Energy Fuels 2011, 25, 656–669 Published on Web 01/19/2011
: DOI:10.1021/ef101336m
Thermodynamic Analysis of Alternative Approaches to Chemical Looping Combustion V. Kalyana Chakravarthy, C. Stuart Daw,* and Josh A. Pihl Energy & Transportation Science Division, Oak Ridge National Laboratory, 2360 Cherahala Boulevard, Knoxville, Tennesee 37932 Received September 30, 2010. Revised Manuscript Received December 21, 2010
In this article, we review and clarify some of the points made by previous authors1,2 regarding chemical looping combustion (CLC). Although much of the recent interest in chemical looping combustion has been associated with carbon sequestration, our primary interest here is its potential to increase the thermodynamic efficiency of converting fuel chemical energy into useful work. We expand on several points about the details of CLC that we feel have not previously been sufficiently explored and suggest alternative (and possibly more practical) approaches that exploit some of the same thermodynamic concepts. We illustrate our key points with first and second law analyses of ideal conceptual processes, which, in addition to CLC, also include isothermal, nonequilibrium, preheated combustion and combustion with thermochemical recuperation. Our results suggest that a significant portion of the potential efficiency benefit of CLC might be achieved without the need to handle and transport large quantities of solid oxygen-storage material. Exploitation of this fact may lead to approaches for power generation from hydrocarbon fuel combustion that can achieve second law efficiencies 10-15% higher than those that are currently possible.
Previous studies have emphasized that the expected efficiency advantage of CLC (relative to conventional unrestrained combustion) is a consequence of constraining the oxidation and reduction reactions to separate gas-solid reaction stages that operate near chemical equilibrium, resulting in lower exergy destruction and higher efficiency.1,2,6,8 Other approaches have been proposed for constraining gasphase reactions to near equilibrium conditions also, but the direct gas-phase reactions require very high temperatures to achieve equilibrium (e.g., typically above 3500 K2,10,11). Because the gas-solid reactions in CLC have much lower equilibrium reaction temperatures, it offers the potential to avoid high temperatures while still reducing exergy destruction. One important aspect that has not always been strongly emphasized in the CLC literature is that thermal equilibrium is also a critical factor. That is, all heat-transfer processes need to occur over very small temperature gradients to avoid generation of entropy and destruction of exergy.10-12 In practice, this means that all gas-gas, solid-solid, or gassolid heat transfer between flowing streams should be counterflow and must be thermally balanced so that temperature pinch points or large gradients do not develop. This places significant constraints on the design of CLC systems and is an extremely important practical consideration. In the following sections, we discuss this issue at length and its relationship to the idea of isothermal combustion where chemical equilibrium may or may not be present. Two other important features of CLC are (1) the generation of two separate exhaust streams of different composition and temperature and (2) the presence of an internal heat flow that can be used to generate additional work. The first feature is exploited in carbon sequestration, but, as far as we are aware, the precise role of this exhaust partitioning on entropy
Introduction Chemical looping combustion (CLC) is a two-stage process of burning fuel.1 The first stage involves oxidation of a solid material using oxygen in the air, and in the second stage, it is reduced back to its original state by a fuel. To complete the process, the reduced oxygen carrier is recycled back to the oxidation stage where it can be reoxidized. The combustion gases leaving the second stage are not diluted with atmospheric nitrogen, and so, they are more concentrated in CO2 than conventional flue gas. In the current day, much of the interest in CLC is driven by its value as a key step in carbon sequestration rather than for its direct benefit to fuel exergy preservation. Analyses of the exergy budget in CLC-based power generation plants have been performed in a few papers.3-9 Most such studies have typically been directed at very specific and complex applications, and thus, they have not been so concerned with clarifying the fundamental thermodynamic impact of CLC relative to conventional combustion. One recent exception to this is the study by McGlashan,2 which provides the kind of basic thermodynamic analysis of CLC that we have found useful as a starting point. In our analyses, we expand on points briefly noted by McGlashan and consider alternative ideal processes that capture some key CLC features. We also point out practical problems that could arise in implementation of these alternative processes. *To whom correspondence should be addressed. E-mail: dawcs@ ornl.gov. (1) Richter, H. J.; Knoche, K. F. ACS Symp. Ser. 1983, 235, 71–85. (2) McGlashan, N. R. Proc. Inst. Mech. Eng., Part C 2008, 222, 1005– 1019. (3) Ishida, M.; Zheng, D.; Akehata, T. Energy 1987, 12, 147–154. (4) Jin, H.; Ishida, M. Energy 1993, 18, 615–625. (5) Ishida, M.; Jin, H. Energy 1994, 19, 415–422. (6) Anheden, M.; Svedberg, G. Energy Convers. Manage. 1998, 39, 1967–1980. (7) Jin, H.; Ishida, M. Int. J. Hydrogen Energy 2000, 25, 1209–1215. (8) Brandvoll, O.; Bolland, O. J. Eng. Gas Turbines Power 2004, 126, 316–321. (9) Wolf, J.; M. Anheden, J. Y. Fuel 2005, 84, 993–1006. r 2011 American Chemical Society
(10) Lutz, A.; Larson, R. S.; Keller, J. O. Int. J. Hydrogen Energy 2002, 27, 1103–1111. (11) Daw, C. S.; Chakravarthy, K.; Conklin, J. C.; Graves, R. L. Int. J. Hydrogen Energy 2006, 31, 728–736. (12) Dunbar, W. R.; Lior, N. Combust. Sci. Technol. 1994, 103, 41–61.
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generation and exergy destruction has not been fully quantified. We include an explicit discussion of this issue in the results presented here. We also point out how the internal heat-flow feature of CLC can be mimicked by other processes, such as thermochemical recuperation (TCR), in which hydrocarbon fuels are first reformed to syngas using exhaust heat prior to being combusted. Taken together, we feel that the above issues should be important considerations in future discussions about CLC and related concepts. Thermodynamic Analysis Work Extraction Calculations. Two kinds of work extraction with heat engines are considered here. The first is work that can be extracted using the heat available from an isothermal reactor, which is limited by the Carnot efficiency. The second is work that is continuously extracted from hot flue gas (produced by combustion) as it isobarically cools to ambient conditions, which is limited by thermodynamic exergy. The Carnot limit and the concept of exergy (for isobaric systems) can easily be shown to be consistent with each other as follows. Assume infinitesimal heat dQ is drawn from hot flue gas at temperature T, which results in its cooling to temperature (T - dT). In the limit of dT being small, heat dQ can be assumed to be drawn at a constant temperature T. The maximum amount of work, dW, that can be drawn from dQ can be determined using the Carnot efficiency.
Figure 1. Performance of an ideal preheated, isothermal, stoichiometric gas-phase combustion of H2 at atmospheric pressure with no heat losses. (a) Exergy budget nondimensionalized by initial exergy (the fraction of exergy that remains after unrestrained stoichiometric combustion is indicated using a symbol for comparison at the adiabatic flame temperature); second law efficiency associated with adiabatic, isobaric combustion is indicated by the symbol at the adiabatic flame temperature. (b) Temperature of the combustion products at the exit of the preheater.
