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Ind. Eng. Chem. Res. 2006, 45, 76-84
Thermal Management in Catalytic Microreactors D. G. Norton,† E. D. Wetzel,‡ and D. G. Vlachos*,† Department of Chemical Engineering, Center for Catalytic Science and Technology, and Center for Composite Materials, UniVersity of Delaware, Newark, Delaware 19716-3110, and Army Research Laboratory, Aberdeen ProVing Ground, Maryland 21005-5069
In reactors with one or two physical dimensions in the microscale and the remaining dimensions in the mesoscale or macroscale, thermal uniformity is often assumed. However, this assumption is not necessarily correct, especially for well-insulated reactors with fast reaction chemistries. In this paper, a catalytic microreactor with “tunable” wall thermal resistance (conductivity and thickness) is fabricated. Changing the thermal properties of the wall of this reactor is found to have a substantial effect on reactor thermal uniformity but only a slight effect on conversion and extinction limits. Further experiments show that the flow rate and feed composition for this reactor have a large effect on the operating temperature but only a moderate effect on thermal uniformity. An analysis of the overall energy balance of this system indicates that more than 50% of the generated heat is lost to the surroundings before the fluid exits the reactor, although this energy exchange becomes less efficient as the flow rate increases. Computational fluid dynamics simulations suggest that the reaction is limited by heat transfer near the entrance and by mass transfer further downstream, despite the small length scales of this system. Introduction A significant amount of experimental and numerical work has been performed on microscale reactors. In the field of portable power generation, the possibility of attaining high power densities with low emissions by harvesting the energy of hydrocarbons has been suggested.1-10 Microreactors have also been explored as a viable route for the production of synthesis gas and hydrogen, to be used in fuel cells.1,11-22 Finally, several biological systems, such as DNA amplification23 and bacterial pathogen detection,24 have been studied in microsystems. For many applications, isothermal or near-isothermal operation is favorable. Examples include measurements for reaction rate parameter estimation, catalyst or drug screening, pharmaceutical manufacturing, integration with thermoelectrics, and elimination of hot spots that can lead to catalyst degradation. The degree of thermal uniformity can adversely or favorably affect the performance.25 For example, temperature nonuniformity in different pores of a monolith reactor can significantly alter the conversion and selectivity. These temperature gradients can be large. As an example, simulations in partial oxidation of ethane in a catalytic monolith reactor show temperature gradients of ∼5 × 104 K/m despite the relatively low reaction exothermicity.26 Temperature nonuniformity and problems associated with hot spots at the macroscale are well recognized. Groppi and Tronconi performed numerical calculations for exothermic reactions in catalytic monolith reactors and found that increasing the support material thermal conductivity reduced the maximum catalyst temperatures and decreased the temperature gradients.19 Tronconi et al. performed experiments and simulations to study the effect of reactor material thermal conductivity on thermal uniformity within honeycomb monolith reactors. They compared traditional cordierite monoliths with novel copper monoliths for * To whom correspondence should be addressed. Tel.: (302) 8312830. Fax: (302) 831-2085. E-mail:
[email protected]. † University of Delaware. ‡ Army Research Laboratory.
CO oxidation. Significant radial thermal nonuniformity existed within cordierite monoliths that was absent within copper monoliths.25 Less work has been devoted to thermal uniformity and thermal management in microreactors. It is often tacitly assumed that microscale reactors are isothermal because of their small length scales and thus their large heat-transfer coefficients. Often, a microreactor has one dimension in the sub-millimeter range, for example, the pore diameter, and the remaining in the mesoscale or macroscale (e.g., pore lengths are often in the centimeter range). As a result, thermal uniformity may be achieved in the microscale dimension, but significant thermal gradients may exist in the nonmicroscale dimensions. Both of these conditions can present challenges for material selection, efficient conversion, and integration with other components or devices. Our previous 2D computational fluid dynamics (CFD) simulations of homogeneous chemistry microreactors combusting methane/air and propane/air mixtures predicted significant differentials in temperature within the reactor, on the order of more than a million degrees per meter. These large gradients and corresponding