NISTIR 7445

January 2002

THE STRUCTURE AND 

EXTINCTION OF LOW STRAIN 

RATE NON-PREMIXED FLAMES 

BY AN AGENT IN MICROGRAVITY

Anthony Hamins 

Matthew Bundy 

Chang Bo Oh 

S. Paul Fuss 

Jennifer Logue

ii

U.S. Department of Commerce 

Carlos M. Gutierrez, Secretary 

Technology Administration 

Robert Cresanti, Under Secretary of Commerce for Technology 

National Institute of Standards and Technology 

 William Jeffrey, Director

September 2007

THE STRUCTURE AND 

EXTINCTION OF LOW STRAIN 

RATE NON-PREMIXED FLAMES 

BY AN AGENT IN MICROGRAVITY

Anthony Hamins 

Matthew Bundy 

Chang Bo Oh 

S. Paul Fuss 

Jennifer Logue

Building and Fire Research Laboratory

NISTIR 7445
THE STRUCTURE AND EXTINCTION OF 

LOW STRAIN RATE NON-PREMIXED FLAMES BY AN AGENT IN 

MICROGRAVITY

A Final Report on NASA Contract NNC04CB07C 

Period of Support (2/19/2004 to 3/31/2007)

by 

The National Institute of Standards and Technology 

Building and Fire Research Laboratory 

Gaithersburg, MD 20899 

 

 

Please note that certain trade names and company products are mentioned in the text to specify adequately

the experimental procedure and equipment used or to identify types of currently available commercial 

products. In no case does such identification imply recommendation or endorsement by the National 

Institute of Standards and Technology, nor does it imply that the products are necessarily the best 

available for the purpose.

iii
FINAL REPORT

THE STRUCTURE AND EXTINCTION OF 

LOW STRAIN RATE NON-PREMIXED FLAMES BY AN AGENT IN MICROGRAVITY 

Table of Contents 

1. Overview ............................................................................................ 1 

2. Description of Suppression Experiments and Experimental Configuration ........... 2 

3. Axisymmetric Flame Simulations investigating Nitrogen as an Agent .................... 4 

4. Axisymmetric Flame Simulations investigating CO2 as an Agent ........................... 5 

5. Suppression Limits of Low Strain Rate 1-g Non-Premixed Propane Flames .................... 5

6. Spectroscopy .......................................................................................................... 6 

7. References .......................................................................................................... 7 

8. List of Publications and Conference Proceedings ........................................................ 8 

9. Appendix 1: The Two-Dimensional Structure of Low Strain Rate Counterflow Nonpremixed

Methane Flames in Normal and Microgravity Flames ............................................................ 9 

10. Appendix 2: Effect of Buoyancy on the Radiative Extinction Limit of Low Strain Rate 

Non-Premixed Methane-Air Flames...................................................................................... 39 

11. Appendix 3: Numerical Study of the Suppression of Methane-Air Flames with 

Carbon Dioxide in Normal Gravity and Microgravity ........................................................ 65

12. Appendix 4: Suppression Limits of Low Strain Rate 1-g Non-Premixed Propane 

Flames....................................................................................................... 87 

13. Appendix 5: A Technique for Extrapolating Absorption Coefficient Measurements to High 

Temperatures .............................................................................. ........................ 116 

14. Appendix 6: Determination of Planck Mean Absorption Coefficients for CF3Br .................. 139 

15. Appendix 7: Determination of Planck Mean Absorption Coefficients for Hydrocarbon Fuels .............................................................................................................................. 149 

iv
THE STRUCTURE AND EXTINCTION OF LOW STRAIN RATE NON-PREMIXED 

FLAMES BY AN AGENT IN MICROGRAVITY

This final report describes the study that has been supported through the National Aeronautics 

and Space Administration (NASA) microgravity grants program from February 2004 through 

March 2007. This work was a continuation of a NASA microgravity grant that was active from

February 1999 through September 2003. The work has used computations and measurements to 

investigate the structure and suppression of low strain rate non-premixed flames. This final 

report, on the latest grant, includes a summary of the work done on the contract and a list of 

publications. 

The work during this grant period focused on both measurements and calculations and included 

the following: 

• Continued development of the two-dimensional simulation with finite rate chemistry. 

[Hamins et al., 2007; Oh et al., 2007]. These publications are provided in Appendices 1 and 

2. These calculations considered extinction of methane-air diffusion flames with N2 added to 

the fuel stream. The work compares calculations with our earlier measurements made in 

normal gravity and in microgravity using the NASA 2.2 s drop tower. 

• Two-dimensional calculations were completed that considered the extinction of methane-air 

diffusion flames with CO2 added to the fuel stream in normal and microgravity. This study is 

provided in Appendix 3. The computations were compared to the N2 addition results 

described in [Hamins et al., 2007; Oh et al., 2007]. 

• Flame extinction studies on the normal gravity extinction of propane were conducted. The 

results of this study are provided in Appendix 4. 

• Finally, spectroscopic measurements were conducted on key flame intermediates to support 

detailed modeling of radiative exchange in near-extinction flames. Wakasuki et al. [2005] 

developed a method to estimate the temperature dependence of absorption coefficients 

through an extrapolation algorithm. This publication is provided in Appendix 5. 

Spectroscopic measurements are important to understand the mechanisms of extinction 

associated with flames being suppressed by agents. CF3Br (Halon 1301) is one of the most 

important suppressants and has been used on space platforms for many years. Suppression 

studies often use CF3BR as a reference or baseline in considering the effectiveness of 

alternative agents. As part of this study, Fourier Transform Infrared Spectroscopy (FTIR) 

measurements were made on CF3Br at elevated temperatures. Those measurements are 

reported in Appendix 6. Appendix 7 presents similar spectral measurements and analysis for 

propane, heptane and propylene. 

 

The seven manuscripts mentioned above are included as appendices in this final report. The 

remainder of the report presents an overview of the work and the main findings of the project. 

Detailed information on each of the topics can be found in the appendices. 

1. Overview 

When the original grant proposal was written in 1997, a review of the literature indicated that 

there had been little work on the extinction of flames by an agent in microgravity, and practically 

none on low strain rate diffusion flame extinction by an agent. There have been several relevant 

studies since that time. 

1
The work of Maruta et al. [1998] is mentioned in several places in this report. Maruta conducted 

the first comprehensive extinction measurements of very low strain nonpremixed flames in 

microgravity using the 10 s JAMIC drop tower in Japan. In that study, the extinction of 

methane-air diffusion flames with N2 added to the fuel stream was measured under microgravity 

conditions. The minimum methane concentration required to sustain combustion was measured 

to decrease as the strain rate decreased until a critical value was observed. As the global strain 

rate was further reduced, the required methane concentration increased. This behavior was 

denoted as a "turning point" and was attributed to the enhanced importance of radiative loss in 

low strain rate flames. In terms of fire safety, the turning point agent concentration represents a 

fundamental limit for nonpremixed flames, assuring suppression for all flow conditions. Bundy 

et al. [2003] observed similar behavior in low strain rate normal gravity flames. 

