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Modeling Tar Handling Options in Biomass Gasification Seethamraju Srinivas, Randall P. Field,* and Howard J. Herzog MIT Energy Initiative, Massachusetts Institute of Technology (MIT), 77 Massachusetts Avenue, Cambridge, Massachusetts 02139, United States ABSTRACT: Biomass gasification is receiving increasing attention as a potential source of sustainable development for the production of power, chemicals, and transportation fuels. A major hurdle, however, despite years of research, is the removal of tar formed during the biomass gasification process. In this paper, we model some of the tar handling options using the process simulator, Aspen Plus. While there is extensive experimental work reported in the literature on the tar conversion and absorption methods, the literature on modeling is sparse. Hence, as a part of the study, we present the challenges in modeling the tar handling methods. We report the performance of three frequently suggested tar handling options: uncatalyzed partial oxidation, catalytic steam reforming, and absorption. Some of the practical difficulties that need to be overcome with each of these methods are also discussed. Hamelinck et al.2 emphasize that primary treatment alone does not suffice to eliminate tars, especially for syngas applications involving production of chemicals or transportation fuels. Secondary treatment of the tar-laden syngas downstream of the gasifier is essential, which is the focus of this work. Milne et al.19 and Neeft et al.20 provide excellent reviews of the secondary treatment measures for tar elimination. Torres et al.21 review the use of basic, acidic, metallic, and redox catalysts for the removal of tars, ammonia, and H2S from biomass gasification syngas under hot conditions. Aravind and de Jong22 present a detailed review on using high-temperature gas cleaning alternatives for using syngas derived from biomass gasifiers in solid oxide fuel cells (SOFCs). Both primary and secondary treatment measures are discussed with regards to the removal of tars, particulates, H2S, HCl, and alkali components and their impact on the performance of the SOFCs. On the basis of the operating temperature (T), Woolcock and Brown23 categorize syngas cleanup into hot (T > 300 °C), cold (T < ∼100 °C), and warm (100 < T < 300 °C) methods and review them in detail. Tar removal options under the three temperature classifications are discussed too. Secondary elimination measures can be classified as physical (scrubbers and filters) or reactive type. Hot gas filters (dry) and absorption (wet) are representative of the former, while partial oxidation and steam reforming are included in the latter. Hot gas cleaning methods, such as partial oxidation or catalytic reforming, are preferred over wet methods, such as absorption, because they transfer the energy present in tar components to syngas (transformation into CO and H2) rather than creating a waste stream (e.g., wastewater), which is difficult to dispose.24 Rabou9 presents an interesting solution, tar recycling within the gasification process, and presents results of tar recycling experiments from a circulating fluidized-bed (CFB) gasifier. Downstream of the gasifier, the tar−water mixture from the electrostatic precipitator (ESP) is recycled back into the gasifier for combustion. The idea is shown to be technically feasible, with
1. INTRODUCTION With the increasing concern over energy security and sustainable development, biomass is being considered as an alternative for the production of power, chemicals, and transportation fuels.1−6 Thermochemical pathways, including gasification and pyrolysis, have received great attention as a means of converting biomass into liquid fuels. Although they have been proven at the demonstration scale [e.g., Gussing plant in Austria that uses wood chips and a fluidized-bed gasifier with Fischer−Tropsch (FT)7], there are no commercial, large-scale biomass-to-liquid (BTL) plants yet because of the high capital costs, uncertainty in some of the processes involved, uncertainty in government policies, and high feedstock costs.8 One of the associated process uncertainties is the removal of tars formed during gasification (e.g., about 10 g/m3 of tar produced in fluidized beds or ∼1−5 wt % of the biomass feed2 or ∼3% of the biomass energy content9). The definition of tars is not consistent across the literature, but it typically includes heavy aromatics and excludes benzene.10 The tars are problematic to the process equipment because they condense at moderate temperatures, cause blocking of filters and accumulation and choking of pipes and valves, are corrosive, and cause catalyst deactivation. It is, hence, necessary to eliminate them by either primary or secondary treatment. The aim of primary treatment is to optimize the gasifier design and operating conditions such that the syngas produced has a minimum tar concentration. Primary treatment is difficult and has limited ability to reduce the tar content at common biomass gasification temperatures. Primary treatment involves proper selection of gasifier operating parameters (pressure, temperature, input steam and oxygen/air flow rates, residence time, etc.) as well as a good design of the gasifier (point of introduction of the feed, degree of channeling or circulation, feed particle size distribution handled, secondary air/oxygen injection in the free-board zone, use of in-bed catalyst, etc.). A representative but not exhaustive list of works that looked at tar formation and destruction is shown in Table 1. Readers are referred to the study by Devi et al.11 for a review on the primary tar treatment options in biomass gasification. © XXXX American Chemical Society
Received: March 5, 2013 Revised: May 6, 2013
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Table 1. Representative Works That Study the Effect of Various Parameters on Tar Formation/Destruction reference Gil et al.12 Brage et al.13 Li et al.14 Corella and Sanz15 Kaushal et al.16 Aljbour and Kawamoto17 Jordan and Akay18
biomass and gasifier used pine wood; bubbling fluidized bed birch wood; pressurized fluidized bed circulating fluidized bed pine wood; circulating fluidized bed wood pellets; bubbling fluidized bed cedar wood; updraft gasifier followed by reforming cane bagasse; downdraft gasifier
remarks effect of gasifying agent used on product distribution and gas quality tar evolution profile as the gasification temperature changes tar concentration mainly depends upon gasifier operating temperature location of biomass feeding (from the top or bottom of the bed) and presence of secondary air or in-bed catalyst affects the cracking of tars formed and, hence, the final content in the exit syngas effect of the temperature and residence time (height of free board zone) on tar cracking; compare simulation results to experimental data effect of operating parameters (residence time, air and steam flow rates, and gasifier temperature) on tar levels and nitrogen and sulfur impurities effect of varying CaO additive amounts on the tar composition, concentration, dewpoint, and syngas yield
