Measurements and analysis of CO and O 2 emissions in CH 4 /CO 2 /O 2 flames

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1 Available online at Proceedings of the Combustion Institute 33 (2011) Proceedings of the Combustion Institute Measurements and analysis of CO and O 2 emissions in CH 4 /CO 2 /O 2 flames A. Amato, B. Hudak, P. D Souza, P. D Carlo, D. Noble, D. Scarborough, J. Seitzman, T. Lieuwen Ben T. Zinn Combustion Laboratory, Georgia Institute of Technology, Atlanta, GA, USA Available online 6 October 2010 Abstract Concerns about green house gas emissions have encouraged interest in hydrocarbon combustion techniques that can accommodate carbon dioxide capture and sequestration. Oxy-fuel combustion, where the fuel is combusted in oxygen diluted with steam or CO 2, is one promising approach for post-combustion carbon capture. In this paper we focus on CO 2 dilution effects and, in particular, on CO and O 2 emissions from these flames. The emissions issue must be considered from a different perspective than conventional power plants as the combustor effluents will be sequestered, and, thus, their interactions with the terrestrial atmosphere are not relevant. CO and O 2 are of interest for these systems as their presence in the exhaust stream represents wasted fuel and oxidizer. In addition, CO 2 pipeline specifications impose limitations on CO and O 2 levels, which also must then be controlled either through the combustion process or post gas cleanup. Equilibrium and kinetic modeling of CH 4 /O 2 /CO 2 combustion systems was performed in order to analyze CO 2 dilution effects upon CO and O 2 emissions level. Companion experiments were also performed in an atmospheric pressure, swirl stabilized combustor. These numerical and experimental results demonstrate the key tradeoffs associated with optimizing these systems, as well as the dependence of emissions on stoichiometry, pressure, CO 2 dilution and residence time. Ó 2010 The Combustion Institute. Published by Elsevier Inc. All rights reserved. Keywords: Oxy-fuel; Carbon dioxide; Carbon capture; Carbon monoxide; Oxygen 1. Introduction Corresponding author. Address: Ben T. Zinn Combustion Laboratory, 635 Strong Street, Atlanta, GA 30318, USA. address: aamato3@mail.gatech.edu (A. Amato). Concerns about green house gas emissions have encouraged interest in hydrocarbon combustion techniques that can accommodate capture and sequestration of carbon dioxide. This can be done by removing carbon by [1] (1) pre-combustion capture via reforming of hydrocarbon fuels into a high hydrogen mixture, or (2) post-combustion capture of CO 2, which can be accomplished with amine membranes (MEA), solvents or condensation processes. This CO 2 can then be sequestered in the ocean or deep saline aquifiers, or used for enhanced oil recovery (EOR) or enhanced coal bed methane recovery (ECBM) [2]. The work described here is directed toward post-flame capture. In such cases, the combustion process typically involves diluted oxygen rather than air, in order to avoid the large volumes of N 2 which would otherwise need to be dealt with. The flame temperature is controlled by diluting /$ - see front matter Ó 2010 The Combustion Institute. Published by Elsevier Inc. All rights reserved. doi: /j.proci

2 3400 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) the oxygen with either steam or CO 2. This largely eliminates NO x formation and enables the exhaust stream to be separated into concentrated CO 2 and water by a simple condensation process. Several gas turbine power cycles utilizing oxy-fuel combustion have been proposed that either consist of slight modifications of existing gas turbine combined cycles [3], completely new concepts (e.g., Graz [4] or Matiant [5] cycles) and integration with gasification processes for coal combustion [6]. Our objective in this study is to focus on the O 2 /CO 2 problem. Using CO 2 as a diluent to burn natural gas impacts the flame through changes in: (1) mixture specific heat and adiabatic flame temperature, (2) transport properties (thermal conductivity, mass diffusivity, viscosity) [7] (3) chemical kinetic rates [8 11], and (4) radiative heat transfer [12 15]. It also raises new combustion challenges for both emissions and operability [16]. The emissions issue is the focus of this paper. Emissions must be considered from a different perspective than conventional power plants as the combustor effluents will be presumably sequestered, and, thus, their interactions with the terrestrial atmosphere are not relevant. However, the