hot spots were due to localized reaction zones, combined with relatively slow heat transfer, in comparison to the chemistry, from the homogeneous reaction zone to and from the walls. However, when highly conductive wall materials were employed, the thermal uniformity of the wall increased substantially, and that of the fluid increased slightly.5,6 When the thermal transport within the walls is relatively efficient, the rate-limiting step in homogeneous combustion in microchannels becomes the heat transfer between the fluid and the walls. These predictions are supported by our recent experiments using prototype catalytic microcombustors, consisting of microchannels within a cast alumina cylinder with wet-deposited platinum catalyst.7,27 In these reactors, temperature differences of several hundreds of degrees were measured over a length of several centimeters. In related work, Tanaka et al. fabricated microcombustors comprised of a microchannel wet-etched into silicon, with platinum on titania particles as the catalyst. They
10.1021/ie050674o CCC: $33.50 © 2006 American Chemical Society Published on Web 11/16/2005
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also reported a hot zone near the reactor inlet and noted that controlling the temperature uniformity is a significant challenge.28 The importance of wall conductivity in mesoscale combustion has also been realized using several simplified models.29-32 In the area of microscale heat exchange, Stief et al. and Hasebe numerically studied the effect of wall thermal conductivity on the efficiency of a countercurrent microscale heat exchanger.33,34 For low wall thermal conductivities, the heat transfer between the hot and cold streams was hindered. For high wall thermal conductivities, the walls become isothermal, therefore limiting the efficiency to that of a cocurrent heat exchanger. Recent simulations in microreactors indicate that the wall thermal conductivity controls the heat-exchange operation mode and the window of operation of integrated (segregated exothermic and endothermic reactions in adjacent channels) microdevices.35,36 From all of these studies, it appears that optimal performance is attained for materials of intermediate thermal conductivities, but the optimum is possibly processspecific. While thermal management is key to the operation of microchemical devices, most fabrication protocols result in rigid, suboptimal structures. In this work, a “tunable” microreactor is fabricated using composite walls with tailorable effective thermal resistance (conductivity and thickness). This design allows for a systematic investigation of the effect of wall thermal properties on thermal uniformity during the highly exothermic combustion of hydrogen or propane in air. Additionally, the effects of feed composition and flow rate on the peak temperature and thermal uniformity are studied. Finally, an overall energy balance analysis is performed, and 2D CFD simulations are carried out to better understand the heat-transfer characteristics of microdevices. Fabrication and the Experimental Apparatus The microreactors utilized in this work are depicted schematically in Figure 1a, and a picture of one such microreactor is shown in Figure 1b. To form the reaction chamber, a 500-µmthick type 316 stainless steel (SS) gasket, comprised of thin SS plates, is sandwiched between two thicker (0.79 mm) SS plates, resulting in a channel gap that is 500 µm thick by 1 cm wide by 5 cm long. The inlet and outlet of the reaction chamber are formed by welding SS tubing to opposite ends of the channel. To eliminate “jetting”, which was observed in previous work,7 fine metal screens are placed at the inlet and exit to achieve more uniform flow patterns over the entire catalyst length. The system is made airtight with a series of nuts and bolts around the perimeter of the channel. Catalytic inserts 100 µm thick by 1 cm wide by 5 cm long are placed in the channel and held against the top and bottom with small pieces of fibrous alumina, resulting in a reaction chamber with dimensions of 300 µm by 1 cm by 5 cm and a volume of 150 µL. The catalytic inserts consist of anodized aluminum with platinum deposited on it. The inserts are fabricated by completely anodizing a 75-µm-thick 1145 alloy aluminum foil to create thin alumina sheets that serve as high-surface-area porous catalyst supports. The foil is immersed in 0.3 M aqueous oxalic acid at a potential of 40 V for 48 h. The system is held at approximately 1-2 °C using an ice bath. Parts a and b of Figure 2 show scanning electron microscopy (SEM) images of the top view and of the cross section of the porous structure. The pore structure is semiordered with approximately 2 × 1014 pores/m2 with an average pore diameter of 50 nm. The pores travel the thickness of the insert. This results in approximately 1500 m2
Figure 1. (a) Schematic of the experimental apparatus. (b) Picture of a microreactor.
