Generalized Analysis of Gasifier Performance using Equilibrium

ARTICLE pubs.acs.org/IECR

Generalized Analysis of Gasifier Performance using Equilibrium Modeling Anapagaddi Ravikiran,† Thiruvengadam Renganathan,*,† Subramaniam Pushpavanam,† Ravi Kumar Voolapalli,‡ and Young Sang Cho§ †

Department of Chemical Engineering, Indian Institute of Technology Madras, Chennai 600036, India Corporate R & D Centre, Bharat Petroleum Corporation Limited, Greater Noida 201306, India § Energy Division, Korea Institute of Science and Technology, Seoul 130650, Korea ‡

ABSTRACT: Thermodynamic modeling of gasification process provides a quick estimate of performance of the gasifier. Most of the earlier work on thermodynamic modeling is restricted to a particular feedstockgasification agent combination and hence the results cannot be generalized. In the present work, the equilibrium modeling based on Gibb’s free energy minimization approach is used to analyze the performance of gasification of any fuel using oxygen or steam. The performance is analyzed at the carbon boundary point at which the cold gas efficiency is maximum. The gasification temperature, amount of gasification agent required, composition of syngas, and cold gas efficiency are predicted using Aspen Plus. The results are presented as contour plots on Van Krevelen coordinates (H/C vs O/C) and interpreted based on simplified gasification reactions. The performance for different feedstocks represented in Van Krevelen diagram is also analyzed. Finally, advantage of cogasification of feedstocks is highlighted.

1. INTRODUCTION Gasification is a process in which a carbonaceous feedstock is converted to fuel gas using a gasifying agent. The carbonaceous feedstock can be a nonrenewable source like coal or a renewable source like biomass. Other feedstocks used include natural gas, petroleum oil, and sewage sludge.1 The gasifying agents typically used are oxygen or air and/or steam. The combination of feedstock and gasifying agent determines the composition of synthesis gas produced and hence its end use. Syngas to be used for power generation requires high concentrations of carbon monoxide and hydrogen and hence oxygen is a suitable gasifying agent. On the other hand, when syngas is to be used for synthesis of chemicals, larger concentrations of carbon dioxide and methane are desired so steam is used as a gasifying agent. Modeling of gasification processes can be done using an equilibrium approach or using a rate-based approach.2 The equilibrium approach is based on the concept of chemical reaction and phase equilibria. It assumes that the rate of reactions is very fast relative to the residence time of reactants in the gasifier so that equilibrium conditions can be attained. The predictions using this approach can be realized only when this assumption is met. On the other hand, the rate-based approach is more comprehensive taking into account the hydrodynamics, transport process, and reaction kinetics. Consequently, the results of a rate-based simulation are closer to experimental results. However, the development of a rate-based model is challenging depending on the level of detail required and this is determined by the type of gasifier used. Though the rate-based model is more realistic compared to equilibrium models, equilibrium models give a quick idea of the limits of operation and hence are useful for preliminary design of gasifiers. Besides, the equilibrium models predict the performance of down draft gasifier satisfactorily, where the assumptions of equilibrium model are valid. r 2011 American Chemical Society

The equilibrium model formation can be based on two approaches.2 The stoichiometric approach is based on a set of selected independent reactions and uses elemental balances and the equilibrium relations of the reactions. The nonstoichiometric approach (also called Gibbs free energy minimization approach) is based on a set of selected species assumed to be present in the syngas. This method minimizes the total Gibbs free energy of the system subjected to the constraints of elemental balances and energy balance. It can be proved that both methods are equivalent.3 In the present work, the Gibbs free energy minimization approach is used to thermodynamically model the gasification process. Literature on modeling of gasifiers using the equilibrium approach is abundant. The existing literature can be classified into three main categories. In the first category, the articles focus on gasification of a specific feedstock (e.g., refs 4, 5) or few feedstocks (refs 6, 7) using a specific gasifying agent. These articles focus on the effect of variables such as amount of gasifying agent, moisture content of the feedstock, gasifying temperature (if isothermal), operating pressure, preheating of gasifying agent on the syngas composition, calorific value of syngas, gasification efficiency, and gasification temperature (if adiabatic), etc. Because of specificity of feedstock and gasifying agent, these results cannot be generalized. In the second category of articles, a modification of equilibrium model has been attempted. As explained above because of the inherent assumptions in equilibrium modeling, the predictions Special Issue: Nigam Issue Received: March 29, 2011 Accepted: June 10, 2011 Revised: June 10, 2011 Published: July 19, 2011 1601

dx.doi.org/10.1021/ie2006406 | Ind. Eng. Chem. Res. 2012, 51, 1601–1611

Industrial & Engineering Chemistry Research

Figure 1. Representation of different feedstocks on Van Krevelen (H/C vs O/C) coordinates.

