Study of Lean Premixed Methane Combustion with CO2 Dilution

Article pubs.acs.org/EF

Study of Lean Premixed Methane Combustion with CO2 Dilution under Gas Turbine Conditions Stéphanie de Persis,*,† Gilles Cabot,‡ Laure Pillier,† Iskender Gökalp,† and Abdelakrim Mourad Boukhalfa‡ †

ICARE-CNRS, 1C, Avenue de la Recherche Scientifique, 45071 Orleans Cedex 2, France CORIA-CNRS-INSA et Universite de Rouen, St Etienne du Rouvray, France

‡

ABSTRACT: Among the existing processes for CO2 capture in postcombustion, membrane separation technologies are of great interest because they can provide separation at a greatly reduced energy cost. This technology can only be envisaged, however, if a relatively high content of CO2 is present in the exhaust gas. Exhaust gas recirculation (EGR) is one solution to increase the CO2 content. In this paper, in a first approach, EGR composed only of CO2 was studied. The study of lean premixed methane combustion with CO2 dilution in gas turbine conditions was carried out through an experimental approach performed in a model gas turbine chamber coupled to a kinetic approach. Modeling was carried out using the CHEMKIN package and Cantera1.7, together with the Gas Research Institute Reaction Mechanism Release 3.0, in order to simulate the combustion conditions in terms of burning velocity, temperature, and pollutant emissions (NO and CO) required for proper operation of the system. Particular attention was paid to the case of CO emissions.

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INTRODUCTION Carbon dioxide capture and storage (CCS) is a promising option for the reduction of greenhouse gas emissions from fossil-fuel power plants.1 Much of the research in this area focuses on minimizing the energy required for CO2 capture.2,3 For integrated coal gasification combined cycle or natural gas combined cycle power plants, precombustion capture offers the most promising alternative, whereas postcombustion capture is usually the only option for other industrial plants (cement, steel, glass, refineries, chemical plants, etc.). In postcombustion processes, the capture step dominates and typically accounts for 60−80% of the overall cost of the capture−transportation− injection−storage chain.4 The aim of the research conducted at present concerns mainly the reduction of this cost. CO2 capture in a postcombustion process is currently envisaged5 either by oxycombustion (i.e., with a high purity oxygen supply to avoid dilution with nitrogen, leading to an “easy” capture of concentrated CO2 after condensation of the water vapor) or by CO2 capture in the exhaust gas of a conventional process with air supply (in this case, CO2 concentration is usually between 4 and 15% depending on the fuel type and the combustion process). In the latter case, absorption in liquids (amine washing), adsorption processes, cryogenic fractionation, or membranes can be used as separation processes. Among these processes, postcombustion membrane separation affords a greatly reduced energy cost6 provided that a relatively high content of CO2 is produced in the exhaust gas (>30% at 1 atm). However, this capture process cannot be directly applied to the most competitive power plant technique (combined cycle), which combines gas and steam turbines. In gas turbines, high air dilution is used in order to keep the turbine inlet temperature (TIT)7 below the metallurgical temperature limit of the first turbine stages. Unfortunately, this dilution leads to a lower CO2 concentration of close to 5% in the exhaust gases. One solution to increase the © 2012 American Chemical Society

CO2 concentration, called EGR (exhaust gas recirculation), is to replace (totally or partially) air dilution by cooled exhaust gases. Moreover, EGR can be used to decrease NOx emissions by controlling the adiabatic flame temperature, a method already used in car engines, and to decrease the compression work by lowering the compressed air mass flow rate. Evulet et al.8,9 confirmed the feasibility of EGR for enhanced CO2 capture in conditions similar to real gas turbine conditions. They performed an experimental study on a staged combustion rig with natural gas as fuel in lean combustion conditions and measured CH4, CO, CO2, NOx, and O2 concentrations in the burned gases. Their study demonstrated high combustion efficiency and lower CO and NOx emissions using EGR. This study was completed by a numerical approach.10 Rokke and Hustad11 performed an experimental study on a 65 kW gas turbine combustor, investigating the effect of adding N2, CO2, and O2 in the combustion process, focusing on flame stability and NOx emissions. They proposed a correlation for each additive based on the effect on NOx emissions and a final correlation for the full EGR process. However, enhancement by EGR only allows the increase of CO2 up to the stoichiometric concentration (i.e., 11.5% with methane) but not to the minimum required for membrane separation processes (i.e., 30% target). The present work focuses on the investigation and understanding of the effect of CO2 addition in a methane/air premixed flame in order to improve capture conditions. This assumption allows us to (i) decouple the effects of the exhaust gas recirculation and focus on the CO2 effect; Received: October 4, 2012 Revised: December 28, 2012 Published: December 29, 2012 1093

dx.doi.org/10.1021/ef3016365 | Energy Fuels 2013, 27, 1093−1103

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gases) were obtained using a PG 250 HORIBA gas analyzer. The suction probe of the exhaust gases was placed at the cone exit, as shown in Figure 1. NO x emissions were detected by a chemiluminescence sensor and CO and CO2 emissions by a nondispersive infrared (NDIR) sensor. The O2 measurement was based on a zirconia method using a galvanized cell. The measurement uncertainties were 0.1% for O2 and CO2, 1 ppm for NOx, and 5 ppm for CO. The flame structure was characterized by detecting CH* chemiluminescence with an ICCD camera and a BG5 filter. Average and root-mean-square parameters were obtained from a sample of 300 images.

