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Chemical reactor modelling of oxy-fuel combustion chamber for semi-closed combined cycle. Vicente Paul Timón, Gregorio Corchero, and José Luis Montañés Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.7b01311 • Publication Date (Web): 28 Aug 2017 Downloaded from http://pubs.acs.org on August 29, 2017
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Chemical reactor modelling of oxy-fuel combustion chamber for semi-closed combined cycle. Vicente P. Timón*, Gregorio Corchero, José L. Montañés. 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 oxyfuel combustion chambers for a semi-closed combined cycle, at a preliminary design level. To this end, simple chemical reactor networks, based on the Well Stirred 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 behaviour 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 were 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 Semi-Closed 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 semi-closed 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 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 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
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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 authors3, 4. Their results are similar, with a 28% by volume O2, and a 2% excess O2, for natural gas fuel3; and 27% O2, 3.8% excess O2, for propane fuel4. 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 combustor5. 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 content6. As has been pointed out8, 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 semi-closed 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. 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 kinetics9-17. Besides, the thermal and transport properties of CO2 are different to that of air. Hence, the temperatures, cooling flow distribution, combustion stability, and the length required for complete combustion will be different to that of air combustion. For example, Amato et al9 have compared N2 and CO2 diluted combustion in an atmospheric pressure, swirl stabilized combustor, showing that stable CO2 diluted combustion requires significantly higher temperatures that air combustion at a given reactants velocity, while Shroll et al17 have 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 numerically8, 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 properties3, 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 study1. Afterwards, 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 amplify the oxidant content, giving much lower actual equivalence ratios and excess oxygen values closer to that reported by other authors. The 1-3 issues presented above will then be studied by means of a simplified model of an ideal oxyfuel 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 behaviour 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 the CO burnout in the Primary Zone is selected as representative of the real conditions. Afterwards, this model will be applied to the conditions of the SCOC-CC design point obtained from a parametric analysis1. 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
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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 1850 K), would get an exit CO concentration near to that of equilibrium. Additional results will be presented, taking into account the possible changes in the primary design variables of the SCOC-CC, 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.
2. DESCRIPTION OF THE SYSTEM. Figure 1 shows a scheme of the Semi-Closed 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 semi-closed, 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 none 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 remainder 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 modelled 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 article1, 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 cycles 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 non-hydrocarbon species, and it is difficult to remove all the H2O by condensation. Since they are required to model the behaviour 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 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 non-dimensional form, dividing them by the air mass flow.
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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. 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: mfuel = Φ farst mair In a similar way, the composition at any station of the gas turbine of the Semi-Closed 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 nondimensionalise 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 = φ fst 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, approximately1. 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 SCOC-CC is a Recirculated Gas (RG) which consists in 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 = r mASU 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 to obtain r at this station. Although the φ, r, representation has been found to be adequate for the 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 YO,RG = (1 - φ) (1 - wRG) / ( 1 - φ - φ ( 1 - fst) (1 - wst) ), 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,RG r) mASU and the actual equivalence ratio is Φ = φ / (1 + YO,RG r). 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,RG r ) - 1 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 Hals5; 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 non-dimensionalize the others, the fuel mass flow can be given
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by an almost arbitrary cycle equivalence ratio, φ, which can be taken 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, 1.15% N2. The end of combustion temperature T4t = 1600 K, gives r = 8.3, Φ = 0.8, 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 a 2% and a 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 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 authors3-5, 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 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 of 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 modelling assumptions are given below; the reader is referred to the specialized literature (e.g.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 micro-mixing, 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 chamber2729 , 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) characteristics30. 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 WSRs31, allowing the 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. All calculations will be made using the python interface of the Cantera package24 and the reaction mechanism GRI-Mech 3.025, 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 results33. Taking into account the air or RG mass flow distribution inside the combustion chamber, two different configurations can be considered34: 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
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continue to burn towards 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 modelled 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 combustion37, 38. It has also been used before to simulate oxy-fuel combustion for some different design conditions39, 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 accuracy40. 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 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 modelled 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. The 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 additional gas injection will be called Secondary Zone (SZ), and the rest of the PFR will be called 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 turbines41. Then, assuming that the static and total conditions are approximately equal, a reference area Aref can be calculated as Aref = m31 √(R31 T31 )/(Mref 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; 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 LPZ = VPZ / Aref and LPFR = VPFR / Aref 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 the 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 homogenous combustion above 1300 K42, or near the liner cool-wall boundary layer43. CO quenching has been related to the presence of film-cooling or fuel-air nonuniform distribution41, 43. In type 2 configurations, the reactions can be considered to proceed towards 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 non-chemical kinetics related effects are similar.
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6. COMBUSTION CHAMBER MODEL BEHAVIOUR 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.
6.1. Primary Zone. A chemical time can be associated to combustion inside a WSR; when the residence time inside the reactor is lower than this chemical time, stable combustion is not possible and 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