of equilibrium models are not close to experimentally observed values for different types of gasifiers such as fluidized bed and entrained bed gasifiers. Hence to “force” the equilibrium model to match the experimental results, different modifications of the equilibrium model have been attempted. These include empirically changing the gasification temperature (e.g., ref 2) or equilibrium constants of the reactions (e.g., 8) or carbon conversion to fit the model predictions to the experimental data. Modifications using a two-stage approach have also been attempted (e.g., 9). These require empirical fitting of fractional steam utilization. All these modifications are empirical and hence are limited in scope to the specific operating conditions viz. feed stocks, gasifying agent, gasifier type, and hence cannot be generalized. The feedstocks used for gasification can be coal or biomass or even waste materials. These feedstocks can be characterized by O/C and H/C ratios (on an atom basis) and are represented on the Van Krevelen diagram shown in Figure 1.10 Unlike the first and second category of papers, the third category includes papers that apply equilibrium models to present generalized results applicable to a wide range of feedstocks considering only a few independent variables. The work in this direction is just emerging. Prins et al.10 present results of the equilibrium model in terms of O/C and H/C ratio of feedstock and cover a wide range of fuels from hard coal (O/C = 0.01, H/C = 0.2) to biomass (O/C = 0.7, H/C = 1.5). The authors consider oxygen as the gasifying agent and determine the oxygen required for complete carbon conversion at which maximum efficiency occurs, both under adiabatic and isothermal conditions. The authors define the efficiency based on second law rather than first law of thermodynamics. Vaezi et al.,11 using a stoichiometric equilibrium model, determined the syngas characteristics for gasification of biomass using air at a fixed equivalence ratio. The gasification temperature, higher heating value, and cold gas efficiency are plotted on C/H vs oxygen content coordinates, covering a wide range of biomass. However, the study is limited to oxygen gasification of biomass feed stocks at a fixed equivalence ratio of 0.4. Recently a significant contribution toward unification of results of the equilibrium model of gasification has been made by Stemmler and Muller.12 The authors used a different set of

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coordinate system (independent variables) taking into account the total oxygen and hydrogen content in the feedstock and gasifying agent. The results are thus applicable to any feedstock and gasifying agent combination. The authors analyze the effect of operating temperature, pressure on composition and calorific value of syngas. However their analysis is limited to isothermal operation. The literature survey reveals that most of the works done so far are limited in scope due to specific feedstock gasifying agent combination and/or empirical modification and only a few attempts have been made to generalize the results of equilibrium model. The objectives of the present work are (i) to make a generalized prediction of optimal performance of gasifiers for any given fuel (characterized by H/C and O/C ratios). (ii) to analyze in detail the influence of H/C and O/C ratios on optimal gasifier performance under adiabatic conditions. The nonstoichiometric equilibrium model is used to predict the performance of the gasifier for two different gasifying agents viz. oxygen and steam. The thermodynamic model based on Gibbs free energy minimization approach and details of simulation are presented in the next section. Following this, the results obtained viz. gasification temperature, gasifying agent required, composition of syngas, and efficiency at carbon boundary point are presented and discussed as contour plots. The significant conclusions drawn from this work are summarized at the end.

2. THERMODYNAMIC MODELING AND SIMULATION DETAILS In the present work, the fuel is assumed to be dry and consist of only carbon, hydrogen, and oxygen. Accordingly the fuel is represented as Cnc HnH Ono . The reasons for taking only these three elements are 2-fold. First, these three elements make up to about 8590% of most fuels. The results and conclusions obtained when only these three elements are considered do not deviate much when other elements are also included. Second, since the objective of the work is to present results valid for different fuels, considering only these three elements enables us to uniquely represent the fuel and present the results in terms of H/C and O/C ratios. The gasifying agent used is either oxygen or steam. The components assumed to be present at equilibrium are C(s), CO, H2, CO2, H2O, and CH4. Based on the above assumptions, the global gasification reaction can be written as CnC HnH OnO + mO2 + sH2 O f nCðsÞ CðsÞ + nH2 H2 + nCO CO + nCO2 CO2 + nCH4 CH4 + nH2 O H2 O Here, s = 0 and m = 0 for gasification using only oxygen and only steam, respectively. The nonstoichiometric model is based on the concept that at equilibrium condition, the total Gibbs free energy of the system is minimum.13 The total Gibbs free energy is a function of temperature, pressure, and the number of moles of ith component (ni) present at equilibrium. Hence, the total Gibbs free energy of the system is given by Gt ¼ f ðT, P, nCðsÞ , nH2 , nCO , nCO2 , nCH4 , nH2 O Þ

ð1Þ

At equilibrium, since the total Gibbs free energy is minimum, dGt ¼ 0 1602

ð2Þ

dx.doi.org/10.1021/ie2006406 |Ind. Eng. Chem. Res. 2012, 51, 1601–1611


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Figure 2. Aspen Plus flowchart used for simulation of equilibrium model of adiabatic gasifier.

For a given pressure, the moles of the different components and temperature are determined such that the total Gibbs free energy is minimum subject to the elemental balances, Element C balance : nC ¼ nCO + nCO2 + nCH4 + nCðsÞ

ð3Þ

nH + 2s ¼ 2nH2 + 4nCH4 + 2nH2 O ð4Þ

Element H balance :

Element O balance : nO + s + 2m ¼ nCO + 2nCO2 + nH2 O ð5Þ and the energy balance, Hprod ðTÞ ¼ Hreact + Q

ð6Þ

where _

Hreact ¼ Hfuel + mH0fO

Z

_

2 , 298

+ s H0fH OðvÞ , 298 + 2

!

Tin 298

CpH2 OðvÞ dT ð7Þ

Hprod ðTÞ ¼

∑ prod

_

ni H0fi , 298

+

∑ ni prod

Z

T

Cpi dT

ð8Þ

298

Enthalpy of fuel is calculated using 1 _ nH H0fH O , 298 ð9Þ 2 2 The simulation is carried out using Aspen Plus and the flowsheet used is shown in Figure 2. In Aspen Plus, coal is represented as a nonconventional component. This nonconventional component has to be “split” into its constituent elements before giving it as input to a reactor. Hence two reactors, viz, the RYIELD and RGIBBS reactors are used to simulate the gasification process. The RYIELD reactor is used to “split” the fuel into its constituent elements. RGIBBS reactor is used to perform multiphase-reaction equilibrium calculation based on the Gibbs free energy minimization approach. Adiabatic condition is ensured by giving the heat output of RYIELD reactor as heat input to RGIBBS reactor. The sensitivity analysis feature of Aspen Plus is used to study the effect of various combinations of H/C and O/C ratios. For specified H/C and O/C atomic ratios, the ultimate analysis in terms of mass percentage of the elements carbon, hydrogen, and oxygen is determined. In each case the total fuel flow rate is kept constant. For each case, a design specification of carbon entering equal to carbon leaving in gaseous species is used to determine the oxygen flow rate required for complete conversion of carbon.