Most recently, Han et al. [2007] reported on the structure of ultra low-stretch (≈2 s-1) methane 

nonpremixed flames. Their study considered radiation effects and extinction in curved flames 

established in normal gravity by flowing diluted fuel in a downward direction through a porous 

spherically symmetric burner with a large radius of curvature. Heat losses to the burner were 

characterized and the computational results were found to be consistent with experimental 

observations. This study was different in the sense that conductive heat losses to the burner were 

always avoided. This impacted the range of strain rates that were investigated, because flame 

thickness increases as the strain rate decreases. Here, the global strain rate (ag) was varied from

7 s-1 to 50 s-1 in microgravity flames and from 14 s-1 to 50 s-1 in normal gravity flames to avoid 

conductive heat losses to the burner. This was confirmed through consideration of the profile of 

temperature near the burner boundaries To maximize the range of strain rates considered in the 

normal gravity flames, the approach used in Bundy et al [2003] was applied in which the 

velocity ratio of the air stream to the fuel stream was adjusted such that the flame location was 

forced towards the center of the flowfield, away from the burner ducts. By avoiding heat losses 

to the burner, it is possible to have an understanding of flame structure while avoiding 

inadvertent flame holder effects, which may be important in the stability of a flame (i.e., flame 

suppression studies). 

2. Description of Suppression Experiments and Experimental Configuration 

The experimental apparatus consisted of a counterflow burner, a gas supply system for the air, 

agent, and fuel, an igniter, an electronic sequencer, and a flash-memory computer processor to 

control the electronics and acquire data. The burner, shown in Fig. 1, consisted of two 15 mm

inner diameter tubes aligned along the same axis with a separation distance of 15 mm. The 

flame was established between the burner ducts. Fine mesh steel screens (200 mesh) were 

mounted at the burner duct exits to straighten the gas flow. The burner and combustion chamber 

design considered the volume of combustion products formed and consequent changes in 

chamber pressure. The key to the suppression measurement was control of the reactant flows, 

which was facilitated by fast response (≈10 ms) pressure controllers positioned upstream of 

selected critical flow orifices. The time response of mass flow controllers was too slow (≈0.7 s) 

for a 2.2 s drop experiment. 

In the experiments, conditions were selected to repeat the work of Maruta et al. [1998], who 

performed experiments in the 10 s JAMIC drop tower. Figure 2 summarizes our suppression and 

2
modeling results. Experiments were in general agreement with the results of Maruta. The 

simulations in Fig. 2 are discussed below. In addition to the flame suppression measurements, 

the time varying temperature was determined in the near-extinction microgravity flames using 

thin filament intensity measurements. The method was similar to that used by Pitts [1998]. 

Hamins et al [2007] describes the temperature measurements. 

Oxidizer Stream 

Porous sintered metal 

Fuel + Agent 

 Streams

L = 15 mm

15 mm diameter 

Stainless Steel Burner 

Flame

Wire Mesh Screen 

Porous sintered metal 

Figure 1. Stainless Steel Opposed Flow Diffusion Flame (OFDF) Burner with agent added to the fuel stream. 

Some experiments will add agent to the oxidizer stream. 

0.74

0.76

0.78

0.80

0.82

0.84

0.86

0.88

0 10 20 30 40 50 60 70 8

Global Strain Rate (1/s)

Critical Volu

me Fractio

n at 

E

xtinctio

n

0

μ-g experiments [Maruta et al., 1998]

Current μg experiments

1-g experiments [Bundy et al., 2003]

0-g, 2D model

1-g, 2D model

0-g, 1D model

Figure 2. Computed and measured critical nitrogen volume fraction in the fuel stream for suppression of CH4 flames 

in microgravity (μ-g) and normal gravity (1-g). Our microgravity N2 fuel stream extinction measurements are 

3
compared with those of Maruta et. al [1998], with Oppdif simulations based on the assumption of an optically thin

radiating gas, and with the 2D model developed here (also assuming the optically thin limit). 

3. Axisymmetric Flame Simulations investigating Nitrogen as an Agent

A transient two-dimensional flame simulation based on one step chemistry was developed to 

simulate the drop tower experiments. The details can be found in Oh et al. [2007] (see Appendix 

1). Figure 3 is an example of the model results. The figure compares the computed flame heat 

release rate with photos of flame shape in normal gravity and microgravity for methane-air 

nonpremixed flames at ag=20 s-1 and a fuel stream N2 volume fraction of 0.80. 

 (a) 1-g (b) 0-g 

Figure 3. Comparison of flame shape in normal gravity and microgravity for methane-air nonpremixed flames at 

ag=20 s-1 and a fuel stream N2 volume fraction of 0.80. The left hand sides are photographs, while the right hand

sides are simulation results. (The thin blue flames are faint, but clearly visible when the photos are viewed in color.) 

The numerical model is a time-dependent axisymmetric two-dimensional (2D) computation that 

considers buoyancy effects and radiative heat transfer. A three-step global reaction mechanism

was used in both the 1D and 2D computations to predict the measured extinction limit and flame 

temperature. Photographic images of flames undergoing the process of extinction were compared 

with model calculations. The axisymmetric numerical model was validated by comparing flame 

shapes, temperature profiles, and extinction limits with experiments and with the 1D 

computational results. The 2D computations yielded insight into the extinction mode and flame 

structure. The contribution by each term in the energy equation to the heat release rate was 

evaluated to investigate the multi-dimensional structure and radiative extinction of the normal 

gravity (1-g) flames. Two combustion regimes depending on the extinction mode were 

identified. Lateral heat loss effects and multi-dimensional flame and flow structure were also 

found. At low strain rates in 1-g flames (“Regime A”), the flame is extinguished from the weak 

outer edge of the flame, which is attributed to a multi-dimensional flame structure and flow field. 

At high strain rates, (“Regime B”), the flame extinction initiates near the flame centerline due to 

an increased diluent concentration in the reaction zone, similar to the extinction mode of 1D 

4
flames. These two extinction modes can be clearly explained by consideration of the specific 

maximum heat release rate. 

Hamins et al. [2007] studied the structure and extinction of the non-premixed flames through 

comparison of the counterflow suppression experiments and temperature measurements (global 

strain rates from 7 s-1 to 50 s-1 ) with the 2D calculations (see Appendix 2). The measurements 

were used to validate the two-dimensional flame simulation. The simulations yielded insight into 

the differences between microgravity and normal gravity suppression results, and also explained 

the inadequacy of the one-dimensional model results to explain the microgravity suppression 

results. Figure 2 highlights these results and is a focus of the study by Hamins et al. [2007]. 

Figure 2 also shows the results of the 1D simulations, which exhibit gross trends similar to the 

measurements. The calculations deviated from Maruta’s measurements for the lowest strain 

conditions, where the calculations overpredict the required agent concentrations. The 2D model 

successfully predicted the microgravity results [Hamins et al., 2007]. The use of the 

axisymmetric simulation yielded valuable information on flame shape and the effects of 

buoyancy on low strain rate normal gravity flames. By providing physical insight, the 

axisymmetric simulations are a useful tool for guiding the design and interpretation of drop 

tower experiments where the gravitational level changes or for simulating flame structure and 

extinction in a relatively moderate gravitational environment such as the moon. 

