heating value, and H/C ratio, as well as cold gas efficiency and hot gas efficiency. The gasifier was designed to operate with air flow in the range between 6 and 18 Nm3 h-1. Wooden cubes (15 × 15 × 15 mm) of Red Lauan and White Lauan woods were used as fuel for the gasification experiments. The burning rate of wooden cubes was found to increase with an increasing air flow rate. The producer gas heating value showed a trend of increase with an increasing air flow rate up to 15 Nm3 h-1, and a further increase in air flow resulted in a decrease in the gas heating value, because of the increase in conversion of CO and H2 with oxygen as the air flow rate was increased. The optimum mean higher heating value (HHV) of 5.68 MJ Nm-3 was obtained in this study when feeding the air unheated and using wooden cubes containing 18% moisture. Preheating the air up to 573 K can effectively shorten the time required to attain the steady-state condition but appears to have marginal effects on the producer gas heating value. The fuel moisture content can significantly lower the fuel conversion rate and the producer gas heating value. The one-dimensional steady-state model developed is able to predict the axial profiles of temperature and gas composition in a downdraft biomass gasifier with reasonable accuracy.

1. Introduction Biomass residues can be converted to a combustible gas composed of H2, CO, and CH4, by a thermochemical process of gasification, for better use as a fuel for heating and power generation. Among the gasification technologies, downdraft gasification is a simple and reliable process that can generate a producer gas with very low tar content. Consequently, downdraft biomass gasifiers are generally favored for coupling to smallscale heat engines to generate power and electricity.1 The downdraft gasifier may be divided into two categories: imbert and stratified gasifiers. A stratified downdraft gasifier does not have any restriction (throat) near the air inlet, for the ease of achieving uniform flows of fuel and air within the fuel bed. Furthermore, the throatless design can easily be scaled to larger diameters.2 Although the downdraft gasifiers are attractive for small-scale power systems, tar removal or cracking remains a * To whom correspondence should be addressed. Telephone: 886-223117722, ext. 2610. Fax: 886-2-23317741. E-mail: allanhsi@ yahoo.com.tw. † Environmental Protection Administration, Executive Yuan of the Republic of China. ‡ Department of Mechanical Engineering, National Taiwan University. § National Pingtung University of Science and Technology. | Graduate Institute of Environmental Engineering, National Taiwan University. (1) Bridgwater, A. V. Fuel 1995, 74 (5), 631–653.

major problem, and a more automated operation is required especially for small-scale industrial applications.3 Fixed-bed downdraft biomass gasifiers have been studied extensively in the past.4-9 Walawender et al.4 carried out experiments using hardwood chips with an open core downdraft gasifier (0.6 m in diameter) to investigate the effect of the fuel feeding rate on the producer gas heating value. Barrio et al.5 investigated the operation of a small-scale stratified downdraft gasifier (0.1 m in diameter and 0.5 m in height) fueled with wood pellets to study the variations of product gas composition and contaminant content as a function of the air supply rate. (2) Reed, T. B.; Das, A. Handbook of Biomass Downdraft Gasifier Engine Systems; The Biomass Energy Foundation Press: Golden, CO, 1988. (3) Maniatis, K. In Progress in Thermochemical Biomass ConVersion; Bridgwater, A. V., Ed.; Blackwell: Oxford, U.K., 2001; pp 1-32. (4) Walawender, W. P.; Chern, S. M.; Fan, L. T. In Fundamental of Thermochemical Biomass ConVersion; Overend, R. P., Milne, T. A., Mudge, L. K., Eds.; Elsevier: Amsterdam, The Netherlands, 1985; pp 911-921. (5) Barrio, M.; Fossum, M.; Hustad, J. E. In Proceedings of the Sixth International Conference on Technologies and Combustion for a Clean Environment, Oporto, Portugal, 2001; Vol. 3, pp 1269-1276. (6) Zainal, Z. A.; Rifau, A.; Quadir, G. A.; Seetharamu, K. N. Biomass Bioenergy 2002, 23, 283–289. (7) Dogru, M.; Howarth, C. R.; Akay, G.; Keskinler, B.; Malik, A. A. Energy 2002, 27, 415–427. (8) Midilli, A.; Dogru, M.; Akay, G.; Howarth, C. R. Int. J. Hydrogen Energy 2002, 27, 1035–1041. (9) Tiangco, V. M.; Jenkins, B. M.; Goss, J. R. Biomass Bioenergy 1996, 11 (1), 51–62.

10.1021/ef800026x CCC: $40.75  2008 American Chemical Society Published on Web 09/18/2008

