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Article Cite This: Energy Fuels 2017, 31, 11348-11361

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Chemical Reactor Modeling of Oxy−Fuel Combustion Chamber for Semiclosed Combined Cycle Vicente P. Timón,* Gregorio Corchero, and José L. Montañeś Escuela Técnica Superior de Ingenieros Aeronáuticos, Universidad Politécnica de Madrid, Pz. Cardenal Cisneros 3, 28040 Madrid, España ABSTRACT: This paper compares standard gas turbine combustion chambers and CO2-diluted oxy−fuel combustion chambers for a semiclosed combined cycle at a preliminary design level. To this end, simple chemical reactor networks, based on the wellstirred reactor plus plug flow reactor scheme, are analyzed using the Cantera package and the GRI 3.0 chemical kinetics mechanism. The focus is put on the CO consumption process and the final CO concentration. The behavior of this model suggests the use of the adiabatic equilibrium temperature to characterize the composition at any station inside the chamber and the incipient lean blow-out equilibrium temperature to fix the well-stirred reactor volume. This model is applied to a feasible design point of a power production cycle (combustion exit temperature 1600 K, combustion pressure 30 bar). The fuel is a natural gas with an 87% by volume CH4 content, the ASU stream is a 95% O2 gas, and the recirculated gas is an 82% CO2 gas. The residence times required for CO burnout are approximately 30% greater than those for air combustion for the same conditions, although the required lengths are much closer. The residence times and lengths would be reduced if the combustion exit temperature or the combustion pressure of the cycle was increased.

1. INTRODUCTION The increasing concern about global warming has led to the development of new power production cycles, designed to facilitate the process of carbon dioxide capture and sequestration. One of the techniques under study is based on oxy−fuel combustion. An air separation unit (ASU) extracts the nitrogen from air, obtaining a high-purity (≥95%) oxygen stream, which is used to burn the fuel. The combustion products consist mainly of carbon dioxide and water, which can be removed by condensation; the resulting high CO2 purity gas can then be processed for sequestration. Since pure oxy−fuel combustion would lead to a very high combustion temperature, a diluent must be used to reduce it. The semiclosed oxy−fuel combustion combined cycle (SCOC-CC),1,2 is a high-efficiency cycle based on oxy−natural gas combustion, with a high CO2 concentration pressurized gas as working fluid. The combustion gases are expanded in a turbine and then passed through a heat exchanger in order to recover more energy by means of a bottoming cycle. After the water condensation and the CO2 capture processes, the remaining gases are recirculated and compressed again. This semiclosed cycle is similar to a standard air Brayton cycle, where combustion gases, consisting mainly of carbon dioxide, have substituted the nitrogen. A parametric analysis of the cycle has given a feasible design point, with a maximum pressure of around 30 bar and a combustion exit temperature, T4t, near 1600 K; these values are similar to that of existing aeroderivative and heavy duty gas turbines. The time and effort required to develop this kind of cycles would be greatly shortened if existing gas turbine preliminary design techniques could be adapted to the working fluid rich in CO2. This paper will compare standard gas turbine combustion chambers and CO2-diluted oxy−fuel combustion chambers for the SCOC-CC, at a preliminary design level. Several qualitative © 2017 American Chemical Society

differences between air and oxy−fuel combustion are relevant to this case. (1) In diluted oxy−fuel combustion there are three streams to consider (recirculated gas, ASU oxidant stream, and fuel). Therefore, one parameter must be added to the flow distribution problem of a standard gas turbine combustion chamber. However, the problem is constrained by the energy consumption of the ASU, requiring that the combustion take place with little excess oxygen. The oxygen concentration in the oxidant and the excess oxygen needed to give similar combustion characteristics to that of combustion with air have been investigated in semi-industrial furnaces by some authors.3,4 Their results are similar, with a 28% by volume O2 and a 2% excess O2 for natural gas fuel3 and 27% O2 and 3.8% excess O2 for propane fuel.4 Other authors have studied a premixer−combustor with a planar sudden expansion, finding that a minimum of 30% O2 is needed for flame sustainability in the combustor.5 Similar studies in swirl-stabilized gas turbine model combustors have shown that O2 content must be greater than 21%6,7 with stability problems for O2 below 25%7 and less stable combustion even for 30% O2 content.6 As has been pointed out,8,9 these results may change if a gas turbine-like combustion chamber is considered, especially at the high pressure (30 bar) required. An added complexity for semiclosed cycles is that any excess oxygen will be recirculated, so that the actual equivalence ratio inside the chamber depends on the recirculated gas mass flow. Received: May 8, 2017 Revised: August 28, 2017 Published: August 28, 2017 11348

DOI: 10.1021/acs.energyfuels.7b01311 Energy Fuels 2017, 31, 11348−11361

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Energy & Fuels

Figure 1. Scheme of the semiclosed oxy−fuel combustion combined cycle.

lower actual equivalence ratios and excess oxygen values closer to that reported by other authors. Issues 1−3 presented above will then be studied by means of a simplified model of an ideal oxy−fuel combustion chamber, consisting in a well-stirred reactor (WSR) followed by a plug flow reactor (PFR). This simple reactor network has the advantage that it does not require the full specification of the combustion chamber geometry nor empirical information about the combustion process. The python interface of the Cantera package24 will be used, together with the GRI-Mech 3.025 chemical kinetics mechanism, comprising 325 elementary reactions and 53 species. This oxy−fuel combustion chamber has three independent input flows: two of them can be the oxidant air and excess air for air combustion and the ASU mass flow and recirculating gas for diluted oxy−fuel combustion; the fuel is the third input mass flow, being natural gas for the case under study. The available recirculating gas or excess air can be distributed along the chamber in order to choose the WSR temperature and the additional dilution along the PFR. The behavior of this model suggests the use of the adiabatic equilibrium temperature Teq to characterize the composition at any station inside the chamber and the incipient lean blow-out equilibrium temperature (Teq,LBO) to fix the WSR volume. For the same Teq and Teq,LBO, the combustion with air and CO2 diluents is qualitatively similar. A point injection of the remaining recirculating gas or excess air in the station after CO burnout in the primary zone is selected as representative of the real conditions. Afterward, this model will be applied to the conditions of the SCOC-CC design point obtained from a parametric analysis.1 The fuel is a natural gas with an 87% by volume CH4 content, the ASU stream is a 95% O2 gas, and the recirculated gas is an 82% CO2 gas, carrying a 3% additional O2. The results of a parametric analysis of this model are compared with those of a similar one for standard air combustion. The residence times required for CO burnout are approximately 30% greater than those for air combustion for the same conditions, although the required lengths are much closer. This analysis also suggests that a configuration similar to that of lean premixed gas turbines, with moderate PZ equilibrium temperatures (around