system s CO and O 2 emissions are of interest because of issues associated with overall cycle efficiency and pipeline specifications. For example, the Weyburn CO 2 pipeline has specifications on O 2 (<50 ppm), CO (<0.1%), H 2 O (<20 ppm), H 2 S (<0.9%), N 2 (<300 ppm) [17], among others. Note that of these species, the combustion process can really only be used to manipulate levels of CO and O 2. CO emissions at a given flame temperature are augmented due to higher equilibrium levels as well as slower CO oxidation rates [18]. CO 2 competes with other reactions for atomic hydrogen, leading to formation of CO, mainly through the reaction CO 2 +HMCO + OH [19]. High levels of CO are, in essence, a loss in efficiency, and results in increased fuel costs for the same power output. In addition, high local CO levels leads to corrosion in piping due to the formation of iron pentacarbonyl and nickel tetracarbonyl [20]. For similar reasons, exhaust O 2 emissions are also of interest because of the cost and power consumption associated with generating O 2, as well as pipeline corrosion concerns discussed earlier [1]. Diluted oxy-combustion systems have a different degree of freedom from conventional air systems, since flame temperature and fuel/air ratio are decoupled. Such systems will be operated at or near stoichiometric to minimize both fuel and O 2 usage, with an O 2 /CO 2 ratio controlled by temperature limitations. Figure 1 plots the two key degrees of freedom of this problem, stoichiometry and flame temperature. As discussed later, high stoichiometries and flame temperatures lead to Fig. 1. Operational space of a premixed, CH 4 /O 2 /CO 2 flame in the equivalence ratio flame temperature space. excessive CO emissions, while low stoichiometries and high flame temperatures with excess O 2 emissions. As such, lower temperature operation considerably broadens the fuel/air ratio sensitivity of these emissions. However, these minimum emissions at low temperature operation are limited by blow-off. The rest of this paper focuses on measurements and calculations of these CO and O 2 emissions. 2. Equilibrium trends Equilibrium calculations provide a useful perspective on emissions trends for a system with a sufficiently long residence time. This section presents example calculations of these equilibrium tendencies obtained used the GasEq solver [21], at pressures of p = 1 and 15 atm and initial temperature T in = 533 K. Figure 2 plots the dependence of equilibrium CO concentration upon flame temperature at several equivalence ratios. Also shown for reference is a methane/air flame, where flame temperature is varied with equivalence ratio (bottom loop is Fig. 2. Dependence of CO equilibrium concentration in wet combustion products on adiabatic flame temperature, pressure and equivalence ratio (T in = 533 K, p =1 and 15 atm).

3 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) lean conditions, top is rich). Note that CO levels are much higher for the CO 2 diluted system than the fuel lean air system at a given temperature. They are also lowest at lean conditions and increase monotonically with equivalence ratio. Increasing pressure and decreasing temperature reduce equilibrium levels for lean mixtures, as might be expected for a dissociation dominant process. These pressure and temperature sensitivities are much lower for rich mixtures, reflecting the shift away from a dissociative mechanism to the water gas shift reaction as the controlling equilibrium process at rich stoichiometries. Similarly, Fig. 3 shows that the concentration of O 2 increases with flame temperature and decreases with equivalence ratio and pressure. The pressure and temperature sensitivity to fuel/air ratio is inverted from the CO case, reflecting the fact that O 2 is a major and trace equilibrium product under lean and rich conditions, respectively, opposite that of CO. 3. Chemical kinetics CO and O 2 calculations Equilibrium calculations are useful for providing insight into limiting emissions levels that are approached given a sufficiently long residence time. In reality, the actual combustor residence time is finite and so emissions levels do not actually reach their equilibrium levels. For reference, a typical value of combustor residence time (defined by ratio of combustor volume to burned gas volume flow rate) for an F-class gas turbine is about 40 ms. As such, kinetic calculations are needed to determine the characteristic times associated with CO and O 2 evolution to determine their values as a function of the actual residence