Air-Blown Fixed-Bed Downdraft Biomass Gasifier

Zainal et al.6 did experiments to study the effect of the equivalence ratio (the actual air/fuel ratio/the air/fuel ratio for complete combustion) on producer gas composition, calorific value, and gas production rate of a downdraft gasifier. Dogru et al.7 studied gasification of hazelnut shells in a 5 kWe downdraft gasifier and showed that an optimum operation could be achieved with an air/fuel ratio between 1.44 and 1.47 Nm3 kg-1 to produce a gas of calorific value of ∼5 MJ m-3. Midilli et al.8 performed downdraft gasification of sewage sludge and obtained producer gas containing 11% hydrogen, with a maximum gross calorific value of 4 MJ m-3. Tiangco et al.9 conducted downdraft gasification experiments of rice hull for reactors with diameters in the range of 16-30 cm and specified the gasifier performance using the parameters including cold gas efficiency and specific gasification rate. In the study, it was found that the cold gas efficiency would increase as the specific gasification rate was increased to ∼200 kg m-2 h-1 and then declined with a further increase in the gasification rate. For the purposes of establishing a rational basis for the improvement of gasifier reactor design and determination of the optimum range of operating parameters, many investigators have engaged in the development of models to simulate the processes of biomass pyrolysis, gasification, and combustion.10-17 Models based on thermodynamic equilibrium were adopted to predict the producer gas composition of a biomass gasifier.10-12 Although a kinetics-free equilibrium model is adequate for the prediction of the exit gas composition, its application on the reactor design is limited.13 Most of the researchers modeled the char reduction zone as a steady state and one-dimensional problem and considered the flaming-pyrolysis region as a lumped system to calculate its temperature and gas composition based on the conservation of mass and energy and chemical reaction equilibrium14-16 or directly assumed the temperature and concentrations of the gas leaving the pyrolysis zone a priori.17 The unsteady model developed by Di Blasi13 has the most sophistication that is able to predict the dynamic behavior of a stratified downdraft gasifier but with a less desirable feature that the yield fractions of the gasified product have to be assumed a priori. A one-dimensional steady-state model was developed in this work to simulate the downdraft biomass gasification processes. One unique feature of this model is that the oxygen distribution inside the fuel bed during combustion and gasification is simulated according to the model developed by Kuo,18 which has been successfully applied to study optimum combustion-gasification process control.19 The regulation of the air supply rate is a key issue in controlling the biomass gasification to achieve a stabilized and continuous mode of operation. In the present study, the model can be applied to analyze the effect of the air flow supply rate and fuel moisture on the pyrolysis and gasification processes in the fuel bed. (10) Zainal, Z. A.; Ali, R.; Lean, C. H.; Seetharamu, K. N. Energy ConVers. Manage. 2001, 42, 1499–1515. (11) Mahishi, M. R.; Goswami, D. Y. Int. J. Hydrogen Energy 2007, 32, 3831–3840. (12) Sharma, A. Kr. Energy ConVers. Manage. 2008, 49, 832–842. (13) Di Blasi, C. Chem. Eng. Sci. 2000, 55, 2931–2944. (14) Reed, T. B.; Levie, B.; Markson, M. L.; Graboski, M. S. Symposium on mathematical modeling of biomass pyrolysis phenomena. Prepr. Pap.Am. Chem. Soc., DiV. Fuel Chem. 1983, 28 (5), 410–420. (15) Milligan, J. B. Downdraft gasification of biomass. Ph.D. Dissertation, University of Aston, Birmingham, U.K., 1994. (16) Jayah, T. H.; Aye, L.; Fuller, R. J.; Stewart, D. F. Biomass Bioenergy 2003, 25, 459–469. (17) Giltrap, D. L.; McKibbin, R.; Barnes, G. R. G. Sol. Energy 2003, 74, 85–91. (18) Kuo, J. T. Combust. Sci. Technol. 1998, 137, 1–29. (19) Hsi, C. L.; Kuo, J. T. Biomass Bioenergy 2008, doi: 10.1016/ j.biombioe.2008.03.008.

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Figure 1. Gasification system: (1) silo, (2) screw feeder, (3) gas damper, (4) air distributor, (5) reactor, (6) rotameter, (7) air blower, (8) ash collector, (9) producer gas sampling port, (10) condensates collector, (11) cooling-water heat exchanger, and (12) igniter.

While there are plenty of studies reporting the producer gas composition at the exit of the reactors as functions of operating parameters, data of the gas composition inside the gasification reactor are still limited. The conditions of the thermochemical processes going on inside a downdraft gasifier can significantly affect the rates and characteristics of the producer gas.13,20 The gas concentration distribution inside the fuel bed reflects the characteristics of combustion and gasification reactions in the reactor. In this work, an air-blown fixed-bed, stratified downdraft biomass gasifier was build for experimental investigation and model validation. The purposes of the investigation include studying the effects of operation parameters, such as air flow rate, air preheating temperature, air/fuel ratio, and fuel moisture content, on fuel conversion rate, specific gasification rate, producer gas heating value, and H/C ratio, as well as the cold gas efficiency and hot gas efficiency. The axial profiles of temperature were measured; gas samples were extracted and analyzed along the fuel bed; and the results were used to validate the model developed in this study. 2. Experimental Setup and Procedure 2.1. Experimental Setup. The downdraft biomass gasification system designed in this work is schematically represented in Figure 1. The system mainly comprises a reactor, fuel feeding system, air blower, an electric air heater, igniter, producer-gas-cooling heat exchanger, condensate water collector, and ash collector. The reactor is a cylinder with an internal diameter of 200 mm and height of 1000 mm, and there is no throat or constriction device in the reactor. The thickness of the reactor vessel wall is 140 mm, and the wall structure, from inside to outside, consists of a refractory lining (50 mm), steel cylinder (10 mm), a blanket of rock wool (30 mm), and a cylindrical steel outer shell (10 mm). To further reduce heat loss, a blanket of fiberglass fabric (40 mm) is wrapped over the outer steel shell. A total of 10 K-type thermocouple temperature sensors were installed at various locations above the grate. Thermocouple CH01 was located at the height of 930 mm above the grate. The other nine thermocouples were spaced 100 mm apart, and the lowest thermocouple (CH10) was located at a distance of 30 mm above the grate. The gasifier is not equipped with a pollutant emission (20) Ross, D.; Noda, R.; Horio, M.; Kosminski, A.; Ashman, P.; Mullinger, P. Fuel 2007, 86, 1417–1429.

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Table 1. Ultimate and Proximate Analyses of Fuel Used in This Work Red Lauan ultimate analysis, dry basis C (%) 45.85 H (%) 5.98 O (%) 48.05 N (%) 0.12 proximate analysis, dry basis combustible matter (%) 99.91 ash content (%) 0.09

White Lauan 45.11 6.35 48.48 0.06

from operation procedures. Total uncertainties of quantities, such as the fuel conversion rate, rates of ash and liquid condensate production, specific gasification rate, gas production rate, air/fuel ratio, and cold and hot gas efficiencies, were calculated using the measured data by eqs 1 and 221,22

