Paper # 070RK-0088 Topic: Reaction Kinetics

8th

U. S. National Combustion Meeting

Organized by the Western States Section of the Combustion Institute

and hosted by the University of Utah

May 19-22, 2013

High-pressure low-temperature ignition behavior of syngas

mixtures

A.B.Mansfield1, M.S.Wooldridge

1,2

1Department of Mechanical Engineering, University of Michigan, Ann Arbor, MI 48109, USA 2Department of Aerospace Engineering, University of Michigan, Ann Arbor, MI 48109, USA

The use of coal-derived syngas has the potential to reduce pollutant emissions associated with electricity

generation (US DOE 2013b). However, the implementation of realistic syngas fuel in lean pre-mixed combustion

strategies is currently limited by an incomplete understanding of the oxidation kinetics and ignition physics at low

temperatures (<1000 K) and high pressures (>10 atm) (Chaos & Dryer 2008). In the present study, the ignition

behavior of simulated syngas mixtures is systematically investigated near these conditions using the University of

Michigan Rapid Compression Facility. Pressure-time history measurements and high-speed imaging of the

ignition process in this facility are used to determine the auto-ignition delay time and observe ignition behavior.

The simulated syngas mixtures are composed of various amounts hydrogen, carbon monoxide, oxygen, nitrogen,

carbon dioxide, and argon. The experiments are conducted at 3 and 15 atm, for temperatures ranging from ~850 –

1200K, φ = 0.1, and dilution of 75%.

The results demonstrate that for experiments with strong ignition behavior the Li et al. (2007) chemical

mechanism applied in a zero-dimensional homogeneous reactor simulation can accurately predict the measured

auto-ignition delay times. The uncertainties in the key reactions in the detailed mechanism were quantified in this

study and shown to be significant at the conditions of interest to gas turbine combustors. Furthermore, a

comparison of simulation methods indicates that heat transfer effects on auto-ignition delay times are negligible at

these conditions ( . For experiments with longer test times this will likely not be the case and

appropriate criteria should be applied to both model and report the experimental data. An evaluation of the

simulation methods indicated that defining effective thermodynamic conditions is likely the most useful, allowing

for the inclusion of first-order heat transfer effects while retaining ease and clarity in reporting the results. In

addition, a close relationship between transitions in ignition behavior and transitions across the classical and

extended 2nd limits on the H2/O2 explosion map was demonstrated using a pressure/temperature map of ignition

behavior. This behavior seems to be largely unaffected by reactant mixture composition, though the existence of

weak ignition behavior may be linked to equivalence ratio.

1. Introduction

Synthesized gas, or syngas, is a mixture composed primarily of H2 and CO that can be used as a chemical pre-cursor in

manufacturing or combusted directly as a fuel. Syngas can be produced via gasification of various carbonaceous sources

such as biomass, municipal solid waste, landfill gas, and most notably coal. As both emissions regulations and resource

scarcity increase there is a great desire to develop “clean-coal” technology, and by changing the combustion strategy of

coal this may be achieved. The proposal of an Integrated Gasification Combined Cycle (IGCC) plant is particularly

promising, whereby a coal gasification process is run in concert with a combined cycle gas turbine system to generate

power. Compared to a pulverized coal power system, an IGCC plant can achieve reductions in emissions of SOx, NOx,

and particulate matter without a significant reduction in plant efficiency. Furthermore, removal of hazardous coal

impurities such as mercury is more easily achieved in this highly controlled process. (US DOE 2013b)

Currently the gas turbine portion of the IGCC system is still in the research and development phase. High hydrogen

content fuel like syngas adds complexity to combustion systems given its unique and extreme behaviors, e.g. high flame

speeds, high diffusivity, wide flammability limits, and increased flame temperatures (US DOE 2013a). With higher

flame temperatures, a NOx control strategy must be enacted which typically involves an increase in dilution by either air

or steam (Lieuwen et al. 2008). For syngas fuel however, large amounts of dilution alone likely cannot reduce the NOx

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emissions levels sufficiently. Therefore, an increasingly common combustion strategy for high hydrogen content fuels

like syngas is to operate in a lean pre-mixed mode (Richards et al. 2001). While pre-mixed combustors do exist for

natural gas, these cannot be operated well with syngas fuel. Given the physical characteristics listed above, issues like

blowout, flashback, auto-ignition, and flame instability are a major concern (Lieuwen et al. 2008). Further adding to the

complexity of pre-mixed syngas combustor design and operation is the highly variable makeup of coal-based syngas

fuels – both in terms of constituents and concentrations, as indicated in Table 1.

Table 1. Coal-based Syngas Composition (United States Department of Energy 2009; Cayan et al. 2008)

Component % by

Volume

H2 25-30

CO 30-60

CO2 5-15

H2O 2-30

CH4 0-5

N2 0-4

Ar, N2, H2S, COS, NH3, Ash 0-1

Trace Impurities

(Fe(CO)5, HCl, Si(OH)4, Metals, etc.)

< 10ppm

In order to develop gas turbine combustors that can fire syngas in a pre-mixed combustion mode, it is therefore necessary

to understand both the combustion physics and chemistry of this fuel at relevant gas turbine conditions and across a wide

range of possible fuel mixtures. The chemistry is of particular importance in this system, as pre-mixed combustion

physics are directly linked to the chemical kinetics of the fuel oxidation process.

A review of the literature for the baseline chemical kinetics of syngas combustion reveals that while they have been

reasonably well established, their overall applicability in real world combustors may be limited. This is because the

kinetics have been developed almost exclusively for fuels containing only H2 and CO, which as evidenced in Table 1

above, is an unrealistically simple composition. Furthermore, gas turbine pre-mixer conditions (T ~600-900 K, P = ~5-

35 atm) lay almost exclusively at high-pressure, low-temperature conditions, the region of least certainty for the

established chemical kinetics.

