Energy Fuels 2010, 24, 1316–1322 Published on Web 12/23/2009
: DOI:10.1021/ef9010077
Modeling the Woody Biomass Combustion in a Vortexing Fluidized-Bed Combustor Kuo-Chao Lo,† Keng-Tung Wu,‡ Chien-Song Chyang,*,† and Kuan-Chang Su† †
Department of Chemical Engineering, Chung Yuan Christian University, Chung Li, Taiwan 32023, Republic of China and ‡ Department of Forestry, National Chung Hsing University, Taichung, Taiwan 40227, Republic of China Received September 10, 2009. Revised Manuscript Received November 28, 2009
This study develops an isothermal one-dimensional model to easily predict the performance of biomass combustion in a vortexing fluidized-bed combustor (VFBC). This model comprises a series of three reactor zones: a fluidized-bed combustion chamber, a lower freeboard, and an upper freeboard. This model can calculate the composition of the flue gas exhausted from a combustor and the concentration profiles of individual gas species at various combustor heights. This study verifies the proposed model by comparing its predictions with actual experimental data obtained from a series of systematic combustion experiments burning pinewood sawdust in a pilot-scale VFBC. This comparison revealed an air-leakage problem in the pilot-scale VFBC. The simulated results of CO2 and O2 concentrations are most sensitive to EA but almost non-sensitive to λ. As for the operating temperature, the simulation results for CO, CO2, and O2 are only sensitive at temperatures lower than 750 °C. The CO concentration in the exhaust is more affected than CO2 and O2 by the secondary air added. The fresh O2 added by secondary air can enhance CO consumption, but too much secondary air shortens the residence time of CO in the upper freeboard. Although there are obvious gaps between experimental and simulation results, the predicted change tendencies of gas concentrations correspond to the actual data. This model makes it possible to obtain more information in the preliminary steps of designing a larger scale VFBC for burning highly volatile fuel, e.g., biomass.
Researchers typically employed three gas-solid models (i.e., plug flow and two- and three-phase models) to describe hydrodynamic behaviors in fluidized-bed reactors. The early studies on gas-solid reactors assumed the plug flow of gas through a bubbling fluidized bed (BFB). However, most of the gas upward bypasses in type of bubbles in BFB. The plugflow model neglects the bubble effects, and therefore, its predictions are usually overestimated.2,3 In 1952, Toomey and Johnstone4 introduced the simple two-phase concept, which divides a fluidized bed into two phases: emulsion and bubble phases. On the basis of this concept, numerous varieties of two-phase models were built to describe the bubbling fluidized bed. The two-phase theory provides an ideal and simplified description of the hydrodynamic phenomenon in BFBs. On the basis of the studies regarding the fluid mechanics in a fluidized bed by Davidson and Harrison5 and Rowe and Partridge,6 a well-known “Davidson bubble” was modeled. Kunii and Levenspiel7,8 developed a BFB model based on the Davidson bubble assumptions, the K-L model, that unifies the concepts of the BFB flow pattern in a useful form. This simple three-phase model of the gas flow through fluidized
1. Introduction Because of its dual contributions to sustainability and environmental protection, biomass is perhaps the most feasible alternative energy source for a primary energy supply. Since people first started to use fire, biomass has played an important role as energy sources throughout human history. For thousands of years of human culture, biomass combustion dominated the energy supply until the mass exploitation of fossil fuels. Direct combustion is the simplest, most common, and the most successful thermo-chemical process for converting biomass into energy. Combustion technologies, including fluidized-bed combustion (FBC), were invented or matured during a period in which fossil fuel was the predominant energy resource. Consequently, many current biomass combustion technologies are based on the concept of fossil fuel combustion. Among the various combustion technologies, FBC technology has attracted significant attention for extracting energy from biomass because of its fuel flexibility. Over the past 50 years, researchers have tried to simulate the combustion behaviors of bubbling fluidized-bed combustors (BFBCs) using mathematical models. The purpose of modeling BFBCs is to correctly and easily predict the fuel combustion process in a fluidized-bed combustor. In the general model of a BFBC with solid fuel combustion, the processes in a fluidizedbed combustor should include bed hydrodynamics, solid fuel combustion and reactions, and phenomena in the freeboard.1
(2) Levenspiel, O. Powder Technol. 2002, 122, 1–9. (3) Mostoufi, N.; Cui, H.; Chaouki, J. Ind. Eng. Chem. Res. 2001, 40, 5526–5532. (4) Toomy, R. D.; Johnstone, H. F. Chem. Eng. Prog. 1952, 48, 220– 226. (5) Davidson, J. F.; Harrison, D. Fluidized Particles; Cambridge University Press: New York, 1963. (6) Rowe, P. N.; Partridge, B. A. Trans. Inst. Chem. Eng. 1965, 43, 157–175. (7) Kunii, D.; Levenspiel, O. Ind. Eng. Chem. Fundam. 1968, 7, 446– 452. (8) Kunii, D.; Levenspiel, O. Ind. Eng. Chem. Process Des. Dev. 1968, 7, 481–492.