time. Calculations were performed using the PRE- MIX module in CHEMKIN 4.1 with the GRI 3.0 mechanism. Spatial profiles were converted to temporal ones by integrating the gas velocity profile across the flame. The s res = 0 point was defined as the point where the heat release rate drops to 10% of its maximum value. The key reactions controlling destruction of O 2 and CO in the post-flame zone of a CH 4 /O 2 /CO 2 flame are listed in Figs. 4 and 5, showing that they are dominated by reactions involving the OH and H radicals. Specifically, CO destruction is dominated by the CO 2 +HM CO + OH reaction. O 2 destruction is dominated by reactions involving H radical, particularly the O 2 +HM O + OH. At higher pressures, three bodies reactions become more active but these two reactions remain the fundamental ones CO emissions We first consider CO concentrations. There is a very important difference between the combustion process considered here and lean, premixed combustion that should be emphasized. This can be understood by considering the ratio between the maximum value of CO and its equilibrium value, as shown in Fig. 6. Note the large CO Fig. 4. Rate of production (ROP) of the principal reactions involving O 2 in the relaxation zone of a premixed CH 4 /O 2 /CO 2 flame (T AD = 1900 K, p = 1 atm, / = 1.0). Fig. 3. Dependence of O 2 equilibrium concentration in dry combustion products on adiabatic flame temperature, pressure and equivalence ratio (T in = 533 K, p =1 and 15 atm). Fig. 5. Rate of production (ROP) of the principal reactions involving CO in the relaxation zone of a premixed CH 4 /O 2 /CO 2 flame (T AD = 1900 K, p = 1 atm, / = 1.0).

4 3402 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) Fig. 6. CO overshoot: ratio between maximum and equilibrium CO concentration (T in = 533 K, p = 1 and 15 atm). overshoot in the flame relative to equilibrium for the fuel lean condition (/ = 0.9) at atmospheric pressure. At 1800 K, for example, peak CO levels in the flame exceed equilibrium levels by a factor of nearly 100. Even larger overshoots are present at leaner equivalence ratio. Increasing pressure increases CO overshoot values at lean equivalence ratios further. This overshoot is a major design challenge for lean, premixed systems which must achieve both acceptable CO levels at low power and NO x at higher power. For the oxy-fuel system running slightly rich, this sensitivity is not present because of the small value of the CO overshoot. While CO levels are substantially higher than lean systems, there is almost no overshoot typical values are 2 or 3. This implies substantially less sensitivity of exhaust CO levels to combustor design details. In other words, similar CO levels will be experienced regardless of quenching by walls or dilution jets. We next consider post-flame relaxation of CO emissions toward equilibrium. Figure 7 plots the dependence of CO concentration on adiabatic flame temperature, under both equilibrium and fixed residence time conditions, where s res = 40 ms. Notice the strong influence of flame temperature upon combustion relaxation time. At high temperatures (T > 2000 K), the fixed residence time and equilibrium values are the same. At lower temperatures, the fixed residence time and equilibrium CO concentrations diverge progressively as the temperature decreases. Note that in all cases, the fixed residence time result is less sensitive to temperature than the equilibrium result. To illustrate the effect of pressure, Fig. 8 plots the CO concentrations at fixed residence time and different pressures. It shows that CO concentrations at 15 atm are considerably lower than those at 1 atm especially in lean conditions, where CO concentrations are lower. This is due to the combination of two effects, (i) the lower CO equilibrium concentration (as described in the previous section) and (ii) the faster CO relaxation kinetics at higher pressures O 2 emissions We next consider the trends for O 2 concentrations. Figure 9 plots the dependence of O 2 concentration on adiabatic flame temperature both at Fig. 8. Dependence of CO concentration at s res =40ms in wet combustion products upon adiabatic flame temperature at two pressures (T in = 533 K, p = 1 and 15 atm). Fig. 7. Dependence of CO concentration in wet combustion products at s res = 40 ms and equilibrium upon adiabatic flame temperature (T in = 533 K, p = 1 atm). Fig. 9. Dependence of O 2 concentration in dry combustion products at s res = 40 ms and equilibrium upon adiabatic flame temperature (T in = 533 K, p = 1 atm).