Figure 3. (a) Effect of oxygen flow rate on gasifier performance for fuel with H/C = 0.9 and O/C = 0.45. (b) Effect of steam flow rate on gasifier performance for fuel with H/C = 0.9 and O/C = 0.45.

The fuel is specified as a non conventional component and higher heating value (HHV) of the fuel is calculated using the correlation proposed by Channiwala and Parikh14 HHV ¼ 341:9xC + 1178:3xH 103:4xO

_

Hfuel ¼ Mfuel HHV nC H0fCO

ð10Þ

2 , 298

The range of validity of the above correlation is 0 < H/C < 1.81 and 0 < O/C < 0.83. The correlation has been developed based on specific H/C and O/C combinations (for the different feedstocks in Van Krevelen diagram). In this work, this correlation is assumed to be valid for all combinations of H/C and O/C. In this study, the cold gas efficiency (CGE) is used to represent the efficiency of the gasification process. It is defined as1 Msyngas LHV syngas Mfuel LHV fuel

ð11Þ

1 Mfuel HHV nH λ298 2 ¼ Mfuel

ð12Þ

CGE ¼ where

LHV fuel _

LHVsyngas ¼ 1603

_

nCO H0CCO , 298 + nH2 H0CH

_

2 , 298

Msyngas

+ nCH4 H0CCH

4 , 298

ð13Þ



3. RESULTS AND DISCUSSION In this work the gasification process is simulated using Gibbs free energy minimization approach using Aspen Plus simulation tool. The operating pressure of the gasifier is taken to be 1 atm. The gasifier is assumed to operate adiabatically and hence Q is taken as zero in the energy balance. Fuel is assumed to enter at 298 K. Oxygen at 1 atm and 298 K or steam at 1 atm and 500 K is used as gasifying agent. Figure 3a and b depict the dependence of adiabatic gasification temperature (henceforth called gasification temperature), fraction of unconverted carbon, composition of synthesis gas, and cold gas efficiency as a function of oxygen flow rate (mole of oxygen per mole of carbon) and steam flow rate (mole of steam per mole of carbon), respectively, for gasification of a typical fuel with O/C = 0.45 and H/C = 0.9. From Figure 3a, it can be seen that, with increase in oxygen flow rate, the fraction of carbon unconverted decreases and for a particular oxygen flow rate (0.27 mole of oxygen per mole carbon) all carbon in the fuel gets converted and no carbon appears in the exit syngas. At this point mole fraction of CO is maximum and that of CO2 is minimum and the CGE is maximum. From Figure 3b, it can be seen that, as the steam flow rate is increased, fraction of carbon unconverted decreases and for a particular steam flow rate of 1.8 mole of steam per mole of carbon, all carbon in the fuel gets converted and no carbon appears in the exit syngas. At this point mole fraction of CH4 is maximum and the CGE reaches a maximum and remains constant thereafter. Because thermodynamically, for complete carbon conversion maximum CGE is achieved, performance of the gasifier is analyzed at this point which is called the carbon boundary point.10 Gasifier performance is analyzed for O/C varying from 0 to 0.9 and H/C varying from 0 to 1.8, the limits obtained from the Van Krevelen diagram (Figure 1). Results are presented on gasification temperature, mole fraction of CO, H2, CO2, H2O, and CH4, molar ratio of H2 to CO, and cold gas efficiency (CGE). Whereas mole fractions of CO, H2, CO2, and CH4 are on dry basis, mole fraction of H2O is on wet basis. The results obtained from Aspen Plus for different H/C and O/C ratios are depicted as contour plots using Matlab. Along the contours the different performance variables are constant. The results are presented as contours on Van Krevelen coordinates namely H/C vs O/C and the effect of H/C and O/C on the simulation results is analyzed in detail. The results are interpreted based on the following simplified reversible chemical reactions which occur in gasification. Water gas shift reaction : CO + H2 O T CO2 + H2 ΔHr ¼ 41 166 kJ=kmol Methane reforming : CH4 + H2 O T CO + 3H2 ΔHr ¼ 205 13 kJ=kmol Methane formation : CS + 2H2 T CH4 ΔHr ¼ 74 520 kJ=kmol Boudouard reaction : CS + CO2 T 2CO ΔHr ¼ 172 459 kJ=kmol Primary water gas reaction : CS + H2 O T CO + H2 ΔHr ¼ 131 293 kJ=kmol The different feed stocks as represented by regions in Van Krevelen diagram are also superimposed on the contour plots to

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Figure 4. Contour plot of gasification temperature for gasification of fuel using oxygen.