The maximum work that can be extracted using heat drawn from an isothermal reaction, at temperature T, is easily computed using the Carnot efficiency. T0 ð2Þ Wiso ¼ - ΔH 1 T
Cp T0 T0 ¼ Cp 1 dT ¼ Cp dT - T0 dT ¼ dH - T0 dS dW ¼ dQ 1 T T T
ð1Þ dH and dS in the above equation are the differential changes in enthalpy (H) and entropy (S) associated with temperature change dT, which, respectively, equal CpdT and Cp(dT/T). Integrating the above equation from T to T0 (ambient temperature) results in the maximum amount of work that can be extracted from hot flue gas using a series of Carnot engines. Integration of (dH - T0dS) from temperature T to ambient temperature T0 equals the exergy of the mixture, [(H - H0) - T0(S - S0)], at atmospheric pressure, where H0 and S0 are the enthalpy and entropy of the mixture at the reference dead state. Preheated, Isothermal Gas-Phase Combustion. We begin our discussion with preheated, isothermal gas-phase combustion, which is the simplest concept analyzed here. In conventional (unrestrained) combustion, reactants enter the combustion chamber at ambient temperature, whereas the products leave at the adiabatic flame temperature. Given that most fuels have ignition temperatures much higher than ambient, the combustion can be sustained only if there is some heat transfer from already formed hot products to reactants. This exchange is driven by a large temperature difference that results in high entropy generation and loss of exergy.12 The gas-phase reaction between fuel and air can theoretically be maintained at isothermal conditions by transferring the heat generated by the combustion reactions to a heat engine and using the combustion products to preheat the reactants to the operating temperature. For a simplified analysis, it is assumed that reactants can be preheated to the reaction temperature using hot reaction products in compliance with both first and second laws. This enables comparison with CLC, which is analyzed along similar lines below.
ΔH is the enthalpy change associated with the reaction. This expression implies that work increases with reaction temperature and is limited by the enthalpy change of combustion and not exergy. This is problematic from a second law perspective if the latter is lower. In such cases, however, there is an upper limit on operating temperature. For T > ΔH/ΔS, the conversion of reactants to products cannot proceed spontaneously because the Gibb’s free energy change is positive. At this limiting operating temperature, Wiso can be shown to equal the exergy of the mixture. The benefits of isothermal combustion over conventional unrestrained combustion are due to the different nature of the heat exchange between products and reactants in the two processes. In the former, heat is exchanged between hot products and reactants in a (counterflow type) preheater where the heat exchange (at any location) is driven by a small temperature gradient between the two streams. Because heat is constantly extracted to maintain the reactions at an isothermal state, there is no heat exchange between reactants and products when they are physically in contact in the reactor. Nevertheless, there is some entropy generation due to composition change, which can only be avoided at the equilibrium temperature of the overall fuel-air reaction (if one exists).2,10,11 We also analyzed preheated, gas-phase isothermal combustion considering losses associated with preheating while using temperature-dependent thermodynamic properties in an effort to understand the extent of deviations from ideality represented by eq 2. Figure 1 shows the theoretical second law efficiency of this process for H2 fuel. 657
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The exergy budget for this process has four components: work extracted from the heat engine, combustion irreversibility (which also includes irreversibility due to mixing), preheating irreversibility, and exergy in the exhaust. The results for this plot are generated as follows. A reaction temperature is chosen, and the equilibrium composition of the products is determined. The product stream is used to preheat reactants to the maximum temperature possible. When the products are not hot enough to preheat reactants all the way to the reaction temperature, the reactants are first assumed to react partially so that the temperature rises to the chosen reaction temperature before heat is drawn to a heat engine. There is minimal exergy loss due to preheating in this scenario. If there is more heat in the hot products than what is needed to preheat the reactants to the required reaction temperature, then the products exit the heat exchanger at higher than ambient temperature. This starts to happen as the chosen reaction temperature starts to exceed the adiabatic flame temperature of H2. Instead of neglecting the exhaust heat, the exergy of the exhaust is added to the work that is extracted from the heat engine (deriving heat from reaction between preheated reactants at chosen reaction temperature). There is, however, exergy loss associated with heat transfer at the colder end of the heat exchanger in such cases. Its increase with temperature is rapid enough beyond this point that useful exergy (work and exhaust exergy combined) decreases with temperature despite decreasing combustion irreversibility, and the 100% second law efficiency at equilibrium temperature, as suggested by eq 2 and the Gibbs free energy constraint earlier, is never reached. The same trend of highest efficiency when the reaction temperature is near the adiabatic flame temperature is observed for other fuels. Preheated, isothermal gas-phase combustion at best has about half the combustion irreversibility associated with unrestrained gas-phase combustion between reactants. This is a valid comparison given that the peak efficiency of the former is roughly around the adiabatic flame temperature, and so, the peak operating temperatures in both cases are comparable. It should be noted that implementing isothermal gasphase combustion could be extremely difficult. At temperatures below the adiabatic flame temperature, gas-phase combustion reactions are highly irreversible and spontaneous. It is difficult to match the rate of heat extraction to the rate of heat generation from highly spontaneous gasphase reactions. Conventional Chemical Looping Combustion. To help clarify the reasons behind the efficiency benefits of CLC, we discuss two specific features that set it apart from conventional combustion. First, we review the importance of generating two separate exhaust streams rather than a single mixed exhaust. We then consider the impact of how the chemical reactions and internal heat transfer are managed. Adiabatic CLC. The most distinctive global feature of CLC highlighted in recent studies is that it generates two separate gaseous exhaust streams, one containing N2 (and O2 under lean conditions) and the other containing H2O and CO2, as depicted in Figure 2. The two streams emerge separately from staged oxidation and reduction reactors. A solid oxygen carrier (typically some type of metal) shuttles between the two reactors in reduced (Me) and oxidized (MeO) forms. When no heat flows between the overall system and the environment, the enthalpy of the fuel and air mixture is split
Figure 2. Simplified schematic of CLC.
Figure 3. Second law combustion efficiency (ratio of exergy after combustion to initial exergy of unmixed fuel and air at ambient conditions) and temperature of outlet streams for the stoichiometric simplified (flue-gas generating) version of chemical looping combustion of H2. The oxidation reactor exhausts only N2, whereas the reduction reactor exhausts H2O.
between the two product streams. The temperatures of the product streams can be different because the heats of reaction for oxidation and reduction can be different. The first law requires only that the net enthalpy of the reactants equals the net enthalpy of two exhaust streams at steady state. We refer to this generic case as adiabatic CLC. As we discuss below, the separation of the exhaust gases reflects an increase in the final exergy that is destroyed in conventional combustion when the exhaust gases are allowed to mix. Without initially defining the details of how each reactor operates, it is informative to consider just the second law impact of exhaust gas separation for a range of possible exhaust temperatures from the two stages. For example, with stoichiometric H2 combustion, the product stream from the oxidation reactor only has N2 at temperature TN2, whereas the product stream from the reduction reactor contains only H2O at a temperature of TH2O. For TH2O ranging from 500 to 3000 K, we computed values of TN2 so that the sum of enthalpies of products equals that of the reactant streams (H2 and air). The sum of the exergies of the two products divided by the initial fuel-air exergy equals the second law efficiency of the process. In Figure 3, the second law efficiency and TN2 are plotted for a range of values of TH2O. The second law efficiency of conventional (isobaric gas-phase combustion resulting in a single mixed product stream at the adiabatic flame temperature) is also shown for comparison. As seen in the above figures, the efficiency benefit from exhaust separation is minimal when the two streams have the same temperature (which equals the adiabatic flame temperature). Even when the exhaust streams are at the same temperature, however, there is also a slight benefit remaining because their compositions are different. At least theoretically, it is possible to exploit this concentration difference to produce work with a concentration engine. The second law efficiency increases as the temperatures of the two streams diverge, which is expected. This is because 658
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Figure 4. Second law combustion efficiency and temperature of outlet streams for lean (equivalence ratio of 0.5), simplified chemical looping combustion of H2. Because the combustion is lean, only half the oxygen is consumed and the rest remains in the exhaust coming out of the oxidation reactor.
