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Comparison of ignition strategies for catalytic microburners
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Combustion Institute
Proceedings of the Combustion Institute 32 (2009) 3027–3034
www.elsevier.com/locate/proci
Comparison of ignition strategies for catalytic microburners Niket S. Kaisare a,b, Georgios D. Stefanidis b, Dionisios G. Vlachos b,* a
Department of Chemical Engineering, Indian Institute of Technology – Madras, Chennai 600-036, India b Department of Chemical Engineering and Center for Catalytic Science and Technology (CCST), University of Delaware, 150 Academy Street, Newark, DE 19716, USA
Abstract Ignition of propane-air combustion in a Pt-coated microburner is numerically investigated. Three startup modes are compared: heating the inlet feed above its ignition temperature, resistive heating using electric power, and spatially distributed (stratified) resistive heating. Depending on wall conductivity, velocity, and inlet temperature (for preheated feed) or power supplied (for resistive heating), the fuel lights off either at the entrance (front-end ignition), towards the exit (back-end ignition) or in the middle of the reactor. The cumulative propane emissions are the highest for resistive heating of microburners. Promoting front-end ignition, via locally heating the initial section of the reactor or by feed preheating, significantly reduces the ignition time and the emissions. Lower conductivity materials show shorter ignition times and lower emissions for all start-up modes. The time to steady state depends on start-up mode and materials’ conductivity. A good start-up strategy would be to ignite the microburner with a low flow rate and then increase it. Ó 2009 The Combustion Institute. Published by Elsevier Inc. All rights reserved. Keywords: Catalytic; Micro-combustion; Ignition; Inlet preheating; Electrical heating
1. Introduction Integrated microdevices are increasingly being explored for application in distributed and portable power generation [1–3] due to the high energy density of fuels. Process intensification is achieved by combining reactors, heat exchangers and/or electricity generation (e.g., thermoelectrics or fuel cells) in a compact, integrated device [4,5]. Since the energy required for autonomous operation is usually provided via combustion, significant effort has been devoted to studying the stability of *
Corresponding author. Fax: +1 302 831 1048. E-mail address: [email protected] (D.G. Vlachos).
homogeneous [6–8] and catalytic [9–11] microburners (characteristic dimension < 1 mm). Using the ‘‘excess enthalpy” of hot effluent gases [12] to preheat the incoming cold feed, via spatial (e.g., Swiss-roll type burners) [13–17] or temporal [18] (reverse-flow microburners) coupling, has also been studied. Aside from safety and stability requirements, these devices should also be fast starting, i.e., the time required for warm-up from cold-start conditions should be short. This is also desirable in order to minimize emissions during start-up. Owing to the large heat capacity of the reactor solid structure, such a device could have a long start-up time [19]. A strategy of reducing the start-up time is to combust the unreacted hydro-
1540-7489/$ - see front matter Ó 2009 The Combustion Institute. Published by Elsevier Inc. All rights reserved. doi:10.1016/j.proci.2008.06.132
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gen from the fuel cell anode off-gases to preheat the inlet stream [20]. Hydrogen, which is self-igniting on platinum (Pt) catalyst, can also be used to start-up hydrocarbon combustion [21]. While microreactor start-up is a relatively new field, ignition and runaway in catalytic monoliths and packed bed reactors have been of interest for a long time [22]. The runaway behavior of catalytic combustion was studied using one-dimensional (1D) heterogeneous models [23]. Improving catalytic converters to meet emission standards became an important driver of light-off research [24,25]. Ramanathan et al. [26] presented an analytical criterion for predicting monolith light-off and developed a 1D, two-phase model to determine optimal catalyst loading [27] to minimize cold-start emissions. More recently, the onset of gaseous combustion in catalytic microburners [28,29] has also been studied. The majority of previous ignition studies have considered warm-up of conventional size devices by preheating the feed stream. Reactor resistive heating is another commonly followed strategy. Our aim is to compare these two start-up modes along with electrically preheating only a fraction of the reactor (third mode) for microscale devices. 2. Model description A parallel plate reactor is considered as illustrated in Fig. 1. The plates are 5 cm long, 1 mm thick, 300 microns apart (gap size), and are coated with Pt catalyst. The reactor is modeled using a pseudo-2D model, discussed in [10]. The model consists of the continuity, species mass balances, and energy balances for the bulk gas and the reactor wall. The energy balance for the wall is herein modified to include possible electric power input as: q s cs
X oT s o2 T s ¼ k s 2 þ g^as DH j rcat;j h1 ^a1 dt ox j ðT s T 1 Þ þ hg ^as ðT g T s Þ þ
h ∞ (T s - T∞ ) + εσ(T s4 - T ∞4 )
Qelec : 2lbw ð1Þ
h∞ (T s - T ∞ ) bw
Yk0 u 0 T g0
y x
catalyst
d/ 2
l Fig. 1. Schematic of the parallel plate burner with planar symmetry.
Here, Tg, Ts, and T1 are the bulk gas, solid wall, and ambient temperatures, respectively; qscs and ks are the solid heat capacity and thermal conductivity; h are coefficients of heat transfer; ^as ¼ ^ a1 ¼ 1=bw are specific surface areas per unit volume; g (= 1.7) is the catalyst area over the geometric surface area; Qelec is the electric power input (for resistive heating only). Other symbols are explained in Fig. 1. Radiation within the reactor is neglected due to the large aspect ratio of the channel [30]. Heat loss via radiation to the cold feed ‘‘reservoir” is accounted for in the inlet wall boundary condition: oT s ks ð2Þ ¼ h1 ðT s T g0 Þ þ er T 4s T 4g0 : ox Since the thermal reservoir at the outlet is at the exit gas temperature, a zero-flux boundary condition is a good approximation there. During the ignition phase, the temperature at the entrance is rather low and the radiation boundary condition is unlikely to affect the inlet temperature but may affect the time needed to reach steady state. The remaining equations are as in [10] and are skipped for brevity. The reduced order kinetic rate expression for propane combustion on Pt of [31], obtained via a posteriori model reduction of a detailed surface reaction model is used. Propane dissociative adsorption is the rate-determining step. The activation energies depend on the surface coverage of O*, which is computed numerically at each reactor location. Gas-phase chemistry is neglected because homogeneous combustion was not observed in earlier work for the small gap, equivalence ratio, and higher flow rates considered herein [21]. Thermodynamic and transport properties are obtained from [32,33]. The Nusselt and Sherwood numbers for transverse heat and mass transfer to the catalytic wall, treated as constants (Nu = 4 and Sh = 3.8) give reasonable agreement with 2D computational fluid dynamics (CFD) simulations [30]. The governing equations are solved using the method of lines using 500 equidistant axial nodes. The resulting differential algebraic equations are solved using the DASPK package [34]. The propane combustion kinetics was validated independently with ignition data from the literature [31]. Despite its simplicity, our model also captures the experimental trends for a stand-alone catalytic microburner [10]. The pseudo-2D model was further verified by comparing the results with 2D CFD simulations (data not shown). Spatially varying Nusselt and Sherwood numbers affects somewhat the actual values of ignition temperature and ignition time, but do not affect the overall ignition trends observed in the microburner. We define the ignition time, tign, as the time required to reach 50% outlet propane conversion.