*To whom correspondence should be addressed: Department of Chemical Engineering, Chung Yuan Christian University, Chung Li, Taiwan 32023, Republic of China. Telephone: þ886-3-256-4119. Fax: þ886-3-463-6242. E-mail:
[email protected]. (1) Oka, S. N. Fluidized Bed Combustion, 1st ed.; Marcel Dekker, Inc.: New York, 2004. r 2009 American Chemical Society
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: DOI:10.1021/ef9010077
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Figure 2. Scheme of the hydrodynamic assumption of the VFBC model.
emissions.13-16 However, few studies have been published on modeling biomass combustion in a VFBC, and a mathematical VFBC model is needed to advance VFBC technology. This paper develops a one-dimensional isothermal VFBC model to address this need and compares the results of pinewood sawdust combustion in a pilot-scale VFBC with the simulation results to determine the suitability of this VFBC model. The experiments in this study use typical operating parameters [temperature, excess air, and air partition ratio (λ)]. The objective of this research is to develop a useful model that can easily predict the combustor performances in the preliminary steps of designing a large-scale VFBC.
Figure 1. Sketch of the oblique wall design of the overbed feeding point.
beds assumes that bubbles are uniformly sized, surrounded by clouds, and followed by wakes, rising through an emulsion of the downward moving solid.8 Kinetic studies of biomass combustion have been developing for a long time, but the large demand for biomass energy has led to new findings in recent years. Biomass is constructed of simpler and more original biological structures. The combustion process of solid fuels can be roughly described as a combination of events: drying, pyrolysis (or devolatilization), homogeneous oxidation of volatile matters, heterogeneous oxidation of solid char, and finally, the burning out of char. These events can happen sequentially or simultaneously depending upon factors such as the fuel size, temperature, and types of combustor. Because of the small fuel size and the rapid heating combustor, e.g., BFBC, the biomass fuel is reasonable to assume the presence of instantaneous devolatilization when biomass is fed into a combustor in this study.9-12 Volatiles and char particles burn with the upward gas flow which flows through the bed and freeboard successively. The high volatile content of biomass fuel emphasizes the need to analyze the reactions and phenomena in the freeboard. It is common practice to add additional air in the freeboard to aid volatile combustion. An advanced BFBC design, called vortexing fluidized-bed combustor (VFBC), was developed to enhance the combustion performance in the freeboard of a conventional BFBC. The distinguishing characteristic of a VFBC is its tangential secondary air injection in the freeboard. The vortex flow acts as an internal cyclone in the freeboard, trapping more unburned char particles and providing sufficient air to improve combustion. Some VFBC experiments demonstrate high combustion efficiency and low pollutant
