High-Pressure Methane Combustion over a Perovskyte Catalyst

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High-Pressure Methane Combustion over a Perovskyte Catalyst Paola S. Barbato,† Almerinda Di Benedetto,‡ Valeria Di Sarli,‡ Gianluca Landi,*,‡ and Raffaele Pirone§ †

Department of Chemical Engineering, University of Naples Federico II, P. le Tecchio 80, 80125 Naples, Italy Institute of Researches on Combustion-CNR, P. le Tecchio 80, 80125 Naples, Italy § Department of Material Science and Chemical Engineering, Politecnico di Torino, Corso Duca degli Abruzzi 24, 10129 Turin, Italy ‡

ABSTRACT: Catalytic combustion has been extensively studied as an alternative route to homogeneous combustion for power generation systems, in particular for gas turbines. Despite the great interest, very little work has been devoted to high-pressure catalytic combustion, i.e., under conditions more relevant for gas turbines. In this work, the effect of pressure on the catalytic combustion of methane on a perovskite-based monolith is investigated both experimentally and numerically. Results show that methane can be ignited by increasing the operating pressure, and this behavior can be reproduced qualitatively and quantitatively by simulating the monolith using simple overall homogeneous and heterogeneous reaction rates. Moreover, numerical results show that only the coupling between catalytic and homogeneous reactions allows correct prediction of methane conversion. As the operating pressure increases, the catalytic reaction is activated, thus behaving as a pilot for sustaining the homogeneous reaction that allows it to overcome the mass transport limitations at the catalytic surface.

1. INTRODUCTION In view of its ability to burn different fuels beyond their flammability limits with high efficiency even at relatively low temperatures, thus producing extremely low levels of pollutants (NOx, CO, and UHC), catalytic combustion (CC) is a promising technique for high-efficiency clean combustion.1 As a consequence, CC has been widely studied for power generation by natural gas-fueled turbines (GT) (see, e.g., ref 2). In this framework, structured catalytic reactors, such as honeycomb monoliths, characterized by high thermal and mechanical resistance and high geometric surface area per unit volume (that allows high catalyst loading and low pressure drops), represent the only viable choice for the reactor configuration.3 Compared with other primary measures of emission control, CC is the only technique achieving high combustion efficiency (99.99%) and (CH3CO2)2Mn 3 4H2O (Aldrich, >99%). The samples were dried in a MW oven and in a stove at 120 °C and calcined at 800 °C for 3 h under flowing air. The process was repeated 10 times to achieve the target loading (≈20 wt % perovskite with respect to the active washcoat layer, monolithic substrate excluded). The adopted procedure allowed us to deposit ∼1.4 g of catalyst onto the substrate. 2.2. Catalytic Tests under Pressure. The catalytic monolith was placed between two mullite foams acting as thermal shields and wrapped in a ceramic wool tape before being inserted into the cylindrical stainless steel reactor depicted in Figure 1. A heating jacket (Tyco Thermal Controls) equipped with a PID controller provided the preheating of the reactor. Two thermocouples were placed inside the reactor at the center of the first thermal shield (or preheater) (Tpre) and at the center of the 7548

dx.doi.org/10.1021/ie201736p |Ind. Eng. Chem. Res. 2012, 51, 7547–7558

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Figure 2. Schematic drawing of the high-pressure experimental setup.

monolith (Tcat). Moreover, a third thermocouple (Tw) was placed close to the external steel wall and the heating jacket, approximately in correspondence to Tcat (Figure 1). The two internal thermocouples were sealed thanks to a multiple-hole ceramic gland (Conax Buffalo, MHC series). Combustion tests were conducted in a lab-scale setup designed to work at pressures of e12 bar (see Figure 2). The plant can be ideally divided into three sections: gas feed preparation and control, reaction zone, and analysis. The latter was at atmospheric pressure. O2, N2, and CH4 from gas cilynders at high purities (99.7% O2, 99.998% N2, and 99.995% CH4) were indipendently controlled through mass flow controllers (Brooks SLA5850; M-1 in Figure 2) operating at 15 bar and each having a downstream two-way electrovalve (V-1 in Figure 2) operated by remote control (i) to block the flow of unused gas and (ii) to interrupt the flow of all gas in case of danger. Besides, a pressure transducer (ABB 261G; S-1 in Figure 2) was placed just downstream from the gas mixing point, thus allowing the on-line monitoring of the pressure in the first section of the rig. A system of three two-way remote-controlled electrovalves simulating a four-way valve (V-4 in Figure 2) allowed the analysis of both the reacting mixture and the reactor off products. A second pressure transducer (ABB 261G; S-1 in Figure 2) was positioned just upstream from the reactor, thus allowing the pressure measurement at the reactor (R-1 in Figure 2) inlet also when the reactor was bypassed. To prevent indesired water condensation, especially at high pressures, the reactor exit line was maintained at 120 °C until the entrance of a condenser (Parker; E-1 in Figure 2), which consisted of two coaxial steel tube coils, the inner for the gas flow and the outer for the countercurrent cooling water flow. The condensed water was collected in a tank. Finally, the dry gas or the reacting mixture, depending on the four-way valve position, flowed through the pressure controller (Brooks SLA5820; M-2 in Figure 2), which regulated the upstream pressure in the range of 015 bar gauge regardless of the total gas flow rate. The remainder of the plant was, then, at a pressure slightly above atmospheric pressure. A constant fraction