σxi2 ) Bxi2 + Pxi2 ∂r ∑ ( ∂x )σ j

99.57 0.43

control system such as a filter and scrubber. A part of tars and moisture were collected by a cooling heat exchanger located downstream the exit of the reactor. The gasification system was operated on a higher-than-atmospheric positive pressure, and the off gas was burned in a flare. 2.2. Experimental Procedure. Wooden cubes made of Red Lauan (RL) and White Lauan (WL) woods of 15 × 15 × 15 mm were used as fuel for the air-blown gasifier. The average RL wood density with 18% moisture content (wet basis) is 790 kg m-3. The average WL wood density (dry basis) is 563.5 kg m-3. The ultimate and proximate analyses of RL and WL woods were listed in Table 1. The test conditions for the selected test runs of the downdraft gasifier are listed in Table 2. Test run numbers 1-5 were performed with unheated air and RL wood cubes containing 18% moisture. Test run numbers 6 and 7 were carried out with the air preheated to 473 and 573 K, respectively. The RL woods used in test number 8 were oven-dried at 378 K for 96 h before the test. The wooden cubes used in test number 9 were initially oven-dried by the same procedure as that used in test number 8, bathed in water for 7 days and then dried by natural convection to the moisture content of ∼33% (wet basis). Before the tests, the wooden cubes were sealed in plastic bags to preserve the moisture content. Tests of run numbers 10-15 were conducted with WL cubes at air flow rates varying between 7 and 17 Nm3 h-1. The wooden cubes were fed into the reactor from the top of the gasifier. The fuel bed was ignited by an igniter located at a distance of 130 mm above the grate. The bed level was maintained above the air inlet port and constantly monitored using a photoelectric sensor. The amount of wooden cubes introduced into the reactor was weighed for estimation of the average fuel consumption rate. The air for gasification was force-fed by a blower, and the flow rate was metered by a rotameter and regulated with a valve. The tests were conducted for over a period of at least 2 h after the stable condition was achieved for gas sampling and analysis. Once the stable condition was established, the gas samples were taken with sampling bottles vacuumed by a suction pump. The gas samples were drawn at positions of the same heights with different meridian angles than those of the thermocouples, and two samples of gas were taken at each location. At the exit of the reactor, as shown in Figure 1, at least 20 samples were obtained for each test condition for determining the producer gas composition and heating value. Although the collected raw gas was not cleaned with a cleanup device, the gas was filtered by glass wool to remove tar and particulates before introduction to gas chromatography analysis. The collected gas samples were analyzed for its composition, which include CO, CO2, CH4, H2, N2, and O2. At the end of the tests, the ash and liquid condensates were collected and weighed for the estimation of gas production rate according to the mass analysis balancing the difference between the mass of air and fuel introduced and the mass of collected ash and condensates. 2.3. Uncertainty Analysis. Identifying the sources of errors affecting the measurement processes and quantifying the uncertainties is essential for improving the precision and accuracy of an experimental work. A detailed calculation procedure of uncertainty analysis related to sewage sludge gasification was presented by Dogru et al.21 In this work, the uncertainties in the measurement of temperature, air flow rate, and gas composition were determined by eq 1 considering the measuring principles, precision, and resolution of the individual instruments, as well as the errors arising

σr2 )

i)1

2

i

xi

(1) 2

(2)

where r ) r(x1, x2,..., xj) is a function of j measured variables xi, σr and σxi are the uncertainties of r and xi, respectively, and Bxi and Pxi are the systematic uncertainty and random uncertainty of the variables xi, respectively. Systematic uncertainty and random uncertainty are defined as the squares root of the sum of the uncertainties of various elemental sources. • Temperature: measured by thermocouples (K type) with an accuracy of (0.75%; uncertainty arising from operation is (1.5%. Consequently, the total uncertainty in the temperature measurement σT is estimated to be (1.68% by eq 1. • Air flow rate: measured by a rotameter with a full scale of 300 NL min-1 and a least count of 10 NL min-1 and a specified accuracy of (5%; uncertainty caused from the reading is (1%. The total uncertainty in the air flow rate measurement σAR is calculated to be (5.10% by eq 1. • Gas composition analysis: measured by a gas chromatographer with an accuracy of (2%; a rather high uncertainty is estimated to be (5% caused by the manual procedure of sampling and injection. Thus, the total uncertainty of gas composition σGC as well as the uncertainty of gas heating value σGHV is (5.39%, estimated by eq 1. • Fuel conversion rate and liquid condensate rate: the amount of fuel of each batch fed into the silo as well as the amount of unused fuel of the last batch at the end of operation were weighed by an electronic balance with a full scale of 30 kg and a least count of 1 g to calculate the total amount of fuel fed into the reactor during the steady-state operation. Uncertainties from the resolution of electronic balance and from weighing of fuel are ∼ (0.16 and (1.5%, respectively. Thus, the uncertainty in measuring the amount of fuel introduced σfuel is (1.51%, estimated by eq 1. The uncertainty of the timer σtimer is (1%. The fuel conversion rate is defined as

m ˙ fuel ) amount of fuel fed into the reactor/ total operation time of steady-state period (3) Consequently, the total uncertainty of the fuel conversion rate σFCR is affected by the uncertainty arisen from determining the amount of fuel σfuel and the uncertainty of the operation time σtimer. It could be derived21 from eq 2 that the total uncertainty of fuel conversion rate σFCR is equal to ((σfuel2 + σtimer2)0.5 ) (1.81%. Liquid condensate was collected from the condensates collector and was weighed by the same electronic balance, the uncertainties deduced from measurement of the liquid amount and from the timer were estimated to be (1.5 and (1%, respectively. The total uncertainty of measuring the liquid condensate rate is (1.81%. • Ash production rate: the amount of ash was collected from the ash collector at the end of the tests and was weighed by a micro-electronic balance with a full scale of 120 g and a least count of 0.1 mg. The uncertainties from the resolution of the micro-electronic balance and from weighing of ash are (0.000 25 (21) Dogru, M.; Midilli, A.; Howarth, C. R. Fuel Process. Technol. 2002, 75, 55–82. (22) Coleman, H. W.; Steele, W. G. Experimentation and Uncertainty Analysis for Engineers; Wiley: New York, 1999.

Air-Blown Fixed-Bed Downdraft Biomass Gasifier

Energy & Fuels, Vol. 22, No. 6, 2008 4199 Table 2. Test Conditions

test number

wood speciesa

wood moisture content (wt %)

fuel conversion rate (kg h-1)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

RL RL RL RL RL RL RL RL RL WL WL WL WL WL WL

18 18 18 18 18 18 18 0 33 18.5 28.8 24.5 30.5 22.3 24

2.72 4.40 6.02 8.18 8.90 6.11 6.75 8.42 4.47 2.75 4.09 5.72 7.22 9.10 9.48

a

air preheat temperature (K)

473 573

air flow rate (Nm3 h-1)

air/fuel ratio (Nm3 kg-1)

dry producer gas HHV (MJ Nm-3)

6 9 12 15 18 12 12 12 12 7 9 11 13 15 17

2.21 2.05 1.99 1.83 2.02 1.96 1.78 1.43 2.68 2.55 2.20 1.92 1.80 1.65 1.79

4.47 5.16 5.42 5.68 4.64 5.49 5.42 5.98 4.77 4.56 4.63 4.78 4.79 5.09 4.94

RL and WL denote Red Lauan and White Lauan woods, respectively.

and (1%, respectively; the uncertainty of the timer is (1%. The total uncertainty of ash production rate is (1.41%, estimated by eqs 1 and 2. • Specific gasification rate: because the specific gasification rate is defined as eq 49 BG ) m ˙ fuel/AG

(4)

thus, the total uncertainty of the specific gasification rate σSGR is affected by the uncertainty of the fuel conversion rate σFCR and the uncertainty arisen from the calculation of the grate area (σAG ∼ (1%); σSGR is calculated to be (2.07% from eq 2. • Air/fuel ratio: because the air/fuel ratio is defined as eq 5 air/fuel ratio ) F˙a/m ˙ fuel