2. Model Description This study proposes an isothermal one-dimensional model to predict the performance of biomass combustion in a VFBC. The proposed model imitates a pilot-plant-scale VFBC, which is designed and built by the Department of Chemical Engineering of Chung Yuan Christian University. The VFBC features a rectangular fluidized-bed combustion chamber with a special design of the feeding point. An oblique-wall design makes solid fuel particles fall into the fluidized bed entirely (see Figure 1). The fuel particles mix and devolatilize in the shallow and vigorous fluidized bed. This model can calculate the composition of the flue gas exhausted from a combustor and the concentration profiles of individual gas species at various combustor heights. The VFBC in this study is modeled as a series of three reactor zones: a fluidized-bed combustion chamber, a lower freeboard, and an upper freeboard (Figure 2). The fluidized bed extends from the air distributor to the bed surface. The lower freeboard (splash zone included) extends from the bed surface to the secondary air injection point. The remainder of the combustor forms the upper freeboard, which extends from the secondary air injection to the combustor exit. The main assumptions and characteristics of the VFBC model are as follows: (i) The model is isothermal, one-dimensional, and steady-state. (ii) The fluidized bed consists of two phases, emulsion and bubble phases. (iii) Gas streams in the emulsion phase, bubble phases, and freeboard are assumed as plug flow. (iv) Inert bed material and fixed carbon particles are wellmixed in the emulsion phase. (v) The bubble diameter is constant as the average bubble size in the fluidized bed. (vi) The bubbles are free of solids. (vii) Biomass fuel is mixed well and immediately devolatilized in the emulsion phase when fed
(9) Jiang, H.; Morey, R. V. Biomass Bioenergy 1992, 3, 431–447. (10) Bryden, K. M.; Ragland, K. W. Energy Fuels 1996, 10, 269–275. (11) Zhou, H.; Jensen, A. D.; Glarborg, P.; Jensen, P. A.; Kavaliauskas, A. Fuel 2005, 84, 389–403. (12) Petersen, I.; Werther, J. Chem. Eng. Proc. 2005, 44, 717–736. (13) Lin, C. H.; Teng, J. T.; Chyang, C. S. Combust. Flame 1997, 110, 163–172. (14) Teng, H.; Chyang, C. S.; Shang, S. H.; Ho, J. A. J. Air Waste Mange. Assoc. 1997, 47, 49–57. (15) Chyang, C. S.; Liu, C. Y.; Chang, Y. D. J. Air Waste Mange. Assoc. 2001, 51, 542–551. (16) Chyang, C. S.; Lo, K. C.; Wang, K. L. J. Chem. Eng. Korea 2005, 22, 774–782.
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into the fluidized-bed combustion chamber. (viii) The volatile gas composition is estimated by the pyrolysis model by Jiang and Morey.9 (ix) Unburned char particles are all trapped by vortexing flow in the freeboard. (x) Unburned char particles are uniformly distributed in the lower freeboard. (xi) Simple kinetic models are used for homo- and heterogeneous reactions. 2.1. VFBC Hydrodynamics. A simple two-phase model describes the hydrodynamic behavior in the fluidized-bed combustion chamber. Homogeneous gas reactions occur in both the emulsion and bubble phases. Meanwhile, concentration differences drive the gaseous mass transfer between phases. Char particles are burned in the fluidized bed and lower freeboard. To simplify the burning out process of char particles, this study defines a factor β as char conversion in the bed. The value of β was obtained by the experimental data. On the basis of the experimental data, the oxygen consumption and the heat release through fuel combustion can be formulated by balance calculations. According to these two balance equations, we can estimate the conversions of char and volatile matter combustion in the fluidized-bed region (see the Supporting Information). The value of β is constant at 0.64 when a VFBC operates under the typical conditions of bubbling fluidization. Approximately, 64% of the char particles are burned in the bed, while the rest burns out