Gasification of Coffee Grounds in Dual Fluidized Bed - American

Energy & Fuels 2006, 20, 2695-2704

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Gasification of Coffee Grounds in Dual Fluidized Bed: Performance Evaluation and Parametric Investigation Guangwen Xu,* Takahiro Murakami, Toshiyuki Suda, Yoshiaki Matsuzawa, and Hidehisa Tani Ishikawajima-Harima HeaVy Industries Co., Ltd. (IHI), Shin-Nakahara-Cho 1, Isogo-ku, Yokohama 235-8501, Japan ReceiVed March 23, 2006. ReVised Manuscript ReceiVed August 28, 2006

Coffee grounds refer to a kind of high-moisture biomass refuses generated in drink works. With a national technical program, we recently worked on converting this biomass waste into middle-caloric product gas. The conversion proceeded in two consecutive steps, via first fuel drying/upgrading and in turn pyrolytic gasification of the dried fuel with the so-called dual fluidized bed gasification (DFBG) technology. The present paper investigated the gasification of dried coffee grounds in a 5.0 kg/h pilot DFBG facility, with the aim of demonstrating the adaptability of the technique to the tested fuel and clarifying its chemically possible efficiency. It was shown that gasifying dried coffee grounds with moisture of about 10 wt % through DFBG at about 1073 K is easy to convert more than 70% of fuel’s C into product gas, and the gas can have a higher heating value (HHV) over 3500 kcal/m3n. Nonetheless, the tar load in the product gas was sometimes up to 50 g/m3n. Increasing the steam/fuel mass ratio and decreasing the fuel particle size reduced the tar yield, but the available reduction degree was limited. Inclusion of a small amount of air into steam (gasification reagent) appeared efficient to lower the tar content of the product gas, but only the O2/C molar ratios below 0.1 are applicable in the view of preventing the HHV of the product gas from becoming lower than 3000 kcal/m3n. As a consequence, other tar elimination techniques were suggested to be necessary for the gasification of coffee grounds via DFBG. Furthermore, the paper demonstrated that for DFBG using steam as the gasification reagent its attainable C conversion likely determines the parameters, such as H conversion, cold gas efficiency, tar content in the product gas, and product gas molar composition.

Introduction There are various kinds of refuses or residual wastes generated in food and drink industries. While some of them are generally reused as feeds for animals, such as pigs, cattle, and horses, the others that are not suitable to be animal feeds are usually disposed as wastes or sometimes as fertilizer or combusted to generate heat. The typical examples of the latter type are coffee grounds, tea grounds, vinegar lees, and bagasse. All of these materials represent a kind of concentrated biomass resource, making them easy be used as an efficient CO2 neutralizer. From this point of view, the development of high-efficiency energy conversion technologies for these biomass materials becomes highly necessary. Under financial support of The New Energy and Industrial Technology Development Organization (NEDO), Japan, a technical program was started since 2003 to work on the conversion of such biomass materials into middle-caloric product gases.1 Because the rude materials contain water of higher than 60 wt %, the concerned technology consists of an upstream drying of rude fuel and a downstream gasification of the dried fuel. A technique of slurry dewatering in kerosene2 was adopted to dry * To whom correspondence should be addressed. Telephone: 0081-45759-2867. Fax: 0081-45-759-2210. E-mail: [email protected]. (1) Central Research Institute of Electric Power Industry (CRIEPI). Summary of Research Achievements for the First Half of 2004 on “HighEfficient Energy Conversion Technology Development for Biomass”; CRIEPI: Tokyo, Japan, 2004. (2) Mito, Y.; Komatsu, N.; Hasegawa, I.; Mae, K. In 2005 International Conference On Coal Science and Technology (ICCS&T); Yamada, O., Ed.; IEA-Clean Coal Center: London, U.K., 2005; paper 2E01.

and also upgrade the rude high-moisture fuel. With drying part of the fatty matters of the fuel can be reformed. The resulting dry fuel thus has not only a lower water content of a few percents but also a higher heating value and higher carbon/ oxygen ratio.2 The so-called dual fluidized bed gasification (DFBG) technology was adopted to gasify the dried fuel to produce gases having heating values above 3000 kcal/m3n. Because the DFBG asks a combination of two fluidized bed reactors, it may have different technical arrangements depending upon the types of employed fluidized beds.3 We recently found that the deployment of fuel gasification into a dense bubbling fluidized bed and char combustion (for endothermic heat) into a pneumatic riser is the superior technique for DFBG.3 The designed DFBG according to this philosophy is similar to the FICFB biomass process4 developed at the Vienna University of Technology, Austria. The development of the DFBG plant was started with gasification tests in a 5.0 kg/h pilot facility by using coffee grounds as a model fuel. Because this laboratory-scale facility cannot run thermally independently, the tests hoped to show the chemically possible fuel conversion efficiency and to evaluate the quality of the resulting product gas for coffee grounds dried/ upgraded in advance. A series of tests were carried out to investigate the parametric influences on the gasification performance expressed as the efficiency converting fuel into product gas and (3) Xu, G.; Murakami, T.; Suda, T.; Matsuzawa, Y.; Tani, H. Ing. Eng. Chem. Res. 2006, 45, 2281-2286. (4) Pfeifer, C.; Rauch, R.; Hofbauer, H. Ing. Eng. Chem. Res. 2004, 43, 1634-1640.