There exists a wide body of experimental and modeling work done to investigate the chemical kinetics of syngas

combustion. This knowledge will be only briefly mentioned here, interested readers are directed to a comprehensive

presentation and discussion of syngas kinetics by Chaos & Dryer (2008). Overall the kinetics are well understood at

high-temperature low-pressure conditions where most previous experiments have been conducted. There is evidence,

however, that they are less accurate at high-pressure low-temperature conditions, specifically between the extended

second and third explosion limits in the H2/O2 explosion map (Lieuwen et al. 2009). This inaccuracy is currently

attributed to a shift in ignition behavior that is not captured in typical modeling and not a failure of the kinetic

mechanism. Yet the results highlight important ignition behavior that could both impact the safe operation of real

combustor systems (Chaos & Dryer 2008) and limit the effectiveness of traditional experimental validation of chemical

mechanisms. This shift in ignition behavior is observed as a transition between “strong” ignition, characterized by

uniform ignition occurring simultaneously throughout the test volume, and “weak” ignition, characterized by the

existence of local ignition sites and flame propagation. These phenomena are seen across a range of experimental

facilities (Chaos & Dryer 2008) and hypotheses as to the cause and effect of the shift between the ignition regimes

remains largely speculative. The existence of uncontrolled changes in ignition behavior poses a clear safety risk in a gas

turbine pre-mixer and will likely require new experimental treatment to ensure the appropriate evaluation of chemical

kinetics in existing facilities. In addition to these behaviors, the effect of CO concentration on the auto-ignition delay

times at high-pressure, low-temperature conditions is not entirely understood, with several studies reporting a weak

correlation (S. M. Walton et al. 2007; Sander Gersen et al. 2012) and one reporting a very strong correlation between the

two (Mittal et al. 2006).

There have been several studies aimed at characterizing the effects of variations in syngas composition and trace

impurities. Mathieu et al. (2012) recently conducted ignition experiments using lean biomass derived syngas mixtures in

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a shocktube and found that while CH4 addition can increase the auto-ignition delay time by OH radical scavenging

through the reaction (CH4 + OH = CH3 + H2O), the addition of H2O and CO2 had little effect. Mathieu et al. (2012) also

found that the addition of ppm levels of NH3 did not affect the auto-ignition delay time due to the lack of reactivity of

NH3 with H2/O2 chemistry. Work by Rasmussen et al. (2008) and Mueller & RA Yetter (1999) in flow reactors

highlighted the potential for NOx to influence syngas ignition chemistry. Mueller et al. found specifically that ppm

levels of NO promoted generation of OH radicals. Singh et al. (2012) found that water addition increased flame speeds

for mixtures with very low ratios of H2:CO; whereas, Das et al. (2011) found that water addition accelerated auto-

ignition at high pressures and decelerated auto-ignition at lower pressures, largely the result of the increased third body

collision efficiency of the water. (Cong & Dagaut 2008) used a jet stirred reactor to study the effect of CO2 addition on

syngas combustion chemistry and found that CO2 inhibits oxidation of CO largely by slowing the CO + OH = CO2 + H

reaction. Rumminger & Linteris (2000) discovered that the presence of Fe(CO)5, a trace impurity in some syngas

mixtures – see Table 1, can reduce the flame speed of syngas by up to 30% through H and O radical scavenging.

Upon review of the previous work, it is apparent that there is a need to continue the investigation of syngas combustion

chemistry at thermodynamic conditions relevant to gas turbine combustors, with a focus on the effects of mixture

variations and trace impurities. Furthermore, it is important that the ignition behavior of syngas mixtures be

systematically characterized across the explosion limit map for a wide range of mixture variations, with a specific focus

on behaviors during the transition across both the extended second explosion limit and the classical second explosion

limit.

It is the overall aim of the present work to quantify auto-ignition behavior of realistic syngas mixtures at conditions

relevant to gas turbine combustors. This is generally accomplished by a series of ignition experiments conducted in the

University of Michigan Rapid Compression Facility (UM RCF), with which it is possible to measure auto-ignition delay

times, directly evaluate syngas kinetic models, and observe ignition behavior as a function of thermodynamic state and

mixture characteristics.

Observations of the ignition process occur from several vantage points in UM RCF experiments. Transient pressure

measurements of the test chamber are used to determine an auto-ignition delay time and characterize thermal energy

changes in the system. Line of sight UV laser absorption measurements are used to determine the time-resolved

concentration of OH in the ignition test chamber. Further, high-speed imaging of the chemilluminescence is used to

characterize the general ignition behavior. These ignition experiments are generally conducted between 5 and 30 atm,

across the widest range of allowable temperatures in the UM RCF (typically 900-1150 K), for syngas mixtures with an

equivalence ratio between 0.1 and 0.5. After firmly establishing data for simple mixtures of H2 and CO, additional

components such as CH4, CO2, H2O, and trace impurities are systematically added.

The goals of this paper are to:

(1) Examine relevant literature and describe the trajectory of this work in that context

(2) Present new experimental findings for mixtures of H2 and CO

(3) Evaluate and recommend methodologies for reporting experimental results and conducting kinetic model

simulations

(4) Map ignition behavior for simple H2/CO mixtures, using new and existing data

Ignition experiments have been completed for lean mixtures of H2 and CO (φ = 0.1) at 3 and 15 atm from ~900-1150 K

using the UM RCF. The exact mixture compositions and the corresponding results for each experiment are provided in

Table A of the Supplemental Material. The auto-ignition delay times and ignition behavior for the experiments are

reported here along with the general experimental and simulation methodologies.

