Combustion and Flame 157 (2010) 1942–1958

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Combustion and Flame

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Effects of hydrogen preconversion on the homogeneous ignition of fuel-leanH2/O2/N2/CO2 mixtures over platinum at moderate pressures

Yohannes Ghermay, John Mantzaras *, Rolf BombachPaul Scherrer Institute, Combustion Research, CH-5232 Villigen PSI, Switzerland

a r t i c l e i n f o a b s t r a c t

Article history:Received 4 December 2009Received in revised form 11 February 2010Accepted 23 February 2010Available online 16 March 2010

Keywords:Homogeneous ignition of hydrogen overplatinumFuel preconversionIn situ Raman and LIF measurements overcatalytic surfacesHetero-/homogeneous combustion conceptCatalytic reactor thermal management

0010-2180/$ - see front matter � 2010 The Combustdoi:10.1016/j.combustflame.2010.02.016

* Corresponding author. Fax: +41 56 3102199.E-mail address: ioannis.mantzaras@psi.ch (J. Mant

The impact of fractional hydrogen preconversion on the subsequent homogeneous ignition characteris-tics of fuel-lean (equivalence ratio u = 0.30) H2/O2/N2/CO2 mixtures over platinum was investigatedexperimentally and numerically at pressures of 1, 5 and 8 bar. Experiments were performed in an opti-cally accessible channel-flow reactor and involved Raman measurements of major species over the cat-alyst boundary layer and planar laser induced fluorescence (LIF) of the OH radical. Simulations werecarried out with a 2-D elliptic code that included detailed hetero-/homogeneous chemistry. The predic-tions reproduced the LIF-measured onset of homogeneous ignition and the Raman-measured transport-limited catalytic hydrogen consumption. For 0% preconversion and wall temperatures in the range900 K 6 Tw 6 1100 K, homogeneous ignition was largely suppressed for p P 5 bar due to the combinedeffects of intrinsic gas-phase hydrogen kinetics and the competition between the catalytic and gas-phasepathways for fuel consumption. A moderate increase of preconversion to 30% restored homogeneouscombustion for p P 5 bar, despite the fact that the water formed due to the upstream preconversioninhibited homogeneous ignition. The catalytically-produced water inhibited gas-phase combustion, par-ticularly at higher pressures, and this kinetic inhibition was exacerbated by the diffusional imbalance ofhydrogen that led to over-stoichiometric amounts of water in the near-wall hot ignitable regions. Radicaladsorption/desorption reactions hindered the onset of homogeneous ignition and this effect was morepronounced at 1 bar. On the other hand, over the post-ignition reactor length, radical adsorption/desorp-tion reactions significantly suppressed gas-phase combustion at 5 and 8 bar while their impact at 1 barwas weaker. By increasing hydrogen preconversion, the attained superadiabatic surface temperaturescould be effectively suppressed. An inverse catalytically stabilized thermal combustion (CST) concepthas been proposed, with gas-phase ignition achieved in an upstream porous burner via radiative and heatconduction feedback from a follow-up catalytic reactor. This arrangement moderated the superadiabaticsurface temperatures and required modest or no preheat of the incoming mixture.

� 2010 The Combustion Institute. Published by Elsevier Inc. All rights reserved.

1. Introduction

Combustion of hydrogen and hydrogen-rich fuels has attractedincreased interest in CO2 capture technologies for large powerplants. The ‘‘front-end” approach deals with fuel decarbonization,followed by combustion of a hydrogen-rich fuel mixture [1,2].For natural gas or gasified coal plants, in particular, by reformingthe fuel to syngas, using a shift reactor to convert CO to CO2, andfinally capturing the formed CO2 before combustion, high H2 con-tent (>80 vol.%) fuels can be cost-efficiently produced. Combustionof hydrogen is also of interest for ‘‘tail-end”, post-combustion CO2

capture strategies. Therein, exhaust gas recycle (EGR) is commonlyused to increase the CO2 content in the flue gas and thus facilitateits subsequent capture [3,4]. With increasing amounts of EGR,

ion Institute. Published by Elsevier

zaras).

however, the reactivity of the diluted fuel mixture is reduced, thuscompromising combustion stability at the desired operationalreactor temperatures [3]. An approach to overcome such stabilityissues is by adding H2 (chiefly produced via catalytic partial oxida-tion, CPO, of a fuel fraction) into the diluted reactive mixture [5,6].Apart from large-scale power generation applications, H2 and H2-rich fuels are also of interest in microreactors for portable powergeneration [7,8].

Combined catalytic (heterogeneous) and gas-phase (homoge-neous) combustion is intensively investigated for large turbines.The first studies focused on the catalytically stabilized thermalcombustion (CST) concept. Therein, part of the fuel is combustedcatalytically in Pd/Pt-coated honeycomb reactors operated underfuel-lean stoichiometry, while the remaining is converted in a fol-low-up homogeneous combustion zone – again under fuel-leanstoichiometry [9–11]. The fuel-lean CST concept, however, posesconcerns for reactive mixtures with large H2 content. This is due

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Nomenclature

b channel half-height, Fig. 1cp specific heat at constant pressuredh channel hydraulic diameterDkm mixture-average species diffusion coefficient, Eq. (6)DT

k species thermal diffusion coefficient, Eq. (6)Fk�j radiation configuration factor between surface elements

k and j, Eq. (11)h, ho

k total enthalpy, chemical enthalpy of the kth gaseousspecies, Eq. (3)

I unity diagonal matrix, Eq. (2)Kg total number of gaseous species, Eq. (4)L, LH catalytic reactor length, Figs. 1 and 3Ms total number of surface species, Eq. (5)N number of discretized channel elements, Eq. (11)p pressureqk radiant heat flux of surface element k, Eq. (11)R universal gas constant, Eq. (7)r, R radial coordinate and channel radius, Fig. 3bReIN inlet Reynolds number_sk heterogeneous molar production rate of kth species, Eq.

(8)T, To temperature and reference temperature, Eq. (7)u, UIN streamwise velocity component, inlet streamwise

velocityv transverse velocity component~Vk species diffusion velocity vector, Eq. (6)W catalytic channel width, Fig. 1Wk, W gas-phase species molecular weight, average molecular

weightXk, Yk mole fraction and mass fraction for kth gaseous species

x, y, z streamwise, transverse and lateral coordinates, Fig. 1

Greek symbolsC surface site density, Eq. (5)d channel wall thickness in honeycomb reactor, Fig. 3bej emissivity of surface element j, Eq. (11)hm surface species coverage, Table 1kg thermal conductivity of gas, Eq. (3)ks thermal conductivity of solid, Eq. (10)l viscosityq densityrm surface species site occupancy, Eq. (5)r Stefan–Boltzmann constant, Eq. (11)u equivalence ratio_xk homogeneous molar production rate of kth species, Eq.

(4)

SubscriptsIN inletig homogeneous ignitiong, s gas, solidk, m indices for gas-phase and surface speciesW wallx streamwise componenty, r transverse component in Cartesian and cylindrical coor-

dinates

AcronymsCST catalytically stabilized thermal combustionCPO catalytic partial oxidation

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1943

to the superadiabatic surface temperatures attained during cata-lytic combustion with deficient reactants having Lewis numbersless than unity [12–14]. A more recent hetero-/homogeneous com-bustion methodology is the ‘‘rich-catalytic/lean-gaseous”, involv-ing CPO of fuel-rich hydrocarbon/air mixtures to syngas followedby fuel-lean homogeneous combustion [5,15,16]. This approachcan also be applied to H2 or H2-rich syngas fuels. For such fuelsthe upstream catalyst does not have a prime CPO function, butrather acts as a preheater and stabilizer for the follow-up homoge-neous combustion zone. While hetero-/homogeneous combustionis an option for large-scale power generation, it appears to be thepreferred route for microreactors [17–20] due to the associatedlarge surface-to-volume ratios that favor catalytic fuel conversion,the multitude of gas-phase combustion instabilities at small con-finements [21–24] and their effective suppression in the presenceof a catalyst [25].

Recent studies have improved the understanding of homoge-neous kinetics for hydrogen and syngas fuels at elevated pressures[26–30]. On the other hand, there is a lack of corresponding high-pressure hetero-/homogeneous combustion studies for H2 or H2-rich fuels. Early numerical studies investigated the homogeneousignition of lean and rich H2/air mixtures over Pt at atmosphericpressure in stagnation flows [31,32]. Appel et al. [13] measuredmajor and minor species concentrations across the boundary layerof a Pt-coated channel-flow reactor, in H2/air atmospheric pressurecombustion; they provided validated hetero-/homogeneous reac-tion schemes for the combustion of fuel-lean H2/air mixtures overPt and clarified the underlying hetero-/homogeneous chemistrycoupling. More recently, Mantzaras et al. [33] extended the laststudy to pressures of up to 10 bar, for non-preheated lean H2/airmixtures. The intricate pressure and temperature dependence ofhydrogen homogeneous ignition chemistry [26] and the competi-

tion between the heterogeneous and homogeneous reaction path-ways led to suppression of gas-phase combustion for pressuresabove 4 bar, for non-preheated lean H2/air mixtures and surfacetemperatures as high as 1200 K [33].

The present works investigates the effect of fractional hydrogenpreconversion on the subsequent homogeneous ignition of fuel-lean H2/air mixtures over Pt surfaces. Fractional fuel preconversioncan be of general industrial interest. For example, in solid oxidefuel cells the unreacted H2-rich fuel at the cell exit must be subse-quently combusted; in autothermal chemical processes, a fractionof the fuel is converted in a dedicated burner to generate heat andan unconverted amount of fuel escaping this burner should besafely combusted; finally, in gas turbines there already exist two-stage, sequential combustion burners [3]. Preconversion resultsin reduced fuel and increased temperatures for the ensuing reac-tive mixture and further provides main combustion products andradicals over the gaseous induction zone. For catalytic combustionsystems, in particular, these factors can significantly affect the het-ero-/homogeneous chemistry coupling that leads to the onset ofhomogeneous ignition.

Experiments with fuel-lean H2/O2/N2/CO2 mixtures (equiva-lence ratio u = 0.30) have been performed in an optically accessi-ble, Pt-coated, catalytic channel reactor. Different fractionalhydrogen preconversions (ranging from 0% to 80%) were achievedby means of a dedicated honeycomb catalytic preburner positionedupstream of the main optically accessible reactor. The surface tem-peratures and pressures spanned the ranges 860 K 6 Tw 6 1250 Kand 1 bar 6 p 6 8 bar, respectively. The main reactor inlettemperatures and equivalence ratios varied between 310–810 Kand 0.078–0.30, respectively, depending on the fractionalpreconversion. One-dimensional Raman measurements providedthe boundary layer profiles of major species over the catalyst

1944 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

surface, and planar laser induced fluorescence (LIF) of the OH rad-ical monitored homogeneous combustion. Computations were car-ried out with an elliptic 2-D CFD code that included elementarycatalytic and gaseous chemical reaction schemes and detailedtransport. The main objectives were to assess the impact of hydro-gen fractional preconversion on the ensuing homogeneous ignitioncharacteristics over Pt surfaces and also on the attained superadi-abatic surface temperatures. Based on the outcome of the previousfundamental studies, a new reactor concept has been designed andtested for facilitating CST of lean H2/air mixtures. Therein, the or-der of combustion modules has been reversed, with the homoge-neous burner preceding the catalytic reactor.