assist in directly estimating the performance for different feedstocks. Co-gasification, which refers to using a mixture of two or more feedstocks (e.g., coal and biomass) as feed to the gasifier, results in O/C, H/C ratios of mixtures lying outside regions of single feed-stocks, as shown by line AB in Figure 1. The performance of cogasification of binary feed stock mixtures is also discussed. 3.1. Gasification Using Oxygen. 3.1.1. Effect of O/C and H/C Ratios. The contour plot of gasification temperature at the carbon boundary point on H/C vs O/C coordinates for gasification of fuel using oxygen is shown in Figure 4. For a given H/C, the gasification temperature decreases with O/C. With increase in O/C, since more oxygen is available in fuel, contribution to gasification by oxygen from external gasification agent reduces. The reactions of fuel with oxygen are exothermic. The net heat released depends on whether the oxygen consumed is from fuel or from the external gasifying agent. With increase in oxygen content in fuel the higher heating value of the fuel decreases (eq 10) and hence the specific enthalpy decreases. As a result the net heat released through gasification by oxygen in fuel is less than that released by oxygen in gasifying agent. So, with increase in O/C, the gasification temperature decreases. For a given O/C, the gasification temperature decreases with H/C. With increase in H/C, due to the increase in hydrogen content of the fuel, formation of more hydrogen is favored. Gasification reactions with hydrogen formation are endothermic. Further, with increase in hydrogen content in fuel, though the higher heating value increases (eq 10), the specific enthalpy decreases (eq 9). So, the gasification temperature decreases with H/C. The decrease in temperature with H/C is more for lower O/C ratios compared to higher O/C ratios. At higher O/C ratios, contribution of exothermic reactions increases with H/C, and this limits the decrease in temperature. The dependency of gasification temperature on O/C and H/C obtained in the present work is similar to that obtained by Prins et al.10 Figure 5 shows the contour plot of oxygen required (mole of oxygen per mole of carbon) for complete carbon conversion on H/C vs O/C coordinates for gasification of fuel using oxygen alone. For a given H/C, oxygen requirement decreases with increase in O/C. With increase in O/C, more oxygen is available 1604



Figure 5. Contour plot of oxygen required (mole of oxygen per mole of carbon) for gasification of fuel using oxygen.

Figure 6. Comparison of H2 concentration vs ROC for isothermal and adiabatic operation.

as part of the fuel and hence the oxygen required as external oxygen decreases with O/C. Oxygen in the fuel and oxygen from external gasification agent are equivalent stoichiometrically as proposed by Stemmler and Muller.12 However this is valid for isothermal operation only, where the rate of heat transfer is determined for a given temperature. For adiabatic operation the above equality is not valid. To prove the above inequality, simulations were carried out for two different fuels, CH1.44O0.66 and CH1.44O0.1 (which have same H/C ratio of 1.44) under isothermal (1200 K) and adiabatic condition (Q = 0) by varying the flow rate of oxygen. The concentration of H2 in syngas is plotted against relative oxygen content (ROC) for the two different fuels in Figure 6. ROC is defined as the ratio of oxygen in fuel and gasifying agent to oxygen stoichiometrically required for complete combustion (not discounting oxygen in fuel). As concluded by Stemmler and Muller12 for isothermal operation, the hydrogen concentration is a unique function of ROC. However under adiabatic condition, the hydrogen concentration is not a unique function of ROC, as can be seen in Figure 6. The same conclusion can also

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Figure 7. Contour plot of mole fraction of hydrogen in syngas for gasification of fuel using oxygen.

be drawn based on eq 8 (equilibrium relation for water gas shift reaction) of Stemmler and Muller12 in which the equilibrium constant depends on temperature which is determined by energy balance for the adiabatic case. Oxygen as a gasifying agent has no effect on the energy balance. However, the oxygen from the fuel influences the energy balance because of the dependency of heating value of fuel on oxygen content of fuel. Hence the gasification caused by one mole of oxygen in fuel is less than that caused by one mole of oxygen as gasifying agent. This difference is significant at sufficiently low temperatures at which exothermic reactions are favored. Consequently the reduction in oxygen as gasifying agent is less pronounced at high O/C ratios. For a given O/C with increase in H/C, oxygen required is almost same for low O/C ratios (less than ∼0.25). It increases with H/C for higher O/C ratios. At low O/C, although there is a decrease in temperature with H/C (Figure 4), the temperature is high enough and endothermic reactions (Boudouard and methane reforming reactions) are favored. This results in effective utilization of oxygen as a gasifying agent and hence oxygen required is independent of H/C. At higher O/C, the gasification temperature is low and decreases with H/C (Figure 4) increasingly favoring exothermic reactions resulting in completely oxidized products. This results in less efficient utilization of oxygen as gasifying agent and hence oxygen required increases with H/C (Figure 5). With increase in H/C, hydrogen can contribute for the gasification through methanation reaction, reducing the oxygen required. However this contribution is much less because methanation reaction being exothermic is not significantly favored even at the lowest temperature obtained for oxygen gasification (about 1000 K). The trends obtained in the present work are equivalent to that obtained by Prins et al.10 where the oxygen required is represented in terms of equivalence ratio. Figures 7 to 11 show the contours of mole fraction of H2, CO, CO2, CH4, and H2O in syngas on H/C vs O/C coordinates for gasification of fuel using oxygen at carbon boundary point. For a given H/C, with increase in O/C, the concentration of H2 and CO is almost constant for low O/C ratios. For higher O/C ratios, the concentration of H2 and CO decreases with O/C. The shift 1605



Figure 8. Contour plot of mole fraction of carbon monoxide in syngas for gasification of fuel using oxygen.

Figure 9. Contour plot of mole fraction of carbon dioxide in syngas for gasification of fuel using oxygen.

from constancy to decreasing trend occurs at a lower O/C with increase in H/C. This point of transition approximately corresponds to ∼1500 K, as can be inferred from Figure 4. Above 1500 K the endothermic reactions (Boudouard reaction, primary water gas reaction, and methane reforming) are favored resulting in formation of H2 and CO (Figures 7 and 8) with relatively no CO2, H2O, and CH4 (Figures 9, 11, and 10). This is the reason for the constant concentration of H2 and CO for low O/C ratios. At temperatures lower than 1500 K exothermic reactions (water gas shift reaction and backward reaction of methane reforming) are favored, resulting in the formation of CO2 (Figure 9) and H2O (Figure 11). As a result concentration of CO2 (Figure 9) and H2O increases (Figure 11) and hence concentration of H2 and CO decreases (Figures 7 and 8) with O/C for higher O/C ratios. For a given O/C ratio, with an increase in H/C ratio the availability of hydrogen increases. For low O/C ratios, increased hydrogen content in fuel results in formation of hydrogen only (Figure 7) without formation of H2O (Figure 11), CO2, and CH4

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Figure 10. Contour plot of mole fraction of methane in syngas for gasification of fuel using oxygen.