(2) The high CO2 concentration gas is much denser than air and air combustion products, altering the mass flow through a given area and the residence time inside the chamber. The high CO2 concentration also changes the characteristic times associated with carbon monoxide and nitrogen chemical kinetics.9−17 Besides, the thermal and transport properties of CO2 are different to that of air. Hence, the temperatures, cooling flow distribution, combustion stability, and length required for complete combustion will be different from that of air combustion. For example, Amato et al.9 compared N2- and CO2diluted combustion in an atmospheric pressure, swirlstabilized combustor, showing that stable CO2-diluted combustion requires significantly higher temperatures than air combustion at a given reactants velocity, while Shroll et al.17 found that the transition between one unstable mode to another is predominantly a function of the adiabatic combustion temperature of the reactive mixture. (3) The removal of nitrogen from air should reduce the production of nitrogen oxides (NOx), although there are chemical effects to consider. Since this issue has been studied by many authors, both experimentally3,10,11,18,19 and numerically,8,10,11,20 this study focuses on the CO consumption process and the final CO concentration as an indicator of the combustion efficiency. (4) The higher concentration of CO2 and H2O in the combustion products, as compared with that of air combustion, significantly changes the emissivity properties.3,4,19,21−23 Since this effect would require a detailed analysis of a given combustion chamber geometry, it is left out of this study. A brief description of the SCOC-CC will be given first, together with the properties of the relevant gases at the design point given by the parametric analysis which has been published by the authors of this study.1 Afterward, as a preliminary step, the influence of the SCOC-CC design variables in the actual equivalence ratio inside the chamber will be presented, showing that the recirculation of combustion products greatly amplifies the oxidant content, giving much 11349


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H2O by condensation. Since they are required to model the behavior of the combustion chamber, which is the focus of this study, the composition and other properties of the relevant gases, including their correspondent mass flow rates (m), are provided in Table 1 for the ASU oxidant, Table 2 for the

1850 K), would get an exit CO concentration near that of equilibrium. Additional results will be presented, taking into account the possible changes in the primary design variables of the SCOCCC, i.e., the combustion exit temperature and combustion pressure. An increase in any of them will require smaller residence times or shorter chambers for CO burnout. The reason behind these results will also be investigated by means of an analysis of the CO, CO2, and CH4 production and destruction rates, the correspondent reaction pathways, and sensitivity.

Table 1. ASU Oxidant Properties and Mass Flow for the Design Point of the SCOC-CC Ar (% volume) O2 (% volume) T (K) P (bar) m (kg/s)

2. DESCRIPTION OF THE SYSTEM Figure 1 shows a scheme of the semiclosed oxy−fuel combustion combined cycle (SCOC-CC), a relatively new power production-cycled designed to facilitate the process of CO2 capture and sequestration. The cycle is similar to a standard combined cycle, with a semiclosed, Brayton, primary cycle and a heat recovery secondary cycle. The first difference with a standard combined cycle is the use of an air separation unit (ASU), which separates the nitrogen from air and generates a high-purity oxygen stream which is used to burn a natural gas fuel, giving mainly carbon dioxide and water as combustion products. After the combustion products are expanded in a turbine to obtain electrical power, the water can then be removed by condensation, making it easier to capture and store the remaining CO2. Because of the need to generate the oxygen stream, with an associated energy penalty, it is desirable to operate with little or no excess oxygen. As near-stoichiometric pure oxy−fuel combustion would give an excessive temperature for the turbine, the remaining combustion gases are recirculated in order to control the combustion chamber end temperature (T4t). Therefore, the primary cycle can be implemented as a gas turbine, where the high-pressure turbine (HPT) drives the high-pressure compressor (HPC) and the ASU oxidant compressor and provides auxiliary power, including the power needed for the CO2 compression and capture system (CCS), while the low-pressure turbine (LPT) provides the power for external use. The heat recovery steam generator (HRSG) and the additional cooler/condenser are responsible for the heat transfer to the secondary (Rankine) cycle and for leaving the recirculated gases at the proper conditions for the entry of the HPC of the primary cycle. The remaining part of the flue gases is derived out and compressed up to 110 MPa for later treatment and CO2 storage. The secondary cycle can be modeled as a single-pressure steam/water Rankine cycle, with a maximum temperature and mass flow rate depending on the conditions at the exit of the LPT of the primary cycle. A parametric analysis of the performance of this system has been published by the authors of this article,1 and the details of the simulation model and its results will not be repeated here, for brevity. This analysis has given a feasible design point, with a maximum pressure around 30 bar and a combustion exit temperature, T4t, near 1600 K; these values are similar to those of existing aeroderivative and heavy duty gas turbines. The time and effort required to develop this kind of cycle would be greatly shortened if existing gas turbine preliminary design techniques could be adapted to the working fluid rich in CO2. Despite other simplified descriptions of the cycle, the recirculated gas is not pure CO2; the ASU oxidant is not pure oxygen, the natural gas can include several nonhydrocarbon species, and it is difficult to remove all of the

4 96 291 1.25 61.5

Table 2. Natural Gas Fuel Properties and Mass Flow for the Design Point of the SCOC-CC N2 (% volume) CO2 (% volume) CH4 (% volume) C2H6 (% volume) C3H8 (% volume) T (K) LHV (kJ/kg) m (kg/s)

1.54 0.98 87.12 9.01 1.34 300 47100 15.3

Table 3. Recirculated Gas Properties and Mass Flow for the Design Point of the SCOC-CC CO2 (% volume) Ar (% volume) O2 (% volume) N2 (% volume) H2O (% volume) T (K) P (bar) m (kg/s)

82.41 6.76 3.24 1.15 6.44 310 1 638.3

natural gas fuel, and Table 3 for the recirculated gas corresponding to the feasible design point obtained by the parametric analysis. The temperatures of the gases after the compression to 30 bar are 751 K for the ASU oxidant and 665 K for the recirculated gas. The ratio of recirculated gas to ASU mass flow is equal to 10.4 at the entry of the HPC, but it is equal to 8.3 at the entry of the combustion chamber, since part of the recirculated gas has been derived out for cooling the HPT.

3. CYCLE STOICHIOMETRY The composition at any station inside a standard gas turbine is usually fixed assuming that it derives from either the mixing or the combustion of a certain mass flow of air, fuel, and ambient water. It is common to give the mass flows in nondimensional form, dividing them by the air mass flow. Therefore, the fuel mass flow is fixed by the fuel−air ratio ( far) mfuel = far mair and the water mass flow by the water−air ratio (war) mwater = war mair 11350


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Energy & Fuels In order to study the combustion properties it is more useful to express the far in terms of the stoichiometric far ( farst) and the equivalence ratio (Φ), which is less than one for lean combustion. Therefore, the fuel mass flow can be expressed as

To complete the description given by the actual equivalence ratio, a dilution ratio (d) can be used to set the ratio of the diluent mass flow to the total oxidant mass flow. The dilution ratio can be shown to be d = (1 + r )/(1 + YO,RGr ) − 1

mfuel = Φfarstmair

The Φ, d representation also simplifies the comparison with combustion with air, which can be interpreted as a diluted combustion using oxygen as oxidant and nitrogen as diluent with a fixed dilution ratio dair. The simplified composition of 79% N2 and 21% O2 gives dair = 3.29. If dair is used with CO2 diluent, the resulting volume fraction of O2, XO2 = 30%, is the minimum value required for stable combustion as reported by Ditaranto and Hals;5 it is also near the values used by Tan et al. (28%)3 and by Andersson and Johnsson (27%).4 Combustion with air can also be expressed in a similar way to the φ, r representation. The total air mass flow can be artificially split into an oxidant mass flow mair,o, and an excess air mass flow, mair,e. If the main oxidant mass flow is used to nondimensionalize the others, the fuel mass flow can be given by an almost arbitrary cycle equivalence ratio, φ, which can be taken to be equal to one for lean combustion, and an excess air−main oxidant mass flow, fixed by r. The value of the cycle equivalence ratio chosen for the design point of the cycle, φ = 0.98, gives wst = 0.41473, fst = 0.2538, and the value of the condensation temperature gives wRG = 0.028 and a RG composition of 82.41% CO2, 6.76% Ar, 6.44% H2O, 3.24% O2, and 1.15% N2. The end of combustion temperature T4t = 1600 K gives r = 8.3, Φ = 0.8, and d = 6.6. As has been said before, the efficiency losses associated with the ASU and the CCS systems are inversely proportional to φ. In this case, the efficiency losses due to the ASU and the CCS have increased 2% and 2.8% with respect to their values for φ = 1, being a second-order effect The actual vs cycle equivalence ratio relationship for these values of wst, fst, and wRG is shown in Figure 2, where it is clear