This article is organized as follows. The experimental tech-niques and the numerical models are introduced in Sections 2and 3, respectively. In Section 4.1, comparisons between measure-ments and predictions are used to assess the impact of hydrogenpreconversion and pressure on homogeneous ignition. Pure homo-geneous ignition characteristics are numerically calculated in Sec-tion 4.2, while the effect of water and radical species – formedeither due to upstream fuel preconversion or due to the catalyticpathway – on homogeneous ignition is presented in Section 4.3.The thermal coupling between the two reaction pathways andthe moderation of the catalyst surface temperatures with increas-ing preconversion are discussed in Section 4.4. Finally, the newhetero-/homogeneous reactor concept is presented in Section 4.5.

2. Experimental

2.1. High-pressure optically accessible catalytic reactor

The test rig comprised an optically accessible channel-flow cat-alytic reactor and a catalytic honeycomb preburner, both posi-

O2, N2, CO2H2,p

Therfeed

Hsu

Flushingair Flow straigh

tening grids

H2,m

xy

Cross sectionA-A

zy

W

Quartzwindows

35

Inconel-steelframe

3

Thermo-couples

Preburner(Pt-coated

honeycomb)

Staticmixers

TCA

TCB

TTCC

(a)

(b)

Pre-heater

35

75

Fig. 1. (a) Schematic of the high-pressure test rig with the channel-flow catalytic reaccatalytic reactor. All distances are in mm.

tioned inside a cylindrical tank that provided the desiredpressurization (Fig. 1). The reactor, also used in earlier studies[13,33–36], consisted of two horizontal non-porous ceramic platesmade of Si[SiC] (L = 300 mm long, 110 mm wide, 9 mm thick, posi-tioned 2b = 7 mm apart) and two vertical quartz glass windows(3 mm thick, 12 mm high and 300 mm long) with a lateral separa-tion W = 104 mm (Fig. 1b). Plasma vapor deposition was used tocoat the inner Si[SiC] plate surfaces with Pt. A 1.5-lm thick non-porous Al2O3 layer was deposited first, followed by a 2.2-lm thickPt layer. Such a thick Pt layer resembled a polycrystalline surfaceand this was verified with Brunauer–Emmet–Teller (BET) total sur-face area measurements using Kr-physisorption, active area mea-surements using CO-chemisorption, and X-ray photoelectronspectroscopy (XPS), as described in [36]. BET showed the absenceof surface porosity whereas both BET and chemisorption revealedthe same total and active areas, suggesting that the surface wasonly covered with Pt. The latter was further confirmed by post-combustion XPS, which confirmed the absence of bulk Al or Si onthe catalyst surface.

The temperature along each catalytic plate was monitored with12 S-type (Pt-10%Rh/Pt) thermocouples, positioned axially alongthe x–y symmetry plane of the reactor (Fig. 1a). The thermocoupleswere embedded 0.9 mm beneath the catalytically coated surfaces,through 1.2-mm in diameter and 8.1-mm deep holes eroded fromthe uncoated, outer Si[SiC] surfaces. During fuel-lean catalyticcombustion of diffusionally imbalanced fuels with Lewis numbersless than unity (for lean hydrogen combustion LeH2 � 0:3) superad-iabatic surface temperatures are attained at the channel entry[13,37]. In a fashion similar to earlier hydrogen catalytic combus-tion studies [13,33], a combined cooling and heating arrangementwas adopted to suppress the high entry surface temperaturesand at the same time to balance the heat losses at the rear of

Powerfeedthroughs

Heatercoils

Pressurethrottle

LIF laser sheet

mocouplethroughs

Watercooling

Pt-coatedsurfaces

Exhaust

2b=7

Quartzwindows

igh pres-re vessel

L=300Insulation

-

Si[SiC]ceramicplates

=104 50

Ceramicsupportrims

A

A

Water cooling

CD

110

9

tor and the honeycomb catalytic preburner. (b) Cross-section of the channel-flow

Fig. 2. Schematic of the Raman and OH-LIF optical setups. All focal lengths are inmm.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1945

the reactor. Therefore, the front sides of the Si[SiC] plates(110 mm � 9 mm) contacted a water-cooled metal plate of thereactor support frame (see Fig. 1a), whereas the central and rearreactor length 100 < x < 300 mm was heated by two resistive coilsplaced 15 mm above each ceramic plate. An insulating porous fiberceramic material was positioned above the heater coils, while thereactor assembly was mounted inside an inconel steel supportframe.

To simulate fractional H2 preconversion, a honeycomb catalyticburner was positioned upstream of the main reactor (Fig. 1a). Theburner consisted of a cylindrical steel tube with an internal diam-eter of 35 mm and a length of 75 mm, wherein corrugated 50-lm-thick FeCr-alloy foils coated with a technical Pt/Al2O3 catalyst(5 wt.% Pt) created a honeycomb structure with individual channelhydraulic diameters dh = 1.2 mm. High-pressure bottles providedO2, N2, H2 and CO2. A small amount of CO2 (comprising �11 vol.%of the main reactor inlet composition) was added only at the high-est preconversion cases (80%) so as to better control the onset ofhomogeneous ignition and thus facilitate the laser-based measure-ments. All gases were supplied at room temperature (296 K).Hydrogen flows were supplied separately to the preburner andmain reactor, further denoted as H2,p and H2,m, respectively(Fig. 1a). The overall H2/O2 stoichiometry (accounting for both H2

flows) was fuel-lean with u = 0.30, while different preconversionscould be obtained by adjusting the H2,p and H2,m flowrates.

The O2/N2/(CO2) flows were preheated in the 30% and 80% pre-conversion cases, using an electric heater. This was to compensatefor the heat losses in the preburner and main reactor inlet section,which increased with rising preconversion (higher preburner tem-perature). The O2/N2/CO2 and H2,p flows were mixed in two sequen-tial static mixers before entering the honeycomb preburner. Thepreburner honeycomb length and inflow velocities were designedso as to generously guarantee complete conversion of H2,p. Thiswas also confirmed experimentally by running only the preburner(turning H2,m off) and observing the absence of a hydrogen Ramansignal in the main reactor. A sheathed K-type (Ni–Cr/Ni) thermocou-ple monitored the inlet temperature of the preburner (TCA in Fig. 1a),while a second similar thermocouple (TCB) ensured no accidentalignition upon injection of H2,m into the hot exhaust gases of thepreburner.

The H2,m main fuel stream was mixed with the preburnerexhaust gases in two sequential static mixers. After mixing, a20-cm long rectangular steel unit (cross-section 104 mm � 7 mm)having two cross-flow grids, yielded a uniform flow at the entry ofthe main reactor. The flows of O2, N2, H2,p, H2,m and CO2, were reg-ulated and measured by five dedicated Brooks mass-flow control-lers with accuracies better than ±0.2%. Two sheathed K-typethermocouples were positioned 40 mm and 2 mm upstream ofthe main reactor entrance (TCC and TCD, respectively, Fig. 1a).The indications of thermocouples TCB, TCC and TCD ensured noaccidental gas-phase ignition within the mixing and flow-straight-ening sections upstream of the main reactor. Thermocouple TCD

also provided the inlet temperature of the main reactor neededfor the numerical simulations. Radiation corrections were appliedto the indications of TCD since it had a partial view to the hot cat-alytic plates. The corrections in TCD even for the most demandingcases with 0% preconversion, and thus lowest inlet temperature,amounted to less than 8 K, given the particular levels of gas inletand maximum catalyst plate temperature (the correction magni-tude was similar to that in earlier methane CPO studies [38]).The accuracy of the corrected inlet gas temperature (TIN) measure-ments at TCD was better than ±6 K.

The high-pressure tank was a stainless steel structure 1.8 mlong and 0.28 m in internal diameter. It was equipped with two350-mm long, 50-mm high and 35-mm thick quartz glass win-dows, allowing optical access from both reactor sides (Figs. 1b

and 2). Two additional quartz windows, one at the reactor exhaustand a second one at the rear flange of the tank (Fig. 1a), provided astreamwise optical access for the LIF excitation beam.

2.2. Laser diagnostics

The planar OH-LIF and 1-D Raman arrangements are depicted inFig. 2. In the LIF experiments, the 532-nm second harmonic radia-tion of a pulsed Nd:YAG laser (Quantel TDL90 NBP2UVT3) pumpeda tunable dye laser (Quantel TDL90), whose output radiation wasfrequency-doubled to 285 nm. This beam had a pulse energy of0.5 mJ, low enough to avoid saturation of the A(v = 1) X(v0 = 0)OH transition. A cylindrical lens telescope and a 1-mm slit masktransformed the 285 nm beam into a thin laser sheet that propa-gated counterflow along the x–y symmetry plane of the reactor(Figs. 1a and 2). The fluorescence from both (1–1) and (0–0) OHtransitions at 308 and 314 nm, respectively, was collected at 90�through the reactor and side tank windows with an ICCD camera(LaVision Imager Compact HiRes IRO, 1392 � 1024 pixels). Channelareas of 100 mm � 7 mm were recorded on a 628 � 44 pixel areaof the ICCD detector chip. The camera was traversed axially tomap the entire 300 mm reactor extent. At each measuring location,400 images were averaged to increase the signal-to-noise ratio.The LIF was calibrated with absorption measurements performedwith the 285 nm beam crossing the reactor laterally (z-direction)through all four reactor and tank side-windows, as in [13,34,39].

The Raman setup was different from that used in previous H2

combustion studies [13,33]. The light source was a frequency-dou-bled Nd:YLF high repetition rate pulsed laser (Quantronix DarwinDuo) at 526.5 nm, operated at 1.5–2 kHz, with a pulse durationand energy of 130 ns and 37–43 mJ, respectively. The signal of20,000–40,000 pulses was integrated on the detector chip whenacquiring an image. The signal-to-noise ratio was thus increasedby a factor of 20 compared to previous experiments that usedthe same Nd:YAG laser for both LIF and Raman [33,34]. The526.5 nm beam was focused through the tank and reactor side-windows into a vertical line (�0.3 mm thick) by an f = 150 mmcylindrical lens. The focal line spanned the 7 mm channel separa-tion and was offset laterally (z = 15 mm) to increase the collectionangle and minimize thermal beam steering. Two f = 300 mm lensescollected the scattered light at a 50� angle with respect to the

1946 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

sending optical path and focused it on a 25 cm imaging spectro-graph (Chromex-250i) equipped with an ICCD camera (PrincetonInstruments PI-MAX1024GIII). The 1024 pixel and 256 pixel longICCD dimensions corresponded to wavelength and transverse (y)channel distance, respectively. The effective Raman cross-sections,which included transmission efficiencies, were evaluated byrecording the signals of pure hydrogen, O2/N2/CO2, and burnt gasesof known composition. Raman data were acquired at differentpositions by traversing axially a table supporting the sending andcollecting optics as well as the Nd:YLF laser (Fig. 2). The measure-ment accuracy was �8% for compositions as low as 0.5 vol.%; con-centrations lower than 0.5 vol.% entailed larger uncertainties.Raman data closer than 0.7 mm to both walls were discarded dueto low signal-to-noise ratios.

It is clarified that the Raman data could not assess the catalyticreactivity, as done in earlier hydrocarbon catalytic combustionstudies [35,36], since the high reactivity of hydrogen on Pt alwaysyielded transport-limited conversion for the present wall temper-atures. Nonetheless, these measurements were important in ensur-ing that no partial or total catalyst deactivation had occurredduring the experiments: such an event could potentially falsifythe gas-phase kinetics due the resulting near-wall hydrogen excess[40].