Figure 5. Schematic of CLC for exothermic oxidation and endothermic reduction reactions. An internal engine generates work by extracting heat from the oxidation reactor and rejecting some of it to the reduction reactor. Some additional heat needs to be extracted from the oxidation reactor by an external heat engine in order to maintain the isothermal oxidation reaction.
some work can be generated by operating a heat engine between the two streams in addition to the work that can be generated by operating a heat engine between the mixture of two streams (when they are combined) and the ambient. More of this additional work can be generated if there are two streams at two different temperatures Even in extreme cases with high temperature differences where peak temperatures exceed 3000 K, the second law efficiency only increases by about 7%, so exhaust gas separation can only account for a relatively modest improvement by itself. Exhaust gas temperature differences are also constrained by materials considerations. As in conventional combustion, materials temperature limits favor the use of lean combustion (high excess oxygen) to keep peak temperatures low. One potential benefit of chemical looping in this regard is that it enables stable combustion below the conventional lean limit for gas-phase combustion. To consider how the above estimates of second law efficiency for adiabatic CLC are affected by lean operation, we also computed them for a global equivalence ratio of 0.5. These results are depicted in Figure 4. Even in this case, the second law efficiency benefit from the exhaust gas separation is only a few percent as long as the peak temperature is restricted to 2000 K. Thus, it is clear that exhaust gas separation alone does not provide a major efficiency benefit in and of itself. If CLC is to be pursued for efficiency reasons, there are other important factors to consider. Classical CLC. Two additional constraints usually associated with CLC are the isothermal and near-equilibrium operation of each reactor. Analysis of these factors requires consideration of individual heat balances for each reaction stage. To be consistent with earlier analysis,2 we assume here temperature-independent specific heats. ΔH1 and ΔS1 represent the enthalpy and entropy changes associated with the oxidation reaction in CLC, whereas ΔH2 and ΔS2 represent those of the reduction reaction. Obviously, ΔH1 þ ΔH2 equals ΔH( T2/T1, does not seem possible to meet with any combination of oxygen carrier and fuel among several analyzed in this study for reasonable values of T2. This is because T2 needs to be high enough for the reduction reaction to be kinetically feasible (there are several material-fuel combinations where the reduction reaction is thermodynamically feasible at subambient temperatures). For reasonable values of T2 (around 600 K), no reduction reaction we found is endothermic enough (i.e., has a high enough Q2 to absorb all of the heat necessary to eliminate the discharge of heat to the environment). Further, the work output from the engine (W = Q1 - Q2) equals -ΔH, the enthalpy change of the overall fuel-air reaction, implying 100% first law efficiency. This is thermodynamically possible only if the enthalpy change of the overall fuel-air combustion is lower than the exergy of the fuel-air mixture. The condition, Q2/Q1 > T2/ T1, if possible, would have to be met with fuels with this feature. Given that there are more practical limitations that supersede this condition (a discussion of which is included later), it is not necessary here to rule out the possibility that this condition could be met. When the above condition is not met, two heat engines, one operating between the two reactors and the other operating between the oxidation reactor and ambient, as shown in Figure 5, are needed to extract the maximum possible amount of work. It is assumed that T01 and T02 in the schematic, which are, respectively, the temperatures of preheated air and fuel streams, equal T1 and T2 in this simplified analysis to be consistent with the analysis of 659
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McGlashan. In reality, however, T01 is lower than T1 because N2 at temperature T1 cannot be used to preheat air, which contains the same amount of N2 and additional O2 to the same temperature. The Carnot efficiencies of the two engines are [1 - (T2/T1)] and [1 - (T0/T1)], respectively. The work extracted from the two engines and the total work can be written as follows. T1 T2 ð3Þ 1Wint ¼ Q2 T2 T1 2
T1 T0 Wext ¼ Q1 - Q2 1T2 T1 W ¼ Wint þ Wext
¼ ðQ1 - Q2 Þ - T0 ¼ - ΔH þ T0
Q1 Q2 T1 T2
ΔH1 ΔH2 þ T1 T2
ð4Þ
Figure 6. Schematic of CLC for both exothermic reactions. Each heat engine extracts heat from a reactor and rejects some of it to ambient.
ð5Þ
obviously, is limited by Teq,1, and this condition used in the above expression results in the following inequality.
Given that ΔH1 e 0 and ΔH2 g 0, the efficiency increases with increasing T1 and decreasing T2. Thermodynamic feasibility conditions (negative free energy changes for the reactions) in addition to the previously noted facts that ΔS1 e 0 and ΔS2 g 0 imply that T1 e Teq,1 and T2 g Teq,2 (where Teq,1 and Teq,2 are equilibrium temperatures defined by McGlashan et al.2) for sustained operation of the reactors. ΔHi Teq;i ¼ ; i ¼ 1;2 ð6Þ ΔSi
W e - ΔH þ T0 ΔS1 þ T0
e - ΔH þ T0 ðΔS - ΔS2 Þ þ T0 e - ΔH þ T0 ΔS þ
T0 T0 - ΔH2 1 W ¼ - ΔH1 1 T1 T2
ΔH1 ΔH2 þ T1 T2
ΔH2 T2
T0 ðΔH2 - T2 ΔS2 Þ ð8Þ T2
Once again, Teq,2 imposes a lower limit on T2 to ensure thermodynamic feasibility. However, because ΔH2 < 0 and ΔS2 > 0, the calculated Teq,2 would be negative, leading to two conclusions: there is no attainable equilibrium temperature for the exothermic reduction reaction, and the reduction reaction should occur spontaneously at all temperatures. There does not seem to be an obvious thermodynamic upper limit on T2, and efficiency increases monotonically with increasing T2. T2 could theoretically be higher than T1. It is worth noting that the overall efficiency is actually increasing as T2 is moving away from the “equilibrium reduction’’ temperature. In reality, T2 is limited by the melting points of the two solid materials involved. Even without these limitations, the total work is limited by fuel-air exergy (-ΔH þ T0ΔS) in compliance with the second law given that the Gibbs free energy change for the reduction reaction ΔH2 - T2ΔS2 e 0 for all values of T2. In summary, it seems thermodynamically possible to achieve 100% second law efficiency using CLC for combinations of fuel and oxygen-storage material that have exothermic oxidation reactions and endothermic reduction reactions. For combinations where both reactions are exothermic, the second law efficiency is limited by the maximum possible temperature where reduction reactions can be sustained. The overall efficiency for such cases, surprisingly, increases as the reduction reaction becomes more irreversible (i.e., T2 moves away from Teq,2). If T1 and T2 in eqs 5 and 7 are assumed to be equal, then the expression for total work for either endothermic or exothermic reduction reactions becomes identical to the expression for preheated, isothermal, gas-phase combustion given by eq 2. Of course, there are different constraints on operating temperatures for the two kinds of combustion (as discussed above).
The maximum efficiency is, therefore, attained when T1 = Teq,1 and T2 = Teq,2, the choices suggested by McGlashan et al.2 The maximum work equals -ΔH þ T0ΔS, the exergy of the fuel-air mixture, and so the second law is never violated. Note that operation of the idealized CLC system at the oxidation and reduction equilibrium temperatures would result in 100% second law efficiency. For the fuel-oxygen carrier combination with exothermic oxidation and reduction reactions, two heat engines are needed to maximize work extraction. A Carnot engine operates between each reactor and the ambient, as illustrated in Figure 6. As before, T01 and T02 are assumed to equal T1 and T2, respectively. The total work extracted is determined by the following expression.