In a similar way, the composition at any station of the gas turbine of the semiclosed oxy−fuel combustion combined cycle (SCOC-CC)1 can be fixed assuming that it is derived from the three different mass flow streams. (1) The oxidant is a 96% O2, 4% Ar (all compositions will be given by volume) stream obtained from air by means of a cryogenic air separation unit (ASU) with a considerable energy cost (about 0.75 MW/(kg/s of oxidant)). The ASU mass flow will be used to nondimensionalize the others. (2) The fuel is a natural gas (NG) with the composition 87% CH4, 1.54% N2, 9% C2H6, 1.34% C3H8, and 0.98% CO2. The NG mass flow can be fixed using the stoichiometric fuel to ASU mass ratio (fst) and a cycle equivalence ratio (φ)

mNG = φfst mASU Due to the high energy cost associated with the generation of the ASU stream, φ should be near one. For the SCOC-CC, the efficiency losses associated with the ASU and the CCS systems can be shown to be inversely proportional to φ, approximately.1 (3) An additional stream is needed to reduce the high temperature that would be generated by near to stoichiometric conditions. The gas used in the SCOCCC is a recirculated gas (RG) which consists of the products of the ASU-NG combustion without the amount of water that is removed by the condensation process. The RG mass flow at any station is fixed by a recirculation ratio (r) mRG = rmASU The RG composition can be obtained by the application of species mass balance along the cycle, as a function of φ and the condensation temperature. The temperature at the exit of the combustion chamber required by the cycle parametric analysis design point allows one to obtain r at this station. Although the φ, r, representation has been found to be adequate for analysis of the cycle, it may be convenient to obtain the actual Φ and the oxygen mass fraction at any station. This can be done by the artificial separation of the RG into an oxidant (O), with the same composition of the ASU stream, and a diluent (D). The mass fraction of oxidant in the RG can be shown to be

Figure 2. Actual equivalence ratio (Φ) as a function of cycle equivalence ratio (φ) and recirculation ratio (r) for wst = 0.41473, fst = 0.2538, and wRG = 0.028.

YO,RG = (1 − φ)(1 − wRG)/(1 − φ − φ(1 − fst ) (1 − wst))

that the high values of r greatly amplify the oxidant content, giving much lower actual equivalence ratios and excess oxygen values closer to that reported by other authors,3−5 with better combustion characteristics. As a limit of the combustion process, a map of the equilibrium temperature and CO concentration as a function of Φ and d is presented in Figure 3 for CO2 diluent and pure CH4 fuel; the curve for standard air (N2 diluent d = 3.29) is also shown to serve as a reference. To achieve low CO emissions

where wRG is the mass fraction of water in the RG and wst is the mass fraction of water in the products of stoichiometric oxidant−NG combustion. Then the total oxidant mass flow is mO = (1 + YO,RGr )mASU

and the actual equivalence ratio is

Φ = φ /(1 + YO,RGr ) 11351


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All calculations will be made using the python interface of the Cantera package24 and the reaction mechanism GRI-Mech 3.0,25 a compilation of 325 elementary equations, 53 species, and the associated rate coefficient expressions and thermochemical parameters.

5. COMBUSTION CHAMBER MODEL Different reactor networks and different methods to adjust its parameters have been used to study gas turbine combustion chambers. For example, the residence time inside each reactor can be fixed using empirical information such as CO, NOx, and UHC emissions32 by means of experimental flow visualization31 or by comparison with CFD results.33 Taking into account the air or RG mass flow distribution inside the combustion chamber, two different configurations can be considered.34 Type 1: Standard aircraft engine combustion chamber design methods35 divide the chamber into a PZ, where the flame is anchored by means of recirculation; an intermediate zone (IZ) where the PZ gases continue to burn toward completion while mixing with air; and a dilution zone (DZ) where any remaining air is mixed with the IZ gas. IZ air is injected to maintain a constant temperature along this zone, equal to the PZ exit temperature. The liner temperature is controlled by means of film cooling, requiring additional air injection. Type 2: In lean premixed gas turbines, about 80% of the available air is injected into the PZ, reducing or eliminating the need for film cooling or IZ dilution holes. Most of the liner cooling is achieved by means of internal convection, impingement cooling, and thermal barrier coatings; combined cycle heavy duty gas turbines can include closed-loop steam cooling for improved efficiency. A generic model, able to simulate both configurations, will be used. Combustion will be modeled with a very simple network consisting of a WSR followed by a PFR. This network is based on Bragg’s classical study in 195336 and continues to be applied to gas turbine combustion.37,38 It has also been used before to simulate oxy−fuel combustion for some different design conditions,39 but a systematic parameter variation is not found in the literature. A greater number of reactors would add more parameters (and geometric constraints) to the model, without necessarily increasing the accuracy.40 The total mass flows available for combustion can be fixed using the parameters defined in the cycle stoichiometry section. (1) For combustion with air, mair and farst will be used. In order to obtain a similar model to that for oxy−fuel combustion, the air mass flow will be split into the oxidant and excess air parts, with a cycle stoichiometry ratio taken to be equal to one. A fraction KPZ of the excess air is assumed to be perfectly mixed with the fuel before it enters the PZ. (2) For diluted oxy−fuel combustion, mASU, φ, fst, and r will be used. A fraction KPZ of the RG is assumed to be perfectly mixed with the fuel and the oxidant mass flows before the mixture enters the PZ. The PZ is modeled using a WSR, with an associated volume VPZ, giving the stability characteristics of the whole chamber. The combustion products at the exit of the PZ are then allowed to burn in a PFR with an associated volume VPFR. Injection of the remaining dilution gas begins at a certain station of the PFR, nd, and can follow any desired mass flow distribution. The part of the PFR between the PZ exit and the beginning of

Figure 3. Equilibrium CO concentration and temperature as a function of the actual equivalence ratio (Φ) and dilution ratio (d) for CO2 diluent and 30 bar pressure. Dashed line corresponds to combustion with air.

using CO2 diluent, Φ lower than 0.9 and T4t lower than 1700 K should be used; however, equilibrium CO is much higher than that resulting from air combustion. The values chosen for the design point of the cycle and the end of combustion temperature give approximately 16 ppmv equilibrium CO.