2.3. Homogeneous and heterogeneous reactor concept

The test rig consisted of a ceramic porous foam structure,wherein homogeneous combustion could be sustained, and a fol-low-up catalytic honeycomb reactor that ignited gas-phase reac-tions in the porous burner via upstream radiative and conductiveheat transfer (Fig. 3a). A single H2 supply was used, while the oxi-dizer stream comprised O2, N2 and CO2. The porous burner and cat-

Thermocouple

O2, N2, CO2

H2

Staticmixers

Pho

20

B CA

rx

Sym

heat conduqrad

qrad

UIN,TIN,YIN

Pt-

adiabati

Pre-heater

Porousfoam

(a)

(b)

Fig. 3. (a) Schematic of the high-pressure test rig with the homogeneous/heterogeneSimulation domain of an individual channel of the honeycomb reactor. All distances are

alytic honeycomb reactors were positioned inside the high-pressure tank described in Section 2.1. The porous foam was analumina-based ceramic disk (porosity 40%), 35 mm in diameterand 15 mm long. The catalytic honeycomb reactor (the same unitas the hydrogen preburner in Fig. 1a) was mounted inside a2.5 mm thick and 35 mm internal diameter holder steel tube. Theprocedure for the ignition of both homogeneous and heteroge-neous reactors will be discussed in Section 4.5.

The temperature at the inlet and outlet of the porous foam aswell as the honeycomb reactor temperatures were monitored bysix K-type sheathed thermocouples (A–F) in Fig. 3a. The beads ofthermocouples B to F were positioned at x = �10, 3, 37, 70 mmand 85 mm, with x = 0 denoting the beginning of the honeycomb.The carrying wires of the four thermocouples B–E were drivencounterflow into the reactor through four honeycomb channels.The three thermocouples inside the honeycomb structure (C, Dand E) provided neither the true surface temperature nor themean-gas temperature but rather a weighted average, which wasonly indicative of the local reactor temperature. Of the three mea-sured true local gas temperatures (A, B and F), radiation corrections(amounting up to 10 K) have been applied to thermocouples A andB that had a view to the reacting porous foam and honeycomb,respectively. The absolute accuracy of the gas temperature mea-surements was ±8 K at positions A and B.

3. Numerical

For the optically accessible channel-flow reactor in Fig. 1a, 2-Dsimulations along the x–y symmetry plane have been carried out.For the honeycomb reactor in Fig. 3a, the same numerical codesimulated a single channel as a cylindrical tube with internal diam-eter equal to the channel hydraulic diameter dh = 1.2 mm (Fig. 3b).

t-coatedneycomb

Dischargenozzle

D E F 35

L =75Holder tube

Thermocouples

R= 0.6

δ /2

L =75

metry axis

ction in solidqrad

catalyst

c outer wall

Insulation

H

H

ous combustion concept (porous foam and honeycomb reactor, respectively). (b)in mm.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1947

3.1. Governing equations

The governing equations were solved in their full elliptic formfor a steady laminar flow with hetero-/homogeneous reactions,either in Cartesian (optically accessible reactor) or cylindrical (hon-eycomb channel) coordinates. The general form of the equations is:

Continuity equation:

r � ðq~uÞ ¼ 0: ð1Þ

Momentum equations:

r � ðq~u~uÞ þ rp�r � l r~uþ ðr~uÞT � 23ðr �~uÞI

� �¼ 0: ð2Þ

Total enthalpy equation:

r � ðq~uhÞ þ r �XKg

k¼1

qYkhk~Vk � kgrT

!¼ 0: ð3Þ

Gas-phase species equations:

r � qYkð~uþ ~VkÞ � _xkWk ¼ 0; k ¼ 1; . . . ;Kg : ð4Þ

Surface species coverage equations:

rm_sm

C¼ 0; m ¼ 1; . . . ;Ms: ð5Þ

Buoyancy was insignificant for the present high Reynolds num-bers (see Section 4.1) and the narrow vertical channel gap (7 mm),and was thus neglected. The diffusion velocities ~Vk in Eqs. (3) and(4) were computed using mixture average diffusion, includingthermal diffusion for the light species H and H2 [41]:

~Vk ¼ �Dkmr½lnðYkW=WkÞ� þ DTk Wk=ðqYkWÞ

h irðln TÞ: ð6Þ

Finally, the system of equations was closed by the ideal gas andcaloric state laws:

p ¼ qRoTW

; hk ¼ hokðToÞ þ

Z T

To

cp;k dT; k ¼ 1; . . . ;Kg : ð7Þ

3.1.1. Optically accessible reactorThe interfacial boundary conditions for the gas-phase species

and temperature at the lower and upper catalytic walls (y = 0and y = 2b, respectively) become:

ðqYkVk;yÞy¼0 ¼Wkð_skÞy¼0; �ðqYkVk;yÞy¼2b ¼Wkð_skÞy¼2b ð8Þ

and

Tðx; y ¼ 0Þ ¼ TW;LðxÞ; Tðx; y ¼ 2bÞ ¼ TW;UðxÞ; ð9Þ

with TW,U (x) and TW,L (x) the temperature profiles of the upper- andlower-walls, respectively, fitted through the 12 thermocouple mea-surements of each plate. No-slip was applied for both velocity com-ponents at the gas–wall interfaces. The inlet conditions wereuniform profiles for the temperature TIN (measured at TCD,Fig. 1a), the axial velocity UIN and the species mass fractions. Atthe outflow (x = L), v = 0 was used for the transverse velocity andzero-Neumann conditions for all other scalars. The governing equa-tions were discretized with a finite volume scheme and the solutionwas obtained iteratively using a SIMPLER [42] method for the pres-sure–velocity field. An orthogonal staggered grid of 420 � 100points (in x and y, respectively) was used, with finer x-spacing clo-ser to the entrance and y-spacing closer to both catalytic walls.

3.1.2. Honeycomb channelThe domain in Fig. 3b simulated a single channel (R = 0.6 mm)

of the honeycomb (Fig. 3a) and further provided the numericalplatform for fundamental studies of the hetero-/homogeneous

thermal coupling (Section 4.4). A staggered grid of 162 � 40 points(in x and r, respectively) was used for the flow; 162 grid points(=N + 2) were also used for the solid, yielding N = 160 solid ele-ments. Axial heat conduction in the solid and surface radiationheat transfer were accounted for. Half of the FeCr-alloy wall thick-ness (d/2 = 25 lm) was included in Fig. 3b, due to consideration ofadjacent channels. For the same reason, the outer horizontal chan-nel surfaces in Fig. 3b were considered adiabatic. The solid energybalance becomes:

ks@2TW

@x2 ðd=2Þþ _qrad� kg@T@r

� �r¼R

þXKg

k¼1

ð_skhkWkÞr¼R

" #2R

2Rþd=2

� �¼0

ð10Þ

with ks = 16 W/mK, the thermal conductivity of the FeCr-alloy. Radi-ation feedback from the catalytic reactor to the inlet was importantfor the concept in Fig. 3a. Radiation heat transfer exchange betweenthe discretized catalytic surface elements themselves and betweeneach surface element and the reactor inlet and outlet was accountedfor by the net radiation method for diffuse-gray areas [43]. For thekth channel surface element, the radiation balance becomes:

1ek

qk ¼XNþ2

j¼1

Fk�jr T4k � T4

j

� �þXNþ2

j¼1

1� ej

ej

� �Fk�jqj; ð11Þ

where qk is the radiant flux, Fk�j is the configuration factor betweenthe finite surface elements k and j, and j runs over the N (=160)channel elements as well as the inlet (j = N + 1) and outlet(j = N + 2). The inlet and outlet sections were treated as black bodies(eN+1 = eN+2 = 1.0), with radiation exchange temperatures equal tothe inlet and mean exhaust gas temperatures, respectively. Allchannel elements had an emissivity ej = e = 0.6, j = 1, . . . , N.

The interfacial boundary conditions for the gas-phase speciesbecome:

ðqYkVk;rÞr¼R þ B_skWk ¼ 0; k ¼ 1; . . . ;Kg : ð12Þ

The factor B is the ratio of the catalytically active to the geomet-rical surface area of the technical Pt/Al2O3 catalyst. For the type ofcoating and preparation method used for the present technical cat-alyst, BET and H2-chemisorption have provided B values rangingbetween 4 and 10 [16,44]. B = 5 was used, however, the resultswere insensitive to the particular B value since, for the high reac-tivity of hydrogen on Pt, the catalytic conversion was trans-ported-limited. Radiative boundary conditions were applied tothe vertical ring-shaped front and rear solid channel faces:

ks@TW

@x

����x¼0¼ er T4

Wðx ¼ 0Þ � T4IN

h i; �ks

@TW

@x

����x¼L

¼ er T4Wðx ¼ LHÞ � T4

OUT

h i; ð13Þ

with TOUT the radially-averaged gas temperature at the outlet. Zero-Neumann conditions were applied to all thermoscalars and the ax-ial velocity at the centerline (r = 0) and outlet (x = LH).

3.2. Chemical kinetics

The detailed heterogeneous reaction scheme for the oxidationof H2 over Pt from Deutschmann et al. [45] (11 irreversible andthree reversible reactions, five surface and six gaseous species)was employed, with a platinum surface site density C = 2.7 �10�9 mol/cm2 (see Table 1). Surface thermodynamic data for thereversible reactions were taken from [46]. It is noted that CO2 doesnot participate in the surface reaction mechanism, since there is noadsorption of this species in the extended scheme for CH4/H2/COoxidation on Pt [45].

Table 1Catalytic reaction scheme for H2 oxidation on Pt.a

A (c) b E

Adsorption reactionsS1. O2 + 2Pt(s) ? 2O(s) 0.023 0.0 0.0S2. O2 + 2Pt(s) ? 2O(s) 1.8 � 1021 �0.5 0.0S3. H2 + 2Pt(s) ? 2H(s) 0.046 0.0 0.0S4. H + Pt(s) ? H(s) 1.0 0.0 0.0S5. O + Pt(s) ? O(s) 1.0 0.0 0.0S6. H2O + Pt(s) ? H2O(s) 0.75 0.0 0.0S7. OH + Pt(s) ? OH(s) 1.0 0.0 0.0

Surface reactionsS8. H(s) + O(s) = OH(s) + Pt(s) 3.7 � 1021 0.0 11.5S9. H(s) + OH(s) = H2O(s) + Pt(s) 3.7 � 1021 0.0 17.4S10. OH(s) + OH(s) = H2O(s) + O(s) 3.7 � 1021 0.0 48.2

Desorption reactionsS11. 2O(s) ? O2 + 2Pt(s) 3.7 � 1021 0.0 213.2–60hO

S12. 2H(s) ? H2 + 2Pt(s) 3.7 � 1021 0.0 67.4–6hH

S13. H2O(s) ? H2O + Pt(s) 1.0 � 1013 0.0 40.3S14. OH(s) ? OH + Pt(s) 1.0 � 1013 0.0 192.8

a From [45]. The surface site density is C = 2.7 � 10-9 mol/cm2. In the surface anddesorption reactions, the reaction rate coefficient is k = ATbexp (�E/RT), A(mol cm K s) and E (kJ/mol). In all adsorption reactions, except S2, A denotes asticking coefficient (c). Reactions S1 and S2 are duplicate. Reaction S3 has an orderof one with respect to platinum. The suffix (s) denotes a surface species and hi thecoverage of surface species i.