(ii) determine the modeling methodology, which will be applied when more realistic dry EGR (composed of N2, CO2, O2, CO) is used. Reducing EGR to CO2 is commonly done in the literature, mainly because although N2 is the major component of EGR gases, its thermophysical properties such as heat capacity and diffusion coefficients are close to those of air but are significantly different from those of CO2. In the present study, we therefore focus on the effects of CO2 dilution instead of N2. Particular attention is paid to the combustion conditions (flame speed and temperature) and to the pollutant emissions (NO and CO). It is important to note that the case of CO emissions is only rarely addressed in the literature;12 this will be extensively studied and discussed in the present paper. To achieve this goal, two approaches performed in conditions similar to those encountered in gas turbines were developed: • an experimental approach carried out in a model gas turbine chamber; • a kinetic approach carried out using CHEMKIN13 and Cantera 1.7,14 together with release 3.0 of the Gas Research Institute Reaction Mechanism,15 in order to compute the combustion conditions in terms of burning velocity, temperature, and pollutant emissions required for proper operation of the system.

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MODELING To reduce NOx formation, many combustors, including gas turbine combustors, tend to operate in the lean premixed combustion mode. In order to model the chemical processes in gas turbine conditions, lean premixed flame simulations were performed with either the PREMIX code 18 from the CHEMKIN package13 or the Cantera 1.7 package.14 The following assumptions were used: • one-dimensional premixed CH4/N2/O2/CO2 laminar flames (CO2-added methane/air flames); • isobaric combustion; • ambient (To = 293 K) or preheated (To = 600 K) inlet gas temperatures. These assumptions were based on the work by Lafay et al.19 with low CO2 dilution of CH4 flames, using the same experimental setup, since Lafay et al. 19 demonstrated correlations between the turbulent flame structure and the calculated laminar flame speed. PREMIX18 is a FORTRAN computer program for the calculation of propagation velocity, temperature profiles, and species mole fraction profiles in premixed laminar flames. CANTERA 1.714 is a suite of object-oriented software tools for problems involving chemical kinetics, thermodynamics, and/or transport processes. The version using Python was employed. The governing equations of the problem solved by these codes are the conservation of energy and the conservation of species (equations are detailed in ref 18). PREMIX18 and CANTERA14 are capable of predicting temperature and species profiles in two laminar premixed flame configurations. The first one is the freely propagating flame. This configuration is used to determine the characteristic laminar flame speed of the gas mixture at a specified pressure and inlet temperature. The second flame configuration is the burner-stabilized flame with a known mass flow rate. In this configuration, two cases can be considered: one where the temperature profile is known and one in which the temperature profile is determined by the energy conservation equation. In the present study, calculations carried out with PREMIX18 were performed with the freely propagating flame option and calculations carried out with CANTERA14 were performed to compute species and temperature profiles in steady-state burner-stabilized laminar flames. In both cases, multicomponent diffusion and thermal diffusion options were taken into account. It is important to note that preliminary calculations showed that CANTERA14 and PREMIX18 give similar numerical results, but that CANTERA14 converges faster than PREMIX18 in the case of burner-stabilized laminar flames. In order to study the influence of the degree of completion (pure kinetics in 0D models), the fluid flow (turbulent calculations), and the temperature, different computer codes

EXPERIMENTAL SETUP

The experimental setup presented in Figure 1 is a gas turbine combustion chamber facility; the flame is premixed and turbulent. The

Figure 1. Axial swirl injector and scheme of the transparent burner. experimental setup is composed of an air−fuel mixture supply line, an axial swirl injector with a swirl number of 0.92, and a combustion chamber equipped with full optical access to allow diagnostics. The injector is composed of six helical vanes at a 50° angle with the flow axis, placed on an 8 mm diameter center body. The outer diameter of the injector is 18 mm. The fuel and air are mixed 300 mm upstream of the dump plane. The mixture can be preheated up to 700 K using a 4 kW electrical heater. Transparency of the combustion chamber is ensured by a Herasil quality quartz tube with an internal diameter of 80 mm and a length of 250 mm. In order to avoid the inlet of air induced by the depression of the internal recirculation zone, the exit of the chamber is terminated by a stainless steel cone. The mass-flow rates of the gases composing the reactive mixture are controlled by a set of thermal mass flowmeters (Bronkorst). Further information concerning burner arrangements is provided in Cabot et al.16 and Taupin et al.17 Measurements of the O2 mole fraction and pollutant emissions (i.e., the CO2 mole fraction and NO and CO emissions in the burned 1094


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Figure 2. (a) Operability domain of the swirl burner for two different air feeding temperatures. (b) Mean flame structure evolution as a function of CO2 dilution b (ϕ = 0.93, To = 293 K). For values of b greater than 1.48, the flame is extinguished; (c) Mean flame structure evolution as a function of equivalence ratio ϕ (b = 0, To = 293 K). For values of ϕ below 0.54, the flame is extinguished.

available in the CHEMKIN-II package13 were used. SENKIN20 was employed to perform a kinetic analysis of the gas phase as a function of the residence time. PSR21 was used to compute the case of very intense turbulence and perfect mixing. SENKIN20 and PSR21 are 0D models that account for finite-rate elementary chemical reactions. SENKIN20 computes the time evolution of a homogeneous reacting gas mixture in a closed system. The model can also perform kinetic sensitivity analysis with respect to the reaction rates. In our case, a constant pressure adiabatic system was used. PSR21 can predict the steady-state of a perfectly stirred reactor characterized by a reactor volume, residence time or mass flow rate, heat loss or gas temperature, the incoming temperature and mixture composition. In a perfectly stirred model, a perfect mixture is ensured by very intense turbulence. The governing equations are a system of nonlinear algebraic relations. Finally, EQUIL22 was used to predict the thermodynamic equilibrium of an ideal gas in adiabatic and isobaric conditions. The reaction mechanism considered in the present work is the Gas Research Institute Mechanism, release 3.0 (GRImech3.0).15 GRImech3.0,15 containing 53 species involved in 325 reversible reactions, is a reference mechanism widely employed in the literature concerning methane combustion. GRImech3.015 includes NOx chemistry. The thermodynamic and transport property files provided with the mechanism were employed. Mazas et al.23 demonstrated that it can be used to

simulate CO2 and H2O-added oxygen-enriched methane flames.