10.1021/ef060120d CCC: $33.50 © 2006 American Chemical Society Published on Web 10/28/2006

2696 Energy & Fuels, Vol. 20, No. 6, 2006

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Figure 1. Schematic diagram of the employed pilot DFBG facility.

the tar evolution with gas conversion. This paper delivers the first comprehensive report on such gasification tests and their results. After the running characteristics and mass balance conditions of the pilot facility were evaluated, a few major operating parameters were varied to elucidate how the parameters affect the product gas composition, fuel conversion efficiency into product gas, and tar content in the gas. All of these data not only characterize the gasification performance of DFBG for the fuel coffee grounds but also guide further studies required to improve the fuel conversion efficiency and suppress the tar evolution. Experimental Section Apparatus and Control. Figure 1 shows a schematic diagram of the 5.0 kg/h pilot DFBG facility employed for the tests reported herein. It consists of a pneumatic riser char combustor and a bubbling fluidized bed (BFB) fuel gasifier. The riser of the facility was 52.7 mm i.d. and 6400 mm high. Its BFB had a rectangular cross-section of 80 × 370 mm2 and a height of 980 mm, with a 700 mm high expanded section of 180 × 370 mm2 in the crosssection area as its freeboard. Both the downcomer between the cyclone and BFB of the riser and the duct bridging the bottoms of the BFB and riser had the same i.d. of 52.7 mm. A seal made in the left-hand side of the BFB prevented the gas leakage from the riser to BFB, while the immersion of the downcomer into the particle bed of the BFB avoided the flowing of gas into the downcomer from BFB.5 The BFB and riser were both electrically heated, and temperatures of up to 1173 K were possible. Both the reactors had their respective exhaust lines consisting of a cyclone, a gas condenser, a bagfiler, and an induction draft fan (IDF). Thus, (5) Xu, G.; Murakami, T.; Suda, T.; Matsuzawa, Y. Ing. Eng. Chem. Res. 2005, 44, 9347-9354.

Table 1. Properties of Coffee Grounds Referred to in this Paper

proximate (wt %) moisture VM FC ash ultimate (wt %) C H N S O HHV (MJ/kg-db) bulk density (kg/m3) particle size

number 1

number 2

0.5 71.8 16.7 1.0

9.3 69.4 19.3 2.0

52.97 6.51 2.80 0.05 36.62 5260

54.9 6.12 3.07 0.03 33.62 5682 around 350 see Figure 2

independent control of the pressures inside the BFB and riser was possible. This was implemented through adjusting the amount of sucked shortcut air of each line. To diminish combustible-gas emission with venting, a gas combustion tube (GCB) was set between the cyclone and condenser of the exhaust line of the gasifier. Listed in Table 1 are the major property parameters of the two kinds of coffee grounds tested in this paper. Of them, fuel number 1 was generally employed unless a statement was made for the use of fuel number 2 (actually only used in Figure 8b). Both of the fuels were dried/upgraded with the aforementioned slurry dewatering technology.2 The major difference of the fuels was particle size. Figure 2 compares the size distributions of the two kinds of fuels measured with sieving. While fuel number 1 had a size d50 of about 520 µm, this size approached 900 µm for fuel number 2. Especially, about 60 wt % of fuel number 2 had sizes over 850 µm. Both of the fuels had a water content of about 10 wt % and were rich in volatiles and oxygen (see Table 1). A table feeder was used to feed fuel quantitatively and continuously, and the fed fuel was carried into the freeboard of the BFB gasifier via an argon stream monitored

DFBG of Coffee Grounds

Figure 2. Particle-size distribution of the tested fuels specified in Table 1.

in a mass flow controller. The argon stream was also a gas tracer employed in determining the volume (i.e., moles) of the produced gas. Steam at about 673 K was the gasification reagent adopted, while air preheated to 573 K in a spot heater was applied to the riser to combust the unreacted char coming from the gasifier. Air feed at room temperature was available also to the BFB gasifier so that during the startup and closedown of the facility both of the reactors could be fluidized with air. Thus, the airflows to both the beds were controlled with their respective rotameters. The flow of steam to the BFB was controlled through measuring the water amount in a water meter pump. To speed up heating, a supply of propane was made additionally for the BFB. Between the riser and BFB, silica sand (d50 ) 190 µm, and Fp ) 2600 kg/m3) was circulated as heat-carrier particles (CHPs). Nonetheless, for the small-scale pilot facility specified in Figure 1, the heat carried from the riser combustor to BFB gasifier via CHPs is unable to maintain the desired temperature of the gasifier (such as up to 1123 K) because of the excessive large heat loss of the facility. Therefore, circulation of HCPs in the present case simulates the chemical features more than the thermal characteristics involved in DFBG. The fact was also that electric heating controlled the temperatures inside the riser combustor and BFB gasifier of the pilot facility, as will be shown in the results via Figure 4. However, how many particles are circulated determines the available residence time of fuel particles inside the BFB gasifier, as fuel particle residence time ter ) particle mount inside the BFB (1) particle circulation rate Gs × A where ter refers to the explicit residence time estimated with a plug granular flow and A is the cross-sectional area of the riser. The residence time ter is critical to the conversion of fuel into product gas in the gasifier. Equation 1 indicates that ter is conversely proportional to the particle circulation rate Gs. As for the dual fluidized bed system shown in Figure 1, the particle circulation rate Gs is a function of not only Ug, the superficial gas velocity in the riser, but also the particle inventory in the whole system and the pressure difference between the BFB and riser.5 In our tests, we thus kept the latter two parameters for constants so that Gs was subject merely to Ug. To maintain a relatively steady pressure difference between the BFB and riser, the pressures in the tops of the BFB and riser were fixed respectively at about -1600 and -500 Pa in operating the beds (see Figure 5). Measurement. The product gas sample was taken via a suction pump just behind the cyclone of the gasifier. The sampled gas entered a tar trap system (detailed in the next paragraph) to have its tar-intake dropped. Until the tar traps, the gas was warmed to above 523 K via external heating with tape heaters to prevent tar deposition. The tar-stripped sample gas was finally led to a microGC for measuring its molar composition after the gas was detected for volumetric flow rate (1-2 Ln/min) in a wet gas volumeter and dewatered in a CaCl2 column. Analysis of the flue gas from the