¼ - ΔH þ T0
ΔH2 T2
ð7Þ
Alternately, a combination of an internal Carnot engine drawing heat from the oxidation reactor and rejecting heat to the reduction reactor and an external Carnot engine drawing heat from the reduction reactor and rejecting heat to the ambient could be used. The above expression for total work can easily be shown to be applicable in this instance as well. It is also worth noting that this expression is exactly the same as the one derived for the combination of exothermic oxidation and endothermic reduction. In this case, however, work output increases with an increase in either T1 or T2. T1, 660
Energy Fuels 2011, 25, 656–669
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Preheating also is different, as there are two exhaust streams and two inlet streams in CLC. Therefore, the differences between the two concepts stem primarily from constraints on operating temperatures2 and implementation issues. As mentioned above, isothermal reactions can be hard to control in preheated, gas-phase combustion. This is not a problem in CLC because the oxidation reaction stops once the reactor temperature exceeds Teq,1 and will resume only after heat is extracted to lower the temperature, so there is a natural barrier against thermal runaway in CLC. Reduction reactions in CLC are spontaneous, but the heat generated/ consumed by these reactions is relatively low for most fuel-oxygen-storage material combinations. A simplified analysis with constant specific heats of CLC and preheated, isothermal, gas-phase combustion has resulted in eqs 5, 7, and 2 for work outputs. A theoretical comparison can be made of the two kinds of combustion for a given fuel if an oxygen carrier is chosen for CLC. These expressions, however, do not reflect the performance of realworld implementations of these idealized combustion processes. The analysis, for example, does not take into account preheating of the reactants. Minimizing the exergy losses associated with preheating is a challenge for both CLC and preheated, isothermal, gas-phase combustion. The transport of solids required in CLC introduces an even more difficult challenge. To achieve the isothermal operation assumed above, the circulating solids must enter the reactor at its operating temperature. Unless the reactors are operating at the same temperature, the solid streams circulating between the reactors must achieve complete heat exchange with minimal losses. That is, the solids exiting the high-temperature reactor must transfer enough heat to the other solid stream to cool to the operating temperature of the low-temperature reactor. MeOðT1 Þ þ MeðT2 Þ w MeOðT2 Þ þ MeðT1 Þ
mMe is the moles of Me involved in CLC. The thermal capacity of the carrier solids (mMeCp) is usually much higher than the thermal capacity of the gaseous reactants (when all materials are in stoichiometric proportions). As a result, CLC without optimal solids heat management ends up having low efficiency when T1 and T2 are very different. The problem of solids heat exchange can be avoided if reactor temperatures are nearly equal, which is possible in cases where both oxidation and reduction are exothermic, as discussed later. Alternative to CLC: Staged Combustion with Oxygen Transfer (SCOT). In addition to the gas and solids preheating issues described above, CLC requires a mechanism for solid circulation between the reactors. This difficulty alone precludes its use in nonstationary applications. It also introduces substantial cost and complexity into the system relative to other combustion approaches. A version of chemical looping combustion with stationary solid reactants, which mitigates some of the problems associated with CLC, was proposed and analyzed by Noorman et al.15,16 A similar concept, referred to as staged combustion with oxygen transfer (SCOT) to distinguish it from conventional CLC, is analyzed here. Unlike in CLC, the solid reactants in SCOT are stationary, and there is no solid-tosolid heat exchange. SCOT utilizes two reactors operating in opposite phases (one is under oxidation while the other is under reduction). Once the solid reactants have been fully converted to products, the inlet gas compositions are switched between air and fuel to change the operating phase of the reactors. Unlike the analyses performed in the previous sections, temperature-dependent thermodynamic properties and preheating inefficiencies are included in the analysis of this new concept. We also revisit preheated, isothermal, gas-phase combustion using this more general approach to provide a basis for comparison with SCOT. Estimates of maximum work that can be extracted for several operating scenarios are provided below. The calculations of these estimates are done as described below. The oxidation and reduction reactions are assumed to occur under (mostly) isothermal conditions at temperatures T1 and T2, respectively. When reactants are in stoichiometric proportions, N2 is the only gaseous product of the oxidation reactor. With the N2 at temperature T1, it will not be possible to completely preheat ambient air (which contains O2 in addition to the same amount of N2 as in the product stream) to the required oxidation reaction temperature (T1). The enthalpy of the CO2 and H2O exhaust from the reduction reactor is generally enough to preheat fuel to the reduction reaction temperature. However, excess heat in the fuel preheater cannot easily be used to make up for the deficit in the air preheater because typically T2 < T1. Even if the two temperatures are equal, designs to do this routing can be very complex. Here, we consider simple heat-exchange processes with outgoing N2 from the oxidation reactor at T1 preheating the air and CO2 and H2O from the reduction reactor at T2
ð9Þ
Implementing such a system would require identical heat capacities of the solid materials, counter-flow heat exchange, and possibly an intermediate fluid to transfer the heat between the streams. Realizing such a system would clearly involve many practical hurdles. Possibly because of these hurdles, many implementations of CLC discussed in the literature do not appear to include low-gradient heat exchange for the carrier solids. Instead, the solids enter each reactor at a temperature far different from the equilibrium reaction temperature. In cases of highly endothermic reduction reactions, the resulting intrareactor temperature swings can be quite high (as the reduction reactions consume large quantities of heat), and so, a highly endothermic reduction reaction is considered undesirable.13,14 On the other hand, minimization of the endothermic stage reaction heat would appear to be counterproductive for efficiency based on the above analyses and the work of McGlashan.2 If the solid-solid heat-exchange process is not managed in an optimal fashion, there is considerable generation of entropy associated with the high-gradient heat transfer. In the limit, the two solids streams simply equilibrate to their thermal capacity weighted average temperature. T1 þ T2 T1 þ T2 þ Me ð10Þ MeOðT1 Þ þ MeðT2 Þ w MeO 2 2
(13) Abad, A.; Mattisson, T.; Lyngfelt, A.; Johansson, M. Fuel 2007, 86, 1021–1035. (14) Zafar, Q.; Abad, A.; Mattisson, T.; Gevert, B.; Strand, M. Chem. Eng. Sci. 2007, 62, 6556–6567. (15) Noorman, S.; van Sint Annaland, M.; Kuipers, J. Ind. Eng. Chem. Res. 2007, 46, 4212–4220. (16) Noorman, S.; van Sint Annaland, M.; Kuipers, J. Chem. Eng. Sci. 2010, 65, 92–97.