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RESULTS AND DISCUSSION In this work, measurements and calculations were carried out for exhaust gas recirculation (EGR) conditions, in a first approximation, composed solely of CO2. The following proportions were used: ϕCH4 + 2*(1O2 + DR N2 + bCO2 )

where ϕ is the equivalence ratio, b is the number of added moles of CO2 for one mole of injected O2, DR is the N2/O2 ratio, denoted dilution ratio. DR is set to the air constant value (3.78). In the following, the CO2 addition will be characterized by b and XCO2, the corresponding mole fraction in the initial mixture. XCO2 is calculated from b and ϕ, as shown in eq 1: XCO2 =

2b ϕ + [2(1 + DR+b)]

(1)

All experiments in the gas turbine combustion chamber were performed under atmospheric pressure (P = 1 atm). Two injection temperatures (To) were considered: the ambient temperature (293 K) and a preheat temperature of 600 K. The major species, O2 and CO2 (%), and the pollutant emissions (i.e., NO and CO emissions) (parts per million, ppm) were measured and calculated in dry burned gases conditions. The 1095


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Figure 3. Experimental dry NO and CO emissions (in ppm) for a CO2-added methane/air flame at various equivalence ratios as a function of b at To = 293 K (left column) and at To = 600 K (right column). P = 1 atm. DR = 3.78.

By normalizing the CH* images to the same maximum intensity, Figure 2b and c shows the evolution of the mean flame structure as a function of CO2 dilution (b) and equivalence ratio (ϕ) at 293 K. For low dilution rates, the flame is stable and anchored to the center bluff-body, the reaction zone licks the wall of the combustion chamber, and no reaction is observed in the corner recirculation zone17 (CRZ). As the dilution rates increase, the flame height and the flame surface increase, and reactions progressively appear in the CRZ. For b = 0.78 (Figure 2b) and ϕ = 0.65 (Figure 2c), the reactive zone fills the whole region of interest (ROI) (D = 80 mm and H = 120 mm) and two intense reacting zones clearly appear in the CRZ. For this structure, the flame is unstable. As the dilution rate increases, the flame is lifted and moves downstream to the ROI outside, and the flame again becomes stable. Then, the flame anchors the bluff-body again, becoming

main objective of the experiments in turbulent gas turbine conditions was to determine the global stable flame conditions and emission levels in order to provide input for chemical kinetics calculations. 1. Experimental Results. Experiments were performed in the following conditions: the equivalence ratio ϕ was varied from 0.93 to 0.60 and b was varied from 0 to a maximum value corresponding to the flame blow-off (Figure 2a). It was observed that, in the richest conditions (e.g., for ϕ = 0.93), the maximum injected CO2 was b = 1.48 (XCO2 = 23%) and b = 2.25 (XCO2 = 30%) at To = 300 K and To = 600 K, respectively. For leaner conditions (e.g., for ϕ = 0.60), the maximum injected CO2 decreased to respectively b = 0.391 (XCO2 = 7%) at To = 293 K and b = 0.704 (XCO2 = 12%) at To = 600 K. 1096


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Table 1. Laminar Flame Speed (SL, cm/s) and Adiabatic Flame Temperature (T, K) for To = 293 K and To = 600 K without and with CO2 Dilutiona SL (cm/s)

T (K)

To= 293 K

a

To = 600 K

To = 293 K

To = 600 K

ϕ

without CO2 b=0

with CO2 b = 0.625

without CO2 b=0

with CO2 b = 0.625

without CO2 b=0

with CO2 b = 0.625

without CO2 b=0

with CO2 b = 0.625

0.7 0.8 0.9 1

19.1 27.4 34.1 38.4

7.8 12.1 15.4 17.4

88.5 109.5 123.0 131.9

45.7 58.1 66.2 70.6

1820 1982 2113 2174

1632 1772 1887 1955

2043 2182 2285 2320

1853 1973 2063 2117

The values of b are respectively 0 and 0.625 and the equivalence ratio is varied from 0.7 to 1.0.