Energy & Fuels, Vol. 20, No. 6, 2006 2697

Figure 3. Particle circulation rate and the corresponding explicit residence time of particles in the BFB gasifier estimated according to the plug granular flow.

riser combustor for molar composition was in an infrared gas analyzer. A sampled gas stream (about 2 Ln/min) was similarly created with gas suction at the exit of the riser’s cyclone. Before entering the gas analyzer, the sampled gas was first dewatered by passing through a CaCl2 column. The adopted tar trap system was made of, in succession, a water condenser and four water bubblers immersed in an ice-water bath. The condenser worked at temperatures below 278 K so that nearly all of the water and most of the tars carried with the sample gas could be dropped in this step. On the other hand, when the gas flow rate (1-2 Ln/min) was controlled at reasonable values, it was guaranteed that the gas residence time inside the tar traps was longer than 2 min. Hence, we believed that the tar trapping from the sample gas was efficient (capture efficiency > 90%), although not thoroughly completed. This was verified by further catching the tars escaping from the mentioned tar trap system into a fabric filter. The trapped tars were extracted by following a procedure of collecting the tarry water, washing the condenser, water bubblers, and silicon rubber tubes (for connection) using acetone, filtrating the acquired water/acetone liquid, vacuum evaporating water and acetone from the filtered liquid (at temperatures below 333 K), and drying the resulting tars, i.e., the solid residual of evaporation, in an airflow of 323 K. The tar content of the product gas reported herein refers to the ratio of the extracted dry tar weight over the standard-state (273 K and 1.0 atm) sample gas volume determined according to the measured total volume of sampled gas in a wet gas volumeter during tar trapping and the molar composition of the sample gas, as tar content ) extracted dry tar weight (2) meter-measured gas volume × (1.0 - tracer gas fraction) The extracted dry tar was also used to analyze its element composition and molecular weight. To determine ter according to eq 1, the particle circulation rate Gs under actual hot running conditions of up to 1173 K was measured in the facility sketched in Figure 1 using a heat-resistant switch valve.6 The acquired results demonstrated that the operating temperature did not evidently affect the particle circulation rate so that Gs was subject to the same control rules as at room temperatures. Thus, Figure 3 replots the average of our measured Gs (left y axis) versus Ug, the superficial gas velocity in the riser under the actual temperature of the riser. In measuring such Gs, the total particle amount loaded to the facility, which is termed particle inventory, was 24.0 kg and the pressures for the BFB and riser were controlled according to their values specified formerly in “Apparatus and Control”. From the pressure drop over the BFB (see Figure 5), we found that the particles held in the BFB were about 20 kg, thus giving the fuel particle residence time ter displayed in Figure 3 with its right-side ordinate (right y axis). Consequently, for all tests (6) Xu, G.; Murakami, T.; Suda, T.; Matsuzawa, Y.; Tani, H. AIChE J. 2005, 52, 3626-3630.


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Figure 4. Bed temperatures inside the BFB gasifier and riser combustor: (a) typical time series and (b) axial profile of the time-averaged value. Table 2. Performance Evaluation for the Test Shown in Figures 4-6 product gas rate from gasifier (Ln/min)a tar content in product gas [g/m3n (with Ar)] tar composition (wt %)

flue gas rate from combustor (m3n/h) CO2 content in flue gas carbon balance (wt %)b carbon conversion to gas (wt %) carbon combusted in riser (wt %) carbon in tars (wt %)

48.2 35.2 moisture, 0.1 C, 64.4; H, 6.2 S, 3.01; N, 11.1 O, 15.2; Cl, 0.11 9.2 8.5 97.3 66.1 27.0 4.2

a The product gas referred to is the raw gas containing Ar tracer and O 2 (about 0.5 mol %) and N2 (about 3.0 mol %) contaminates. b The left 2.7 wt % should exist mainly in the dust elutriated from the gasifier. The actual measurement shown that this part of C amounted to 3-5 wt %.

reported herein, only Ug was specified because one can simply determine the corresponding Gs and ter according to Figure 3.

Performance Evaluation Figures 4-7 and Table 2 characterize the running status of a test at about 1073 K. In this test, the fuel feed rate was 3.27 kg/h and the steam rate into the gasifier was 3.4 kg/h (thus, F/S ) 1.04). The latter led to a superficial gas velocity of 0.16 m/s in the particle bed of the BFB gasifier at its operating temperature of 1070 K. The superficial gas velocity Ug inside the riser combustor was 5.1 m/s, giving a fuel residence time ter of about 120 s in Figure 3. The fuel feed was started at the plotted time 0, and the test lasted for about 55 min. Running Characteristics. Figure 4 shows a few typical transient temperature profiles (Figure 4a) and the axial distributions of temperatures (Figure 4b) pertaining to both the reactors. The fuel feed induced a slight temperature decrease in the particle bed of the BFB (shown with T3), indicating just the fact that endothermic fuel gasification reactions occurred inside the bed. In the top of the BFB, there was an initial increase in

Figure 5. Pressures inside the BFB gasifier and riser combustor for the same test specified in Figure 4: (a) typical time series and (b) axial profile of the time-averaged differential pressure drop.