4. Axisymmetric Flame Simulations investigating Carbon Dioxide as an Agent

The objective of this study was to numerically investigate the structure and extinction 

characteristics of low strain rate counterflowing methane-air nonpremixed flames using carbon 

dioxide as the agent. An axisymmetric 2D computational model, including buoyancy effects and 

radiative heat transfer was developed. The model was validated through comparison of the 2D 

model with experiments and 1D computational results for counterflow nonpremixed flames diluted 

by nitrogen. Once validated, the calculation was used in this study to explore the near-extinction 

flame structure for flame suppressed by CO2. The details of this study are presented in 

Appendix 3. 

The simulations yielded insight into the effectiveness of CO2 as an agent, highlighting differences 

in the radiative loss when CO2 rather than N2 is used as a suppressant. The differences between 

microgravity and normal gravity suppression results were modeled. The results lends confidence in 

the viability of the axisymmetric model, extending its application to cases beyond methane-air 

flames with N2 as an agent. 

5. Suppression Limits of Low Strain Rate Non-Premixed Propane Flames

The suppression of low strain rate non-premixed propane flames was investigated experimentally 

in a counterflow configuration. The detailed results of this study are presented in Appendix 4. 

The critical concentration of N2, CO2, and CF3Br added to the fuel and the critical concentration 

of N2 added to the oxidizer stream required to obtain extinction of propane-air non-premixed 

flames was measured as a function of the global strain rate. In agreement with previous 

measurements performed with methane-air diffusion flames, limiting non-premixed flame 

5
extinction behavior in which the agent concentration obtained a value that insures suppression 

for all global strain rates was observed. 

A series of extinction measurements varying the air: fuel velocity ratio showed that the critical 

N2 concentration was invariant with this ratio, unless conductive losses were present. In terms of 

fire safety, the measurements confirm the existence of a fundamental limit for suppressant 

requirements in normal gravity flames as was previously seen in methane-air flames. This can 

be thought of as analogous to agent flammability limits in premixed flames. The critical agent 

volume fraction in the propane fuel stream assuring suppression for all global strain rates was 

measured to be 0.923 ± 0.003 for N2, 0.879 ± 0.004 for CO2, and 0.517 ± 0.01 for CF3Br. These 

results were somewhat higher than the results for methane. The critical agent volume fraction in 

the oxidizer stream assuring suppression for all global strain rates in the propane flames was 

measured as 0.354 ± 0.003 for N2.. 

 

Measurements using thin filament pyrometry measures flame temperatures for diluted propaneair flame near extinction for several different strain rates. Insight was gained into the near 

extinction flame temperatures as a function of strain rate. 

6. Spectroscopy

To support the calculation of the structure and extinction of diffusion flames suppressed by 

Halon 1301, spectral transmission measurements were conducted on C3H8, C7H16, C3H6, and 

CF3Br at elevated temperatures using Fourier Transform Infrared Spectroscopy [Wakasuki et al., 

2007; Fuss et al., 2001]. A Mattson model 7020 FTIR was arranged so that the infrared (IR) 

beam passed through the interferometer but is then diverted outside the spectrometer, bypassing 

the standard sample compartment, through a quartz flow cell located inside a three zone tube 

furnace and onto an external detector. A detailed description of the apparatus and procedure is 

given in Wakasuki et al. [2007] (see Appendix 5) and Fuss et al. [2001]. The entire optical path 

was purged with N2 to eliminate absorption from ambient water vapor and CO2. The flow cell 

provided a measurement path length of 0.32 m and the ends were sealed with Zinc Selenide 

(ZnSe) windows, which transmit in the spectral region of interest. This experimental 

configuration measured spectra at temperatures as high as ≈700 K. It will be possible to 

extrapolate this data to higher temperatures using the approach developed in Appendix 5. That 

methodology fits the absorption coefficient parameters in a simplified expression, and uses the 

parameters to extrapolate data to higher temperatures [Wakasuki et al., 2005]. The approach was 

validated using absorption data for water vapor, CO, and CO2 from the HITEMP database 

[Hitran, 1998]. 

After acquiring transmission measurements, the temperature dependent Planck mean absorption 

coefficients for CF3Br were calculated for the first time (see Appendix 6). Previously, Planck 

mean absorption coefficients were determined for HF and HBr, species that are important 

products when CF3Br is present [Fuss and Hamins, 2002]. These species participate in radiative 

exchange in a flame or fire when CF3Br is present. HF and HBr are the major F and Br

containing molecular products [Hamins et al., 1994b] when CF3Br is added to a flame. In 

preparation for radiative calculations involving CF3Br, we have used previously developed high 

temperature spectral information for HBr and HF [Rothman et al., 1992] to calculate the Planck 

absorption-emission coefficients for these molecules [Fuss and Hamins, 2002]. For molecules 

6
not in the high temperature spectral database such as C3H8 and CF3Br, the infrared spectra was 

measured using a high temperature furnace in conjunction with an FTIR instrument. Planck 

mean absorption coefficients for C3H8 and CF3Br are useful in this regard, allowing narrowband 

spectral parameters to be determined for these molecules to provide input data for the RADCAL 

radiation sub-model. Similar measurements and analysis (Planck determination) are presented 

for C3H8, C7H16, and C3H6 in Appendix 7. 

In CF3Br inhibited flames, Seshadri [1975] and Masri [1994] report that the volume fraction of 

CF3Br is comparable to that of CO in the high temperature reaction zone. For a flame with the 

CF3Br suppressant added to the air stream of a diffusion flame, the concentrations of CO2, H2O, 

and CO will fall as a function of distance from the high temperature reaction zone, and CF3Br 

will be larger than any other species that participates in radiative exchange. In cool flame 

regions, CF3Br may absorb flame radiation, and block its emission to the surroundings. The 

inclusion of CF3Br in radiative emission calculations of CF3Br inhibited hydrocarbon flames 

may be significant and ought to be considered, depending on the combustion configuration. 

Estimates of the impact of CF3Br in the cool flame region showed that, for optically thin 

conditions, CF3Br was found to account for approximately 80 % of the total emission in the 

region approximately 1.5 mm from the reaction zone, where the temperature is 675 K. In 

summary, CF3BR is expected to play a significant role in the total radiative emission from the 

flame, which will impact the structure and extinction of these flames. 

6. References

Bundy, M., Hamins, A., and Lee. K.Y., Suppression Limits of Low Strain Rate Non-Premixed 

Methane Flames, Combust. Flame, 133, 299-310 (2003). 

Fuss, S.P., and Hamins, A., Determination of Planck Mean Absorption Coefficients for HBr, 

HCl, and HF, ASME J. Heat Transfer, 124, 26-30 (2002). 

Hamins, A., Trees, D., Seshadri, K., and Chelliah, H., Combust. Flame, 99, 221 (1994b). 

Hamins, A., Bundy, M., Puri, I. K., McGrattan, K. B. and Park, W. C., Proc. 6th Int’l

Microgravity Combustion Workshop, 101 (2001). 

Hamins, A., Bundy, M., Oh, C.B., Kim, S.C., “Effect of Buoyancy on the Radiative Extinction 

Limit of Low Strain Rate Non-Premixed Methane-Air Flames,” Combust Flame, to 

appear, 2007; (see Appendix 2). 

Han, B., Ibarreta, A.F., C.J. Sung, J. T’ien, Combust. Flame, 149 (2007) 173-190. 

HITRAN (high-resolution transmission molecular absorption database), .J Quant. Spec. Rad. 

Tran., 60, 665-710 (1998). 