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At tign, the inlet preheating or the electric power is turned off. We found this definition to give robust prediction of when preheating may be turned off while still reaching the ignited steady state branch. The time for steady state is denoted as tSS. The flow rate is 0.98 SLPM (6.0 m/s at Tg0 = 300 K), unless otherwise stated. Other parameters are equivalence ratio / = 0.75; heat loss coefficient, h1 = 20 W/m2/K; inlet emissivity e = 0.7. These are just representative values, chosen based on comparison with experiments presented in [10]. For preheated feeds, the inlet temperature, Tg0, increases from 300 K with a concomitant increase of inlet velocity to maintain the same flow rate. The wall thermal conductivities considered, 2, 20, and 200 W/m/K, are representative of low conductivity materials (e.g., ceramics), low conducting metals (e.g., steel) and highly conducting metals (e.g., copper), respectively. The solid heat capacity is qs cs = 2800 900 J/m3/K. 3. Inlet feed preheating mode
Max. Temperature, T
s,max
(K)
Figure 2 shows the maximum wall temperature at steady state vs. the inlet feed temperature for three wall thermal conductivities. The maximum wall temperature increases gradually as the feed temperature increases from low values until the ignition temperature Tign (a turning point in bifurcation nomenclature), beyond which propane light-off causes a large temperature increase. When the feed temperature is brought back to ambient along the ignited branch, combustion is still sustained. Under these conditions, typical hysteresis is observed. Microburners with less conducting walls ignite at a lower inlet feed temperature (heated feed in
2000 1700
2 W/m/K
A B C
1400 200 W/m/K 1100
20 W/m/K
800 500 200 300
200 W/m/K 400 500 600 700 Inlet temperature (K)
Fig. 2. Maximum wall temperature vs. inlet gas temperature for three wall thermal conductivities and 0.98 SLPM. Owing to greater heat losses, the ignition temperature increases at higher wall conductivities.
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Table 1; e.g., 535.5 vs. 576.5 K for 2 and 200 W/ m/K, respectively) due to their localized reaction zone and lower heat losses. The equivalent amount of power Qign required to heat the feed from ambient to Tign is noted in the second row of Table 1. The effect of wall conductivity on ignition temperature is contrary to that of adiabatic monoliths [26]. While not quantitatively a dramatic effect, heat loss from microscale devices results in qualitatively different behavior from their large scale, nearly adiabatic counterparts. Indeed, when we simulated an adiabatic channel, burners with highly conducting walls ignited at a lower inlet temperature (456.5 vs. 431.5 K for 2 and 200 W/m/K, respectively). The qualitatively different behavior is a function of the heat loss coefficient, but microburner behavior is expected to be as shown for reasonable heat loss coefficients [35]. Note that the solid heat capacity has no effect on steady state solutions, as expected from Eq. (1). Inlet feed temperatures exceeding Tign (see Table 1) are necessary for microburner ignition. Figure 3 compares the transient warm-up response for three wall thermal conductivities (marked A, B, and C in Fig. 2) when inlet feed temperatures are slightly above (5 K) the respective ignition temperatures. Comparing temperature profiles 100 s after start-up (dashed lines in all panels of Fig. 3), in case A (2 W/m/K) the temperatures are higher over larger distances, despite the inlet temperature being the lowest, due to the lowest heat losses. A back-end ignition is observed for 2 W/m/K with Tg0 = 540 K, wherein light-off occurs near the exit after which, the reaction zone moves progressively upstream. Middle and frontend ignition are observed for 20 and 200 W/m/ K, respectively. The inlet velocity also affects ignition location, with front-end ignition favored at lower and back-end ignition at higher flow rates (data not shown). The feed temperature has the strongest effect on ignition location. For example, for ks = 2 W/m/K and Tg0 = 582 K, middle ignition occurs; for a higher Tg0 = 615 K, front-end ignition occurs for all conductivities. The dotted lines, marked as ‘‘Ignited” in Fig. 3, correspond to the ignition time tign (50% propane conversion). For highly conducting walls, the entire reactor gets uniformly preheated. Consequently, tign is longer for highly conducting walls and decreases with decreasing wall conductivity. Following a back-end ignition, the reaction zone creeps upstream and a long time is required for microburners with less conductive walls to reach steady state. In contrast, although the time to steady state (tSS) is higher for conducting walls, the time required for the reaction zone to reach its steady state location after ignition (i.e., tSS tign) is lower for conditions favoring frontend ignition. Table 1 shows that at a higher inlet feed temperature of Tg0 = 640 K, tign, and tSS
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Table 1 Comparison of microburner start-up modes Heated feed
Resistive heating
ks = 2
ks = 20
ks = 200
Tign (K)a Qign (W)a
535.5 5.5c
550.0 5.9c
576.5 6.5c
tign (s)b tSS (s)b Emission (gm)b
50 231 0.048
122 338 0.113
238 464 0.226
ks = 2
Spatially distributed heatingd
ks = 20
ks = 200
Front
Middle
Rear
7.37
7.05
6.71
5.61
5.65
6.24
200 971 0.194
264 491 0.25
276 467 0.261
66 316 0.062
70 652 0.066
70 932 0.083
a
Steadystate ignition temperature (K) or electric power (W). Transient results for conditions considerably above the minimum ignition requirements: 8.0 W power (electric mode) or 640 K feed temperature (inlet preheating mode). c Equivalent power required to preheat the feed to Tign. d ks = 2; ks in W/m/K.