temperature until 15 min. This was rather due to the fact that the freeboard of the gasifier did not reach its set temperature of 1103 K before the fuel feed was started. Figure 4a clarifies that the temperatures T1-T4 remained to be stable during the test, except for an initial period of about 15 min in which some dynamic variations could be identified. Thus, temperature values averaged in the last 40 min of the tested time were plotted in Figure 4b to typify the axial distributions of temperatures inside the riser (O) and BFB (b). There was a gradual increase in temperature with the bed height in the 2.0 m long bottom section of the riser, whereas in the upper section the temperature was almost axially uniform. The result complies with the facts that heating both of the particles from the gasifier and fresh air supplied to the riser proceeded in the riser bottom until they both reached the uniform temperature of the riser in the upper bed section. The nonuniformity of temperature inside the particle bed of the BFB was within 10 K, with the lower value at a lower bed elevation (see the lowest three b in Figure 4b). Heating freshly supplied steam and moving particles from the gasifier to the combustor to create the particle circulation of the bed had to cause the temperature in the BFB to vary in such a way. The freeboard temperature T4 in the BFB was subject basically to its setting so that little fluctuation can be identified for it in Figure 4a. Hence, the temperature averaged over the four lower points in the BFB and that over all of the points above 2.0 m in the riser were adopted to be the operation temperatures, T0 and Tg, of the BFB and riser, respectively. Figure 5 displays the transient pressures measured at the top and bottom of the riser and BFB (Figure 5a) and the axial profiles of averaged differential pressure drops in both of the beds (Figure 5b). Evidently, during the tested time, all of the displayed pressures were basically stable, with plus and minus values at the bottom (P1 and P3) and top (P2 and P4), respectively. The pressures in the riser (P1 and P2) exhibited much higher fluctuations than in the gasifier (P3 and P4), indicating that the riser was in a relative dense condition. The axial profile of the time-averaged differential pressure drop in Figure 5b verified this by showing a denser bottom (O, below 1.0 m). As


Figure 6. Time series of molar concentrations of the major gas components in the effluent gases from the (a) BFB gasifier and (b) riser combustor for the test specified in Figures 4 and 5.

it should be, the bottom of the BFB gasifier had the highest differential pressure drop (b) to indicate that the bed was in the dense conditions close to bubbling/turbulent status. According to Figure 5a, we can confirm that in this test the pressures at the tops of the gasifier (P4) and riser (P2) were about -1600 and -400 Pa, respectively. These indicate rightly the conditions stated in the Experimental Section. Furthermore, the pressure drop over the BFB gasifier, which is P3-P4, was about 7000 Pa, thus determining a particle holdup of about 21 kg in the BFB, if supposing that the pressure drop is fully due to the weight of particles. While this complies with the loaded particles (inventory) of 24.0 kg, it also explains why in Figure 3 we used a particle holdup of 20 kg to calculate ter. Figure 6 shows the time series of molar concentrations of major gas components in the effluent gases from the BFB gasifier (Figure 6a) and riser combustor for the same test. The right-side ordinate (right y axis) of Figure 6a also mentions the higher heating value (HHV) of the product gas from the gasifier. Obviously, the reactions inside the gasifier quickly reached their quasi-steady state so that the resulting product gas composition manifested little variation from 15 min after the fuel feed started at time 0 [particularly on the basis of the CO concentration (4) and the HHV of the gas (×)]. The response to the flue gas concentration at the combustor exit exhibited a time delay (Figure 6b), because of the long sampling line made for this reactor. The detected CO2 (left y axis) and O2 (right y axis) concentrations, however, became almost steady after 12 min of fuel feed. Between 25 and 48 min, the CO2 content exhibited some higher values. The decreases in the airflow rate into the riser should be the cause because this flow was controlled via a rotameter so that unsteadiness in the flow rate was inevitable. Notwithstanding, throughout the tested time, the sum of O2 and CO2 concentrations in the flue gas remained to be a constant close to 21 vol % (right y axis in Figure 6b). The result actually demonstrates that the combusted matters inside the combustor was C exclusively, implicating that the BFB gasifier enabled the fuel fed into the reactor to be completely pyrolyzed to release


its volatiles as product gas and tars. As for the quoted test, the product gas was rich in CO (see Figure 6a), which reached 28 vol % or so in the raw gas (i.e., the gas with argon tracer). Succeeding to CO, the concentrations of the other gas species presenting in the product gas adhered to an order of gradual decreasing according to H2, CH4, CO2, C2H4, C2H6, and C3H6. The total amount of hydrocarbons (CmHn) was higher than 20 vol %. On the other hand, summing all of the displayed gas species gives a molar content of about 74 vol % so that the product gas had a HHV of about 3800 kcal/m3n. These demonstrate just the characteristic of the DFBG technology, which is the production of middle-caloric product gas that is little diluted with N2 and CO2 from in-situ combustion, even though air is employed to generate the endothermic heat required for fuel gasification reactions. Further Evaluation. Table 2 lists a few other parameters measured, including the tar content in the product gas (free of Ar tracer), tar element composition, and flow rates of product gas and flue gas. The flow rate of flue gas was supposed to be equal to the rate of supplied air, while that of the product gas was determined according to the given flow rate and measured molar concentration of the argon tracer. The raw product gas with the Ar tracer contained tars as high as 35.2 g/m3n, a content much higher than that reported for the 8 MW FICFB biomass gasification plant4,7 but in the same order shown in Caballero et al.8 for a laboratory-scale pilot BFB gasifier. The higher tar evolution in our case may be relative to the fuel itself, which is highly tar-productive, and to the fuel feed into the freeboard of the gasifier. Corella et al.9 showed earlier that the feed of biomass fuel into fluidized particles can greatly facilitate gasification and tar destruction reactions to allow for higher gasification efficiency and lower tar production. When the element composition of tars in Table 2 are compared with that of fuel in Table 1 (number 1) we can see that the tars contained much more S and N, evidently reduced O, slightly increased C, and weakly decreased H. These actually clarify that during gasification elements S and N are likely to be concentrated into tars, while the high-molecular-weight tars may have fewer carboxylic groups than the virgin fuel. The higher C and lower H contents in the tars than in the fuel imply that the release of fuel H into gas is quicker than that of C, complying with the fact that in the gasification of a fuel the conversion into the product gas is usually higher for H than for C, as one will see from Figures 8 and 9 reported in the coming section. Table 2 also shows a mass balance evaluation for the quoted test with C as the representative element, where the C conversion was estimated according to