(5)

where F˙a is the air flow rate fed into the reactor. The uncertainty of the air/fuel ratio σAFR is affected by σAR and σFCR and is estimated to be (5.41% from eq 2, that is, the square root of the sum of σAR and σFCR. • Gas production rate: the dry gas production rate F˙g,dry is derived by means of the nitrogen balance between the inlet and outlet of the reactor, as given in eq 6 F˙g,dry ) F˙a(3.76/4.76)/yN2

(6)

where yN2 is the nitrogen mole fraction of the dry producer gas. Consequently, its uncertainty σGPR is affected by σAR and σGC and is estimated to be (7.42% from eq 2. • Gas efficiencies: cold gas efficiency and hot gas efficiency are calculated as ηcold )

gas production rate × LHV of producer gas fuel conversion rate × LHV of fuel

(7)

gas production rate × LHV of producer gas + sensible heat of hot gas (8) ηhot ) fuel conversion rate × LHV of fuel The uncertainties of cold and hot gas efficiencies are (9.40 and (8.92%, respectively, which were determined by eq 2, associated with the uncertainties of the variables listed on the right-hand side of eqs 7 and 8. The estimated total uncertainties of variables are listed in Table 3. It is seen that the uncertainties of the gas production rate and gas efficiencies are rather higher than ∼ (7%. These high uncertainties mainly resulted from the inaccuracy of the rotameter

Table 3. Total Uncertainties variables

uncertainties (%)

temperature air flow rate gas composition fuel conversion rate liquid condensates rate ash production rate specific gasification rate air/fuel ratio gas production rate cold gas efficiency hot gas efficiency

(1.68 (5.10 (5.39 (1.81 (1.81 (1.41 (2.07 (5.41 (7.42 (9.40 (8.92

used in this work, as well as from the uncertainty of the gas composition measurement conducted with manual sampling and injection.

3. Modeling On the basis of the consideration that the burning rate of the solid fuel bed is related to the consumption rate of oxygen as the air is flowing through the fuel bed, Kuo18 developed a model for the estimation of the burning rate of refuse bed in a combustion furnace. The gross reaction of solid fuel in the pyrolysis and flaming combustion regions can be described as follows:18 CaHbOc · f(H2O) + rˆofO2 f a(1 - rCO)CO2 + arCOCO + (f + b ⁄ 2)H2O (9) where rˆof denotes the oxygen/fuel mole ratio and rCO denotes the fraction of carbon converted to carbon monoxide. Assuming the oxygen concentration in the reactor is varying one-dimensionally along the fuel bed, for the diffusioncontrolled reaction under steady-flow conditions, the oxygen distribution in the fuel bed can be represented as follows:18 dXO2

j XO n(1 - arXO ) ) -K (10) 2 2 dy* j ) -(Rd/ar)ln(1-arXO2), y* ) y/L, where L denotes where K the bed depth, ar is the ratio of change in gaseous moles to the change in oxygen moles, and n is the order of the reaction. Rd denotes the dimensionless reaction rate factor18 Rd ) apL

jD

(11) Sc2⁄3 where ap ) 6(1 - ε)/dp is the fuel bed solid surface area per unit volume. ε denotes the void fraction of the fuel bed; jD is

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the Chilton-Colburn j factor for oxygen transfer; and Sc is the Schmidt number. For packed beds, the value of jD is ∼0.1 and that of Sc is approximately of the order of 1.18 The void fraction of the fuel bed in this study is approximately 0.5, and Rd is estimated to be ∼20. The mole ratio of the rate of oxygen depletion and rate of solid fuel gasified and burned according to eq 9 can be described as

( )( ) ( )

dXO2 Fgug r˙FP / ) rˆof ˆg ˆ fuel dy M M

(12)

Therefore, the corresponding fuel combustion rate is expressed as follows: r˙FP ) -

dXO2 M ˆ fuel 1 Fgug ˆg rˆof dy M

(13)

Kuo and Hwang23 found that during pyrolysis or flaming combustion of a single wooden sphere, the atomic ratios H/O of the remaining solid varied within a narrow range between 2.2 and 2.6. Therefore, the H/O atomic ratio of wooden cubes was assumed to be approximately constant during pyrolysis and flaming combustion. Experimental results of combustion of single wooden particles also suggest the C/H atomic ratio of the wood could be assumed to vary linearly with the conversion of the solid during thermal decomposition. In the gasification region where the char is converted into gases, the endothermic reactions involved in this conversion are expressed as follows:24 C + CO2 f 2CO ∆hG,CO2 ) -14372 kJ/kg

(14)

C + H2O f CO + H2 ∆hG,H2O ) -14609 kJ/kg

(15)

The carbon conversion rates for the reactions of eqs 14 and 15 are taken from24 ˆ C/M ˆ CO )hDkG,CO /(hD + kG,CO ) r˙G,CO2 ) FCO2ap(M 2 2 2

(16)

ˆ C/M ˆ H O)hDkG,H O/(hD + kG,H O) r˙G,H2O ) FH2Oap(M 2 2 2

(17)

where hD ) 0.8Dg(2.0 + 1.1Re0.6Sc1/3)/dp is the mass-transfer ˆ denotes the molecular weight. The symbols coefficient and M kG,CO2 and kG,H2O, given in eqs 16 and 17, representing the rates of reaction, are described in the Arrhenius form by kG,CO2 ) 3.42Ts exp(-1.56 × 104/Ts) and kG,H2O ) 1.67kG,CO2, respectively. An exothermic reaction that involves char and hydrogen to produce methane is considered in this model. C + 2H2 f CH4 ∆hG,H2 ) 74.93 kJ/mol

(18)

The reaction rate is given by r˙G,H2 ) apCH2hDkG,H2/(hD + kG,H2)

(19)

where CH2 denotes the hydrogen molar concentration and hD denotes the mass-transfer coefficient; they can be estimated using the following correlations:25 hD ) 2.06(1/ε)ugRe-0.575Sc-2/3

(20)

kG,H2 ) 104 exp(-26095/Ts)

(21)

(23) Kuo, J. T.; Hwang, L. H. Combust. Sci. Technol. 2003, 175, 665– 693. (24) Bryden, K. M.; Ragland, K. W. Energy Fuels 1996, 10, 269–275. (25) Di Blasi, C. AIChE J. 2004, 50 (9), 2306–2319.

