Experimental and Modeling Analysis of the Thermal Behavior of an

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Experimental and Modeling Analysis of the Thermal Behavior of an Autothermal C3H8 Catalytic Partial Oxidation Reformer Dario Livio, Alessandro Donazzi, Alessandra Beretta, Gianpiero Groppi,* and Pio Forzatti Laboratory of Catalysis and Catalytic Processes, Dipartimento di Energia, Politecnico di Milano, Piazza Leonardo da Vinci 32, 20133 Milano, Italy ABSTRACT: In this work, a spatially resolved sampling technique is applied to characterize the performance of a C3H8 CPO reformer and to compare it with that of a CH4 reformer. The case of Rh-coated honeycomb catalysts is examined. The axial profiles show that higher temperatures are reached in C3H8 CPO, especially at the reactor inlet. Surface hot-spot temperatures around 950 °C lead the catalyst to rapid loss of activity. A detailed model analysis is also applied to better understand the reasons for the observed differences of the thermal behavior. On one hand, the heat release via oxidation reactions is controlled by O2 mass transfer rate and thus proportional to O2 inlet concentration, which is ∼20% higher in the C3H8/air mixture at equal C/O ratio. On the other hand, while CH4 steam reforming is partly chemically controlled, C3H8 steam reforming is mainly limited by gassolid diffusion. Thus, a less efficient balance between exo- and endothermic reactions occurs in the case of C3H8 CPO, and this results in much higher hot-spot temperatures. As a consequence, specific strategies for the optimization of the thermal behavior are required depending on the fuel. Modeling of the C3H8 CPO results shows that an increased catalyst load or a suitable aspect ratio of the reactor, combined with a decrease of the flow rate, produces a beneficial moderation of the hot-spot temperature of the catalytic wall.

1. INTRODUCTION In the last two decades the catalytic partial oxidation (CPO) of methane has been extensively studied both experimentally and theoretically. A significant amount of data is available in the literature concerning the catalytic materials, the catalyst stability, the reaction mechanism, the impact of diffusive limitations and the reactor design.113 Recently, the focus of research has shifted toward the study of the CPO of heavier fuels, such as LPG or logistic fuels (gasoline, kerosene, diesel)1416 with interest in potential commercial applications such as the on-board and distributed production of H2 and syngas.17 For these fuels, it is generally understood that complete conversion of reactants and high syngas selectivity can be obtained by using Rh-based catalysts supported on honeycomb or foam monoliths, and that major challenges concern the catalyst deactivation, in terms of coke production, and the thermal behavior of the reactor, in terms of temperature of the surface hot spot.15 Concerning the thermal behavior of the reformer, in a previous study on CH4 CPO18 we have shown that at high flow rates (20 NL/min) and at 350 °C preheat temperature the maximum surface temperature rises well above 900 °C, leading the catalyst to rapid deactivation, likely due to rhodium sintering.19 Such deactivation is an autocatalytic process that starts in the first part of the monolith: the loss of activity promotes the temperature rise, which in turn causes a further loss of activity, which spreads across the whole reactor. In the case of CH4, in order to minimize the hot-spot temperature, prevent catalyst sintering, and extend the stable operation of the reactor, criteria for the reactor design have been proposed. These are based on the modification of the balance between the rate of the exothermic reactions (kinetically limited by external mass transfer) and the rate of endothermic steam reforming (kinetically controlled by the surface chemistry). For honeycomb-supported catalysts, we have shown that the r 2011 American Chemical Society

sensitive design parameters are the channel opening, the reactor aspect ratio, the catalyst load, and the positioning of the front heat shield.18,20,21 In the case of C2+ fuels, it may be expected that the thermal stability of the catalyst becomes even more critical. On a purely thermodynamic basis an important increase of adiabatic temperature rise is associated with the use of a fuel heavier than methane (e.g., ΔTad,C3H8 = 783 °C vs ΔTad,CH4 = 650 °C, for fuel/air mixture at C/O = 0.9, TIN = 25 °C, P = 1 atm), mainly because of the increasing concentration of O2 in the feed mixture with increasing number of C atoms, for a given C/O ratio. Thanks to its wide availability and possibility of being stored as a liquid, propane can be regarded as a case molecule for the study of the CPO of light hydrocarbons. Investigating Rh-based catalyst supported over FeCrAlloy and Al2O3 foams, Holmen and co-workers22,23 reported temperatures exceeding 900 °C both for CPO and OSR (oxidative steam reforming) experiments. At these temperatures, the authors found that loss of metal dispersion occurred. A small fraction of CH4 and C2H4 were detected in the reaction products and were associated with the activation of gas phase chemistry. Coherently with these results, we have recently reported the formation of coke precursors, such as C2H4 and C3H6, in the CPO of C3H8 over Rhbased catalysts supported on a honeycomb: by application of the spatially resolved sampling technique, peaks of C2+ hydrocarbons have been observed in the first millimeters of the channel, that is, in correspondence of the hot spot (950 °C).24 Special Issue: Russo Issue Received: September 13, 2011 Accepted: December 19, 2011 Revised: December 16, 2011 Published: December 19, 2011 7573