of surface area available for every geometric square meter of support or approximately 14 m2/g. To deposit the platinum catalyst onto the support, the inserts are immersed in 0.007 M aqueous dihydrogen hexacholoroplatinate(IV) for 2 h. The pH of the solution is adjusted to a value of 3 with drops of either 1 M sodium hydroxide or 1 M hydrochloric acid. The inserts are then removed from the solution and dried, and the dihydrogen hexachloroplatinate(IV) is finally reduced to platinum metal with H2 at 600 °C for 3 h. To effectively manipulate the thermal resistance of the reactor walls, metallic plates (referred to herein as “thermal spreaders”) of different thermal conductivities and thicknesses are placed between the reactor walls and an outer, 6.4-mm-thick fibrous alumina insulating jacket. The thermal spreaders are adhered to the exterior of the reactor with colloidal silver paste, obtained from Structure Probe International. The silver paste is hardened by drying overnight under ambient conditions, followed by curing at 600 °C for 2 h. The drying/curing process of the paste is important to achieving acceptable reproducibility between reactors. If the initial drying step is too short, bubbles form in the paste during the high-temperature curing step, drastically changing the thermal conduction between the reactor and the conducting spreaders. We have therefore ensured that good reproducibility can be obtained between various reactors by developing a consistent drying/curing protocol. The microreactors were constructed with very thin walls in order to minimize the intrinsic thermal conduction that occurs within the reactor walls. This leads to a wide range of thermal resistance that can be studied by simply changing the thermal spreader properties. A copper thermal spreader is used to obtain a very low thermal resistance, while no thermal spreader, i.e., an air gap between the reactor wall and the insulating jacket, provides a very high thermal resistance. For intermediate thermal resistance, a SS thermal spreader is used. All thermal spreaders
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Figure 3. (a) Temperature profiles for H2/air combustion, with an equivalence ratio of 0.6, for different thermal spreaders with a thickness of 3.2 mm. As the material conductivity increases, temperature uniformity improves. (b) IR photograph of the microreactor during combustion of H2/ air at an equivalence ratio of 0.6. The location of the catalytic insert is outlined by a box for ease of visualization.
Figure 2. SEM images of (a) the porous structure of the anodized alumina wafer and (b) the cross section of the pores in the anodized alumina. The pores have diameters on the order of 50 nm. The pores start on either side of the wafer and meet at the center.
have the same width and length as the catalytic insert and a thickness of 3.2 mm (this is also the geometry of the air gap for the air thermal spreader case), unless otherwise stated. All surfaces of the reactor, in addition to the reactor walls, are enclosed in an insulating 6.4-mm-thick fibrous alumina jacket. An advantage of this “tunable” setup is that one keeps all other geometric, catalyst surface area, and flow parameters fixed to ensure maximum reproducibility. Thermocouples are spot-welded to the exterior of the reactor down the middle of the channel in the axial direction to obtain temperature profiles. Because the reactor walls are very thin, the temperatures measured on the exterior of the reactor should be very close to those along the interior channel wall. CFD simulations presented later confirm this assumption. Additionally, thermocouples were attached in the transverse direction in some runs, and the reaction channel was found to be nearly isothermal in the transverse direction. Finally, infrared (IR) imaging was used to interrogate spatial temperature profiles for noninsulated microreactors. The experiments described in this work are performed using hydrogen/air and propane/air mixtures. The former burns faster than the latter, and thus, comparison of the results for two fuels provides insight into the effect of the speed of the chemistry on thermal uniformity. Mass flow controllers control the flow
rate and the composition. The exhaust gases are sampled, and the composition is determined by gas chromatography, using a thermal conductivity detector and a flame ionization detector. The equivalence ratio is defined as the fuel-to-air ratio normalized by the stoichiometric fuel-to-air ratio. Equivalence ratios of less than 1 are denoted as fuel-lean, whereas ratios greater than 1 are denoted as fuel-rich. The flow rate is set at 2 standard liters per minute (SLPM) unless otherwise stated, and the outlet pressure is atmospheric. Ignition can be achieved during startup either via external heating, such as by using a resistive heater or preheating, or via a suitable fuel/catalyst system that self-ignites. In this work, the last method was followed. Specifically, H2/air feeds were found to be self-igniting at room temperature, as discussed in more detail in previous work.7 C3H8/air feeds were ignited by first using self-igniting H2/air feeds to sufficiently heat the reactor to allow C3H8/air feeds to ignite. Details along with optimal startup strategies can be found elsewhere.27 Effect of the Wall Thickness and Conductivity on Thermal Uniformity Figure 3a shows the temperature profiles for H2/air combustion at an equivalence ratio of 0.6 for various thermal spreader materials with a thickness of 3.2 mm. Figure 3b shows thermal imaging of the microreactor using an IR camera for a case where the insulation was completely removed from all sides of the microreactor to enable the IR camera to directly view the plates (not directly comparable with the top graph). Figure 4 shows the corresponding temperature profiles for stoichiometric C3H8/ air combustion. The error bars in these and all of the following
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Figure 4. Temperature profiles for stoichiometric propane/air combustion, for conducting thermal spreaders of different materials with a thickness of 3.2 mm. Also plotted are model predictions for air in the exterior (dotted line).
