Thermodynamic Limits and Actual Product Yields and Compositions in

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Ind. Eng. Chem. Res. 2006, 45, 834-843

Thermodynamic Limits and Actual Product Yields and Compositions in Biomass Gasification Processes Nader Jand, Vincenzo Brandani, and Pier Ugo Foscolo* Chemical Engineering Department, UniVersity of L’Aquila, 67040 L’Aquila, Italy

A method is described for significantly improving the predictive capability of equilibrium-based calculation tools for the estimation of fuel gas composition in high-temperature biomass gasification processes. It is based on acknowledgment of the fact that the overall thermochemical conversion occurs in two stages: fast biomass-particle devolatilization, followed by the much slower, and hence often nonequilibrium, conversion of methane and char. Inputs to the equilibrium calculation routine, modified in relation to this latter phenomenon, are proposed that can be readily implemented in commercial flow sheet calculation routines for chemical reactors. The method provides estimates for the yields of specific compounds, such as hydrogen and carbon monoxide, and is shown to predict well the performance trends relating to changes in operating conditions. None of the available kinetic models is able to offer comparable predictive capability. Nor too do thermodynamic equilibrium model methods that fail to take into account the two-stage nature of the gasification process and, as a result give rise to substantial deviations from available experimental data: underprediction of both the methane content in the product gas and the unconverted carbon in the solid phase; negative findings similar to these have been reported in analogous literature studies of coal gasification. 1. Introduction Biomass gasification has come to be regarded as a potentially competitive and sustainable process for the production of chemicals and gaseous fuels for power generation. It exhibits unquestionable environmental advantages with regard to the development of clean waste-to-energy conversion systems and makes no contribution to greenhouse gas accumulation in the atmosphere; it also leaves intact the available reserves of fossil fuel. The chemical efficiency of the gasification process (the ratio of the heating value of the gas produced to that of the fuel input) exceeds 70%, allowing for a potential overall electrical efficiency of over 40% in the more sophisticated applications, such as high-pressure gasification coupled with fuel-cell generation and combined-cycle operation.1 The critical effect of temperature on gasification product distribution (permanent gases, condensable organic vapors, and charcoal) is well established,2 so that modern processes for the generation of gaseous fuels now operate at above 800 °C. In addition to temperature, the nature and composition of the gasification agent (steam, air, oxygen, and combinations of these three components), the presence of appropriate catalysts,3,4 and the gasifier architecture itself can all greatly influence the gas yield, composition, and calorific value, spanning from a low heating value (LHV) producer gas, suitable for internal combustion engines, to a hydrogen-rich synthesis gas, amenable to constitute the anode feed of a high-temperature fuel cell (molten carbonate or solid oxide), as well as a sustainable source of hydrogen, after appropriate purification and separation stages. Steam gasification is a complex thermochemical conversion process, involving biomass particle devolatilization, steam reforming of the hydrocarbons thereby produced, and heterogeneous decomposition of the char resulting from particle pyrolysis. The gas yield and its hydrogen content are maximized by high-temperature operation, which promotes rapid heating of the biomass particles and the virtual absence of heavy, condensable organic molecules (collectively referred to as tar) in the product stream. * To whom correspondence should be addressed. Tel.: +39-0862434214. Fax: +39-0862-434203. E-mail: [email protected].

The overall endothermic nature of the pyrolysis process, together with the requirement of high-temperature operation, gives rise to the need for a substantial input of heat; this is traditionally provided by the addition of some air, resulting in the exothermic combustion of some of the system species. A major drawback of this procedure, however, is that it results in the dilution by nitrogen of the product gas, an effect that can be reduced by the use of oxygen-enriched air. Alternatively, the gasification and combustion zones, and their respective output streams, can be kept separate by adopting a circulating particle arrangement, as is employed in the fast internally circulating fluidized bed (FICFB) gasifier design.5 Fluidizedbed gasifiers are well suited to biomass gasification: they provide for temperature homogeneity, good gas and solid mixing, rapid heating of the biomass feedstock, and the possibility of including catalyst particles in the bed inventory to enhance the reforming reactions. A major goal of current research is the development of flexible models for gasification processes, capable of providing reliable quantitative predictions of product yield and composition. In a previously reported study, the dynamic behavior of biomass particles of various sizes dropped into a hot, inert-gas fluidized bed was investigated in some detail; this work resulted in the development of a highly simplified model for estimating the global yield of volatiles as a function of time, bed temperature, and particle size.6 More sophisticated kinetic modeling approaches have also been attempted, directed at predicting the yields of permanent gases, condensable heavy organic vapors, and charcoal, under various operating conditions.7 The problem, however, is that these approaches require the knowledge of a very large number of physical and kinetic parameters, because of the complexity of the pyrolysis reactions involved, many of which are simply not available. An even more comprehensive modeling approach has also been reported, in which the additional influence of the gasifier design can be taken into account; this involves including the kinetic and thermal aspects of the process in a computational fluid dynamic (CFD) simulation code representation of a fluidized-bed gasifier.8 Despite these and other ongoing research efforts, no method has yet emerged whereby the gasification yields of specific gas

10.1021/ie050824v CCC: $33.50 © 2006 American Chemical Society Published on Web 11/15/2005