dx.doi.org/10.1021/ie202098q | Ind. Eng. Chem. Res. 2012, 51, 7573–7583

Industrial & Engineering Chemistry Research

Figure 1. Photo of the catalytic honeycomb monolith with a continuous inert front heat shield at the beginning.

Thus, given the known consequences of high temperatures, which influence both the catalyst morphology and the chemical mechanism, it is important to characterize and fully rationalize the thermal behavior of C3H8 CPO. This has been little treated in the literature. In fact, the experimental investigations concerning the CPO of C2+ hydrocarbons have mostly focused on the integral performances of the reactors, in terms of temperature and composition of the syngas.1416,2529 Some works have also measured the evolution of the temperature profile along the axis of the catalyst by means of sliding thermocouples.30,31 Only a few experimental works have focused on the temperature and concentration profiles within the catalyst,8,32,33 but none, to our knowledge, has addressed the CPO of C3H8. The present paper addresses an experimental and numerical study of C3H8 CPO over Rh-coated honeycomb monoliths. The aim of this work is the analysis and the rationalization of the thermal behavior of the reactor by means of a spatially resolved sampling technique. Also, potential optimization strategies for minimizing the hot spot temperatures are proposed.

2. EXPERIMENTAL AND MODELING 2.1. Catalytic Materials and Adiabatic Lab-Scale Reactor. C3H8 CPO experiments were performed over 2 wt % Rh/αAl2O3 catalysts, supported onto 400 cpsi cordierite honeycomb monoliths (diameter = 24 mm, length = 40 mm). The catalyst was prepared by incipient wetness of α-Al2O3 with an aqueous solution of Rh(NO)3 and by dipping the honeycomb into a slurry of the powders, followed by blowing the excess of the slurry with air. The catalyst was deposited over a ∼25 mm length of the support (Figure 1). The remaining inert part (∼15 mm) was left uncoated and acted as a continuous front heat shield. As shown in a recent study,21 this heat shield allowed the preservation of the adiabaticity of the system and minimization of axial heat dispersion by radiation at the front face of the catalyst. A catalyst load of 600650 mg was estimated by weight difference before and after coating the monolith, which results in ∼3.9 g/L Rh load referred to monolith volume. The thickness of the layer (∼14 μm) was calculated assuming a washcoat density of 1.38 g/cm3, independently derived from dedicated measurements of weight and thickness on flat FeCrAlloy slabs, washcoated with the same procedure. The experiments were carried out in a lab-scale adiabatic reactor. The thermal insulation was realized by wrapping the reactor with a very thick layer of quartz wool. The catalytic monolith was placed in between a FeCrAlloy foam monolith and a cordierite honeycomb, which act as thermal shields and flow mixers. To avoid C formation, the catalyst and the heat shields were inserted in a quartz tube. The reactor was equipped with separated electric heaters for preheating the reactants. Once the light-off of the reaction occurred, the preheating system was turned off in order to perform autothermal tests without any