figures represent the 95% confidence limits from a minimum of two and typically four to six measurements. The gray area denotes the catalytic zone. In this and subsequent plots of the experimental data, lines are guides to the eye, unless otherwise stated. The trends in both graphs are similar. Upstream of the catalytic region and thermal spreader, there is a sharp increase in temperature. The maximum temperature is reached shortly after the catalytic region starts, and the temperature drops as the flow continues downstream. There are small variations in the temperature in the transverse direction. It appears that most of the reaction completes in a small area near the start of the catalytic zone, causing the sharp temperature rise, and heat losses cause the temperature drop downstream. When no conducting thermal spreader is used, a pocket of air within the insulation is present. The thermal gradients in this case are large, whereas when a more conducting material, such as SS, is used as a thermal spreader, some thermal smoothing is observed. When copper is used, the zone in contact with the thermal spreader is nearly isothermal. For all three cases and for both fuels, the outlet temperatures are fairly close, with the copper cases being slightly higher than the less conducting thermal spreaders, suggesting smaller overall heat losses to the surroundings. However, these differences are minimal in comparison to the large difference between the exit temperature and the adiabatic flame temperature (the adiabatic flame temperatures for H2/air and C3H8/air are 2123 and 1977 °C, respectively). Therefore, the heat losses from the reactors to the surroundings are similar for all thermal spreaders of a given fuel. A more detailed analysis of the overall energy balance is presented in a later section. Figure 5a shows the maximum reactor temperature versus the equivalence ratio for H2/air (points at the left) and C3H8 combustion (points at the right), for thermal spreaders of different materials with a thickness of 3.2 mm. H2/air mixtures are self-igniting and self-sustaining down to the leanest equivalence ratios. C3H8/air mixtures are self-sustaining down to an equivalence ratio of approximately 0.6. As the system approaches the fuel-lean limit of extinction, the maximum temperature drops considerably with decreasing C3H8 mole fraction in the feed. Near extinction, the scatter in the maximum temperature is larger. For both fuels, the maximum temperatures are highest without a thermal spreader (pocket of air outside), lower with a spreader of type 316 SS, and lowest with a spreader of copper. Again, this can be attributed to the thermal smoothing associated with the lower thermal resistance of the copper walls, which diffuses the locally generated heat along the length of the reactor wall.
Figure 5. Measured (a) maximum temperature and (b) fuel conversion in the microreactor versus the equivalence ratio for H2 and C3H8 combustion, for conducting thermal spreaders of different materials with a thickness of 3.2 mm. The model predictions for the maximum temperature with air in the exterior are also shown (dotted line). SS stands for type 316 SS.
Figure 5b shows the corresponding fuel conversion versus the equivalence ratio for H2/air and C3H8/air combustion, for thermal spreaders of different materials with a thickness of 3.2 mm. The conversion of hydrogen is nearly complete for all but the leanest mixtures. Because the highest temperatures are observed at the first thermocouple within the catalytic zone (Figure 3), it can be inferred that complete conversion of hydrogen also occurs very early in the catalytic zone. The thermal spreader material has a slight effect on the conversion of the leanest H2/air mixtures, with the best thermal spreader, copper, appearing to give lower maximum temperatures and slightly lower conversions. This behavior could be due to the elimination of a hot spot, which may help to locally stabilize the reaction. The conversion of propane is nearly complete at higher equivalence ratios but drops off quickly as fuel-lean mixtures approach extinction. This behavior is likely due to the lower reaction temperatures observed for fuel-lean conditions. Additionally, under all fuel-lean conditions, no production of CO was measured in the combustion of C3H8, which, in addition to the lack of emission of unburned fuel, greatly increases the safety of these devices in enclosed spaces. The maximum temperature for near-stoichiometric mixtures is measured at the first thermocouple in the catalytic zone. For more fuel-lean mixtures, the maximum temperature is found further downstream. In comparison, the highest temperature for H2/air combustion is always measured at the first thermocouple in the catalytic zone. This behavior is indicative of the faster reaction rate of hydrogen oxidation. Although thermal spreader material appears to have some influence on the conversion for fuel-lean mixtures, these differences are likely within the large experimental uncertainty associated with measurements near extinction. To quantify thermal uniformity, we utilize the mean deviation in temperature normalized by the average temperature in the catalytic zone
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normalized mean deviation (NMD) )
∫0L|T - Tave| dz LTmax
(1)
where 0 and L are the start and end of the catalytic and conducting insert regions, respectively,
Tave )
∫0LT dz/L
(2)
is the average temperature, and Tmax is the maximum measured temperature in the catalytic zone. The above integrals are calculated using the trapezoidal method based on the experimental temperature profile data. The heat transfer within the system can be summarized by the following Biot number, Bi, which we define in terms of thermal resistances as
Bi ) Rcond/Rconv
(3)
Figure 6. Normalized NMD versus the Biot number for H2/air combustion at equivalence ratios of 0.27 (solid circles), 0.42 (solid squares), and 0.60 (solid triangles) and for C3H8/air combustion at equivalence ratios of 0.8 (open circles), 0.9 (open squares), and 1.0 (open triangles), using thermal spreaders of various materials and thicknesses. The flow rate for the runs with H2/air using SS spreaders was varied between 1 and 4 SLPM. The flow rate of other runs was held constant at 2 SLPM.