Table 1. Operating Conditions Adopted for Experimental Tests preheating temperature (°C) CH4 (%) O2 (%)

460 2.54.5 10.0

N2 (%)

balance

heating value (kJ/Nl)

0.81.4

Qtot (splh)

3188

ReIN, @STP

1133

GHSV, @ STP

1.36.7 104

P (bar)

1/11

of the gas flow rate was further dried by means of a CaCl2 chemical trap before entering the analysis system (ABB AO2000; S-2 in Figure 2), provided with four systems for the on-line and continuous analysis of the main gas species (CH4, CO2, and CO by infrared detectors and O2 by a paramagnetic detector), and equipped with a cross-sensitivity correction. Experiments were conducted under fixed preheating conditions with the pressure increasing from ∼1 to 1011 bar. The mass flow rate was kept constant (i.e., the volumetric flow rate was decreased with an increase in pressure). The conditions adopted are summarized in Table 1. It is trivial that the effects of pressure on the phenomena occurring at a fixed mass flow rate are different from those detected at a fixed inlet velocity (see, e.g., ref 34).

3. MODEL A two-dimensional CFD model was developed to simulate the coupling of the fluid flow and the chemical processes at the gassolid interface and in the gas phase for lean methane/air combustion. The central channel of the monolith reactor was simulated, which is the more adiabatic one. The channel was modeled as a right circular cylinder so that axisymmetric cylindrical coordinates could be used.35 The model solves the mass, 7549



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momentum, chemical species, and energy conservation equations in the fluid (coupled to the ideal gas equation), along with the energy equation in the solid wall. Steady-state simulations were conducted. At the inlet of the channel, a fixed flat velocity profile was assumed. For species and energy, Danckwerts boundary conditions were used. At the exit, the static pressure was imposed and far-field conditions were specified for the remaining variables. At the fluidwall interface, a no-slip boundary condition was assigned (the fluid has zero velocity relative to the boundary) that was coupled to the species balances (the mass flux of each species, FJi, is equal to its rate of production and/or consumption, ωy,i) FJi ¼ ωy, i

ð1Þ

and the energy balance λ

∂T ∂Tw ¼ λw þ ωh ∂y ∂y

ð2Þ

where λ is the fluid thermal conductivity, λw is the solid thermal conductivity, and ωh is the heat production rate of the surface. All reactor walls were considered adiabatic. However, when the role of heat losses was studied, Newton’s law of convection was used at the outer surface of the walls (whereas heat losses from the ends of the combustor were not considered): q ¼ hðTw;ext Ta;ext Þ

atmospheric pressure by a factor (P/P0)0.53 to take into account a sort of inhibiting effect of pressure on combustion kinetics.28,34 Expression 5, even if suitable for rough calculations of combustor performance, is valid over a wider range of conditions with respect to a simple power law expression, because it was derived from a reaction mechanism. For the molecular viscosity, the temperature dependence reported by Canu39 was adopted. The fluid specific heat and thermal conductivity were calculated by a mass fraction-weighted average of species properties. The species specific heat was evaluated as a piecewise fifth-power polynomial function of temperature. The model equations were discretized using a finite volume formulation on a structured mesh built by means of the Gambit preprocessor of the Fluent package. 40 Grid-independent solutions were found using cells with a dimension equal to 2.5 102 mm. The spatial discretization of the model equations used firstorder schemes for all terms, except for the diffusion terms that were treated with a second-order central difference scheme. Computations were performed by means of the segregated solver of the ANSYS FLUENT code (release 13)40 that adopts the SIMPLE method for treating the pressurevelocity coupling. All residuals were always smaller than 1.0 107.