referred to section 4.2 for a further discussion on the mechanism aspects. 2.2. Steam Reforming. Catalytic steam reforming is also useful in lowering the tar content in syngas from biomass gasification. Both metallic (Ni) and non-metallic (dolomite) catalysts have been used in situ as a primary means of tar elimination in biomass gasification. However, the in situ results were not encouraging, owing to issues of coking/deactivation (Ni) and friability (dolomite).19,24 To overcome this problem, secondary treatment using these catalysts has been suggested and has proven to be more effective. The strategy used by most researchers has been to use a guard bed of calcined dolomite followed by a second bed having steam-reforming catalysts, such as Ni.19,26,35 The dolomite guard bed reduces the tar concentration to within acceptable limits that the Ni catalytic bed can handle. We have chosen this two-bed strategy in our simulation model configuration, as described in section 5. Wu and Williams36 review several types of Ni-based catalysts used for tar reduction in biomass gasification with regards to their synthesis, reaction conditions, and stability. Olivine and CeO2 are typical promoters that are added to these catalysts.24 On the basis of their testing of many Ni-based commercial catalysts for steam gasification of biomass, Aznar et al.35 concluded that naphtha reforming catalysts are better suited for tar reduction than methane reforming catalysts. Yung et al.37 review the various catalyst formulations that have been used in steam reforming of tars to condition biomass-derived syngas. They also discuss the role of catalyst additives and deactivation mechanisms for tar reforming catalysts. An interesting comparison between fixedbed and fluidized-bed operation for tar reforming is presented too. 2.3. Absorption. Scrubbing or absorption by a liquid medium is one of the physical strategies to remove tar from syngas. One might use water as the scrubbing liquid [e.g., Harboore plant as described in the report by Zwart25], which effectively cools the syngas and removes the tar species. However, it produces large amounts of wastewater, requiring downstream treatment; in other words, the problem has shifted from tar removal to wastewater treatment! Rabou et al.38 present three possible schemes tried out for water scrubbing and their associated problems. A major deficiency of water as a scrubbing medium is that most of the tar components are hydrophobic in nature. Hence, a better alternative is to use an organic solvent, such as vegetable oil, biodiesel, or diesel.39,40 At the Gussing demonstration plant in Austria, a rapeseed methyl ester (RME) absorber operating at 5 °C was shown to successfully remove tars from syngas.25 Research at Energy Research Centre of the Netherlands (ECN) over the last 2 decades has led to the
the reported tar destruction higher than expected for thermal cracking, but doubts remain over its economic viability. Partial oxidation and reforming are the most promising technologies that the scientific and engineering community is looking at for tar removal, followed by adsorption and/or absorption. Note that absorption has been proven on a demonstration-scale successfully.25 Each of these secondary tar treatment options is briefly discussed in section 2. Note that the focus is on partial oxidation, catalytic steam reforming, and absorption in the rest of this paper after section 2. Following the introduction to the available tar handling options, the two major inputs required for modeling them, the tar component characterization and choice of model components and the reaction kinetics, are discussed in sections 3 and 4, respectively. The simulation methodology and results are presented in section 5, with the succeeding section describing some of the practical issues with each of the three chosen tar handling options. Concluding remarks and potential future work form part of the last section.
2. SECONDARY TREATMENT OPTIONS 2.1. Partial Oxidation (POX). Non-catalytic partial oxidation can be used to lower the tar content in syngas from biomass gasification. This requires high temperatures in the range of 900−1150 °C with a residence time varying between 1 and 12 s.26 For a similar process, 1250 °C with a 0.5 s residence time is reported by Brandt and Henriksen.27 More recently, tar destruction in a partial oxidation environment has been studied by different authors using different biomass feeds (e.g., see refs 28−30). Uncatalyzed methane reforming requires high temperatures and is a part of the POX process. Bruggemann et al.31 investigated the influence of pressure and temperature on noncatalytic POX and reforming of natural gas in pilot-scale experiments. They observed that methane conversion increased with the temperature but did not approach equilibrium under the conditions used (T, 1450 °C; P, 30−100 bar, and t, 5−20 s). To achieve such high operating temperatures in a POX reactor, part of the chemically bound energy in the gas is used, thereby lowering the efficiency of the process.32 Houben et al.33 and van der Hoeven34 investigated partial combustion (oxidation) for tar reduction with an emphasis on the burner design and geometry using naphthalene as the model tar component. They studied the effect of varying oxygen and hydrogen levels. On the basis of the oxygen and hydrogen contents in the syngas, the tar components can either crack into light gases or polymerize and grow into larger molecular weight polyaromatic hydrocarbons (PAHs). Su et al.30 report both of these paths too and suggest that their mechanisms are different from inert conditions. Readers are B
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high temperatures, there are others that condense only at very high concentrations and low temperatures. Li and Suzuki47 reiterate the same in their overview of tar property and analysis for tar elimination by catalytic reforming. However, another approach classifies tars as being primary, secondary, or tertiary based on their ease of destruction. A detailed list of the organics present in primary, secondary, and tertiary tars is given in the report by Milne et al.19 Tar is formed as a part of the devolatalization process in the gasifier. The tars thus formed can undergo secondary reactions in either the vapor phase (homogeneous) or the solid phase by reacting with char (heterogeneous). It is, hence, essential to have a good devolatalization model as well as a kinetic model to describe the secondary reactions. There are many models presented in the literature to model biomass gasifiers of all types, fixed bed (both updraft and downdraft) and fluidized bed (both bubbling and circulating). Gomez-Barea and Leckner48 present an exhaustive review on fluidized-bed biomass gasification, while Puig-Arnavat et al.49 discuss biomass gasification models in general. Most of these models use an equilibrium approach, which does not model tar components at all, neither its formation nor its destruction.49 Using the kinetic approach, there are some models that consider tar but use a single component to represent it, either a true component or a pseudo-component.50,51 While such an approach may suffice in the context of overall gasifier modeling, the authors believe that one should consider a broader range of components to represent tar when looking at downstream tar handling options. Moreover, published experimental works on tar removal tend to consider both approaches, using a single component, such as naphthalene, to perform experiments or using tars formed in situ and then classifying them. Because the currently available models cannot adequately predict tar formation (quantity and quality) under typical gasifier operating conditions, it was decided to make use of the experimental data of measured tar component concentrations to define the syngas stream used in our tar handling simulations. Evans et al.52 present data and results from extensive experimental work on oxygen-blown, pressurized gasification of woody biomass in CFBs under a variety of conditions (gasifier temperature, pressure, biomass feed rate, feedstock type and moisture, steam/biomass ratio, and fluidized-bed height). A total of 22 tests were conducted in the program. Naphthalene and benzene are reported to be the major species present in the oils (tar) produced. No secondary tar treatment is considered in their work, and the tars formed are well-characterized. Their data are, hence, used as the inlet basis in our modeling work. Note that this work formed a part of a larger scope of work that looked at pressurized, O2-blown biomass gasification for fuel synthesis (hence, our interest in the data from Evans et al.52). To the best of our knowledge, this is a first attempt at trying to compare the three chosen tar handling options under pressurized conditions. Most of the studies in the literature dealing with these options (partial oxidation, catalytic reforming, or absorption) are at atmospheric pressure, which is the typical operating pressure for most existing biomass gasifiers. Such a study is interesting because elevated pressure introduces challenges for each of these tar handling options (e.g., increased condensation during absorption making dew point control difficult in OLGA process, increased soot formation in POX reactors, increased operating temperature for reforming with dolomite catalyst, owing to decalcination concerns, etc.). Although tar has a large number of components, only a few can be chosen for developing a simulation model for reasons of