Figure 11. Contour plot of mole fraction of H2O for gasification of fuel using oxygen.

(Figures 9 and 10) due to high temperature (Figure 4). Therefore, with increase in H/C, the hydrogen concentration increases with corresponding decrease in CO concentration. For higher O/C ratios increase in hydrogen content of fuel results in the formation of H2 and also formation of CO2 (Figure 9), H2O (Figure 11), and CH4 (Figure 10) due to low temperature (Figure 4). Therefore, for higher O/C rations, with increase in H/C, concentration of H2, CO2, H2O, and CH4 increases (Figures 7, 9, 11, and 10) and concentration of CO decreases (Figure 8). Figure 12 shows the contour plot of molar ratio of H2 to CO for gasification of fuel using oxygen. For a given H/C ratio, H2/CO ratio is almost independent of O/C ratio. The value is almost equal to molar ratio of hydrogen to carbon in the fuel. Concentration of both H2 and CO remain constant and then decrease with O/C (Figures 7 and 8) giving the constant H2/CO ratio with respect to O/C ratio. For a given O/C ratio, H2/CO ratio increases with H/C ratio due to increase in H2 concentration (Figure 7) and decrease in CO concentration (Figure 8). 1606



Figure 12. Contour plot of molar ratio of H2 to CO for gasification of fuel using oxygen.

Figure 13. Contour plot of cold gas efficiency (CGE) for gasification of fuel using oxygen.

The contour plot of cold gas efficiency (CGE) for gasification of fuel using oxygen at carbon boundary point is shown in Figure 13. For a given H/C ratio, the plot of CGE shows a maximum with respect to the O/C ratio. With increase in O/C, the concentration of H2 and CO remains constant and then decreases (Figures 7 and 8). With an increase in O/C ratio, lower heating value (LHV) of fuel decreases (eq 12). With increase in O/C, the mass of syngas produced per unit mass of fuel decreases. As a result of these effects the CGE shows a maximum with respect to O/C ratio (Figure 13). With increase in H/C ratio, the point of maxima in CGE shifts toward lower O/C ratio. This is due to the formation of CO2 and H2O at lower O/C ratios with increase in H/C as explained earlier. For a given O/C, for lower O/C ratios with increase in H/C, the cold gas efficiency increases. With an increase in H/C, the hydrogen concentration increases (Figure 7) and concentration of CO decreases (Figure 8). The LHV of fuel increases with H/C (eq 12). Mass of syngas produced per unit mass of fuel decreases with H/C. The net effect is an increase in CGE with H/C ratio. For higher O/C

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Figure 14. Contour plot of gasification temperature for gasification of fuel using steam.

ratios, with increase in H/C, the CGE decreases. With an increase in H/C, the concentration of H2 increases and the concentration of CO decreases. LHV of fuel increases with H/C. Mass of syngas produced per unit mass of fuel increases with H/C. The net effect is a decrease in CGE with H/C ratio. For an O/C ratio of about 0.4, CGE is almost independent of H/C and is about 80%. Prins et al.10 show a dependency of CGE on O/C and H/C similar to that obtained in the present work. 3.1.2. Effect of Feedstock. The different feedstocks characterized by H/C and O/C ratios are represented as regions in Figures 5 to 13. From anthracite coal to biomass, the gasification temperature decreases from ∼3400 K to ∼1050 K due to higher O/C and H/C for biomass. The oxygen required for complete carbon conversion decreases from ∼0.5 mole oxygen/mole of carbon for anthracite coal to ∼0.25 mole of oxygen/mole of carbon for biomass due to rich oxygen content of biomass feedstock. The concentration of H2 in syngas obtained by gasifying anthracite coal varies from 5 to 15%. However, due to higher hydrogen content in biomass, concentration of H2 in syngas obtained by gasifying biomass is about 40%. The concentration of CO in syngas is about 8595% when anthracite coal is gasified. When biomass is gasified, the concentration of CO in syngas drops to about 55%. Accordingly the molar ratio H2 to CO varies from about 0.1 for anthracite coal to about 0.8 for biomass. The syngas obtained from the anthracite coal is relatively free of CO2. But the syngas obtained from biomass has about 10% CO2 due to lower gasification temperatures. While the concentration of CH4 in syngas is negligible for anthracite coal, syngas obtained by gasifying biomass contains about 1% CH4. The syngas obtained from biomass is relatively wet compared to that obtained from anthracite coal. The cold gas efficiency (CGE) for gasification of feedstock using oxygen varies from about 70% for gasification of coal to about 80% for gasification of biomass. In Figures 513, the point A represents a typical anthracite coal (O/C = 0.02, H/C = 0.2) and point B represents a typical biomass (O/C = 0.7, H/C = 1.5). The line joining A and B represents feed consisting of a mixture of anthracite coal and biomass of varying proportions. The gasification temperature 1607



Figure 15. Contour plot of steam required (mole of steam per mole of carbon) for gasification of fuel using steam.