4. CHEMICAL REACTOR NETWORKS Chemical reactor models represent the combustion chamber by a network of interconnected stirred and plug flow reactors. This approach can manage complex chemical kinetics with a low computational cost, making it very useful at the preliminary phase of combustion chamber design. The modeling assumptions are given below; the reader is referred to the specialized literature (e.g., ref 26) for more details. A well-stirred reactor (WSR) is a zero-dimensional chemical reactor of constant volume VWSR in which the composition, temperature, and pressure are uniform. The following assumptions are made in this study: perfect macro- and micromixing, constant pressure, adiabatic process, and steady state. Instead of the WSR volume, a mean residence time defined as tres,WSR = ρWSR VWSR /mWSR where ρWSR is the density inside the reactor, can be used to define the WSR together with the reactor inlet mass flow, mWSR. Although a single WSR can be used to model the entire combustion chamber,27−29 it is more frequently used to simulate the primary zone (PZ) of a gas turbine combustion chamber, in order to fix its lean blow-out (LBO) characteristics.30 A plug flow reactor (PFR) is a one-dimensional chemical reactor that has uniform properties in the direction perpendicular to the flow but permits no mixing in the axial direction. If the changes in axial momentum and kinetic energy are neglected, the equations of species and energy conservation can be applied to a differential slice in the flow direction. Together with the assumptions of constant pressure, steady state, and adiabatic process, the PFR can be represented by a series of differentially small WSRs,31 allowing injection of dilution gas at any station. To define the PFR size, the total PFR volume can be given as the sum of the volumes of each WSR in the series. Alternatively, the PFR size can be fixed by the total residence time inside it, being the sum of the residence time inside each WSR. 11352


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Figure 4. Representation of a generic diluted combustion chamber model: O and R represent oxidant air and excess air for air combustion, respectively. O and R represent ASU mass flow and recirculating gas for diluted oxy−fuel combustion, respectively. F is for fuel, natural gas in the case under study.

6.1. Primary Zone. A chemical time can be associated with combustion inside a WSR; when the residence time inside the reactor is lower than this chemical time, stable combustion is not possible and the blow-out phenomenon occurs. The minimum residence time for stable combustion will be called blow-out residence time and abbreviated as tres,LBO to emphasize that only the lean combustion zone (Φ < 1) is of practical interest in this case. For air combustion, this condition is usually represented as a single curve in a plot of equivalence ratio versus the inverse of the residence time. The region where stable combustion is possible is located to the left of this curve. However, diluted oxy−fuel combustion has the amount of diluent as an additional parameter, giving a family of stability characteristics. Figure 5 (top) shows the stability characteristics of the PZ for a fixed pressure (30 bar) and inlet temperature (700 K). For an easier comparison between air and diluted oxy−fuel combustion, the oxidant is pure O2, the fuel is CH4, the

additional gas injection will be called the secondary zone (SZ), and the rest of the PFR will be called the dilution zone (DZ). A scheme of this generic chamber model is shown in Figure 4. A reference Mach number, Mref, can be assigned to the gas entering the PZ, a usual procedure in gas turbines.41 Then assuming that the static and total conditions are approximately equal, a reference area Aref can be calculated as A ref = m31√(R31T31)/(M ref P31√(γ31))

where the 31 station is defined as the beginning of the PZ, just before combustion, m31 is the total mass flow entering the PZ, T31 and P31 are the temperature and pressure at the exit of the mixer; and R31 and γ31 are the corresponding ideal gas constant and specific heat ratio of the unburned mixture. The volume of the PZ and the volume of each of the WSRs comprising the PFR are known, so their lengths can be estimated as L PZ = VPZ/A ref

and L PFR = VPFR /A ref

At the end of the chamber, the temperature should be the design point exit temperature, and ideally, the combustion products should have reached the equilibrium composition. In type 1 configurations, the reactions are usually considered to be frozen after injection of dilution air. This effect is not due to the low kinetics rate associated with the reduced temperature after dilution, since CO consumption is fast in homogeneous combustion above 1300 K,42 or near the liner cool−wall boundary layer.43 CO quenching has been related to the presence of film cooling or fuel−air nonuniform distribution.41,43 In type 2 configurations, the reactions can be considered to proceed toward equilibrium along the chamber. Since this study only takes into account chemical kinetics effects, it cannot predict the exact CO concentration. That concentration should be between that of equilibrium at PZ exit temperature for a type 1 configuration and that of equilibrium at chamber exit temperature for a type 2 configuration. The comparison between oxy−fuel and air combustion chambers should serve to diminish this uncertainty, assuming that nonchemical kinetics-related effects are similar.

6. COMBUSTION CHAMBER MODEL BEHAVIOR In this section, the combustion chamber model working with air and under oxy−fuel conditions will be studied in order to choose design parameters suitable for both cases, making easier the comparison between them.

Figure 5. PZ stability characteristics as a function of the actual equivalence ratio (Φ) and dilution ratio (d) (top) and as a function of the equilibrium temperature (Teq) and d (botom) for air and CO2diluted combustion. 11353


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Energy & Fuels diluent is pure CO2 or N2, and the inlet composition is fixed using the Φ, d representation. The dotted line in Figure 5 is that of air combustion with N2 diluent and d = dair = 3.29; a representative residence time for standard gas turbines is between 0.1 and 1 ms.30 The solid lines are for CO2 diluent, requiring more residence time than air combustion for the same d and Φ; increasing d greatly increases LBO residence time for the same Φ. However, a comparison of the residence times for the same value of Φ and d is not adequate for the SCOC-CC. In an air combustion chamber, ΦPZ is between 0.5 and 0.6,35,41 but it changes with the amount of recirculating gas in the oxy− fuel case, although the cycle stoichiometric ratio should be near to one to minimize ASU energy consumption.1 Alternative stability characteristics can be obtained if Φ is substituted by the equilibrium temperature (Teq), which would be obtained for each Φ, d point. This temperature is representative of the real temperature inside the PZ, controlling the combustion process via chemical reaction rates. Figure 5 (bottom) shows the alternative stability characteristic curves of Figure 5 (top), showing a much lower effect of d alone. Stable combustion with CO2 diluent, in the 1−0.1 ms tres range, requires d ≤ 6 and an Teq slightly greater than that for combustion with air. In view of these results, it seems adequate to use the equilibrium temperature of the PZ (Teq,PZ) as a design parameter, fixing the fraction of air/RG gas which is mixed with the fuel (KPZ). The stability behavior of the chamber will be given by the PZ volume and the associated blow-out equilibrium temperature (Teq,LBO). Design Teq,PZ must be greater than Teq,LBO, and the difference between them is the stability margin (ΔTLBO). This stability margin can be used instead of the PZ volume to fix the PZ behavior. 6.2. Secondary Zone. The lean-premixed combustor described by Snyder et al.44 will be used as a basis to compare the evolution of air and CO2-diluted combustion in the SZ. Base load (BL) operating conditions for CH4 combustion with air are inlet pressure 19 bar, inlet temperature 720 K, PZ equivalence ratio ΦBL = 0.51 (Teq,BL = 1820 K), and total residence time 6 ms. The LBO equivalence ratio is also given, ΦLBO = 0.45 (Teq,LBO = 1700 K). Since there is no additional dilution this combustor can be simulated using the combustion chamber model with a composition at the PZ inlet given by Φ and d = dair = 3.29 using N2 as diluent. A similar model can be obtained for CO2 diluent; the same d = dair will be used in order to simplify the comparison. Each VPZ will be fixed to match the given Teq,LBO. As a preliminary step, the results of the air combustion model are compared with the experimental results.44 (1) Figure 6 (top) shows the NOx emissions of the base lean-premixed combustor using two different premixers44 and the results of the air combustion model as a function of Teq,PZ. The results of the model are very near the experimental results for combustion chambers with nearly perfect premixers and different operating conditions,45 while the other experimental series represent different levels of unmixedness, with increased emission levels for increased unmixedness. In fact, the difference in emissions can be used to evaluate the level of unmixedness of a combustion chamber, as suggested by Barnes and Mellor.46 (2) Figure 6 (bottom) compares the base lean-premixed combustor experimental CO44 emissions and the results