where Rcond is the conductive resistance and Rconv is the convective resistance. For this analysis, conduction refers to conduction in the walls of the reactor in the axial direction, while convection refers to the transfer of heat from the hot internal gases to the inner surface of the reactor channel walls. Rcond can be written as
Rcond ) L/kA
(4)
where L is the channel length (5 cm) and, for cases with no thermal spreader (air outside the housing), A is the crosssectional area of the reactor wall (0.79 mm × 1 cm) and k is the thermal conductivity of the wall (SS). For cases with a thermal spreader, because the spreader is much more diffusive than the reactor wall, A is the cross-sectional area of the spreader (spreader thickness × 1 cm) and k is the conductivity of the spreader. The values for k are interpolated from tabulated values over a range of temperatures.37 Rconv can be written as
Rconv ) 1/hLW
(5)
where h is the convective heat-transfer coefficient and W is the channel width (1 cm). h is given by
h ) kfluidNu/B
(6)
where kfluid is the thermal conductivity of the flowing gases (calculated from lookup tables37 as a function of the average temperature), B is the channel height (500 µm), and Nu is the Nusselt number. For fully developed flow between parallel plates, the Nusselt number approaches a value of 7.541.38 Combining eqs 3-6 allows the calculation of the Biot number for each set of experimental conditions. Figure 6 shows NMD versus the Biot number for H2/air combustion and C3H8/air combustion at different equivalence ratios and for thermal spreaders of different materials and thicknesses. Increasing the Biot number results in a logarithmic increase of NMD. As the Biot number decreases to unity, the system becomes increasingly isothermal because the time scale for conduction in the walls approaches the time scale for heat transfer between the wall and the fluid. Increasing the equivalence ratio increases the thermal nonuniformity slightly. This result can be attributed to the increase in temperature with the equivalence ratio, which causes higher reaction rates and leads to more localized heat generation. Conversely, fuel-leaner feeds have lower reaction rates over a broader reaction zone, leading
Figure 7. Temperature profiles for C3H8/air combustion at an equivalence ratio of 0.8 for several different total flow rates. In all cases, a 3.2-mmthick SS thermal spreader was used.
to more uniform temperature profiles.7 The choice of fuel does not play a significant role in determining NMD. Effect of the Flow Rate on Thermal Uniformity Figure 7 shows the temperature profiles measured in combusting C3H8/air mixtures with an equivalence ratio of 0.8 at different total flow rates, using a 3.2-mm-thick type 316 SS thermal spreader. This equivalence ratio was chosen to enable a wide range of flow rates. For flow rates below approximately 1.3 SLPM, the system extinguished because of excessive heat losses relative to the power generation rate. Under these conditions, the long reactant residence times allow for efficient heat removal from the system, resulting in low operating temperatures, which destabilize the reaction. The maximum and average temperatures increase with increasing flow rate. Above a flow rate of 3 SLPM, the temperatures exceed that of the melting temperature of the silver paste. Temperatures increase with flow rates primarily because of the increased ratio of the heat generation rate to the heat loss rate. The former is roughly proportional to the flow rate, whereas the latter is only roughly proportional to the operating temperature (see also the next section). In all cases, there is a sharp rise in the temperature as the fluid reaches the catalytic zone with a maximum at approximately 1 cm downstream of the start of the catalytic zone. The temperature drops steadily from that point onward because of ambient heat losses. The location of the maximum temperature does not shift significantly as a function of the flow rate for this rather limited range of flow velocities. This suggests
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Figure 8. Fraction of enthalpy lost and estimated external surface temperature versus the equivalence ratio for H2/air combustion. Increasing the equivalence ratios increases both the fraction of enthalpy lost and the external surface temperature. In all cases, a 3.2-mm-thick SS thermal spreader was used. The flow rate was held constant at 2 SLPM.
that the time scales for reaction and heat transfer between the fluid and the channel walls are shorter than the residence time. NMD increases slightly with increasing flow rate for both fuels (see Figure 6 for the H2/air data) in comparison to the changes caused by varying the Biot number. The residence time decreases with increasing flow rate, and so the ratio of the residence time to the time scale of conduction in the thermal spreader decreases, causing less thermal spreading.