2. Methods

2.1 Experimental

The UM RCF is a unique experimental apparatus for creating uniform high temperature and high pressure conditions,

through an isentropic compression process (M. T. Donovan et al. 2004). A detailed description of the UM RCF and

results of studies characterizing its performance can be found in M. T. Donovan et al. (2004); He et al. (2006). Briefly,

the UM RCF consists of a long cylinder, the Driven Section, in which a gas mixture is rapidly compressed by the motion

of a free piston (Sabot). Prior to compression, the test volume is evacuated with a pump and then filled with a specific

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test gas mixture. Upon firing, the Sabot travels the length of the Driven Section compressing the test gas mixture into

the Test Section – a small cylindrical volume located at the end of the Driven Section. As the Sabot reaches its final

position near the Test Section, it achieves an annular interference fit, thereby sealing the test gas mixture in the Test

Section. At this point, the Test Section is filled with a uniform and isentropically compressed test gas mixture at the

desired high-pressure, high-temperature condition. This is achieved in large part because cool boundary layer gases from

the Driven section are trapped in an external volume formed by the geometry of the Sabot (M. Donovan et al. 2004; S.

Walton et al. 2007).

For this study, the test section was instrumented with a piezoelectric transducer (6125B Kistler, Amherst, NY) and

charge amplifier (5010, Kistler, Amherst, NY) for pressure measurements, and a transparent polycarbonate end-wall to

permit high-speed imaging of the ignition process. High-speed color imaging was taken using a digital video camera

(V711-8G-MAG-C, Vision Research, Phantom) with a Navitar 50mm lens (F0.95), a Hoya 62mm lens (+2 zoom), and a

Hoya 62mm UV(0) filter. Video sequences were recorded at 25,000 frames/second with a CCD resolution of 512 x 512

pixels. These settings result in an exposure time of 39.3 μs. During each experiment the pressure-time history is

recorded using the pressure transducer at 100 kHz sampling frequency and the chemilluminescence is recorded by the

high-speed camera.

All test gas mixtures were made using a dedicated stainless steel tank and the mixture composition was determined by

measurement of relative partial pressures of the components. After filling, the tank is left closed for approximately one

hour before the test gas is used for an experiment, during which it is assumed that the mixture homogenizes. The

mixture compositions for each experiment can be seen in Table A in the Supplemental Material. Error in the mixture

compositions is assumed to be negligible and have negligible effect on the ignition results.

As mentioned above, for these experiments the end-of-compression pressure (PEOC) was toggled between 3 and 15 atm

while the EOC temperature (TEOC) was varied from ~900-1150 K. These conditions were achieved controlling both the

initial pressure in the UM RCF Driven Section and the composition of the diluent mixture (N2, Ar, and CO2). A typical

pressure time history for this series of experiments is shown in Figure 1, illustrating the compression process, induction

period, and ignition. There is a noticeable decline in the pressure during the induction process, which can be attributed

to heat loss from the UM RCF test chamber. Note here that the pressure time history shown in Figure 1 has been

averaged with a 150-point moving average smoothing algorithm for illustrative simplicity, though the unfiltered raw

pressure time history data was used for all quantitative calculations.

Figure 1. Typical pressure time history for strong ignition event at the experimental conditions TEOC =1073 K, PEOC

=13.37 atm, φ = 0.1, H2:CO = 0.7, and Dilution = 75%.

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From each pressure time history, PEOC is determined by manually identifying the time of the maximum pressure (tEOC)

during the compression process and placing appropriate time error bounds on either side of this value to incorporate

uncertainty in this selection. PEOC is defined as the average of all pressure values within the time error bounds with

pressure uncertainty equal to the standard error of the mean of the pressure values. The time of the end-of-compression

(tEOC) value has uncertainty equal to the standard deviation of the time values between the time error bounds selected.

Assuming an isentropic compression process: knowledge of the mixture composition, initial pressure and temperature,

and final pressure (PEOC) then allows for the calculation of TEOC through an iterative process described in M. T. Donovan

et al. (2004). The uncertainty of PEOC is propagated through these calculations to assign uncertainty values to TEOC. The

NASA thermodynamic database was used to supply the necessary thermodynamic data used in all calculations (McBride

et al. 2002).

The auto-ignition delay time (τign) is defined here as the time from end-of-compression (tEOC) to the average time of the

ignition event (tign-avg). The average time of the ignition event is determined by manually bounding the time of maximum

rate of pressure rise, indicated in Figure 1, and taking the average time in this range. τign is then calculated by taking the

difference between the tEOC and the tign-avg, with the uncertainty of each value propagating through the calculation to

define the error bounds for τign. τign was only calculated for experiments that exhibited strong ignition, given the

difficulty in separating chemical kinetic effects from physical effects during weak ignition events.

Still images from two typical high-speed imaging videos are shown in Figure 2a and b, exemplifying the

chemilluminescence that emanates from the combusting syngas mixture during ignition (shown in grey scale, but

typically a blue color). In both images, the presence of an “adiabatic core” and a cooler boundary layer region can be

seen. The small discolorations or bright spots seen in both images are commonly found and are attributed to debris from

the sealing-ring material from the Sabot. During weak ignition it is typical for flames to initiate and emanate from the

bright spots.

(a) Strong: Uniform ignition, no flame propagation (b) Weak: Multiple ignition sites, flame propagations

Figure 2. Typical high-speed imaging results for (a) “strong” ignition behavior, for the experimental conditions TEOC

=1050 K, PEOC = 2.87atm, φ = 0.1, H2:CO = 0.7, Dilution = 75% (b) “weak” ignition behavior for the experimental

conditions TEOC =1019 K, PEOC =2.85 atm, φ = 0.1, H2:CO = 0.7, Dilution = 75%. White arrows point to flame fronts.

For each experiment, the high-speed video of the chemilluminescence during ignition was reviewed. The ignition was

classified as “strong” if there was a clear uniform ignition that occurred over nearly the entire volume at the same instant.

The ignition was classified as “weak” if there were local ignition sites and flame propagation, and there was no uniform

ignition event. The ignition was classified as “weak transitioning to strong” if there were local ignition sites and flame

propagation, followed by a strong ignition event. The ignition was classified as “indeterminate” if the images were too

dim to identify the ignition characteristics.