5 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) equilibrium and at s res = 40 ms. Similar conclusions as discussed for CO can be seen here as well. An important point that should be re-emphasized is the substantially reduced temperature sensitivity of the fixed residence time results relative to the equilibrium, even more the case here than for CO. To illustrate, consider / = Equilibrium considerations suggest that O 2 emissions can be substantially reduced by decreasing flame temperature. For example, there is a two order of magnitude difference between levels at T AD = 1800 and 2200 K. The important dependence of relaxation times upon temperature substantially changes the fixed residence time result. Note that the difference between O 2 emissions at fixed residence time for these two temperatures changes from a 100 reduction to a 2 reduction. To illustrate pressure effects, Fig. 10 compares the temperature dependence of O 2 concentration at fixed residence time at two pressures. Again, similar results as in the CO case occur (see Fig. 8): there are significant differences between the two pressures due to faster relaxation times and lower equilibrium values at high pressure. 4. Experimental facility Fig. 10. Dependence of O 2 concentration at s res =40ms in dry combustion products upon adiabatic flame temperature for different pressures (T in = 533 K, p =1 and 15 atm). Fig. 11. Combustor configuration used for emissions sampling (dimensions in mm). Parallel measurements were obtained to study emissions trends in a swirl stabilized, atmospheric combustor shown in Fig. 11. This system is a duplicated from an experimental rig described by Williams et al. [22]. The facility consists of a swirler/nozzle able to operate both premixed and nonpremixed conditions, a combustor, and exhaust sections; for the emissions measurements described in this paper, only fully premixed conditions were tested. In this configuration, the fuel (CH 4 ) is injected 150 cm upstream into the preheated oxidizer O 2 /CO 2 mixture in order to achieve a premixed condition. Then, the premixed gases flow through a swirler section consisting of a six-vane, 45 swirler, which is located in the annulus between the centerbody and nozzle wall. The nozzle is an annular tube with an outer diameter of 28 mm while the center body has an outer diameter of 20 mm. The overall flow area is 3.0 cm 2 and remains constant inside the nozzle annulus. The theoretical swirl number is The combustor consists of a 508 mm long quartz tube, with a 115 and 120 mm inner and outer diameter, respectively. It rests in a circular groove in a base plate. The O 2,CO 2, and fuel flow rates are measured with mass flow meters. Reactants mixtures were preheated to 533 ± 20 K. Particular attention was paid in order to eliminate air leakage, which is an important contaminant in the low O 2 exhaust. The combustor was insulated with ceramic cloth and sealed at the bottom and top edges of the quartz tube using an 8 00 steel pipe. The 8 00 pipe features flanges on either end for attachment of the combustor base plate and an exhaust plate to allow for sealing of the quartz from the atmosphere. Without these sealing provisions, we found that atmospheric air is pulled into the quartz tube, leading to inaccurate O 2 measurements. O 2 emissions were obtained with an Advanced Instruments Inc. GPR-1200 Portable ppm Oxygen Analyzer. This O 2 analyzer has four different scales, 0 10, 0 100, ppm, and 0 25%, allowing for the full range of O 2 emissions of interest to be measured with the same device. The sensor was zeroed with pure nitrogen and calibrated at 800 ppm oxygen in a nitrogen bath calibration gas. CO emissions were obtained with a LAND Lancom Series II Combustion Products Analyzer. This analyzer was zeroed with air and calibrated with 3000 and 30,000 ppm CO in nitrogen bath calibration gases. The combustion gases were sampled with a water quenched probe in order to rapidly freeze the O 2 and CO reactions. The emissions sample was then drawn out by means of a vacuum pump; additionally, during O 2 measurements a coalescing filter was used to extract any water present in the sample. After