in the lower freeboard. Because tangential secondary air traps all of the char particles, only gas combustion reactions occur in the upper freeboard. The following paragraphs list the correlations of hydrodynamics and mass transfer in this model. Minimum fluidization velocity:17 ! pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μg ð 33:72 þ 0:0408Ar -33:7Þ Umf ¼ ð1Þ Fg dp Superficial velocity in the emulsion phase:18 1 Ue ¼ Umf þ ðUsv - Umf Þ 4 Voidage of the emulsion phase:19 ðUsv -Umf Þ εe ¼ ðεmf þ 0:2Þ -0:059 exp 0:429 Volume fraction of the bubble phase:19 -ðUsv -Umf Þ δ ¼ 1 -0:166 - 0:534 exp 0:413
Mass-transfer coefficient between the bubble and emulsion phases:18 2:7Ue ð8Þ K be ¼ 4 a ¼
pyrolysis
biomass sf volatiles ðH2 , CO, CH4 , CO2 , C2 H4 , N2 Þ þ tar ðCHx Oy Nz Þ þ char ðCÞ
ð4Þ
Average bubble diameter:20
Afb ðUsv - Umf Þ 0:4 db0 ¼ 0:8713 ND
ð6Þ
dbm ¼ 1:6377ðAfb ðUsv - Umf ÞÞ0:4
ð7Þ
ðR1Þ
This study also uses temperature-dependent correlations given by Jiang and Morey to attain volatiles, tar, and volatile composition yields.9 Char consists of fixed carbon, and the amount of char yield equals the fixed carbon composition of the biomass. The parameters x, y, and z in the chemical formula for tar are computed by atomic conservation. The proposed model also accounts for the homogeneous combustion of volatiles and tar and heterogeneous char particle combustion. The homogeneous reactions dominate the biomass combustion process. H2, CH4, and C2H4 are oxidized to CO and H2O, and char is oxidized to CO and CO2. The primary CO converts to CO2 by further oxidation. Table 1 lists the reactions mentioned above. 2.3. Mass Balance. The proposed model differentially divides a VFBC into several control volumes along the vertical axis. Each control volume is dz in height. The temperature in one control volume is equivalent in both phases, and there is no concentration gradient in each phase. The proposed model is therefore based on the mass balance of each control volume. The general forms of steady-state mass balances for one control volume and phase in the fluidized bed are formulated as ðrate of mass flow inputÞ - ðrate of mass flow outputÞ
ð3Þ
ð5Þ
ð9Þ
2.2. Chemical Reactions. Predicting the volatile composition of primary pyrolysis is extremely difficult. The factors related to volatile composition include the heat rate, pressure, particle size, and temperature. This study simplifies the sawdust combustion model by assuming that volatile matters release instantaneously when the biomass fuel is introduced into the hot fluidized bed. This study adopts a simple stoichiometric model of biomass pyrolysis to figure out the volatile products (R1).
ð2Þ
-0:15hmf db ¼ dbm - ðdbm - db0 Þexp dfb
6δ db
þ ðmass change by chemical reactionÞ þ ðmass change by interphase diffusionÞ ¼ 0
ð10Þ
The differential form of mass balance for species j in the ith control volume of the emulsion phase in the fluidized-bed combustion chamber is expressed as 1 dnj , i, e 1 ¼ -Kbe aðCj , i, e - Cj , i, b Þ þ nj, py - ð1 - δÞεe Afb dz hfb X β νj rj þ nj, i, fc ð11Þ hfb
(17) Wen, C. Y.; Yu, Y. H. Chem. Eng. Prog. Symp. Ser. 1966, 62, 100–111. (18) Tepper, H. Zur vergasung von rest- und abfallholz in wirbelschichtreaktoren f€ ur dezentrale energieversorgungsanlagen. Ph.D. Dissertation, Otto-von-Guericke University, Magdeburg, Germany, 2005 (in German). (19) Cui, H.; Mostoufi, N.; Chaouki, J. Chem. Eng. J. 2000, 79, 133– 143. (20) Jafari, R.; Sotudeh-Gharebagh, R.; Mostoufi, N. Chem. Eng. Technol. 2004, 27, 123–129.