conclusion that mixing of dust, tar, and water is not good and helped in the development of an oil-based scrubbing process called the OLGA process, which uses a proprietary solvent. A detailed description and principle of the OLGA process can be found in ECN publications,41,42 including a discussion on the oil recovery system (ORS) used, which separates the heavy tars from the collector oil loop while maintaining the light fraction of the tars within the loop. The recovered tars are recycled back to an appropriate zone in the gasifier to be destroyed (gasified). Methane passes through the OLGA process unchanged. 2.4. Tar Cracking. Cracking of tars can be performed with heat alone (thermal) or in the presence of both heat and catalysts (catalytic). There is some amount of both of these processes present in a biomass gasifier because of the heat inside as well as the catalytic effect of ash components present in biomass. Fjellurup et al.26 discuss in detail the decomposition and cracking of tars during biomass gasification. A summary of tar cracking catalysts with their relative merits and demerits can be found in the report by Zwart.25 The high-temperature operation for tar cracking leads to a penalty in the cold gas efficiency (CGE). Researchers from ECN report that a minimum temperature of 1150 °C is necessary to thermally crack tars to a level below 100 mg/m3, resulting in an approximate 8% CGE loss.25 Three pathways of tar cracking, thermal, catalytic, and plasma, are discussed by Rabou et al.38 2.5. Adsorption. Paethanom et al.40 consider adsorption as an alternative tar removal technique. It is interesting to note that they differentiate between the physical nature of different tars; light tars are assumed to be in the vapor form, and heavy tars are assumed to be in the form of liquid droplets. Accordingly, they propose absorption and adsorption for heavy and light tar removal, respectively. Using a combination of vegetable oil scrubber and rice husk char adsorbent bed, they report ∼95% removal of gravimetric tar. The disadvantage of adsorbents, though, is the need to regenerate them by an appropriate method. Note that rice husk char may not be regenerated, owing to its ease of availability, but other adsorbents might need to be regenerated depending upon the cost, availability, etc. 2.6. Catalytic Filtration. Ceramic or metallic filters are conventionally employed for particulate removal in syngas cleanup systems. A recent trend has been to manufacture filter elements that have a catalytic action too. This serves two purposes, removal of dust and tar contaminants, both at relatively high temperatures (750−900 °C). The review by Heidenreich43 covers the fundamental aspects of filtration at higher temperatures with a discussion on tar contaminant removal from biomass gasifiers also. Use of a ceramic candle filter containing a Ni-based tar cracking catalyst is an example of this kind of system.44 Nacken et al.45 tested catalytic filtration using Ni with a variety of supports to achieve complete conversion of benzene and naphthalene (model tar components) at 900 °C. However, such catalytic filters have only reached the “proof of concept” stage,23 and there are challenges yet to be resolved to make them commercial.
3. COMPONENT REPRESENTATION: TAR CHARACTERIZATION van Paasen and Kiel46 characterize tars into five categories based on their condensation behavior and water solubility (see Appendix 1). Note that in determining the effect of tars on downstream processes, both the individual concentrations of the components as well as the tar dew point are important. While some classes of tars condense at very low concentrations and C
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Table 2. Representative Works Showing Surrogate Tar Components reference Kinoshita et al.53 Jess54 Houben et al.33 Zhang et al.24 Gerun et al.28 Fourcault et al.55 Gomez-Barea and Leckner48 Su et al.30 Wu et al.56 Sarioglan57 Damartzis et al.58 Abdelouahed et al.59
model tar components considered benzene, naphthalene, one-ring compounds, two-ring compounds, three- and four-ring components, and phenolic compounds benzene, toluene, and naphthalene naphthalene only, in both the modeling and experimental parts of the work benzene and toluene
comments use six lumps as shown in the adjacent column; benzene and naphthalene are treated separately from the one- and two-ring components, respectively investigate kinetics of thermal conversion in the presence of H2 and steam being a two-ring aromatic hydrocarbon, naphthalene can exhibit both reaction paths of interest (the aim of the experiment), crack into lighter or polymerize into heavier investigate steam reforming of tar components over Ni/olivine catalysts in a fixed-bed reactor investigate partial oxidation in a two-stage downdraft gasifier study tar removal using a high-temperature plasma torch
benzene, phenol, and naphthalene benzene, toluene, and naphthalene representative of primary, secondary, and tertiary tars, respectively acetol, anisole, and phenol for primary tars; toluene as components listed are from cross-references within; naphthalene chosen as the overall an alkyl tertiary product; and naphthalene as a PAH model tar compound because it is hard to destroy and is expected to give an estimate of compound the overall tar concentration benzene, toluene, phenol, and naphthalene compare experimental data to model results for performance of tar destruction using partial oxidation benzene, toluene, styrene, phenol, and naphthalene use rice/straw as biomass feed to study tar evolution; identify the components shown as most typical tar species present under all decomposition conditions benzene, toluene, and xylene study tar removal performance by catalytic steam reforming over dolomite and metal catalysts benzene, toluene, and naphthalene biomass gasifier modeling using Aspen Plus for CHP applications; tar oxidation alone is considered benzene, toluene, phenol, and naphthalene detailed modeling of dual-fluidized-bed biomass gasifier using Aspen Plus with an emphasis on the water−gas shift reaction kinetics