Bulk temperature, Tg (K)
b
Steady State (>780 s)
1700
Steady State (>940 s)
1400 1100
580 s 380 s (a) 2 W/m/K
800
Ignited 750 s
Ignited 280 s 100 s 0s
500 200
(c) 200 W/m/K
(b) 20 W/m/K
0
1 2 3 4 Axial distance, x (cm)
100 s 5
Steady State (>1130 s) 1000 s
800 s
Ignited 940 s 500 s
500 s 100 s
0s
1 2 3 4 Axial distance, x (cm)
5
0s 1 2 3 4 Axial distance, x (cm)
5
Fig. 3. Temporal evolution of bulk gas temperatures for the three cases marked in Fig. 2, with inlet temperatures (540, 555, and 582 K for 2, 20, and 200 W/m/K) exceeding the ignition temperatures by 5 K. Back-end ignition occurs for case A, middle of the reactor ignition for B, and front-end ignition for C. The times to ignition and steady state are considerably longer than those of Table 1 due to lower preheating used here.
decrease sharply when compared to that of Fig. 3, where Tg0 was 5 K higher than Tign. As mentioned before, front-end ignition is observed in all three cases at this Tg0 and the transient time after ignition (tSS tign) is uniformly lower for all wall conductivities. In summary, increasing preheating temperature shortens the ignition and steady state times and changes the ignition type, i.e., it moves the ignition zone upstream. Lower conductivity materials are preferred for ignition. Heat losses from microburners give qualitatively different trends from their large scale, adiabatic counterparts regarding materials’ choice, for relatively high heat losses [35]. 4. Electric resistive heating mode External heating required for ignition can alternatively be provided via resistive heating of the burner wall (e.g., via a heating tape or embedded resistances within walls). Figure 4 shows steady state solutions (typical hysteresis) in maxi-
mum wall temperature vs. electric power supplied to the microburner. The power required for ignition (Qign) increases as the wall conductivity decreases and is higher than the equivalent power required in the feed preheating mode (second row of Table 1). The wall thermal conductivity has again a strong effect on start-up. The transient ignition response is compared in Fig. 5 for three different wall conductivities when the electric power supplied is greater than the respective ignition turning point by a constant small amount (i.e., Qelec Qign is constant at 0.1 W). For 2 W/m/K, when Qelec exceeds slightly ignition conditions by 0.1 W, back-end ignition occurs, as with feed preheating. In contrast to feed preheating, doubling the electric power still yields a back-end ignition. tign (410 s) is the shortest for poorly conducting walls but the slow movement of the thermal wave from the exit to the front results in a large tSS (>1200 s). For 20 W/m/K, tign is longer and a back-end ignition occurs. However, the thermal wave now takes lesser time to reach its upstream steady state location due to higher axial wall heat
2000 2 W/m/K
1700
Max. Temperature, T
s,max
(K)
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
1400 200 W/m/K
20 W/m/K
1100 800 500 200
0
5 10 Power supplied (W)
15
Fig. 4. Maximum wall temperature vs. power supplied via electric resistive heating. Ignition points are marked with arrows.
conduction. Further increasing wall conductivity to 200 W/m/K gives a larger tign (1120 s), though 99% conversion and steady state are achieved in another 35 and 200 s, respectively. An alternate, practically relevant way is to compare transient ignition behavior at a constant electric power input. Table 1 shows that at a higher Qelec = 8 W, tign is still longer for more conducting walls (though the difference in tign is lower), whereas highly conducting walls have the shortest tSS. Figure 6 indicates that aside from wall conductivity, the reaction is also important on start-up. The reaction zone is rather narrow for 2 W/m/K and becomes broader for ks = 20 W/m/K. The combined effect of higher axial heat recirculation and heat release over a wider zone is responsible for significantly lower warm-up times for mid conductivity materials. Highly conducting walls are uniformly preheated (Fig. 5c), with the reaction zone spread over the entire reactor at tign. Subsequently, as the reactor warms up, most of the
reaction occurs upstream resulting in ignition reminiscent of front-end. In this ignition mode, more conductive walls cause slower ignition but lead faster to steady state thereafter (lower tSS tign). The heat capacity (qscs) of the solid structure affects the transient response of the microburner. Dimensional analysis and simulations (not shown) for various combinations of density and specific heat indicate that at a constant flow rate, a suitable characteristic time scale of the system (others can also be used) is based on heat conduction s = lbw/as. Here as = ks/qscs is the thermal diffusivity of the solid. With this characteristic time, the transient responses for different qscs collapse when plotted vs. the dimensionless time t/s. Thus, the higher the heat capacity of the solid structure, the longer tSS is. Simulations have been performed for various flow rates. We have found (data not shown) that the higher the flow rate, the greater the electric power required for ignition. For example, the power at ignition increases from 3.9 to 7.05 to 13.0 W as the flow rate increases from 0.33 to 0.98 to 1.96 SLPM for ks = 20 W/m/K. Thus, a good start-up strategy would be to ignite the reactor with a low flow rate, after which the flow rate is ramped up to the desired level. In passing, we should note that in the feed preheating mode, the ignition temperature decreases with increasing velocity due to slightly higher heat release prior to ignition for sufficiently large heat losses (not shown). However, the amount of power required for heating the feed to the ignition temperature shows similar increase with increasing flow rate. Detailed parametric studies of steady state ignition are given in [35]. 5. Spatially distributed (stratified) heating mode To better understand the differences between modes, instead of heating the entire reactor, the (b) 20 W/m/K
Bulk temperature (K)
1700 Steady State (>1200 s)
(b) 200 W/m/K
Steady State (>1140 s)
1400
Steady State (>1320 s)
1100
(a) 2 W/m/K
800
650 s
0s 0
940 s
Ignited 410 s 300 s
500 200
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100 s
1 2 3 4 Axial distance, x (cm)
5
1200 s Ignited 880 s 500 s 0 s 100 s
1 2 3 4 Axial distance, x (cm)
Ignited 1120 s 500 s 0 s 100 s 5
1 2 3 4 Axial distance, x (cm)
5
Fig. 5. Temporal evolution of axial bulk gas temperatures for different wall thermal conductivities for resistive heating with electric power exceeding the ignition ones by 0.1 W. Dotted lines represent ignited time, after which, the electric heating is switched off. Back-end ignition occurs for low and medium conductivities; a thermal wave creeps slowly upstream for 2 W/m/K and much faster for 20 W/m/K. The times to ignition and steady state are longer than those of Table 1 due to lower power input used here.
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(a) 2 W/m/K
Reaction rate (mol/m2 /s)
(b) 20 0.25
1700 Rear
Entire reactor 410 s
SS (c) 200
0.2
SS
0.15
940 s
880 s
0.1 1200 s
0.05 0
1350
Mid 0.05
1000
Front 0.025
650
Heated feed
1120 s 0
(b)
650 s
Exit Temperature (K)
SS
1 2 3 4 5 Axial distance, x (cm)
Fig. 6. Axial profiles of catalytic combustion rates for cases shown in Fig. 5. At the ignition time, the reaction zone spans the entire reactor for highly conducting walls (dotted lines) and gets progressively narrower with less conductive walls. Although back-end ignition is observed in panels a and b, the latter attains steady state in shorter time because the heat of combustion is released over a wider zone.