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2.2 Simulation

The auto-ignition delay time measurements were compared with numerical simulations conducted using CHEMKIN

software (Reaction Design 2010) and the Li et al. (2007) chemical mechanism for H2 and CO fuels mixtures. Within the

software, the zero-dimensional homogeneous reactor model with adiabatic boundary conditions was used with the

constant volume form of the conservation of energy equation. The mixture composition, along with an initial

temperature and pressure value were input to the simulation and a corresponding pressure time history and auto-ignition

delay time was output. The auto-ignition delay time was defined in the simulation as the time at which the time

derivative of the temperature reached its maximum. All simulations were begun from an already-compressed

thermodynamic state at the initial time i.e. the compression stroke was not simulated.

An important consideration in these simulations is the impact of uncertainty in the Li et al. (2007) chemical mechanism.

After running an “A-factor sensitivity” analysis, a standard option in CHEMKIN, it was determined that the reaction for

which variation in the A-factor of the kinetic rate coefficient had the largest impact on τign, was Reaction 1 (R1): H + O2

= OH + O. Given that this was the dominant source of error in the chemical mechanism, the effect of error from all other

chemical mechanism parameters was assumed to be negligible. The error bounds of the A-factor for (R1) were

experimentally determined in Hessler (1998) and these values were used to generate three A-factor values that were used

in the simulations (maximum, minimum, and nominal). In this way, for a given mixture composition and initial

thermodynamic state, three simulated auto-ignition delay times were calculated, a maximum, minimum, and nominal

value.

In an ideal experimental combustion system there would be no heat transfer losses to the test chamber and the chamber

would behave exactly like a zero-dimensional constant volume adiabatic system as modeled in the CHEMKIN

simulation. However, as seen in the typical pressure time history in Figure 1, and in all similar combustion facilities

(Lee & Hochgreb 1998; Das et al. 2012; S Gersen et al. 2008), heat transfer from the test volume to the surrounding

environment results in a continuous decrease in the pressure after it reaches a maximum value at the end-of-compression.

While it has been established that an “adiabatic core” exists within the reaction chamber volume (M. Donovan et al.

2004; Lee & Hochgreb 1998), this adiabatic core experiences an expansion as surrounding boundary layer gases cool,

resulting in a pressure decrease (Lee & Hochgreb 1998). As this pressure decrease results in a corresponding

temperature decrease, the expansion of the adiabatic core can impact the chemical kinetics and thereby affect the auto-

ignition delay time. However, the magnitude of the impact will vary depending on the rate of pressure decrease and

properties of the reacting mixture (ratio of specific heats, initial thermodynamic state, and auto-ignition delay time).

In addressing the well-known issue of heat transfer losses in RCF experiments, three simulation methods have been used

in previous work, each with a different treatment of the situation (Mathieu et al. 2012; S Gersen et al. 2008; S. M.

Walton et al. 2007). In the following section each method and the details of its application to this work are discussed. A

comparison of simulated pressure time histories is plotted against the corresponding experimental pressure time history

in Figure 3.

Method 1: Constant volume at end-of-compression conditions

In this, the simulations are conducted using adiabatic constant volume boundary conditions with the initial

thermodynamic values set as the end-of-compression conditions (PEOC, TEOC). It is assumed in using this model that

there are no significant effects from the experimental pressure decrease on the auto-ignition delay time, lending to its

common use in simulating shocktube experiments (Kalitan et al. 2007; Mathieu et al. 2012). The major advantage of this

method is its simplicity and the major disadvantage is its inaccuracy in simulating an experiment with heat loss

significant enough to effect the auto-ignition delay time. From a reporting perspective, if the ignition behavior of an

experiment is adequately described using this method, then it is appropriate to report the measured and simulated auto-

ignition delay times directly on a classic isobaric, isothermal plot at the EOC conditions. An example of such a plot can

be seen in Figure 5 and in Eric L. Petersen et al. (2007), where the auto-ignition delay time is plotted as a function of

inverse temperature at a constant pressure. It is then valid to directly compare any data that is appropriately placed on

such a plot, even if it was generated in facilities with differing heat transfer characteristics. This method was applied to

every experiment and a characteristic simulated pressure time history created using this method can be seen in Figure 3.

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Method 2: Specific volume trace at end-of-compression conditions

In this, the simulations are conducted using an adiabatic boundary condition with the initial thermodynamic values set as

the end-of-compression conditions (PEOC, TEOC). A “specific volume trace” is applied in this model, which forces the

simulated test mixture to undergo the same pressure decrease as was measured experimentally (Sander Gersen et al.

2012; Mittal et al. 2006; Tanaka et al. 2003; Lee & Hochgreb 1998). The specific volume trace is often generated from

an “inert” pressure time history which is meant to represent the pressure decrease that would be observed experimentally

had the mixture not ignited, essentially capturing heat transfer effects only. This “inert” pressure time history can be

measured directly by conducting the same experiment without oxidizer (Mittal et al. 2006; Tanaka et al. 2003)

presuming pyrolysis or other chemical reactions are negligible at the experimental conditions, or it can be simulated

using a mathematical model calibrated to experimental data (Sander Gersen et al. 2012). It is assumed in using this

model that there are significant effects from the experimental pressure decrease on the auto-ignition delay time. Also, it

must be assumed that the pressure decrease is purely a function of heat transfer and is not convolved with endo- or

exothermic chemical reactions (Lee & Hochgreb 1998). The major advantage of this method is that it can simulate

essentially all heat transfer effects in the experiment and the major disadvantages are its complexity and the requirement

to conduct a series of “inert” experiments. From a reporting perspective, if the ignition behavior of an experiment can be

described only by this method, then it is not appropriate to report the measured and simulated auto-ignition delay times

on a classic isobaric, isothermal plot like the one mentioned above. The reporting method must therefore take some

other form, possibly including sets of experimental and simulated pressure time histories. Therefore, while this method

allows for a direct evaluation of the chemical mechanism, it may not be appropriate to directly compare results from

different experimental facilities with differing heat transfer characteristics.