The first term on the right-hand side is the mass transfer between emulsion and bubble phases. The second term describes the instantaneously homogeneous devolatilization P of biomass fuel in the emulsion phase. The term νj rj summarizes all of the j-related reaction rates, while νj is the stoichiometric coefficient of species j. Because the products of char combustion are CO and CO2, the last term of eq 11 1318
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Table 1. Chemical Reactions for Biomass Combustion Simulation reaction
rate expression (mol m-3 s-1)
rate constant (s-1)
2H2 þ O2 f 2H2O CO þ 1/2O2 f CO2 CH4 þ 3/2O2 f CO þ 2H2O C2H2 þ O2 f 2CO þ 2H2 CHxOyNz þ (1 þ x/4 - y/2)O2 f CO2 þ x/2H2O þ N2
rH2 = kH2CH21.5CO2 rCO = kCOCCOCO20.5CH2O0.5 rCH4 = kCH4CCH40.7CO20.8 rC2H4 = kC2H4CC2H40.9CO21.18 rtar = ktarCtar0.5CO2
kH2 = 51.8 T1.5 exp(-3420/T) (R2)11 kCO = 3.25 107 exp(-15098/T) (R3)11 kCH4 = 1.6 1010 exp(-24157/T) (R4)11 kC2H4 = 3.71 1012 exp(-25163/T) (R5)12 ktar = 2.9 105T exp(-9650/T) (R6)11
Table 2. Properties of Pine Sawdust proximate analysis
wt %
moisture volatile matters fixed carbon ash
13.6 72.18 13.82 0.4
ultimate analysis
The bed temperature was controlled by a tubular heat exchanger. During operation, flue gas from the sawdust combustion entered a baghouse filter after cooling. The baghouse then separated dust materials from the flue gas, which was induced into a wet scrubber to remove gaseous pollutants. The clean flue gas was then exhausted to the air. Five gas sampling probes detected CO/CO2/O2 concentration profiles along the height of the VFBC combustor. The CO/CO2/ O2 profiles and concentrations at the outlet of VFBC were collected to compare actual results with predicted results. The O2 (accuracy = 0.3%) and CO (accuracy = 1 ppm) concentrations were detected by a TSI-CA 6230 flue gas analyzer (electrochemical gas sensor). The CO2 concentration was detected by a MTI-M200 GC. A one-factor-at-a-time (OFAT) experimental design was conducted to obtain the results of various operating parameters. Three typical operating parameters [bed temperature, excess air, and air partition ratio (λ)] were investigated in the experimental tests. The bed temperatures were controlled by changing the heatexchanger immerged area in the bed. The tests of the EA effect were carried out with a given primary air (fluidizing air), the stoichiometric O2 requirement, and various secondary air. The various excess air was supplied by secondary air. The air partition ratio is defined as a ratio of primary air to the total air, with a fixed excess air (EA = 60%). The total air was kept the same in the tests of the λ effect. A middle level operating condition was T = 700 °C, EA = 60%, and λ = Q1/(Q1 þ Q2) = 0.625. The tested factors increased and decreased on the basis of the middle level condition. Table 3 lists the experimental array.
daf wt %
C H O N HHV
48.39 5.80 44.21 0.21 18.2 MJ/kg
equals 0 if j 6¼ CO or CO2. In the bubble phase, mass balance is expressed as X 1 dnj , i, b νj rj ð12Þ ¼ Kbe aðCj , i, e - Cj, i, b Þ - δ Afb dz In the lower freeboard, a gas-combustible species burns continuously in the upward stream. Because the particles are entrained with the upward gas flow and the downward particles, trapped by secondary air, they exist simultaneously. For simplification, this model assumes that a wellmixed particle flow is applied in this region. The last 36% of char particles burn out in this region. The differential form of species j in the ith control volume in the lower freeboard is formulated as X 1 dnj, i ð1 - βÞ nj, i, fc ð13Þ ¼νj rj þ Alf dz hlf
4. Results and Discussion The purpose of this study is to develop a useful mathematical model to easily predict the performance of biomass combustion in a VFBC. This study measures VFBC performance in terms of flue gas composition. The proposed model simulates gas concentration changes at various heights above the air distributor. The concentration of CO is an essential part of determining combustion efficiency. The generally accepted theories for the mechanisms of hydrocarbon combustion state that the formation of CO is the earlier step of carbon oxidizing into CO2. The CO concentration in exhausting flue gas levels below 2000 ppm at 6% O2 (air pollutant control regulations of Taiwan) implies that all of the hydrocarbons were burned out. Therefore, we can determine the optimum performance of the combustor by comparing the CO concentrations of different simulation cases. Figures 3 and 4 depict the O2 and CO concentration profiles versus VFBC height under the middle level operating conditions. The CO and O2 profiles of the emulsion and bubble phases are plotted separately in the fluidized-bed combustion chamber. The two gas streams merge into one above the bed surface. In Figures 3 and 4, in the fluidized-bed region, the emulsion phase has a lower O2 concentration than the bubble phase and a contrary behavior of CO. The proposed model assumes that the sawdust decomposes in the emulsion phase. Then, the volatile combustibles simultaneously burn in the emulsion phase and transport into the bubble phase.