same power is applied only to 1 cm length of the reactor (and zero in other locations) at (i) the front: 0–1 cm, (ii) the center: 2–3 cm or (iii) the rear: 4–5 cm. This distributed heating results in decreased ignition power (bifurcation graphs not shown), as shown in Table 1 (ca. 7.37 W when the entire reactor is heated). Similar trends were observed for other wall thermal conductivities and inlet velocities, although the difference in the power required for ignition decreased for higher wall conductivities. The transient response with localized heating is shown in Fig. 7 and the times are summarized in Table 1, with the same total power input to facilitate comparison of heating modes. The heating location has a strong impact on time scales. Locally heating the reactor front promotes front-end ignition; tign reduces to 66 s and tSS is only 316 s, owing to the rapid heating by the propane combustion. Preheating the central 1 cm promotes middle ignition (tign = 70 s), after which the thermal wave creeps upstream toward its steady state location. Preheating the rear 1 cm promotes back-end ignition and tSS is comparable to that of heating the entire reactor (Table 1). However, tign is lower for heating the reactor rear than the entire reactor, resulting in lower propane emissions (Fig. 7a). The dash-dotted lines in Fig. 7 represent feed preheating, with a feed temperature Tg0 = 640 K for which the power supplied to the feed is 8.0 W, equal to Qelec. The transient response for resistively heating the initial 1 cm of the reactor closely follows that of inlet feed preheating (the two curves are almost undistinguishable), with tSS being the lowest for less conductive walls. A
300
0
200
Exit C3H8
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(a) 0 0
100 200 300 400
400 600 800 Time (sec)
1000
Fig. 7. Transient profiles of (a) propane mass fraction and (b) temperature at the reactor exit for four cases at an electric power Qelec = 8 W and ks = 2 W/m/K. The warm-up time is longest when the entire reactor or the final rear 1 cm is electrically heated, and shortest when the initial 1 cm of the reactor is electrically heated. Feed preheating, with a feed temperature Tg0 = 640 K (dashdotted line), is shown for comparison.
summary of several simulations reveals that ignition time and the time to reach steady state are longer whenever there is a back-end ignition and shorter whenever front-end or middle ignition occurs. Consequently, as the region of stratified heating is moved closer to the inlet, both tign and tSS get shorter. The effect of stratified heating is most prominent at low wall conductivities and the least for highly conducting walls. 6. Cumulative emissions and time scales vs. power supply Under steady state operation, ‘complete’ conversion of propane is obtained (<1 ppm for nominal conditions). Propane emissions are therefore the highest during the warm-up phase. In order to quantify the effect of power input when resistively heating the entire reactor, Fig. 8a shows tign, tSS, and cumulative emissions vs. power for three conductivities. As the power input increases above ignition (at which tign ? 1), tign and tss first decrease sharply and then more gradually. The cumulative propane emissions decrease monotonically with increasing power supply. Less conductive walls exhibit the lowest tign and cumulative emissions, due to their localized reaction zone and lower heat losses, but the longest tSS. tign falls below 1 min for power input exceeding 16 W, whereas the break-off point where the total energy (Qelec tign) is the lowest is 22 W. Figure 9 compares various modes of ignition for low conductivity walls: viz. heated feed, resis-
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
2 W/m/K 20 W/m/K
550
250 200 W/m/K 600
(b)
450 300 150
Cumulative propane emission (gm)
0 1
(c)
Total time (s)
850
517 1150
Ignition time (s)
(a)
900
20 W/m/K
0.015
Preheated feed
150 360 240 120 0 1
20
Fig. 8. Effect of electric power supplied on (a) time required for steady state, (b) ignition time, and (c) cumulative propane emission for electrically heating the entire reactor, for various wall thermal conductivities. Increasing the electric input beyond 20 W does not yield much improvement in the steady state and ignition times.
tive heating and spatially distributed heating. Propane emissions, tign, and tSS are lowest for the preheated feed mode (dash-dotted lines). Resistive preheating the entire reactor has the longer tign and higher cumulative propane emissions than any of the stratified heating cases. Heating the front 1 cm of the reactor is the best ignition strategy amongst all resistive heating options, with times and propane emissions being comparable to the preheated feed. In practical situations, front local reactor heating might be preferred since it eliminates an additional (preheating) device. 7. Conclusions Inlet feed preheating and electric resistive heating modes were compared for ignition of propane-air combustion in a Pt-coated catalytic microburner. A transition from back-end to middle of the reactor to front-end ignition was observed as the wall thermal conductivity increases, for both modes of microburner start-up. High inlet feed temperatures promote front-end ignition even for low conductivity walls. For the same power sup-
Mid Front (b)
(c) Entire reactor Rear
0.1
2 W/m/K
10 15 Power supplied (W)
Entire reactor
400
200 W/m/K
0.1
Inlet feed temperature (K) 1092 720 911 (a) Rear
650
Cumulative propane emissions (gm)
Ignition time (s)
Total time (s)
1150
3033
Preheated feed 0.01 5
10 15 Power supplied (W)
Mid
Front 20
Fig. 9. (a) Time required for steady state, (b) ignition time, and (c) cumulative propane emission vs. total power input into the system via inlet feed preheating (dashdotted line), resistive heating (solid lines), and spatially distributed resistive heating (dashed lines) for ks = 2 W/m/K. The inlet feed temperatures corresponding to the respective power are shown at the top for reference.
plied during warm-up, ignition times and cumulative propane emissions are the lowest for preheated feed mode. The time to reach steady state is especially large for back-end ignition because a thermal wave forms at the reactor exit that slowly creeps upstream. Front-end ignition via preheating the initial reactor section provides significant reduction in ignition time and propane emissions. The effect of wall thermal conductivity on power requirements and transients depends on heating mode, as summarized in Table 1. For example, less conductive walls lead to faster ignition times and lower emissions in all modes and in faster times for steady state for front-end ignition (i.e., inlet preheating and resistively preheating the initial 1 cm of the reactor).
Acknowledgment This work was supported in part by the NSF (CBET-0729714).
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[19] E.R. Delsman, M. de Croon, G.J. Kramer, Chem. Eng. J. 101 (2004) 123–131. [20] S.K. Ryi, J.S. Park, S.H. Cho, S.H. Kim, J. Power Sources 161 (2006) 1234–1240. [21] D.G. Norton, D.G. Vlachos, Proc. Combust. Inst. 30 (2005) 2473–2480. [22] G. Groppi, E. Tronconi, P. Forzatti, Catal. Rev. Sci. Eng. 41 (1999) 227–254. [23] L.C. Young, B.A. Finlayson, AIChE J. 22 (1976) 331–343. [24] C.P. Please, P.S. Hagan, D.W. Schwendeman, SIAM J. Appl. Math. 54 (1994) 72–92. [25] T. Kirchner, G. Eigenberger, Chem. Eng. Sci. 51 (1996) 2409–2418. [26] K. Ramanathan, V. Balakotaiah, D.H. West, Chem. Eng. Sci. 58 (2003) 1381–1405. [27] K. Ramanathan, D.H. West, V. Balakotaiah, Catal. Today 98 (2004) 357–373. [28] M. Reinke, J. Mantzaras, R. Schaeren, R. Bombach, A. Inauen, S. Schenker, Proc. Combust. Inst. 30 (2005) 2519–2527. [29] R. Schwiedernoch, S. Tischer, O. Deutschmann, J. Warnatz, Proc. Combust. Inst. 29 (2003) 1005–1011. [30] N.S. Kaisare, G.D. Stefanidis, D.G. Vlachos, in: A. Renken, Y. Wang (Eds.), Handbook of Microprocess Engineering, Part A: Fundamentals, Operations and Catalysts. Wiley-VCH, Berlin, in press. [31] S.R. Deshmukh, D.G. Vlachos, Combust. Flame 149 (2007) 366–383. [32] R.J. Kee, F.M. Rupley, J.A. Miller, The CHEMKIN thermodynamic data base, Sandia National Laboratories Report, SAND87-8215B, 1991. [33] R.J. Kee, G. Dixon-Lewis, J. Warnatz, M.E. Coltrin, J.A. Miller, A FORTRAN computer code package for the evaluation of gas phase multicomponent transport properties, Sandia National Laboratories Report, SAND86-8246, 1990. [34] L.R. Petzold, SIAM J. Sci. Stat. Comput. 4 (1983) 136–148. [35] G.D. Stefanidis, N.S. Kaisare, D.G. Vlachos, Chem. Eng. Technol. 31 (2008) 1170–1175.