6 3404 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) the vacuum pump, the sample was then run to either the O 2 sensor or CO sensor. Care was taken to use all stainless steel tubing and to minimize leaks or losses in order to acquire quality O 2 measurements. Calculations using Henry s law [23] show that, assuming an average temperature of 350 K inside the probe, up to 0.3% of the oxygen can dissolve in the condensing water which would not then be measured by the O 2 sensor, a negligible error. Flame temperatures measurements were obtained using a grounded, K-type thermocouple from Omega Engineering, model number KMQXL-125G-24: the sheath material of this thermocouple is Nickel-Chrome based, providing good performance in heavily oxidizing environments. Radiation losses were minimized by inserting the diameter thermocouple into a diameter ceramic sleeve. The estimated temperature uncertainty is ±50 K. Measurements indicated typical differences between theoretical adiabatic and measured flame temperature of 300 and 400 K at the measurement point 40 cm above the combustor exit plane, due to heat losses from the combustor. No measurable sensitivity of this temperature difference to stoichiometry at a fixed adiabatic flame temperature was observed. Figure 12 plots the dependence of the measured CO emissions on temperature for different equivalence ratios. It can be observed that the measured CO emissions are in good agreement with the model at high equivalence ratios. Near stoichiometric conditions, where the sensitivity to experimental uncertainties is greater, the difference between numerical and experimental values is larger. However, the same qualitative trends are observed both for numerical and experimental measurements; i.e., the CO emissions monotonically rise with flame temperature and stoichiometry. Figure 13 plots the dependence of the measured O 2 emissions on adiabatic flame temperatures at different equivalence ratios. Similar to CO emissions, the measured O 2 emissions differ from the numerical model primarily in the regions of high sensitivity (where values are kinetically controlled, rather than by equilibrium) described previously. However, the same qualitative trends 5. Comparison of experimental and numerical results Comparison of the one-dimensional model results described earlier and experimental data for this swirling burner is not straightforward, primarily because of uncertainties associated with residence time, such as flow recirculation in the vortex breakdown bubble and backward facing step. In addition, heat losses cause the actual flame temperature to be significantly lower than the theoretical adiabatic flame temperature as stressed at the end of the previous paragraph. This latter effect was partially compensated for by using the actual measured temperature at the probe locations when comparing numerical and experimental results. Finally, CO and O 2 levels are both quite sensitive to variations in stoichiometry and temperature in parameter space regions where their concentrations are low, implying that small uncertainties in temperature and flow rate can lead to significant variations in predicted O 2 and CO levels (see Figs. 7 and 9). All the data reported here were measured at a fixed probe location, 40 cm from the base plate in the combustor center as shown in Fig. 11. The flow rates were adjusted at each flame temperature in order to give the same average burned combustor velocity of 10 m/s, based upon a calculated adiabatic flame temperature. As such, this probe location was estimated to correspond approximately to a bulk residence time of 40 ms. Fig. 12. Comparison between calculated and experimental data for CO concentration in wet products: CO levels are represented as function of flame temperature at different equivalence ratio (T in = 533 K, p = 1 atm). Lines with small symbols refer to numerical data while large full symbol to experimental data. Fig. 13. Comparison between calculated and experimental data for O 2 concentration in dry products: O 2 levels are represented as function of flame temperature at different equivalence ratio (T in = 533 K, p = 1 atm). Lines with small symbols refer to numerical data while large full symbols refer to experimental data.