Figure 6. Comparison between experimental44 and predicted NOx (top) and CO emissions (bottom).

obtained with the model, which gives a near to equilibrium exit concentration for all temperatures. The trend of the experimental results is that of equilibrium for temperatures greater than 1800 K, although the values are slightly lower than expected. The incomplete combustion shown by the experimental results below 1800 K is typical of gas turbines and can be related to several causes not included in the perfectly mixed WSR/ PFR model;41,43 it could be simulated assuming a much lower residence time inside the PZ (in the order of 1 ms). The evolution of CO, UHC, and NOx inside the SZ for N2 and CO2 diluents is shown in Figure 7. The process is similar

Figure 7. Evolution of contaminants in the secondary zone for CO2 and N2 diluents, d = dair = 3.19, Teq,PZ = Teq,BL = 1820.

with a very fast consumption of the fuel followed by the slow consumption of CO. Due to the asymptotic nature of this process, complete CO burnout will be considered when 1.1 times the equilibrium CO is reached.39 The leftmost point of each series corresponds to the exit of the PZ, showing that the residence time required for stable combustion at the same Teq,LBO is greater for the CO2 diluent, as expected from the LBO results. The exit concentration of CO is much higher in the case of CO2 diluent, a result explained by chemical equilibrium, as shown in Figure 3. The total length required for 11354


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Energy & Fuels CO burnout at a given exit temperature T4t, denoted as Leq,CC, will be used to compare different combustion chambers. 6.3. Dilution Zone. A proper additional dilution distribution should be the result of a detailed study of the liner shape and cooling requirements, something beyond the scope of the current study. A very simple distribution will be used instead; the number of parameters will be reduced to a minimum to allow an easy comparison between air and diluted oxy−fuel combustion. The evolution of combustion in the SZ suggests two characteristic points to begin the dilution: the exit of the PZ, typical of type 1 configurations, and the station of the SZ where the CO burn-out condition is reached (1.1 of equilibrium CO at Teq,PZ). Besides, the mass flow distribution shape and total length must be selected. A separate study of the effect of these parameters on the total length required for CO burnout has been carried out, but its full results will not be shown here for brevity. The distribution that gives the shortest length has been found to be a point injection in the station after the CO burn-out condition has been reached in the SZ. Since this study tries to obtain qualitative results, this dilution distribution will be used as a basis of comparison of the minimum chamber length required for complete combustion. Although the point distribution may seem to be unrealistic, it has the advantage that it does not require one to select the value of any additional parameter. Moreover, the length required for CO burnout that it predicts is not very far from more realistic distributions. An example is shown in Figure 8

for a combustion chamber model corresponding to the SCOCCC inlet and outlet conditions, Teq,PZ = 1850 K, ΔTLBO = 50 K, and Mref = 0.05. This figure shows the evolution of the temperature and the CO concentration as a fraction of the equilibrium CO for several uniform distributions of dilution; its lengths ΔLdilution vary between point injection and 0.1 m. The length required for CO burnout (i.e., to reach the 0.1 value of the y axis) increases with the dilution distribution length but not very much, even for ΔLdilution = 0.1 m.

7. COMPARISON OF AIR AND SCOC-CC COMBUSTION CHAMBERS In this section, the previous results will be used to estimate the behavior of air combustion chambers designed for typical heavy duty gas turbines inlet conditions (P = 30 bar, T = 720 K) and combustion chambers designed for the SCOC-CC conditions. Air and diluted oxy−fuel combustion chamber models are constructed to give the same exit mass flow (587.4 kg/s) and temperature (1600 K) of the SCOC-CC design point. The dilution mass flow scheme consisting of instant mixing after CO burnout in the SZ will be used in all cases. A PZ inlet reference Mach number (0.05) is assumed to estimate the size. Teq,PZ and ΔTLBO remain as parameters of the models; once the value of both is selected, the reference area, PZ volume, and length, the total CO burnout length can be computed and used as an estimation of the required combustion chamber length. The effect of Teq,PZ with a fixed ΔTLBO = 100 K is presented first. Numeric results for the SCOC-CC chamber are shown in Table 4, while the results for air combustion are shown in Table 5. The values of the obtained lengths are similar to those found in the literature. Even near the wall (at about 1100 K), CO may reach its equilibrium value in less than 0.3 m;44 this length can be considered typical of aero engines combustion chambers.45 The lengths of stationary gas turbines are usually greater; the family of stationary gas turbines of General Electric MS70009000 uses combustion chambers ranging from 0.76 (F family) to 0.97 m (E family).47 However, these turbines are designed to be able to burn several different types of fuels, and vaporization or ignition processes may dominate the whole process.46 The CO values of the air combustion chambers are also typical. For Teq,PZ < 1800 K, COeq is less than 10 ppm, which is a typical value given by gas turbine manufacturers.43 However, the maximum value given for latest gas turbines is less than 6 ppm,43 while the result of field tests is in the order of 2 ppm.34 The last value is near the COeq for 1700 K. The values obtained for the SCOC-CC conditions are much greater, as expected from the equilibrium computations, due to the high CO2 content. In both cases, the true value of CO should be between COeq,PZ and COeq,CC. In low-emission gas turbines, the existence of out of equilibrium CO is mainly attributed to defects in the air−fuel

Figure 8. Evolution of temperature and CO above equilibrium in the SCOC-CC chamber for uniform distributions of increasing length ΔLdilution, Teq,PZ = 1850 K, ΔTLBO = 50 K, and Mref = 0.05.

Table 4. Results for SCOC-CC Combustion Chamber with ΔTLBO = 100 Ka

a

Teq,PZ (K)

VPZ (m3)

mPZ (kg/s)

Aref (m2)

Uref (m/s)

COeq,PZ (ppmv)

tres,PZ (ms)

LPZ (m)

COeq,CC (ppmv)

tres,PZ+SZ (ms)

LPZ+SZ (m)

tres,CC (ms)

LCC (m)

1700 1800 1900 2000

0.068 0.031 0.016 0.008

523 471 426 388

1.17 1.06 0.96 0.88

20.7 20.8 20.8 20.9

54.0 161.0 426.0 1015.0

1.30 0.62 0.32 0.18

0.06 0.03 0.02 0.01

15.9 15.9 15.9 15.9

2.06 1.04 0.57 0.33

0.11 0.06 0.03 0.02

2.94 2.21 1.95 1.87

0.16 0.13 0.13 0.14

CC refers to the whole combustion chamber lengths and residence times. 11355


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Energy & Fuels Table 5. Results for Air Combustion Chamber with ΔTLBO = 100 Ka

a

Teq,PZ (K)

VPZ (m3)

mPZ (kg/s)

Aref (m2)