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external heat input; that is, the inlet temperature of the feed mixture was equal to room temperature (TIN = 25 °C). The spatial sampling technique was applied to collect temperature and concentration profiles along the axis of the reactor. The setup is described in detail elsewhere.20,34 For the purpose of this work it is worthy to note that, to realize the measurement, a fused silica capillary was inserted in the central channel of the catalytic monolith. The capillary was moved along the channel with a linear actuator. A submillimetric K-type thermocouple and an optical fiber (45° polished tip) connected to a narrow band IR pyrometer (Impac Infrared, IGA 5-LO) were used to collect the temperature profiles. The thermocouple measurements were taken as representative of the temperature of the gas phase, while the pyrometer measurements were representative of the temperature of the catalyst surface. To measure the composition of the gas phase, the capillary was connected to a micro-GC (Agilent 3000A). All the CPO experiments were carried out at atmospheric pressure, with 10 NL/min flow rate and C/O ratio of 0.9. The measured thermal efficiency (estimated as the ratio of the experimental and the theoretical adiabatic temperature rise) was always higher than 0.98. 2.2. Mathematical Model of the Reactor. The experimental results were quantitatively analyzed by a 1D, dynamic, heterogeneous, fixed-bed, single-channel model of the adiabatic reactor. The model is described elsewhere35 and consists of mass, enthalpy and momentum balances for the gas and solid phase. Heat and mass transfer coefficients were estimated according to specific correlations for square channels.36 The model included both homogeneous and heterogeneous kinetic schemes. Gas phase reactions for C1C3 species were taken into account according to the detailed scheme by Ranzi and co-workers.37 A molecular kinetic scheme was adopted to describe the heterogeneous chemistry of C3H8 CPO. This scheme was independently derived on the basis of a study performed in an isothermal annular reactor, which represents an extension of previous works on CH4 CPO38,39 and will be subject of a dedicated paper. Experiments of CPO and steam reforming of C3H8 were carried out within the temperature range 300850 °C, at varying space velocity (GHSV = 7 105 to 9 106 h1), C/O ratio (0.50.9), reactants dilution (C3H8 = 0.254%), and cofeed of products (H2O = 12%, H2, CO = 0.51%). The kinetic scheme was derived by analyzing the experimental data with a 1D mathematical model of the annular reactor,38 and it consisted of the whole set of CH4 CPO reactions (CH4 total oxidation, CH4 steam reforming, direct and reverse water gas shift, H2 and CO oxidation) plus several additional reaction steps, namely C3H8 total oxidation, C3H8 steam reforming, CO methanation and the steam reforming of some C2C3 intermediates (C2H6, C2H4, C3H6). The complete set of rate expressions and kinetic parameters is reported in Table 1. In line with the rate expressions for CH4 conversion, C3H8 oxidation and steam reforming were found to be first order dependent on C3H8 partial pressure, and independent of the concentration of the coreactant (O2 or H2O). It was also assessed that the rate constants of the oxidation and the steam reforming of C3H8 are about 2.5 times greater than those of CH4, with comparable activation energy. The kinetic parameters of C2C3 intermediates were set equal to those of C3H8 both in total oxidation and steam reforming. The importance of including heterogeneous conversion steps for the hydrocarbons species such as propylene, ethylene, and ethane produced by gasphase reactions has been discussed in a previous paper.24 The simplifying character of the kinetics herein adopted for such steps 7574

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Table 1. Rate Equations and Kinetic Parameters in C3H8 CPO over 2% Rh/α-Al2O3 Catalyst ratei [mol gcat1s1]

reaction ROxC3 H8 ¼

kOxC3 H8 PC3 H8 1 þ KadsH2 O PH2 O σ O2

C3H8 steam reforming C3H8 + 3H2O T 3CO + 7H2

RSRC3 H8 ¼

kSRC3 H8 PC3 H8 ð1 ηSRC3 H8 Þ 1 þ KadsCO PCO þ KadsO2 PO2 σ H2 O

CH4 total oxidation CH4 + 2O2 f CO2 + 2H2O

ROxCH4 ¼

C3H8 total oxidation C3H8 + 5O2 f 3CO2 + 4H2O

CH4 steam reforming CH4 + H2O T CO + 3H2

ki873K [mol atm1 gcat1s1]

EACT [kJ mol1]

2.500 101

80.00

2.486 10

kOxCH4 PCH4 1 þ KadsH2 O PH2 O σ O2

kSRCH4 PCH4 ð1 ηSRCH4 Þ RSRCH4 ¼ σH O 1 þ KadsCO PCO þ KadsO2 PO2 2 ηSRCH4 < 1

1

84.63

1.030 101

91.96

1.027 101

91.80


RSRC2 H4 ¼


2.486 101

84.63


RSRC2 H6 ¼


2.486 101

84.63


RSRC3 H6 ¼


2.486 101

84.63

Water gas shift CO+H2O T CO2 + H2

RWGS ¼ kWGS PH2 O ð1 ηWGS ÞσCO ηWGS < 1

6.831 103

74.83

Reverse water gas shift CO2 + H2 T CO + H2O

RRWGS ¼ kRWGS PCO2 ð1 ηRWGS Þσ H2 ηRWGS < 1

1.277 102

62.37

H2 oxidation H2 + (1)/(2)O2 f H2O

ROxH2 ¼ kOxH2 PH2 σ O2

2.666 103

61.65

CO oxidation CO + (1)/(2)O2 f CO2

ROxCO ¼ kOxCO PCO σ O2 RMetCO ¼ kMetCO PH2 ð1 ηMetCO ÞσCO ηMetCO < 1 Ki873K [atm1]

1.937 101

76.07

CO methanation CO + 3H2 T CH4 + H2O surface adsorption O2 CO H2O

5.461 100 2.114 102 8.974 100

(equaled to the rate of steam reforming of propane) is due on one side to the absence at this stage of the work of specific pieces of evidence, and on the other side to the satisfactory response of the model. The numerical analysis herein reported is fully predictive, with no parameter adjustment. The only input data of the calculations were the catalyst amount, the Rh load and dispersion (20%, as estimated experimentally by H2 chemisorption measurements), the geometrical parameters, and the physical properties of the honeycomb support. A detailed list of the relevant input parameters of the simulations is reported in Table 2.