unstable (respectively, b = 1.32, ϕ = 0.57) before blowing-off (respectively, b = 1.48, ϕ = 0.54). In the latter two modes of flame structure, it was observed that CO emissions drastically increase. The same evolution was also observed for the case of 600 K preheating but for larger CO2 or air dilution rates (not shown here). Pollutant emissions (NO and CO) measured at the exit of the cone are reported in Figure 3 for the two inlet temperatures as a function of the equivalence ratio (ϕ) and the CO2 dilution (b). At a constant equivalence ratio, Figure 3 shows the expected decrease of NOx emissions by the flame cooling effect as the CO2 dilution increases until the lean blow-off limit (LBO). At ϕ = 0.89, NOx emissions diminish by a factor of 10 between the minimum and the maximum CO2 dilution rates. As the equivalence ratio decreases, NOx emissions are lower for the same CO2 dilution rate. The effect of preheated air temperature on NOx emissions is to double the emissions between 293 K (left column) and 600 K (right column); the same evolutions are observed with equivalence ratio and CO2 dilution. The behavior of CO emissions differs from that of NOx emissions. At a constant equivalence ratio, Figure 3 shows an overall increase in CO emissions as CO2 dilution increases. For low b values, the increase in CO emissions is low and rises drastically for higher b values. This behavior is observed whatever the equivalence ratio, and the slope change is directly linked to the change from an anchored flame to a lifted flame (see Figure 2b). As the equivalence ratio is lowered, CO emissions decrease for low b values (low CO2 dilution) and increase for high b values. The effect of air preheating is shown by a slight rise in CO emissions, which can be attributed to higher thermal dissociation. 2. Numerical Results. 2.1. Results with Freely Propagating Flame Modeling. Modeling of freely propagating flames using PREMIX18 from CHEMKIN-II13 was used to determine laminar flame parameters, such as the laminar flame velocity (SL, cm/s) and the flame temperature (T, K) of the gas mixture, as well as O2 and CO2 mole fractions and pollutant (CO and NO) emissions. In the calculations carried out in the free flame configuration, a sufficiently high maximum distance from the burner was chosen in order to reach the stationary state corresponding to the burned gases. 2.1.1. Laminar Flame Parameters. The results are reported in Table 1 as a function of the equivalence ratio for two cases: with CO2 dilution (b = 0.625, XCO2 = 10%) and without CO2 (b = 0, XCO2 = 0%) at To = 293 K and To = 600 K, under 1 atm. As expected, Table 1 shows that the coupled effect of the decrease in the equivalence ratio and the increase in the added

CO2 leads to a decrease in the laminar flame velocity. For a given equivalence ratio, both flame temperature (T) and laminar flame speed (SL) decrease when the CO2 dilution rate increases. This result is in good agreement with experimental results obtained in the literature.24 For a given fraction of added CO2, adiabatic flame temperature and laminar flame speed increase as the equivalence ratio increases. The same variations of SL and T are observed when the inlet mixture is preheated. For an inlet preheating from 293 to 600 K, the adiabatic temperature varies by about 200 K while the laminar flame speed is increased 4- to 5-fold. 2.1.2. O2 and CO2 Mole Fractions. Table 2 shows the calculated and measured O2 and CO2 mole fractions in the dry Table 2. Calculated and Experimental Dry O2 and CO2 Mole Fractions for a CO2-Added Methane/Air Flame As a Function of b. To = 293 K, P = 1 atm, ϕ = 0.80, DR = 3.78 dry O2 mole fraction (% mol)

dry CO2 mole fraction (% mol)

b

XCO2 (% mol)

calc.

meas. (±0.1%)

calc.

meas (±0.1%)

0 0.313 0.625 0.938

0 5.70 10.76 15.33

4.71 4.54 4.09 3.73

4.37 4.05 3.87 3.19

8.86 14.80 19.90 25.90

8.70 14.50 19.90 25.40

burned gases (computations with the freely propagating flame option in conjunction with GRIMech3.0) for a CO2-added methane/air flame at ϕ = 0.80 and DR = 3.78 as a function of four values of b (from 0 to 0.938; i.e., respectively: 0%, 5.70%, 10.76%, and 15.33% of CO2 in mole fraction) (To = 293 K). An overall good agreement between calculations and experiments is observed. This agreement is obtained whatever the equivalence ratio. At a given equivalence ratio, O2 concentration in the exhaust gas decreases when the CO2 dilution rate increases. On the contrary, increasing b (XCO2) makes it possible to increase the CO2 concentration in the exhaust gas. Table 2 suggests that the mole fraction of CO2 can be multiplied by a factor of 2.92 with b = 0.938. This shows that EGR could be a potentially effective method for increasing CO2 concentration in exhaust gas. Moreover, EGR could compete with other technologies. In a recent study, Li et al.25 compared four technologies enabling increased CO2 concentration: exhaust gas recirculation (EGR), humidification (EvGT), supplementary firing (SFC), and external firing (EFC). They showed that EGR has the greatest ability to change the CO2 mole fraction, as it can be multiplied by 2.6 in the case of stoichiometric combustion. 1097


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Figure 4. Experimental (left-hand side scale) and calculated (right-hand side scale) pollutant emissions for CO2-added methane/air flame as a function of the CO2 dilution ratio: (a) dry NO and (b) dry CO. To = 293 K. P = 1 atm. ϕ = 0.80. DR = 3.78.

2.1.3. NO and CO Emissions. Figure 4 shows the comparison between experimental and calculated CO and NO mole fractions in ppm in the dry burned gases as a function of b, for an equivalence ratio of 0.80, at To = 293 K and P = 1 atm. Note that, in order to compare the qualitative trends of experimental and calculated results, the data are plotted with two different scales (left-hand side scale for experimental results and right-hand side scale for calculated results). Figure 4a shows that the increase in CO2 dilution leads to a decrease in NO emissions. This trend is correctly predicted by modeling and a good quantitative agreement between experiments and modeling is observed. The NOx calculations are overestimated by only a factor of 2, which is satisfactory considering the assumptions used for calculations. The same trends are obtained for the other conditions of equivalence ratio and temperature. Figure 4b shows a large discrepancy between experimental and calculated CO emissions from both the quantitative and qualitative points of view. The trends are totally contradictory: whereas the experimental results show that CO emissions increase when the CO2 dilution rate increases, calculated results show the opposite. This very poor qualitative agreement between experimental and numerical results for CO is also observed for the other conditions of equivalence ratios and temperature. The results are not satisfactory and show that the flame configuration used here for predicting CO emissions (freely propagating flame) is ill-suited to reproduce the chemistry of our swirled and turbulent gas combustion chamber. To understand the reasons for the opposite evolutions of experimental and calculated CO emissions, more kinetic calculations are required, as shown in the next section. However, it should be pointed out that, in the literature on pollutant emissions in gas turbines, NOx is always addressed, but CO emissions are usually not discussed, indicating that CO prediction is an arduous task. 2.2. Kinetic Study. The kinetic study presented in this section aims at understanding the possible reasons for the experimental and calculated discrepancies observed previously in the case of CO. The possible effects of the degree of completion, the temperature, and the fluid flow were studied. 2.2.1. Influence of the Degree of Completion. To check the influence of the degree of completion, the simulated CO and

NO mole fraction profiles calculated with the freely propagating flame option as a function of the distance from the burner are plotted in Figure 5 for a CH4/CO2/O2/N2 flame for an equivalence ratio of 0.80 and b ranging from 0 to 0.391 (i.e., for XCO2 ranging from 0 to 7%).