and oxygen required for different composition of feed (% biomass in feed) can be obtained along the line AB. Similarly the composition of syngas obtained from cogasification of anthracite coal and biomass can be obtained by following line AB in Figures 513. Line AB in Figure 13 gives the variation of CGE for cogasification. Whereas the CGE is about 70% and 80% for anthracite coal and biomass, respectively, it shows a maximum of 82% for feed composition of 60% biomass (which corresponds to O/C = 0.79 and H/C = 0.33). This shows that the cold gas efficiency on cogasification can be higher than that obtained from individual feedstocks. 3.2. Gasification Using Steam. 3.2.1. Effect of O/C and H/C Ratios. Figure 14 shows the contour plot of gasification temperature on H/C vs O/C coordinates for gasification of fuel using steam at carbon boundary point. For a given H/C, the gasification temperature increases with O/C. With increase in O/C the relative contribution of oxygen in fuel to gasification compared to that of steam increases. Because reactions with oxygen are exothermic, the gasification temperature increases with O/C. For a given O/C with increase in H/C, the gasification temperature decreases. With increase in H/C, hydrogen content in the fuel increases, contributing to the gasification of fuel through methanation reaction. Because the methanation reaction is exothermic, the temperature is expected to increase with higher contribution by methanation reaction. However with increase in H/C, the enthalpy of fuel decreases much more than the heat released by exothermic reaction. So the gasification temperature decreases with H/C. The contour plot of steam required (mole of steam per mole of carbon) for complete carbon conversion on H/C vs O/C coordinates for gasification of fuel using steam is shown in Figure 15. For a given H/C, steam required decreases with O/C. With increase in O/C, increased oxygen content in fuel contributes to gasification of fuel. Hence the steam required for gasification decreases with O/C. For a given O/C, with increase in H/C steam required decreases. With increase in H/C, more hydrogen is available in fuel which contributes to the gasification of fuel through methanation reaction. Hence, steam required decreases with H/C. In contrast to gasification using oxygen, contribution of hydrogen to gasification is significant in

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Figure 16. Contour plot of mole fraction of hydrogen in syngas for gasification of fuel using steam.

Figure 17. Contour plot of mole fraction of carbon monoxide in syngas for gasification of fuel using steam.

steam gasification due to relatively low temperatures which prevail. Figure 16 shows the contour plot of mole fraction of hydrogen on H/C vs O/C coordinates for gasification of fuel using steam. For a given H/C, with increase in O/C the concentration of hydrogen increases. For a given O/C with increase in H/C the concentration of hydrogen decreases. It can be seen from Figures 14 and 16 that the concentration of hydrogen depends directly on the temperature. With increase in temperature the endothermic methane reforming and primary water gas shift reactions are favored resulting in higher concentration of hydrogen. The decrease in concentration of hydrogen due to backward reaction of water gas shift reaction is not significant because it is relatively less endothermic. The contour plot of mole fraction of CO for gasification of fuel using steam is shown in Figure 17. For a given H/C, with increase in O/C, the concentration of CO increases steeply. With increase in O/C, the gasification temperature increases (Figure 14). Hence, the endothermic primary water gas shift and methane 1608



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Figure 18. Contour plot of mole fraction of carbon dioxide in syngas for gasification of fuel using steam.

Figure 20. Contour plot of mole fraction of H2O for gasification of fuel using steam.

Figure 19. Contour plot of mole fraction of methane in syngas for gasification of fuel using steam.

Figure 21. Contour plot of molar ratio of H2 to CO for gasification of fuel using steam.

reforming reactions are favored resulting in the steep increase in the concentration of CO. For a given O/C, with an increase in H/ C the concentration of CO decreases. With increase in H/C, the gasification temperature decreases (Figure 14). Hence backward reaction of the above endothermic reactions is favored resulting in decrease in concentration of CO. Figure 18 shows the contour plot of mole fraction of CO2 on H/C vs O/C coordinates for gasification of fuel using steam. For a given H/C, with increase in O/C, the concentration of CO2 increases. With increase in O/C, contribution to gasification by oxygen in fuel increases resulting in the increased concentration of CO2. For a given O/C, with increase in H/C the concentration of CO2 decreases. With increase in H/C, contribution to gasification by H2 increases, reducing the steam requirement consequently the contribution of oxygen to gasification reduces resulting in lower concentration of CO2. The contour plot of mole fraction of CH4 on H/C vs O/C coordinates for gasification of fuel using steam is shown in Figure 19. For a given H/C, the concentration of methane

decreases with increase in O/C. For given O/C, concentration of CH4 increases with increase in H/C. Based on Figure 19 and Figure 14, it can be seen that the concentration of CH4 is inversely related to temperature. With an increase in temperature, the endothermic reverse reaction of methane formation and the methane reforming reactions are favored resulting in a decrease in CH4 concentration. Figure 21 shows the contour plot of molar ratio of H2 to CO on H/C vs O/C coordinates for gasification of fuel using steam. For a given H/C, molar ratio of H2 to CO decreases with increase in O/C. With increase in O/C, though the concentrations of CO and H2 increase, the rate of increase in concentration of CO is much higher than that of H2. So the molar ratio of H2 to CO decreases. For a given O/C, with increase in H/C molar ratio of H2 to CO increases. With increase in H/C, concentration of CO decreases more than that of H2 and hence molar ratio of H2 to CO increases. The contour plot of gasification efficiency on H/C vs O/C coordinates for gasification of fuel using steam is shown in 1609



Figure 22. Contour plot of cold gas efficiency (CGE) for gasification of fuel using steam.