3. RESULTS AND DISCUSSION 3.1. Thermal Behavior and Stability in C3H8 CPO. Figure 2a shows the axial temperature profiles measured by the pyrometer and the thermocouple in a series of autothermal C3H8 CPO experiments performed at increasing reactants concentration. The tests were carried out at 10 NL/min flow rate by maintaining the C/O ratio at 0.9 and progressively increasing the concentration of C3H8 from 4 to 11% v/v. The temperature profiles had the typical features of a CPO experiment: the temperature recorded by the pyrometer showed a hot spot in the first part of the catalyst (010 mm), associated with the occurrence of the exothermic oxidative chemistry, which was followed by a decrease toward the exit section, due to the prevailing role of the endothermic chemistry. The temperature measured by the thermocouple had a sharp rise, passed through a maximum and finally matched the pyrometer temperature, in line with the thermodynamic equilibrium. Extensive discussions on the measurement and the analysis of the temperature profiles can be found in the literature.20,34

1.200 10

3

88.02 ΔHADS [kJ mol1] 72.83 37.15 57.48

Table 3 reports the conversion of C3H8, the selectivity of syngas and the temperature measured by the thermocouple at the outlet of the catalyst (∼26 mm in Figure 2a) compared with the thermodynamic equilibrium values and with the model predictions. The outlet performances were very close to the adiabatic equilibrium, except for the experiment at 4% C3H8. In this latter case, the reaction was controlled by kinetics due to the low temperatures. In line with the temperature rise, the conversion of C3H8 increased, accompanied by an increase in the selectivity of syngas. Importantly, at increasing C3H8 concentration, the hot spot measured by the pyrometer became sharper and reached 950 °C with the stoichiometric mixture. Such a high local temperature was detrimental for the catalyst activity, as revealed by reference CH4 CPO experiments (Figure 2b); such experiments were performed after each C3H8 CPO run at increasing concentration under conditions (27% vol CH4, 0.9 C/O, 10 NL/min flow rate) that guarantee stable operations and are suitable to follow the occurrence of the deactivation. Noteworthy, as reported by Beretta et al.18 for CH4 CPO, the catalyst deactivation is expected to cause a marked increase of the hot-spot temperature, while the outlet temperature and reactor performance in terms of fuel conversion and syngas selectivity maintain almost constant. The results of Figure 2b show that no deactivation was evident after the experiments at 4% and 6% C3H8, whereas, after the experiment at 9% C3H8, the hot spot in CH4 CPO was 40 °C higher. Additionally, a dramatic increase of 100 °C of the hot spot in CH4 CPO was apparent after the run under stoichiometric C3H8 conditions, accompanied by a 3 mm shift of the peak inside the channel, which strongly suggests a loss of activity in the front part of the catalyst. In line with the results of Beretta et al.,18 no change of CH4 conversion and syngas selectivity was observed at the outlet of the reactor (Table 4). 7575



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The comparison of Figure 2 clearly shows that, with stoichiometric fuel/air mixtures, the C3H8 CPO experiment is characterized by much higher temperatures than the CH4 CPO experiment and this is ther cause of the rapid loss of activity. In p 3.2. Species and Temperature Profiles in C3H8 CPO and CH4 CPO. The kinetic factors can be well evidenced by focusing on the axial profiles of temperature and concentration measured in CPO experiments with stoichiometric mixtures. Figure 3 compares the results of a C3H8 CPO experiment (panels a and c) with the results of a CH4 CPO experiment (panels b and d). A Table 2. Relevant Geometrical and Physical Properties of the Honeycomb Monolith Honeycomb diameter (mm)

24

length (mm)

40

channel opening (mm)

1.092

void fraction (-)

0.74

cordierite density (g cm3)

2.3

cordierite thermal conductivity (W m1 K1)

2.5

Catalyst density (g cm3) Rh load (% w/w)

1.38 2

Rh dispersion (%)

20

Catalyst A length (mm)

25

weight (mg)

600

thickness (μm)

14

Catalyst B length (mm)

27

weight (mg)

650

thickness (μm)