1948 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

For homogeneous chemistry, the scheme of Li et al. [26] wasused (21 reversible reactions and nine species, see Table 2) withits accompanying gas-phase thermodynamic data. This mechanismhas been validated against shock tube, flow reactor, and laminarflame speed measurements at pressures of up to 87 bar. The chem-ical effect of added CO2 appears in the third body efficiencies ofcertain reactions (Table 2). For the moderate temperatures of the

Table 2Homogeneous chemical reaction mechanism for H2.a

A b E

H2/O2 reactionsR1 H + O2 = O + OH 3.55 � 1015 �0.41 69.45R2 O + H2 = H + OH 5.08 � 104 2.67 26.32R3 H2 + OH = H2O + H 2.16 � 108 1.51 14.35R4 O + H2O = OH + OH 2.97 � 106 2.02 56.07

H2/O2 dissociation–recombinationR5 H2 + M = H + H + M 4.58 � 1019 �1.40 436.73R6 O + O + M = O2 + M 6.16 � 1015 �0.50 0.00R7 O + H + M = OH + M 4.71 � 1018 �1.0 0.00R8 H + OH + M = H2O + M 3.80 � 1022 �2.00 0.00

HO2 formation–consumptionR9 H + O2 + M = HO2 + M 1.48 � 1012 0.60 0.00

H + O2 + M = HO2 + M 6.37 � 1020 �1.72 2.18R10 HO2 + H = H2 + O2 1.66 � 1013 0.00 3.43R11 HO2 + H = OH + OH 7.08 � 1013 0.00 1.26R12 HO2 + O = O2 + OH 3.25 � 1013 0.00 0.00R13 HO2 + OH = H2O + O2 2.89 � 1013 0.00 �2.09

H2O2 formation–consumptionR14 HO2 + HO2 = H2O2 + O2 4.22 � 1014 0.00 50.12R15 HO2 + HO2 = H2O2 + O2 1.30 � 1011 0.00 �6.82R16 H2O2 + M = OH + OH + M 2.95 � 1014 0.00 202.63

H2O2 + M = OH + OH + M 1.20 � 1017 0.00 190.37R17 H2O2 + H = H2O + OH 2.41 � 1013 0.00 16.61R18 H2O2 + H = H2 + HO2 4.81 � 1013 0.00 33.26R19 H2O2 + O = OH + HO2 9.55 � 106 2.00 16.61R20 H2O2 + OH = H2O + HO2 1.00 � 1012 0.00 0.00R21 H2O2 + OH = H2O + HO2 5.80 � 1014 0.00 40.00

a From [26]. Reaction rate k = ATbexp (�E/RT), A (mol cm K s), E (kJ/mol). Thirdbody efficiencies in reactions R5–R8 and R16 are x(H2O) = 12.0, x(H2) = 2.5,x(CO2) = 3.8; in reaction R9 x(H2O) = 11.0, x(H2) = 2.0, x(O2) = 0.78, x(CO2) = 3.8.Reaction pairs (R14, R15) and (R20, R21) are duplicate. Reactions R9 and R16 areTroe reactions centered at 0.8 and 0.5, respectively (second entries are the lowpressure limits).

present study, reactions involving CO2 were insignificant as shownby comparing the simulations to those using an extended syngasmechanism [47]. Gas-phase and surface reaction rates were evalu-ated with CHEMKIN [48] and Surface-CHEMKIN [49], respectively.Finally, the CHEMKIN transport database [41] was used to calculateall transport parameters.

4. Results and discussion

The impact of fractional hydrogen preconversion on homoge-neous ignition is firstly assessed with the aid of the OH-LIF mea-surements. Detailed simulations clarify the effect of major andminor species formation on homogeneous ignition and on the het-ero-/homogeneous chemistry coupling. The attained surface tem-peratures are then numerically investigated in honeycombreactors as a function of the H2 preconversion. Finally, experimentsin the reactor concept of Fig. 3a follow suit and the advantages ofthe new combustion approach are outlined.

4.1. Comparison between experiments and predictions

The experimental conditions at the main reactor inlet areshown in Table 3. Pressures of 1, 5 and 8 bar were investigatedfor a fuel-lean stoichiometry u = 0.30, based on both H2,p andH2,m flows. Preconversions of 0%, 30% and 80% (1 and 5 bar) and0% and 30% (8 bar) were obtained by adjusting the H2,p and H2,m

flow rates; stable combustion at 80% preconversion and 8 bar couldnot be achieved due to autoignition in the mixing section. The totalmass throughput ( _m in Table 3) was not kept constant, but it in-creased for the 30% and 80% preconversions, so as to avoid flameflashback for these higher inlet temperature cases. This arrange-ment also exemplified the impact of preconversion on homoge-neous ignition as will be further discussed in Section 4.2. Theinlet Reynolds numbers (ReIN in Table 3), calculated from the uni-form inlet properties and the channel hydraulic diameter(=13.1 mm), ranged from 1680 to 2652. Nonetheless, all casescould be treated with a laminar model since recent turbulent cat-alytic combustion studies [50] have shown that the strong flowlaminarization induced by the hot catalytic walls guaranteed lam-inar conditions at ReIN considerably higher than 3000.

The measured gas temperatures at the inlet of the main reactor,TIN, ranged from 310 to 811 K. For the 0% preconversion Cases 1, 4and 7, TIN was up to 16 K higher than the 296 K flows of O2, N2, andH2 due to upstream conduction of heat in the solid walls of theflow-straightening section. The computed temperatures at thereactor inlet, Tad,IN, accounting for adiabatic equilibrium combus-tion in the preburner (for H2,p and O2/N2/CO2 flows at 296 K) andsubsequent adiabatic, infinitely fast mixing (without chemicalreactions) of the combustion products with the 296 K H2,m stream,are also shown in Table 3. The O2/N2/CO2 preheat has been accord-ingly adjusted to compensate for the heat losses and thus achievecomparable values for TIN and Tad,IN, mainly for the 30% preconver-sion case. For the 80% preconversion cases, TIN were kept �80 Klower than Tad,IN to eliminate the danger of flame flashback. The in-let species composition in Table 3 was computed by consideringinfinitely fast mixing of the preburner adiabatic equilibrium prod-ucts with the 296 K H2,m stream. Radicals were not considered, buttheir inclusion had a small impact for most cases, as will be shownin Section 4.3.

Comparisons between measured and predicted OH distribu-tions (ppmv) are illustrated in Fig. 4 for all conditions in Table 3.The 2-D OH maps have been constructed by using successive,100-mm-long overlapping LIF images. The modest asymmetry ofthe flames (e.g., in Fig. 4(1a and 1b) and (2a and 2b)) is due to tem-perature differences between the upper and lower catalytic walls.

Table 3Experimental conditions for optically accessible reactor.a

Case p (bar) Preconv. (%) uIN H2 (%) O2 (%) H2O (%) CO2 (%) UIN (m/s) ReIN TIN (K) Tad,IN (K) _m (g/s)

1 1 0 0.30 10.9 18.5 – – 2.4 1680 310 296 1.82 1 30 0.23 7.9 17.2 3.4 – 9.6 2427 572 572 3.93 1 80 0.077 2.0 13.0 8.1 11.6 14.2 2231 811 886 4.54 5 0 0.30 10.9 18.5 – – 0.5 1750 310 296 1.85 5 30 0.23 7.9 17.2 3.4 – 2.2 2652 588 572 4.36 5 80 0.078 1.9 12.1 7.7 11.1 3.0 2395 799 886 4.87 8 0 0.30 10.9 18.5 – – 0.4 2241 312 296 2.48 8 30 0.23 7.8 16.9 3.3 – 1.2 2406 575 572 3.9

a Pressure, fractional H2 preconversion, equivalence ratio and species volumetric composition at the entry of the main reactor (the balance is N2), velocity, Reynoldsnumber, measured temperature at the inlet of main reactor, the theoretical adiabatic inlet temperature, and the mass throughput.

0 64 128

0 335 670

0 484 968

7 mm

p = 1 bar

0 0.9 1.8

0 1.5 3.0

0 150 22575x (mm)

0 0.5 1.0

0 0.6 1.2

(1a)

0%

30%

80%

(1b)

(2a)

(2b)

(3a)

(3b)

(4a)

(4b)

(5a)

(5b)

(6a)

(6b)

(7a)

(7b)

(8a)

(8b)

p = 5 bar

p = 8 bar

30%

0%

80%

30%

0 1.2 2.4

0%

300

Fig. 4. (a) LIF-measured and (b) numerically-predicted OH maps for all cases inTable 3. The arrows in Cases 1–3, 5 and 8 define the onset of homogeneous ignition.The provided scales in the color bars span the minimum and maximum predictedOH in ppmv.

0.0

0.4

0.8

800900100011001200

0.0

0.6

1.2

1.8

800900100011001200

0.0

0.9

1.8

2.7

800900100011001200

G

G

C

C

C

p=5bar

0.0

0.4

0.8

Wal

l tem

pera

ture

(K)

800900100011001200

G

0.0

1.0

2.0

800900100011001200

0.0

1.0

2.0

800900100011001200

G

G

C

C

C

(1) 0% preconv.

p = 1bar

Tw

CG

Hyd

roge

n co

nver

sion

rate

(g/m

2 s) p = 5bar

0.0

0.8

1.6

2.4

800900100011001200

0 50 100 150 200 250 3000.0

0.8

1.6

2.4

800900100011001200

G

G

C

C

p = 8bar

x (mm)

Tw

Tw

Tw

Tw

Tw

Tw

Tw

xig

xig

xig

xig30% preconv.

(2)

(3) 80% preconv.

(4) 0% preconv.

(5) 30% preconv.

(6) 80% preconv.

(7) 0% preconv.

(8) 30% preconv.

Fig. 5. Temperature profiles of upper-wall (dashed-dotted lines) and lower-wall(dashed double-dotted lines), fitted through the thermocouple measurements(upper-wall: triangles; lower-wall: circles) for all cases in Table 3. The computedcatalytic (C) and gas-phase (G) hydrogen conversion rates are shown with solid anddashed lines, respectively. The arrows in Cases 1–3, 5 and 8 define the onset ofhomogeneous ignition.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1949

The maximum local temperature differences between the twowalls ranged from 10 K (Case 5) to 74 K (Case 4), as seen fromthe temperature profiles in Fig. 5. The homogeneous ignition loca-tion (xig), marked with arrows in Fig. 4, for all cases where a strongflame was established (Cases 1–3, 5 and 8), has been defined inboth predictions and experiments as the far upstream locationwhere OH reaches 5% of its maximum value inside the reactor.The flames in Fig. 4 form two separate branches close to the chan-

nel walls, which was a result of the diffusional imbalance of hydro-gen [13].

(a) Case 1 (0 % preconversion, p = 1bar)

x=10mm0.0

3.5

7.0

Mole fraction H2, H2O

Cha

nnel

hei

ght y

(mm

)

0.0

3.5

7.0

0.0

3.5

7.0

0.1 0.2 0.1 0.20.0 0.1 0.2

(b) Case 2 (30% preconversion, p = 1bar)

(c) Case 6 (80% preconversion, p = 5bar)

(a1) (a2)x=40mm

(a3)x=75mm

x=10mm(b1) (b2)

x=40mm(b3)

x=75mm

x=10mm(c1) (c2)

x=40mm(c3)

x=75mm

2*H2 2*H2 2*H2

H2H2O

Fig. 6. Raman-measured (symbols) and numerically-predicted (lines) transverseprofiles of H2 and H2O mole fractions at three selected axial positions for (a) Case 1,(b) Case 2 and (c) Case 6 in Table 3. H2 (solid lines, triangles), H2O (dashed-lines,circles). The axial positions are: x = 10, 40 and 75 mm. For clarity, in (c) the H2 molefraction is multiplied by a factor of two.

2 3 4 5 6 7 8 9 101

Inve

rse

igni

tion

dela

y (s

-1)

101

102

103

104

105

Pressure (bar)

1000 K

1050K

1100K

1150K

950K

1200K1250 K

Fig. 7. Predicted inverse ignition delays for a u = 0.30 H2/air mixture in a pure gas-phase, constant pressure batch reactor, at different pressures and temperatures.