Lee, K.Y., Cha, D.J., Hamins, A., and Puri, I.K., Combust. Flame, 104, 27 (1996). 

Lutz, A., Kee, R. J., Grcar, J. and Rupley, F. M., “A Fortran Program Computing Opposed Flow 

Diffusion Flames,” SAND96-8243, Sandia National Laboratories, Livermore, CA (1997). 

Maruta, K., Massaharu, Y. Hongsheng, G., Ju, Y., Niioka, T, Combust. Flame, 112, 181 (1998). 

Masri, A.R., “Chemical Inhibition of Nonpremixed Flames of Hydrocarbon Fuels with CF3Br,” 

Combust Sc Tech, 96, pp. 189-212 (1994). 

McGrattan, K. B., Baum, H. R., Rehm, R. G., Hamins, A., Forney, G. P., Floyd, J. E., and 

Hostikka, S., “Fire Dynamics Simulator Technical Reference Guide,” v.2, NISTIR 6467, 

National Institute of Standards and Technology, Gaithersburg, MD (2001). 

7
Macdonald, M.A., Jayaweera, T.M., Fisher, E.M., and Gouldin, F.C., Proc. Twenty-Seventh 

Symp. (Int.) on Combustion, The Combustion Institute, Pittsburgh, pp. 2749-2756, 1998. 

Oh, C.B., Hamins, A., Bundy, M., Park, J., “The Two-Dimensional Structure of Low Strain Rate 

Counterflow Nonpremixed Methane Flames in Normal and Microgravity,” Combust 

Theory Modeling, to appear, 2007; (see Appendix 1). 

Pitts, W.M., Smyth, K. C., and Everest, D. A., Proc. Twenty-Seventh Sym. (Int.) on Combustion, 

The Combustion Institute, Pittsburgh, pp. 563-569, 1998. 

Puri, I.K. and Hamins, A., A Numerical Investigation of Radiative Effects in Near-Extinction 

Methane-Air Nonpremixed Flames, submitted to Combustion & Flame, 2002. 

Rothman, L., Gamache, R., Tipping, R., Rinsland, C., Smith, M., Benner, D., Malathy, D., Flaud, 

J.M., Peyret-Camy, C., Perrin, A., Goldman, A., Massie, S., Brown, L., and Toth, R., J. 

Quant. Spectroc. Radiat. Transf., 48, 469 (1992). 

Seshadri, K. and Williams, F.A., Halogenated Fire Suppressants. ACS Symposium Series 16. 

April 23-24, 1975, San Antonio, TX, American Chemical Society, Washington DC, (Ed.: 

R.G. Gann) pp. 149-182, 1975. 

Smyth, K., and Everest, D., Proc. Twenty-Sixth Symp. (Int.) on Combustion, The Combustion 

Institute, Pittsburgh, pp. 1385-1392, 1996. 

Wakatsuki, K., Jackson, G.S., Kim, J., Hamins, A.. M.R. Nyden, and S.P. Fuss, Determination 

of Planck Mean Absorption Coefficients for Hydrocarbon Fuels, in review, Combust. Sci. 

Tech., 2007. 

7. List of Publications

Refereed Publications:

1. K. Wakatsuki, S.P. Fuss, A. Hamins, M.R. Nyden, G. S. Jackson, and J. Kim, “A technique 

for extrapolating absorption coefficient measurements to high temperatures,” Proceedings of 

the 30th (Int.) Sym. Combust., 1565-1573 (2005) ; see Appendix 5. 

2. A. Hamins, M. Bundy, C.B. Oh, S.C. Kim, “Effect of Buoyancy on the Radiative Extinction 

Limit of Low Strain Rate Non-Premixed Methane-Air Flames,” Combust Flame, to appear, 

2007; see Appendix 2. 

3. C.B. Oh, A. Hamins, M. Bundy, J. Park, “The Two-Dimensional Structure of Low Strain 

Rate Counterflow Nonpremixed Methane Flames in Normal and Microgravity,” Combust

Theory Modeling, to appear, 2007; see Appendix 1. 

4. K.Wakatsuki, G.S. Jackson, J.Kim, A.Hamins, M.R. Nyden, and S.P. Fuss, Determination of 

Planck Mean Absorption Coefficients for Hydrocarbon Fuels, in review, Combust. Sci. Tech., 

2007; see Appendix 7. 

Conference Presentations

1. Hamins, A., Bundy, M., Oh, C.B., Park, and Puri, I.K., “Suppression of Low Strain Rate 

Nonpremixed Flames by an Agent,” presented at the Microgravity Combustion Workshop, 

Cleveland, OH, June, 2004. 

8
Appendix 1 

The Two-Dimensional Structure of Low Strain Rate Counterflow 

Nonpremixed Methane Flames in Normal and Microgravity 

Abstract 

The structure and extinction of low strain rate nonpremixed methane-air flames was studied 

numerically and experimentally. A time-dependent axisymmetric two-dimensional (2D) 

model considering buoyancy effects and radiative heat transfer was developed to capture the 

structure and extinction limits of normal gravity (1-g) and zero gravity (0-g) flames. For 

comparison with the 2D modeling results, a one-dimensional (1D) flamelet computation 

using a previously developed numerical code was exercised to provide information on the 0-g 

flames. A 3-step global reaction mechanism was used in both the 1D and 2D computations to

predict the measured extinction limit and flame temperature. Photographic images of flames 

undergoing the process of extinction were compared with model calculations. The 

axisymmetric numerical model was validated by comparing flame shapes, temperature 

profiles, and extinction limits with experiments and with the 1D computational results. 

The 2D computations yielded insight into the extinction mode and flame structure. A specific 

maximum heat release rate was introduced to quantify the local flame strength and to 

elucidate the extinction mechanism. The contribution by each term in the energy equation to 

the heat release rate was evaluated to investigate the multi-dimensional structure and 

radiative extinction of the 1-g flames. Two combustion regimes depending on the extinction 

mode were identified. Lateral heat loss effects and multi-dimensional flame and flow 

structure were also found. At low strain rates in 1-g flames (“Regime A”), the flame is

extinguished from the weak outer edge of the flame, which is attributed to a multidimensional flame structure and flow field. At high strain rates, (“Regime B”), the flame

extinction initiates near the flame centerline due to an increased diluent concentration in the 

reaction zone, similar to the extinction mode of 1D flames. These two extinction modes can 

be clearly explained by consideration of the specific maximum heat release rate. 

9
1. Introduction 

The counterflow flame geometry is a convenient configuration for control of the flow field 

strain rate as well as the flame position [1]. From an experimental perspective, the 

counterflow configuration has been used to acquire well-defined data on the structure and 

extinction behavior of flames. Furthermore, the counterflow configuration is amenable to 

mathematical and numerical analysis, and has provided useful insight into local “flamelet” 

structure [1-4]. 