Figure 5. Calculated dry (a) CO and (b) NO mole fractions as a function of the distance from the burner (in cm). PREMIX with FREE option in conjunction with GRIMech3.0 was used. To = 293 K. P = 1 atm. ϕ = 0.80. 1098


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flames, and at thermodynamic equilibrium (computed with EQUIL22). In the latter case, the temperature profile was determined by the energy conservation equation and the flow rate used was the experimental inlet flow rate divided by 100. Figure 7 shows that the temperature in the burned gases differs significantly as a function of the flame configuration.

Figure 5a clearly shows that the variations in CO mole fractions with b depend on the distance from the burner surface. There are two inflection points (denoted 1 and 2 in Figure 5a). For instance, for a distance less than 0.3 cm, the CO mole fraction increases with b, as was experimentally observed, whereas for a distance greater than 0.3 cm, the CO mole fraction decreases when b increases. This calculation takes into account both kinetics and diffusion. In order to separate the two phenomena, only pure kinetics calculations were carried out with SENKIN20 in an isobaric and adiabatic closed system. Results (not presented here) also showed the existence of a residence time in which the variations of CO as a function of b change. This demonstrates that the chemical kinetics is responsible for this behavior. As shown in Figure 5b, this effect does not affect the qualitative evolution of NO emissions, because whatever the distance from the burner, the NOx mole fractions always decrease as b increases. However the NO emissions increase with the distance from the burner. To check whether the experimental degree of completion was under- or overestimated by the calculations, thermodynamic equilibrium calculations were also carried out with EQUIL22 from CHEMKIN-II.13 From a kinetic point of view, thermodynamic equilibrium corresponds to an infinite time or an infinite distance from the burner. Figure 6 shows dry CO

Figure 7. Comparison between temperatures calculated in the case of the thermodynamic equilibrium, a free flame configuration and a burner-stabilized flame. ϕ = 0.80 and for b that ranges from 0 to 0.5. To = 293 K, P = 1 atm. GRIMech3.0 was used.

The same behavior can be observed between the free flame configuration and the thermodynamic equilibrium. It is important here to check that temperatures predicted in free flame configurations correspond to the adiabatic temperatures computed at thermodynamic equilibrium (computed with EQUIL22). The temperature in the burned gases decreases when the CO2 dilution rate increases and a 400 K decrease was observed for a CO2 dilution of b = 1. As the calculations with the FREE flame configuration show the opposite evolution to that encountered experimentally, it seems that (i) in our experimental conditions, the adiabatic temperature is not reached; (ii) the temperature variations are not the same. This could explain the drastic increase in CO which may be due to a lower conversion of CO2 as the temperature decreases when b increases. Figure 7 also shows that in a burner-stabilized flame in which the temperature is computed by the resolution of the energy conservation equation, the slope of the temperature in the burned gases is very gentle. The decrease is very small. In this case, the temperature is quasi independent of b, and the evolution of CO emissions predicted by the calculation is identical to that determined experimentally. Whatever the flame configurations, NO mole fractions always decrease with b, as it was observed experimentally. 2.2.3. Influence of the Fluid Flow. Another possible reason for the discrepancy between experiments and calculations is the effect of the turbulence flow. In the previous calculations, turbulence was not taken into account. To understand the role of turbulence, calculations were performed in a perfectly stirred reactor, in which the turbulence is infinite to ensure mixing. In our calculations, the pressure, the equivalence ratio, and the volume of the reactor were respectively 1 atm, 0.80, and 200 cm3, and two cases of mixing temperature were tested: the

Figure 6. Calculated dry NO (lines) and CO (dashed lines) mole fractions as a function of b in isobaric and adiabatic conditions. EQUIL in conjunction with GRIMech3.0 was used. To = 293 K. P = 1 atm. ϕ = 0.80.

and NO mole fractions as a function of b for an equivalence ratio of 0.80 at To = 293 K. The CO and NO concentrations decrease as a function of the CO2 dilution. This result is observed whatever the equivalence ratio and temperature inlet, showing that the experimental degree of completion is lower than that calculated in the case of a freely propagating flame configuration, which considers adiabatic conditions. This means that the calculated CO mole fractions should be considered at a lower residence time value or at a lower distance from the burner than in the experimental configuration. 2.2.2. Effect of the Temperature. In this section, the effect of the temperature on CO emissions was studied by comparing the temperature of the burned gases as a function of b obtained in different flame configurations (ϕ = 0.80 and To = 293 K): in free flame option and in steady-state burner-stabilized laminar 1099


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Figure 8. (a) Calculated dry CO mole fractions in ppm with Tf = 2000 K, (b) calculated dry CO mole fractions in ppm with T = Tadia, (c) calculated dry NO mole fractions in ppm with Tf = 2000 K, (d) calculated dry NO mole fractions in ppm with T = Tadia as a function of residence time τ and b. PSR in conjunction with GRIMech3.0 was used. P = 1 atm. ϕ = 0.80.