1950 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

Axial profiles of the computed catalytic (C) and gaseous (G)hydrogen conversion rates are shown in Fig. 5. The C plots ac-counted for the contribution of both catalytic surfaces, while theG plots were constructed by integrating the volumetric gaseousreaction rates over the 7 mm channel height. To facilitate compar-isons of the C and G curves, the maxima in the axes of the conver-sion rate plots in Fig. 5 have been scaled with the inlet H2 massthroughputs ðYH2 ;IN

_mÞ of each case. The predicted xig in Fig. 4 arealso provided in Fig. 5, clearly showing that the rise in OH corre-sponds to the rise in homogeneous hydrogen conversion G.

The experimentally deduced maximum OH levels were within20% of the predicted values for all cases where a strong flame couldbe established in the reactor, thus allowing calibration of the LIFsignal with absorption measurements. Hence, a good agreementwas established between measurements and predictions in termsof homogeneous ignition location, flame shapes and absolute OHlevels. This is a first important outcome of the present study,clearly demonstrating the aptness of the employed gas-phase reac-tion mechanism in the presence of catalytic reactions, with highreactant preheats and H2 preconversions, hence extending earlierresults without preheat/preconversion [33]. For the atmosphericpressure Cases 1–3, strong OH signals could be detected for all pre-conversions (Fig. 4). However, for higher pressures and 0% precon-version (Cases 4 and 7), no OH signal could be detected above thethermal radiation noise of the hot catalytic plates. The maximumpredicted OH in Cases 4 and 7 (2.4 and 1.0 ppmv, respectively)were too low for detection with planar LIF, when also consideringthe thermal radiation interference from the hot catalytic walls. Thesuppression of gaseous combustion for the p P 5 bar Cases 4 and 7,is identical to that reported in [33]. The present study, however,shows that a 30% preconversion of H2, with its resulting higher in-let temperatures, leads to appreciable gas-phase combustion atelevated pressures. The reason for this behavior and the implica-tions for a hetero-/homogeneous system will be clarified in Sec-tions 4.2 and 4.3.

The Raman measurements were important in assessing the cat-alytic processes preceding homogeneous ignition. An incorrect pre-diction of the catalytic fuel conversion over the gas-phaseinduction length could greatly affect the homogeneous ignitionlocation [40] and thus falsify the gaseous kinetics. Comparisons be-tween Raman-measured and predicted transverse profiles of H2

and H2O are shown in Fig. 6 for three cases in Table 3 at three axialpositions. The numerical predictions at the far upstream positionx = 10 mm indicate a catalytic conversion close to the transport-limit, evidenced by the very low hydrogen levels near both walls(y = 0 and 7 mm). This behavior is also captured by the measure-ments, despite the lack of the Raman data at distances closer than0.7 mm to both walls. The presence of gaseous combustion for Case1 in Fig. 6(a3) is manifested by the lack of H2 over extended zonesnear both walls and by the corresponding leveling of the H2O pro-files. The same behavior is also attested in Fig. 6(a2), whereby gas-eous combustion predominantly occurs close to the upper-wall(y = 7 mm) due to the delayed ignition at the lower-wall flamebranch (Fig. 4(1a and 1b)). Transport-limited hydrogen catalyticconversion is finally evident for all locations of the 80% preconver-sion Case 6 in Fig. 6c, whereby the homogeneous combustion con-tribution is minimal (Fig. 5(6)).

4.2. Computed pure homogeneous ignition characteristics

Pure gas-phase ignition characteristics of hydrogen are firstlyexamined. Ignition delays for a u = 0.30 H2/air mixture were com-puted in a constant pressure batch reactor using the SENKIN pack-age of CHEMKIN [51] for initial temperatures ranging from 950 to1250 K. Such high initial temperatures mimic the presence of thehot catalytic walls, which significantly preheat the flowing reacting

gas. The inverse of the ignition delays – a quantity proportional tothe hydrogen gaseous reactivity – are plotted in Fig. 7 as a functionof pressure with parameter the temperature. At a moderate initialtemperature of 950 K, the gaseous reactivity initially decreasesrapidly with rising pressure, while above 2 bar it only changesmodestly. At 1000 K 6 T 6 1200 K, the gaseous reactivity initiallyincreases with rising pressure and then drops, with the turningpoint shifted to higher pressures for higher temperatures. Finally,at T = 1250 K there is a monotonic increase of reactivity from 1to 10 bar. This rich behavior of hydrogen ignition characteristicsis due to the competition between the chain branching reactionH + O2 = O + OH the chain terminating reaction H + O2 + M =HO2 + M, and also of the chain branching sequence HO2 + H2 =

Fig. 8. Predicted inverse ignition delays for an initial u = 0.30 H2/air mixture as afunction of H2 preconversion for three pressures (circles: 1 bar, triangles: 5 bar,diamonds: 8 bar). The resulting temperatures and equivalence ratios at various H2

preconversions are also shown.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1951

H2O2 + H and H2O2 + M = 2OH + M that overtakes the stability ofHO2 in the termination step at higher temperatures [52].

For the spatially inhomogeneous channel-flow combustionexperiments in Fig. 4, the effective temperatures shown in Fig. 7are a weighted average between TIN and Tw with a likely addedweight on the latter. For the 0% preconversion cases (TIN � 310 K)with corresponding wall temperatures ranging between 900 and1200 K over the induction length x < xig in Case 1 (Fig. 5(1)) or overthe length whereby weak homogeneous conversion is attained inCases 4 and 7 (Fig. 5(4, 7)), the resulting effective temperaturesare moderate enough, such that the gaseous reactivity is highestfor pressures close to 1 bar, dropping rapidly for pressures above2–3 bar (see e.g. the 950 and 1000 K curves in Fig. 7). The implica-tion for a hetero-/homogeneous combustion system is that the cat-alytic pathway has the opportunity to consume significantamounts of H2 during the elongated gas-phase induction zones atp P 5 bar, depriving fuel from the gaseous pathway and thus sup-pressing flame formation for Cases 4 and 7. The suppression ofhomogeneous combustion for p P 5 bar at 0% preconversion is anoutcome not only of intrinsic H2 gas-phase kinetics but also ofcompetition between gaseous and catalytic reactions. The catalyticpathway is very effective in consuming H2 due the large moleculardiffusivity and the very high reactivity of this species on platinum,even at moderate surface temperatures.

For 30% preconversion (uIN = 0.23), TIN is high enough suchthat the gaseous reactivity drops only modestly with rising pres-sure according to Fig. 7. The result is the formation of noticeablystronger flames for the 30% preconversion Cases 5 and 8 (seeFig. 4(5) and (8) as well as Fig. 5(5) and (8)). Moreover, the 30%preconversion flames in Cases 5 and 8 have been established de-spite the fact that the mass throughput has been substantially in-creased (the reactor residence times have decreased) compared tothe 0% preconversion Cases 4 and 7 (see Table 3). Cases 5 and 8exhibit combined catalytic C and gaseous conversion G, evenwell-downstream of homogeneous ignition (Fig. 5(5) and (8)).This is in contrast to the p = 1 bar Case 2, whereby G conversiondominates shortly after homogeneous ignition (Fig. 5(2)). Such acombined catalytic and gaseous combustion is a result of theinterplay between transport and gaseous reactivity [13,50]: theweaker gaseous combustion at 5 and 8 bar compared to 1 bar al-lows for hydrogen leakage through the near-wall gaseous reactionzone and for subsequent catalytic conversion of the escaping fuelon the channel walls.

Fig. 7 has considered only the effect of rising temperature on theignition characteristics of a constant u mixture. The combined ef-fects of rising temperature, drop in H2 concentration, and presenceof major and minor products with increasing preconversion, areillustrated in Fig. 8. Initial conditions for the 0% preconversion cor-respond to a u = 0.30 H2/air mixture at 700 K (this temperaturemimics the average between TIN and Tw for the inhomogeneousreactor results in Fig. 4). Ignition delays at a given preconversionhave been subsequently computed using as initial condition theadiabatic equilibrium temperature and composition (the latterincluding radicals) of the preconverted H2 fraction, mixed adiabat-ically and infinitely fast with the 700 K remaining H2. For a givenpressure, the H2 reactivity clearly increases with rising preconver-sion. Furthermore, at intermediate preconversions (<50%), thereactivity is highest at 1 bar whereas at larger preconversions thereactivity is highest at 5 and 8 bar. Direct comparisons of these re-sults with the inhomogeneous, non-adiabatic, combined gaseous/catalytic combustion experiments in Fig. 4 cannot be made. None-theless, the sharp increase in reactivity with rising preconversion,particularly for the 5 and 8 bar plots in Fig. 8, is in qualitativeagreement to the findings in Fig. 4, whereby the suppression ofgaseous combustion for 0% preconversion at 5 and 8 bar has largelybeen removed for 30% preconversion.

The impact of the water and radicals present in the reactivemixture due to preconversion is calculated next. The results ofFig. 8 have been computed anew by replacing the initial H2O byan artificial species H2O*, which had the same thermodynamicproperties as H2O but did not participate in any chemical reaction.Furthermore, Fig. 8 was recomputed by removing all minor speciesOH, H, O, HO2 and HO2. The results are summarized in Fig. 9, pro-viding the ratios of the inverse ignition delays to those in Fig. 8. Theinlet water (Fig. 9a) clearly inhibits the onset of ignition, at least forpreconversions greater than 20%. This inhibition is particularlystrong for intermediate preconversions (40–60%) and higher pres-sures. Radical species significantly accelerate ignition (Fig. 9b), par-ticularly at high preconversions (>40%). Removal of the OH radicalalone, accounts for more than 99% of the ignition delay increaseshown in Fig. 9b.

4.3. Hetero-/homogeneous chemistry coupling

This section analyzes numerically the chemical impact of cata-lytically-produced major and minor species on gas-phase ignition.The catalytically-induced near-wall depletion of hydrogen (seeFig. 6) inhibits homogeneous ignition, as discussed in Section 4.2and shown theoretically in [40]. However, this is a masked trans-port and not a true chemical effect, since the catalytic conversionis practically transport-limited (see Fig. 6).

4.3.1. Impact of water on homogeneous ignitionThe 30% preconversion Cases 2, 5 and 8 have been recomputed

for two different conditions. At first, the inlet H2O due to H2 pre-conversion has been substituted by an artificial species H2O*,which had the same thermodynamic and transport properties asH2O but did not participate in any catalytic or gas-phase chemicalreaction. In a second step, the inlet water was restored to normalH2O, but the hetero-/homogeneous H2O coupling was switchedoff by replacing the gaseous H2O involved in the adsorption/desorption reactions S6 and S13 (see Table 1) with H2O*. Fig. 10provides computed gas-phase H2 conversion rates G for the afore-mentioned two different treatments of H2O, along with the original

(a)

(b)

Fig. 9. (a) Predicted ratios of inverse ignition delays calculated with chemically-inert initial H2O* to those calculated in Fig. 8 with reacting H2O. (b) Predicted ratiosof inverse ignition delays calculated without the inclusion of radicals to thosecalculated in Fig. 8 with radical inclusion, for various H2 preconversions. Pressure:1 bar (black bars), 5 bar (gray bars) and 8 bar (dark gray bars).

0.0

1.0

2.0

x (mm)0 50 100 150 200 250 300

H2

gas-

phas

e co

nver

sion

rate

(g/m

2 s)

0.0

0.5

1.0

0.0

1.0

2.0

(a) Case 2 1 bar, 30% preconv.

(b) Case 5 5 bar, 30% preconv.