Uref (m/s)

COeq,PZ (ppmv)

tres,PZ (ms)

LPZ (m)

COeq,CC (ppmv)

tres,PZ+SZ (ms)

LPZ+SZ (m)

tres,CC (ms)

LCC (m)

1700 1800 1900 2000

0.057 0.024 0.010 0.005

519 464 418 379

1.34 1.20 1.08 0.98

26.1 26.0 26.0 25.9

2.0 7.0 22.0 63.0

0.79 0.34 0.16 0.08

0.04 0.02 0.01 0.01

0.5 0.5 0.5 0.5

1.54 0.69 0.34 0.18

0.10 0.05 0.02 0.01

2.33 1.68 1.49 1.49

0.15 0.12 0.12 0.14

CC refers to the whole combustion chamber lengths and residence times.

premixing,35,44 giving rise to locally fuel-rich zones, which can interact with turbulence and/or the walls, freezing the CO consumption reactions. Some authors have proposed measuring the quality of premixing by the amount of out of equilibrium CO.42 Another possible cause is the mixing of the primary zone combustion gases with that of film cooling or dilution air. The trend in stationary, low-emission, gas turbine combustion chambers has been to replace film cooling of the liner by impact and convective cooling.41 This has been possible by the use of moderate temperatures in the primary zone (below 1850 K), a condition required also by the low NOx emission criteria. Figure 9 presents a more detailed comparison of the residence times inside each zone of the SCOC-CC and air

Figure 10. Variation with Teq,PZ of the length of each zone of the SCOC-CC and air combustion chambers, with ΔTLBO = 100 K; PZ, SZ, and DZ fitted to an exponential function.

for both cases, much lower than the value obtained for minimum residence time inside the combustion chamber. The preceding results have been obtained for a fixed ΔTLBO = 100 K. This parameter has a very great influence on the size of the PZ, but its influence on the whole combustion chamber size or length is not so great, as shown in Figure 11, where the difference between the minimum length for ΔTLBO = 5 and 150 K is only about a 10%. The value of Teq,PZ corresponding to this

Figure 9. Variation with Teq,PZ of the residence time in each zone of the SCOC-CC and air combustion chambers, with ΔTLBO = 100 K; PZ, SZ, and DZ fitted to an exponential function.

combustion chambers, with Teq,PZ between 1700 and 2000 K and ΔTLBO = 100 K. PZ and SZ results approximately follow exponential functions diminishing with Teq,PZ. However, greater values of Teq,PZ, lead to higher values of COeq,PZ, which must be allowed to react until CO burnout at T4t is reached. The residence time inside the DZ approximately follows an exponential function increasing with Teq,PZ. The residence times of the SCOC-CC cases are greater than that of air combustion for all zones, especially for the PZ. The results for the whole combustion chamber are approximately 30% higher for the SCOC-CC case. It is also interesting to notice that the air combustion times for the whole chamber present a minimum near 1950 K, while the SCOC-CC case reaches its minimum for 2000 K. Figure 10 presents the same comparison for the lengths of each zone. The results of each zone are qualitatively similar to that of the residence times, but while the length of the PZ of the SCOC-CC case is again much greater than that of the air case, the size of the DZ of the SCOC-CC cases is slightly smaller than those of the air cases, giving much closer lengths for the whole chamber. It can also be seen that the Teq,PZ corresponding to the minimum chamber length is near 1850 K

Figure 11. Variation with Teq,PZ and ΔTLBO of the length of the PZ (top) and the whole combustion chamber (bottom) for SCOC-CC; PZ results are fitted to an exponential function. 11356


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Energy & Fuels minimum does not change very much for ΔTLBO ≤ 100 K. These results suggest that the value of ΔTLBO for the SCOCCC can be chosen to be slightly greater than those of similar air combustion chambers without an excessive impact in the combustion chamber size and that ΔTLBO = 100 K can be used to give representative values.

8. EFFECT OF CHANGES IN THE SCOC-CC DESIGN PARAMETERS Since the analysis of the SCOC-CC gave a possible design point and studied the influence of possible changes of its main design parameters,1 this section will show how the changes in the combustion exit temperature (T4t) and combustion pressure (P3t) of the cycle would affect the combustion chamber. The effect of P3t is shown in Figure 12. As was to be expected, an increase in pressure enhances the CO con-

Figure 13. Variation with Teq,PZ and T4t of the residence time (top) and length (bottom) of the SOCC-CC combustion chamber, with ΔTLBO = 100 K.

Two different stations will be considered, corresponding to the exit of the WSR which forms the primary zone and to the beginning of the dilution zone, after mixing with the additional dilution gas. The trends shown by the comparison of oxy−fuel and air chambers will be also be qualitatively related to the species rates. The reaction numbers will correspond to the order in which they appear in the GRI-Mech 3.0 mechanism data file.25 The differences between air and oxy−fuel combustion at the exit of PZ are considered first for Teq,PZ = 1850 K. Although the relative effect of the different reactions varies with Teq,PZ, the comparison does not change too much, and the effect of this variable will be considered afterward for oxy−fuel combustion only. The temperature at the exit of the PZ by is near Teq,PZ, being slightly higher for the air case with 1807 K, while the oxy−fuel case reaches only 1750 K. The corresponding molar densities are very similar; that of the oxy−fuel case is only 3% higher. As has been pointed out by many other authors, e.g.,12 the greater rate of CO oxidation to CO2 is due to the reaction with OH

Figure 12. Variation with Teq,PZ and P3t of the residence time (top) and length (bottom) of the SOCC-CC combustion chamber, with ΔTLBO = 100 K.

sumption process. A further effect not shown here is that a greater pressure implies lower values of COeq. Conversely, a selection of a smaller pressure would increase COeq and require much greater values of tres and lengths. The position of Teq for minimum tres and length changes very little in both cases. Figure 13 shows the effect of an increase in T4t. The required values of tres and lengths greatly decrease with T4t increases. The position of Teq for minimum tres and length are clearly shifted to higher values when T4t is increased.

OH + CO ↔ H + CO2

(R99)

However, CO2 is also produced by the reaction of CO with HO2 HO2 + CO ↔ OH + CO2

(R120)

The importance of the forward step of R120 for CO2 production from CO is similar for both cases, being 4% for the air case and 6% for oxy−fuel. The overall CO to CO2 production rate of the oxy−fuel case is 1.8 times that of the air case. Although the reaction rate coefficients are slightly higher for the air cases due to the higher temperature, the molar concentration of CO is 3.3 times that of the air case, and the OH concentration in the oxy−fuel case is

9. SPECIES PRODUCTION AND DESTRUCTION RATES In order to shed light on the previous results, the chemical kinetics of the air and diluted oxy−fuel combustion chambers will be studied in this section by analyzing the CO, CO2 (and CH4 for the PZ) production and destruction rates through the different reaction pathways together with a sensitivity analysis. 11357


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A sensitivity analysis of CO/CO2 with respect to the reaction rates for the air case shows that they are most sensitive to the R99 reaction, but the other most important reactions are those of the O/H radical pool, in order of importance

only 0.56 times that of the air case, making the forward step of R99 much faster. Since the molar concentration of HO2 for the oxy−fuel case is 1.1 times that of the air case, the rate of the forward step of R120 is again greater for oxy−fuel combustion, and it gains relative weight. Additional CO2 is formed directly from the triplet state of methylene, CH2, through the reaction CH 2 + O2 → 2H + CO2