It is common engineering practice to develop and apply models that are as simple as possible,

while maintaining the salient physics. Viewed as a flamelet, the counterflow flame is 

typically modeled as one-dimensional (1D). Experimental observations under microgravity 

conditions have confirmed that counterflow flames are nearly flat [5]. Under normal gravity 

conditions, high strain rate flames are also flat near the flow field centerline as buoyancy is 

dominated by convection. As the strain rate decreases in normal gravity, however, buoyancy 

forces become relatively more important and counterflow flames are clearly curved. Onedimensional flame codes such as OPPDIF [6], which are based on a similarity solution that

neglects buoyancy fail to describe the multi-dimensional structure of low strain rate normal 

gravity flames, nor do they capture the dynamics of a flame during the transition from normal 

gravity to microgravity. To gain physical insight, a numerical method is needed that is 

capable of simulating realistic flames that are multidimensional due to the influence of

buoyancy. The price of using a multidimensional modeling approach is simplification of the 

complex flame chemistry, such that the duration of a computation is practical. Such

approximations are acceptable provided that the most significant observables are adequately 

represented. 

There have been many studies on the structure and extinction of moderately strained 

counterflow nonpremixed flames [1-4, 7]. Relatively few studies, however, have considered 

low strain rate flames. Maruta et al. [5] conducted the first comprehensive extinction 

measurements of very low strain nonpremixed flames in microgravity using the 10 s JAMIC 

drop tower in Japan. In that study, the extinction of methane-air diffusion flames with N2

added to the fuel stream was measured under microgravity conditions. The minimum 

methane concentration required to sustain combustion was found to decrease as the strain rate

decreased until a critical value was observed. As the global strain rate was further reduced,

the required methane concentration increased. This behavior was denoted as a "turning point" 

and was attributed to the enhanced importance of radiative loss in low strain rate flames. In 

terms of fire safety, the turning point agent concentration represents a fundamental limit for 

nonpremixed flames, assuring suppression for all flow conditions. Bundy et al. [8] observed 

similar behavior in low strain rate normal gravity flames. Recently, the extinction behavior of 

ultra low-stretch, buoyant counterflowing nonpremixed flames established in a large Tsujitype burner was investigated [9]. In that study, information about the extinction limit was not 

provided although multi-dimensional effects on flame extinction were discussed. 

The objective of this study was to investigate the structure and extinction characteristics of 

low strain rate, counterflowing methane-air nonpremixed flames in normal gravity, which 

was observed but not fully clarified in previous work [8]. In this study, an axisymmetric (2D) 

computational model including buoyancy effects and radiative heat transfer was developed 

and validated for counterflow nonpremixed flames by comparison with experiments and 1D 

computational results. The flame structure and flow field were investigated using the 

computational results. Extinction limits were also investigated. 

10
2. Experiments 

Experiments for counterflow flames were performed using a burner with an inner burner duct 

diameter (D) of 23.4 mm and a duct separation distance (L) of 15.0 mm. Four 80 mesh/cm 

stainless steel screens were positioned 1.6 (± 0.1) mm from the duct exits and at 1.6 mm 

intervals to impose a top-hat velocity profile. A water-cooled jacket about the top burner duct 

prevented heating of the metal burner and preheating of the supplied reactants. The reactant 

flows were controlled using mass flow controllers that were calibrated using a dry cell 

primary flow meter with an uncertainty of 1 %. The details of the burner system and 

experimental methods were described previously [8]. 

Flame temperature measurements were conducted using thin filament pyrometry [10]. The

details of the measurement and calibration methods are given in Refs. [8, 10]. Still images 

were taken with a 35 mm camera (without optical filters), using a shutter time of 1 s and an 

f 2.8 aperture to compare the steady flame shapes with computations. Images of 

extinguishing flames were also recorded with a digital video camera to clarify the difference 

between flame extinguishing modes at low and high strain rates, while the nitrogen mole 

fraction in the fuel stream (Ca) was varied from 0.79 to the critical nitrogen mole fraction at 

extinction (“the extinction limit; Caex”). Experimental data for the extinction limits as a 

function of a global strain rate for methane-air normal gravity (1-g) counterflow flames was

compared with previous results [8]. 

The global strain rate, ag, was defined as [11]: 

 ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = + ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = +

r O

r F F

O O

O F F

g

V ρ

ρ

L

V V

V ρ

V ρ

L

V 1 2 1 2

a (1) 

where the V denotes the velocity and ρ the density of the reactant streams at the duct 

boundaries, L is the duct separation distance, and the subscripts O and F represent the air and 

fuel streams, respectively. The velocity ratio, Vr, is defined as equal to VO/VF. The global 

strain rate concept allows quantification of a characteristic flame residence time or flow time. 

3. Model formulation and numerical methods 

3.1. Numerical methods for a two-dimensional model 

A transient two-dimensional simulation of the 1-g counterflow diffusion flames formed

between two opposed circular ducts of finite dimensions was developed. A coupled set of 

governing equations with a low Mach number approximation [12] was used in an 

axisymmetric configuration. The governing equations were written as follows: 

+ ∇ ⋅( ) = 0

∂

∂

ρu

t

ρ (2) 

( uu) ( ) u ( ) u ( ) u I ( )g

u T

1 0 3

( ) 2

ρ p μ ρ ρ

t

ρ + − ⎥

⎦

⎤ ⎢

⎣

⎡ + ∇ ⋅ = −∇ + ∇ ⋅ ∇ + ∇ − ∇ ⋅ ∂

∂ (3) 

( ) ( ) , ( 1, 2, ) ( )

ρ Y ρD Y W ω i N

t

ρY

i im i i i

i + ∇ ⋅ = ∇ ⋅ ∇ + = ⋅⋅⋅ ∂

∂

u & (4) 

( ) ( ) r

N

i 1

N

i 1

p pi im i i i i T λ T ρ c D Y T W h ω q

t

T

ρc ⎟ = ∇ ⋅ ∇ + ∇ ⋅∇ − & + & ⎠

⎞ ⎜

⎝

⎛ + ⋅∇

∂

∂ ∑ ∑ = =

0 u (5) 

11
 ∑= ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ =

N

i 1 i

i

W

Y

p0 ρR0T (6) 

where and represent the thermodynamic and hydrodynamic pressure, respectively; the

superscript T is the transpose tensor; I is the identity tensor, u is the velocity vector, T is the

temperature, is the volumetric radiative heat loss rate, ρ is the density; μ is the viscosity; 

is the thermal conductivity, R0 is the universal gas constant, and Yi, hi

0

, 

p0 p1

qr&

λ ωi & , and cpi are the 

mass fraction, heat of formation, production rate, and specific heat of species i, respectively.

Dim is the mixture-averaged diffusion coefficient, which is written in the following form: 

∑ ≠

− = N

j 1 j ji

i

im

X /D

Y

D

1 (7) 

where Dji are the binary diffusion coefficients obtained from the collision integrals [13] and 

Xj are the volume fractions of species j. A detailed description of the transport model can be 

found elsewhere [13]. The buoyancy term was included in the momentum equations to 

capture 1-g flames features. The Soret and Dufour effects were not considered. 

QUICK [14] and a second-order central difference schemes were used to discretize the 

convective and diffusive terms in the governing equations. A predictor-corrector scheme

nearly identical to that of Najm et al. [15] was used for time integration of the governing 

equations. For the time integration of the species and energy equations, a second-order 

Adams–Bashforth scheme was used for the predictor step and a second-order quasi–Crank–

Nicolson scheme was used for the corrector step. A second-order Adams–Bashforth scheme

was used for both the predictor and the corrector steps in the time integration of the 

momentum equation. Efficient algebraic relaxation for the velocity-pressure correction was

performed using the HSMAC method [16], which was modified to consider density variation. 