in the laminar case. The key parameter seems to be the temperature. Indeed, as shown previously, an infinite turbulence cannot counterbalance the effect of the temperature. Furthermore, the experimental temperature seems to be lower than the adiabatic temperature and the variation of the temperature as a function of the CO2 dilution rate seems to be smaller than that obtained in adiabatic conditions. However, as no temperature measurements could be carried out in our model gas turbine, it was not possible to experimentally determine the decrease of the temperature due to CO2 dilution. To avoid estimating the temperature in the burned gases without any experimental information, our final choice was therefore to model the gas turbine combustion chamber using a burner-stabilized flame configuration (with CANTERA1.714) in which the temperature is computed by the resolution of the energy equation conservation. The results are presented in the next section as well as the methodology employed. 2.3. Modeling of Burner-Stabilized Flames. For each condition, kinetic modeling of a burner-stabilized laminar premixed methane/air flame configuration was carried out in

adiabatic temperature (which is a function of CO2 dilution) and an arbitrary constant value (2000 K) corresponding to the maximum temperature encountered in these conditions (b = 0, ϕ = 0.8). The results of CO and NO emissions are plotted in Figure 8 as a function of the residence time (τ). Figure 8a and b shows that CO emissions decrease as the residence time increases. For the adiabatic temperature case (Figure 8b), the evolution of the CO mole fraction is similar to that of Figure 5a. For a low residence time (low distance from the burner, cf. Figure 5), the CO emissions increase as CO2 dilution increases, while for a higher residence time (high distance from the burner) this behavior inverts. In the case of constant temperature, this change of behavior is never observed (Figure 8a): CO emissions always increase as CO2 dilution increases, whatever the residence time and the equivalence ratio. NO mole fractions (Figure 8c and d) always decrease with b whatever the imposed temperature and residence time. 2.2.4. Summary. As a conclusion, taking turbulence into account does not lead to different results from those obtained 1100


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Figure 9. Calculated (lines) and experimental (symbols) dry NO in ppm for a CO2-added methane/air flame and as a function of b at (a) To = 293 K and (b) To = 600 K. CANTERA 1.7 in conjunction with GRIMech3.0 was used. P = 1 atm. ϕ = 0.80. DR = 3.78.

conjunction with the GRI mechanism,15 version 3.0 with a known mass flow rate (experimental mass flow rate divided by 100). The temperature profile was determined by the energy conservation equation and the calculations included multicomponent diffusion. Calculated and experimental O2 and CO2 mole fractions (not presented here) were compared and a good agreement between calculations and experiments was obtained whatever the equivalence ratio and the inlet temperature. Figure 9 shows the calculated and experimental NO emissions as a function of b, for an equivalence ratio ϕ = 0.8, and for the two temperatures To = 293 K (left-hand side) and To = 600 K (right-hand side). Whatever the inlet temperature, the calculations successfully reproduced the decrease in NO emissions as the CO2 dilution increases. However, the predicted slope of the NO decrease is less pronounced than in the experimental condition, and the shape of this decrease is less well reproduced than in the case with a freely propagating flame (see Figure 4a). This is probably due to the temperature profiles used in our calculations, which are only marginally sensitive to CO2 dilution. From a quantitative point of view, the orders of magnitude between experimental and calculated dry NO are similar. For the leanest conditions, NO emissions predicted by simulations appear to represent a maximum value compared to the experimental ones. On the contrary, for the richest conditions, the predicted value seems to be a mean value compared to the experimental ones. For comparison, experimental laser induced fluorescence (LIF) NO measurements obtained by Thomsen et al.26 in the burned gases of a counterflow premixed methane/air flame are reported in Figure 10 as a function of the equivalence ratio. The calculations show good agreement between both experimental results. Figures 9 and 10 thus demonstrate that the present calculations can be used as a predictive tool in order to quantify the amount of NO in the burned gases of a model gas turbine chamber. As shown in Figure 11, where two different scales are used to plot the results, only the qualitative trends of experimental CO

Figure 10. Calculated (lines) and experimental (symbols) NO in ppm for a methane/air flame at various equivalence ratios. CANTERA 1.7 in conjunction with GRIMech3.0 was used. DR = 3.78. To = 293 K. P = 1 atm.

emissions are reproduced by the simulations. Both experimental and calculated results show that CO emissions increase with CO2 dilution. From a quantitative point of view, experimental and calculated CO emissions vary by a factor ranging from 2 to 60 depending on the equivalence ratios. The difference between experimental and calculated CO emissions increases when equivalence ratio decreases, as shown by Amato et al.12 For instance, in our conditions, at ϕ = 0.8 and b = 0, a factor of 30 between experiments and modeling is observed, at ϕ = 0.93 and b = 0, this factor is reduced to 23. Those discrepancies can be mainly attributed to the high sensitivity of the CO/CO2 equilibrium to temperature at low equivalence ratio, as also pointed out by Amato et al.12 In their work a factor of 30 is observed when the equivalence ratio decreases from 1.1 to 0.9. This quantitative disagreement can be minimized by the fact that computed CO emissions systematically overestimate experimental ones: simulations thus give 1101


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Figure 11. Calculated (lines, right-hand side scale) and experimental (symbols, left-hand side scale) dry CO in ppm for a CO2-added methane/air flame as a function of b at (a) To = 293 K and (b) To = 600 K. CANTERA 1.7 in conjunction with GRIMech3.0 was used. P = 1 atm. ϕ = 0.80. DR = 3.78.