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products (hydrogen, methane, carbon monoxide, etc.) can be predicted. Equilibrium models, which take into account only thermodynamic limitations, thereby disregarding specific reaction mechanisms and chemical and transport rate phenomena, have long been applied to the prediction of product yields and compositions in reforming processes involving hydrocarbons with possible carbon deposition.9 For high-temperature reacting systems, for which the behavior of the predominant gas phase closely approximates that of an ideal gas, their application is particularly straightforward and can be expected to yield predictions in close accord with experimental observation. Of particular value is the ability of such models to yield reliable predictions of the influence of changes in the operating conditions of real systems, thereby providing an invaluable tool for process design and development purposes. Applications of thermodynamic equilibrium models to gasification processes have been reported in the literature with reference to coal conversion10 and to a number of specific biomass gasification processes.2,11,12 The latter have been mainly concerned with parametric studies of the influence of operating conditions and fuel characteristics on the heating value of the product gas and on the overall behavior of particular gasification plant installations; they have not addressed the potential predictive capability of such methods with regard to such objectives as, say, the enhancement of the product-gas hydrogen content or the production of a syngas with a given CO/H2 ratio. It is worth pointing out at this point a feature of biomass that makes its thermal conversion somewhat different from the corresponding conversion of coal. When biomass particles are rapidly heated at high temperature (above 550 °C), more than 80% of their (dry) mass is rapidly converted into permanent gases and organic vapors, leaving only a small amount of char and very few mineral ashes in the solid phase; the gasification agent (steam, carbon dioxide) then ensures that the char is further consumed. Thus, as a result of its high proportion of volatiles, biomass is much more amenable to rapid conversion into a gaseous product than is coal. The above-mentioned biomass devolatilization study6 indeed confirmed biomass pyrolysis to be a relatively rapid phenomenon, even for particle sizes comparable to those of typical industrial applications, which can be on the order of centimeters. In the work described herein, the outputs of equilibrium models for biomass gasification are first examined, starting from the application of the “series reactor method” for heterogeneous systems to an ideal cellulose gasification process, for which all needed thermodynamic data are readily obtainable; the major compounds formed at equilibrium are calculated over a wide range of temperatures. Evaluations relating to an extended number of chemical species are then performed with reference to the gaseous phase alone, by means of the “Gibbs energy minimization” method applied to high-temperature gasification of biomass fuels of given elemental composition. The theoretical predictions are compared to the experimental results of benchscale tests performed under controlled conditions and over a wide range of variation of operating parameters. Modified inputs to the calculation method are then proposed on the basis of the discrepancies between the precise thermodynamic limits and the actual results; the method utilizes standard routines for chemical reactors at equilibrium, available in commercial software packages for process flow sheet calculations. Example applications of this method are also provided, with reference to two fluidized-bed gasification pilot plants; a comparison is

made between predictions and actual plant performance, given by the results of typical runs published in the literature. 2. Application of Thermodynamic Equilibrium Models to Biomass Gasification In the following discussion, thermodynamic calculations are applied to the determination of equilibrium compositions in gasification systems at constant temperature and pressure. Although in the majority of practical applications the gasifier operates continuously, under close-to-constant and, in the case of fluidized-bed reactors, also uniform, temperature and pressure conditions, the solid biomass fuel itself is being continuously heated from ambient to the gasifier temperature level; as a result of this and also of the presence of intraparticle heat- and masstransfer resistances, heterogeneous and homogeneous reactions occur over a wide range of temperatures. For these reasons, the equilibrium simulations considered in this section are not confined to the most common range of operation of industrial gasifiers (T ) 800-900 °C), but are extended down to about 300 °C, which represents the lower limit for devolatilization phenomena to occur; variable oxygen and steam concentrations are also accommodated in these simulations. In the case of simultaneous reaction systems, two different approaches to the solution of the simultaneous equilibria problem have been used: the series reactor method and the G-minimization method. 2.1. Series Reactor Method. The series reactor method13 is conceptually the easier of the two to understand; it is based on the fact that the Gibbs free energy of a system, G, at constant temperature and pressure, will always be reduced when any one reaction is allowed to proceed to equilibrium. A clear illustration of this statement is provided by Modell and Reid.14 It follows that, in a iterative procedure, if the reactions are allowed to proceed to equilibrium “sequentially”, G will always converge to a minimum value that corresponds to the equilibrium condition for the simultaneous series. The series reactor method is an iterative procedure in which we envisage r reactors in series, one for each independent reaction. In any given reactor, all species that do not take part in that reaction are treated as inert substances. In each reactor, the specific reaction is allowed to proceed to equilibrium. The procedure is repeated until the extent of reaction in each reactor has passed below some predetermined minimum value. As a first step, this procedure is applied to the air and steam gasification of pure cellulose, (C6H10O5)x, for which thermodynamic data are predicted from literature sources by application of a group-estimation method based on its structure.15 The input species are air, steam and cellulose, in fixed ratios. On the basis of experimental evidence and in view of a simplified description of such systems, it is assumed that hydrogen, carbon monoxide, carbon dioxide, methane, naphthalene, and graphitic carbon make their appearance in the evolution of the system toward equilibrium through the following set of independent stoichiometric relations:

(C6H10O5)x ) 5xCO + 5xH2 + xC(s)

(1)

CO + H2O ) CO2 + H2

(2)

CO + 3H2 ) CH4 + H2O

(3)

CH4 ) C(s) + 2H2

(4)

C10H8 + 10H2O ) 10CO + 14H2

(5)

2CO + O2 ) 2CO2

(6)

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Figure 1. Equilibrium constants for stoichiometric relations 1-6 as a function of temperature.

Figure 2. Gasification of cellulose: influence of temperature, SBR, and ER on the presence at equilibrium of a solid phase containing carbon.