14

common trend was evident: in the first 7 mm of the channel, the reactants were consumed with the production of syngas, H2O, and CO2. Once O2 was completely depleted, the fuel was further converted at the expenses of H2O by steam reforming, with the additional production of syngas. In the case of C3H8, a small fraction of CH4 and of other intermediates (Figure 3c, inset) was also formed, as a consequence of a side gas phase-reaction pathway, as extensively discussed elsewhere.24 Concerning the temperature profile, a difference of nearly 100 °C between the two tests was maintained along the axis of the monolith, which grew to 200 °C at the hot spot. Overall, the catalysts were extremely active: the profiles showed a flat trend (corresponding to the reaching of the thermodynamic equilibrium) within a few mm from the entrance, and a strong superposition of the oxidative and the reforming reactions was apparent. The model simulations (solid lines) were satisfactory, especially with respect to the temperature profiles. A good match was indeed obtained, which clearly showed that the thermocouple provided a close estimate of the gas phase temperature, while the pyrometer provided an accurate description of the catalyst surface temperature. The deviation observed outside the catalyst (between 0.5 and 0 mm) was due to a measurement artifact, as discussed in a previous work dedicated to the application of the optical fiber pyrometer in CPO.34 Concerning the concentration profiles, a good match was found for all the major species. We observed some deviations between measured and calculated profiles of the hydrocarbon gasphase intermediates at the reactor inlet, which however have a negligible impact on the consumption of the reactants and on the thermal behavior of the reactor. A source of inaccuracy could be the adoption of a lumped description of cross sectional concentration and temperature profiles in the honeycomb channel. Given the accordance with the experimental data, the numerical analysis can be used to provide a reliable picture of the reaction mechanism, in terms of axial evolution of the local rates of reactants consumption and production of H2 and CO. The rates per m3 of reactor (Figure 4) are calculated at the surface temperature and composition. In line with previous results of

Figure 2. Axial temperature profile for autothermal CPO experiments. (a) C3H8 CPO at increasing concentration of the reactants: C/O = 0.9, C3H8 = 411% v/v, flow rate = 10 NL min1, TIN = 25 °C. (b) CH4 CPO at reference conditions: C/O = 0.9, CH4 = 27% v/v, flow rate = 10 nL min1, TIN = 25 °C. The experiments are performed after each run in C3H8 at increasing concentration. 7576



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Table 3. Autothermal C3H8 CPO at Increasing Concentration of the Reactants. C/O = 0.9, C3H8 = 411% v/v, Flow Rate = 10 NL min1, TIN = 25 °C. C3H8 Conversion, Syngas Selectivity, and Temperature at the Outlet of the Catalysta

4% C3H8

6% C3H8

9% C3H8

11% C3H8

χ C3H8 [%]

σ CO [%]

σ H2 [%]

TOUT [°C]

exp

89

60

67

542

mod

81

63

74

554

eq

100

55

60

564

exp

98

78

80

650

mod

97

81

84

657

eq

100

80

80

659

exp mod

100 99

88 89

89 90

712 724

eq

100

89

89

724

exp

100

92

94

787

mod

100

94

94

788

eq

100

94

94

822

a

Comparison of experimental (exp), calculated (mod), and thermodynamic equilibrium (eq) results.

Table 4. Autothermal CH4 CPO at Reference Conditionsa χ CH4 [%]

σ CO [%]

σ H2 [%]

TOUT [°C]

reference after 4% C3H8

84.48 84.41

86.50 86.45

91.60 91.61

665 667

after 6% C3H8

84.19

86.61

91.68

665

after 9% C3H8

84.73

86.85

91.78

665

after 11% C3H8

84.89

87.28

91.60

666

equilibrium

85.79

86.03

91.71

678

a Flow rate = 10 NL min1, C/O = 0.9, CH4 = 27% v/v, TIN = 25°C. CH4 conversion, syngas selectivity, and temperature at the outlet of the catalyst. Comparison of experimental and thermodynamic equilibrium results. The experiments are performed after each run in C3H8 at increasing concentration.