Proceedings of the
Combustion Institute
Proceedings of the Combustion Institute 32 (2009) 3027–3034
www.elsevier.com/locate/proci
Comparison of ignition strategies for catalytic microburners Niket S. Kaisare a,b, Georgios D. Stefanidis b, Dionisios G. Vlachos b,* a
Department of Chemical Engineering, Indian Institute of Technology – Madras, Chennai 600-036, India b Department of Chemical Engineering and Center for Catalytic Science and Technology (CCST), University of Delaware, 150 Academy Street, Newark, DE 19716, USA
Abstract Ignition of propane-air combustion in a Pt-coated microburner is numerically investigated. Three startup modes are compared: heating the inlet feed above its ignition temperature, resistive heating using electric power, and spatially distributed (stratified) resistive heating. Depending on wall conductivity, velocity, and inlet temperature (for preheated feed) or power supplied (for resistive heating), the fuel lights off either at the entrance (front-end ignition), towards the exit (back-end ignition) or in the middle of the reactor. The cumulative propane emissions are the highest for resistive heating of microburners. Promoting front-end ignition, via locally heating the initial section of the reactor or by feed preheating, significantly reduces the ignition time and the emissions. Lower conductivity materials show shorter ignition times and lower emissions for all start-up modes. The time to steady state depends on start-up mode and materials’ conductivity. A good start-up strategy would be to ignite the microburner with a low flow rate and then increase it. Ó 2009 The Combustion Institute. Published by Elsevier Inc. All rights reserved. Keywords: Catalytic; Micro-combustion; Ignition; Inlet preheating; Electrical heating
1. Introduction Integrated microdevices are increasingly being explored for application in distributed and portable power generation [1–3] due to the high energy density of fuels. Process intensification is achieved by combining reactors, heat exchangers and/or electricity generation (e.g., thermoelectrics or fuel cells) in a compact, integrated device [4,5]. Since the energy required for autonomous operation is usually provided via combustion, significant effort has been devoted to studying the stability of *
Corresponding author. Fax: +1 302 831 1048. E-mail address: [email protected] (D.G. Vlachos).
homogeneous [6–8] and catalytic [9–11] microburners (characteristic dimension < 1 mm). Using the ‘‘excess enthalpy” of hot effluent gases [12] to preheat the incoming cold feed, via spatial (e.g., Swiss-roll type burners) [13–17] or temporal [18] (reverse-flow microburners) coupling, has also been studied. Aside from safety and stability requirements, these devices should also be fast starting, i.e., the time required for warm-up from cold-start conditions should be short. This is also desirable in order to minimize emissions during start-up. Owing to the large heat capacity of the reactor solid structure, such a device could have a long start-up time [19]. A strategy of reducing the start-up time is to combust the unreacted hydro-
1540-7489/$ - see front matter Ó 2009 The Combustion Institute. Published by Elsevier Inc. All rights reserved. doi:10.1016/j.proci.2008.06.132
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gen from the fuel cell anode off-gases to preheat the inlet stream [20]. Hydrogen, which is self-igniting on platinum (Pt) catalyst, can also be used to start-up hydrocarbon combustion [21]. While microreactor start-up is a relatively new field, ignition and runaway in catalytic monoliths and packed bed reactors have been of interest for a long time [22]. The runaway behavior of catalytic combustion was studied using one-dimensional (1D) heterogeneous models [23]. Improving catalytic converters to meet emission standards became an important driver of light-off research [24,25]. Ramanathan et al. [26] presented an analytical criterion for predicting monolith light-off and developed a 1D, two-phase model to determine optimal catalyst loading [27] to minimize cold-start emissions. More recently, the onset of gaseous combustion in catalytic microburners [28,29] has also been studied. The majority of previous ignition studies have considered warm-up of conventional size devices by preheating the feed stream. Reactor resistive heating is another commonly followed strategy. Our aim is to compare these two start-up modes along with electrically preheating only a fraction of the reactor (third mode) for microscale devices. 2. Model description A parallel plate reactor is considered as illustrated in Fig. 1. The plates are 5 cm long, 1 mm thick, 300 microns apart (gap size), and are coated with Pt catalyst. The reactor is modeled using a pseudo-2D model, discussed in [10]. The model consists of the continuity, species mass balances, and energy balances for the bulk gas and the reactor wall. The energy balance for the wall is herein modified to include possible electric power input as: q s cs
X oT s o2 T s ¼ k s 2 þ g^as DH j rcat;j h1 ^a1 dt ox j ðT s T 1 Þ þ hg ^as ðT g T s Þ þ
h ∞ (T s - T∞ ) + εσ(T s4 - T ∞4 )
Qelec : 2lbw ð1Þ
h∞ (T s - T ∞ ) bw
Yk0 u 0 T g0
y x
catalyst
d/ 2
l Fig. 1. Schematic of the parallel plate burner with planar symmetry.