In applying this method to the present work a mathematical model for the pressure during the induction period was

selected and evaluated using non-igniting experiments, seen in Table A. The modeled pressure time history was then

translated to a specific volume time history using isentropic expansion relations, assuming a constant ratio of specific

heats calculated at TEOC. The mathematical model found to have good agreement with all non-igniting pressure time

histories is a combination of exponential functions as follows,

( ( (

) ( (

) (

where,

(

)

and P1, τ1 τ2 are fitting parameters that can be adjusted to match the modeled pressure time history to the experimental.

This model was developed based on the general shape of the pressure time history measurements observed in non-

igniting experiments. It is a summation of two exponential functions, each designed such that at the initial time the

pressure equals the end-of-compression value and at very long times the pressure equals the approximate pressure of the

reaction chamber when it achieves room temperature ( . The two exponential functions were necessary to describe

the rapid pressure decrease that occurs immediately after end-of-compression and the more shallow pressure decrease

that occurs later. Typical values for each of the fitting parameters were: . For each simulation several points were selected along the measured pressure time history and the fitting

parameters were adjusted until sufficient agreement existed between the measured and modeled pressure time history at

these points. This method was only applied to experiments that exhibited some measurable pressure decrease during the

induction period. A characteristic simulated pressure time history created using this method can be seen in Figure 3.

Method 3: Constant volume at effective thermodynamic conditions

In this, the simulations are conducted using adiabatic constant volume boundary conditions with the initial

thermodynamic values set at “effective” conditions (Peff, Teff) (S. Walton et al. 2007; S. M. Walton et al. 2007; He et al.

2005). These conditions are generally calculated by first defining a time-averaged pressure, Peff, and then deriving Teff

using the compression stroke physics, as if Peff were achieved by compression from initial uncompressed conditions.

The bounds over which Peff is calculated can vary, but typically they are chosen as the time at end-of-compression and

the time at which the pressure is minimized before the ignition event (He et al. 2005). The goal of this method is to

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simulate the first-order effects of the experimental pressure decrease while retaining the simplicity of constant volume

adiabatic conditions. It is assumed in using this model that there are effects from the experimental pressure decrease on

the auto-ignition delay time and that these effects can be basically accounted for in a kinetic simulation by a simple

decrease in the initial pressure and temperature. The major advantage of this method is its ability to capture first-order

effects of the observed pressure decrease on the auto-ignition delay time, while remaining a relatively simple process.

The major disadvantage with this method is the necessary assumption regarding the effect of pressure and temperature

changes on the chemical reaction rates. From a reporting perspective, it is appropriate to report the measured and

simulated auto-ignition delay times directly on a classic isobaric, isothermal plot at the “effective” conditions, if the

assumption stated above regarding the effect of pressure decrease on the chemical reaction rates is validated. As with

Method 1, these data may then be compared directly to other data on such a plot, even if they were generated in facilities

with differing heat transfer characteristics.

In applying this method to the present work, a time averaged pressure over the induction period was calculated between

the time at end-of-compression and the time at which the pressure is minimized before the ignition event. The time of

minimum pressure was determined manually using the raw pressure time history data. The effective pressure values

calculated can be seen in Table A, though only for experiments which exhibited some measurable pressure decrease

during the induction period. It was found that typically

and

A characteristic simulated

pressure time history created using this method can be seen in Figure 3.

Figure 3. Comparison of experimental and simulated pressure time histories for a strong ignition event at the

experimental conditions TEOC =1073 K, PEOC =13.37 atm, φ = 0.1, H2:CO = 0.7, Dilution = 75%. Time = 0

corresponds to the end-of-compression. Each simulated pressure time history pictured here used the nominal value of

the A-factor for reaction (R1). Descriptions of each method are provided in the text.

Figure 3 presents a comparison of the experimental data for a strong ignition condition with the three modeling methods

described above. The simulated pressure time histories exhibit notable differences during the induction period; however,

the ignition delay times (as defined using the maximum rate of pressure rise for each simulation) do not differ

significantly. Moreover, if the uncertainty in the reaction mechanism is considered in each simulation by varying the A-

factor of (R1), then the calculated auto-ignition delay times are not statistically different (i.e. the error bounds for each

calculated auto-ignition time are overlapping).

As noted above, there are advantages and drawbacks to each simulation method, with a direct tradeoff between universal

applicability and simplicity. In the literature, all of these methods have been found to be generally successful (Sander

Gersen et al. 2012; S. Walton et al. 2007; Kalitan et al. 2007); however, also described above, there are assumptions

underlying each method that are not always valid. If these assumptions are invalid, then the comparison between

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experimental results and the chemical mechanism, facilitated by the kinetic simulation, lose meaning. Rather than

explicitly choose a simulation method to use in the present work, all three methods were used to facilitate comparison

and understanding.

3. Results and Discussion

A summary of the results of the current work is presented in Table A of the Supplemental Material, which includes the

mixture composition, thermodynamic conditions, measured ignition delay time, ignition behavior, and simulated ignition

delay times for each experiment. In the following section, the data are illustrated and discussed using Figures 4 and 5.

In Figure 4, the auto-ignition delay time measurements and simulation results are plotted as a function of inverse

temperature. In Figure 5, the ignition behavior (strong, weak, etc.) is plotted as a function of pressure and temperature,

and include ignition behavior data from S. M. Walton et al. (2007) and Kalitan et al. (2007) for comparison.

Figure 4. Auto-ignition delay time (τign) as a function of inverse temperature. Vertical error bars on the experimental

data are the uncertainty in the measured ignition delay time and horizontal error bars are uncertainty in the calculated

temperature. Simulation data are presented using the end of compression temperature for Methods 1 and 2, and

effective temperature for Method 3. Vertical error bars on the simulation data are the limits of the simulation results

when considering the uncertainty in the A-factor of (R1) (Li et al. 2007). The dashed and dotted lines represent the error

bounds of adiabatic constant volume simulations for 3 atm and 15 atm as defined by the uncertainty in reaction (R1).