(c) Case 8 8 bar, 30% preconv.H2O (ad-des)*

H2O (IN)*

Fig. 10. Predicted axial profiles of transversely-integrated gaseous hydrogenconversion rates, G, for the three cases in Table 3 with 30% H2 preconversion: (a)Case 2, 1 bar, (b) Case 5, 5 bar, and (c) Case 8, 8 bar. Solid lines: G rates from Fig. 5;dashed lines: G rates computed when the inlet water (Table 3) is substituted withchemically-inert H2O*; dashed-dotted lines: G rates computed when the gas-phaseH2O in the adsorption–desorption reactions S6 and S13 (Table 1) is replaced withchemically-inert H2O*.

1952 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

G conversion rates (Fig. 5). The chemical impact of the inlet wateris appreciable for all cases and inhibits homogeneous ignition, inagreement with the pure gas-phase studies of Section 4.2 (Fig. 9a).

The impact of catalytically-produced water is much strongerthan that of the inlet water, particularly for the higher pressureCases 5 and 8 (Fig. 10(b and c)). The preferential diffusion of hydro-gen leads to an effective surface equivalence ratio roughly twice

that of the bulk gas-phase, as shown in [13,32]. This effect isresponsible for the superadiabatic surface temperatures discussedin Section 2.1 and also for the production of water at the gas–wallinterface in amounts roughly twice as high as the ones expected bystoichiometric considerations alone. For example, the inlet compo-sition in Table 3 for the 30% preconversion Cases 2 or 5, shouldhave dictated a volumetric H2O content at the catalyst surfaceequal to 11.7% (when H2 is completely converted at the catalyst).While nearly complete conversion of H2 at the catalyst surface isachieved (see the vanishingly small wall levels of H2 inFig. 6(b1)), the predicted volumetric H2O content in Fig. 6(b1) is19.0% and 18.5% at the lower- and upper-walls, respectively. Thewater-induced inhibition is further accentuated by the fact thatgaseous combustion is initiated and sustained in the near-wallhot ignitable zone whereby the H2O content is the highest (seeFig. 6). An increase in pressure intensifies the H2O inhibition asseen in Fig. 10, due to the increasing importance of the reactionH + O2 + M = HO2 + M at elevated pressures. Homogeneous ignitioninhibition due to catalytic water formation has also been reportedin stagnation flow simulations of atmospheric pressure H2/air com-bustion over Pt [31,53]. Finally, it can be shown that the impact ofcatalytically-produced water is still appreciable at preconversionsas high as 80%.

4.3.2. Effect of radicals on homogeneous ignitionAll cases in Table 3 have been recomputed with the inclusion of

radicals, by mixing infinitely fast the full equilibrium compositionof the preburner products to the H2,m stream. This assumption cer-tainly entails inaccuracies, since gas-phase and wall radical recom-bination reactions are deemed to occur in the mixing and flow-straightening sections upstream of the reactor. Nonetheless, thisresulted in a conservative upper estimate for the radical concentra-tion at the inlet. Calculations have shown that the homogeneousignition predictions in Fig. 4 were not impacted by the presenceof radicals, a reassuring result given the lack of any informationon the inlet concentrations of minor species for the 30% and 80%preconversion cases. This is in contrast to the results of Fig. 9b,the reason being that the temperatures determining the impactof the incoming radicals for the spatially inhomogeneous reactorin Fig. 1 are the inlet temperatures, which are in turn substantiallylower than the initial temperatures used for the batch reactor sim-ulations in Figs. 8 and 9. For the 80% preconversion Case 6 reactionflux analysis reveals that reaction H2 + OH = H2O + H (R3 in Table 2)depletes a very small fraction of the inlet hydrogen mass flow rate(�0.01%) over the initial 0.02 mm length and also practically con-sumes the entire amount of inlet OH (the catalytic pathway hasno impact on the depletion of OH over the initial 0.02 mm length).Inlet radicals have an appreciable effect only for the 1 bar 80% pre-conversion Case 3, whereby over the initial �1 mm reaction R3amounts to 8% conversion of the incoming H2 mass flow rate.Nonetheless, the OH distribution and the main ignition locationxig in Fig. 4(3b) are hardly affected, since gas-phase combustionis initiated in the near-wall zone where the H2 levels are alreadyvery low due to catalytic depletion.

The impact of radical adsorption and desorption reactions is fi-nally discussed. All cases in Table 3 have been recomputed byremoving the adsorption of H and O and the adsorption/desorptionof OH (reactions S4, S5, S7 and S14). The new calculations indicateda promotion of homogeneous ignition for all cases, which wasmore pronounced p = 1 bar. For example, in Cases 1, 2 and 3 theignition distances reduced to xig = 8 mm, 24 mm and 98 mm,respectively, compared to the original xig = 22 mm, 48 mm and124 mm in Fig. 4. At higher pressures, however, radical adsorp-tion/desorption reactions had a weaker impact. For example, inCase 5 (5 bar), the new computations yielded xig = 66 mm com-pared to xig = 73 mm in Fig. 4, while for Case 8 (8 bar) the corre-

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1953

sponding numbers were xig = 62 mm and 70 mm. Although thelocation of homogeneous ignition is only mildly affected at 5 and8 bar, the ensuing gas-phase combustion at x > xig is strongly sup-pressed by the presence of radical adsorption/desorption reactions.Earlier studies have shown that the inhibition of homogeneousignition comes mainly via the OH radical adsorption/desorptionreactions S7 and S14 [13]. Moreover, detailed study of the OHadsorption/desorption and gas-phase reactions has shown[13,36,54] that while the catalytic pathway is a poor producer ofOH radicals so as to meaningfully affect the gaseous ignition chem-istry, the catalyst itself becomes an efficient sink of homoge-neously-produced OH. This combined behavior leads to theoverall inhibition of homogeneous ignition.

The origin of this inhibition is illustrated in Fig. 11, providingcomputed transverse OH profiles near the lower-wall for Cases 2and 5 in Table 3, at selected streamwise positions x / xig. Forthe 5 bar Case 5 (Fig. 11b), over most of the induction zone(x < 60 mm) the OH wall fluxes are net-desorptive (negative dXOH/dy at y = 0) turning to net-adsorptive (positive slope) at x P 60 mm,i.e. sufficiently close to the ignition position x = 73 mm; this deter-mines the weak inhibition of homogeneous ignition. However, thenet-adsorptive OH wall fluxes at post-ignition locations x > xig

strongly suppress gaseous combustion. Characteristically, the inte-grated G rates for 30% preconversion at 5 and 8 bar increase by afactor of 3 and 10, respectively, when the adsorption/desorptionreactions of OH are removed. On the other hand, for the p = 1 barCase 2 the OH wall fluxes are always net-adsorptive (Fig. 11a), sincethe higher homogeneous reactivity at 1 bar allows for a faster build-up of radicals in the gas-phase that in turn shifts the OH adsorption/desorption equilibria to adsorption. The result is a stronger inhibi-tion of homogeneous ignition via OH adsorption/desorption reac-tions at p = 1 bar. However, once ignited, gas-phase combustion isvigorously sustained for x > xig since the reactivity at 1 bar is highand not appreciably affected by the catalytically-induced loss ofOH radicals.

4.4. Hetero-/homogeneous thermal coupling

The thermal coupling between the heterogeneous and homoge-neous pathways has been previously suppressed by using aprescribed wall temperature for the computations in Fig. 10. More-over, the experiments in Fig. 4 pertained to a highly non-adiabatic

0.0 1.0e-5 2.0e-5 3.0e-5

y (m

m)

0.0

0.1

0.2

OH mole fraction0 2e-7 4e-7 6e-7 8e-7

y (m

m)

0.0

0.1

0.2

(a) Case 2 1 bar

(b) Case 5 5 bar

x =1mm

x =5mm

x=10mm

x=20mm

x=30mm

x=40mm

10mm

x=40mm

x=50mm

x =60mmx =70mm

x =75mmx=

Fig. 11. Computed transverse profiles of OH mole fraction, extending 0.2 mm fromthe lower catalytic wall (y = 0), for (a) Case 2 and (b) Case 5 in Table 3.

system with an external heating/cooling arrangement that maskedthe impact of chemical heat release. For this reason, computationsat three different preconversions (0%, 30% and 80%) and pressures(1, 5 and 8 bar) were performed using as numerical platform thegeometry in Fig. 3b, at flow conditions relevant to gas turbines(Table 4). For the 1 bar 0% preconversion Case H1, TIN = 310 K. For30% and 80% preconversions, the inlet temperatures werecalculated by considering adiabatic equilibrium combustion ofthe converted H2 (radicals included) followed by adiabatic, infi-nitely fast mixing with the remaining 310 K H2. The mass through-put, qINUIN, was held constant as the inlet pressure or temperatureincreased. This approach facilitated comparisons at different pres-sures and same preconversions, by maintaining the same inlet ReIN

(Table 4): analytical studies have shown [40,55] that for the sameReIN the catalytic fuel conversion and the onset of homogeneousignition only depend on the hetero-/homogeneous kinetics (whenother parameters, e.g. geometrical and transport, are the same).Wall temperatures are presented in Fig. 12 when only catalyticreactions are included and when both catalytic and gas-phase reac-tions are present.

4.4.1. Surface temperatures due to catalytic reactionsAt 0% and 30% preconversions the wall temperatures computed

with heterogeneous chemistry alone exhibit a strong superadiabat-ic peak close to the entry (Fig. 12). For pure catalytic combustionthe theoretically maximum attainable wall temperature obtainedunder adiabatic and transport-limited fuel conversion is [37]:

TW ¼ T IN þ Le�2=3DT; ð14Þ

with Le the Lewis number of the deficient reactant (hydrogen) andDT = Tad � TIN the adiabatic combustion temperature rise. WithLeH2 � 0:32 for lean H2/air combustion, the 0% preconversionCase H1 gives DT = Tad � TIN � 888 K, such that Eq. (14) yields amaximum attainable wall temperature of �2208 K. The predictedpeak temperatures for the 0% preconversion cases with pure cata-lytic combustion range from 1649 to 1670 K, and are �550 K lowerthan the theoretical maximum value, mainly due to radiation heatlosses towards the colder inlet. The wall temperature drops whenapproaching x = 0, due to more pronounced radiation heat losses to-wards the entry and also due to finite-rate surface chemistry effectsat x � 0. The radiation losses are stronger for the 0% preconversioncases due to their higher wall temperature and the correspondinglower inlet temperature. At the rear end of the reactor, the walltemperatures should theoretically approach Tad when the reactoroperates adiabatically, without redistribution of thermal energyvia solid heat conduction and with complete conversion of H2 atthe outlet [54]. The wall temperatures at x = 75 mm are somewhathigher than Tad mainly due to transfer of heat from the upstreamhotter parts via conduction in the wall and secondly due to a smallH2 breakthrough at the reactor exit. The peak wall temperatures inthe pure catalytic cases with 0% preconversion are 1649 K (CaseH1), 1668 K (Case H4) and 1670 K (Case H7), while the front walltemperatures (x = 0) for these cases are 1337 K, 1448 K and1464 K, respectively. The reason for the lower temperatures in CaseH1 (1 bar) is the enhancement of the catalytic reactivity withincreasing pressure, as also shown in a recent numerical investiga-tion of syngas combustion over Pt [14]. Peak wall temperatures of�1650 K, such as those shown in Fig. 12 for 0% preconversion, cancause reactor meltdown and catalyst deactivation. For the pure cat-alytic combustion cases, an increase in hydrogen preconversionappreciably reduces the maximum attained wall temperature, thusgreatly facilitating reactor thermal management.