The lost work potential can be determined easily as [mMeCp(T1 - T2)(T0/T2 - T0/T1)/2] per mole of fuel, where 661
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ambient. The net work extracted can simply be estimated as follows. T0 W1 ¼ Q1 1 ð13Þ T1
preheating the fuel. The reaction schemes in the two reactors can be symbolically denoted as follows. airðT10 Þ þ MeðT2 Þ w N2 ðT1 Þ þ MeOðT1 Þ þ Q1 ðT1 Þ
ð11Þ
T0 W2 ¼ Q2 1 T2
fuelðT20 Þ þ MeOðT1 Þ w CO2 ðT2 Þ þ H2 OðT2 Þ þ MeðT2 Þ þ Q2 ðT2 Þ
ð12Þ
W ¼ W1 þ W2 T0 T0 þ Q2 1 ¼ Q1 1 T1 T2
T01 and T02 are the temperatures of the air and fuel as they enter the oxidation and reduction reactors, respectively, and, as explained above, may not equal T1 and T2. Q1 and Q2 denote the heat extracted from or added to the oxidation and reduction reactors, respectively. The signs of Q1 and Q2 do not necessarily indicate whether the oxidation and reduction reactions are exothermic or endothermic, as explained below. It is also worth noting that Q1 and Q2 do not necessarily add up to -ΔH because T01 and T02 do not equal T1 and T2. At the end of each stage of combustion, the solid material is at a temperature that is different from what it should be for the next stage. Therefore, neither the gasphase nor the solid-phase reactants are typically at the target operating temperatures at the beginning of the combustion stage. This complication is handled using the following approach. The oxidation reaction is assumed to proceed at temperatures below T1 until enough heat is generated to raise the temperature of the reactor to T1. When the reactor reaches temperature T1, heat is drawn continuously from it to maintain the reaction under isothermal conditions until the reaction is complete. If the solids and gases in the oxidation reactor begin with temperatures much lower than the chosen T1, the heat of the oxidation reaction alone may not be sufficient to raise the temperature of the reactor to T1 and additional heat may be necessary; that is, Q1 is negative. Because T1 > T2, this heat cannot be provided from within the system, so an external heat pump is assumed to supply the heat required to achieve temperature T1. The external heat pump uses work generated from the reduction reactor (Q2 is necessarily positive in such cases). Alternately, Q2 is sometimes positive even if the reduction reaction is endothermic (it is always positive for an exothermic reduction reaction) because the solid-phase reactant starts out at a much higher temperature than T2. Therefore, heat has to be drawn to lower the temperature to T2 (and maintain it there). For a conservative estimation of work that can be extracted, we assume that all of the heat is extracted at T2. If the reduction reaction is sufficiently endothermic to drop the reactor temperature to T2 at some point during the reduction phase, heat is subsequently supplied at T2 using an internal engine operating between the two reactors to maintain isothermal conditions. In cases with endothermic reduction, the decision to add or remove heat has to be made (based on the sign of Q2) prior to the beginning of the combustion phase. This precludes the possibility of drawing heat out while the reaction temperature is above T2 and adding heat when it drops below. Once Q1 and Q2 are calculated, the following method is used to compute the net amount of work that can be generated. If both Q1 and Q2 are positive, then we assume that each reactor is supplying heat at its nominal operating temperature to a heat engine that is rejecting heat to the
ð14Þ
ð15Þ
The above expression suggests that higher values of T1 and T2 lead to higher efficiency. However, it is to be noted that Q1 and Q2 in SCOT are not independent of T1 and T2, as evident from eqs 11 and 12. When T1 and T2 are very different, the oxidation reaction starts with Me at T2, and no heat is extracted until the reactor temperature reaches T1. The heating of the solids during this process leads to irreversibility that reduces the second law efficiency of the overall process. Efficiency is, therefore, expected to increase as T2 is taken closer to T1 when both Q1 and Q2 are positive. If Q1 is positive and Q2 is negative, an internal heat engine is used to supply heat to the reduction reactor just as in the case of traditional CLC. Work generated by this engine is estimated assuming Carnot efficiency. T1 T2 ð16Þ 1Wint ¼ - Q2 T2 T1 The heat drawn by the internal heat engine is limited by the ability of the reduction reactor to accept the heat by this engine while maintaining isothermal conditions. Any heat not drawn from the oxidation reactor by the internal heat engine is used to generate work using an external heat engine that rejects heat to ambient. The heat left over to power this external engine equals [Q1 þ Q2(T1/T2)], and the work it generates is estimated using its Carnot efficiency [1 - (T0/ T1)]. T1 T0 1ð17Þ Wext ¼ Q1 þ Q2 T2 T1 T0 T0 Wint þ Wext ¼ Q1 1 þ Q2 1 T1 T2
ð18Þ
For this combination (Q1 g 0 and Q2 e 0), the dependence of efficiency on T1 and T1 is complicated by two opposing effects. First, it can be argued that it is beneficial to generate most of the work using the internal heat engine because that minimizes the heat rejected to the ambient by the external heat engine. This requires T1 and T2 to be far apart. However, the heat-transfer processes required to reach the reactor operating temperature at the beginning of the two combustion phases generate entropy and reduce efficiency. The minimization of this entropy generation requires small differences between T1 and T2. Therefore, the optimal combination of temperatures is determined by the interaction between these opposing effects. This work expression also suggests that the optimal work extraction could be achieved by generating work using all the heat from the oxidation reactor in an external heat engine and using part of it to power a heat pump to supply heat to the reduction reactor. 662
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Table 1. Enthalpy Change (ΔH1 in J) for Oxidation Reactions Me
ΔH1 (J)
Ce2O3 Ni Mn3O4
-70 808 -101 055 -42 587
Table 2. Enthalpy Change (ΔH2 in J) for Reduction Reactions with Various Fuels fuel
As explained earlier, we also encounter cases where Q1 is negative while Q2 is positive. In this instance, work is generated using heat drawn from the reduction reactor (Q2) using an external heat engine and heat is supplied to the oxidation reactor using a heat pump. The above equation for net work output is found to be applicable for this case as well. Just as in the CLC case, if all four temperatures are the same, then Q1 þ Q2 = -ΔH (i.e., the heat in the combustion products from two reactions is enough to preheat the air and fuel) and the expression for work extracted from SCOT (for all combinations of Q1 and Q2) is reduced to the following form. T0 W ¼ - ΔH 1 ð19Þ T1
MeO
H2
CH4
iso-octane
CeO2 NiO Mn2O3
-30 760 -512 -58 980
-13 454 16 794 -41 674
-14 873 15 374 -43 094
Table 3. Percentage of Heat Released during Oxidation Stage fuel MeO
H2
CH4
iso-octane
CeO2 NiO Mn2O3
70 100 42
84 120 51
83 118 50
Table 4. Equilibrium Temperatures (in K) of Oxidation Reactions Me
Teq,1 (K)
Ce2O3 Ni Mn3O4
1609 2300 1096
This is essentially the same expression as was derived for preheated, gas-phase isothermal combustion under similar assumptions. Therefore, when T1 equals T2, the efficiency of SCOT should be comparable to that of preheated, isothermal, gas-phase combustion. Differences, if any, would result from the way preheating is conducted in the two cases. Note that there are two product streams in SCOT as in CLC, whereas gas-phase combustion has only one. The SCOT concept is analyzed for combinations of three fuels, H2, CH4, and C8H18 (iso-octane) and three oxygenstorage materials, Ni/NiO, Mn3O4/Mn2O3, and Ce2O3/ CeO2. These materials were chosen to provide high melting points, readily available thermodynamic data, and a wide range of enthalpies for the reduction reactions (discussed in more detail below). Ni/NiO and Mn3O4/Mn2O3 have been studied extensively for chemical looping (e.g., refs 14 and 17-22) with the former receiving most attention. Ceria has been used extensively as an oxygen-storage material in automotive catalysis. There is a large volume of data on redox kinetics, thermal aging, and durability of this material that could prove useful for SCOT application development. Typical temperatures of chemical looping, which are higher than in automotive catalysts, will not be a problem because Ce2O3 and CeO2 have very high melting points. Other materials that have multiple oxidation states, such as those based on iron, have been considered in CLC research (e.g., refs 9, 13, 17, and 23). The simultaneous existence of multiple oxides at equilibrium complicates the analysis, and therefore, these materials are not included here. Mn can form multiple oxides in different temperature ranges, 14 but the Mn3O4/Mn2O3 couple is considered here for reasons discussed below. More comprehensive discussions on the thermodynamics, kinetics, and stabilities of potential materials for CLC can be found elsewhere.22,24-27
The oxygen-storage materials considered here span a range of possibilities regarding exothermicity of the oxidation reaction relative to the overall exothermicity of fuel-air combustion. The heats of oxidation of the materials are shown in Table 1, whereas the heats of reduction with various fuels are shown in Table 2. For easy reading, the relative exothermicities of oxidation reactions (ΔH1/(ΔH) for all combinations are shown in Table 3. With Ce2O3/CeO2, both phases are exothermic, though most of the heat is released during oxidation. For Ni/NiO, the reduction reaction is endothermic, making this a good case for testing the advantage of having an internal heat engine. For Mn3O4/Mn2O3, the heat release is split roughly evenly between the oxidation and reduction reactions, making it a good candidate for an isothermal reactor with no temperature swings between combustion phases. Tables 4 and 5 list the equilibrium temperatures for the oxidation and reduction reactions, respectively, computed using eq 6. The fractions of the more highly oxidized state in the oxygen storage material at equilibrium are plotted for the three materials considered in Figure 7. It turns out that the equilibrium oxidation temperatures as determined using heats of reaction and entropies of oxidation reactions under standard conditions and listed in the above table are indeed good estimates for temperatures where conversion becomes equilibrium-limited. Note that both Ni and NiO melt below the estimated equilibrium temperature, Teq,1. Liquid-phase thermodynamic properties are used to compute equilibrium compositions above melting points. Equilibrium compositions were also computed when higher oxides are exposed to different reductants in amounts that would allow for complete reduction to the desired lower
(17) Cho, P.; Mattisson, T.; Lyngfelt, A. Fuel 2004, 83, 1215–1225. (18) Mattisson, T.; Johansson, M.; Lyngfelt, A. Fuel 2006, 85, 736– 747. (19) Johansson, M.; Mattisson, T.; Lyngfelt, A. Chem. Eng. Res. Des. 2006, 84, 807–818. (20) Sedora, K. E.; Hossain, M. M.; de Lasa, H. I. Chem. Eng. Sci. 2008, 63, 2994–3007. (21) Gayan, P.; Dueso, C.; Abad, A.; Adanez, J.; de Diego, L. F.; Garcia-Labiano, F. Fuel 2009, 88, 1016–1023. (22) Chandel, M. K.; Hoteit, A.; Delebarre, A. Fuel 2009, 88, 898–908.