4. TAR DESTRUCTION CHEMISTRY: KINETICS AND REACTION PATHWAYS 4.1. Kinetics. Three agents contribute to tar degradation in the partial oxidation and reforming reactors: heat (temperature) resulting in thermal cracking, steam leading to steam reforming (catalytic or non-catalytic), and oxygen (or air) causing oxidation. Each of them affect tar reduction in a different way. The physics of the phenomenon that affects tar formation or destruction is the same in both the gasifier and the downstream secondary tar reduction reactor. Hence, the gasifier environment is referred to in some places in the text. 4.1.1. Effect of the Temperature. It is commonly believed that tars thermally crack into lighter gases. However, this is true for cracking of primary tars only. Tertiary tars tend to grow in molecular weight and polymerize as reaction severity increases.19,33,38,56,60 In other words, as temperature increases, the total amount of tar decreases but the concentration of compounds with a larger number of rings increases.61 Elliott62 points out this surprising feature of increasing temperature (severe gasification conditions), which results in a more refractory kind of tar that may be harder to destroy compared to primary or secondary tars that are produced at lower gasification temperatures. 4.1.2. Effect of the Steam Addition. The addition of steam aids in the production of fewer refractory tars, increases phenol formation, lowers the concentration of other oxygenates, and produces “tars” that are easier to reform catalytically.12,19 From results of gasification experiments using sawdust in a drop-tube furnace with different gasification agents, Zhang et al.63 observed that steam does not influence the tar species destruction significantly but oxygen does, and oxygen is more reactive than steam for tar destruction. Similar conclusions have been reported by Striugas et al.64 when comparing steam reforming and partial oxidation of biomass pyrolysis tars over activated carbon catalyst in a fixed-bed reactor. Jess54 also reports little or no influence of steam on thermal conversion of aromatics. Contrary to previous statements, Evans et al.52 report a lowering of oxygenated oil components and overall phenols yield as the steam/biomass ratio
simplicity and kinetic data availability. Table 2 lists some of the works that represent tar using different components. The following four components are chosen as model tar compounds in our study for the reasons stated: (a) phenol, represents oxygenates, an easy-to-destroy component; (b) benzene, represents primary tar, an intermediate product formed during aromatic group cracking or thermal decomposition of higher tar species, such as naphthalene (despite not being a tar component, benzene is included, owing to its importance in the reaction mechanism and also because of its consideration by many groups as shown in Table 2); (c) toluene, represents secondary tar, a primary pyrolysis tar component, and has an alkyl group that can react easily; and (d) naphthalene, represents tertiary tar, a multi-ring component that can crack into lighter species or polymerize to heavy molecular weight components and a hard-to-destroy component. Note that the components that we chose are also used by other research groups as shown in Table 2. Further, kinetics for these four components is available (see section 5). The tar components in the report by Evans et al.52 are characterized as one-, two-, three-, four-, and five-ring and higher aromatic compounds, oxygenates, N- and S-containing compounds, and unidentified compounds. They are mapped into the chosen model tar components as follows: (i) phenol, all oxygenates from original data52 lumped together; (ii) benzene, benzene from original data; (iii) toluene, toluene and xylene and all other one-ring aromatics from original data; and (iv) naphthalene, all two-, three-, four-, and five-ring and higher aromatic hydrocarbons from the original data lumped together. N- and S-containing compounds and unidentified components from the original data are not considered. The composition of CH4 includes CH4, C2H4, C2H6, and C3H8 from the original data. This simplification is justified because there is not enough experimental data available for validating the fate of these species during tar destruction. However, if one wants to include them, the model would need to include kinetics of partial oxidation or steam reforming for these C2 and C3 hydrocarbons. D
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lumps is not specified. The final kinetic models chosen for this work are discussed in section 5. 4.2. Reaction Pathways. Numerous reactions are possible during tar destruction because of the complex nature of gasification gas with many components (CO, H2, CO2, H2O, tars, CH4, etc.) and their interactions under hot conditions.68 This complicates the tar chemistry modeling for hot gas cleaning options, such as POX and catalytic reforming, and is challenging to create a reaction set that closely mimics the real behavior. To add to this, one also needs the right values of the corresponding kinetic parameters for the chosen set of reactions under the conditions of interest, which, in itself, is another challenge. Another important consideration is that the chosen mechanism should achieve a balance between accuracy and computation time.30 Tar destruction chemistry is a combination of kinetic and equilibrium reactions (e.g., thermal decomposition of tars and reactions such as water−gas shift, respectively). A representative set of such reactions as presented by Simell et al.68 are given in Appendix 2, with toluene as the model tar component. Chen et al.60 report that oxygen initiates vital reactions necessary for tar reduction because it is a very good initiator of free radicals. Oxygen also combusts part of the syngas, releasing heat, which is useful for endothermic reforming reactions by steam or aids chain growth, and chemically reactive species (free radicals) to propagate other reactions. From their partial oxidation studies on nascent tar from sawdust pyrolysis, Hosokai et al.29 conclude that oxygen consumes hydrogen and light gases selectively while leaving the heavy tars intact. The tar reduction reactions are then a result of the reactions initiated by free radicals formed during oxidation of hydrogen and other light gases. Similar conclusions are reported by Su et al.,30 who investigated tar destruction under partial oxidation conditions using rice straw as the biomass feed. They conclude that tar destruction is influenced significantly by the levels of oxygen present, and oxidizing conditions are more suitable than inert conditions for tar removal. Both hydrogen and oxygen play an important role in the tar reduction chemistry. van der Hoeven34 discusses this in detail. The oxygen level in the reactor dictates if the initial ring rupture leads to cracking (further ring breakage) or polymerization (ring growth). By competitive inhibition, H2 hinders polymerization reactions and provides increased driving forces for cracking to occur. Owing to fast reaction rates of H radicals with unsaturated molecules, the unsaturated molecules cannot recombine; viz., soot formation is suppressed. However, the chemistry is more complicated and is dependent upon the rate-determining step in the cracking path. Jess54 shows that the effect of H2 may not always be positive on the formation of cracking products. Sarioglan57 observed hydro- and steam dealkylation of alkyl aromatics as well as selective reforming of alkyl groups on both dolomite and metal-based catalysts. The author indicates favorable temperature ranges for different reactions, 500−600 °C for steam dealkylation reaction and above 700 °C for thermal degradation/polymerization. Interestingly, they note that the presence of tar and methane together is advantageous in increasing methane conversion but leads to the formation of undesired (and immeasurable) soot particles. On the basis of a comparison between steam and dry reforming reactions, Simell et al.68 conclude that dry reforming is thermodynamically more favorable as the temperature increases. They suggest that both reforming and thermal decomposition reactions of tar components are important over both dolomite and Ni catalysts.