element conversion Xi ) moles of the element i in the produced gas × 100%, moles of the element i in the supplied fuel with i ) C or H (3) Principally, the element C fed into the system with fuel is converted into three streams: the product gas, flue gas, and tars. The part of C converted into the product gas is shown with the C conversion computed from eq 3, while the C in the flue gas can be determined with the flue gas flow rate and CO2 (7) Results of the 100 kW Pilot FICFB Plant, http://www.ficfb.at/ results.htm. (8) Aznar, M. P.; Caballero, M. A.; Gil, J.; Martı´n, J. A.; Corella, J. Ind. Eng. Chem. Res. 1997, 36, 5227-5239. (9) Corella, J.; Herguido, J.; Alday, F. L. In Research in Thermochemical Biomass ConVersion; Bridgwater, A. V., Kuester, J. L., Eds.; Elsevier Applied Science: London, U.K., 1988; pp 384-397.


Figure 7. Molecular-weight distribution of the tars sampled in the test characterized in Figures 4-6.

concentration in this gas (e.g., 8.5 vol % on average from Figure 6b). In the test, we also confirmed that CO in the flue gas was less than 500 ppm. The carbon with tars is equal to the product of the tar content in the produced gas, produced gas volume, and C percentage in the tars, of which all are available in Table 2. Consequently, we finally had the C fractions shown in the lower side of Table 2, which are defined against the imported C with the fuel feed. They are 66.1 wt % in the product gas, 27.0 wt % in the flue gas, and 4.2 wt % presenting in tars. The overall sum of these percents was 97.3 wt %, showing a good mass balance condition for this test. On the other hand, collecting gasifier ash and further analyzing its C intake corroborated that the escaped C from the gasifier was rarely over 5.0 wt % of the C supply. Shown in Figure 7 is the molecular-weight distribution of the tars from this test. It was measured with gel-permeation chromatography (GPC) at a column temperature of 313 K. In the tars, there are molecules with a weight of up to 10 000 but most have weights below 3000. Generally, the lower the molecular weight, the more molecules in the tars (shown with a higher intensity). This leads the average molecular weight weighted with molecule numbers (Mn) to be 420 and that weighted with molecular weight (Mw) to be 830. Parametric Investigation Herein, the influences of the gasifier temperature, fuel particle size, steam/fuel ratio, explicit fuel residence time in the gasifier, and air supply into the gasifier will be examined. Table 3 summarizes the testing conditions and product gas compositions of the involved tests R1-R9. Their realized C and H conversions, cold gas efficiencies, and tar contents in the produced gases are compared in Figures 8 and 9. The quoted cold gas efficiency ηe refers to an energy conversion efficiency based on HHV, as

cold gas efficiency ηe ) total HHV of produced gas × 100% (4) total HHV of supplied fuel Gasifier Temperature. The involved tests are R1 and R2, whose performance data are plotted in Figure 8a. Both the tests had the same fuel feed rate F and steam/fuel mass ratio S/F (see Table 3). In comparison with the gasifier temperature T0, the characteristic temperature Tg of the riser combustor was usually 15 K higher, making it be about 1000 and 1110 K corresponding to the tested T0 of 993 and 1093 K, respectively. Because the tests adopted fixed rates of airflow to the riser and steam to the BFB, the resulting actual gas velocities inside the riser (Ug) and BFB (U0) were slightly different between R1 and

Xu et al.

R2. Overall, however, we can see that U0 was about 0.15 m/s and Ug was slightly below 3.0 m/s. Comparing the gas composition for R1 and R2 in Table 3 clarifies that the pure product gas excluding the Ar tracer, N2, and O2 contained more CO and hydrocarbons (C2 and C3 species) but less H2 at the lower gasifier temperature T0. This is indicative of the fact that the produced gas from a lower gasification temperature had a higher heating value. However, the conversions of fuel C and H (i.e., Xc and Xh) and the available cold gas efficiency ηe in Figure 8a were conversely lower at the lower temperature, implying less gas production from treating every kilogram of fuel. Furthermore, the tar content in the produced gas free of the Ar tracer (right y axis in Figure 8a) was much higher at the lower T0. All of these showed just what we anticipated and clarified essentially that increasing gasification temperature facilitated fuel conversion and suppressed tar evolution, although it slightly reduced the heating value of the produced gas. Furthermore, one can see that, under the typified gasifier temperature of about 1093 K and the usual steam/fuel mass ratio of about 1.0, gasifying the fuel coffee grounds with the adopted DFBG technology can convert up to 70% of fuel C into product gas. The corresponding energy recovery efficiency approached 75%. However, it is noteworthy that the produced gas was contaminated with a relatively high tar content of up to 40 g/m3n. Fuel Particle Size. The influences of the fuel particle size can be seen from the results of tests R3 and R4 shown in Table 3 and Figure 8b. The tested fuel sizes were those specified in Figure 2 (numbers 1 and 2). Although the test R4 regarding the coarser fuel number 2 had a slightly higher gasifier temperature T0 and lower gas velocity Ug in the riser, Figure 8b clarifies that gasifying this larger size fuel had led to evidently lower conversions of fuel C and H and lower cold gas efficiency. The tar content in the produced gas was also distinctively higher for the test R4. Notwithstanding, there was not a great difference in the normalized product gas composition. From Table 3 we may see strictly that the produced gas from the finer fuel number 1 (R3) had a slightly lower H2 fraction and somehow higher CO and hydrocarbon concentrations. When recognizing the gasifier temperature influences delineated above (Figure 8a), we may suggest that such slight differences in gas composition are rather relative to the lower gasifier temperature for R3 (1070 K against 1085 K for R4). Therefore, the fuel particle size likely affects more fuel conversion efficiency and tar production than product gas composition. One may also note that the fixed-carbon (FC) ratio was about 2 wt % higher for the coarser fuel number 2. The realized C conversion in Figure 8b, on the other hand, was roughly 6% lower for this fuel, even if its T0 was slightly higher. Therefore, the preceding efficiency difference between the tests R3 and R4 should be due to the different sizes of the fuels rather than their different FC contents. Steam/Fuel Ratio. The tests R5 and R6 were conducted to demonstrate the influences of the steam/fuel mass ratio (S/F). Figure 8c compares the acquired gasification performances. The two tests had almost the same condition parameters except that U0 inside the gasifier was differentiated with their employed different S/F ratios. At the higher steam/fuel ratio of 1.45, U0 was surely higher, which led to a shorter residence time of gas inside the particle bed of the gasifer. Notwithstanding, Figure 8c clarifies that elevating the S/F ratio not only facilitated fuel (Xc and Xh) and energy (ηe) conversions but also lowered the tar content in the produced gas. In response to the elevation of the S/F ratio from 0.84 to 1.45, the actually achieved increments