As in eq 11, the last term of eq 13 is considered only in the cases of j = CO or CO2. The mass balance in the upper freeboard is expressed as X 1 dnj , i ¼νj rj ð14Þ Auf dz Integrating these differential equations produces the concentration profiles of the considered gas species along the height of a VFBC. This study uses the semi-implicit third Runge-Kutta method to conduct numerical integration. 3. Experimental Section Biomass combustion experiments were conducted in a pilotscale VFBC consisting of a 0.4 0.8 m rectangular fluidized-bed combustion chamber and a 0.75 m diameter cylindrical freeboard. The combustor is 4.5 m in height. Secondary air is tangentially injected into the freeboard by four nozzles positioned 1.75 m above the air distributor. Pinewood sawdust served as the biomass fuel. The bed material is 160 kg of silica sand with a diameter of about 500 μm. The Umf of the bed material is about 0.08 m/s at 800 °C. The height of the expansion bed is about 0.7 m. Table 2 shows the proximate and ultimate analyses of the sawdust. Sawdust was fed into the combustion chamber by a chute at a position 0.85 m above the air distributor, and the feed rate was controlled by varying the revolutions per minute (rpm) of the screw feeder. The feed rate was fixed at 37 kg/h ca. 150 kWth. 1319
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Table 3. OFAT Experimental Array T
EA
λ = Q1/(Q1 þ Q2)
650 700 750 800
60% 60% 60% 60%
0.625 0.625 0.625 0.625
700 700 700 700 700
10% 20% 40% 60% 80%
0.909 0.833 0.714 0.625 0.556
700 700 700 700 700
60% 60% 60% 60% 60%
0.5 0.625 0.75 0.875 1
Figure 4. CO profiles in the VFBC (T = 700 °C, EA = 60%, and λ = 0.625).
due to an O2-dificient environment. In the upper freeboard, the CO concentration drops at the secondary air injection, because fresh air enhances the CO combustion and dilution effect. However, the great differences between the detected data and model-predicted results in Figures 3 and 4 are somewhat confusing. The higher detected O2 concentration implies that the combustion is not as good as expected. However, the CO data do not support this opinion, because poor combustion conditions should produce a high CO concentration. To explain this phenomenon, this study examines not only model construction factors but also practical experimental conditions. The air-leakage phenomenon has been demonstrated by cold experiments. We measured the gas flow rate at the exit pipe of the VFBC in the condition of a given air supply and the regular negative operating pressure. The results indicated that there is additional air leaking into the VFBC. To ensure the safety of the combustor operator, the operating pressure in the combustor was kept lower than the atmospheric pressure surrounding the combustor. This may have given ambient air the opportunity to leak into the combustor if it was not airtight. In this case, a hand hole, located in the fluidized-bed region, and an air-lock valve on the feeding chute provided paths for air leakage. These air-leakage paths are near the fluidized-bed region and the lower freeboard. Resulting in oversupply of air enhanced CO combustion and increased the O2 concentration in these two regions. The great differences between experimental data and simulation results in Figure 4 are not only attributed to the air leakage. The temperature profile in the combustor is another possible reason. The isothermal assumption controls the temperature in the lower freeboard in simulations. However, in practice, a higher temperature took place in the lower freeboard. Both the faster combustion rate caused by the higher temperature and the dilution effect, which is caused by the air leakage, make a big difference between experimental data and simulation results. To eliminate the effect of air leakage and further modeling studies with non-isothermal should improve agreements between the experimental data and simulation results. Figures 5-7 compare the experimental data and modelpredicted results for the CO, CO2, and O2 concentrations of flue gas exhausted from the combustor. The dilution effect behaves through the obvious divergences between the simulation and
Figure 3. O2 profiles in the VFBC (T = 700 °C, EA = 60%, and λ = 0.625).