Here, Tg, Ts, and T1 are the bulk gas, solid wall, and ambient temperatures, respectively; qscs and ks are the solid heat capacity and thermal conductivity; h are coefficients of heat transfer; ^as ¼ ^ a1 ¼ 1=bw are specific surface areas per unit volume; g (= 1.7) is the catalyst area over the geometric surface area; Qelec is the electric power input (for resistive heating only). Other symbols are explained in Fig. 1. Radiation within the reactor is neglected due to the large aspect ratio of the channel [30]. Heat loss via radiation to the cold feed ‘‘reservoir” is accounted for in the inlet wall boundary condition: oT s ks ð2Þ ¼ h1 ðT s T g0 Þ þ er T 4s T 4g0 : ox Since the thermal reservoir at the outlet is at the exit gas temperature, a zero-flux boundary condition is a good approximation there. During the ignition phase, the temperature at the entrance is rather low and the radiation boundary condition is unlikely to affect the inlet temperature but may affect the time needed to reach steady state. The remaining equations are as in [10] and are skipped for brevity. The reduced order kinetic rate expression for propane combustion on Pt of [31], obtained via a posteriori model reduction of a detailed surface reaction model is used. Propane dissociative adsorption is the rate-determining step. The activation energies depend on the surface coverage of O*, which is computed numerically at each reactor location. Gas-phase chemistry is neglected because homogeneous combustion was not observed in earlier work for the small gap, equivalence ratio, and higher flow rates considered herein [21]. Thermodynamic and transport properties are obtained from [32,33]. The Nusselt and Sherwood numbers for transverse heat and mass transfer to the catalytic wall, treated as constants (Nu = 4 and Sh = 3.8) give reasonable agreement with 2D computational fluid dynamics (CFD) simulations [30]. The governing equations are solved using the method of lines using 500 equidistant axial nodes. The resulting differential algebraic equations are solved using the DASPK package [34]. The propane combustion kinetics was validated independently with ignition data from the literature [31]. Despite its simplicity, our model also captures the experimental trends for a stand-alone catalytic microburner [10]. The pseudo-2D model was further verified by comparing the results with 2D CFD simulations (data not shown). Spatially varying Nusselt and Sherwood numbers affects somewhat the actual values of ignition temperature and ignition time, but do not affect the overall ignition trends observed in the microburner. We define the ignition time, tign, as the time required to reach 50% outlet propane conversion.
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
At tign, the inlet preheating or the electric power is turned off. We found this definition to give robust prediction of when preheating may be turned off while still reaching the ignited steady state branch. The time for steady state is denoted as tSS. The flow rate is 0.98 SLPM (6.0 m/s at Tg0 = 300 K), unless otherwise stated. Other parameters are equivalence ratio / = 0.75; heat loss coefficient, h1 = 20 W/m2/K; inlet emissivity e = 0.7. These are just representative values, chosen based on comparison with experiments presented in [10]. For preheated feeds, the inlet temperature, Tg0, increases from 300 K with a concomitant increase of inlet velocity to maintain the same flow rate. The wall thermal conductivities considered, 2, 20, and 200 W/m/K, are representative of low conductivity materials (e.g., ceramics), low conducting metals (e.g., steel) and highly conducting metals (e.g., copper), respectively. The solid heat capacity is qs cs = 2800 900 J/m3/K. 3. Inlet feed preheating mode
Max. Temperature, T
s,max
(K)
Figure 2 shows the maximum wall temperature at steady state vs. the inlet feed temperature for three wall thermal conductivities. The maximum wall temperature increases gradually as the feed temperature increases from low values until the ignition temperature Tign (a turning point in bifurcation nomenclature), beyond which propane light-off causes a large temperature increase. When the feed temperature is brought back to ambient along the ignited branch, combustion is still sustained. Under these conditions, typical hysteresis is observed. Microburners with less conducting walls ignite at a lower inlet feed temperature (heated feed in
2000 1700
2 W/m/K
A B C
1400 200 W/m/K 1100
20 W/m/K
800 500 200 300
200 W/m/K 400 500 600 700 Inlet temperature (K)
Fig. 2. Maximum wall temperature vs. inlet gas temperature for three wall thermal conductivities and 0.98 SLPM. Owing to greater heat losses, the ignition temperature increases at higher wall conductivities.
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Table 1; e.g., 535.5 vs. 576.5 K for 2 and 200 W/ m/K, respectively) due to their localized reaction zone and lower heat losses. The equivalent amount of power Qign required to heat the feed from ambient to Tign is noted in the second row of Table 1. The effect of wall conductivity on ignition temperature is contrary to that of adiabatic monoliths [26]. While not quantitatively a dramatic effect, heat loss from microscale devices results in qualitatively different behavior from their large scale, nearly adiabatic counterparts. Indeed, when we simulated an adiabatic channel, burners with highly conducting walls ignited at a lower inlet temperature (456.5 vs. 431.5 K for 2 and 200 W/m/K, respectively). The qualitatively different behavior is a function of the heat loss coefficient, but microburner behavior is expected to be as shown for reasonable heat loss coefficients [35]. Note that the solid heat capacity has no effect on steady state solutions, as expected from Eq. (1). Inlet feed temperatures exceeding Tign (see Table 1) are necessary for microburner ignition. Figure 3 compares the transient warm-up response for three wall thermal conductivities (marked A, B, and C in Fig. 2) when inlet feed temperatures are slightly above (5 K) the respective ignition temperatures. Comparing temperature profiles 100 s after start-up (dashed lines in all panels of Fig. 3), in case A (2 W/m/K) the temperatures are higher over larger distances, despite the inlet temperature being the lowest, due to the lowest heat losses. A back-end ignition is observed for 2 W/m/K with Tg0 = 540 K, wherein light-off occurs near the exit after which, the reaction zone moves progressively upstream. Middle and frontend ignition are observed for 20 and 200 W/m/ K, respectively. The inlet velocity also affects ignition location, with front-end ignition favored at lower and back-end ignition at higher flow rates (data not shown). The feed temperature has the strongest effect on ignition location. For example, for ks = 2 W/m/K and Tg0 = 582 K, middle ignition occurs; for a higher Tg0 = 615 K, front-end ignition occurs for all conductivities. The dotted lines, marked as ‘‘Ignited” in Fig. 3, correspond to the ignition time tign (50% propane conversion). For highly conducting walls, the entire reactor gets uniformly preheated. Consequently, tign is longer for highly conducting walls and decreases with decreasing wall conductivity. Following a back-end ignition, the reaction zone creeps upstream and a long time is required for microburners with less conductive walls to reach steady state. In contrast, although the time to steady state (tSS) is higher for conducting walls, the time required for the reaction zone to reach its steady state location after ignition (i.e., tSS tign) is lower for conditions favoring frontend ignition. Table 1 shows that at a higher inlet feed temperature of Tg0 = 640 K, tign, and tSS
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Table 1 Comparison of microburner start-up modes Heated feed
Resistive heating
ks = 2
ks = 20
ks = 200
Tign (K)a Qign (W)a
535.5 5.5c
550.0 5.9c
576.5 6.5c
tign (s)b tSS (s)b Emission (gm)b
50 231 0.048
122 338 0.113
238 464 0.226
ks = 2
Spatially distributed heatingd
ks = 20
ks = 200
Front
Middle
Rear
7.37
7.05
6.71
5.61
5.65
6.24
200 971 0.194
264 491 0.25
276 467 0.261
66 316 0.062
70 652 0.066
70 932 0.083
a
Steadystate ignition temperature (K) or electric power (W). Transient results for conditions considerably above the minimum ignition requirements: 8.0 W power (electric mode) or 640 K feed temperature (inlet preheating mode). c Equivalent power required to preheat the feed to Tign. d ks = 2; ks in W/m/K.