The results illustrated in Figure 4 are an indication that the three simulation methods considered here yield similar

results. Furthermore, the results indicate that the Li et al. (2007) chemical mechanism is describing the syngas

combustion kinetics within the experimental uncertainties at these conditions. As evidenced by the guidelines in Figure

4, the uncertatinty in the chemical mechanism can result in a rather large band of expected ignition delay values –

especially for the 3 atm experiments. It is therefore to be expected that the experimental and simulation results for 3 atm

have correspondingly large error bounds. It is important to note that the guidelines mark the boundaries for expected

behavior at 3 atm and 15 atm. As seen in Table A in the supplemental data, the pressures in the experiments vary around

3 and 15 atm (e.g. the experimental data span 2.6 – 2.87 atm and 13.4 to 17.2 atm), explaining some of the scatter in the

experimental data.

It is useful to compare the results of the three simulation methods to identify which is most appropriate to use at these

conditions. However, given that a goal of these experiments and simulations is to evaluate the chemical mechanism,

10

care must be taken not to convolve the evaluation of the chemical mechanism with the evaluation of the simulation

method. The procedure of evaluating different simulation methods has not appeared in literature, as in most cases a

single method is chosen and assumed to be appropriate. Key to this evaluation process is the consideration of the

uncertainty in the chemical mechanism.

To evaluate each simulation method, the key underlying assumptions of each were considered. For Method 2, there is an

assumption that the pressure decrease during the induction period is the result of heat losses to the environment only and

there is no significant exo- or endothermic chemistry during the induction period. The majority of the present

experiments were located in the “branched chain-explosion” regime of the H2/O2 explosion map (i.e. temperatures

greater than ~1050 K at 15atm and temperatures greater than ~1000K at 3atm (Lieuwen et al. 2009)), indicating the test

gas mixtures are dominated by the explosive nature of HO2 chemistry, not exothermic heat addition. Therefore, it can be

assumed that the underlying assumption for Method 2 is valid. It is important to note, however, that since the explosion

map regimes are defined by the chemical mechanism itself, this validation is not entirely independent from the reaction

chemistry.

For Method 1, there is an underlying assumption that the chemical kinetics are not affected by the experimentally

observed pressure decrease. This assumption can be validated through a comparison of the results of Method 1 and

Method 2. Because simulations using Method 2 incorporate heat transfer effects, the comparison indicates the extent to

which heat transfer affects the ignition delay time. As seen in Figure 4, the auto-ignition delay time results of Methods 1

and 2 are within the error bounds of the simulation methods. Thus, heat transfer effects are not sufficiently large to

differentiate between the two modeling approaches. It follows that the experimental results should be reported at the

end-of-compression thermodynamic conditions, as seen in Figure 4. Given the short auto-ignition delay times reported

here (< 10 ms) this result agrees with findings of S Gersen et al. (2008) who found heat transfer effects for experiments

with short auto-ignition delay times (< 2-3ms) could be neglected. Again it is important to note that the chemical

mechanism is not entirely independent from this validation, though agreement with other experimental findings supports

the argument.

For Method 3, there is an underlying assumption that the major effects of the pressure decrease on the chemical reaction

rates can be represented by average pressure and temperature values that are shifted to values slightly lower than the end

of compression conditions. This assumption can be validated by comparison of the results of Method 3 and Method 2.

Because simulations using Method 2 incorporate the time dependent heat transfer effects, this comparison indicates

whether a bulk shift in the initial state conditions sufficiently captures these effects. As seen in Figure 4, the ignition

delay time results of Methods 3 and 2 are also within the error bounds of the simulation methods (where the predicted

values for ign are reported at Teff for Method 3). It follows that the experimental results should be reported at the

effective thermodynamic conditions. However, given that Method 1 was shown to be valid, i.e. the heat transfer effects

can be neglected, the validation for Method 3 is trivial. This highlights an important behavior though, given that in some

cases both Methods 1 and 3 are valid, implying that the experimental data could be reported at either end-of-compression

or effective thermodynamic conditions.

While for the present experiments all three simulation methods were shown to yield essentially the same results

(predicted ignition delay times have overlapping error bounds, see Table A), this will not likely be the case for

experiments with longer ignition delay times. As illustrated in the literature (S. M. Walton et al. 2007; Lee & Hochgreb

1998; Mittal et al. 2006; S Gersen et al. 2008), heat losses for longer test times can reach levels where they significantly

impact the ignition delay time (error bounds for Method 1 will fall outside of those for Method 2 and 3). For these

experiments, this process of evaluating different simulation methods need not be conducted each time, as done in this

work. Simulations methods could be evaluated using selected experiments, generally representing the limiting test

conditions, and once the appropriate assumptions have been validated, the methods could be extended to other

experimental conditions.

Upon review, it is clear that given the simplicity of Method 1 and the ease of comparing results across experimental

facilities, it is desirable to use this method whenever possible. However, Method 1 is likely restricted to experiments

with very short (< 10 ms) auto-ignition delay times. Given the ability of Method 2 to capture heat transfer effects, it is

desirable to use this method for cases with significant heat transfer. However, Method 2 introduces difficulty in how to

meaningfully report the results. Method 3 seems therefore to be the most useful, as it allows for a consideration of heat

transfer effects, like Method 2, while retaining ease and clarity in reporting, like Method 1.

11

Figure 5. Ignition behavior as a function of pressure and temperature, includes data from Kalitan, et al. (Kalitan et al.