CHEMKIN-II [17] and TRANFIT [13] were adopted for the calculation of thermodynamic

and transport properties. The numerical code used in this study was validated previously by 

comparing the results with experiments for steady counterflow flames [18]. The detailed 

numerical scheme has been described elsewhere [19]. 

The experimental inlet duct geometry and computational domain are shown in Figure 1. The 

fuel and air issued from the lower and upper ducts, respectively. The inner duct diameter was 

23.4 ± 0.1 mm. The separation distance between the two duct exits and the nozzle rim 

thickness (W) were 15 mm and 3.2 mm, respectively. In a previous study [8], it was shown 

that the flame temperature could affect the inflow boundary temperature of the upper duct at 

low strain rates due to buoyancy effects, and that the flame location could be affected by the 

inflow temperature of the upper duct. The nozzle protrusion length (Lp) is the length of the 

inner rim, just before the location of the fine mesh metal screens in the experimental

apparatus, which were described previously. A sensitivity analysis showed that the value of

Lp affects the character of the flow field and therefore must be included in the model. In the 

experiment and the computation, Lp was equal to 1.2 mm. The computational domain in the 

axial and radial directions was 40 mm and 50 mm, respectively. 

For the computation, a 268 x 90 grid system was used. A uniform grid spacing of 

0.15 mm and 0.3 mm were used in the axial and radial directions, respectively. Because a 

global reaction mechanism was used in the computations, the smallest length scale in these 

laminar flames is the heat release rate region. The heat release zone for conditions considered in 

this study is relatively thick, due to the low strain rate conditions. In Figs. 11 and 13, it is seen that the

width of the heat release rate region is approximately 1.7 mm to 2.0 mm. Generally, at least 10 grid

12
points are needed to resolve the heat release rate region. A grid sensitivity study was conducted for a 

number of limiting flame conditions in which 3 grids sizes (dx) were tested: 0.20 mm, 0.15 mm and 

0.10 mm. The results for dx=0.15 mm were within 0.4% and 0.3% for the maximum temperature and

the extinction limits, respectively, of the results using dx=0.1 mm for simulation of normal gravity

flames (ag = 60 s-1 and 15 s

-1). Consequently, a grid system with dx=0.15 mm was considered

sufficient to resolve the flames considered in this study. 

Symmetry boundary conditions were applied for the velocities, species mass fractions, 

and temperature on the centerline. Uniform axial velocity without a radial velocity 

component was imposed on the fuel and air duct exits. As in the experiment, N2 flowed from

the lower concentric duct (Side A) with an axial velocity of 0.03 m/s [8], creating a curtain 

about the reactants. A slip boundary condition was applied to Side B. Side C was treated as an 

outflow boundary, because the incoming flow mainly escaped from the upper boundary in the 

normal gravity flame computations. No-slip and Neumann boundary conditions were applied 

for the velocity and its gradient, respectively, on the burner walls. The inflow temperatures 

for the fuel, air and curtain streams were set to be 298 K, and the wall temperature was also 

taken to be 298 K. The air stream was composed of 21% oxygen and 79% nitrogen by 

volume, and the fuel stream was composed of methane diluted by nitrogen. The ambient

curtain flow was set to pure nitrogen to prevent secondary combustion of fuel. 

The computational time required to obtain the initial steady flame was typically 24 CPU 

hours on a 2 GHz PC. 

3.2. Chemical reaction mechanism and radiation model 

A 3-step irreversible reaction mechanism [20] for methane oxidation was used in the 2D 

computation, because the mechanism was relatively less stiff compared to other global

mechanisms that were considered. The pre-exponential factor and the activation energy in the 

reaction mechanism were slightly adjusted, so that the 2D numerical code reasonably 

predicted the experimentally determined extinction limit (see Fig. 5) and the flame

temperature profile of a moderately strained methane-air nonpremixed flame (ag=40 s-1). A

moderately strained flame was selected so that radiative heat losses would not significantly 

influence the calculation results. The reaction mechanism used in this study was: 

CH4 +1.5O2 →CO + 2H2O (8)

5O2 CO2 CO + 0. → (9)

2 5O2 CO →CO + 0. (10)

and the associated reaction rates were: 

( ) 0.8

2

0.7

4

11.68

4 − d[CH ]/ dt =10 exp − 23,500 /T [CH ] [O ] (11) 

( ) 0.25

2

0.5

2

12.35 1.0 − d[CO]/ dt =10 exp −19,200/T [CO] [H O] [O ] (12) 

( ) 0.25

2

0.5

2

12.50 1.0

2 − d[CO ]/ dt =10 exp − 20,500 /T [CO] [H O] [O ] (13) 

with the reactions rates in units of kmole/m3

-s. 

Experiments (Figs. 7 and 9) have shown that the near-extinction N2 diluted flames considered

here are completely non-luminous (band radiation only). This may be due to the relatively 

cool flame temperatures, precluding the formation of soot particles that otherwise occur in 

undiluted methane flames. The assumption of no thermal radiation by soot in the N2 diluted 

flames was thereby warranted. Therefore, an optically-thin radiation sub-model [21] was

implemented, which considered heat loss from the participating gaseous species CH4, CO2, 

H2O and CO. 

13
The volumetric rate of radiative heat loss in the energy equation was written as: 

( ) 4 4 = − 4 T − T∞ qr σK p & (14) 

where σ is the Stefan-Boltzmann constant, T and T∞ are the local flame and the ambient 

temperature, respectively. The term Kp is the Planck mean absorption coefficient of the 

mixture, which is expressed as: 

Kp =PCH4KCH4 + PCO2KCO2 + PH2O KH2O + PCO KCO (15) 

where and denote the partial pressure and the Planck mean absorption coefficient of 

species i, respectively. Data for Kp were taken from Ref. [21]. 

Pi Ki

3.3. One-dimensional model computation 

For comparison with the two-dimensional (2D) model results, a previously developed onedimensional (1D) flamelet code [6] was used to compute the near-extinction flame structure 

and the agent extinction requirements in low and moderate strain rate diluted methane-air 

diffusion flames. The numerical code solves the conservation equations using detailed models

of molecular transport. An optically thin radiation model [21], the same as in the 2D model, 

was also adopted in the 1D code. Both, the detailed chemistry of GRI-3.0 [22] and the 3-step 

global reaction mechanism (Eqs. 8-13) were used in the 1D computations. The counterflow 

code does not include a buoyancy term in the momentum equation. In this context, the 

simulations represent zero gravity conditions, but are also a good representation of moderate 

and high strain rate 1-g conditions. The inlet boundary conditions at the burner ducts are the 

same as the 2D computation. 

4. Results and discussion 

4.1. Validation of the computational results 

The computed flame shape and temperature profiles were compared with measurements to

validate the 1D and 2D models. Figure 2 compares photographs of the thin non-luminous 1-g 

flames and computed flame shapes using the 2D model for the fuel stream diluted with N2, 

such that the flames were near extinction. The high heat release rate region represents the 

high temperature zone and the peak heat release rate can be considered to be near the flame 

surface. The computed flame shape and the flame surface location were similar to those 

shown in the Fig. 2 photographs. The photographs in Fig. 2 show that the flame size, the 

shape, and the flame surface location of the 1-g flames were reasonably well predicted by the 

2D computations for a variety of global strain rate and velocity ratio conditions (see Eq. 1). 