Naphthalene is assumed to represent all of the organic compounds that are formed as a result of the primary pyrolysis processes and are expected to undergo partial oxidation by water in accord with the generalized form of eq 5:

pyrolysis of cellulose proceeds to completion. The first important result (which is in accord with experimental evidence accumulated with a large variety of biomass fuels) is therefore that the major effect associated with changing biomass type (we are considering here dry, ash-free, i.e., daf, biomass) is restricted to changing the ratios among the quantities of C, H, and O in the system. Figure 2 shows the influence of temperature, steam/biomass (cellulose in this case) ratio (SBR), and air/cellulose ratio (the latter expressed as the “equivalence ratio”, or ER, the actual value of the oxygen/cellulose ratio divided by that corresponding to complete biomass combustion) on the presence at equilibrium of a solid phase containing carbon. The isotherms represent the locus of operating conditions where 0.5% of the carbon in the feedstock is left in the solid phase: at the corresponding equilibrium temperature, higher percentages of solid carbon, C(s), are obtained for systems falling below the curve, whereas it is practically absent for systems above the curve. It is worth noting that the C/O ratio in the system largely determines this threshold limit: the intersections of each curve with the x and y axes (representing steam and air gasification, respectively) are both characterized by very similar values of this ratio, as well as the intermediate points, thanks to the almost straight trend of the curves. The figure thus illustrates the second important finding of this analysis: that the solid phase becomes negligible at high temperature. As shown in the next section, the behavior of real systems does not coincide with these thermodynamic predictions; nevertheless, equilibrium simulations of gasifiers operating above 800 °C can be carried out by considering solely a single gaseous phase. In a similar way, Figure 3 shows the influence of the operating conditions on the methane yield at equilibrium. This time, the curves represent systems characterized by the presence of 5 mol of CH4 per kilogram of daf biomass (cellulose in this case), a condition chosen to help compare these data with experimental results in the next section: at each temperature, systems falling above the corresponding curve are characterized by a lower equilibrium yield of methane, whereas those below the curve exhibit higher methane yields. From the thermodynamic predictions reported in the figure, it is clear that methane is expected to be a significant component of the gas phase at low temperature, especially when the hydrogen concentration is high (the case for steam gasification), and to progressively decrease

(

CnHx + nH2O ) nCO + n +

x H 2 2

)

(7)

An additional portion of these organic compounds contributes to the condensable organic phase, collectively called “tar”, by means of secondary, complex polymerization processes that are not taken into account here. The presence of a carbon-containing solid phase is the result of biomass primary pyrolysis (see eq 1), and also derives from a reaction process in competition with the production of hydrogen and carbon monoxide from the primary pyrolysis products (illustrated in principle by eq 4); these reactions have been studied widely in relation to the catalytic steam reforming of hydrocarbons and the associated problem of loss of catalyst activity due to progressive accumulation of carbon on the active surfaces.16 As far as the chemical nature of this solid phase is concerned, graphitic carbon is typically obtained by the disproportion of carbon monoxide (linear combination of eqs 2-4, the Boudouard reaction),

2CO ) C(s) + CO2

(8)

whereas coke is preferentially obtained by decomposition and condensation of hydrocarbons and is the result of polymerization processes, such as those involved in tar formation.17 For the sake of simplicity, the presence of a solid phase consisting entirely of graphitic carbon has been assumed here. The standard Gibbs free energy of formation as a function of temperature has been obtained from literature data18 for most of the species considered here; otherwise, it has been calculated by means of the Gibbs-Helmholtz equation based on standard Gibbs free energy and enthalpy of formation at 298 K and the standard heat capacity.19 The most significant results of equilibrium simulations are now presented with the help of a few figures. First, Figure 1 shows the equilibrium constants, KP, for stoichiometric relations 1-6 as a function of temperature; for eq 1, ln KP is always greater than 100, so that, under all investigated conditions, the

Ind. Eng. Chem. Res., Vol. 45, No. 2, 2006 837 Table 1. Elemental Analysis of Biomass Utilized in the Experimental Tests almond shells moisture (wt %) ash (wt % dry) elemental analysis (wt % daf) carbon hydrogen oxygen nitrogen LHV (kJ/kg daf)

sawdust

7.9 1.26

6.33 1.86

51.7 6.1 41.4 0.76 18306

48.2 6.4 45.9 0.23 18467

picture, distinguishing between char and tar compounds, appears infeasible at this level of description. 2.2. G-Minimization Method. The second method for solving the simultaneous chemical equilibrium problem, the G-minimization technique,21,22 does not require knowledge of an independent set of chemical reactions. For a system at constant temperature and pressure, the criterion of equilibrium is given by c

Figure 3. Gasification of cellulose: influence of temperature, SBR, and ER on the methane yield at equilibrium.

in abundance as the temperature is increased, in favor of further hydrogen formation. This picture is in general qualitative agreement with the experimental findings, although the methane yield in the product gas of high-temperature gasification processes is considerably underestimated. The mole fraction of naphthalene is found to be insignificant over the whole temperature range: at high temperatures, the equilibrium constant associated with eq 5 becomes very high, whereas at low temperatures, the mole fractions of steam, carbon monoxide, and hydrogen are such to result in a negligible presence of naphthalene, despite a rapidly decreasing thermodynamic constant. This result is quite general: it is linked neither to the specific ratios among carbon, hydrogen, and oxygen in the system (it is obtained also with a biomass different from cellulose) nor to the choice of naphthalene to represent organic vapors in the system. The values of the Gibbs free energies associated with the different species show that the evolution of the system toward equilibrium can be well described by considering the primary products of pyrolysis to be very unstable, tending to result in a solid phase of essentially carbon and a gaseous phase of small molecules (permanent gases) that becomes predominant at high temperature. This is the reason rapid quenching and separation of the vapor phase is strongly recommended for “flash pyrolysis” processes, aimed at the production of bio-oils.20 A simple means for confirming the above hypothesis is to substitute a high-molecular-weight hydrocarbon, for instance, anthracene (C14H10) for graphitic carbon, C(s), in the above eqs 1-6. Equations 1 and 4 thus become

x x x (C6H10O5)x ) xCO + CO2 + 10 H2O + C14H10 (1a) 3 3 3 14CH4 ) C14H10 + 23H2

(4a)