(23) Mattisson, T.; Lyngfelt, A.; Cho, P. Fuel 2001, 80, 1953–1962. (24) Abad, A.; Adanez, J.; Garcia-Labiano, F.; de Diego, L. F.; Gayan, P.; Celaya, J. Chem. Eng. Sci. 2007, 62, 533–549. (25) Hossain, M. M.; de Lasa, H. I. Chem. Eng. Sci. 2008, 63, 4433– 4451. (26) Fang, H.; Haibin, L.; Zengli, Z. Int. J. Chem. Eng. 2009, 2009, 1–16. (27) Adanez, J. Oxygen Carrier Material for Chemical Looping Processes - Fundamentals. 1st International Conference on Chemical Looping, Lyon, France, 2010; keynote address.
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Table 5. Equilibrium Temperatures (in K) of Reduction Reactions with Various Fuels fuel MeO
H2
CH4
iso-octane
CeO2 NiO Mn2O3
-1461 -24 -3717
-329 411 -1168
-307 317 -995
Figure 8. Equilibria of NiO reduction reactions. The equilibrium predictions are made with the fully oxidized state of oxygen-storage material and just enough reductant required for complete reduction to the desired reduced state. Table 6. Second Law Efficiencies of Unrestrained Gas-Phase Stoichiometric Combustion of the Three Fuels
Figure 7. Equilibria of oxidation reactions. The equilibrium predictions are made with the reduced state of oxygen carrier material and just enough amount of air for its complete oxidation.
fuel
η2
H2 CH4 iso-octane
0.77 0.70 0.69
In the results shown below, a series of T1 values are selected subject to these constraints and T2 is varied between T0 and T1. Potential kinetic limitations are not considered here. The exergy budget is computed with contributions from work, irreversibility of reactions (which includes heating/cooling the reactants in both phases to the target reaction temperature as well as entropy generation due to the reactions themselves), preheater irreversibility, and residual exergy in the exhaust streams. The last component is always nonzero for SCOT because the two exhaust streams have different compositions. Even if both are at ambient temperature, the composition difference between these unmixed streams has some exergy associated with it. For comparison, the second law efficiencies associated with unrestrained (nonpreheated, noncompressed) gasphase stoichiometric combustion of the three fuels under consideration here are listed in Table 6. These efficiencies are associated with corresponding adiabatic flame temperatures, which are all higher than 2200 K. The efficiency plots for SCOT using Ce2O3/CeO2 with the three fuels are shown in Figure 9. For comparison, the second law efficiencies associated with preheated, isothermal, gas-phase combustion are also included. The peak efficiency always occurs when T2 is equal to T1 for reasons pointed out earlier for cases where both oxidation and reduction reactions are exothermic. Also, peak efficiency increases with increasing T1 (due to the nature of the Carnot efficiency) except when equilibrium starts to limit the extent of oxidation reaction. This is evident from the drop in efficiency as T1 is increased from 1400 to 1600 K. At 1600 K, oxidation is equilibrium-limited and so about 14% of the fuel exits the reactor unutilized. Higher efficiency could be achieved by reducing the amount of fuel fed to the reactor to what is actually consumed in reaching equilibrium. As seen from the plots, slightly higher than conventional efficiencies (shown in Table 6) are achieved using SCOT for all fuels when T1 is close to the equilibrium temperature. The efficiency advantage would further be enhanced if heat losses are included in the analysis given that the peak temperatures for the SCOT cases analyzed here are far lower than the
oxide. In the case of Mn2O3, excess reductant led to formation of MnO, but not elemental Mn. MnO is, in fact, the chosen reduced state in some CLC studies,19,28 but the choice of Mn3O4 was based on relative heats of oxidation and reduction reactions, as explained earlier. As expected, reduction of CeO2 and Mn2O3 is always thermodynamically possible at all temperatures above 300 K. The reduction of NiO to Ni is, however, not possible below the corresponding Teq,2 values in Table 5 when using CH4 or iso-octane. The fractions of NiO at equilibrium under reducing conditions are plotted in Figure 8. Less noticeable in the plots is the fact that reduction seems limited due to the presence of CO at equilibrium at higher temperatures when using CH4 or iso-octane. The entropies of reduction reactions, as mentioned previously, are always positive, so, the sign of Teq,2 can be used to determine if a reduction reaction is exothermic or endothermic. A positive Teq,2 indicates an endothermic reaction, whereas a negative Teq,2 indicates an exothermic reaction. Lower values of Teq,2 indicate more exothermicity of reduction. It is worth noting that no fuel-material combination has a Teq,2 that is high enough to thermodynamically limit the reduction reaction. At equilibrium, both reactants and products coexist. Complete oxidation is not possible unless the higher metal oxide is being removed continuously. Another option to achieve complete conversion is to run the oxidation reactor at a T1 slightly less than Teq,1. To prevent melting, T1 also needs to be lower than the melting points of both solid-phase compounds. These two constraints provide an upper limit for T1. Most of the oxygen-carrier materials used in the analysis of McGlashan2 have Teq,1 and Teq,2 that are thermodynamically convenient for CLC but are higher than their respective melting points. As a result, those materials cannot be considered for practical implementation of CLC or SCOT. Among the materials chosen here, the oxidation temperature is limited by the melting point only for Ni/ NiO (which melts around 1700 K). (28) Ryden, M.; Lyngfelt, A.; Mattisson, T.; Chen, D.; Holmen, A.; Bjørgum, E. Int. J. Greenhouse Gas Control 2008, 2, 21–36.