The higher volatile matters consume more O2 and produce more CO in the emulsion phase. Contrary to the bubble phase, less volatile matter makes higher O2 and lower CO concentrations than the emulsion phase. No fresh air was added in lower freeboard region until the secondary air injection at the bottom of the upper freeboard. Secondary air provides enough fresh O2 to enhance combustion. The remaining combustible gases burn faster and more completely after the secondary air injection. The O2 concentration changes little in this region, implying that a significant part of the combustion process is completed in the fluidizedbed combustion chamber and lower freeboard. Experimental data detected in the upper freeboard seem to support the predicted results (Figure 3). Figure 4 shows that the maximum CO concentration occurs at the interface between the fluidized bed and the lower freeboard. This position is defined as a splash zone in practical fluidization operation. The model in this study assumes that the lower freeboard region includes the splash zone. The increasing concentration of CO in the fluidized bed is caused by the well-mixed solid fuel in the emulsion phase. Fresh volatiles enter the emulsion phase of every control volume in the fluidized-bed region. The CO consumption rate is faster than its production rate after the gas stream leaves the fluidized-bed region. Therefore, the CO profile presents a decreasing tendency above the bed surface. A retarding decreasing tendency near the top of the lower freeboard is 1320
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: DOI:10.1021/ef9010077
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Figure 5. Comparisons of predicted results with experimental data of CO, CO2, and O2 concentrations in the flue gas at various temperatures (EA = 60% and λ = 0.625).
Figure 7. Comparisons of predicted results with experimental data of CO, CO2, and O2 concentrations in the flue gas at various air partition ratios (T = 700 °C and EA = 60%).
Figure 5, which shows the effect of the operating temperature, indicates that the predicted results are sensitive at lower temperatures, especially for CO. At temperatures over 750 °C, the CO concentration is zero and O2 and CO2 change little. The reaction rate of CO converting into CO2 is very sensitive to temperature. Lower temperatures tend to produce more CO and lead to higher CO emission. In the low-temperature combustion conditions, lower than 750 °C, the CO decrement caused by temperature increasing is obvious. Contrary, the CO emissions of high-temperature combustion are too low to reflect the effect of the temperature. However, the experimental data in this figure do not seem to fit the predicted results. This is because of the air-leakage effect. Lower CO2 concentrations were caused by the dilution effect, whereas higher O2 concentrations were caused by superfluous air supply. In experimental tests with various excess air levels, the superficial velocities in the bed were kept at a constant level to meet 100% of the stoichiometric O2 demand. The amount of excess air was provided only by secondary air. Figure 6 shows the effect of EA and comparisons between experimental and model-predicted results. As mentioned above, the difference in magnitude between the experimental and simulation results is caused by air leakage. The decreasing tendency of CO2 and increasing tendency of O2 correspond to the simulation results and experimental data, which demonstrate that the proposed model is not far from reality. The simulated CO curve first decreases and then increases when EA > 40%. This is because more EA implies a higher superficial velocity in the upper freeboard. Although this region provides more O2, there is not enough time to consume the CO. Figure 7 shows the experimental and model-predicted results at various air partition ratios (λ). In the tests of various λ, the simulation results of O2 and CO2 are not sensitive. According to the definition of λ, similar results at various λ are achieved when the simulations are carried out at the same operating temperature and EA. The O2 and CO2 concentrations are mostly dictated by the overall mass balance in the tests of the λ effect. Therefore, the O2 and CO2 concentrations in flue gas almost do not change with the different λ. The CO curve, first decreasing and then increasing when λ > 0.6, in Figure 7 is attributed to the various residence times in the upper freeboard and the different reaction environments caused by secondary air injection. A smaller λ means less
Figure 6. Comparisons of predicted results with experimental data of CO, CO2, and O2 concentrations in the flue gas at various excess air (T = 700 °C and λ = 0.625).