used to benchmark that the effect of pressure on SL is much greater than that of b. Table 4 sums up the effect of the pressure

the upper limit and consequently, calculations could be used as a predictive tool. These results demonstrate that these calculations could be used to extrapolate NO and CO emissions to other conditions that are more difficult or impossible to investigate experimentally. 3. Extrapolation to Gas Turbine Conditions. The previous results show that calculations can be used as a predictive tool in order to quantify not only O2 and CO2 mole fractions in the burned gases of a model gas turbine chamber but also NO and CO emissions. In this section, calculations were carried out to simulate gas turbine conditions, that is, conditions with a preheated temperature of To = 600 K when the pressure is increased from 1 to 4 and 8 atm. Table 3 shows the laminar flame speed of a CO2-added methane/air flame at various b corresponding to various CO2 mole fractions (XCO2 = 0, 10, 20, and 30%) at To = 600 K and P = 1, 4, and 8 atm. As mentioned elsewhere,24 Table 3 can be

Table 4. CO and NO Emissions (in ppm) Calculated at To = 600 K; 1, 4, and 8 atm for ϕ = 0.7 as a Function of b (XCO2)a CO (ppm)

a

b=0 (XCO2 = 0%)

b = 0.570 (XCO2 = 10%)

b = 1.281 (XCO2 = 20%)

b = 2.197 (XCO2 = 30%)

1 atm 4 atm 8 atm

88.48 48.40 33.93

45.71 22.94 15.58

9.76 6.85

-

ϕ = 0.8

b=0 (XCO2 = 0%)

b = 0.576 (XCO2 = 10%)

b = 1.296 (XCO2 = 20%)

b = 2.221 (XCO2 = 30%)

1 atm 4 atm 8 atm

109.50 63.35 45.7

58.08 30.93 21.23

28.44 13.64 9.269

4.12

ϕ = 0.9

b=0 (XCO2 = 0%)

b = 0.583 (XCO2 = 10%)

b =1.311 (XCO2 = 20%)

b = 2.248 (XCO2 = 30%)

1 atm 4 atm 8 atm

123.40 75.25 55.33

66.17 36.98 25.78

33.49 16.45 11.17

13.50 6.00 4.56

b

XCO2

1 atm

4 atm

8 atm

1 atm

4 atm

8 atm

0 0.078 0.235 0.313 0.391 0.467 0.625 0.704

0.00 1.50 4.38 5.75 7.09 8.38 10.87 12.07

36.43 46.48 69.23 82.15 96.21 111.67 144.89 165.87

3.63 4.66 5.94 6.70 7.39 8.13 9.59 10.16

1.34 1.63 2.17 2.45 2.72 2.98 3.48 3.74

3.62 3.59 3.52 3.49 3.46 3.43 3.38 3.35

1.47 1.46 1.42 1.41 1.39 1.38 1.35 1.33

0.65 0.64 0.62 0.61 0.61 0.60 0.59 0.58

CANTERA 1.7 in conjunction with GRIMech3.0 was used.

on CO and NO emissions for ϕ = 0.7 as a function of XCO2. Table 4 clearly shows that an increase in the pressure leads to a decrease in NO and CO emissions. O2 and CO2 mole fractions (not shown here for reasons of clarity) are nearly constant as a function of the pressure for a given CO2 dilution rate. For instance, when the CO2 dilution rate is equal to b = 0.625 (i.e., XCO2 = 10.87%), the O2 mole fraction is always 6.12% and the CO2 mole fraction is 19.17% in the burned gases. The results reported by Thomsen et al.26 showed a lesser effect of the pressure on LIF NO measurements. The peak NO concentrations measured by Thomsen et al.26 in the burned gases of a counterflow premixed methane/air flame, ϕ = 0.60, 0.65, 0.75, and 0.70, showed that NO formation was not strongly influenced by pressure. For all the cases studied by Thomsen et al.,26 the total spread in the data was less than 30% of the peak value. The strongest pressure sensitivity occurred at ϕ = 0.75.

Table 3. Laminar Flame Speed (SL, cm/s) of a CO2-Added Methane/Air Flame at XCO2 = 0%, 10%, 20%, and 30%, To = 600 K at 1, 4, and 8 atm at ϕ = 0.7, 0.8, and 0.9a ϕ = 0.7

NO (ppm)

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CONCLUSION In this paper, the effect of CO2 addition was studied in CO2added methane/air flames. This study is a first approach to the study of the dry EGR (exhaust gas recirculation) effect. It was

a

PREMIX with FREE option in conjunction with GRIMech3.0 was used. The symbol - means that no convergence was obtained due to too low flame velocity value (below 10 cm/s). 1102