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Figure 9. Comparison of second law efficiencies of SCOT using Ce2O3/CeO2 and preheated, isothermal, gas-phase combustion (symbols). T1 is the temperature at which the oxidation reaction is isothermally constrained.
the oxidation reactor using a heat pump in such cases; that is, Q1 is negative. The work in such cases is generated using heat drawn from the reduction reactor. When T2 is low (and close to T0), the efficiency of this process is very low and much of the heat is rejected to the ambient. As a result, for very low values of T2, the efficiency could even drop to negative values, as seen in the plot. When T2 is close to T1 (where efficiencies are higher), however, both Q1 and Q2 are positive. Peak efficiency increases with T1 until equilibrium limits the extent of the oxidation reaction, and fuel utilization begins to drop just as for the Ce2O3/CeO2 system. The efficiency plots for Ni/NiO in Figure 11 show very different trends than those for the other two materials. This is due to the fact that the NiO reduction reactions are neutral or endothermic for all three fuels. In the case of H2, the reduction is only slightly exothermic and Q2 is positive (this is because the reduction reaction starts out with NiO at T1, which is higher than the target reaction temperature of T2). For the other two fuels, Q2 is actually negative and SCOT with these fuels (using Ni/NiO) is close to the chemical looping combustion case where an internal engine takes heat from the oxidation reactor and rejects heat to the reduction reactor. For such cases, the optimal efficiency using CLC is achieved by operating the reactors at corresponding equilibrium temperatures. This seems to be true for SCOT even when, unlike CLC, the solid-phase reactants in both reactors start out at temperatures different from the intended target temperatures. The peak efficiency for all values of T1 occurs when T2 is closer to the equilibrium reduction temperature than in the case of the other two oxygen carrier materials considered. As T1 approaches Teq,2, the extent of the reduction reaction and fuel consumption become equilibrium-limited, as evident from Figure 8. The efficiency drops as a result of unused fuel exiting the reactor.
adiabatic flame temperatures (the peak temperatures during conventional combustion) of the fuels. For comparison, the efficiencies achieved through isothermally constrained gas-phase combustion are also shown in the plots. Excluding the differences in preheating, the efficiencies should be nearly the same as in SCOT when T1 equals T2. The plots show a slightly higher efficiency for isothermal gas-phase combustion. This is due to the fact that the enthalpy in the N2 exhaust from the oxidation reactor is not sufficient to preheat the inlet air to T1, so some of the heat generated by the reactions is used to bring the oxidation reactor temperature to T1 before heat is drawn from it. The gaseous products of the reduction reaction have more heat than what is needed to preheat fuel to T2 (which equals T1), but the excess is not used to help in preheating air. The gasphase combustion does not have this problem because a single product stream is used to preheat both air and fuel. The qualitative trends of SCOT with Mn3O4/Mn2O3, shown in Figure 10, are similar to those observed with Ce2O3/CeO2. This is because both oxidation and reduction reactions are exothermic when using either of these materials. The peak efficiencies for each value of T1 are also comparable to those of preheated, isothermal, gas-phase combustion at the same temperature. There is, however, a difference that is not fully evident in the efficiency plots. The Ce2O2 oxidation reaction has an enthalpy change that is above 80% of the overall enthalpy of fuel-air combustion, whereas for Mn3O4, it is near 50%. Therefore, the oxidation reaction is much less exothermic in the case of Mn3O4. The oxidation reaction starts out with the higher oxide being at T2, and, as explained earlier, if T2 is much lower than T1, the oxidation reaction involving some solid materials may not be exothermic enough for the oxidation reactor to reach the target temperature T1. Therefore, heat needs to be added to 665
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Figure 10. Comparison of second law efficiencies of SCOT using Mn3O4/Mn2O3 and preheated, isothermal, gas-phase combustion (symbols). T1 is the temperature at which the oxidation reaction is isothermally constrained.
Figure 11. Comparison of second law efficiencies of SCOT using Ni/NiO and preheated, isothermal, gas-phase combustion (symbols). T1 is the temperature at which the oxidation reaction is isothermally constrained.
One noteworthy fact in Figure 11 is that the efficiency of SCOT never actually exceeds that of preheated, isothermal, gas-phase combustion for comparable peak temperatures. It is important to recognize that the abscissa is the reduction temperature and not
the peak temperature for SCOT, whereas in the case of preheated, isothermal, gas-phase combustion, it is the peak temperature. In summary, SCOT provides a way to harness some of the efficiency advantages (over unrestrained gas-phase combustion) 666
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We first provide a simplified analysis of this novel process; a more detailed exploration with temperature-dependent thermodynamic properties and iso-octane fuel follows. Just as in our simplified analysis of preheated, isothermal combustion and CLC, we assume that there is enough heat in the combustion products to preheat both reactant streams to the target combustion temperature (TC). The additional heat, QR, coming from the internal heat engine is assumed to equal the heat of the reformation reaction. Thus, the combination of the heat in the combustion products and QR is sufficient to facilitate entry of air and syngas into the combustion chamber at temperature TC. The heat of the syngas combustion reaction would, therefore, equal -ΔH þ QR. This amount of heat needs to be removed from the combustion chamber on a continual basis for maintaining isothermal conditions. Assuming that the internal engine operates at Carnot efficiency, the following relationships are easily derived. TC - 1 QR ð20Þ Wint ¼ TR
Figure 12. Simplified schematic of proposed isothermal combustion with reforming. This process differs from preheated, gas-phase, isothermal combustion in that fuel in it is reformed in the process of getting preheated and also the heat rejected from the heat engine is utilized (in addition to heat from combustion products) for preheating.
" Wext ¼
of CLC without the problems associated with moving solids and heat exchange between them. Although the efficiency of SCOT is lower than that of isothermal, preheated, gas-phase combustion, the former does offer some practical advantages. SCOT has a natural barrier against thermal runaway because the oxidation reaction, during which much of the heat release occurs, becomes thermodynamically limited as the temperature increases beyond its equilibrium temperature, which, in the case of most material-fuel combinations, is well below the adiabatic flame temperature. There is also no lean flammability limit associated with SCOT. There are, however, some practical problems with SCOT. It is not a steady-state process. Thermal stresses and fatigue of solidphase materials resulting from temperature swings can pose a challenge for long-duration use. In the case of SCOT with an internal engine, heat needs to be removed and added at different times to the same location. The phasing of such alternating heat transfer is hard to achieve. These problems are not so acute if temperature swings are minimized and perfectly isothermal SCOT can be achieved. Two of the oxygen-carrier materials seem to be good candidates for such a process. Alternative to CLC: Thermochemical Recuperation (TCR). As discussed above, some of the efficiency benefits of CLC and SCOT are derived from the use of an internal heat engine between the two reactors, thereby avoiding or minimizing heat rejection to the ambient. There may be other processes that can be used to achieve this effect. Consider, for example, the schematic of thermochemical recuperation illustrated in Figure 12. The process is similar to preheated, isothermal combustion except that the fuel is reformed to syngas during preheating. The fuel stream needs to include water for steam reformation when using fuels other than methanol. Given that hydrocarbon steam reformation is endothermic, the preheater requires more heat than what is available from the combustor exhaust gases. This heat is provided by an internal heat engine that is drawing heat from the isothermal combustor at temperature (TC) and rejects heat (QR) at temperature TR to the fuel preheater/ reformer.
# TC T0 - ΔH þ 1 QR 1 TR TC
ð21Þ
If TR is chosen as per the following equation, there is no need for an external engine. TC QR ¼ 0 - ΔH þ 1 ð22Þ TR All the work is generated by the internal engine, and it equals -ΔH, which means a 100% first law efficiency. This, as stated previously, is possible if the exergy of the fuel is higher than its heating value, which is the case with most of the hydrocarbon fuels. The assumptions made in the above simplified analysis may, however, be invalid in practice. The temperatures of the two product streams as they exit the two preheaters (T00 and T000 ) in Figure 12 may be different from ambient temperature for the preheating and reforming to be compliant with both laws of thermodynamics. To explore this possibility, we perform a simple analysis of the process with iso-octane fuel for a range of combustion-chamber temperatures. For each combustion-chamber temperature, a range of TR values is chosen and overall second law efficiencies are computed. It turns out that, for each value of TC, there is a minimum value of TR below which the preheating and reforming of the fuel stream to syngas is impossible for any value of QR. This minimum value of TR varies little with TC and is roughly around 640-650 K. This, coincidentally, is roughly the temperature at which the degree of steam reformation of iso-octane at equilibrium reaches 50%.29 Figure 13 shows the dependence of overall second law efficiency on TR for a range of TC values. The efficiency is highest when TR is close to the minimum possible temperature (of 650 K) for all values of TC. This is because the fraction of combustion heat (which is fixed for a given TC) extracted by internal engine is highest in this limit. Though the efficiency of the internal engine is lower, the heat rejected is utilized internally within the system for reformation. On the other hand, heat rejected by the external heat engine is (29) Chakravarthy, V. K.; Daw, C. S.; Pihl, J. A.; Conklin, J. C. Energy Fuels 2010, 24, 1529–1537.