CH4 CPO,40,41 also in the case of C3H8 CPO, once the reactor lit off, oxygen is almost completely consumed by H2 oxidation, while the consumption of the fuel exclusively occurs via steam reforming, which is also responsible for the production of CO and H2. WGS plays a minor role and leads to the formation of CO2 in the first part of the catalyst. Such numerical analysis reveals that under the operating conditions herein tested, the C3H8 CPO process can be very well approximated to the coupling of H2 oxidation, C3H8 steam reforming, and WGS equilibrium. 3.3. Role of Mass Transfer Limitations in C3H8 CPO and CH4 CPO. A deeper insight into the evolution of the reactants is obtained by analyzing the concentration profiles at the catalyst wall predicted by the model (Figure 5). In both the experiments, the consumption of O2 was governed by external mass transfer, as indicated by the zero concentration of O2 at the catalyst wall. Coherently, the length for complete O2 consumption kept almost unchanged, being independent of the O2 concentration and very weakly dependent on the gas temperature. A different situation emerged when focusing on the profiles of the fuel consumption. The wall concentration of CH4 was initially lower than the concentration in the gas bulk and slowly decreased until reaching the equilibrium value at the outlet of the monolith. Instead, the C3H8 consumption showed a trend similar to that of

O2: the wall concentration dropped to zero immediately after the catalyst entrance, while the consumption of C3H8 in the gas phase spread over 10 mm. To better rationalize this difference, which suggests the presence of different controlling regimes, the Carberry number for each reactant can be introduced (eq 1). Cai ¼

CBi Cwi eq CBi Ci

ð1Þ

In the equation, CBi is the concentration of the ith species in the bulk of the gas phase, Cwi is the concentration at the catalyst wall and Ceq i is the concentration calculated assuming local equilibrium at the composition and temperature of the catalyst surface. According to this definition, the external mass transfer regime corresponds to Ca f 1 and the chemical regime is represented by Ca f 0. The axial evolution of the Ca numbers is plotted in Figure 5c. As expected, this analysis confirmed that O2 consumption was totally limited by external mass transfer. Instead, a higher Ca number was found for C3H8 compared with CH4. Specifically, the curves showed that CH4 consumption was controlled by a mixed chemical-diffusive regime that approached the chemical regime by the end of the channel, while C3H8 consumption was more strongly hindered by external mass transfer along the entire axis of the catalyst. This is due to the low diffusion coefficient of C3H8 in the gas mixture, which is nearly half of the diffusion coefficient of CH4 (DC3H8,N2/DCH4,N2 = 0.53 according to the Fuller correlation42). The thermal behavior observed in C3H8 CPO and CH4 CPO is strictly related to the occurrence of the different regimes that govern the fuel consumption. This can be shown by analyzing the rates of consumption of the reactants via total oxidation and steam reforming, as well as the rates of heat removal and heat release. As previously discussed, the O2 consumption is mainly due to the oxidation of H2 and was limited by the diffusion of O2 from the gas bulk to the surface. For C3H8 and CH4, the rate of consumption was estimated as the rate of steam reforming. The rates of heat release and removal were taken as the product of the rate of consumption ri and the reaction enthalpy ΔH0R, and therefore had the dimension of a power density. In line with the full control by external mass transfer limitations, the rate of oxygen consumption was about 20% higher in the case of C3H8 (Figure 6a,c), namely due to the different inlet concentration of O2 (18.7% in C3H8 CPO vs 15.3% in CH4 CPO). Thus, the heat released by the two fuels followed the same ratio, given that in both processes H2 combustion is the O2 consuming reaction (ΔH0R = 483 kJ/molO2). On opposite, despite of a higher intrinsic kinetic rate of steam reforming (consider that kC3H8,SR is about 10-fold higher than kCH4,SR at the solid hot spot temperature), the ratio of the local rates of fuel consumption was nearly 0.5, exclusively due to the occurrence of larger mass transfer limitations for C3H8 (Figure 6b,c). Since the reaction enthalpy of steam reforming per mol of C3H8 is more than twice that of CH4 (+497 kJ/molC3H8 vs +206 kJ/molCH4), the power density locally removed in C3H8 CPO was slightly higher. Figure 6d reports the total power density, which was calculated as ∑NR i = 1 ri(ΔR,i), where NR is the number of the reactions of the kinetic mechanism. As a consequence of the coupling between oxidation and steam reforming, the reactor can be divided in two zones: in the first part (04 mm) of the catalyst, where the hot spot was located, the rate of heat release was larger than that of heat removal and the total power density of C3H8 CPO was 7577



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Figure 3. Spatially resolved profiles of temperature and composition for autothermal CPO experiments: (a and c) C3H8 CPO, catalyst A. (b and d) CH4 CPO, catalyst B. Symbols and thin lines are experimental data. Thick lines are model predictions.