Figure 8. Influences of the (a) gasifier temperature T0, (b) fuel particle size, (c) steam/fuel ratio S/F, and (d) explicit fuel residence time ter in the gasifier on the gasification efficiency and tar evolution. The test IDs R1-R7 in the parentheses refer to the run IDs compiled in Table 3. Table 3. Conditions and Normalized Composition of Resulting Product Gases for All Tests Referred to in the Parametric Investigation run ID

R1

R2

R3

R4

R5

Test Conditions number 2 number 1 0.0 0.0 3.1 3.3 1.10 0.84 1103 1105 1085 1093 4.8 2.9 0.16 0.13

R6

R7

R8

R9

number 1 0.0 3.3 1.45 1105 1093 2.9 0.23

number 1 0.0 3.7 0.92 1103 1080 2.9 0.16

number 1 0.0 2.8 1.21 1073 1053 2.9 0.15

number 1 0.09 2.8 1.21 1073 1073 2.9 0.20

22.09 37.38 12.31 17.76 6.82 3.55 0.09 0.0

21.21 37.29 12.08 15.87 8.08 4.12 0.37 0.0

19.26 37.74 16.54 14.51 7.71 3.91 0.33 0.0

fuel O2/C F (kg/h) S/F Tg (K) T0 (K) Ug (m/s) U0 (m/s)a

number 1 0.0 3.8 1.04 1008 993 2.6 0.15

number 1 0.0 3.8 1.04 1108 1093 2.9 0.16

number 1 0.0 3.27 1.04 1085 1070 5.1 0.16

H2 CO CO2 CH4 C2H4 C2H6 C3H6 C3H8

16.71 41.62 12.80 16.58 7.17 3.70 1.31 0.11

23.24 37.80 12.55 16.89 6.27 3.24 0.02 0.0

Normalized Product Gas Composition (mol %)b 20.96 23.94 22.74 24.31 38.04 37.32 38.14 37.39 12.05 12.21 12.06 12.31 17.34 16.14 17.18 15.68 7.62 6.74 6.50 6.77 3.85 3.55 3.30 3.44 0.14 0.01 0.08 0.10 0.0 0.0 0.0 0.0

a For steam and the adopted sand particle, the U was estimated to be about 0.04 m/s at 1073 K. b The normalization of gas molar composition was made mf against all of the listed gas species so that the listed concentration values can be summed to 100% for each of the compared cases.

in the fuel and energy conversions were 5-10%, while the tar content in the product gas decreased by about 7 g/m3n. Thus, adopting a higher steam/fuel mass ratio would chemically enhance the gasification reactions to increase gas production. In practical gasification processes that run thermally independently, however, the increase in the steam feed ratio may have to decrease the energy conversion efficiency.7 This is because the higher steam feed requires more energy to vaporize water, but the latent heat with the steam can hardly be recovered behind the gasification reactor. The normalized gas composition in Table 3 does not have great differences between the tests R5 and R6, but it is obvious that the higher S/F ratio made the concentrations of H2, CO2, and hydrocarbons (above C2) slightly higher and those of CO and CH4 somehow lower. These indicate

possibly that the water-gas-shift reaction was greatly enhanced with the increased steam feed, whereas at the higher S/F ratio the shorter residence time of produced gas inside the gasifier did not allow CmHn (above C2) to be efficiently reformed. Accordingly, the cause for the relatively lower CH4 concentration at the higher S/F ratio may be that CH4 was hardly reformed in both the tested cases so that the totally same CH4 production from fuel pyrolysis had to appear with a lower molar concentration when the total volume of produced gas increased with increasing S/F. Fuel Residence Time. The compared data in Figure 8d are based on the tests R3 and R7, whose superficial gas velocities Ug in the riser combustor, in turn, their explicit residence times ter of fuel particles in the gasifier determined according to eq 1,


Xu et al.