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We have previously analyzed and reported thermochemical recuperation for reducing irreversibility in an earlier study.29 There are several differences between the previously analyzed concept and the one we propose here. The combustion in the earlier concept was adiabatic instead of isothermal as in the present instance. The reformation in the previous study was analyzed under far from ideal conditions. Here, we have attempted to study reformation in the limit of minimal exergy loss (which occurs when TR approaches 650 K). There was also an intercooler in the previous study that was used to cool the reformate (to keep the peak temperatures low), and this cooling introduces an exergy loss into the system. As a consequences of all these differences, the efficiencies reported here are much higher. The practical problems associated with reforming pointed out in our earlier study along with the difficulty in maintaining isothermal combustion are significant challenges to be overcome to realize the present concept.
Figure 13. Dependence of isothermal combustion with preheated, reformed reactants on operating temperatures.
Summary and Conclusions The main thermodynamic efficiency benefits of CLC depend on maintaining the isothermal reaction (and associated heat release and work generation) and low-temperature-gradient heat exchange between the reactants and products of both the oxidation and the reduction stages. In the absence of these features, CLC only provides a way of generating two exhaust-gas streams with different temperatures and compositions. Although beneficial for carbon sequestration, exhaust-gas partitioning alone does not offer a significant exergy savings over conventional unconstrained combustion. For CLC, oxygen carriers that support endothermic reduction reactions produce higher efficiencies because of the internal heat sink that is made available as an alternative to rejecting heat to the environment. For oxygen carriers that involve exothermic reduction reactions, it is not generally possible to achieve a feasible equilibrium temperature in the reduction reactor. Nevertheless, it appears that the efficiency can still increase as the reduction reactor temperature increases, and the reaction becomes more irreversible. Thus, it appears that requiring the reactions to occur closer to chemical equilibrium is not, by itself, sufficient to improve efficiency. The need for transporting large masses of oxygen-carrier solids and small-temperature gradient heat exchange between counterflowing oxygen-carrier solid streams and between volumetrically unbalanced gas streams seems to be one of the biggest practical drawbacks in implementation of CLC. A modified form of CLC that utilizes a fixed bed of solids (referred to here as SCOT) overcomes the need for transporting solids. However, SCOT suffers from a significant exergy penalty associated with heating and cooling the stationary oxygen-carrier solids during each gas modulation cycle. It is not clear how this drawback can be overcome when endothermic and exothermic reactions at different temperatures are required. Isothermal combustion with simultaneous work generation and carefully constrained heat exchange between products and reactants are the keys to achieving most of the efficiency benefits expected from CLC. It appears likely that these effects could be achieved without the need of circulating oxygencarrier solids or utilizing impractical oxygen-storage materials that have the endothermic properties required by CLC. Analysis of the simple isothermal process proposed here indicates that overall efficiencies similar to those projected
Figure 14. Variations of efficiencies of isothermal combustion of iso-octane with and without reforming during preheating with combustion temperature. Efficiency variation of preheated, isothermal combustion of H2 is also shown for comparison.
lost to the surroundings. The exergy loss in the fuel preheater/reformer also decreases as TR decreases. It approaches zero as TR approaches the 650 K. Therefore, the preheating losses are also lower in this limit. The only significant exergy loss then is due to combustion of the reformate. The optimal efficiency (with TR fixed at 650 K) is plotted as a function of TC in Figure 14. Also shown in this figure for comparison are the efficiencies associated with preheated, isothermal, gas-phase combustion of iso-octane (without reformation) and H2. The efficiency with thermochemical recuperation is higher because of lower combustion reversibility of syngas when compared to that of iso-octane. Combustion irreversibility and preheater losses are the only exergy losses in preheated, isothermal combustion, with the former being the most significant when preheating is done optimally. Efficiency increases with temperature because combustion irreversibility decreases with temperature. Because H2 and syngas have similar combustion irreversibilities, when reforming is done optimally (with TR close to 650 K), iso-octane can be combusted with the same efficiency as that of preheated, isothermal H2 combustion. This is evident from Figure 14. At lower combustion temperatures, however, differences between combustion efficiencies of reformed iso-octane and H2 are evident. This is because the reformation is thermodynamically limited at lower temperatures and is complete only around 1000 K.29 Note that TR is the temperature at which heat is supplied to the reformer/ preheater but is not the temperature at which reforming occurs; that is, reforming is not an isothermal process. 668
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for CLC can be achieved with relatively modest peak combustion temperatures. One of the most challenging aspects of isothermal combustion is likely to be simultaneous matching of heat generation and work output required to maintain constant temperature. It is to be noted here that this idea was analyzed computationally in past studies30-32 that reported no efficiency benefits. That was due to the fact that there was no preheating of reactants in any of these studies and there was no exhaust heat recovery by any other means. Constant-temperature combustion might be achieved practically with an intelligent combustion and engine control system, but this may be extremely difficult in a load-following context. On the other hand, it may be possible to use a modified SCOT reactor that operates with an oxygen carrier having both a highly exothermic oxidation reaction and a mildly exothermic reduction reaction. Materials such as cerium oxide might be feasible oxygen-storage candidates in this case. During the oxidation phase, the chemical equilibrium of the oxidation reaction would effectively limit heat release and prevent thermal overshoot when work extraction is diminished. Combining TCR with isothermal reaction appears to offer an additional way to obtain some of the theoretical exergy
benefit of CLC. Like the endothermic reduction reaction in CLC, hydrocarbon reforming can provide an internal heat sink for generating work that does not result in external heat transfer to the environment. The TCR concept itself has been of interest also for a number of years, but as far as we know, its similarity to this aspect of CLC has not been recognized. When combined with an isothermal reaction and heat-transfer optimization, TCR appears to offer important advantages. Together, these advanced combustion approaches may be able to achieve second law efficiencies 10-15% higher than those that are currently possible. Acknowledgment. This research was initially sponsored by the Laboratory Directed Research and Development Program of Oak Ridge National Laboratory, managed by UT-Battelle, LLC, for the U.S. Department of Energy and continued later with support from the U.S. Department of Energy (DOE) under Contract No. DE-AC05-00OR22725 with the Oak Ridge National Laboratory, managed by UT-Battelle, LLC. The authors specifically thank Gurpreet Singh of the DOE’s Office of Vehicle Technologies for sponsoring this work. The authors also thank Drs. K. Dean Edwards and Charles Finney of Oak Ridge National Laboratory for enlightening discussions and insights. This research was sponsored by the US Department of Energy under contract number DE-AC05-00OR22725 with the Oak Ridge National Laboratory, managed by UT-Battelle, LLC. The authors specifically thank Gurpreet Singh of DOE’s Office of Vehicle Technologies for sponsoring this work. The authors also thank Dr. K. Dean Edwards and Dr. Charles Finney of Oak Ridge National Laboratory for enlightening discussions and insights.
(30) Rice, M. J. M.Sc. Thesis, Virginia Polytechnic Institute and State University, Blacksburg, VA, 2004. (31) Druecke, B. C. M.Sc. Thesis, University of Wisconsin, Madison, WI, 2006. (32) Teh, K.-Y.; Edwards, C. F. J. Dyn. Syst., Meas., Control 2008, 130, 130-139.
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