Figure 9. Fraction of enthalpy lost versus the total volumetric flow rate for both H2/air and C3H8/air combustion at equivalence ratios of 0.42 and 0.80, respectively. In all cases, a 3.2-mm-thick SS thermal spreader was used.
very high surface area-to-volume ratios. In this case, the surface area-to-volume ratio of the channel is approximately 6.9 × 103 m-1. This finding is true for most microchemical systems that do not employ special insulation techniques, such as a vacuum. To validate the enthalpic losses measured in these experiments, an average temperature of the exterior of the insulation can be calculated
Text )
[H ˆ (Tin,xin) - H ˆ (Tout,xout)] m˘ + Tamb hAext
(8)
Heat Loss from a Single Microreactor Channel Microburners would generally be used to generate heat for utilization by other components. This transfer of thermal energy would likely occur in one of two ways. In one approach, the system would be highly insulated, so that all of the enthalpy would stay within the flowing gases. These heated gases would then be transferred to another component, such as a heat exchanger, for utilization of the enthalpy of the gases. The other approach would be to design the reactor itself to directly conduct heat from the reacting gases to another device. In this design, the reactor has partial or minimal insulation, and thermal losses through the walls would be used directly. For our system, the adiabatic flame temperatures are significantly higher than the maximum or exit temperatures observed in these reactors, suggesting that significant heat is being transmitted through the walls of the device. The fractional enthalpy lost, µlost, can be computed from the enthalpy lost to the surroundings over the enthalpy lost if the flow reached ambient temperatures.
µlost )
H(Tin,xin) - H(Tout,xout) H(Tin,xin) - H(Tamb,xout)
(7)
Here Tin is assumed to be room temperature and xjout are the mole fractions measured at the outlet of the reactor. The enthalpies are calculated using polynomial fits as a function of the temperature and composition, utilizing data from NASALewis and Technion archives.39,40 Equation 7 depicts how well the enthalpy released in the reaction is being transferred from the microburner to a possible adjacent device. Figure 8 shows µlost as a function of the equivalence ratio. Depending on the flow rate and composition, between 60 and 80% of the enthalpy released by the reaction is lost to the surroundings by the time the products reach the end of the catalytic zone. These high enthalpy release values are observed because, although the system is well-insulated, such microscale devices inherently have
where Aext is the external surface area of the insulation, m˘ is the mass flow rate, and h is the external convective heat-transfer coefficient. Using an h value of 60 W/m2/K, Figure 8 shows the external surface temperature as a function of the equivalence ratio. The specific convective heat-transfer coefficient value was chosen to produce surface temperatures approaching 120 °C, which matches reasonably well with observed external temperatures measured during device operation. Note that this h value falls within the range for typical forced-convective heat-transfer coefficients (25-250 W/m2/K).37 Figure 9 shows the fractional enthalpic loss as a function of the total volumetric flow rate for H2/air at an equivalence ratio of 0.42 and C3H8/air combustion at an equivalence ratio of 0.8. For both fuels, the fractional enthalpic losses decrease as the flow rate increases. This behavior is a consequence of more heat being generated into the system with increasing flow rate. As far as complete conversion can be achieved, most of the heat generated by combustion is carried out as the enthalpy of the fluid because of the decreased residence time rather than being lost to the surroundings. The combination results in less heat lost to the surroundings per unit flow rate of reactant. This is an advantage from the stability point of view but a disadvantage for integrated microdevices. Specifically, these results suggest that, in the design a microburner, if sufficiently fast flows are required and a significant amount of heat is to be harvested from the flows, ample reactor surface area will be necessary to achieve the necessary heat exchange. Obviously, substantial increases of the flow rate would eventually lead to incomplete conversion and breakthrough. CFD Modeling To gain a better understanding of the system, CFD simulations for C3H8/air combustion without a thermal spreader have been performed. In our previous work, CFD simulations were performed on microburners utilizing homogeneous combustion
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Figure 10. Schematic illustrating the boundary conditions used in the CFD simulations.