is increased in the gasifier. Thus, more data and studies are possibly needed to elucidate the effect of steam addition on tar reduction in biomass gasification. 4.1.3. Effect of the Oxygen (Air) Addition. Oxygen or air added, in combination with steam, helps produce less reactive (more refractory or difficult to destroy) tars at lower levels and increases conversion of primary tars.12,19,65 Selective oxygen addition to various stages of the gasifier bed results in preferential tar oxidation.19 Houben et al.33 noticed that naphthalene (model tar component considered) polymerizes in a partial oxidation environment as the equivalence ratio (amount of air added) is increased. Owing to these three major contributors, most of the kinetic models in the literature for tar reduction deal with tar cracking, reforming, and oxidation reactions. Note that non-catalytic reforming requires increased severity (higher temperature) compared to catalytic reforming. 4.1.4. Models Available. A good discussion on tar cracking kinetics can be found in the study by Di Blasi.66 A table of kinetic constants from different literature sources for the global reaction of tar cracking based on either total or reactive mass concentration of the vapor-phase tar is presented with the difference in the kinetic parameter values attributed to the different rate expressions used. However, most of the sources listed in the table consider tar as a single lump component of CaHbOc type or do not specify its chemical formula and, hence, is not appropriate for the modeling work performed in our study. Boroson et al.50 studied the homogeneous vapor-phase cracking of wood pyrolysis tar at low molar concentrations in the temperature range of 773−1073 K and reactor residence time between 0.9 and 2.2 s. The observed tar conversion was between 5 and 88%. The resulting experimental data were used to develop kinetics for tar cracking and product formation. They represent tar as a pseudo-component with formula CxHyOz, which after cracking yields H2, CO, CO2, CH4, and a secondary tar, Cx1Hy1Oz1. This kinetic model or its variant is widely used in modeling of biomass gasifiers.16 However, this kinetic model is not suitable for our work because primary and secondary tar components are not explicitly defined; it is a lumped component. Being a hydrocarbon, tar is also subject to combustion in the presence of oxygen or air in the gasifier. Bryden and Ragland51 developed a kinetic model for the partial combustion of tarry materials representing tar as CH1.522O0.0228 with a molecular mass of 90 kg/kmol, yielding CO and H2O. Aznar et al.35 remark that a simple first-order kinetic model with tar represented as a single lump needs improvement and one should consider the varying reactive nature of the different tar components in future models. Hence, Corella et al.67 proposed two advanced kinetic models for catalytic tar removal in biomass gasification. They consider tar as either a continuous mixture or being composed of six different lumps. While the first kinetic model provides information regarding the mean molecular weight of tar and the variance of tar molecular weight distribution, the second kinetic model uses a set of 6 kinetic equations and 11 kinetic constants to describe the behavior. It is interesting to note that the second model permits ranking of tar species according to their reactivity. The authors also point out that the kinetic constants estimated are specific for the given gasifier, feedstock, and process and extrapolation of these values to different operating conditions might not be safe. Despite the novel approach, this model is not used in our work because the single component that is used to represent each family in the six E
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Jess54 investigated the thermal conversion of aromatic hydrocarbons in the presence of steam and hydrogen. It may be noted that the experiments had higher amounts of H2 because the author was trying to mimic coke-oven gas. Temperatures of ∼1200 °C are proposed for the thermal conversion of the aromatics, and further higher temperatures of 1400 °C are proposed for the conversion of the soot and CH4 (by uncatalyzed steam reforming) into CO and H2. In proposing mechanisms of primary and consecutive reactions for the species involved, it is reported that benzene is the key component of thermal decomposition of aromatic hydrocarbons and soot formation is mainly from naphthalene. Thus, soot formation appears inevitable in the partial oxidation/reforming environment and forms the subject of discussion in the next subsection. 4.3. Soot Formation. Tesner and Shurupov69 investigated soot formation during pyrolysis of naphthalene, anthracene, and pyrene and suggested the following order for sooting tendency relative to methane at 1350 °C: methane, ethylene, p-xylene, toluene, benzene, acetylene, diacetylene, pyrene, anthracene, and naphthalene (1, 4, 4, 5.5, 7.4, 7.6, 50, 74, 91, and 112). Joo and Gulder70 experimentally investigated the influence of pressure on soot formation in laminar methane−oxygen diffusion flames up to 100 atm in a high-pressure combustion chamber. They report an increase in soot levels with an increased pressure between 10 and 40 atm and a decrease in soot concentrations with a further increase in pressure. Thus, the effect of the pressure on combustion (and soot formation) cannot be neglected. It is interesting to note that Bruggeman et al.31 report no coke formation in their high-pressure partial oxidation reactor using natural gas as the reactant; this might be possibly because of the oxygen levels used, the flame conditions (diffuse or premixed), burner geometry, and the high-pressure operation (30−100 bar). Svensson et al.32 studied soot formation during partial oxidation of producer gas to identify the effect of different components (H2, CO, CO2, CH4, tar, and H2O) present in it. They observed less impact of steam and H2 and greater impact of methane and CO2 on soot formation. Higher levels of methane increase the sooting tendency, while an increase in inlet CO2 levels suppresses soot formation. The authors also point out that both the amount of oxidant (combined amount of O2, CO2, and H2O) and the kinetics play an important role during soot formation. Jess54 reports inhibition of soot formation by H2, possibly because of the high amounts of H2 used in the feed gas. Suppression of soot inception by hydrogen has also been reported by Du et al.,71 despite an increase in the flame temperature, and it is attributed to the large diffusivity of H2. Roesler et al.72 performed studies on the role of CH4 on the growth of aromatic hydrocarbons and soot in fundamental combustion processes. They conclude that an increase in the CH4 content of the fuel in a flame containing small amounts of aromatics results in the conversion of carbon into soot. This feature of CH4 combining with other aromatics to form PAHs (soot) is attributed to its ability to produce methyl radicals. Also, Roesler et al.72 found that benzene, naphthalene, and pyrene exhibit the strongest sensitivity toward methane. Sarioglan57 and Houben et al.33 also reached the same conclusion regarding the effect of CH4 on soot formation from their studies on catalytic reforming and partial oxidation, respectively, for tar elimination. van der Hoeven34 observed that steam gasification of soot starts at 1100 °C, reaches a maximum at ∼1250 °C, and decreases later on to finally level off at ∼1400 °C. The reason is possible
changes in the soot structure into less reactive soot. A similar behavior is reported by Aoki et al.73
5. SIMULATION: METHODOLOGY AND RESULTS The strategy used to model the reactors is to have two reactors in series, as shown in Figure 1. The motivation to adopt such a
Figure 1. Two-bed reactor strategy used in the simulation work.