experimental data of O2 and CO2 in these figures. However, the experimental data of CO seem to fit the simulation results well. Actually, the dilution effect also took place on the present experimental results. The CO concentrations without dilution must be higher than the present data. The simulation results are based on some assumptions of ideal combustion conditions, i.e., isothermal and well mixing. However, the real combustion conditions are not exactly the same as the simulations. The combustion intensity is lowering and losing heat in the region of the end of the VFBC. The temperature is usually lower here than the other regions. Therefore, the real CO concentrations should be higher than the simulated results. Kouprianov and Permchart reported some CO concentration data of sawdust combustion in a conical FBC.21,22 The results of their experiments indicate that the exhausting CO concentration ranges from 300 to 1000 ppm. To compare the results of the present study with the literature, the experimental CO concentrations of this study are lower than the data in the literature. Therefore, the dilution effect causes the experimental results to fit the simulated results well. (21) Kouprianov, V. I.; Permchart, V. Appl. Energy 2003, 74, 383– 392. (22) Permchart, V.; Kouprianov, V. I. Bioresour. Technol. 2004, 92, 83–91.
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primary air and more secondary air. The CO concentration increases as the residence time decreases (λ < 0.6). When λ > 0.6, a higher λ means less O2 enters the upper freeboard as additional fresh air. The different combustion atmospheres in the upper freeboard cause the CO concentration to increase slightly with λ.
Ar = Archimedes number C = concentration (mol/m3) d = diameter (m) EA = excess air ratio h = height (m) k = reaction rate constant (s-1) Kbe = mass-transfer coefficient (m/s) n = mole (mol) ND = number of holes of the air distributor Q1 = flow rate of primary air (N m3/min) Q2 = flow rate of secondary air (N m3/min) r = chemical reaction rate (mol/s) T = temperature (°C) U = velocity (m/s) z = vertical axial distance (m)
5. Conclusions This study proposes an isothermal one-dimensional model to understand biomass combustion in a VFBC. The hydrodynamics in the fluidized-bed combustion chamber are assumed on the basis of the two-phase theory. This model divides a VFBC into several control volumes along its vertical axis. The proposed model integrates the differential mass balance of each control volume using the semi-implicit third Runge-Kutta method. In comparison to the experimental results from a pilot-scale VFBC, it can be concluded that the model-predicted values follow the same tendency. The air leakage caused the obvious divergence between experimental and simulation results. Despite the isothermal assumption not being exactly the same as real combustion conditions, the simulation results can reasonably describe the real combustor behaviors. This model provides some preliminary knowledge to design a commercial VFBC as well as ideas to modify existing BFBCs. The proposed model is a simplification of reality intended to increase our understanding of biomass combustion in a VFBC. However, future studies on the essential function of vortexing flow are needed to further improve this VFBC model.
Greek Letters β = char conversion in the bed δ = volume fraction of the bubble phase ε = voidage λ = air partition ratio ν = stoichiometric coefficient F = density (kg/m3) Subscripts b = bubble phase b0 = initial bubble bm = maximum bubble e = emulsion phase fb = fluidized-bed region fc = fixed carbon g = gas i = ith control volume j = jth species lf = lower freeboard mf = minimum fluidization p = particle py = pyrolysis sv = superficial velocity uf = upper freeboard
Acknowledgment. This research was supported by the Chung Yuan Christian University, Taiwan, under Grant CYCU-98-CR-CE. Supporting Information Available: How to estimate the β (char conversion in a fluidized bed) value. This material is available free of charge via the Internet at http://pubs.acs.org.
Nomenclature 2
A = area (m ) a = specific area (m2/m3)
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