of either methane or propane.5,6 The insights gained from these previous works are not necessarily applicable to the catalytic systems studied in this work. For example, temperatures in gasphase combustion are too high in comparison to those in catalytic combustion. Furthermore, while gaseous radicals are typically quenched on walls, causing flame extinction, catalysts serve mainly to form radicals that drive surface chemistry. Heat transfer is also different. For gaseous microburners, heat transfer of the heat generated in the gas phase to the walls can be slow in comparison to the chemistry time scale. On the other hand, the heat is liberated on the wall in our catalytic microreactors, accelerating heat transfer within the walls. As a result of such differences, a model incorporating catalytic combustion was developed. A more detailed simulation study will be reported elsewhere. A schematic of the geometry modeled and boundary conditions is shown in Figure 10. The problem solved is 2D flow between parallel catalytic plates. There are two different regions in the computational domain, the fluid and the solid. The domain is 6 cm long. The gap thickness studied is 300 µm, and the reactor wall thickness is 793 µm. The mesh consists of 155 axial nodes by 50 transverse nodes in the fluid and 20 transverse nodes in the wall. The nodes are nonuniformly spaced to allow for better discretization near the inlet. As in the physical system, the first and the last 0.5 cm of the wall are noncatalytic. Within the fluid, the 2D species, energy, and momentum conservation equations are solved, and within the wall, the 2D energy equation is solved. On the exterior edges of the wall, radiative and convective cooling is taken into account. Within the fluid, a symmetry boundary condition is used to reduce the computational domain in half. The specific heat, viscosity, and thermal conductivity are calculated using a mass fraction weighted average of the species properties. The species specific heats are computed using a piecewise polynomial fit of the temperature. The species viscosities and thermal conductivity are determined from kinetic theory. Multicomponent diffusion is considered in this system, where the binary diffusion coefficients are determined from kinetic theory. The emissivity of the wall’s exterior surface is set as 0.7, based on literature values for oxidized SS.37 The convective heat loss coefficient for the exterior of the walls is set as 20 W/m2/K, a value in the range of free convection. The convective heat loss coefficient used here is distinct from the effective heat loss coefficient used in the previous section. This value refers to the heat transfer from the reactor through the air pocket to the insulation, whereas the previous value referred to the overall transfer of heat from the outside of the insulation to the surroundings. The value of thermal conductivity of the wall is set as 60 W/m/K, which is significantly higher than that of SS (∼15 W/m/K). This adjustment is intended to account for the effective increase in axial thermal conduction provided by the mechanical hardware (screws and nuts) surrounding the reactor channel.
A one-step, irreversible, catalytic reaction mechanism is used to model the propane/air catalytic combustion chemistry41
r ) A0e-E/RTPC3H8RPO2β
(9)
where T is the wall temperature, A0 ) 81.6 s-1‚kmol/Pa1.15/m2, E ) 7.1128 × 107 J/kmol, R ) 1.15, and β ) 0. PC3H8 and PO2 are the partial pressures of C3H8 and O2 adjacent to the catalyst surface. The available catalyst surface area is 3 × 1021 active sites/m2, approximately 300 times that of a platinum foil. This value was determined by fitting the experimentally determined location of the maximum temperature to the model predictions. Figure 4 compares model predictions with the experimental temperature profiles for C3H8/air combustion, for the case of an air gap (no thermal spreader). There is reasonably good agreement between the measured and predicted profiles. In particular, the peak temperature value and location are captured relatively well by the model. However, the predicted exit temperature is lower than the measured value. Figure 5a compares model predictions with experimental measurements of the maximum temperature versus the equivalence ratio for C3H8/air combustion, for the case of an air gap. Again, there is fairly good agreement with the experimental data. The model predicts a wider range of stability than what was observed experimentally, giving stabilized combustion down to an equivalence ratio of ∼0.5, with higher temperatures at lower equivalence ratios. However, given that a relatively simple, onestep chemistry model is employed, the comparison is reasonable. Figure 11 shows predicted contour plots of temperature and conversion within the reactor wall and within the fluid. Figure 12 plots the temperature and conversion along the centerline and along the internal surface of the inner wall next to the fluid. In the transverse direction, the SS wall housing is nearly isothermal. The reaction rate at the wall is fast, leading to nearcomplete conversion very near to the start of the catalytic region. Within a centimeter, all of the reactants are consumed at the wall. Mass-transfer limitations, typical of catalytic combustion, are observed here, despite the small length scales involved. Near the inlet, the wall is at a higher temperature than the fluid. Heat transfer from the wall to the fluid heats the fluid until the maximum temperature is reached. After this point, the cooling of the wall is greater than the heat generation from the reaction, and the wall cools as the flow travels downstream. In this downstream area, the fluid is at a slightly higher temperature than the wall, supplying energy to the walls that is lost to the surroundings. Thus, consistent with previous CFD simulations for gaseous microburners,5,6 the wall serves a dual role: it brings the heat upstream to cause ignition and reaction stabilization but, at the same time, it serves as a heat sink controlling heat losses to the surroundings. The large temperature gradients in the transverse direction prior to entering the catalytic zone indicate that the system is initially heat-transfer-limited. The
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Figure 11. Predicted contours of (a) temperature [°C] and (b) fuel conversion for stoichiometric C3H8/air combustion at 2 SLPM with air in the exterior. The y dimension is scaled by factors of (a) 10 and (b) 40 for ease of visualization. The solid thick lines indicate the boundaries of the microreactor.