strategy was explained earlier in this paper. For the case of POX, reactor 2 in Figure 1 is redundant. In the case of reforming, both the reactors are present. The different cases modeled are shown in Table 3. The reactions and the corresponding kinetic parameters Table 3. Reactor Modeling Options Used option uncatalyzed partial oxidation (POX) autothermal reforming (ATR) steam reforming (SR)
reactor 1
reactor 2
source of heat
uncatalyzed
none
partial oxidation
uncatalyzed
Ni catalyzed Ni catalyzed
partial oxidation
dolomite catalyzed
radiant heat exchange
used in the reactor models are shown in Tables A4.1 and A4.3 and Tables A4.2 and A4.4, respectively (see Appendix 4). On the basis of the discussion in the previous section, the set of reactions and the corresponding values of the kinetic parameters are chosen from refs 28, 30, 73, and 74. A few points to note in this context are as follows: (i) Soot oxidation is not considered in the kinetic model proposed by Fourcault et al.55 because the oxygen in the feed is sub-stoichiometric and is assumed to be consumed before soot is produced. For the same reason, we have also not included soot oxidation in our reactor kinetic model. (ii) Gerun et al.28 point out that two reactions are necessary to describe naphthalene conversion: one for its polymerization yielding soot and another for cracking to lighter species. However, owing to the lack of sufficient kinetic data to differentiate the two mechanisms, a single reaction is considered for naphthalene degradation. (iii) References that looked at tar destruction28,30,55 do not report the methane or soot (carbon) concentration in the exit syngas in any of their studies. Hence, it was not possible to validate the exit concentrations of CH4 and soot from our results. (iv) POX reactor has uncatalyzed steam methane reforming as one of its reactions, for which limited kinetic data are available in the literature. The parameters for this reaction were chosen from Aoki et al.,73 who express this reaction rate as being first-order in methane partial pressure. (v) Methane decomposition is another possible reaction under high-pressure and high-temperature environments. As discussed previously, methane decomposes into carbon (soot), which is again regasified by steam into CO and H2. Kinetic parameters for methane decomposition are chosen from Kondratiev,74 who studied this reaction under high-pressure (∼20 atm) and high-temperature (∼1300−1500 °C) conditions, those that are of interest to us too. (vi) Three reactions, steam dealkylation, dry reforming, and hydrocracking, are not considered for any of the tar components (benzene, toluene, naphthalene, or phenol). (vii) Although hydrodealkylation is not explicitly F
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considered for naphthalene and phenol, it forms a part of the thermal cracking reactions (see reactions 8 and 10 in Table A4.1 of Appendix 4). (viii) Carbon formation is not explicitly considered for benzene, naphthalene, and phenol; it appears in the form of steam reforming or thermal cracking reactions (see reactions 11, 10, and 9, respectively, in Table A4.1 of Appendix 4). Note that toluene is not associated with any carbon formation because of the lack of kinetic data. (ix) Steam reforming is included in the POX reaction set for two reasons: the feed gas has steam present in it, and steam is formed as a product of the oxidation reactions. On the basis of the components chosen as described in section 3, the syngas feed composition for the base case is as follows: H2, 5.93; CO, 5.37; CO2, 8.93; CH4, 4.64; N2, 6.7; H2O, 64.7; Ar, 3.58; C6H6, 0.15; C7H8, 1.99 × 10−5; C6H5OH, 3.92 × 10−6; C10H8, 1.59 × 10−4 (all in mol %). A total of 816 m3/h of syngas at 20.2 atm and 816 °C completes the base case feed conditions, corresponding to run GT-11 from the IGT gasifier data.52 5.1. Uncatalyzed POX. Table A4.1 of Appendix 4 shows the list of reactions considered in the model, and Table A4.2 of Appendix 4 shows the list of parameters used. The temperature considered for the POX reactor simulations is in the range of 900−1400 °C. Simell et al.68 suggest temperatures of 900 °C and above for dolomite and Ni catalysts to be active for tar decomposition in the gasification gas. Hence, this temperature is chosen as the lower limit for simulating the reforming cases. To keep the same basis for POX and reforming, the same temperature is set as the lower limit for POX cases too. Svensson et al.32 consider temperatures in the range of 1400 °C in their simulations, which is the upper bound that we have considered in our work. Note that some of the reactions are associated with negative exponents in their rate expressions (benzene in reaction 3 and hydrogen in reactions 10 and 11). These can create mathematical problems when the species concentration approaches very low values. Hydrogen, however, is not an issue because it is a major constituent of syngas and is not expected to be consumed fully. These reaction rates are derived under conditions of a H2-rich environment.54 Although its concentration can be quite low, benzene is also not a problem because its oxidation reaction (reaction 3) proceeds at a lower rate than other oxidation reactions (reactions 1, 2, and 5) and is limited by oxygen availability. Reactor simulations are performed using the plug flow kinetic reactor model (RPLUG) in Aspen Plus75 with the selected components and kinetics for the base case feed at a pressure of 20 atm. Oxygen flow to the reactor is adjusted to reach a specified exit reactor temperature (900−1400 °C), and the residence time is fixed at 5 s, typical of values reported in the literature. The results of this sensitivity study are shown in Table 4. It can be seen from the results that the O2 demand increases with the reactor temperature as expected. To have tar conversions of 99% and above, a temperature of 1300 °C and above is necessary. Besides tar conversion, methane conversion also dictates the reactor temperature. Bruggeman et al.31 investigated the influence of the temperature and pressure in a high-pressure reactor for the noncatalytic partial oxidation of natural gas. They observed an increase in methane conversion with the temperature but comment that equilibrium was not reached (conversion lower than the equilibrium level) in the range of parameters studied (reactor exit temperature of ∼1450 °C, residence time between 5 and 20 s, and pressure of 30−100 bar). Although the tar conversion is >99% at 1300 °C in Table 4, the methane conversion is only ∼40%. It improves to ∼84% at 1400 °C and further to ∼93% at the same temperature and an increase in residence time to 10 s (last column of Table 4).