H + O2 ↔ O + OH

OH + HO2 ↔ O2 + H 2O

(R290)

This reaction is responsible for more than 9% of the CO production in the air case but only for 2% for the oxy−fuel case. This can be explained by the much smaller molar concentration of O2 for the oxy−fuel case, only 0.41 times that of the air case, and the also smaller concentration of CH2, 0.92 times that of the air case. As also pointed out,12 the reverse step of R99 is responsible for most of the chemical effects of CO2 for the oxy−fuel case, since its rate is 0.645 times of the CO to CO2 total rate, while it is only 0.08 times the CO to CO2 total rate for the air case. The effect is that the net CO2 production for the oxy−fuel case is 30% lower than that for the air case, although the CO concentration is much higher. The greater rate of CO production for the air case is due to the HCO reactions HCO + H 2O ↔ H + CO + H 2O

(R166)

HCO + M ↔ H + CO + M

(R167)

HCO + O2 ↔ HO2 + CO

(R168)

(R176)

CH 2 + O2 → OH + H + CO

(R135)

CH 2(s) + O2 ↔ H + OH + CO

(R144)

CH 2(s) + O2 ↔ CO + H 2O

(R145)

O + CH3 → H + H 2 + CO

(R284)

(R287)

H + O2 + N2 ↔ HO2 + N2

(R35)

H + O2 + H 2O ↔ HO2 + H 2O

(R36)

OH + HO2 ↔ O2 + H 2O

(R87)

A similar analysis for the oxy−fuel case shows that CO and CO2 are most sensitive to R38 reaction, followed by R99 and R287. The next most important is OH + H 2O2 ↔ HO2 + H 2O

(R89)

followed by R87 and 2OH ↔ H 2O2

(R85)

The most important reactions for both cases are very similar to those for CO combustion reported by Yetter et al.48 Special importance is attributed to the chain branching reaction R38, although, in concordance with Glarborg et al.,12 no evidence is found of the competition for atomic hydrogen between R38 and R99 , which was reported by Liu et al.49 In fact, the molar concentration of H of the oxy−fuel case is 1.42 times that of the air case. Of course, fuel consumption is also important for a WSR. The rate of CH4 to CH3 decomposition is even greater than that of CO to CO2 for the air case. This is achieved mostly by

Its rate is 0.6 times the CO to CO2 production rate given by the forward steps of R99 and R120 for the air case, with R168 accounting for 42% of it. The rate for the oxy−fuel case is 0.34 times its CO to CO2 production rate, giving a similar actual value to that of the air case. The HCO concentration is also similar, while the H2O concentration of the oxy−fuel case is almost 2 times that of the air case, compensating for the lower O2 concentration. This change in concentration reduces the importance of R168 to 16%, so that the oxy−fuel case generates much more H through this route than the air case. Other hydrocarbon species like HCCO, CH2, CH2(s) (singlet state of methylene), and CH3 contribute to the CO production rate for both cases through reactions HCCO + O2 ↔ OH + 2CO

(R38)

OH + CH4 ↔ CH3 + H 2O

(R98)

with a smaller contribution of O + CH4 ↔ OH + CH3

(R11)

H + CH4 ↔ CH3 + H 2

(R53)

The relative contribution of the equations is R98, 87%, R11, 11%, and R53 2%. The rate of CH4 to CH3 decomposition for the oxy−fuel case is only slightly smaller than that for the air case, and their relative contribution is slightly different, with R98, 89%, R11, 5%, and R53, 5%. More important is the difference due to the reverse step of these equations together with H + CH3 ↔ CH4

(R52)

The rate of CH3 to CH4 transformation is 14% of the direct rate for the air case but more than 31% for the oxy−fuel case thanks to the reverse step of R98 and R53 and the direct step of R52. The net effect is that the CH4 destruction rate of the oxy− fuel case is 23% lower than that of the air case. To compare the effect of the CH4 destruction rates in the residence time inside the PZ, the classical mass-based species conservation equations for a WSR, e.g.,26 can be rewritten in terms of the molar concentrations as

The contribution of each of these species is similar for the air case, with a total CO production of 2/3 of the production through the HCO reactions. Since the concentration of HCCO for the oxy−fuel case is 2.5 times that of the air case, the contribution of R176 is similar for both cases. The concentration of CH3 is also higher for the oxy−fuel case, but the concentration of O is only 0.28 times that of the air case, so that the production rate of R284 is much smaller for the oxy−fuel case. The concentration of CH2 and CH2(s) is smaller for the oxy−fuel case, so that the corresponding rates are much smaller also.

tres = ([Xi] − ρ[X i,in]/ρin )/ωi

(E1)

which must be true for any species Xi. In E1, [Xi] and [Xi,in] represent the molar concentrations of the species Xi inside the reactor and at the reactor inlet, respectively, ρ and ρin are the corresponding (mass) densities inside the reactor and at the 11358


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Energy & Fuels reactor inlet, and ωi = d[Xi]/dt represents the rate of change of the molar concentration of Xi inside the reactor. For the particular case of CH4 at the PZ, since its exit concentration is much less than its inlet concentration, E1 can be written as tres,PZ ≈ −σPZ[CH4,mixer]/ωCH4,PZ

Of course, the SZ is a PFR, and this estimation is very far from the true value, but it captures the dependence on 1/ ωCO2,PZ of the true solution, as can be seen in Figure 16. This

(E2)

where

σPZ = ρPZ /ρmixer The effect of each term of E2 for Teq,PZ = 1850 K can be expressed as σPZ,oxy/σPZ,air = 1.54, [CH4,mixer,oxy]/[CH4,mixer,air] = 0.96, and ωCH4,air/ωCH4,PZ,oxy = 1.3, giving tres,PZ,oxy/tres,PZ,air ≈ 1.92. This expression can also be applied to changes in Teq,PZ and P3t. Figure 14 shows the variation with Teq,PZ of each term of E2 Figure 16. Variation with Teq,PZ of each term of E3 with respect to its value for Teq,PZ = 1850 K, for P3t = 30 bar, for the oxy−fuel cases compared with the result of the full simulation model.

figure shows the variation with Teq,PZ of each term of E3 with respect to its value for Teq,PZ = 1850 K, for P3t = 30 bar, for the oxy−fuel cases, compared with the result of the full simulation model. It can be seen that the dependence with Teq,PZ of tres,SZ can be mainly attributed to different CO2 production rate and that E3 follows the same trend as the true value. However, the ratio between the actual value of tres and that given by E3 is about 4. This approximation also captures well the behavior with P3t changes, although the results will not be shown here for brevity. The differences between air and oxy−fuel combustion at the beginning of the dilution zone are also considered for Teq,PZ = 1850 K. In this section the temperature is very close to T4t, which is equal for all cases, and the molar density of both cases is almost the same. Although the quantity of unburned hydrocarbons is expected to be very low, the concentration of CO and radicals is much different. The CO oxidation process is dominated by R99 reaction, and the ratio of the molar concentrations of the participant species of the oxy−fuel case to that of the air case are OH, 1.02, CO, 24.1, H, 2.7, and CO2, 17.7. The different relation between these concentrations leads to different ratios between the reverse and the direct step of R99 , which is equal to 0.584 for the oxy−fuel case and to 0.305 for the air case, so that the net rate of CO destruction for the oxy−fuel case is only 13.7 times that of the air case. A sensitivity analysis at this station gives similar results to those of the PZ, although the relative importance of the reactions changes. The CO for the air case is most sensitive to R99 followed by R38, R35, and R36

Figure 14. Variation with Teq,PZ of each term of E2 with respect to its value for Teq,PZ = 1850 K, for P3t = 30 bar, for the oxy−fuel cases.

with respect to its value for Teq,PZ = 1850 K, for P3t = 30 bar, for the oxy−fuel cases. The variation is clearly dominated by the change in the CH4 destruction rate, which is much greater for higher temperatures; similar results are obtained for air combustion. Figure 15 shows the variation with P3t of each

Figure 15. Variation with P3t of each term of E2 with respect to its value at P3t = 30 bar, for Teq,PZ = 1850 K, for the oxy−fuel cases.