Figure 9. Influences of a small amount of air fed to the gasifier on the gasification efficiency and tar evolution (test R8 against test R9).

were distinctively different from each other. According to Figure 3, we can actually have the ter values, which are 120 and 950 s for Ug of 5.1 m/s (R3) and 2.9 m/s (R7), respectively. Although the residence time ter increased by about 8 times for the test R7, Figure 8d indicates that the achieved improvement on fuel and energy conversion efficiencies was not very evident. This result complies with the facts that for biomass gasification the available fuel conversion (C and H) is subject exclusively to fuel pyrolysis, which should finish in several tens of seconds.10,11 In our case, the fuel pyrolysis would be definitely completed in both of the examined cases such that prolonging ter from 120 to about 1000 s was influential only to char steam gasification. The kinetic rate of this reaction, on the other hand, is not very quick at 1083 K, thus leading to the little upgraded fuel and energy conversion efficiencies shown in Figure 8d. The longer residence time ter, however, resulted somehow in an obvious suppression on tar evolution, which caused the produced gas to have about a 20% decrease in the tar content (right y axis in Figure 8d). The cause may be that inside the gasifier there were more char particles at the longer ter. With the catalytic action of char particles on tar destruction reactions, it might allow more tars to be destroyed via reforming and decomposition. From Table 3, we can see also that the produced gas at the longer ter had slightly higher concentrations of H2, CO2, and CH4 but lower contents of CO and hydrocarbons (above C2 species). These possibly show again an enhanced effect of char catalysis on tar and hydrocarbon reforming/ decomposition reactions at the longer ter. Enhancing these reactions produces more H2 and CO2, while reforming/decomposing tars would release CH4 as well. The formed CH4, on the other hand, can hardly be reformed or decomposed under the tested conditions, hence making its higher concentration in the produced gas. Nonetheless, all of these variations in gas composition are negligibly small, as will be further clarified in Figure 10. Air Feed into the Gasifier. How a feed of a small amount of air into the gasifier influences the gasification performance was demonstrated in the tests R8 and R9, whose performance data were compared in Figure 9. Air feed for R9 was according to a ratio of 0.09 of oxygen to fuel C. The tests R8 and R9 (10) Murakami, T.; Xu, G.; Suda, T.; Matsuzawa, Y.; Tani, H.; Fujimori, T. Fuel 2007, 86, 244-255. (11) Bingyan, X.; Chuangzhi, W.; Zhengfen, L.; Xiguang, Z. Solar Energy 1992, 49, 199-204.

Figure 10. Exclusive dominance of fuel C conversion over the (a) other performance parameters and (b) normalized gas composition for steam gasification of a given biomass fuel.

were under the same fuel feed rate F, steam/fuel mass ratio S/F, and gas velocity Ug in the riser, but the fed air for R9 caused a slightly higher gasifier temperature T0, 1073 K against 1053 K for R8. From Figure 9, we can see that the application of air to the gasifier increased the C conversion but the corresponding H conversion conversely exhibited a tendency to decrease. Consequently, the product gas compositions in Table 3 revealed lower contents of H2, CH4, and hydrocarbons but obviously higher concentrations of CO2 for the test R9 with air intake. On the other hand, the overhead table of Figure 9 clarifies that in the raw product gas both the tests R8 and R9 had similar O2 content (about 0.3 mol %), whereas the inclusion of air in R9 caused an obviously higher N2 concentration of about 18.5 mol %. Hence, the test R9 had an evidently lower HHV of about 3000 kcal/m3n compared to about 3800 kcal/m3n for R8. The higher CO2 content shown in Table 3 for the test R9 with air feed reveals that the burning of C truly occurred in the gasifier. This combustion should occur to H2 as well, but it did not reduce the total H2 production at the examined O2/C ratio very much so that two very similar H conversions are displayed in Figure 9. Because the normalized CO content was slightly higher for R9, its corresponding lower concentrations of CH4 and hydrocarbons might be due to the gas volume that increased with burning C into CO2 and CO. As expected, with air input, the tar content in the produced gas obviously decreased by about 60%. Furthermore, the air inclusion elevated the cold gas efficiency ηe at the examined O2/C ratio of 0.09. Consequently, the application of an airflow to gasifier according to O2/C ratios of about 0.1 likely creates some benefits to biomass steam gasification. While it increases the energy conversion efficiency and suppresses tar production, the resulting product gas can maintain also its HHV at about 3000 kcal/m2n. Cross Correlation. As a summary of the preceding parametric investigations, Figure 10 intercorrelates the gasification performance parameters (Figure 10a) and normalized molar concentrations of major gas species (Figure 10b) with fuel C conversion. The correlation was made basically for steam gasification. Thus, the C and H conversions and normalized H2 and