2007) and Walton, et al. (S. M. Walton et al. 2007)

The results illustrated in Figure 5 form a map of ignition behavior, indicating the category of ignition (weak or strong)

for a given thermodynamic condition. These results include data from previous studies on syngas (H2 and CO only)

ignition at similar conditions (S. M. Walton et al. 2007; Kalitan et al. 2007), which together with the present results span

a range of ϕ from 0.1-0.5, H2:CO from 0.05 – 4.0, and dilution from 65-75%. Though Kalitan et al. did not explicitly

determine ignition behavior, any point which was found to exhibit “detonation-like ignition preceded by early OH*

emission” were deemed to have weak ignition behavior for the purposes of this work (Kalitan et al. 2007). Walton et al.

determined ignition behavior in the same manner as this work, so data from that study can be directly compared.

However, to maintain consistency, data from that work is reported here at end-of-compression thermodynamic

conditions, not at effective conditions as originally reported (S. M. Walton et al. 2007).

A goal of this analysis was to evaluate potential connections between ignition regimes and transitions in explosion

regimes on the explosion map. These regimes, as described in Lieuwen et al. (2009), are generally classified as follows:

(1) Branched-chain explosion, dominated by H, O, OH chemistry, to the high temperature side of the classical 2nd

limit,

(2) Branched-chain explosion, dominated by HO2, H2O2, OH chemistry, between the classical 2nd

limit and the extended

2nd

limit, (3) Thermal-chain explosion, dominated by HO2, H2O2 exothermic chemistry, between extended 2nd

limit and

3rd

limit, (4) No explosion, to the low temperature side of the 3rd

limit. It is important to note that these limits describe

the H2/O2 chemistry only; any effects of CO chemistry are not included.

The classical and extended 2nd

limits plotted in Figure 5 were calculated according to formulae listed in (Lieuwen et al.

2009), using the Li et al. (2007) chemical mechanism with the nominal values for all reaction rate parameters. These

limits are purely a function of reaction rate parameters and do not rely on mixture composition, thereby allowing a direct

comparison of ignition data that has widely varying mixture compositions. The 3rd

limit was calculated using a

CHEMKIN kinetic simulation with the Li et al. (2007) chemical mechanism, for a simple mixture with ϕ = 0.5, H2:CO =

0.7, dilution = 75% N2 .

Upon examination of the results it is clear that ignition behavior makes a distinct change as it crosses either the classical

2nd

or extended 2nd

explosion limit, at all pressures considered. At a given pressure as the temperature is decreased and

the 2nd

limit is crossed (for 3atm – the classical, for 15atm – the extended), the ignition changes from strong behavior to

12

weak behavior/ no ignition. This suggests that the ignition behavior of syngas, seen in multiple experimental facilities, is

strongly tied to the thermodynamic conditions of the experiment, and the relative location of the state conditions on the

H2/O2 explosion map. Furthermore, the ignition behavior is likely tied to the dominant chemistry of that thermodynamic

location.

An important detail is revealed upon comparing of the results from the present study to data from S. M. Walton et al.

(2007). As the extended 2nd

limit is crossed for the present data at 15atm, the ignition characteristics transition from

strong ignition to no ignition. Whereas, for data at similar pressures from Walton et al. there is a transition from strong

ignition to weak ignition behavior across this limit, followed by an eventual transition to no ignition as the temperature is

decreased further. The experiments in Walton et al. were conducted at nearly the same conditions as the current work,

except the equivalence ratio was set at = 0.5, higher than the equivalence ratio of the present data ( = 0.1). This may

be an indication that the existence of weak ignition behavior at these conditions is dependent on equivalence ratio. This

is consistent with the chemistry of this regime, given that in the thermal-explosion regime explosions are driven by a

temperature increase from exothermic HO2, H2O2 chemistry (Lieuwen et al. 2009). At low equivalence ratios, heat loss

to the cool walls of the test volume could exceed the heat addition from exothermic reactions and the mixture may not

experience sufficient temperature increase to achieve explosion. This result has practical importance from an operational

safety perspective in a gas turbine combustor pre-mixer, as it may indicate a limit on equivalence ratio below which

stable non-igniting conditions can be achieved.

Another interesting behavior seen in Figure 5 is that the transition across the classical 2nd

limit seems to affect ignition

behavior at pressures below 5 atm, but has little effect above 5 atm. This is likely because the pressure is not sufficiently

high to allow for explosive HO2 chemistry, which is driven by 3rd

body collision chemistry ( H2O2(+M) = OH + OH(+M)

(Li et al. 2007) ). What is not indicated in the results above is the pressure at which the transition between strong and

weak ignition moves from the classical 2nd

limit to the extended 2nd

limit.

This map was formed using experiments that were experiencing weak ignition events, likely stemming from local

disturbances like thermal stratification induced by gas dynamic effects or impurities. Therefore it cannot be known if

this map indicates the sensitivity of the ignition behavior to a disturbance at a given thermodynamic condition, or if it

indicates an increased likeliness to create a disturbance at that condition. However, the map is valuable in that the results

show clear ignition behavior changes that do not appear to be random. Furthermore, these ignition behaviors would be

seen in real applications which experience disturbances likely of greater magnitude than what exists in well-controlled

laboratory experiments.

Conclusions

The current work presented the results of new syngas ignition experiments using the UM RCF, where ignition delay

times were measured and the ignition behavior was classified. The results demonstrate that for experiments with strong

ignition behavior the Li et al. (2007) chemical mechanism applied in a zero-dimensional homogeneous reactor

simulation can accurately predict the measured auto-ignition delay times. The uncertainties in the key reactions in the

detailed mechanism were also quantified in this study and shown to be significant at the conditions of interest to gas

turbine combustors.

Three simulation methods from the literature were compared with the experimental results using the Li et al. chemical

mechanism. Each method, with a different treatment of heat transfer effects, was shown to yield similar results (within

quantified error bounds) for the experimental conditions studied, indicating that the effects of heat transfer on these

experiments were negligible at these conditions. For other experiments with longer test times, this will likely not be the

case, and appropriate criteria should be applied to both model the experimental data and for reporting the experimental

results. An evaluation of the simulation methods indicated that defining effective thermodynamic conditions is likely

the most useful, allowing for the inclusion of first-order heat transfer effects while retaining ease and clarity in reporting

the results.