2OH ↔ O + H 2O(R86)

and R87. The CO for the oxy−fuel case is most sensitive to eq R99, followed by R35

term of E2 with respect to its value at P3t = 30 bar, for Teq,PZ = 1850 K, for the oxy−fuel cases. Although [CH4,mixer] is increased with P3t, the destruction rate increases much faster. A similar procedure can be used to estimate the residence time inside the SZ as if it were a single WSR, with the equilibrium CO concentration at the exit and a CO destruction rate equal to that of CO2 generation at the exit of the PZ and assuming that the density is the PZ density tres,SZ ≈ ([COPZ ] − [COeq,PZ ])/ωCO2,PZ(E3)

(R86)

H + O2 + M ↔ HO2 + M

(R33)

R38 OH + H 2 ↔ H + H 2O

(R84)

and R287. A simple estimation procedure based on E3 can be constructed as

(E3) 11359


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Energy & Fuels tres,DZ ≈ ([CODZ0 ] − [COeq,DZ ])/ωCO2,DZ0

combustion pressure 30 bar). The fuel is a natural gas with 87% by volume CH4 content, the ASU stream is a 95% O2 gas, and the recirculated gas is 82% CO2, gas. The results have been compared with those of a similar model working with air. The following conclusions can be drawn from the study. (1) As expected from Le Châtelier’s principle, the equilibrium CO concentration is much greater for CO2-diluted combustion. (2) The residence times required for CO burnout are approximately 30% greater than those for air combustion for the same conditions. This effect can be attributed to the much increased reverse steps of the CH4 to CH3 and CO to CO2 transformation reactions, mainly OH + CH4 ↔ CH3 + H2O and OH + CO ↔ H + CO2. However, the required lengths of both cases are much closer. (3) A configuration similar to that of lean premixed gas turbines, with moderate PZ equilibrium temperatures (around 1850 K), would get an exit CO concentration near the equilibrium CO with a size comparable to that of standard air combustion chambers. (4) The residence times and lengths would be reduced if the combustion pressure of the cycle were increased. This effect is related to the increased net rates of CH4 to CH3 and CO to CO2 conversion. The residence times and lengths would also be reduced if the combustion exit temperature were increased, with a smaller residence time inside the DZ, due to the higher COeq,DZ.

(E4)

where DZ0 is the station at the beginning of the DZ and the density has been assumed to remain approximately constant along the DZ. While this approximation again captures well the effect of P3t on tres, it fails to predict the effect of Teq,PZ for constant Teq,PZ, as can be seen in Figure 17. While the full

Figure 17. Variation with Teq,PZ of each term of E4 with respect to its value for Teq,PZ = 1850 K, for P3t = 30 bar, for the oxy−fuel cases compared with the result of the full simulation model.

model predicts a greater residence time inside the DZ for increasing Teq,PZ, which was expected since the temperatures and [COeq,DZ] are practically the same for all cases, the initial [CO] is that of equilibrium at Teq,PZ and greatly increases with Teq,PZ, E4predicts a residence time decreasing with Teq,PZ. This error can be attributed to the enhanced ωCO,DZ0 due to the greater initial CO concentration and the linear behavior of E4, which is very far from the underlying PFR true model. The behavior of the DZ of the cases with varying T4t for a fixed Teq,PZ is more complex. Although the PZ and SZ are the same, the amount of recirculated gas needed to achieve T4t is different for each case, so that the temperature, molar density, and composition of each case are different. Again, the CO oxidation process is dominated by R99; although its rate coefficients increase with T4t, the coefficient for the reverse reaction increases faster, giving a lower net rate for the DZ0 conditions, but this effect is apparently compensated by the much higher [CO]/[COeq,DZ] ratio.

■

AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. ORCID

Vicente P. Timón: 0000-0002-1217-6834 Funding

This research has been partially funded by the Spanish Framework programs (CENIT) under the OPENAER research project, led by ITP, specifically under WP 1.3 “an evaluation of industrial gas turbines for cero CO2 emissions: oxy−fuel combustion”, led and granted by ENDESA Generación SA, and Universidad Politécnica de Madrid (UPM). Notes

10. CONCLUSIONS This paper has compared standard gas turbine combustion chambers and CO2-diluted oxy−fuel combustion chambers for the SCOC-CC at a preliminary design level. The design variables of the cycle amplify the oxygen content inside the chamber, giving an actual equivalence ratio much lower than those usually associated with diluted oxy−fuel combustion. A simplified oxy−fuel combustion chamber model consisting of a well-stirred reactor (WSR) followed by a plug flow reactor (PFR) with three independent input flows have been used to compare air and oxy−fuel combustion. The adiabatic equilibrium temperature (Teq) has been used to characterize the composition at any station inside the chamber and the incipient lean blow-out equilibrium temperature (Teq,LBO) to fix the WSR volume. A point injection of the remaining recirculating gas or excess air in the station after the CO burnout in the primary zone has been selected as representative of the real conditions. This model has been applied to a feasible design point of a power production cycle (combustion exit temperature 1600 K,

The authors declare no competing financial interest.

■

REFERENCES

(1) Corchero, G.; Timón, V. P.; Montañeś , J. L. A natural gas oxyfuel semiclosed combined cycle for zero CO 2 emissions: a thermodynamic optimization. Proc IMechE part A. Proc. Inst. Mech. Eng., Part A 2011, 225, 377−388. (2) Franco, F.; Mina, T.; Woolatt, G.; Rost, M.; Bolland, O. Characteristics of Cycle Components for CO2 Capture. Proceedings of 8th International Conference on Greenhouse, Gas Control Technologies, Trondheim, Norway; 2006. (3) Tan, Y.; Douglas, M. A.; Thambimuthu, K. V. CO2 capture using oxygen enhanced combustion strategies for natural gas power plants. Fuel 2002, 81, 1007−1016. (4) Andersson, K.; Johnsson, F. Flame and radiation characteristics of gas fired O2 CO2 combustion. Fuel 2007, 86, 656−668. (5) Ditaranto, M.; Hals, J. Combustion instabilities in sudden expansion oxy−fuel flames. Combust. Flame 2006, 146, 493−512. (6) Nemitallah, M. A.; Habib, M. A. Experimental and numerical investigations of an atmospheric diffusion oxy-combustion flame in a gas turbine model combustor. Appl. Energy 2013, 111, 401−415. 11360


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