Figure 22. For a given H/C, with increase in O/C the CGE decreases. With increase in O/C, LHV of fuel decreases. With increase in O/C, the concentration of CO and H2 increases but the concentration of CH4 decreases. With increase in O/C, the mass of syngas produced per unit mass of fuel decreases. The net effect is a decrease in CGE with O/C. For a given O/C, with increase in H/C, CGE increases. With increase in H/C, while the concentration of CO and H2 decreases, concentration of CH4 increases. With increase in H/C, the LHV of fuel increases. The mass of syngas produced per unit mass of fuel decreases with H/C. The net effect is an increase in CGE with H/C. It should be noted that the gasification temperature at the carbon boundary point obtained when using only steam as a gasifying agent is in the range of ∼550 to 750 K (Figure 14). At such low temperature the assumption of equilibrium is questionable. However the results of the present analysis help us in comparing the performance of gasification using oxygen and steam. For instance it can be concluded that if methane is desired in the syngas steam has to be the gasifying agent and not oxygen. 3.2.2. Effect of Feedstock. Similar to gasification using oxygen, the different feedstocks ranging from anthracite coal to biomass are represented as closed regions in Figures 1422. The temperature of gasification at carbon boundary point is about 600 K for anthracite coal and increases to ∼700 K for gasification of biomass. Compared to gasification using oxygen, the gasification temperatures are relatively much lower due to endothermic reaction between steam and carbon. The steam required for complete carbon conversion for gasification of anthracite coal is about 2.5 mole of steam per mole of carbon. However, the steam required reduces to about 1.5 mole of steam per mole of carbon for gasification of biomass due to higher oxygen content in biomass. The hydrogen concentration in synthesis gas obtained by gasifying anthracite coal is 10%. Higher concentration of hydrogen about 20% is obtained by gasifying biomass. The concentration of hydrogen in syngas obtained by using steam gasification is much lower compared to that obtained using oxygen gasification. Similarly the concentration of CO in syngas obtained by steam gasification is very low ranging from 0.1% to 1% depending on the feedstock. The molar ratio of H2 to CO for gasification of anthracite coal has very high value of about 70 and

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even reaching up to 100 for bituminous coal. This ratio drops to the range of 2040 for biomass. For a given feed stock, higher molar ratio of H2 to CO can be obtained using steam as gasifying agent rather than oxygen in which molar ratio of H2 to CO is in the range of 0.10.8 (Figure 12). The concentration of CO2 has almost the same value of about 42% for any feed stock when gasified using steam. The concentration of CH4 varies in the range of 40% to 30% depending on the feedstock. The concentration of CO2 and CH4 in steam gasification is much higher compared to that in oxygen gasification because of lower gasification temperatures favoring exothermic reactions. The syngas obtained by steam gasification is more wet than that obtained by oxygen gasification as can be seen in Figure 20 where the water content varies from 55% to 45% depending on the feedstock. The CGE decreases marginally from 97% for anthracite coal to 94% for biomass. The CGE for gasification of anthracite coal by steam gasification (97%) is much more than that obtained using oxygen gasification (70%). In the case of biomass gasification, the CGE achieved using steam gasification is 9395% and that achieved using oxygen gasification is 80%. In Figures 1422, the points A and B represent anthracite coal (O/C = 0.02, H/C = 0.2) and biomass (O/C = 0.7, H/C = 1.5), respectively. Along the line AB, the percentage of biomass varies from 0 to 100%. The gasification temperature and steam required for different composition of feedstock can be obtained by following the line AB in Figures 14 and 15. Similarly the line AB in Figures 1620 gives the composition of syngas for different composition of feedstock. The efficiency as a function of composition of feedstock can be obtained from Figure 22. Unlike oxygen gasification no maximum in efficiency with respect to feedstock composition is obtained in steam gasification.

4. CONCLUSIONS Thermodynamic equilibrium modeling of gasifier is useful in the preliminary design of gasifiers. In this work an equilibrium model based on a non-stoichiometric approach has been used to predict the gasifier performance. Whereas most of the existing studies focus on a specific fuel, the performance analysis carried out in the present work is applicable for any fuel. The simulation has been carried out for an adiabatic gasifier operating under atmospheric condition using Aspen Plus for O/C varying from 0 to 0.9 and H/C varying from 0 to 1.8. Attention is focused at the carbon boundary point at which maximum CGE is achieved. At this point, the gasification temperature, amount of gasification agent required, syngas composition, and CGE are determined for gasification using oxygen or steam. For gasification using oxygen, the gasification temperature decreases with both O/C and H/C. The oxygen required for complete carbon conversion decreases with O/C but increases with H/C at higher O/C ratios. Free oxygen is relatively more efficient in gasification than oxygen in fuel. The concentration of H2 shows a small decrease with O/C and increases with H/C. The concentration of CO shows a small decrease with O/C and decreases with H/C. The molar ratio of H2 to CO is almost independent of O/C and increases with H/C. The concentrations of CO2, CH4, and H2O are negligible for lower O/C ratios, but increase with H/C at higher O/C ratios. The CGE shows a maximum with O/C and increases and decreases with H/C at lower and higher O/C ratios, respectively. For gasification using steam, the gasification temperature increases with O/C and decreases with H/C. The steam required for complete 1610


Industrial & Engineering Chemistry Research carbon conversion decreases with both H/C and O/C. The concentration of both CO and H2 increases with O/C and decreases with H/C. The molar ratio of H2 to CO shows opposite trend. The concentration of CO2 decreases with H/C and increases with O/C. The concentration of CH4 increases with H/C and decreases with O/C. The CGE decreases with O/C and increases with H/C. In general, the gasification temperature is much higher when oxygen is used as gasifying agent as compared to when steam is used as a gasifying agent. Syngas from oxygen gasification is rich in concentration of CO and H2 and syngas from steam gasification is rich in concentration of CO2 and CH4. Higher molar ratio of H2 to CO is obtained in steam gasification compared to oxygen gasification. CGE is higher for steam gasification compared to oxygen gasification. The results presented in this work can also be applied to cogasification of feedstocks such as coal and biomass. An optimum proportion of biomass and coal exits to get maximum CGE when oxygen is used as a gasifying agent. However, no such optimum exits for gasification using steam. The results presented in this work will be useful in the preliminary design of gasifiers for any fuel or feedstock or combination of feedstocks using oxygen or steam as gasifying agents. Using the results of the present work, the gasification temperature, amount of gasifying agent required for complete carbon conversion, and composition of syngas can be obtained for optimal performance of gasifier for any feedstock. For cogasification of feedstocks (coal and biomass) using oxygen, the optimum proportion of biomass in feedstock can also be obtained. The results will be particularly relevant for the design of down-draft gasifiers where the assumptions of equilibrium model are valid. Gasifiers can be kinetically limited and hence operate far from equilibrium (like fluidized bed gasifiers). For these cases the results of the present work can be used as an initial guess while using rate-based models based on iterative methods.