The water-gas shift reaction is included in this study and is considered to be reversible, as shown in25 CO + H2O f CO2 + H2 ∆hWGS ) 41.2 kJ/mol

(22)

The kinetic rate of the water-gas shift reaction r˙WGS is obtained from25 r˙WGS ) εkWGS(CCOCH2O - CCO2CH2/KE)

(23)

kWGS ) 2.78 exp( - 1513/Tg)

(24)

KE ) 0.0265 exp(3966/Tg)

(25)

The balances of mass and energy conversion between the solid and gas phases are described in one-dimensional form along the axial direction of the downdraft gasification reactor. The mass balance of the solid phase, gas phase, and gaseous species are given as follows: d(Fsus) ) -(r˙FP + r˙G,CO2 + r˙G,H2O + r˙G,H2) dy

(26)

d(Fgug) ) r˙FP + r˙G,CO2 + r˙G,H2O + r˙G,H2 dy

(27)

d(Fg,iug,iYg,i) ) dy

∑b

˙j i,jr

(28)

j

where Y denotes the mass fraction of the gas species i (i ) CO, CO2, CH4, H2O, and O2) and bi,j is the yield of species i from reaction j. The energy balances of the solid and gas phases within the fuel bed could be formulated as follows:

( )

d(Fsushs) d dTs ) k + hsgap(Tg - Ts) + 4hws(Tw - Ts)/D + dy dy s dy r˙G,CO2∆hG,CO2 + r˙G,H2O∆hG,H2O + r˙G,H2∆hG,H2 + 0.5r˙FP∆hFP (29)

( )

d(Fgughg) dTg d ) kg + hsgap(Ts - Tg) + 4hwg(Tw - Tg)/D + dy dy dy r˙WGS∆hWGS + 0.5r˙FP∆hFP (30) where D denotes the reactor internal diameter and hsg ) 2.06(1/ ε)CpgFgugRe-0.575Pr-2/3 is the heat-transfer coefficient between the solid and gas phases.26 The wall heat-transfer coefficients hws and hwg are also obtained from ref 26. Characterizing the thermochemical energy release processes during flaming pyrolysis as the flaming combustion of the fuel particle that is fueled by the pyrolysis gasification of internal fuel mass and sustained by the heat release from external combustion, ∆hFP, is bisected in equal proportion to the solid and gas phases. The higher heating value and lower heating value of producer gas as well as the H/C ratio of combustible gas are defined as follows: HHV of gas ) yH2 × 12.769 + yCO × 12.622 + yCH4 × 39.781 MJ Nm-3 (31) LHV of gas ) yH2 × 10.788 + yCO × 12.622 + yCH4 × 35.814 MJ Nm-3 (32) (26) Hobbs, M. L.; Radulovic, P. T.; Smoot, L. D. Prog. Energy Combust. Sci. 1993, 19, 505–586. (27) Waldheim, L.; Nilsson, T. Heating value of gases from biomass gasification. Report prepared for IEA Bioenergy Agreement, Task 20Thermal Gasification of Biomass, 2001 (available on the website http:// www.gastechnology.org/webroot/app/xn/xd.aspx?it)enweb&xd)iea/publications.xml).

Air-Blown Fixed-Bed Downdraft Biomass Gasifier

Energy & Fuels, Vol. 22, No. 6, 2008 4201

H/C of gas ) (yH2 × 2 + yCH4 × 4)/[(yCO + yCH4) × 12] (33) where yi values are mole fractions of H2, CO, and CH4 of the dry producer gas, and the heating value of gaseous components are cited from ref 27. 4. Results and Discussion Figure 2 shows the temperature histories measured at various locations inside the gasifier in test number 14. Because the fuel bed was ignited by an igniter located at a distance of 130 mm above the grate, the temperature at this location rose first up to the peak value of 1160 K in about 30 min and then declined to ∼850 K because of the endothermicity of the gasification reactions. The patterns of temperature history at 30-430 mm above the grate are similar. The temperature at the position of the air inlet, i.e., 530 mm above the grate, reached the first peak value ∼1430 K at ∼130 min time mark and fluctuated between 950 and 1460 K, caused by the feeding of the wooden cubes. It is evident that the flame front propagated upward from the bottom of the fuel bed to the location of the air inlet resulted from the fact that the flaming combustion rate is significantly higher than the char gasification rate. When the time reached the range of 120-180 min, the entire fuel bed was in steadystate condition. The axial temperature profiles of the fuel bed recorded at various stages of test number 14 is shown in Figure 3. The peak temperature stayed at the position of the air inlet after the flame front reached this height, indicating that the downdraft gasifier was operated in the top-stabilized mode. The temperature measured at the top region of the reactor and that at a height of 773 mm above the grate are generally lower than 400 K, which implies that the wooden cubes in this region were under the moisture evaporation stage. The wooden cubes near 630 mm above the grate began to thermally decompose when the temperature fluctuated between 450 and 550 K. During the steady-state operation, the fuel particle was observed flaming at the height of 530 mm and glowing at 430 mm above the grate. Viewing the combustion condition inside the gasifier through the thermocouple port and judging the pattern of temperature profiles recorded in the above grate region, it can be concluded that, when a steady condition was established, the fuel bed was stratified in sequence into drying, pyrolysis, flaming-pyrolysis combustion, and char gasification regions from

Figure 2. Temperature histories at various locations inside the fuel bed in test number 14.

Figure 3. Axial temperature profiles at various stages inside the fuel bed in test number 14.

the top to the bottom of the reactor. The other test runs listed in Table 1 qualitatively presented the same fuel bed stratification pattern. Because the air was fed into the reactor at a height of 530 mm above the grate rather than from the top of the reactor and the measured temperature of the fuel bed was generally lower than 550 K (Figure 2) in the upper region at the height of 630 mm, the upper boundary of the flaming-pyrolysis zone was set at the position of 630 mm above the grate. Equations 26-30 were integrated from y* ) 0 downward, starting from the upper boundary of the flaming-pyrolysis zone for simulation of the combustion and gasification processes in the fuel bed. The initial solid velocity is estimated from the measured fuel feeding rate by dividing it with the product of the grate area and wood density. The boundary condition of the energy equation at the grate (y* ) 1) is set by assuming the bottom surface of the fuel bed radiating to the surroundings at a known temperature. The rate of wood drying in the region above y* ) 0 can be analyzed by eq 34,28 as follows: r˙m ) Am(Fs - Fs,dry)exp(-EmR/Ts)(1 - ε) h-1

(34) kmol-1.28

where Am is 2 × and Em is 8.78 × kJ Figures 4 and 5 illustrate the measured and calculated profiles of the temperature and dry producer gas composition of test number 14 under the steady-state condition. The oxygen distribution of the product gas predicted by the model with different values of n is given in Figure 6. The measurement showed that the oxygen was rapidly depleted within a narrow flaming combustion region near the air inlet, as indicated in Figure 5; thus, the order of the reaction in eq 10 was selected as n ) 0 to simulate the oxygen profile in the fuel bed. A constant value of rCO ) 0.65 was determined to achieve the best agreement between the predicted and measured CCO and CCO2 in the flaming combustion region. The agreement between the measured and predicted temperatures is good, except for the scattering at the location of the air inlet, which was attributed to the disturbance caused by fuel feeding, as mentioned previously. The increase in the carbon monoxide and carbon dioxide concentrations in the thin flaming combustion region, as reported in Figure 5, is mainly attributed to the result of the pyrolysis and flaming combustion reactions. In the gasification 1010

104

(28) Purnomo Aerts, D. J.; Ragland, K. W. Proc. Combust. Inst. 1990, 23, 1025–1032.