The new system of independent equations (eqs 1a, 2, 3, 4a, 5, and 6) furnishes for anthracene, under equilibrium conditions, results qualitatively similar to those reported previously for C(s) (see Figure 2), although this time, the presence of the less-stable anthracene is limited to a much more restricted range of operating conditions. It is clear from this demonstration that, for real systems, C(s) can be assumed to represent all of the carbon that is not converted into gaseous fuels; a more complete

dG )

µi dni ) 0 ∑ i)1

(9)

where the variations in ni are not independent: they must conform to the element balances, which are treated as constraint equations. Let Ak be the “total” g-atom of the kth element present in the system, as determined by the initial composition of the system. Let aik be the number of atoms of the kth element present in a molecule of chemical species i. The element balance can then be written as c

ni aik - Ak ) 0 ∑ i)1

(k ) 1, 2, ..., m)

(10)

where the constituents are made up of m elements. On the basis of the conclusions of the previous section, this method is now applied to the single-phase, ideal gaseous system that results from the high-temperature gasification of biomass fuels of known elemental composition. Using the Lagrangian method of undetermined multipliers, the constraint equations for the variations in ni can be solved simultaneously with eq 9. Thus, the resulting system of equations to be solved becomes m

∆G f0i + RT ln(yi P) + c

∑ πk aik ) 0

(i ) 1, 2, ..., c) (11)

k)1

A

k yi aik ) -∑ n i)1

(k ) 1, 2, ..., m)

(12)

Of the (c + m) unknowns, (c - 1) represent the yi; m is the number of Lagrangian multipliers; and the last unknown is n, the total number of moles. The numerical procedure proposed by Smith and Missen23 to solve the above equations for single-phase systems has been coded in MATHCAD11.24 It was applied to the gaseous phase obtained as a result of steam gasification of almond shells (the composition of which is given in Table 1) at T ) 800 °C, with SBR in the range of 0.3-1; under these operating conditions, the series reactor method does not predict any solid carbon phase formation. In addition to CO, CO2, H2, and CH4, a variety of light (C1-C3) and heavy (C6-C14) hydrocarbons have been considered as possible species in the system; this is quite easily accomplished with this method as a result of it being unnecessary to specify the reaction network. In addition, nitrogen-

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Figure 4. Steam gasification of almond shells at 800 °C: yields of different gaseous species as a function of SBR.

containing species are mainly ammonia (NH3) and hydrogen cyanide (HCN),25 and these have been considered here, together with N2. Figure 4 shows the most significant species in the system under equilibrium conditions as a function of SBR. The results of the calculations are expressed as moles per kilogram of daf biomass. Among all of the hydrocarbons, only methane has a tangible presence in the producer gas, although never exceeding 2 mol/kg of daf biomass. As the steam concentration is increased, the trends of hydrogen, carbon monoxide, carbon dioxide, and their respective ratios are in qualitative agreement with the experimental evidence: on a volume basis, hydrogen is the most abundant gas component, and syngas (H2 + CO) represents about 80% of the overall dry gas yield and maintains a fairly constant value over the whole range explored. The calculated water conversion is on the whole higher than in real systems, always above 50% and exceeding 90% at low values of SBR. This result is not unexpected, because the thermodynamic predictions consider complete conversion of carbon into gaseous compounds, thus contributing to enhanced steam consumption. Similar calculations were performed for the air gasification of sawdust (the composition of which is also given in Table 1), and the results are presented in Figure 5. In this case, the results reveal an increase of the overall dry gas yield with the equivalence ratio, ER, because of the increasing amount of nitrogen, which more than compensates for the corresponding reduction in syn gas production. As in the previous simulation, the methane yield is underestimated in comparison with the experimental evidence, and ammonia is predicted to be the most abundant nitrogen compound, when molecular N2 is excluded. 3. Comparison with Experimental Results In this section, previously published results,4,26,27 obtained in our laboratory, together with those of additional experimental tests are compared with the results of the thermodynamic limit computations described above. The additional experimental data cover more comprehensively the situation of catalytic air and steam biomass gasification, in which the gasifier bed inventory consists of mineral substances that are catalytically active for the cracking and reforming reactions. The same experimental setup and procedure as that used to obtain the previously published data was employed: a continu-

Figure 5. Air gasification of sawdust at 800 °C: yields of different gaseous species as a function of ER.

Figure 6. Scheme of the experimental apparatus: A, fluidized-bed gasifier; B, biomass feeder; C, steam generator; D, cyclone and ceramic filter; E, condensers and tar sampling; F, gas sampling and analyzers.

ous, bench-scale fluidized-bed biomass gasifier, shown schematically in Figure 6, capable of processing a biomass throughput of about 0.5 kg/h (almond shells for the steam gasification tests, sawdust and wood scraps for the air gasification tests). Each test ran for about 2 h following the establishment of steadystate conditions. The gasifier was placed in a cylindrical furnace that enabled the operating temperature to be controlled and maintained uniform in the wind box, fluidized bed, and freeboard. To maintain uniform fluidization conditions with varying gas inputs, a controlled flow of nitrogen, appropriately determined for each test, was employed to supplement the fluidizing gas: the producer gas yield and composition data, reported in the tables and figures below, do not include this quantity. All tests were carried out in the absence of any secondary gas cracking or reforming step downstream of the gasification reactor; particulate matter in the producer gas was removed in a cyclone and a ceramic candle filter; the product was cooled to separate water and tar and analyzed on-line to determine H2, CO, CO2, and CH4 concentrations. A flame ionization detector