higher than that of CH4 CPO. In the second zone (420 mm), endothermic steam reforming became the prevailing reaction, and the total power density of C3H8 CPO reached slightly negative values, but comparable to those of CH4 CPO. These pieces of evidence fully explain why under stoichiometric conditions higher temperatures are observed in C3H8 CPO compared with CH4 CPO. On one hand, the rate of heat released by the oxidation reactions is slightly faster for C3H8, given that the rate of oxygen consumption is controlled by external mass transfer and O2 concentration is lower for CH4. On the other hand, the rate of heat removal by C3H8 steam reforming was slowed down by the diffusive limitations. To better appreciate the impact of diffusive limitations on temperature profiles, a simulation of C3H8 CPO was performed by considering a diffusion coefficient of C3H8 equal to that of CH4 (red lines). As shown in Figure 7, by increasing the C3H8 diffusivity, the hot spot temperatures of both the solid and the gas phase became less sharp, and the maximum temperatures decreased by ∼90 °C and ∼80 °C, respectively. The shapes of the temperature profiles were very similar to those of a CH4 CPO simulation, indicating that the occurrence of mass transfer limitations plays a major role in determining the catalyst overheating in the first portion of the monolith. 3.4. Effect of Design Parameters on Temperature Profiles. The interplay between surface kinetics and mass transfer largely controls the thermal behavior of CPO reactors and needs to be considered when optimization strategies of the temperature

profile are proposed. In the case of CH4 CPO, catalyst load and channel opening are effective design parameters and can be tuned to favorably change the local balance between the exo- and the endothermic reactions and realize an optimal temperature distribution along the axis of the catalyst. As experimentally and theoretically shown in a previous investigation,20 the mixed chemical-diffusive regime that controls the reforming activity in CH4 CPO is such that the temperature profile can be smoothened by increasing the catalyst load. In fact, an increase of the catalyst load, which means an increase of the active Rh metal area, promotes the rate of steam reforming reaction and the heat removal, without affecting the oxidation rate. As well, a considerable moderation of the hot spot can be achieved by enlarging the channel opening (that is, by decreasing the honeycomb cpsi): in this way, the reduction of the external mass transfer coefficient locally reduces the rate of heat release by oxidation, while it influences to a much lesser extent the rate of steam reforming and of heat removal, causing a net flattening of the whole temperature profile. Herein it is interesting to understand to what extent the same design parameters can be exploited in C3H8 CPO. Figure 8a shows the simulations of C3H8 CPO experiments at varying the catalyst load from 250 to 750 mg (i.e., from 1.6 to 4.5 g/L of Rh). The calculations were performed considering autothermal conditions, stoichiometric C3H8/air feed, and constant honeycomb volume (diameter = 24 mm, length = 20 mm). According to this choice, the thickness of the catalyst layer was different in each simulation. The simulations show that there is still a moderate, 7578



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Figure 4. Axial evolution of the rate of consumption of reactants and production of H2 and CO, calculated at the surface composition and temperature. Conditions as in Figure 2.

Figure 5. Consumption profiles of the reactants for the experiments of Figure 2. Filled symbols are experimental data. Solid lines are model predictions of the molar fraction in the gas bulk. Dashed lines are model predictions of the molar fraction at the catalyst wall: (a) C3H8 CPO; (b) CH4 CPO; (c) axial evolution of Carberry numbers for O2, CH4, and C3H8.

beneficial effect of the catalyst load on the temperature profile, with a 55 °C reduction of the hot spot. This means that the control by external diffusion is not complete, in line with the value of the C3H8 Carberry number (Figure 5c), which does not reach the unity. However, in the calculations, the impact of internal diffusive limitations was assumed negligible and, therefore, the effect of the catalyst load on the temperature profile can be even smaller.

Figure 8b reports the simulations of C3H8 CPO experiments at decreasing the honeycomb cpsi from 400 to 90 (∼1 to 2.5 mm channel opening) and constant catalyst load (250 mg). Differently from CH4 CPO, the increase of the channel opening has a negative impact: even if smoother temperature gradients are observed, especially in the gas phase, the surface hot spot is unaltered and the temperature progressively grows to higher levels. At the lowest cpsi, the reaction does not even reach the 7579



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Figure 6. (a) Axial evolution of the rate of O2 consumption via H2 oxidation; (b) axial evolution of the rate of fuel consumption via steam reforming; (c) axial evolution of the ratio of the rates of oxidation and steam reforming. The ratio is calculated only in the first 8 mm of the catalyst, before complete conversion of the reactants or thermodynamic equilibrium conditions are reached; (d) axial evolution of the ratio of the total power density. Conditions as in Figure 2.