Bulk temperature, Tg (K)
b
Steady State (>780 s)
1700
Steady State (>940 s)
1400 1100
580 s 380 s (a) 2 W/m/K
800
Ignited 750 s
Ignited 280 s 100 s 0s
500 200
(c) 200 W/m/K
(b) 20 W/m/K
0
1 2 3 4 Axial distance, x (cm)
100 s 5
Steady State (>1130 s) 1000 s
800 s
Ignited 940 s 500 s
500 s 100 s
0s
1 2 3 4 Axial distance, x (cm)
5
0s 1 2 3 4 Axial distance, x (cm)
5
Fig. 3. Temporal evolution of bulk gas temperatures for the three cases marked in Fig. 2, with inlet temperatures (540, 555, and 582 K for 2, 20, and 200 W/m/K) exceeding the ignition temperatures by 5 K. Back-end ignition occurs for case A, middle of the reactor ignition for B, and front-end ignition for C. The times to ignition and steady state are considerably longer than those of Table 1 due to lower preheating used here.
decrease sharply when compared to that of Fig. 3, where Tg0 was 5 K higher than Tign. As mentioned before, front-end ignition is observed in all three cases at this Tg0 and the transient time after ignition (tSS tign) is uniformly lower for all wall conductivities. In summary, increasing preheating temperature shortens the ignition and steady state times and changes the ignition type, i.e., it moves the ignition zone upstream. Lower conductivity materials are preferred for ignition. Heat losses from microburners give qualitatively different trends from their large scale, adiabatic counterparts regarding materials’ choice, for relatively high heat losses [35]. 4. Electric resistive heating mode External heating required for ignition can alternatively be provided via resistive heating of the burner wall (e.g., via a heating tape or embedded resistances within walls). Figure 4 shows steady state solutions (typical hysteresis) in maxi-
mum wall temperature vs. electric power supplied to the microburner. The power required for ignition (Qign) increases as the wall conductivity decreases and is higher than the equivalent power required in the feed preheating mode (second row of Table 1). The wall thermal conductivity has again a strong effect on start-up. The transient ignition response is compared in Fig. 5 for three different wall conductivities when the electric power supplied is greater than the respective ignition turning point by a constant small amount (i.e., Qelec Qign is constant at 0.1 W). For 2 W/m/K, when Qelec exceeds slightly ignition conditions by 0.1 W, back-end ignition occurs, as with feed preheating. In contrast to feed preheating, doubling the electric power still yields a back-end ignition. tign (410 s) is the shortest for poorly conducting walls but the slow movement of the thermal wave from the exit to the front results in a large tSS (>1200 s). For 20 W/m/K, tign is longer and a back-end ignition occurs. However, the thermal wave now takes lesser time to reach its upstream steady state location due to higher axial wall heat
2000 2 W/m/K
1700
Max. Temperature, T
s,max
(K)
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
1400 200 W/m/K
20 W/m/K
1100 800 500 200
0
5 10 Power supplied (W)
15
Fig. 4. Maximum wall temperature vs. power supplied via electric resistive heating. Ignition points are marked with arrows.
conduction. Further increasing wall conductivity to 200 W/m/K gives a larger tign (1120 s), though 99% conversion and steady state are achieved in another 35 and 200 s, respectively. An alternate, practically relevant way is to compare transient ignition behavior at a constant electric power input. Table 1 shows that at a higher Qelec = 8 W, tign is still longer for more conducting walls (though the difference in tign is lower), whereas highly conducting walls have the shortest tSS. Figure 6 indicates that aside from wall conductivity, the reaction is also important on start-up. The reaction zone is rather narrow for 2 W/m/K and becomes broader for ks = 20 W/m/K. The combined effect of higher axial heat recirculation and heat release over a wider zone is responsible for significantly lower warm-up times for mid conductivity materials. Highly conducting walls are uniformly preheated (Fig. 5c), with the reaction zone spread over the entire reactor at tign. Subsequently, as the reactor warms up, most of the
reaction occurs upstream resulting in ignition reminiscent of front-end. In this ignition mode, more conductive walls cause slower ignition but lead faster to steady state thereafter (lower tSS tign). The heat capacity (qscs) of the solid structure affects the transient response of the microburner. Dimensional analysis and simulations (not shown) for various combinations of density and specific heat indicate that at a constant flow rate, a suitable characteristic time scale of the system (others can also be used) is based on heat conduction s = lbw/as. Here as = ks/qscs is the thermal diffusivity of the solid. With this characteristic time, the transient responses for different qscs collapse when plotted vs. the dimensionless time t/s. Thus, the higher the heat capacity of the solid structure, the longer tSS is. Simulations have been performed for various flow rates. We have found (data not shown) that the higher the flow rate, the greater the electric power required for ignition. For example, the power at ignition increases from 3.9 to 7.05 to 13.0 W as the flow rate increases from 0.33 to 0.98 to 1.96 SLPM for ks = 20 W/m/K. Thus, a good start-up strategy would be to ignite the reactor with a low flow rate, after which the flow rate is ramped up to the desired level. In passing, we should note that in the feed preheating mode, the ignition temperature decreases with increasing velocity due to slightly higher heat release prior to ignition for sufficiently large heat losses (not shown). However, the amount of power required for heating the feed to the ignition temperature shows similar increase with increasing flow rate. Detailed parametric studies of steady state ignition are given in [35]. 5. Spatially distributed (stratified) heating mode To better understand the differences between modes, instead of heating the entire reactor, the (b) 20 W/m/K
Bulk temperature (K)
1700 Steady State (>1200 s)
(b) 200 W/m/K
Steady State (>1140 s)
1400
Steady State (>1320 s)
1100
(a) 2 W/m/K
800
650 s
0s 0
940 s
Ignited 410 s 300 s
500 200
3031
100 s
1 2 3 4 Axial distance, x (cm)
5
1200 s Ignited 880 s 500 s 0 s 100 s
1 2 3 4 Axial distance, x (cm)
Ignited 1120 s 500 s 0 s 100 s 5
1 2 3 4 Axial distance, x (cm)
5
Fig. 5. Temporal evolution of axial bulk gas temperatures for different wall thermal conductivities for resistive heating with electric power exceeding the ignition ones by 0.1 W. Dotted lines represent ignited time, after which, the electric heating is switched off. Back-end ignition occurs for low and medium conductivities; a thermal wave creeps slowly upstream for 2 W/m/K and much faster for 20 W/m/K. The times to ignition and steady state are longer than those of Table 1 due to lower power input used here.