Ind. Eng. Chem. Res., Vol. 45, No. 2, 2006 839 Table 2. Steam Gasification of Almond Shells with Ni-Olivine Bed Inventory (dP ≈ 500 µM) gasifier temperature, °C steam/biomass dry water conversion, % gas yield, Nm3 dry/kg daf tar content, g/Nm3 dry char residue,a g/kg daf carbon conversion, % H2 (vol % dry gas) CO (vol % dry gas) CO2 (vol % dry gas) CH4 (vol % dry gas) [H2][CO2]/[CO][H2O] KP water-gas shift a

820 1 45.0 1.83 0.5 37.3 92.5 51.8 24.4 19.2 4.7 1.07 0.98

820 1 45.5 1.88 0.4 31.6 93.7 52.4 24.1 19.2 4.3 1.14 0.98

820 0.5 59.0 1.77 0.2 45.2 91.2 50.8 34.4 10.6 4.2 1.08 0.98

750 0.5 43.0 1.32 1.8 127.1 75.5 45.5 27.9 19.8 6.9 1.19 1.31

700 0.5 26.0 0.99 17.7 170.1 63.5 38.8 26.5 24.0 10.8 0.75 1.65

Calculated by imposing the closure of the carbon mass balance.

Table 3. Air Gasification of Sawdust with Olivine Bed Inventory (dP ≈ 500 µM) gasifier temperature, °C equivalence ratio, % gas yield, Nm3dry/kg daf tar content, g/Nm3 dry char residue, g/kg daf carbon conversion, % N2 (vol % dry gas) H2 (vol % dry gas) CO (vol % dry gas) CO2 (vol % dry gas) CH4 (vol % dry gas) [H2][CO2]/[CO][H2O] KP water-gas shift

800 0 1.14 19.0 70.4 81.2 0 35.9 44.8 8.1 11.2 0.73 1.06

811 9.9 1.27 28.6 69.0 77.7 27.3 20.1 29.9 14.2 8.5 0.54 1.01

805 15.1 1.52 16.8 62.8 83.0 34.8 16.3 25.3 16.7 6.8 0.63 1.04

816 18.7 1.59 24.1 59.2 79.1 41.5 13.9 24.8 13.0 6.8 0.41 0.99

829 20.0 1.63 24.7 58.9 79.8 43.2 13.2 19.9 18.3 5.5 0.62 0.94

816 22.3 1.74 14.8 49.0 84.9 45.2 10.7 20.7 18.8 4.7 0.50 0.99

(FID) module was also available to detect total H-C, expressed as the propane equivalent concentration; however, under the experimental conditions employed (high temperature, presence of reforming catalysts), this quantity always remained negligibly small once the measured methane concentration was taken into account. The dry gas yield was measured by means of a volumetric gas meter. The amount of organic compounds condensed with water was determined by means of total organic carbon (TOC) analysis, and the mass of tar was estimated assuming naphthalene to be its representative component; when the tar content was higher, a solid organic phase was also formed, the weight of which was added to the previous quantity. The overall recovery of water and its conversion were also determined. The amount of char elutriated with the gas flow and accumulated in the bed was also measured (by burning it off with air) at the end of each air gasification test, whereas in the steam gasification tests, it was estimated by imposing closure of the carbon mass balance; data calculated in this way were found to be consistent with the measured quantities and previously detected values.27 The elemental analysis of biomass feedstock utilized in these tests is reported in Table 1, and the experimental results are summarized in Tables 2 and 3. Altogether, as shown in Figure 7, previous experimental investigations and recently performed steam gasification tests show that, at high temperature (above 750 °C) and in the presence of nickel catalysts, the molar concentration ratio [H2][CO2]/[CO][H2O] in the gas phase approaches the equilibrium value for the water-gas shift reaction; this happens either when the catalyst is placed directly in the gasifier or when a secondary catalytic reactor is added downstream of the gasifier itself. Similar behavior is found with steam gasification tests in the presence of natural catalysts (olivine, dolomite) with gasification temperatures above 800 °C. On the other hand, when the gasifier bed inventory is sand and there is no downstream catalytic conversion step, the gas

Figure 7. Water-gas shift reaction: experimental vs equilibrium values of the molar concentration ratio at different operating conditions.

Figure 8. Methane yield per kilogram of daf biomass at different operating conditions.

composition is far from equilibrium: under such conditions, the hydrogen yield and steam conversion both decrease substantially, and the experimental values of the concentration ratio relevant for the water-gas shift reaction become smaller than the corresponding theoretical ones. Additional data points of this type, falling far off the 45° straight line, are available, although not included in the figure. Slight deviations from equilibrium of the water-gas shift reaction characterize most of the air gasification tests; the bed inventory is of natural catalytic mineral matter (olivine), yet the very limited presence of steam in the gas phase (the biomass feedstock is rather dry, and no additional input of water is provided) reduces the extent of the reaction of carbon monoxide toward production of hydrogen. Figure 8 reports the methane yield measured at different operating conditions. The maximum methane yield is between 5.5 and 6 mol per kilogram of biomass daf, and it was obtained with gasifier bed inventories of inert substances (sand) or natural catalysts (dolomite, olivine) and in the absence of air. Ni catalysts are able to reduce substantially the methane yield, by