ð3Þ

4. EXPERIMENTAL RESULTS AND MODEL VALIDATION

where h is the exterior convective heat transfer coefficient, Tw,ext is the temperature at the exterior wall surface, and Ta,ext is the external temperature (300 K). The reaction rate for homogeneous methane combustion was calculated according to the single-step reaction rate by Westbrook and Dryer:36

4.1. Experimental Results. In Figure 3, the methane conversion is plotted as a function of the operating pressure. The temperatures are also reported as measured with the thermocouples shown in the reactor scheme in Figure 1. It is observed that the methane conversion increases by increasing pressure. This evidence can be explained by considering that a pressure increase at a fixed mass flow rate results into a longer contact time and faster reaction rates (see eq 5), both positively affecting fuel conversion. Two main branches can be identified: at a pressure of 5 bar, a change in slope is found and complete methane conversion is attained, suggesting that between 4 and 5 bar ignition occurs. The temperatures of the external steel wall (Tw) and the temperature in the first thermal shield (Tpre) exhibit the same trend as methane conversion. Furthermore, in the second branch, Tw increases from 517 to 550 °C and Tpre increases from 450 to 490 °C, suggesting that the combustion reaction is heating the surroundings. The catalyst temperature (Tcat) increases up to 5 bar and, in the second branch (P > 5 bar), starts decreasing. As the operating pressure increases, the reaction front moves along the channel length shifting upstream. When the reaction front is stabilized close to the channel inlet, the remaining length of the reactor acts as a heat exchanger, thus explaining the temperature decrease of the catalyst.41,42 The effect of the inlet methane concentration on the sensitivity of the methane conversion to the operating pressure is shown in Figure 4. Under the leanest conditions (2.5%), the methane conversion steadily increases to ∼80%, without evidence of ignition. The change in methane conversion with operating pressure is qualitatively different at higher methane concentrations (34.5%): at pressures ranging from 4 to 6 bar, a change in slope is observed and the methane conversion jumps to another branch where complete conversion is attained.

RHom ¼ 2:119E þ 11 2:027E þ 8 exp ðCCH4 Þ0:2 ðCO2 Þ1:3 ðkmol m3 s1 Þ RT ð4Þ where the activation energy is in units of joules per kilomol and the concentrations are in kilomoles per cubic meter. The catalytic reaction rate used in this work was obtained from an independent experimental campaign on the LaMnO3/La-γ-Al2O3 catalyst performed in the pressure range of 112 bar:33

RCat

1:155E þ 8 1:01E þ 3 exp PCH4 RT ¼ ðkmol m2 s1 Þ 2:336E þ 7 1 þ 1:59 exp PCH4 RT

ð5Þ where the activation energy is in units of joules per kilomole and the partial pressure is in bars. Expression 5 describes a less than linear dependence of the reaction rate on fuel concentration, but it returns an almost linear dependence at atmospheric pressure. Moreover, even if heterogeneous reaction pathways can be described by an elevated number of elementary reaction steps, for engineering purposes simple reaction rates, for example, derived from reduced schemes, are generally preferred.28,34,37,38 For these reasons, the effect of pressure is often ascribed to simply multiplying a linear expression of the reaction rate at

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Figure 3. Conversion and temperature as a function of the operating pressure. The total flow rate was 31 slph, in 10% O2, 3.7% CH4, and balanced N2.

Figure 4. CH4 conversion as a function of operating pressure for different inlet methane concentrations. Conditions: Qtot = 31 slph, 10% O2, and balanced N2.

The different behavior found for different inlet methane concentrations can be attributed to the heat losses that are not counterbalanced by the combustion heat under the leanest condition (2.5%). It is worth noting that, as the operating pressure increases, the effect of the inlet methane concentration on conversion becomes stronger. At atmospheric pressure, the methane conversion is substantially independent of inlet concentration. Conversely, when the operating pressure increases, the richer mixtures show higher conversions. This behavior cannot be directly predicted from the rate expression (eq 5), reporting a less than linear effect of fuel concentration, and can be related to the fact that no negligible heat is generated by the fuel conversion. In this case,

Figure 5. Comparison of experimental (b and O) and predicted (2 and 4) conversion during CH4 combustion as a function of operating pressure. Experiments and simulations performed by increasing P (b and 2) and decreasing P (O and 4). Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, and balanced N2. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

the mean temperature of the reactor is higher because of the heat released by methane combustion, and as a consequence, the reaction is faster. From the results of Figures 3 and 4, it can be concluded that, in all cases, two different behaviors can be identified in the reactor for different operating pressures. At low pressure values (first branch), the methane conversion slightly increases with pressure and remains lower than 100%. As the operating pressure increases, the passage to a second branch is observed with the methane conversion that increases more rapidly with pressure, reaching values equal to 100%. 7551