Table 4Conditions for honeycomb channel simulations.a

Case p (bar) Preconv. (%) uIN H2 (%) O2 (%) H2O (%) Radicals (%) UIN (m/s) ReIN TIN (K)

H1 1 0 0.30 11.2 18.6 – – 40.0 2558 310.0H2 1 30 0.23 7.8 17.2 3.5 2.3 � 10�4 75.7 1629 596.7H3 1 80 0.08 2.2 14.8 9.4 8.8 � 10�4 127.2 1134 1032.9H4 5 0 0.30 11.2 18.6 – – 8.0 2558 310.0H5 5 30 0.23 7.8 17.2 3.5 3.2 � 10�4 15.1 1631 597.2H6 5 80 0.08 2.2 14.8 9.4 2.4 � 10�4 25.5 1132 1034.9H7 8 0 0.30 11.2 18.6 – – 5.0 2558 310.0H8 8 30 0.23 7.8 17.2 3.5 3.6 � 10�4 9.5 1629 598.8H9 8 80 0.08 2.2 14.8 9.4 9.5 � 10�4 16.0 1128 1040.3

a Pressure, fractional H2 preconversion, equivalence ratio, major species and radical (sum of OH, H, O, HO2, H2O2)% volumetric compositions (the balance is N2), velocity,Reynolds number, and temperature at the inlet of each honeycomb channel. In all cases the adiabatic equilibrium temperature is 1198 K and the mass flow rate in eachchannel is 0.046 g/s.

1100

1200

1300

1400

1500

1600

1700(a) 1 bar

Tad

Wal

l tem

pera

ture

(K)

1100

1200

1300

1400

1500

1600

1700

x (mm)15 30 45 60 75

1100

1200

1300

1400

1500

1600

1700

(b) 5 bar

(c) 8 bar

Tad

Tad0

Case H1(0% preconv.)

Case H2(30% preconv.)

Case H3(80% preconv.)

Case H4(0% preconv.)

Case H5(30% preconv.)

Case H6(80% preconv.)

Case H7(0% preconv.)

Case H8(30% preconv.)

Case H9(80% preconv.)

Fig. 12. Computed wall temperature profiles for the cases in Table 4. Dashed lines:only catalytic chemistry included; solid lines: both catalytic and gas-phasechemistry included.

0.0

0.4

0.8

1.2

1.6

Hyd

roge

n ga

s-ph

ase

conv

ersi

on ra

te (g

/m2 s)

0.0

2.0

4.0

6.0

8.0

x (mm)0 15 30 45 60 75

0.0

2.0

4.0

6.0

8.0

(a) 1 bar

(b) 5 bar

(c) 8 bar

Case H1 (0% preconv.)

Case H2 (30% preconv.)

Case H3

Case H 4 (0% preconv.)

Case H 5 (30% preconv.)

Case H 6 (80% preconv.)

Case H 7 (0% preconv.)

Case H 8 (30% preconv.)

Case H 9 (80% preconv.)

Fig. 13. Computed gas-phase hydrogen conversion rates (integrated over thechannel radius) for the cases in Table 4.

1954 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

4.4.2. Surface temperatures due to catalytic and gas-phase reactionsThe inclusion of gaseous chemistry lowers the wall tempera-

tures (Fig. 12). The underlying reason is that the presence of aflame moderates the reactor wall temperature by shielding the cat-alytic surfaces from the hydrogen-rich channel core and hencereducing the heterogeneous conversion which is responsible forthe superadiabatic temperatures [13,50]. Thus, contrary to manyCST premises, the onset of homogeneous ignition is actually desir-able in fuel-lean H2 catalytic combustion. Gas-phase conversionrates are presented for all cases of Table 4 in Fig. 13. For the1 bar cases in Fig. 13a, vigorous homogeneous combustion is initi-ated at x ’ 10 mm for 0% and 30% preconversions, such that thegaseous combustion zone shields the catalyst only downstreamof the peak wall temperature (see Fig. 12a). Therefore, the peak

temperatures for the 1 bar cases are nearly the same for both com-putations, with or without inclusion of gaseous chemistry.Fig. 14(a1) provides 2-D distributions of the OH radical for CaseH2 (30% preconversion at 1 bar) clearly illustrating flames stillconfined very close the walls, despite the higher confinement ofthe honeycomb channel (diameter 1.2 mm) compared to the7-mm channel height in Fig. 4. For 1 bar and 0% preconversion,the wall temperatures at x ’ 10 mm drop by as much as 50 K whengas-phase chemistry is included (Fig. 12a). For the 80% preconver-sion Case H3 in Fig. 12a, the wall temperatures are only modestlyaffected by the presence of gaseous chemistry.

At 5 and 8 bar, gaseous chemistry substantially lowers the peakwall temperatures for the 0% preconversion (Fig. 12(b and c)). Thereason is that homogeneous combustion is now initiated close toor upstream of the peak wall temperature in the purely catalyticsimulations (see Fig. 13(b and c) and the 2-D OH distribution inFig. 14(b1)), allowing for significant suppression of the superadia-batic surface temperatures. Contrary to the optically accessible

(a) Case H2, 1 bar

0 30 754515x (mm)

(a1) OH

(a2) H

(b1) OH

(b2) H

1.2 mm

(b) Case H5, 5 bar

60

2

2

min max

Fig. 14. Computed 2-D distributions of OH (1) and hydrogen (2) mole fraction, for:(a) Case H2 and (b) Case H5 in Table 4 for the channel geometry in Fig. 3b. Theminimum and maximum levels are: 0.0 and 9.53 � 10�4 for OH; 1.2 � 10�4 and0.078 for H2.

x (mm)15 30 45 60 75

Wal

l tem

pera

ture

(K)

1200

1300

1400

1500

1600

Tad

p = 5 bar30 % preconv.

d=1.2mmd=3.0mm

d=1.8mm

d=2.4mm

d=1.8mmd=3.0mm

d=2.4mmd=1.2mm

0

Fig. 15. Computed wall temperature profiles for 30% preconversion at 5 bar andfour different channel diameters. Dashed lines: only catalytic chemistry included;solid lines: both catalytic and gas-phase chemistry included. The 1.2 mm diameterresults refer to Case H5 in Fig. 12.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1955

reactor experiments in Fig. 4, the present simulations show thateven at 0% preconversion (non-preheated mixtures) the 5 and8 bar cases exhibit vigorous homogeneous combustion. This isdue to the particularly high wall temperatures (>1500 K over theinitial reactor length), leading to high effective temperatures inFig. 7 that in turn warrant a continuous increase of the gaseousreactivity with rising pressure from 1 to 8 bar. However, it shouldbe noted that reactor operation with 0% preconversion and no cool-ing is detrimental due to the high wall temperatures. Even thoughthe moderation in peak wall temperatures due to gas-phase reac-tions is considerable for 0% preconversion at 5 and 8 bar (by123 K and 102 K, respectively), the resulting peak wall tempera-tures are 1545 and 1569 K, respectively, still posing concerns forthe reactor and catalyst integrity. On the other hand, for a 30% pre-conversion the peak wall temperatures at 5 and 8 bar are 1494 and1478 K, respectively. Considering that hot spots of up to �1450 Kcan be tolerated by many ceramic reactors and/or catalysts for ex-tended times, an appropriate preconversion between 30% and 80%can altogether negate reactor thermal management issues.

Hydrogen preconversion (%)0 20 40 60 80

Rad

iatio

n flu

x at

inle

t (W

/cm

2 )

2

4

6

8

10

12

14

16 1 bar5 bar8 bar1 bar,LeH2 = 1

Fig. 16. Computed radiation flux at inlet as a function of hydrogen preconversion,for all cases in Table 4. Computations are also shown for the 1 bar cases when theLewis number of hydrogen is artificially set to unity.

4.4.3. Reactor thermal managementA strategy to further suppress the superadiabatic peak wall

temperatures is to increase the impact of gaseous chemistry byreducing the channel surface-to-volume ratio. Fig. 15 depicts com-puted wall temperatures (with and without inclusion of gaseouschemistry) at p = 5 bar and 30% preconversion for channel diame-ters ranging between 1.2 and 3.0 mm. The mass throughput wasmaintained constant when increasing the diameter by reducingthe inlet velocity. The enhancement of homogeneous combustionwhen increasing the diameter from 1.2 to 3.0 mm leads to a dropin the peak wall temperature by nearly 100 K. Computations inthe absence of gaseous reactions (dashed lines in Fig. 15) revealpeak wall temperatures with a non-monotonic dependence ondiameter. This is because radiation losses to the entry are favoredby increasing diameter but on the other side finite-rate surfacechemistry effects near the entry are more pronounced at smallerdiameters. Channels with larger hydraulic diameters are thus pref-erable for moderating the surface temperatures during fuel-leanH2/air catalytic combustion.

Since hydrogen preconversion moderates the wall tempera-tures, reactor designs utilizing the preconversion benefits are

desirable. Crucial for the particular CST combustion concept pre-sented in the following section is the radiation heat transfer fromthe reactor to the inlet section. The high wall temperatures atx � 0 (Fig. 12) favor large radiative heat transfer to the inlet. Thecombined radiative heat flux, originating from the 160 dicretizedsolid wall elements, towards the inlet is presented in Fig. 16 (sim-ulations using hetero-/homogeneous chemistry). The radiationfluxes qIN for the 5 and 8 bar cases are greater than 14 W/cm2 forpreconversions of up to 30%, while the corresponding values at1 bar are 2–3 W/cm2 lower. This is because the wall temperaturesat x � 0 are higher at elevated pressures. The radiation configura-tion factors Fx�IN, with x denoting the location of a discrete channelelement, are in turn strongly dependent on axial distance (seeFig. 17), leading to lower qIN at 1 bar.

Contrary to hydrogen catalytic combustion, the radiative feed-back during hydrocarbon combustion is much weaker. Inlet radia-tion heat fluxes are also presented in Fig. 16 for 1 bar, computed bysetting LeH2 ¼ 1 (mimicking methane). The resulting wall temper-ature for the 0% preconversion is provided in Fig. 17, clearly show-ing that the reduced qIN is a result of the lower-wall temperaturesat the entry. The temperature profile in Fig. 17 is reminiscent tothat of methane catalytic combustion [56], ramping from lowerto higher values with increasing axial distance.

x (mm)0 15 30 45 60 75

Con

figur

atio

n fa

ctor

Fx-

IN (-

)

1e-7

1e-6

1e-5

1e-4

1e-3

1e-2

1e-1

1e+0

Wal

l tem

pera

ture

(K)

900

950

1000

1050

1100

1150

1200

Fig. 17. Computed radiation configuration factors Fx�IN between the channel wallelements and the inlet for the geometry in Fig. 3b, and wall temperature for Case H1in Table 4 (1 bar, 0% preconversion) when the Lewis number of H2 is artificially setto unity. Time (s)

0 150 300 450 600

Tem

pera

ture

(K)

300

500

700

900

1100

AB

C

D

F

E

Tad

T2

1

2

B FC D EA FCC DD EE

T1 T3

Fig. 18. Measured temperatures for the porous burner and catalytic reactorconfiguration in Fig. 3a at 5 bar. Inlet volumetric composition: 6.1% H2, 9.1% O2,77.6% N2 and 7.2% CO2 and UIN = 2.3 m/s. The time intervals T1, T2, and T3 denotethe fuel-ramping to its final composition, quasisteady operation, and ramping of themixture preheat (thermocouple A), respectively.

1956 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

4.5. Homogeneous and heterogeneous combustion concept

In conventional fuel-lean CST, the catalytic burner converts partof the fuel while the remaining is combusted in a follow-up homoge-neous zone, which is stabilized by the hot exhaust gas products ofthe catalytic reactor. However, this arrangement can be detrimentalfor the reactor and catalyst as it always leads to superadiabatic sur-face temperatures (see Fig. 12, Cases H1, H4 and H7). The foregoingsections have clearly shown the benefits of H2 preconversion insubstantially moderating the catalytic reactor temperatures. Ifhydrogen preconversion is to be accomplished catalytically, theupstream catalytic reactor would also encounter superadiabatic sur-face temperatures. It is thus preferable to achieve the preconversionvia gaseous combustion.