Figure 3 compares the temperature profiles computed by the 1D and 2D models with 

measurements [10] of methane-air nonpremixed counterflow flames for ag = 20 s-1 in both 

microgravity and 1-g. The flames were highly diluted and at near-extinction conditions with 

Ca = 0.81. For the 0-g conditions, the 1D and 2D computations, using the 3-step mechanism,

resulted in nearly identical calculation results for the temperature profiles. For both the 

microgravity and the 1-g flames, the temperature profiles computed by the 2D model were in 

good agreement with previous measurements obtained using thin filament pyrometry [10].

The flame locations were also reasonably predicted by the 2D computation. The 1D model

computations (which neglect buoyancy), using either the 3-step reaction mechanism or

14
detailed chemistry, could not accurately predict the location of the 1-g flame, although the 

flame location and temperature profile in microgravity were well-predicted. The measured 

and computed temperature profiles show that the flame was broader in microgravity than in 

1-g, and that the location of the flame shifts towards the center of the flow field in

microgravity. 

Figure 4 compares the axial velocity profiles computed by the 1D and 2D models for the 0-g 

flame. The 2D computation for the 1-g flame is also shown. The velocity profile of the 

microgravity flame computed by the 2D model was nearly identical to that computed by the 

1D model, providing confidence that the current 2D model is reasonable. There was little

difference in the 1D results, using the full or simplified chemistry model in predicting the

velocity profiles. It should be noted, however, that the 1D model could not simulate the low 

strain rate 1-g flame structure reasonably, because it does not consider buoyancy effects. The 

2D model developed here provides a more complete description of the flame physics as seen 

in the flame structure, including the flame temperature, location, and velocity profile at low 

strain rates. As expected, the 1-g velocity profile was completely different than the 0-g profile, 

with the stagnation point shifting almost 4 mm towards the air duct, similar to the shift of the 

temperature maxima. 

4.2. Extinction limits and characteristics 

The extinction limits were determined by increasing the agent (N2) flow and simultaneously 

decreasing the fuel flow, while maintaining a constant global strain rate, until a critical agent 

volume fraction in the fuel stream was attained and the flame was experimentally observed to 

extinguish or the calculation did not converge to a solution. This efficiency of this process was

optimized by beginning the calculation or experiment at near-extinction, but stable flame conditions.

The rate of change of reactants was sufficiently slow such that the extinction conditions were not 

affected. A comparison between the measured [8] and computed extinction limits is shown in

Figure 5. The 1D model cannot reasonably predict the extinction limits of 1-g flames at low 

strain rates, when buoyancy effects become dominant, since the buoyancy term was not 

included in this model. The 1D model yielded nearly the same result using either the 3 step 

chemistry or GRI-3.0. The results of the 2D model for the 1-g flames are in excellent 

agreement with the measurements at both low and high strain rates. The results also show that

the extinction limit of microgravity flames (1D) is higher than that of 1-g flames at low strain

rates (ag < 35 s

-1), whereas those of both microgravity and 1-g flames are nearly identical at 

high strain rates. In addition, the turning point (i.e., maxima of Ca) in the 1-g flame shifts

towards much higher strain rates than in microgravity. 

The turning point in the microgravity extinction limit curve is a well-known phenomenon, 

which is attributed to radiative heat loss effects [5, 23]. The turning point in 1-g flames differs 

phenomenologically from the microgravity case. The turning point of 1-g flames is not only 

due to radiative heat loss effects, but is also attributed to multi-dimensional flame extinction

effects for the geometry considered in this study. The 1-g flames can be classified into two 

extinction modes, denoted here as Regimes A and B. It is possible to differentiate these two 

modes based on the global strain rate, at a value of about ag ≈ 35 s-1. The nature of these two 

regimes is described below.

Figures 6 and 7 show the computed results and experimental photos of the temporal evolution 

of 1-g flame extinction at a strain rate of ag = 60 s-1. “Regime B” is characterized by the 

extinction mechanism, such that an increasing diluent (nitrogen) concentration in the reaction 

15
zone extinguishes the flame everywhere, over the entire flowfield. The nitrogen mole 

fraction in the fuel stream of the 1-g flame (Ca) was increased at time equal zero from an 

initial steady value of 0.79 to 0.825, its value at the extinction limit. The left hand side of the 

computational results shows the computed flame temperature, while the right hand side 

shows the computed local heat release rate. The right hand side of the figures also shows the

flow streamlines. Figure 6 shows that the extinction of the 1-g flame initiates near the

centerline as the increased nitrogen volume fraction in the fuel stream reaches the flame 

surface. The zone around the centerline becomes weak as the heat release rate (HRR) 

decreases and the entire flame abruptly extinguishes. The photographs of the flame extinction 

process shown in Fig. 7 substantiate these features. The radiant intensity of the flame appears

to decrease as Ca increases, until the flame suddenly extinguishes. This flame extinction 

mechanism is typical of a moderately or highly strained 1D counterflow flame. 

The computed results and photographs for the temporal evolution of 1-g flame extinction at a 

low strain rate (ag=20 s-1) are shown in Figs. 8 and 9, respectively. The flame, characterized 

by what is denoted here as multi-dimensional extinction, belongs to “Regime A”. Note that 

the extinction of the low strain rate 1-g flame initiates at the outer flame edge. That is, the 

flame shrinks radially from the outer flame edge towards the center of the flowfield. 

Photographs of the actual flame extinction process (shown in Fig. 9) confirm the calculated 

flame behavior. This unique extinction mechanism is different than that of the high strain

flame shown in Figs. 6 and 7. It is evident that the 1-g flames have a multi-dimensional flame 

structure. Thus, at low strain rates, a 1D model cannot reasonably predict the structure, nor 

the extinction of 1-g flames. 

To investigate flame extinction characteristics in more detail, a “specific maximum heat

release rate” (SMHRR) is introduced as defined by Sung et al. [24] as the maximum heat

release rate along a normal through the flame divided by the local strain rate (al). The 

parameter SMHRR is a measure of the flame strength, and can be used to compare the local 

flame strength or stability along a flame surface. A smaller value of SMHRR implies that the 

flame is weak or easier to extinguish. In obtaining the SMHRR, a local strain rate was 

defined as the normal velocity gradient to the flame surface at the maximum heat release rate 

location, consistent with the definition given in Ref. [25]. Table 1 compares the local flame 

strength at selected three locations A, B and C. The location A represents the centerline 

position. Location C represents the outer flame edge region, which has one-half the value of 

the heat release rate of the centerline. Location B represents a radial position midway

between the centerline and the outer flame edge. Table 1 shows that the calculated local flame 

strength at the outer edge is weaker than at the centerline for the case ag = 20 s-1, whereas this 

is not the case for the case ag = 60 s-1. For ag = 60 s

-1, both the local strain rate and the heat

release rate at the outer flame edge are simultaneously lower, as compared to the centerline.

However, for the case ag = 20 s-1, the local strain rate at the outer flame edge is not lower than 

at the centerline, even though the local heat release rate is lower than at the centerline. Thus, 

the SMHRR at the outer flame edge becomes small, and flame extinction initiates from the 

weak outer flame edge at low strain rates (ag=20 s-1). These calculation results are consistent

with experimental observations as shown, for example, in Fig. 9. 