4202 Energy & Fuels, Vol. 22, No. 6, 2008

Figure 4. Measured and predicted axial temperature distributions of test number 14.

Figure 5. Measured and calculated gas compositions of test number 14.

Figure 6. Oxygen distribution in the fuel bed for various orders of reactions.

region, the hydrogen concentration increased continuously, resulting from both the char-H2O gasification (eq 15) and the water-gas shift reactions. The measurement shows that CO

Hsi et al.

increases as a result of the gasification reactions (eqs 14 and 15) and then slightly decreases because of the forward water-gas shift reaction. The concentrations of CH4 and O2 in the gasification zone were underestimated. Because the air is regulated at a level lower than the stoichiometric rate for complete combustion, the oxygen found in the gasification zone could either be due to the downward mixing inlet air or the leakage caused during the gas sampling. In general, the discrepancy between the measured and calculated producer gas compositions is partially due to the fact that the unsteady single particle reactions are not precisely modeled with uncertain wood property data and lacking appropriate rate constants of kinetics. Nevertheless, the trends exhibited by the estimated dry gas composition profiles are reasonably simulated. The calculated dry producer gas HHV at the outlet of the reactor is 4.74 MJ Nm-3, which is close to the HHV of 5.09 MJ Nm-3 obtained from the measured gas composition. The fuel conversion rates listed in Table 2 could be converted to the specific gasification rates as expressed in eq 4. The specific gasification rate is between 86.6 and 301.9 kg m-2 h-1, as shown in Table 4. At maintained fuel moisture content, the fuel conversion rate would increase with an increase in the air flow rate. The air/fuel ratio of the test runs showed a trend of decrease with an increase in the air flow rate, owing to the fact that a larger amount of fuel is burned in the flaming combustion zone as a result of the increase in air feeding. Figure 7 confirms the proportionality between the specific air flow rate and the specific gasification rate. However, for test run numbers 1-5, the higher heating value of dry producer gas increased to 5.68 MJ Nm-3 with an increase in the rate of unheated air from 6 to 15 Nm3 h-1 and then sharply declined to 4.64 MJ Nm-3 as the air feeding rate was further increased to 18 Nm3 h-1, as shown in Table 2. Producer gas composition at the exit of the reactor is itemized in Table 5. In Figure 8, the dry gas composition and the corresponding HHV under steady conditions in test runs numbers 1-5 are illustrated as functions of the air flow rates. The concentrations of CO and H2 increased with an increase in the air flow rate from 6 to ∼15 Nm3 h-1 and then decreased with a further increase in air flow, whereas the concentrations of N2 and CO2 showed the opposite trends. This implies that a larger amount of CO and H2 would be combusted with oxygen and was converted to CO2 and H2O when the air flow was increased and as a result the degradation of the product gas heating value. Therefore, maintaining the optimum air flow rate is essential for obtaining high-quality producer gas. Despite the fact that the moisture content of the producer gas was not measured in this study, the dry gas production rate can be estimated by means of the nitrogen balance between the inlet and outlet of the reactor, allowing the gas efficiency to be calculated. Table 4 indicates that the cold gas efficiency varied within 70 and 89% and the hot gas efficiency ranged between 74 and 96%. Moreover, the dependence of gas efficiency on air flow rate was quite similar to the relation between gas heating value and air flow. That is, for test numbers 1-5, the cold gas efficiency increased as the feeding rate of air increased until the optimum value of 88% was reached and then steeply declined to 70% as the air flow rate increased to 18 Nm3 h-1. Within the range of gasification rates studied in this work, the corresponding air/fuel ratio is between 1.43 and 2.68 Nm3 kg-1. The producer gas heating value is shown to be inversely proportional to the air/fuel ratio in Figure 9. A dependence of the heating value on the producer gas H/C ratio is also evident. Combustible gas with higher H/C ratios was obtained at a lower air/fuel ratio. The ratios of H/C of combustible gas produced

Air-Blown Fixed-Bed Downdraft Biomass Gasifier

Energy & Fuels, Vol. 22, No. 6, 2008 4203 Table 4. Gasification Results

test number

air flow rate (Nm3 h-1)

air/fuel ratio (Nm3 kg-1)

specific gasification rate (kg m-2 h-1)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

6 9 12 15 18 12 12 12 12 7 9 11 13 15 17

2.21 2.05 1.99 1.83 2.02 1.96 1.78 1.43 2.68 2.55 2.20 1.92 1.80 1.65 1.79

86.6 140.1 191.6 260.4 283.3 194.5 214.9 268.0 142.3 87.5 130.2 182.0 223.0 289.5 301.9

specific gas production rate (Nm3 m-2 h-1)

dry gas HHV (MJ Nm-3)

dry gas LHV (MJ Nm-3)

H/C ratio of combustible gas (kg/kg)

cold gas efficiency (%)

hot gas efficiency (%)

292.5 471.6 650.3 852.0 910.6 651.9 668.9 728.8 576.5 330.7 453.8 588.5 724.0 852.6 940.2

4.47 5.16 5.42 5.68 4.64 5.49 5.42 5.98 4.77 4.56 4.63 4.78 4.79 5.09 4.94

4.18 4.80 5.05 5.23 4.34 5.07 5.02 5.56 4.42 4.27 4.31 4.40 4.36 4.68 4.53

0.116 0.136 0.135 0.172 0.119 0.158 0.156 0.139 0.147 0.113 0.134 0.176 0.223 0.180 0.191