840


steam reforming a variable portion of the number of moles of CH4 originally formed in the biomass devolatilization process.4 With a secondary fixed bed made of fresh commercial nickel catalyst downstream of the gasifier and a large excess of steam, the methane content in the producer gas became negligible,26 confirming the validity of the thermodynamic predictions; with a perovskite catalyst specifically developed for “tar cleaning” in biomass gasification processes and utilized in the same way, a maximum methane conversion of about 60% (calculated with reference to the above-mentioned top yield) has been obtained at 900 °C.4 Natural mineral substances (dolomite, olivine) are effective as catalytic agents for the destruction of tar, but not for methane reforming: as shown in a previous work,27 they lead to increasing overall gas yields, in comparison with tests where sand is the bed material, but the number of moles of methane decreases very slightly. Published experimental findings generally confirm this point, by reporting a tiny reduction of methane volumetric fraction (against an increase in the overall gas yield) for air gasification that includes a downstream catalytic treatment with dolomite28 and negligible activity for methane reforming of similar treatments (with dolomite, magnesite, calcite) applied to steam gasification.29 It is worth noting here that the maximum methane yield, obtained from steady-state gasification tests, is in very good agreement with the total amount of methane that was detected for devolatilization of beech wood spheres of different size, dropped in a high-temperature (750 °C) fluidized bed.6 In that case, the instantaneous gas composition was measured as a function of time, and cumulative values were obtained by integration; the experimental conditions were also carefully designed to minimize the influence of reactive processes in the gaseous phase surrounding the biomass particles. This procedure identified an average methane yield of around 5 mol per kilogram of beech wood (daf). The consistency of the results obtained from continuous gasification and dynamic devolatilization tests shows that methane is indeed formed in the devolatilization phase of the process and provides experimental justification for assumptions of this kind, often put forward in the literature to explain the presence of methane in coal and biomass gasification processes.10,12,30 Similar conclusions can be drawn from the air gasification tests: when the air flow rate is brought to zero (a plain pyrolysis process in a nitrogen atmosphere), the methane yield is close to the value of 5.5 mol/kg of daf biomass, as found in the devolatilization tests, and it decreases slightly in the presence of oxygen, because of partial burning and/or H2O reforming. With reference to the conversion of carbon to gaseous products, our experimental investigations show that the simultaneous action of temperature and catalytic reforming processes is found to decrease the presence of condensable heavy organic compounds in the gas produced by steam gasification to very low levels (less than 1 g/Nm3 of dry gas; see Table 2), which, as a result, play a negligible role in the overall carbon mass balance. This is not the case for air gasification (see Table 3), where, despite temperature and the presence of an olivine bed inventory, the tar content in the producer gas is more than 2 orders of magnitude greater, with a contribution of 4-8% to the overall carbon balance. These values confirm the well-known important role of steam in tar reforming processes. On the other hand, in all experimental tests, the presence of char was observed in the solid particulate elutriated from the fluidized-bed gasifier, as well as in the bed itself, when discharged at the end of a biomass gasification campaign. Altogether, with gasification temperatures above 800 °C and

steam/biomass ratios in the range of 0.5-1, an average carbon conversion into permanent gases of 92% was obtained, and it remained close to 80% in an inert atmosphere (see Tables 2 and 3). With reference to air gasification, increasing the equivalence ratio, ER, slightly improves the carbon conversion, as has also been noticed in pilot-plant installations.31,32 From the above-reported data and analyses, it appears that a straightforward application of thermodynamic models to isothermal, high-temperature biomass gasifiers leads to substantial discrepancies with actual behavior, in terms of carbon conversion, methane yield, and tar formation; these gaps are reduced quite considerably for cases of catalytic steam gasification processes, although even there, some corrections become necessary. Obvious kinetic arguments can be invoked to justify the residual presence of those compounds that are predicted to be absent (or present in negligible amounts) at gasifier operating conditions, yet such arguments cannot explain why these species have appeared in the first place. On the other hand, the thermodynamic analysis, extended over a wide range of temperature levels, is able to highlight effects linked to the actual, nonisothermal gasification process experienced by fuel particles while they are heated from ambient conditions to about 900 °C. The final products retain a memory of this thermal history, because devolatilization is a relatively fast process, taking place mainly during the particle heating phase (i.e. at temperature levels that are lower than those of the reactor), whereas subsequent heterogeneous reactions involving charcoal are relatively much slower and unlikely to reach completion within their residence time in the gasifier, and methane reforming requires the presence of appropriate catalysts. Such phenomena, previously studied dynamically,6 are cast here into the appropriate thermodynamic framework. In this respect, to increase the yield of the permanent gases and carbon conversion, a critical factor clearly relates to the availability of reactor configurations able to minimize the heating time of the fuel particles, such as fluidized-bed reactors, whereas temperature levels of 800-900 °C appear adequate to maximize the gas yield: in fact, any increase of temperature would be detrimental when a hydrogen-rich fuel gas is desired, because of the unfavorable trend of the water-gas shift equilibrium. 4. Semiempirical Calculation Tool The practical interest in the successful application of equilibrium-type models for reacting systems lies in the availability of software routines for estimating the performance of chemical reactors, offered as standard options in commercial process flow sheet calculation tools. We propose here a simple method for characterizing the output of a biomass gasifier operating at high temperature, based on the CHEMCAD software package. The system is considered to consist of four elemental species (C, H, O, N). It is assumed that all chlorine and sulfur, present in the biomass as trace compounds, give rise to HCl and H2S, respectively, and therefore, these are not included in the equilibrium calculations; in any case, the molar abundance of Cl and S is at least 2 orders of magnitude lower than that of the major three elements (C, H, O). The global reactive process occurring in the gasifier can be characterized as

CaCHaHOaONaN + x1H2O(l) + x2H2O(g) + x3(O2 + γN2) f z1H2 + z2CO + z3CO2 + z4CH4 + z5N2 + z6NH3 + z7H2O(g) + z8C10H8 + z9C(s) (13)


where the g-atoms of carbon, hydrogen, oxygen, and nitrogen in the biomass raw formula are obtained from the fuel elemental analysis: x1 is given by the fuel humidity; (x1 + x2) and x3 are given by the values fixed for the steam/biomass ratio, SBR, and the equivalence ratio, ER, respectively; and γ is chosen according to the nature of the gasification agent (air, enriched air, or pure oxygen). The list of chemical species on the RHS of eq 13 has been restricted to the most significant ones, as found in the previous analyses: in the computation trials, a much more extensive list of compounds (light and heavy hydrocarbons, additional nitrogen compounds) is considered, without difficulty, because their thermodynamic data are included in the software library. To account for the presence of substantial amounts of char and methane in the gasifier output, which are largely underestimated in a rigorous equilibrium picture of the system, it is proposed that the equilibrium calculations be carried out after a modification of the elemental balance constraint conditions for the carbon and hydrogen atomic species as follows:

AC ) AC0ΧC - nCH4dev(1 - ΧCH4)

(14)

AH ) AH0 - 4nCH4dev(1 - ΧCH4)

(15)

where AC0 and AH0 are the true values of the carbon and hydrogen g-atoms in the system, respectively, and AC and AH are the corresponding corrected quantities to utilize in the calculations. In addition, a fictitious inert gaseous compound is introduced, having the sole function of replacing the methane contribution in the calculation of the total moles (total pressure, molar concentrations, etc.) in the gas phase:

nIg ) nCH4dev(1 - ΧCH4)

(16)

The amounts of methane and solid carbon withdrawn before embarking on the equilibrium calculations are then restored to the system so as to compute its final composition. As noted in a previous section, solid carbon, C(s), denotes here the overall amount of carbon not converted into gaseous fuel, because a distinction between char and tar is not feasible. For a characterization of gasification systems in terms of yields and efficiencies, this is not a serious problem because both are considered undesired byproducts. As a result, three empirical parameters need to be fixed: (i) The first is the carbon conversion into gaseous products, XC. According to the experimental evidence reported above, XC is generally to be found in the range 0.8-0.9, and it increases with temperature and steam/biomass ratio, because of a substantial reduction in tar yield; beyond the operating conditions, the gasifier design can significantly affect this value. (ii) The number of moles of methane produced in the devolatilization step, per kilogram of biomass daf, nCH4dev, must also be fixed. This quantity is affected only marginally by the operating conditions, and the value of 5.5 mol/kg of biomass daf is suggested here as a result of the experimental evidence presented above. (iii) Finally, the methane conversion by steam reforming, XCH4, should be fixed. This quantity becomes nonzero in cases when a Ni catalyst (or similar catalyst) is utilized in the gasification process; in such cases, a good approximation could be XCH4 ) 1/3, with higher values (up to XCH4 ) 2/3) becoming appropriate when very active reforming catalysts are employed. A slight conversion of methane (XCH4 e 1/4) can be taken into account in the gasification with air.

Table 4. Comparison between Experimental Data and Results of Simulations for a Typical Gasification Test with the 500 KWth ECN Gasifier (T ) 851 °C) plant calculation straightforward results33 tool equilibrium gas yield, Nm3 dry/kg daf biomass H2, vol % CO, vol % CO2, vol % CH4, vol % C2, vol % C6+, vol % N2, vol % NH3, vol % H2O, vol % wet char + tar, g C/kg daf biomass

2.04 11.9 14.2 16.4 4.0 1.45 0.03 51.4 n.d. 12.4 69.2a

2.10 11.9 15.4 15.7 6.5 0.0 0.0 50.5 0.001 12.0 74.7

2.64 24.5 28.7 6.6 0.01 0.0 0.0 40.3 0.002 5.9 0

a Obtained by subtracting total carbon in the gas from carbon content in the biomass feedstock.

The above assumptions are considered more consistent with the experimental evidence than those suggested elsewhere, in approaches of a similar kind, tailored to describe specific gasification processes and a closer range of parameters variability. For instance, in a previously cited paper,12 the fraction of carbon converted into gaseous species is given as a strongly dependent, inverse function of ER, in marked contrast to the experimental evidence.31,32 Table 4 shows the application of the computing procedure outlined above to the prediction of the performance of the 500kWth fast fluidized-bed biomass gasifier operated by ECN (Netherlands Energy Research Foundation); in their paper,33 the authors give full details of biomass composition, flow rate and inlet temperature of air, operating conditions, and gas analysis, for what they refer to as a typical experiment, thus providing all input data needed for calculations and comparison between experimental data and the results of simulations. The high value of the gasification temperature (851 °C) and the presence of moisture in the feed lead us to choose XC ) 0.85 (the middle value in the range of variation of this quantity found in our experimental investigations), and we assume the methane conversion to be negligible, because of the absence of catalytic substances, either natural or chemically synthesized. Table 4 shows the very good predictive capability of the computing tool and, at the same time, a significant improvement compared to the results obtained when the CHEMCAD software routine for chemical reactors is utilized without any correction for the equilibrium availability of carbon and hydrogen g-atoms in the system (AC ) AC0 and AH ) AH0 in eqs 14 and 15, respectively). It is also worth noting the perfect agreement between experimental and calculated values of the steam content in the gasifier output, a very sensitive indication of the extent of gasification reactions in the system. Last but not least, good agreement is found for the energy balance: the enthalpy difference between inlet and output streams obtained by the thermodynamic calculations is close to the heat loss declared by the authors.33 To check the sensitivity of the simulation tool, the carbon and methane conversions were varied in the intervals 0.8 e XC e 0.9 and 0 e XCH4 e 0.33, respectively. It was found that the values predicted by the model for all combinations of XC and XCH4 span a range that always includes the corresponding experimental data reported in Table 4, with a maximum deviation of (45%. An additional, well-known gasifier design is that developed by the Vienna University of Technology (TUV): a dual fluidized-bed system with circulation of the bed inventory between a bubbling, steam-fluidized bed and an air combustion zone (riser). In this case, the carbon-containing solid phase