A methodology of achieving preconversion by utilizing an in-verse CST combustion sequence is herein proposed. An upstreamhomogeneous combustion zone preconverts part (or nearly all) ofthe fuel, while a downstream catalytic module ensures completeconversion of the remaining fuel. Since gas-phase ignition of H2 re-quires sufficiently high temperatures (in excess of 673 K for pres-sures up to 8 bar [52]), an efficient mechanism of heat transferfrom the catalytic module to the gaseous reaction zone is necessaryto achieve ignition. In the proposed concept, gas-phase combustionis achieved in an inert porous burner, positioned in front of the cat-alytic reactor (Fig. 3a). Upstream heat transfer to the porous burnercan be achieved via heat conduction in the solid frame supportingthe honeycomb and porous burners and via surface radiation fromthe honeycomb reactor. Both heat transfer mechanisms are veryefficient due the nature of hydrogen catalytic combustion thatleads to high reactor entry temperatures (Fig. 12).

The start-up procedure entails the flow of H2/air mixturethrough the porous/catalytic burners, achieving initially light-offof the catalytic honeycomb reactor. Subsequently, the hot honey-comb transfers heat upstream to the porous structure and initiatesgaseous combustion within it. In this arrangement, the amount ofH2 preconversion in the porous burner cannot be controlled, lead-ing in many cases to nearly complete H2 conversion in the up-stream gaseous zone. To achieve gaseous ignition in the porousburner, some preheat of the incoming H2/air mixture may also benecessary. The start-up process for the porous/catalyst modulesis illustrated in Fig. 18, providing the time evolution of the mea-sured temperatures. The pressure is 5 bar, the volumetric inletcomposition (location A) 6.1% H2, 9.1% O2, 77.6% N2 and 7.2% CO2

(u = 0.33) while the inlet velocity (referring to an inlet temperature360 K at A) is 2.30 m/s. The zone marked T1 in Fig. 18 denotes the

ramping of hydrogen to its final content (6.1 vol.%) while maintain-ing the inlet temperature (thermocouple A) at 360 K; zone T2 des-ignates quasisteady operation with the established final mixturecomposition and inlet temperature still at 360 K; finally T3 indi-cates the zone of inlet temperature (thermocouple A) ramping, un-til gaseous ignition is achieved in the porous burner. The adiabaticequilibrium temperatures based on the inlet temperatures (A) arealso shown in Fig. 18. Strong gas-phase combustion is achieved in-side the porous structure for inlet temperatures above �430 K (seetime record of thermocouple A), as indicated by the porous burnerexit temperature (thermocouple B). When the honeycomb reactoris replaced by an identical inert (non-catalytic) structure, no gas-phase ignition could be achieved in the porous burner even for in-let temperatures (A) as high as 720 K.

Simulations for a representative channel of the honeycombstructure have been performed at the time instances marked (1)and (2) in Fig. 18. For these two times, Fig. 19 provides the pre-dicted quasisteady wall temperatures and mean-gas temperatures,the adiabatic equilibrium temperatures, and the thermocouplemeasurements at C, D and E. Condition (2) has been computed byconsidering a 71% gas-phase conversion in the porous burner (cal-culated from the �340 K temperature rise between thermocouplesA and B). For condition (1) it can be argued that the modest tem-perature rise between thermocouples A and B (�60 K) is not dueto H2 conversion but rather due to heat recirculation via conduc-tion in the solid walls of the metal holder tube wherein the porousburner and honeycomb reactor are mounted. Thus, for condition(1) a 0% preconversion has been considered, while the adiabaticequilibrium temperature Tad,1 shown in Fig. 19 has been calculatedbased on the catalytic reactor inlet temperature (position B). Forcondition (1), the predicted wall temperature at the location ofthermocouple C is substantially higher than the correspondingmeasured value. This is due to the large difference between the lo-cal mean-gas and wall temperatures (Tgas,1 and Twall,1 in Fig. 19).For condition (2), the upstream hydrogen gas-phase preconversionleads to mean-gas and wall temperatures closer to each other(Tgas,2 and Twall,2 in Fig. 19), rendering the thermocouple measure-ment at C more reliable. It is nonetheless clear that the increasedH2 preconversion for condition (2) leads to lower predicted walltemperatures at the front honeycomb section x < 4 mm.

x (mm)0 4 8 12 30 45 60 75

Tem

pera

ture

(K)

600

700

800

900

1000

1100

Tgas,1

Tgas,2Twall,1

Twall,2 Tad,2

Tad,1

Tad,1

EDC

Tad,2

Fig. 19. Computed wall (Twall) and mean-gas and (Tgas) axial temperature profilesfor a catalytic channel (Fig. 3b) at the time instances marked (1) and (2) in Fig. 18.Measured thermocouple temperatures at positions C, D and E are indicated bysymbols (circles: time instance (1), squares: time instance (2)). For clarity, the first12 mm are expanded.

Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958 1957

The burner concept in Fig. 3 has been demonstrated for inlettemperatures (A) above 400 K. Such preheat temperatures are eas-ily attainable in many power generation systems. In any case, therequired preheat depends on the specific geometrical configurationand positioning of the porous burner and honeycomb reactors, thereactivity of incoming mixture, the mass flow rate, solid heat con-ductivity, etc. Tests at 8 bar were carried out with practically room-temperature inlet (A at 308 K) by lowering the nitrogen flow suchthat the total mass throughput was eventually reduced down to58% of its initial value at time interval T1 (see Fig. 20). Temperaturemeasurements at positions A, B and F clearly indicate that gas-phase ignition is possible without preheat at the inlet. It is finallynoted that the concept in Fig. 3 is of interest not only to hydrogenbut also to hydrogen-rich fuels (e.g. syngas).

5. Conclusions

The impact of fractional hydrogen preconversion on the ensu-ing homogeneous ignition characteristics of fuel-lean (equiva-lence ratio u = 0.3) H2/O2/N2/CO2 mixtures over Pt has been

Time (s)0 200 400 600 800 1000

Tem

pera

ture

(K)

300

500

700

900

1100

Tad,1

Tad,2

A

BF

Tad,3

T1 T2 T3 T4

Tad,4

Fig. 20. Measured time evolution of temperature at positions A (inlet), B (porousburner outlet) and E (catalytic reactor outlet) for the configuration in Fig. 3a at8 bar. The gas composition in zone T1 is the same as in Fig. 18 and UIN = 1.4 m/s. Inthe intervals T2, T3 and T4 the N2 flow rate has been lowered such that the inletmass throughput was reduced to 61%, 60% and 58%, respectively, of the initial massflow rate at T1. The adiabatic equilibrium temperatures, based on the correspond-ing compositions and the inlet temperature (thermocouple A), are also provided.

investigated experimentally and numerically using a 2-D ellipticcode at pressures of 1, 5 and 8 bar. Experiments were performedin an optically accessible catalytic reactor and involved theassessment of catalytic fuel conversion and homogeneous igni-tion with Raman and OH-LIF, respectively. Computations in atechnical honeycomb catalytic reactor determined the impactof preconversion on the reactor temperature. The following arethe key conclusions of this study.

(1) The employed hetero-homogeneous chemical reactionschemes reproduced the measured transport-limited cata-lytic H2 conversion and the onset of homogeneous ignitionat preconversions of 0–80% and moderate pressures up to8 bar. These preconversions entailed inlet temperatures ashigh as 810 K and equivalence ratios as low as 0.078.

(2) For 0% hydrogen preconversion and catalytic surface tem-peratures in the range 900 K 6 Tw 6 1100 K, homogeneousignition was largely suppressed for p P 5 bar due to com-bined effects of intrinsic gas-phase hydrogen kinetics andcompetition between the catalytic and gas-phase pathwayfor fuel consumption. However, an increase in preconversionto 30% (and the resulting higher inlet temperatures) sharplyincreased the gaseous reactivity and restored homogeneouscombustion for p = 5 and 8 bar.

(3) For all preconversions, gaseous combustion was weaker at 5and 8 bar compared to 1 bar. This led to hydrogen leakagethrough the near-wall gaseous reaction zone and subse-quently to heterogeneous conversion of the escaping fuelon the catalytic wall, hence resulting in combined hetero-/homogeneous combustion even well-downstream the onsetof homogeneous ignition.

(4) The chemical impact of water formation due to hydrogenpreconversion (inlet water) was appreciable for preconver-sions of 30% and 80%, resulting in inhibition of homogeneousignition. This effect was more pronounced at higher pres-sures, due to acceleration of the chain terminating stepH + O2 + M = HO2 + M in which H2O was an efficient thirdbody. Moreover, the inhibition of homogeneous ignitiondue to catalytically-produced water was proportionallymuch stronger than that of the inlet water, particularly athigher pressures. This was due to the diffusional imbalanceof hydrogen, which led to H2O levels in the near-wall hotignitable zone nearly twice as high as the ones expectedby stoichiometric considerations alone.

(5) For inlet temperatures up to 810 K, the inlet radical com-position due to hydrogen preconversion had practically noimpact on the ensuing homogeneous ignition at pressuresof 5 and 8 bar. Only at p = 1 bar and 80% preconversion,the inlet radicals could initiate a modest gas-phase H2

consumption near the catalytic reactor inlet. However, thishad no noticeable effect on the main homogeneous igni-tion event occurring farther downstream near the catalyticwall. The hetero-/homogeneous chemical radical couplingvia adsorption/desorption reactions (notably OH) inhibitedthe onset of homogeneous ignition, with the inhibitionbeing stronger at p = 1 bar and modest at 5 and 8 bar.On the other hand, over the post-ignition reactor length,radical adsorption/desorption reactions significantly sup-pressed gas-phase combustion at 5 and 8 bar while theirimpact at 1 bar was weaker.

(6) For 0% hydrogen preconversion, catalytic channels exhibitedsuperadiabatic surface temperatures at the far upstreampositions that endangered the reactor and catalyst integrity.By increasing preconversion, the catalytically-induced sup-eradiabaticity was progressively suppressed. Moreover, thepresence of strong gaseous combustion at all pressures and

1958 Y. Ghermay et al. / Combustion and Flame 157 (2010) 1942–1958

intermediate preconversions (30%) offered an additionalmechanism for surface temperature moderation by shield-ing the catalyst from the hydrogen-rich channel core.

(7) The diffusional imbalance of hydrogen led to higher surfacetemperatures at the reactor entry, in contrast to hydrocar-bon combustion whereby highest temperatures are typicallyencountered at the reactor rear. This characteristic of hydro-gen catalytic combustion allowed for efficient transfer ofheat upstream of the catalytic reactor via surface radiationand solid heat conduction.

(8) A combustion concept has been investigated that tookadvantage of the surface temperature moderation withincreasing hydrogen preconversion. An inverse catalyticallystabilized thermal combustion concept was presented,wherein the homogeneous combustion zone preceded thecatalytic reactor. It was shown that this arrangementallowed for lower catalyst surface temperatures and thatgas-phase ignition in the porous burner could eventuallybe achieved without preheat of the incoming mixture.

Acknowledgments

Support was provided by the Swiss Commission for Energy andInnovation (KTI), Swiss Competence Center of Energy and Mobility(CCEM), and Alstom Power of Switzerland. We thank Mr. Rolf Scha-eren and Mr. Rene Kaufmann for the help in the experiments.

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