72.1 82.9 86.1 88.0 70.3 89.0 79.0 86.2 86.1 88.6 79.3 73.5 70.0 71.3 73.0

77.1 89.5 94.2 96.5 74.4 94.5 85.3 90.0 90.3 93.0 85.6 80.0 78.6 79.9 83.6

in the downdraft wood gasifier were between 0.113 and 0.223. The air/fuel ratio reflects the amount of air used to gasify the wood fuel; it would be desirable to keep it as low as possible for better conversion efficiency. However, when air flow is too low, the gasification process will not be self-sustainable, meaning that the heat generated in combustion would not be able to sustain the gasification reactions. On the other hand, when the air flow is excessive, the reactions in the fuel bed would switch from less gasification to more combustion and

this will lead to the dilution of producer gas and a decrease in the heating value. Ideally, the air flow should be regulated at a level to maintain a proper air/fuel ratio to keep the fuel bed reactions in top-stabilized mode and obtain optimum concentrations of H2 and CO and a heating value of the producer gas. Table 2 shows that the optimum air flow to operate the gasifier

Figure 8. Dry gas composition and higher heating values as functions of the air flow rate in test run numbers 1-5. Figure 7. Specific gasification rate as functions of the specific air flow rate. Table 5. Dry Gas Composition and the Residues Collected residues (kg h-1)

dry gas composition (vol %) test number

H2

O2

N2

CO

CO2

liquid CH4 condensates

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

10.99 15.15 16.45 19.85 12.13 17.80 17.20 17.47 14.59 11.16 13.23 16.43 19.05 18.30 17.99

1.44 1.31 1.88 1.10 1.47 1.37 1.61 1.57 0.97 2.25 2.31 2.38 2.76 1.93 2.49

56.34 52.25 51.51 48.25 55.02 49.56 49.94 48.24 53.62 55.79 54.46 52.42 51.07 49.46 50.73

18.93 20.74 22.19 20.49 19.68 20.52 20.28 23.68 18.39 19.64 18.67 16.79 14.83 17.94 16.68

10.60 9.04 6.66 8.90 10.18 9.18 9.29 7.14 10.97 9.51 9.86 10.58 11.05 11.14 10.75

1.69 1.52 1.31 1.41 1.52 1.58 1.68 1.91 1.47 1.64 1.48 1.41 1.22 1.23 1.36

0.710 1.181 1.573 1.899 2.024 1.403 1.534 0.579 1.803 0.840 1.740 1.603 1.680 1.755 2.043

ash 0.010 0.011 0.013 0.019 0.014 0.013 0.014 0.019 0.013 0.016 0.016 0.012 0.022 0.042 0.017

Figure 9. Dry producer gas HHV and H/C ratio of combustible gas as functions of the air/fuel ratio.

4204 Energy & Fuels, Vol. 22, No. 6, 2008

Figure 10. Variation of the specific gas production rate as a function of the air flow rate.

is ∼15 Nm3 h-1. At this flow, a lower air/fuel ratio or a higher fuel conversion rate could be obtained by drying the feedstock and preheating the air. The specific gas production rate, defined as the rate of wet gas production per unit grate area, has been used as a parameter to specify the gasification reactor dimension and capacity.2 Figure 10 shows that the specific gas production rate is linearly proportional to the specific air flow rate and varied in a range between 292 and 940 Nm3 m-2 h-1. When the downdraft gasifier was operated under top-stabilized condition, a higher specific gas production rate was achieved at a lower air/fuel ratio, signifying that the gasification rate would increase with an increase in the air supply rate. The turn-down ratio is a parameter used in the scaling of gasifiers and is defined as the ratio of the highest practical gas production rate to the lowest rate.2 Although the turn-down ratio has not been determined directly in this work, tests have been carried out to investigate the variation of the fuel consumption rate with the air supply rate. Adjusting the air flow rate could achieve the regulation (turn-down or turn-up) of the fuel combustion/gasification rate. A linear correlation between the fuel conversion rate and air flow rate was observed (see Figure 7). Ash discharged and liquid condensates collected at the drainage port of the water-cooled heat exchanger were given in Table 5. The ash production rate varied between 0.002 and 0.006 kg of ash/kg of fuel and slightly declined as the air flow rate is increased. The percentage of moisture contained in producer gas can be estimated from the nitrogen balance between the inlet and outlet of the reactor. The calculated result shows that condensation occurring in the heat exchanger would remove about 50% of the moisture in the producer gas. Test runs numbers 6 and 7 were conducted with air preheated to the temperatures of 473 and 573 K, respectively. The heated air stream promotes thermal radiation and propagation of ignition front in a downdraft gasifier.29 The flame spreading speed, which specifies the propagation of ignition front, is defined as the average speed of the flame front propagating from the bottom of the fuel bed to the location where the top-stabilized flame front is maintained. During the gasifier start-up period, the preheated air enhances thermal radiation, speeds up the (29) Saastamoinen, J. J.; Horttanainen, M.; Sarkomaa, P. Combust. Sci. Technol. 2001, 165, 41–60.

Hsi et al.

evaporation of moisture in the woods, and shortens the time to achieve a steady state. Preheating the air to 473 and 573 K accelerated the flame spreading speed to 0.086 and 0.128 mm s-1, respectively, in comparison to the result of test 3 (0.067 mm s-1), where the air used was not preheated. The time required to attain the steady condition of the gasifier operated with air at 473 and 573 K was ∼100 and 90 min, respectively, which were shorter than the time required to reach the steady situation for test number 3 (∼120 min). On the other hand, air preheating appeared to have a marginal effect on the heating value of the producer gas. As demonstrated in Table 1, comparing the results of test run numbers 3 and 6, the mean higher heating value of the producer gas increased slightly from 5.42 to 5.49 MJ Nm-3 when the air was preheated from room temperature to 473 K. The elevation of the sensible heat of the producer gas by preheating the air streams has only a marginal effect on the producer gas heating value as compared to the effect of varying the air flow rate. Jayah et al.16 simulated the operation of a downdraft wood gasifier and observed that the conversion efficiency slightly increased from 56 to 57% when the inlet air temperature was increased from 300 to 600 K. The moisture content in the feedstock can affect the temperature in the fuel bed and consequently influence the heating value of the producer gas in a downdraft biomass gasifier. The wooden cubes used in test number 8 were oven-dried at 378 K for a period of 96 h to remove the moisture. The test results listed in Table 2 show that the conversion rate of oven-dried wooden cubes is significantly higher than those used in test number 3. The time required for the gasifier to attain the steady condition was ∼ 50 min in the test conducted with oven-dried wooden cubes, which is shorter than those of the other test runs, and a flame spreading speed of 0.18 mm s-1 was achieved. In test number 9 where the fuel fed had a moisture content of 33%, the time required for the gasifier to reach steady condition was more than 3 h and resulted in a significantly lower flame spreading speed of 0.018 mm s-1. The moisture contained in the wood would also adversely affect the heating value of the producer gas because a large amount of energy would be consumed in moisture evaporation. An increase of the fuel moisture content could increase the energy needed for vaporization and decrease the temperatures in solid and gas phases, as a result, the rate of heat transferred to the surface of the solid is suppressed. Di Blasi et al.30 performed pyrolytic experiments of thick wooden cylinders and showed that the yields of CO and CH4 increase with the radiant heat flux. Table 5 shows that test number 8 fueled with ovendried wooden cubes produced a syngas containing more H2 (17.47%) and CO (23.68%) than that of the producer gas generated in test number 9 (14.59% of H2 and 18.39% of CO). The fuel conversion rate and producer gas heating value of test number 9 were particularly low, as shown in Table 2, mainly because of the relatively high fuel moisture content (∼33%). 5. Conclusions An experimental investigation of a small-scale fixed-bed, stratified downdraft biomass gasifier was performed. The single most significant factor that influences the gasifier performance is the regulation of air flow. The conversion rate of wooden cubes and the specific gasification rate would increase with increasing air flow rates. The producer gas heating value is (30) Di Blasi, C.; Branca, C.; Santoro, A.; Hernandez, E. G. Combust. Flame 2001, 124, 165–177.