Figure 7. Simulated effect of an increase of C3H8 diffusion coefficient on axial temperature profiles: C/O = 0.9, C3H8 = 11% v/v, flow rate = 10 NL min1, TIN = 25 °C.

thermodynamic equilibrium, with loss of C3H8 conversion and syngas selectivity. This result is again due to the diffusive limitations that affect C3H8 steam reforming: in this case, slowing down the mass transfer rate not only reduces the local heat release, but also hinders the rate of heat consumption, thus

counterbalancing the positive effect on the oxidation rate. Overall, the results of Figure 8 suggest that, in the case of C3H8 CPO, different strategies must be adopted to moderate the hot spot. A possible solution is to reduce the heat release by decreasing the total flow rate. In this respect, the simulations of Figure 9a show the effect that a reduction of the flow rate from 15 to 5 NL/min has on the temperature profiles of the gas and the solid phase. The operating conditions are the same as in Figure 8a. The main effect of lower flow rates is a moderation of the surface hot spot of about 100 °C, while the maximum of the gas phase decreases of 25 °C with a 3 mm shift toward the catalyst inlet. However, it is crucial to note that the reduction of the flow rate also lowers the syngas productivity of the reactor. An optimization strategy that does not affect the syngas productivity is then preferable and can be accomplished by decreasing the aspect ratio of the reactor (i.e., the ratio between length and diameter, L/D). The results are reported in Figure 9b. The reference case (black lines) is characterized by an aspect ratio of 0.8 (diameter = 24 mm, length = 20 mm), while the optimal ratio (red lines, L/D = 0.1) is obtained with a 2-fold diameter (48 mm) and onefourth of the length (5 mm). In the calculations, the catalyst weight was maintained constant. With the disk shape configuration, a marked decrease of the temperature is obtained along the entire length of the catalytic monolith and a reduction of ∼130 °C of the surface hot spot is apparent. This effect happens because upon decreasing the aspect ratio, the linear velocity of the gas decreases and the heat conducted from the catalytic monolith to the front heat shield (back dispersion) becomes increasingly 7580



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Figure 8. Simulated axial temperature profiles for autothermal C3H8 CPO experiments. C/O = 0.9, C3H8 = 11% v/v, flow rate =10 NL min1, TIN = 25 °C: (a) effect of the catalyst weight (400 cpsi honeycomb); (b) effect of the channel opening (constant catalyst weight 250 mg).

Figure 9. Simulated axial temperature profiles for autothermal C3H8 CPO experiments. C/O = 0.9, C3H8 = 11% v/v, catalyst weight = 250 mg, TIN = 25 °C: (a) effect of the total flow rate; (b) effect of the reactor aspect ratio (flow rate = 10 NL min1).

important compared to the heat removed by convection, thus resulting in a moderation of the surface temperatures. Finally, we recall a preliminary discussion in a previous study:21 another promising choice is the optimization of the internal layout of the reactor in terms of distance between the front heat shield and the catalyst entrance, which causes a reduction of the hot spot via heat loss by radiation toward the walls of the reactor.

4. CONCLUSIONS Overheating is a critical issue for a Rh-based catalyst when performing the CPO of hydrocarbon fuels. In this work, we

address an experimental and theoretical investigation of the thermal behavior of a C3H8 CPO reformer with reference to the case of Rh-supported honeycomb catalysts. Spatially resolved CPO experiments with stoichiometric C3H8/air mixtures showed that temperatures as high as 950 °C were reached on the catalyst surface, which caused rapid deactivation. The comparison with reference CPO experiments carried out with stoichiometric CH4/air mixtures revealed that much higher temperatures were reached and sharper gradients were present in the case of C3H8. The reasons for these differences were rationalized, and catalyst design criteria for moderating the hot spot were analyzed. The spatially resolved measurements and the model analysis of the axial concentration profiles confirm that the 7581


Industrial & Engineering Chemistry Research balance between the rate of the exothermic oxidation (fully limited by O2 mass transfer) and the rate of endothermic steam reforming plays a pivotal role in determining the final temperature profile. In fact, in the case of C3H8 CPO the rate of heat release is larger due to the higher O2 concentration in the inlet mixture, whereas the rate of heat removal is slowed down because of the higher external diffusive resistances that affect the steam reforming reaction. In light of these results, strategies different from CH4 CPO are required to minimize the hot spot. Indeed, C3H8 CPO simulations show that enlarging the channel opening of the honeycomb does not produce beneficial effects, while only a moderate temperature decrease (55 °C) is obtained by increasing the catalyst load. The moderation of the hot spot can instead be accomplished by reducing either the total flow rate or the aspect ratio of the reactor. Combined with these solutions, the adoption of a more dissipative reactor configuration is also suggested to optimize the axial temperature profile, for instance by separating the front heat shield from the catalytic monolith in order to enhance heat dissipation by radiation.

’ AUTHOR INFORMATION Corresponding Author

*E-mail: [email protected].

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Experimental and Modeling Analysis of the Thermal Behavior of an

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