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Figure 6. Channel axis temperature, wall temperature, bulk gas temperature, and methane conversion as a function of the axial position along the channel. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, balanced N2, and 9 bar. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

Figure 7. Maps of methane molar fraction, temperature, and homogeneous reaction. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, balanced N2, and 9 bar. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

4.2. Model Validation. The model developed here is a singlechannel model based on the assumption of adiabaticity. Indeed, a single catalytic channel can be representative of the entire system only under assumptions that include adiabaticity.43 As a consequence, the experimental tests chosen for model validation are those characterized by the highest degree of reactor adiabaticity that was obtained at the highest total flow rate.41,44 In Figure 5, the results for combustion of a 3.7/10/86.3 CH4/ O2/N2 mixture at 88 slph in terms of methane conversion as a function of operating pressure are shown as obtained by

experiments (b) and simulations (2) under the hypothesis of adiabaticity (h = 0). The simulations have been performed by the continuation method. The filled symbols represent data from experiments and simulations performed by increasing the operating pressure, while the empty symbols represent data from experiments and simulations performed starting from an ignited steady state (at P = 12 bar) and by decreasing the operating pressure. Also in this figure, it is possible to identify two different branches: the first branch at lower pressures and the second at 7552



Figure 8. Methane conversion as a function of operating pressure as obtained by the full model (triangles) and by artificially neglecting the homogeneous reaction (circles). Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, and balanced N2. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

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higher pressures. The pressure at which the passage from the first branch to the second branch (i.e., ignition) occurs is ∼7.5 bar in the experiments and 8 bar in the simulations. At a slower flow rate (Figure 3), the passage occurs at a lower pressure (5 bar). This effect can be attributed to the increased contact time obtained at the slow flow rate. The model underpredicts the methane conversion on the bottom branch but elucidates well the passage from the first to the second branch. It is worth noting that, in both experiments and simulations, a steady-state multiplicity is found. At the same value of the operating pressure (for example, 5 bar), two stable steady states are present: one corresponding to the nonignited solution (bottom branch) and the other corresponding to an ignited solution (top branch). The existence of steady-state multiplicity leads to a hysteresis: it is then possible to determine complete methane conversion also at low pressures provided that an adequate choice of the initial conditions is made. This behavior has been previously found for both macrocombustors45,46 and microcombustors.47,48 The occurrence of quenching (i.e., passage from the ignited to the nonignited state) can be detected only in experiments at atmospheric pressure,

Figure 9. Wall and bulk gas temperature (top) and CH4 conversion (bottom) as a function of the axial position in the channel for different operating pressures. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, and balanced N2. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1. 7553



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Figure 10. Maps of temperature at difference pressures as obtained by simulating the top branch of Figure 9. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, and balanced N2. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

Figure 11. Pressure drops for different operating pressures (model results). Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, and balanced N2. In the simulations, λw = 3 W m1 K1 and h = 0 W m2 K1.

while at the same pressure, a simulation starting from an ignited state shows only a reduction in CH4 conversion. This difference is due to the exchanging behavior of the reactor used in the experiments. In Figure 6, the wall temperature, the bulk gas temperature, and the gas temperature at the axis of the channel are plotted as a function of the axial position as obtained from simulations run at 9 bar. The wall temperature very rapidly reaches its maximal value (the wall behaves as an isothermal wall). The bulk gas temperature increases more slowly, slightly overcoming the wall temperature (see the enlargement in Figure 6) because of the occurrence of the homogeneous reaction. The profile of the channel axis temperature shows that the homogeneous reaction occurs when x ∼ 15 mm. In Figure 6, the methane conversion is also shown, and it appears that it follows the trend of the bulk gas temperature, suggesting that the occurrence of the homogeneous reaction is quite important in affecting methane conversion.

Figure 12. Wall and bulk gas temperature (top) and methane conversion (bottom) as a function of the axial position for two values of the support thermal conductivity. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, balanced N2, and 9 bar. In the simulations, h = 0 W m2 K1.