Air-Blown Fixed-Bed Downdraft Biomass Gasifier

strongly dependent upon the air flow rate as well as the fuel moisture content. The producer gas heating value was observed to increase with increasing air flow rates up to 15 Nm3 h-1, and a further increase in air feeding results in a decrease in the gas heating value, because of the increase of conversion of CO and H2 with oxygen as the air flow rate was increased. The optimum mean HHV of the dry producer gas obtained in this study by feeding the air unheated and wooden cubes containing 18% moisture is 5.68 MJ Nm-3, which is comparable to the results of Walawender et al.4 and Barrio et al.5 The estimated cold gas efficiency varies between 70 and 89%, and the hot gas efficiency ranges between 74 and 96%. The H/C ratios of combustible gas produced in the downdraft biomass gasifier are in the range between 0.113 and 0.223. Although the specific gasification rate and the specific gas production rate would increase positively with the air flow rates, the cold gas efficiency, hot gas efficiency, and H/C ratio would achieve a peak value and beyond that a further increase in the air supply would result in a decline in the cold gas efficiency, hot gas efficiency, and H/C ratio. Consequently, there exists an optimum air flow rate for high system conversion efficiency and high-quality producer gas. The fuels having lower moisture content are more desirable, because it could shorten the time for the fuel bed to reach the steady-state condition, produce a producer gas with a high heating value, and achieve a higher rate of fuel conversion and a lower air/fuel ratio. Preheating the air streams can also shorten the time required to achieve the steady-state condition but appears to have marginal effects on the producer gas heating value. A one-dimensional, steady-state model based on the conservations of mass and energy has been developed for simulation of the axial profiles of the temperature and gas composition of downdraft biomass gasifiers. Generally, this model performs well in predicting the axial temperatures of solid and gaseous phases. The discrepancy between the measured and predicted gas composition profiles is mainly due to the non-equilibrium thermal decomposition reactions undergoing inside the wooden particles when it is heated unsteadily, which is not considered in the model. Experimental results obtained in this work indicate that, when operating a downdraft stratified fixed-bed biomass gasifier, regulating the air flow rate is effective to control the gasification rate and allows for adjustment of combustible gas composition and the system efficiency. The optimum operating air flow would be to maintain a low air/fuel ratio, which is sufficient to sustain the top-stabilized mode to obtain the desired producer gas H2 and CO composition, high heating value, and conversion efficiency. Acknowledgment. This study was supported by the National Science Council of the Republic of China (NSC 94-2218-E-002070 and 95-2218-E-002-035) and Dai-East Incinerators, Inc., Taiwan.

Nomenclature a, b, c, and f ) atomic number representations of carbon, hydrogen, and oxygen, respectively, and that of the moisture in the chemical formula of the fuel

Energy & Fuels, Vol. 22, No. 6, 2008 4205 AG ) area of the grate, m2 Am ) pre-exponential factor in eq 34, h-1 ap ) fuel bed solids surface per unit volume, m-1 ar ) ratio of change in gaseous moles to change in oxygen moles B ) systematic uncertainty BG ) specific gasification rate, kg m-2 h-1 C ) molar concentration, kmol m-3 Cp ) specific heat, kJ kg-1 K-1 D ) internal diameter of reactor, m Dg ) gas diffusion coefficient, m2 s-1 Em ) activation energy, kJ kmol-1 F˙a ) air flow rate, Nm3 h-1 F˙g,dry ) dry gas production rate, Nm3 h-1 h ) specific enthalpy, kJ kg-1 hws and hwg ) heat-transfer coefficients of the wall-solid and wall-gas, respectively, W m-2 K-1 hD ) mass-transfer coefficient, m s-1 ∆h ) heat of reaction, kJ kg-1 jD ) Colburn-Chilton mass-transfer factor k ) thermal conductivity, W m-1 K-1; reaction rate constant, m s-1 j ) average reaction rate in eq 10 K L ) fuel bed height, m m ˙ fuel ) fuel conversion rate, kg h-1 ˆ ) molecular weight M n ) order of reaction in eq 10 P ) random uncertainty Rd ) coefficient of the reaction rate Re ) Reynolds number Sc ) Schmidt number r˙ ) reaction rate per unit volume, kg m-3 s-1 rCO ) fraction of carbon converted to carbon monoxide rˆof ) stoichiometric oxygen/fuel mole ratio T ) temperature, K u ) velocity, m s-1 XO2 ) mole fraction of oxygen y ) axial space, m; mole fraction of gases F ) density, kg m-3 ε ) fuel bed porosity ηcold ) cold gas efficiency, % ηhot ) hot gas efficiency, % σ ) uncertainty Subscripts AR ) air flow rate a ) air AFR ) air/fuel ratio C ) char FCR ) fuel conversion rate FP ) flaming pyrolysis GC ) gas composition GHV ) gas heating value GPR ) gas production rate g ) gas G, CO2 ) char-CO2 gasification reaction G, H2O ) char-H2O gasification reaction G, H2 ) char-H2 gasification reaction m ) moisture s ) solid T ) temperature w ) reactor wall WGS ) water-gas shift reaction EF800026X