842


Table 5. Comparison between Experimental Data and Results of Simulations for Steam Gasification Tests with the 100-kWh TUV Gasifier (SBR ) 0.63) bed inventory (temperature) olivine (850 °C)

gas yield, Nm3 dry/kg dry H2, vol % CO, vol % CO2, vol % CH4, vol % C2, vol % N2, vol % NH3, vol % [H2Oin - H2Oout]/[biomass dry]

plant results34

calc tool

0.95 38.9 29.1 17.5 11.4 2.0 n.d. n.d. 0.044

1.13 47.7 19.4 21.9 10.7 0 0.08 0.0091 0.077

Ni-olivine (838 °C) plant results34 0.99 43.9 27.2 18.8 8.3 1.3 n.d. n.d. 0.072

calc tool 1.20 52.7 20.2 20.2 6.9 0.0 0.08 0.0002 0.082

resulting from biomass devolatilization is mechanically transported from the gasification zone and burned in the riser (together with some additional fuel), at a slightly higher temperature, to allow for the necessary input of thermal energy to the gasification reactions, by means of solid recirculation. A 100-kWh prototype of this system has been operated with a bed inventory of olivine and of Ni-olivine; extensive information on these tests is available in the literature.34 Because of the design and operation of this system, the fraction of solid carbon to be subtracted for the purpose of equilibrium calculations is assumed to be equal to that removed from the gasification zone: this is easily obtainable from the available experimental data, by consideration of the carbon g-atoms in the biomass feedstock and their total found in the gaseous fuel. The conversion of methane formed in the biomass devolatilization step is fixed in accord with the findings of the previous section: it is assumed to be negligible with the olivine bed inventory, whereas the choice of XCH4 ) 1/3 appears to be most appropriate for the experimental test with the nickel catalyst. The simulation results for the gasification zone alone obtained with these assumptions are reported in Table 5, together with the experimental gas yield and composition. Although the predictive capability of the simulation tool is less accurate in this case, the major trends found experimentally by changing the bed inventory from natural olivine to Ni-olivine are correctly reproduced. It should be considered that, in addition to solid circulation, some gas leakage between the gasification and combustion zones occurs in the operation of the TUV gasifier, which is neglected in our calculations, because it is not quantifiable by means of the available data. 5. Conclusions None of the available kinetic models for biomass gasification, regardless of complexity, is able to predict the yield, composition, and calorific value of the resulting gaseous fuel. For this reason, a different approach has been followed in the present work, by applying equilibrium models and comparing the thermodynamic limitations with the actual results available in the literature and additional, purposely performed experimental tests. It has been found that, at the high temperature levels of interest for most practical applications of biomass gasification processes, hydrogen, steam, carbon monoxide, and carbon dioxide in the gaseous phase match closely the thermodynamic equilibrium of the water-gas shift reaction and that the concentrations of high-molecular-weight hydrocarbons and oxygenated compounds are brought to very low values by catalytic substances in common use, again in good agreement with thermodynamic calculations. On the other hand, in clear disagreement with predictions, methane (and light hydrocar-

bons), formed as a result of biomass particle devolatilization, is found in the producer gas, generally with a molar fraction in the range of 5-8% (on a dry, nitrogen-free basis); in addition, a fraction of solid, unconverted carbon (on the order of 1020% of that originally present in the biomass) remains in the bedsa further clear indication that the heterogeneous reactions involving char are not completely accomplished within the residence time range of most gasifiers. On the basis of these findings, a semiempirical model is proposed to predict quantitatively the major products of a gasification process; it is very easy to apply in practice as it makes use of standard routines for calculations involving chemical reactors under equilibrium conditions that are available in commercial software packages for process flow sheet calculations. Two examples of application are provided for fluidized-bed gasifiers that are very suitable for simulation with equilibrium-derived models because of their high level of operating temperature and almost isothermal behavior. The calculation method is based on an estimate of the displacement of real systems from thermodynamic equilibrium conditions, so that, in principle, it is also applicable to gasifiers of different kinds, moving- or fixed-bed reactors, provided that a reduced number of empirical inputs is identified and quantified to accommodate for such deviations. With updraft gasifiers, where the product gas leaves the reactor in contact with the biomass feedstock and at relatively low temperature, equilibrium calculations are expected to furnish predictions rather far from plant data, whereas with down-draft gasifiers, characterized by sufficient residence times of the gaseous output stream at high temperature and in contact with mineral ashes providing beneficial catalytic effects, the equilibrium composition should be approached more closely. Acknowledgment This work was funded in part by the European Commission (Contract ENK5-CT2000-00314) and by the Italian Ministry for Research (MIUR). We are grateful to G. Antonelli for his assistance in the experimental tests. Nomenclature aik ) number of atoms of the kth element in a molecule of species i Ak ) total number of g-atoms of element k in the system c, i, k, m ) generic chemical species (atom or molecule) C(s) ) carbon-containing solid phase daf ) dry, ash free ∆G 0f ) free energy of formation at system temperature (cal/ mol) ER ) equivalence ratio (defined in the text) γ ) nitrogen-to-oxygen ratio in the gasification agent (air, enriched air, etc.) G ) system Gibbs free energy (cal) KP ) thermodynamic equilibrium constant LHV ) low heating value (kJ/Nm3 of dry product gas) µi ) molar free energy of species i (cal/mol) n, x, z ) number of g-atoms or moles P ) pressure (atm) π ) Lagrange multiplier r ) reactor number in the series reactor method R ) gas-law constant [cal/(mol K)] SBR ) steam-to-daf biomass ratio T ) temperature (°C or K) Χ ) chemical conversion y ) molar fraction


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ReceiVed for reView July 13, 2005 ReVised manuscript receiVed October 3, 2005 Accepted October 18, 2005 IE050824V

Thermodynamic Limits and Actual Product Yields and Compositions in

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