The temperatures shown in Figure 6 significantly differ from the experimental ones that are much more influenced by the exchange of heat with the surroundings and generally show a maximum in their profiles. These discrepancies originate from the adiabaticity assumption that is very useful for model validation under ignition conditions, where the reactor is quite adiabatic (inside and furnace temperatures are practically equal) especially for high flow rates. In Figure 7, the maps of temperature (T), CH4 molar fraction, and homogeneous reaction (Rv) are shown as obtained at 9 bar. It is found that, a few millimeters from the 7554



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Figure 13. Effect of the heat transfer coefficient. Conditions: Qtot = 88 slph, 10% O2, 3.7% CH4, balanced N2, and 9 bar. In the simulations, λw = 3 W m1 K1.

channel entrance, the homogeneous reaction starts at the wall (Rv). The fronts of methane molar fraction and temperature well reproduce the front of the homogeneous reaction, suggesting that the increase in methane conversion and temperature mainly occurs through the front of the homogeneous reaction. The wall temperature is kept constant at the adiabatic value thanks to the occurrence of the catalytic reaction and the axial diffusion of the heat through the wall. The high wall temperature allows the homogeneous reaction to be sustained, which is crucial for reaching the methane global conversion. To weight the role of the homogeneous reaction with respect to the catalytic reaction, we performed simulations by artificially neglecting the homogeneous reaction and simulations by artificially neglecting the catalytic reaction. We found that in the absence of the catalytic reaction, in the range of the operating pressure investigated (112 bar), methane conversion is negligible.

In Figure 8, methane conversion is plotted as a function of the operating pressure as obtained by including (red triangles) and neglecting (blue circles) the homogeneous reaction. In the absence of the homogeneous reaction, the passage from the bottom branch to the top branch occurs at the same value of the ignition pressure of the full model (8 bar), suggesting that the passage from the first branch to the second branch depends exclusively on the catalytic reaction. However, it is worth noting that, in the absence of the homogeneous reaction, even at pressures of >8 bar, the methane conversion is not complete. This behavior has to be attributed to the fact that the surface reaction is controlled by the transport of methane toward the catalytic surface, which becomes the step limiting the methane conversion at the high temperatures obtained in our simulations. Conversely, in the presence of both catalytic and homogeneous reactions, methane conversion is complete, thus 7555


Industrial & Engineering Chemistry Research suggesting that methane depletion is realized only in the presence of the coupling between catalytic and homogeneous reactions.35 From these results, it can be concluded that the catalytic reaction, through the increase in the wall temperature, leads to the activation of a stable homogeneous reaction front that allows the completion of methane conversion, overcoming the mass transport limitations. This result allows us to understand the experimental results of Figures 3 and 4. The different behavior of the methane conversion as a function of operating pressure, which leads to two branches, can be viewed as the result of the activation of the homogeneous reaction that allows the completion of methane conversion. In Figure 9, the axial profiles of the wall and bulk gas temperatures (top) and methane conversion (bottom) are shown for different operating pressures in the steady states of the top branch of Figure 5. It is worth noting that the position of the complete conversion shifts downstream with a decrease in pressure. At 1.3 bar, a steady-state condition with a conversion equal to ∼80% is found. In this case, the reaction front is partially outside the catalytic monolith, as shown in Figure 10, where the maps of temperature are reported for different pressures. The discrepancy with respect to the experimental results is due to the adiabaticity hypothesis used in the simulations. Nevertheless, simulation could help to identify the occurrence of quenching in the experimental tests as a blowout,49 because the heat losses may contribute to the extension of the preheating zone, and as a consequence, a shift of the reaction front toward the end of the reactor occurs. In Figure 11, the pressure drops as obtained from the simulations are plotted as a function of the operating pressure of the top branch. We found that the sizes of the pressure drops decrease with an increase in pressure. A similar trend of pressure drops was already reported by Budzianowski and Miller.34

5. MODEL SIMULATIONS Once the model had been validated, we performed simulations varying the wall thermal conductivity (λw) and the heat transfer coefficient (h). 5.1. Effect of Thermal Conductivity. The support thermal conductivity plays a crucial role in affecting the thermal behavior of catalytic macrocombustors50 and microcombustors.51,52 We simulated the effect of thermal conductivity by changing the nature of the support and then by using SiO2 (λw = 3 W m1 K1), Si (λw = 30 W m 1 K 1 ), and SiC (λw = 90 W m1 K 1 ) at 9 bar. In Figure 12, the profiles of temperature and methane conversion are reported for different thermal conductivities. The simulation results obtained for a λw of 90 W m1 K1 are not reported in Figure 12, because no ignition has been found under the conditions investigated. It is worth noting that, in the case of the more conductive material (λw = 30 W m1 K1), the wall temperature is too low to activate the homogeneous reaction and to provide complete methane conversion. Indeed, when λw = 30 W m1 K1, the gas phase reaction is strongly inhibited, the bulk gas temperature being

High-Pressure Methane Combustion over a Perovskyte Catalyst

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