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(10) Li, H.; ElKady, A. M.; Evulet, A. T. AIAA 2009, 226, 2361− 2372. (11) Rokke, P. E.; Hustad, J. E. Int. J. Thermodyn. 2005, 8 (4), 167− 174. (12) Amato, A.; Hudak, B.; Souza, P. D’; Carlo, P. D’; Noble, D.; Scarbourough, D.; Seitzman, J.; Lieuwen, T. Proc. Combust. Inst. 2010, 33, 3399−3405. (13) Kee, R. J.; Rupley F. M. and Miller J. A. Chemkin-II: A Fortran Chemical Kinetics Package for the Analysis of Gas Phase Chemical Kinetics, Sandia Report SAND89-8009B; Sandia National Laboratory: Albuquerque, NM, 1989. (14) Cantera 1.7; Cantera Developers: 2006; http://cantera.github. com/docs/sphinx/html/index.html (15) Smith, G. P.; Golden, D. M.; Frenklach, M.; Moriarty, N. W.; Eiteneer, B.; Goldenberg, M.; Bowman, C. T.; Hanson, R. K.; Song, S.; Gardiner , W. C., Jr.; Lissanski, V. V.; Qin, Z.GRI-Mech, Ver. 3.0; http://www.me.berkeley.edu/gri_mech. (16) Cabot, G.; Vauchelles, D.; Taupin, B.; Boukhalfa, A. Exp. Therm. Fluid Sci. 2004, 28 (7), 683−690. (17) Taupin, B.; Martins, G.; Cabot, G.; Boukhalfa, A. Combust. Sci. Technol. 2007, 179 (1−2), 117−136. (18) Kee, R. J.; Grcar, J. F.; Smooke, M. D.; Miller, J. A. A Fortran Program for Modeling Steady Laminar One-Dimensional Premixed Flames, Sandia Technical Report SAND85-8240; Sandia National Laboratory: Albuquerque, NM, 1985. (19) Lafay, Y.; Taupin, B.; Martins, G.; Cabot, G.; Renou, B.; Boukhalfa, A. Exp. Fluids 2007, 43 (2−3), 395−410. (20) Lutz, A. E.; R.J. Kee R. J. and J.A. Miller, J. A. Senkin: A Fortran Program for Predicting Homogeneous Gas Phase Chemical Kinetics with Sensitivity Analysis, Sandia National Laboratories Report, SAND878248 UC-4; Sandia National Laboratory: Albuquerque, NM, 1988. (21) Glarborg, P.; Kee, R. J.; Grcar, J. F.; Miller, J. A. PSR: A FORTRAN Program for Modelling Well-Stirred Reactors, SAND868209; Sandia National Laboratories: Albuquerque, NM, 1992. (22) Reynolds, W. D. STANJAN, Ver. 3; Stanford University: Stanford, CA, 1986. (23) Mazas, A. N.; Lacoste, D. A.; Schuller, T. Proc. ASME Turbo Expo, Glasgow, U.K., 2010. (24) Cohe, C.; Chauveau, C.; Gökalp, I.; Kurtulus, D. F. Proc. Combust. Inst. 2009, 32, 1803−1810. (25) Li, H.; Ditaranto, M.; Berstad, D. Energy 2011, 36, 1124−1133. (26) Thomsen D. D. Measurement and modeling of nitric oxide formation in counterflow, premixed, CH4−O2−N2 flames. PhD thesis, Purdue University, West Lafayette, IN, 1999.

shown that CO2 dilution could be an efficient method for increasing CO2 concentration in exhaust gas, thus making its capture easier. In this work, special attention was paid to the kinetic approach and in particular to CO emissions. A modeling methodology was developed to simulate experimental results obtained in the model gas turbine chamber. In this calculation, release 3.0 of the Gas Research Institute Reaction Mechanism15 was employed. PREMIX13 in free flame configurations was used in order to simulate the combustion conditions in terms of burning velocity and adiabatic temperature. The pollutant emissions were simulated by CANTERA14 in the burnerstabilized flame configuration and successfully compared to experimental measurements. The comparison between experimental and calculated results shows the predictive capability of the simulations performed in this paper not only for the major species (O2 and CO2 mole fractions) but also for the NO pollutant emissions. The CO emissions remain difficult to estimate; however, we can now predict qualitatively the trends of CO emissions as a function of CO2 dilution but with a systematic overestimation. This study is a preliminary effort to assess the feasibility of a low cost CO2 capture process in EGR combustion of fossil fuels. In future papers, the effect of both the initial mixture and the amount of EGR, as well as the effect of oxygen-enriched combustion, will be presented. The global objective of this future paper will be to demonstrate that oxygen-enriched air combustion of fossil fuels in association with CO2 capture by membrane processes could be an innovative solution for low cost CCS technologies.

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AUTHOR INFORMATION

Corresponding Author

*Phone: +33 2 38 25 54 88. Fax: +33 2 38 69 60 04. E-mail: [email protected] . Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors thank the Energy Program of the CNRS for supporting their work (COCASE program). The authors are very grateful to E. Favre and J. M. Most for helpful discussions.

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REFERENCES

(1) Herzog, H. Environ. Sci. Technol. 2001, 35 (7), 148−153. (2) Deschamps, P.; Pilavachi, P. A. Oil Gas Sci. Technol. 2004, 59 (3), 323−330. (3) Wall, T. F. Proc. Combust. Inst. 2007, 31, 31−47. (4) Davidson, O.; Metz, B.; de Coninck, H.; Loos, M.; Meyer, L. IPCC Special Report: Carbon Dioxide Capture and Storage; International Panel on Climate Change/Cambridge University Press: Geneva, Switzerland/Cambridge, U.K., 2005; http://www.ipcc-wg3.de/specialreports/special-report-on-carbon-dioxide-capture-and-storage. (5) Figueroa, J. D.; Fout, T.; Plasynski, S.; McIlvried; Srivastava, R. D. Int. J. Greenhouse Gas Control 2008, 2, 9−20. (6) Bounaceur, R.; Lape, N.; Roizard, D.; Vallières, C.; Favre, E. Energy 2006, 31, 2256−2570. (7) Soares, C. Gas Turbines: A Handbook of Air, Land, and Sea Applications; Elsevier: San Diego, CA, 2008; ISBN 978-0-7506-7969-5. (8) ElKady, A. M.; Evulet, A. T.; Brand, A. R.; Ursin T. P.; Lynghjem, A. ASME Turbo Expo 2008: Power for Land, Sea and Air, Berlin, Germany, June 2008; GT2008-51152. (9) Evulet, A. T.; ElKady, A. M.; Brand, A. R.; Chinn, D. Energy Procedia 2009, 1 (1), 3809−3816. 1103


Study of Lean Premixed Methane Combustion with CO2 Dilution

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