Table 4. Effect of the Exit Temperature on POX Reactor Performancea reactor temperature O2 used (TPH) tar conversion (%) tar concentration (mg/m3)c CH4 conversion (%) H2/CO ratio H2 + CO flow (TPH) soot formed (%)d
1100 °C 1200 °C 1300 °C 1400 °C 1400 °Cb ∼21 ∼1 20700
∼31 ∼26 14300
∼51 >99 2.1
∼69 >99 1.6
∼68 >99 1.6
0.5 1.07 25.1 0.03
5.5 0.99 22.1 0.37
39.7 1.06 31.4 2.36
83.9 1.11 45.9 2.59
92.5 1.11 54.2 1.44
Feed conditions: H2/CO ratio, 1.1; H2 + CO flow, 43.3 TPH; tar concentration, 26 200 mg/m3; and biomass feed, ∼11 TPD. b Residence time is 10 s. cGas volume flow is at streams P and T. Stream P is 20 atm, and stream T corresponds to the reactor temperature. dAs a percentage of incoming carbon. a
The tar concentration in the exit syngas stream is ∼2 mg/m3, which is a reasonable target (99 1.45 81.7 1.21
1250 1043 5 3 ∼39 >99 0.1 >99 1.38 77.4 1.13
900 900 5 3
950 900 5 3
>99 11.0 >99 1.66 99.1 1.13
>99 5.2 >99 1.66 99.1 1.13
Feed conditions: H2/CO ratio, 1.1; H2 + CO flow, 43.3 TPH; tar concentration, 26 200 mg/m3; and biomass feed, ∼11 TPD. bGas volume flow is at streams P and T. Stream P is 20 atm, and stream T corresponds to the reactor temperature. cAs a percentage of incoming carbon. a
specify a tar concentration limit of about 2000 mg/m3 of tars after the dolomite bed to avoid catalyst deactivation in the Ni catalytic bed. Our simulation model also ensures a value close to this limit as the syngas enters the Ni catalyst bed. Note that the tar concentration shown in Table 5 is the final concentration of the exit syngas from the Ni catalytic reactor. Note that there is no additional steam added to the reforming reactors. This is because the inlet feed gas already carries a lot of steam in it. In fact, the steam/carbon ratio of the inlet feed is ∼8, which is a much higher number than that typically used in catalytic reformers (3−5). The total residence time in the reactors is fixed at 8 s in all of the cases in Table 5. Given an exit temperature of 1200 °C after the POX reactor in case ATR-1, it is observed that the final exit temperature after reforming is 934 °C. Both the tar and CH4 conversions are above 99%, and the exit tar concentration is ∼10 mg/m3. To lower it further, case ATR-2 was simulated with slightly more severe (higher temperature) conditions. As expected, the O2 consumption increases with the temperature. However, a part of the useful components in syngas is lost, with a slight decrease in the combined flow rate of H2 and CO (compare columns 2 and 3 in Table 5). Note once again that there is soot H
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Figure 2. Schematic of the absorption flowsheet simulated.
Using diesel, vegetable oil, and biodiesel (with compositions as mentioned previously) and NRTL as the vapor−liquid equilibrium (VLE) model, simulations are performed for the base case feed. The absorber is represented using the stage-bystage distillation model (Radfrac) in Aspen Plus,75 with eight stages and operating at 20 atm. The simulation results in Table 6
benzene is removed from the solvent. Representative conditions for one of the cases simulated are shown in Figure 2. Three solvents are tested in our study for their absorption performance: vegetable oil, biodiesel, and (conventional) diesel. Canola oil is chosen as the vegetable oil and is represented as a mixture of the following pure components: oleic, linoleic, linolenic, palmitic, and stearic acids. The respective weight percentage of these components is 61, 21, 11, 4, and 3 (http://en.wikipedia.org/wiki/ Canola). Biodiesel is typically considered to be a methyl ester of vegetable oil (fatty acids) and is represented in our simulation model as a mixture of methyl oleate, methyl palmitate, and methyl stearate. The respective weight percentage of these components is 74, 14, and 12.77 Mueller et al.78 determined optimal surrogate formulations for diesel fuel by blending eight pure compounds using the following design properties: fuel composition and carbon bond types, ignition quality, volatility, and density. The main hydrocarbon classes used in their surrogate palette are n-alkanes, isoalkanes, cycloalkanes, aromatics, and naphthoaromatics. To keep the number of components limited in the flowsheet, we only chose four of their eight components for our modeling work, which are n-hexadecane (13.6), trans-decalin (43.5), 2,2,4,4,6,8,8-heptamethylnonane (11.8), and 1,2,4-trimethylbenzene (30.8); the numbers in parentheses indicate the weight percentage of each of the components. Another important input needed to study scrubber performance is a suitable phase equilibrium model. Bassil et al.79 use model molecules of tar (benzene, toluene, and phenol) and methyl palmitate (a component of biodiesel) to obtain liquid− liquid equilibrium (LLE) data through experiments and regress for non-random two-liquid (NRTL) binary interaction parameters. They conclude that NRTL works reasonably well and so does universal quasichemical (UNIQUAC) functional-group activity coefficients (UNIFAC). Because the components that we chose in our work are similar to the components used by Bassil et al.,79 NRTL is chosen as the phase equilibrium model in simulating the absorbers. The missing binary interaction parameters were estimated using the UNIFAC method.