CO2 concentrations from the test R9 with an air feed into the gasifier were not included into the plot (because these parameters obviously varied with the air feed). Figure 10 demonstrates that the variations of all of the plotted parameters with the operating parameters examined above are likely subject to the realized C conversion Xc exclusively. That is, when gasifying a given kind of fuel with steam, as long as the C conversion is known, it is possible to determine roughly the other parameters based on the relationships clarified in Figure 10. The figure shows that for coffee grounds the tar content in the produced gas gradually decreases, whereas the H conversion Xh and cold gas efficiency ηe proportionally increase with raising the C conversion. Meanwhile, Figure 10b shows that the normalized H2 content linearly increases and the normalized concentrations of CH4, CO2, and hydrocarbons (represented with C2H4) remain to vary little with the rise of C conversion. The normalized CO fraction generally decreases with increasing the C conversion, but it possesses a gradually lower speed to decrease when Xc gets higher. This made the CO concentration tend to be a constant when Xc is over 60%. Although the specified relationships in Figure 10 are valid probably only to steam gasification of coffee grounds at steam/fuel mass ratios of about 1.0, the implicated methodology of correlation should be generally applicable to arbitrary fuels and gasification technologies. Conclusions Gasification of coffee grounds, a kind of biomass wastes, was investigated in a 5.0 kg/h pilot DFBG facility to demonstrate the gasification characteristics of the fuel in chemical aspects. The investigated DFBG consists of a BFB gasifier and a pneumatic riser char combustor, between which silica sand particles were externally circulated to carry combustion heat from the combustor to gasifier to maintain the endothermic fuel gasification reactions occurring in the gasifier. Steam was the gaseous reagent fed to the BFB gasifier, while air was applied to the riser combustor to combust unreacted char from the gasifier. The obtained principal results from the work are as follows: Overall. Evaluation of the performance of the pilot gasification facility demonstrated that the employed DFBG technology worked stably with the fuel of coffee grounds that was predried to a water content of about 10 wt %. Under the conditions of a gasifier temperature of about 1073 K, an explicit residence time of about 120 s of fuel particles inside the gasifier, and a steam/ fuel mass ratio of about 1.0, more than 70% of fuel C was found to be able to be converted into product gas in the BFB gasifier, while up to 25% was burned into flue gas in the char combustor. In the char combustor, little H was combusted, causing the CO2 and O2 molar concentrations in the flue gas to have a sum of about 21 mol %. The produced raw gas (with tracer gas) had a HHV of about 3800 kcal/m3n, but it was with a heavy tar load of up to 50 g per cubic meter of tracer-free gas. The tars were rich in S and N but with an O content lower than that in the treated fuel, clarifying that during steam gasification the elements N and S contained in the virgin fuel are likely concentrated into tars and in the tars there are fewer carboxylic groups than in the fuel. The molecular weights of the tars were shown to be mostly below 3000, with their majority around 1000. Fuel Conversion. Parametric studies demonstrated that the fuel conversion into product gas through DFBG is subject mainly to the gasifier temperature and fuel particle size, while raising the steam/fuel mass ratio chemically increased but would thermally decrease the fuel conversion efficiency. In thermally independent gasification plants, the practically possible seam/

fuel mass ratio would be hardly over 1.0. Prolonging the explicit residence time of fuel particles inside the gasifier from 120 to about 1000 s did not cause an evident increase in fuel conversion, suggesting that the necessary residence time of the fuel coffee grounds inside the gasifier can be below 120 s. The implicated essence is that for biomass gasification the available fuel conversion is subject to fuel pyrolysis exclusively, which can finish in several tens of seconds. Besides, feeding air into the gasifier with an O2/C ratio of about 0.1 revealed an obvious enhancement to fuel C and energy conversions, while it did not decrease the fuel H conversion very much. Tar Production. Feeding air into the gasifier according to O2/C ) 0.1 reduced the tar release by about 60%, while it kept still the heating value of the product gas (HHV) at about 3000 kcal/m3n. Another major means to reduce the tar yield with steam gasification was to raise the gasifier temperature. Making fuel finer, increasing the steam/fuel mass ratio, and prolonging the fuel residence time in the gasifier all suppressed the tar release to a certain degree, but the available effects were small and limited. For the tested coffee grounds, the identified lowest tar content in the product gas was 25 g/m3n (tracer-free gas), thus suggesting that other advanced tar elimination techniques are required to upgrade the examined DFBG. Product Gas Composition. On a normalized base against gas species H2, CO, CO2, CH4, C2H4, C2H6, C3H6, and C3H8 (i.e., the concentration sum of all of these species are 100%), we found that obvious variations in the normalized molar concentrations of all of the gas species occurred mainly with C conversion. Consequently, the gasifier temperature and air feed into the gasifier appeared to be the major parameters affecting the normalized composition of product gas. Even so, for the fuel coffee grounds, its H2/CO ratios varied in a narrow range of 0.55-0.65 and its normalized CO concentration fluctuated merely in 37 and 38 mol %, provided that the gasification was with pure steam at about 1073 K and under steam/fuel mass ratios of about 1.0. Cross Correlation. For steam gasification of a given kind of fuel, the available C conversion likely determines the H conversion, cold gas efficiency, tar content in the product gas, and normalized gas composition measured under various operating conditions and for differently sized fuels. It was shown that the tar content in the product gas gradually decreased, whereas the H conversion and cold gas efficiency proportionally increased with increasing the C conversion. As for coffee grounds, raising the C conversion linearly increased the normalized H2 concentration but caused the normalized CO concentration to gradually decrease until a C conversion of about 60% was reached. No big variations were identified for the normalized CH4 and hydrocarbon concentrations (above C2 species) in the range of C conversion from 55 to 78%. Acknowledgment. The work was conducted during a technical program on biomass upgrading and gasification financed by The New Energy and Industrial Technology Development Organization (NEDO), Japan. The authors are grateful to Mr. Minoru Asai and Ms. Kumiko Uchida of the same company for their help in the experiment.

Nomenclature d50 ) particle diameter at a mass accumulation fraction of 50 wt %, m F ) fuel feed rate, kg/h Gs ) particle circulation rate, kg m-2 s-1

2704 Energy & Fuels, Vol. 20, No. 6, 2006 Pi ) pressure at a local position i ()1, 2, 3, or 4) corresponding to the riser bottom, riser top, BFB bottom, and BFB top, kg m-1 s-2 S ) steam feed rate into the BFB gasifier, kg/h ter ) explicit residence time of fuel in the BFB gasifier, s T0 ) characteristic temperature of the BFB gasifier, K Tg ) characteristic temperature of the riser combustor, K Ti ) temperature at a local position i ()1, 2, 3, or 4) corresponding to the riser bottom, riser top, BFB bottom, and BFB top, K U0 ) superficial gas velocity in the BFB gasifier at its operating temperature, m/s

Xu et al. Ug ) superficial gas velocity in the riser combustor at its operating temperature, m/s Xc ) conversion of fuel C into product gas, % Xh ) conversion of fuel H into product gas, % Greek Symbols ηe ) cold gas efficiency, % Fp ) particle density, kg/m3 EF060120D

Gasification of Coffee Grounds in Dual Fluidized Bed - American

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