Figure 12. Temperature [°C] and fuel conversion versus the displacement along the reactor for the centerline (solid lines) and the inner wall surface (dashed lines) for stoichiometric C3H8/air combustion at 2 SLPM.
most interesting finding from these CFD simulations is the existence of mass-transfer limitations within most of the reaction zone despite the small size of the microreactor gap. Thus, further reduction of the gap size should increase the rate of combustion and is worth exploring in future work. Conclusions The catalytic microreactors used in this study achieved almost 100% conversion to complete combustion products over a wide range of operating conditions for the fuels studied in this work. Using “thermal spreaders” to augment wall thermal conductivity enables uniform reactor wall temperatures, with higher thermal conductivity and thicknesses providing more efficient thermal smoothing. Thermal uniformity, however, plays only a minor role in determining the reaction conversion and selectivity, over the range of conditions studied. Increasing the fuel composition results in higher temperatures and greater thermal nonuniformity. The choice of fuel affects thermal uniformity but to a lesser extent. The flow rate plays a moderate role in determining thermal uniformity. Specifically, higher flow rates increase reactor temperatures and reduce thermal uniformity. While low flow rates improve uniformity, they also can induce extinction, at least in C3H8/air combustion, because of increased heat loss to heat generation ratios. An analysis shows that approximately 60-80% of the enthalpy liberated from the exothermic reactions is lost from the system to the ambient before the flow leaves the reactor, despite the insulation on the reactor exterior. The large fractional heat losses are due to the high surface area-to-volume ratio
inherent to these microscale systems, and these numbers should apply to most single-channel microchemical systems. This quality is beneficial for heat-exchanger applications but detrimental for the design of adiabatic systems. A CFD model indicates that the surface reaction is sufficiently fast to achieve complete conversion, at least at stoichiometric conditions. The modeling results also show that mass transfer limits fuel conversion downstream, whereas heat transfer from the walls to the gases is rate-limiting near the reactor entrance. The choice of fuels used in this work has been motivated in part from the fast chemistry/large exothermicity of these fuels, which makes this study a worst-case scenario, and in part from the application itself, namely, the need of microcombustion for portable power generation. Obviously, in these systems under fuel-lean to stoichiometric conditions, conversion is the main concern, if one avoids slip of unburned hydrocarbons. In our systems, conversion is nearly complete or extinguished. Similarly, the selectivity to complete combustion is generally 100%. At this point, the effects of thermal uniformity on reactions whose selectivity is affected by temperature gradients have not been explored but are expected to be significant for some reactions. The tunability of our microreactors provides an easy laboratory tool to enable the study of such reactions. In contrast to other methods used for controlling temperature uniformity, for example, catalyst dilution in fixed-bed reactors or feed dilution, a variable thermal bandwidth (wall conductivity and thickness) allows study in open, relatively short channels with a low pressure drop of undiluted feeds to determine experimentally suitable materials for reactors, such as monoliths, with controlled temperature uniformity. The assumption that a system is thermally uniform simply because it is a microreactor is not necessarily a valid one. Our experimental results demonstrate that thermal uniformity of microreactor systems is highly dependent on, and can be controlled by, selection of reactor materials and geometry and the reaction sets involved, as well as the feed composition and flow rate. The specific design of a microreactor can therefore be tailored depending on the particular application. For mildly exothermic or endothermic reactions, it may be possible to achieve the desired level of thermal uniformity easily, but for highly exothermic or endothermic reactions, appropriate attention must be paid to the thermal bandwidth of the system to determine whether it is adequate. In a forthcoming paper, we will investigate microcombustor design for integration with thermoelectric devices for the production of electricity.
84 Ind. Eng. Chem. Res., Vol. 45, No. 1, 2006
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ReceiVed for reView June 9, 2005 ReVised manuscript receiVed October 4, 2005 Accepted October 18, 2005 IE050674O