Table 6. Comparison of Solvent Performancea tar removed (%) solvent needed (TPH) tar concentration (mg/m3)b solvent loss (%) stripper heat duty (MW) solvent needed (m3/h)
diesel
vegetable oil
biodiesel
99 116 2520 1.11 25 147
99 194 2460 0.002 46 226
99 197 2500 0.001 42 244
a Feed conditions: H2/CO ratio, 1.1; H2 + CO flow, 43.3 TPH; tar concentration, 26 200 mg/m3; and biomass feed, ∼11 TPD. bGas volume flow is at stream conditions (see Figure 2).
show that the performance order is diesel > vegetable oil > biodiesel. The same conclusions were shown by Phuphuakrat et al.39 in their experimental study with the same solvents. However, their experiments were with a constant volume of solvent and at atmospheric pressure. The exact absorption temperature in their experiments is not known. Note that methane goes through the scrubbing process with almost no absorption in the solvents. The combined flow of H2 and CO also remains unaltered at its feed condition (∼42 TPH), except for a marginal decrease, owing to the solubility effect. The flow rates shown in Table 6 are for the case without recycle. In practice, the solvent is recycled from the stripper to the absorber, and the makeup solvent rate is expected to be much less than the solvent circulation rate. Reboiled stripper for solvent recycle shows promise and is the option considered for solvent regeneration in our case. Stripping agents, such as N2, fuel gas, or off gas, were not used because they need to be heated before using in the stripper and their usage leads I
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6.2. Catalytic Steam Reforming. The foremost challenge with this option is the absence of a proven commercial catalyst that exhibits high conversion for both tars and methane, without considerable deactivation. Most of the catalysts discussed in the literature are at bench- or pilot-scale testing only. The inlet gas to the reformer may be laden with particles (char, sand, in-bed catalyst fines elutriated from the bed, etc.), and hence, a dust filter operating at high temperature is necessary for their removal to avoid catalyst deactivation. Similar to the POX reactor, hightemperature operation in this case also requires the use of expensive alloys and special metallurgical considerations in designing the reactor. Soot formation needs to be kept in mind, and a judicious choice of operating parameters (reactor temperature, inlet steam flow, etc.) is necessary, so that catalyst deactivation is at a minimum. Most of the catalytic reforming work reported in the literature is performed in reactors operating at or close to atmospheric pressure. However, we have considered high-pressure operation in our work assuming that the biomass gasifier operates at a high pressure too. Morris80 cautions that the use of dolomite as a catalyst in highpressure systems may be less effective, owing to its decalcining nature, and there exists a minimum operating temperature dictated by the pressure and possibly the CO2 level in the feed (e.g., corresponding decalcination temperature at 25 bar is in the range of 950−1000 °C). Parent et al.81 describe a process to synthesize attrition-resistant fluidizable reforming catalysts that can be used for reforming tars in a fluidized-bed environment. Apanel and Wright82 discuss the operation of such a fluidized-bed gasifier with a catalyst inside it. Such patent literature suggests progress being made in the field of catalysts suitable for tar reforming. 6.3. Absorption. The main challenge associated with this method is its inability to recover methane from the inlet syngas. While this does not hinder applications of such a syngas for combustion applications [such as in an engine or combined heat and power (CHP) plant], it is a problem for fuel or chemical synthesis applications. In most of the reactors involved in fuel or chemical synthesis, methane acts as an inert and needs to be purged or reformed to keep its concentration to a permissible limit in the reactor feed; this calls for additional modifications in the process, which might prove to be expensive. The possibility of tar condensation before the scrubber, treatment of the water removed by syngas cooling, and the levels of residual tar in the syngas are issues that cannot be overlooked and need to be addressed. The solvent requirement, as shown in section 5.3, can be quite high to reach very high levels of purity in the syngas. Thus, the economics involving solvent cost and its regeneration need to be considered in making a decision. Note also that the syngas leaving the absorber is ∼80 °C and needs to be reheated before going to the shift reactors, a typical process downstream of the tar removal operation in most fuel production processes. The POX and catalytic reforming cases result in a hot syngas, which needs cooling to remove any dust or entrained solids in a filter before going to the shift reactors. The gas temperature after cooling is dictated by the temperature that the filter material can withstand and is typically in the range of 350− 550 °C. Thus, the absorption option has an energy penalty in terms of syngas reheating compared to the POX and reforming cases. 6.4. Other Impurities. Non-woody biomass has impurities, such as chlorine, sulfur, nitrogen, and ash, that are in higher quantities compared to woody biomass. To handle these contaminants, one may need additional gas-cleaning measures to avoid emission or corrosion issues, downstream catalyst deactivation, etc. For example, in-bed dolomite can become deactivated because of the presence of chlorine impurity in biomass
to a dilution of the tail gas stream. With the reboiled stripper, no additional gases are added and only those dissolved from the initial syngas are released. Owing to the presence of relatively higher volatile components in diesel, we observe a higher solvent loss in its case compared to vegetable oil and biodiesel. To conclude, scrubbing syngas with one of the solvents used in this study helps to lower the tar concentration. However, to lower it to the extent needed for fuel synthesis applications, the solvent requirement is quite high, as shown next. As seen in Table 6, the exit tar concentration from the absorption system is ∼2500 mg/m3. Note that this value is high, despite 99% tar removal specification, because the stream flow rate changes considerably as a result of water removal. This limit needs to be decreased further to mg/m3 (