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
(a) 2 W/m/K
Reaction rate (mol/m2 /s)
(b) 20 0.25
1700 Rear
Entire reactor 410 s
SS (c) 200
0.2
SS
0.15
940 s
880 s
0.1 1200 s
0.05 0
1350
Mid 0.05
1000
Front 0.025
650
Heated feed
1120 s 0
(b)
650 s
Exit Temperature (K)
SS
1 2 3 4 5 Axial distance, x (cm)
Fig. 6. Axial profiles of catalytic combustion rates for cases shown in Fig. 5. At the ignition time, the reaction zone spans the entire reactor for highly conducting walls (dotted lines) and gets progressively narrower with less conductive walls. Although back-end ignition is observed in panels a and b, the latter attains steady state in shorter time because the heat of combustion is released over a wider zone.
same power is applied only to 1 cm length of the reactor (and zero in other locations) at (i) the front: 0–1 cm, (ii) the center: 2–3 cm or (iii) the rear: 4–5 cm. This distributed heating results in decreased ignition power (bifurcation graphs not shown), as shown in Table 1 (ca. 7.37 W when the entire reactor is heated). Similar trends were observed for other wall thermal conductivities and inlet velocities, although the difference in the power required for ignition decreased for higher wall conductivities. The transient response with localized heating is shown in Fig. 7 and the times are summarized in Table 1, with the same total power input to facilitate comparison of heating modes. The heating location has a strong impact on time scales. Locally heating the reactor front promotes front-end ignition; tign reduces to 66 s and tSS is only 316 s, owing to the rapid heating by the propane combustion. Preheating the central 1 cm promotes middle ignition (tign = 70 s), after which the thermal wave creeps upstream toward its steady state location. Preheating the rear 1 cm promotes back-end ignition and tSS is comparable to that of heating the entire reactor (Table 1). However, tign is lower for heating the reactor rear than the entire reactor, resulting in lower propane emissions (Fig. 7a). The dash-dotted lines in Fig. 7 represent feed preheating, with a feed temperature Tg0 = 640 K for which the power supplied to the feed is 8.0 W, equal to Qelec. The transient response for resistively heating the initial 1 cm of the reactor closely follows that of inlet feed preheating (the two curves are almost undistinguishable), with tSS being the lowest for less conductive walls. A
300
0
200
Exit C3H8
3032
(a) 0 0
100 200 300 400
400 600 800 Time (sec)
1000
Fig. 7. Transient profiles of (a) propane mass fraction and (b) temperature at the reactor exit for four cases at an electric power Qelec = 8 W and ks = 2 W/m/K. The warm-up time is longest when the entire reactor or the final rear 1 cm is electrically heated, and shortest when the initial 1 cm of the reactor is electrically heated. Feed preheating, with a feed temperature Tg0 = 640 K (dashdotted line), is shown for comparison.
summary of several simulations reveals that ignition time and the time to reach steady state are longer whenever there is a back-end ignition and shorter whenever front-end or middle ignition occurs. Consequently, as the region of stratified heating is moved closer to the inlet, both tign and tSS get shorter. The effect of stratified heating is most prominent at low wall conductivities and the least for highly conducting walls. 6. Cumulative emissions and time scales vs. power supply Under steady state operation, ‘complete’ conversion of propane is obtained (<1 ppm for nominal conditions). Propane emissions are therefore the highest during the warm-up phase. In order to quantify the effect of power input when resistively heating the entire reactor, Fig. 8a shows tign, tSS, and cumulative emissions vs. power for three conductivities. As the power input increases above ignition (at which tign ? 1), tign and tss first decrease sharply and then more gradually. The cumulative propane emissions decrease monotonically with increasing power supply. Less conductive walls exhibit the lowest tign and cumulative emissions, due to their localized reaction zone and lower heat losses, but the longest tSS. tign falls below 1 min for power input exceeding 16 W, whereas the break-off point where the total energy (Qelec tign) is the lowest is 22 W. Figure 9 compares various modes of ignition for low conductivity walls: viz. heated feed, resis-
N.S. Kaisare et al. / Proceedings of the Combustion Institute 32 (2009) 3027–3034
2 W/m/K 20 W/m/K
550
250 200 W/m/K 600
(b)
450 300 150
Cumulative propane emission (gm)
0 1
(c)
Total time (s)
850
517 1150
Ignition time (s)
(a)
900
20 W/m/K
0.015
Preheated feed
150 360 240 120 0 1
20
Fig. 8. Effect of electric power supplied on (a) time required for steady state, (b) ignition time, and (c) cumulative propane emission for electrically heating the entire reactor, for various wall thermal conductivities. Increasing the electric input beyond 20 W does not yield much improvement in the steady state and ignition times.
tive heating and spatially distributed heating. Propane emissions, tign, and tSS are lowest for the preheated feed mode (dash-dotted lines). Resistive preheating the entire reactor has the longer tign and higher cumulative propane emissions than any of the stratified heating cases. Heating the front 1 cm of the reactor is the best ignition strategy amongst all resistive heating options, with times and propane emissions being comparable to the preheated feed. In practical situations, front local reactor heating might be preferred since it eliminates an additional (preheating) device. 7. Conclusions Inlet feed preheating and electric resistive heating modes were compared for ignition of propane-air combustion in a Pt-coated catalytic microburner. A transition from back-end to middle of the reactor to front-end ignition was observed as the wall thermal conductivity increases, for both modes of microburner start-up. High inlet feed temperatures promote front-end ignition even for low conductivity walls. For the same power sup-
Mid Front (b)
(c) Entire reactor Rear
0.1
2 W/m/K
10 15 Power supplied (W)
Entire reactor
400
200 W/m/K
0.1
Inlet feed temperature (K) 1092 720 911 (a) Rear
650
Cumulative propane emissions (gm)
Ignition time (s)
Total time (s)
1150
3033
Preheated feed 0.01 5
10 15 Power supplied (W)
Mid
Front 20
Fig. 9. (a) Time required for steady state, (b) ignition time, and (c) cumulative propane emission vs. total power input into the system via inlet feed preheating (dashdotted line), resistive heating (solid lines), and spatially distributed resistive heating (dashed lines) for ks = 2 W/m/K. The inlet feed temperatures corresponding to the respective power are shown at the top for reference.
plied during warm-up, ignition times and cumulative propane emissions are the lowest for preheated feed mode. The time to reach steady state is especially large for back-end ignition because a thermal wave forms at the reactor exit that slowly creeps upstream. Front-end ignition via preheating the initial reactor section provides significant reduction in ignition time and propane emissions. The effect of wall thermal conductivity on power requirements and transients depends on heating mode, as summarized in Table 1. For example, less conductive walls lead to faster ignition times and lower emissions in all modes and in faster times for steady state for front-end ignition (i.e., inlet preheating and resistively preheating the initial 1 cm of the reactor).
Acknowledgment This work was supported in part by the NSF (CBET-0729714).
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