’ AUTHOR INFORMATION Corresponding Author

*E-mail: [email protected]. Tel.: +914422574186. Fax: +914422570509.

’ ACKNOWLEDGMENT S.P. and V.R. thank Corporate R & D centre, Bharat Petroleum Corporation Limited, India, and T.R and Y.S.C. thank Korea Institute of Science and Technology, Korea, for financial assistance to carry out this work. ’ NOMENCLATURE Cpi = specific heat of ith component, kJ/kmolK Gt = total Gibbs free energy, kJ/kmol ^ C0 ,298 = standard specific heat of combustion of ith component H i at 298 K, kJ/kmol Hfuel = enthalpy of fuel entering the gasifier, kJ/s ^ 0f ,298 = standard specific heat of formation of ith component at H i 298 K, kJ/kmol Hprod = enthalpy of product stream, kJ/s Hreact = enthalpy of reactant streams, kJ/s H/C = hydrogen to carbon atomic ratio of fuel Mfuel = mass flow rate of fuel, kg/s Msyngas = mass flow rate of syngas, kg/s m = molar flow rate of oxygen, kmol/s

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nC, nH, nO = molar flow rate of atomic carbon, hydrogen, and oxygen entering through fuel, kmol/s ni = molar flow rate of ith component in syngas, kmol/s O/C = oxygen to carbon atomic ratio of fuel P = operating pressure of gasifier, atm Q = external heat transfer rate, kJ/s s = molar flow rate of steam, kmol/s T = gasification temperature, K Tin = temperature of steam entering gasifier, K X = fractional conversion of carbon xC, xH, xO = mass percentage of carbon, hydrogen, and oxygen in fuel (from ultimate analysis) yi = mole fraction of ith component in syngas λ298 = latent heat of vaporization of water, kJ/kmol

’ ABBREVIATIONS CGE = cold gas efficiency HHV = higher heating value, kJ/kg LHV = lower heating value of fuel/syngas, kJ/s ROC = relative oxygen content ’ REFERENCES (1) Higman, C.; van der Burgt, M. Gasification, 2nd ed.; Elsevier: UK, 2008. (2) Li, X.; Grace, J. R.; Watkinson, A. P.; Lim, C. J.; Ergudenler, A. Equilibrium modeling of gasification: A free energy minimization approach and its application to a circulating fluidized bed coal gasifier. Fuel 2001, 80, 195–207. (3) Smith, R. W.; Missen, W. R. Chemical Reaction Equilibrium Analysis: Theory and Algorithms; Wiley Interscience: New York, 1982. (4) Melgar, A.; Perez, J. F.; Laget, H.; Horillo, A. Thermochemical equilibrium modelling of a gasifying process. Energy Convers. Manage. 2007, 48, 59–67. (5) Altafini, C. R.; Wander, P. R.; Barreto, R. M. Prediction of the working parameters of a wood waste gasifier through an equilibrium model. Energy Convers. Manage. 2003, 44, 2763–2777. (6) Salaices, E.; Serrano, B.; Lasa, H. Biomass Catalytic Steam Gasification Thermodynamic Analysis and Reaction Experiments in a CREC Riser Simulator. Ind. Eng. Chem. Res. 2010, 49, 6834–6844. (7) Ptasinski, K. J.; Prins, M. J.; Pierik, A. Exergetic evaluation of biomass gasification. Energy 2007, 32, 568–574. (8) Jarungthammachote, S.; Dutta, A. Thermodynamic equilibrium model and second law analysis of a downdraft waste gasifier. Energy 2007, 32, 1660–1669. (9) Nguyen, T. D. B.; Lim, Y.; Song, B.; Kim, S.; Joo, Y.; Ahn, D. Two-stage equilibrium model applicable to the wide range of operating conditions in entrained-flow coal gasifiers. Fuel 2010, 89, 3901–3990. (10) Prins, M. J.; Ptasinski, K. J.; Janssen, F. J. J. G. From coal to biomass gasification: Comparison of thermodynamic efficiency. Energy 2007, 32, 1248–1259. (11) Vaezi, M.; Passandideh-Fard, M.; Moghiman, M.; Charmchi, M. Modeling Biomass Gasification: A New approach to utilize renewable sources of Energy. ASME International Mechanical Engineering Congress and Exposition, Boston, MA, October 31November 6, 2008. (12) Stemmler, M.; Muller, M. Theoretical Evaluation of Feedstock Gasification Using H2/C Ratio and ROC as Main Input Variables. Ind. Eng. Chem. Res. 2010, 49, 9230–9237. (13) Smith, J. M.; Van Ness, H. C.; Abbott, M. M. Chemical Engineering Thermodynamics, 6th ed.; McGraw-Hill: New York, 2003. (14) Channiwala, S. A.; Parikh, P. P. A unified correlation for estimating HHV of solid, liquid and gaseous fuels. Fuel 2002, 81, 1051– 1063. 1611


Generalized Analysis of Gasifier Performance using Equilibrium

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