Finally, a close relationship between transitions in ignition behavior (strong/weak/no ignition) and transitions across the

classical and extended 2nd

limits on the H2/O2 explosion map was demonstrated using a pressure/temperature map of

ignition behavior. This behavior seems to be largely unaffected by reactant mixture composition, though the existence of

weak ignition behavior may be linked to equivalence ratio.

13

Acknowledgements

The authors acknowledge the generous financial support of the Department of Energy, National Energy Technology

Laboratory via the University Turbine Systems Research Program with DOE Project Manager Mark Freeman and the

Department of Mechanical Engineering at the University of Michigan. The authors also acknowledge Mohammad

Fatourie for his valuable insights.

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16

Supplemental Material

Table A. Summary of experimental conditions and results (f)

* CHEMKIN (Reaction Design 2010) simulation using: (a) Method 1, (b) Method 2, (c) Method 3

(d) Balance Ar, (e) error of (± 0.0) indicates that the error is less than the significant digits of the nominal value

(f) Error reported in same units as nominal value

φ H2:

CO

Test Gas Composition [%]d PEOC

[atm]

TEOC

[K]

PEff

[atm]

TEff

[K]

Ignition delay time [ms]

χH2 χCO χO2 χN2 χCO2 χAr τign τsim-EOCa τsim-vol

b τsim-effc

STRONG IGNITION

0.1 0.7 1.7 2.5 20.8 68.3 0 6.7 17.2

(±0.2)

1133

(±2)

16.55 1122 2.1

(±0.8)

1.7

(+0.8,-0.6)

3.5

(+3.1,-1.7)

1.6

(+1.3,-0.5)

0.1 0.7 1.7 2.5 20.8 68.3 0 6.7 15.4

(±0.1)

1142

(±2)

- - 1.4

(±0.3)

1.2

(+0.6,-0.5)

- -

0.1 0.7 1.7 2.5 20.8 68.3 2.0 5.0 14.5

(±0.1)

1096

(±2)

13.83 1083 5.2

(±1.4)

4.0

(+1.7,-1.2)

11.4

(+12.1,-6.1)

5.1

(+2.1,-2.0)

0.1 0.7 1.7 2.5 20.8 68.1 2.0 5.0 14.3

(±0.1)

1094

(±1)

13.89 1086 5.6

(±2.2)

4.2

(+1.7,-1.3)

7.7

(+3.1,-1.7)

4.8

(+2.0,-1.9)

0.1 0.7 1.7 2.5 20.8 68.1 2.0 5.0 13.4

(±0.1)

1073

(±1)

12.89 1064 7.1

(±1.3)

6.6

(+2.5,-1.9)

9.1

(+3.1,-1.7)

7.9

(+2.9,-2.3)

0.1 0.7 1.7 2.4 20.8 67.9 3.9 3.2 2.6

(±0.1)

1025

(±6)

- - 4.0

(±1.2)

1.7

(+7.0,-1.2)

- -

0.1 0.7 1.7 2.4 20.8 63.2 4.0 7.9 2.87

(±0.1)

1050

(±6)

- - 1.9

(±1.1)

0.5

(+2.2,-0.3)

- -

WEAK STRONG IGNITION

0.1 0.7 1.7 2.4 20.8 67.9 3.9 3.2 2.85

(±0.0)(e)

1019

(±4)

- - - - - -

0.1 0.7 1.7 2.4 20.8 74.8 0.3 0 2.79

(±0.0)

1034

(±3)

- - - - - -

0.1 0.7 1.7 2.4 20.8 67.9 3.9 0 3.14

(±0.0)

1033

(±3)

- - - - - -

INDETERMINATE IGNITION

0.1 0.7 1.7 2.6 20.8 74.8 0 0 3.12

(±0.0)

1035

(±3)

- - - - - -

0.1 0.7 1.7 2.4 20.8 74.8 0.3 0 2.65

(±0.0)

1020

(±2)

- - - - - -

0.1 0.7 1.7 2.4 20.8 68.9 5.2 0 3.24

(±0.1)

1012

(±4)

- - - - - -

NO IGNITION

0.1 0.7 1.7 2.4 20.8 50.3 24.7 0 3.1

(±0.0)

888

(±1)

- - - - - -

0.1 0.7 1.7 2.4 20.8 50.3 24.7 0 3.21

(±0.1)

895

(±4)

- - - - - -

0.1 0.7 1.7 2.5 20.8 62.8 12.3 0 2.99

(±0.0)

954

(±2)

- - - - - -

0.1 0.7 1.7 2.5 20.8 62.8 12.3 0 3.07

(±0.0)

960

(±3)

- - - - - -

0.1 0.7 1.7 2.5 20.8 62.6 12.4 0 14.44

(±0.0)

946

(±0)

- - - - - -

0.1 0.7 1.7 2.5 20.9 62.5 12.4 0 15.9

(±0.0)

969

(±1)

- - - - - -

0.1 0.7 1.7 2.4 20.8 69.9 5.1 0 2.99

(±0.0)

1002

(±2)

- - - - - -

0.1 0.7 1.7 2.4 20.8 69.8 5.2 0 13.85

(±0.0)

985

(±1)

- - - - - -

0.1 0.7 1.7 2.4 20.8 69.8 5.2 0 15.11

(±0.0)

1006

(±0)

- - - - - -

0.1 0.7 1.7 2.4 20.8 74.8 0.3 0 14.28

(±0.1)

1042

(±1)

- - - - - -

0.1 0.7 1.7 2.4 20.8 74.8 0.3 0 13.8

(±0.0)

1032

(±1)

- - - - - -