7 A. Amato et al. / Proceedings of the Combustion Institute 33 (2011) Emissions concerns associated with high CO and O 2 levels place limitations on the flame temperature and feasible stoichiometries of operation of premixed methane oxy-fueled systems. This paper has shown that all qualitative measured sensitivities of CO and O 2 concentrations to stoichiometry and temperature can be captured with a one-dimensional calculation. However, there are differences in quantitative values under conditions where either CO or O 2 is a trace specie, a discrepancy that is likely due to the non-adiabatic, swirling flow. This introduces uncertainty in residence time and temperature time histories for calculations. Key trends are the following. Equilibrium CO emissions are higher than with combustion in air due to higher CO 2 levels. In addition, at a fixed residence time, CO emissions are also higher because of slower burnout of the intermediate CO formed in the flame. Equilibrium CO emissions are an exponentially increasing function of flame temperature. However, this temperature sensitivity is substantially reduced at a fixed residence time. Calculations predict that increasing pressure will lower CO emissions because of faster chemical relaxation toward equilibrium, and decreased equilibrium levels. O 2 trends are similar to CO, except inverted about / =1. Fig. 14. Comparison between calculated and experimental data for CO and O 2 concentration: emission levels are represented as function of equivalence ratio at fixed flame temperatures T = 1600 K. Lines with small symbols refer to numerical data while large full symbols refer to experimental data. are observed both in numerical and in experimental results; i.e., the O 2 emissions monotonically rise with flame temperature and decrease with stoichiometry. To facilitate the visualization of these trends, Fig. 14 plots these same experimental and numerical data as a function of equivalence ratio at a fixed flame temperature. Note that the data are bounded by the 40 ms residence time and equilibrium calculations. Parallel calculations were also performed over a range of other residence times. These calculations showed that all of the data could be fit quite well to calculations with a 100 ms residence time. Computations of the reacting flow field would be of interest in order to obtain further insight into this issue of residence time and entrainment into the recirculating flows in the vortex breakdown region and wake flow in the backward facing step. 6. Concluding remarks References [1] B. Metz, IPCC Special Report on Carbon Dioxide Capture and Storage, Cambridge University Press, New York, [2] G. Heddle, H. Herzog, M. Klett, The Economics of CO 2 Storage, Laboratory for Energy and the Environment, MIT, [3] R. Gabbrielli, R. Singh, J. Eng. Gas Turb. Power 125 (2003) 940. [4] W. Sanz, H. Jericha, B. Bauer, et al., J. Eng. Gas Turb. Power 130 (2008) [5] P. Mathieu, R. Nihart, Trans. ASME 121 (1999) [6] B.T. Chorpening, K.H. Casleton, G.A. Richards, Proc. Inst. Mech. Eng. A J. Power Eng. 219 (2005) [7] R.C. Reid, M.J. Prausnitz, B.E. Poling, The Properties of Gasses and Liquids, Mc Graw Hill, New York, [8] F. Liu, H. Guo, G.J. Smallwood, Combust. Flame 133 (2003) [9] D. Yossefi, M.R. Belmont, S.J. Makell, et al., Fuel 78 (3) (1998) [10] D. Wicksall, A. Agrawal, Effects of Fuel Composition on Flammability Limit of a Lean, Premixed Combustor, ASME Turbo EXPO, New Orleans, [11] F. Halter, F. Foucher, L. Landry, et al., Combust. Sci. Technol. 181 (2009) [12] Y. Ju, G. Masuya, P. Ronney, Proc. Combust. Inst. 2 (1998) [13] J. Ruan, H. Kobayashi, T. Niioka, et al., Combust. Flame 124 (2001) [14] Z. Chen, X. Qin, B. Xu, et al., Proc. Combust. Inst. 31 (2007) [15] K. Maruta, K. Abe, S. Hasegawa, et al., Proc. Combust. Inst. 31 (2007) [16] A. Amato, B. Hudak, T. Lieuwen, et al., J. Eng. Gas Turb. Power (2011), doi: / [17] G. Pipitone, O. Bolland, Int. J. Greenhouse Gas Control 3 (5) (2009) [18] T.C. Williams, C.R. Shaddix, R.W. Schefer, Combust. Sci. Technol. 180 (2008) [19] P. Glarborg, L.B. Bentzen, Energy Fuel 22 (2008) [20] W. Schäfer Fresenius, J. Anal. Chem. 335 (7) (1989) [21] C. Morley, Gaseq: A Chemical Equilibrium Program for Windows, available at dsl.pipex.com/. [22] T.C. Williams et al., Rev. Sci. Instrum. 78 (3) (2007). [23] R. Sander, Compilation of Henry s Law Constants for Inorganic and Organic Species of Potential Importance in Environmental Chemistry, 1999.