4.3. Multi-dimensional flame structure 

The left hand side of Fig. 10 shows the distributions of the 2D modeling results for methane 

and oxygen mole fractions in 1-g counterflow flames. Flow streamlines are presented on the 

16
right-hand side of each figure. The white background represents regions, in which the volume

fractions of fuel and oxygen are below 0.05. The white gap between the fuel and oxygen 

roughly corresponds to the heat release rate region. In the figures, A, B and C are located at 

the same locations defined in Table 1. For ag=60 s

-1, the fuel encloses the oxygen distribution 

region. However, for low strain rates (ag=20 s-1), the fuel does not enclose the entire oxygen 

distribution region, and the fuel is not located beyond the burner duct (r>R). The two 1-g 

flames (ag=20 s-1 and 60 s-1) exhibit a curved multi-dimensional flame structure. Table 1 

shows for the case ag=20 s-1, that the local strain rate at location C (where r/R=1.1; see Fig. 

10) is much higher than at location A (r/R=0). Thus, the low strain rate flame becomes weak 

at the outer flame edge. For this case (ag=20 s-1), the fuel concentration drops rapidly beyond 

the flame edge near location C. The streamlines become dense in the downstream region 

because there is no heat release there. The streamline distributions are associated with a high

local strain rate around location C. 

To investigate the local 1-g flame in more detail, the flame structure and the local flow field 

normal to the flame surface at locations, A, B, and C (shown in Fig. 10) are considered in 

Figs. 11 and 12 (for ag=60 s-1), and Figs. 13 and 14 (for ag=20 s-1). For each condition, the 

radial location for locations A, B, and C are specified in Table 1. The abscissa represents the

normal distance from the flame surface at each selected location. The location x = 0 mm 

refers to the flame surface location defined by the maximum heat release rate. The local strain

rate was defined using the flame stretch concept [26], rather than the oxidizer side maximum 

velocity gradient [25] because the oxidizer side maximum velocity gradient is not easily

defined at all flame locations. In Fig. 11, the flame structure, as represented by the local 

species mole fractions and temperature, at location B, is nearly identical to that on the 

centerline at location A. The flame structure at location C is different than at locations A and 

B. The temperature gradient at location C is smaller than at the centerline, and the fuel 

distribution is bell-shaped, which is not typical of 1D flames. On the other hand, the oxidizer 

distribution in the mixing layer is similar to that at the centerline. The normal velocity 

component to the flame surface and the local strain rate are shown in Fig. 12. The trends of

the normal velocity component and the local strain rate profiles relative to the flame surface

are related to the temperature profiles shown in Fig. 11, and the flow field at location C is 

different from those at locations A and B as seen in Fig. 12. That is, the local strain rate at

location C is much lower than at the centerline. In a typical nonpremixed flame, as the strain 

rate increases, the heat release rate also increases, because the amount of fuel and oxidizer 

supplied to the flame surface increases. Figure 11 shows that at the flame edge (location C),

the fuel concentration gradient in the mixing layer is smaller than at the center, and the heat 

release rate is relatively smaller than the centerline. In total, the flame strength at the outer

flame edge as measured by the SMHRR is similar to that at the centerline for ag=60 s-1. 

The case ag = 20 s-1 is shown in Figs. 13 and 14. The flame structure and the local flow field

at locations A and B are similar to those at the centerline for ag=60 s-1. The fuel concentration

gradient in the mixing layer is similar across the flame (locations A, B, and C), even though 

the local strain rate is much larger at the flame edge (location C). This implies that the fuel 

supplied to the flame surface at the flame edge is relatively small, and does not follow the 

profile of a typical nonpremixed flame. Figure 13 shows that the heat release rate at the 

flame edge is relatively small, and Table 1 shows that the SMHRR at the flame edge is weak

compared to the centerline. Consequently, the 1-g flame is extinguished from the relatively 

weak flame edge at low strain rates, while the region near the centerline is strong enough to 

sustain combustion. This also explains why the extinction limit of the multi-dimensional 1-g 

flame is somewhat lower than that of a typical one-dimensional 0-g flame. 

17
It is useful to consider the different processes that contribute to the heat release rate and flame

stability. For this reason, the relative contribution by each term in the energy equation is

considered. The energy equation can be expressed as follows: 

∑ ∑ ∑ = = =

+ − ∂

∂

∂

∂

+

∂

∂

∂

∂

+

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

∂

∂ ⎟ +

⎠

⎞ ⎜

⎝

⎛

∂

∂

∂

∂

+

∂

∂ − ∂

∂ ⎟ = − ⎠

⎞ ⎜

⎝

⎛

∂

∂

N

i 1

i

0

r i i

N

i 1

i

pi im

N

i 1

i

pi im

p p p

q W h ω

r

T

r

Y

ρc D

x

T

x

Y

ρc D

r

T

rλ

x r r

T λ

r x

T

ρc v

x

T

ρc u

t

T

ρc

& &

1

0 Cx Cr Dx Dr

Mx Mr Rs Cs

 (16) 

where Cx and Cr are the axial and radial convection terms, Dx and Dr are the axial and radial 

diffusion terms, Mx and Mr are the axial and radial inter-diffusion terms, Rs is the radiation

heat loss rate, and Cs is the chemical heat release rate. Each term in the energy equation has 

units of energy per unit time per unit volume. 

Figure 15 shows the computed fractional contribution by each term in the energy equation

normalized by the heat release source term (Cs in Eq. 16) in the 0-g flames as a function of 

the global strain rate. The relative contribution of each term was obtained by integrating over 

the flame surface at the centerline. The contribution of the radial convection and interdiffusion terms are also negligible. Axial conduction and diffusion are the largest heat loss 

mechanisms. The contributions of the axial conduction and diffusion terms to the heat loss 

are of similar importance in both 0-g and 1-g flames. Thus, radiation and radial conduction 

mainly contribute to heat loss from the centerline. The relative contribution of the radiative 

heat loss increases exponentially as the global strain rate decreases in 0-g flames. In 1-g 

flames, however, the relative contribution of the radiative heat loss increases linearly as the

global strain rate decreases and becomes constant at low strain rates (ag < 20 s-1). Radial

conduction plays an import role in the heat loss mechanism at low strain rates and its 

contribution becomes similar to that of the radiative heat loss. The reason that the radial 

conductive heat loss (“lateral heat loss”) becomes important at low strain rates (ag < 20 s-1) is 

that the flame thickness in the radial direction is small compared with high strain rate flames. 

As shown in Table 2, the local strain rate in the 1-g flame is much higher than the 0-g flame 

due to buoyancy effects for low global strain rates (ag ≤ 30 s

-1; Regime A in Fig. 5), and the 

thickness of a 1-g flame is smaller than the 0-g flame. This is also seen in Fig.3. Figure 16 

shows that a detailed accounting of the relative contribution of the radiative heat loss from 

the 1-g flame shows that it is smaller than in the 0-g flame. Finally, this small radiative heat 

loss contribution leads to a maximum temperature at the centerline of the low strain (ag = 20 

s

-1) 1-g flame that is higher than in the analogous 0-g flame. This is consistent with previous 

studies that have indicated that radiative heat loss is an important heat loss mechanism in low 

strain rate microgravity flames [5]. 

5. Summary of results 

A fundamental study of the structure and extinction of normal gravity nonpremixed methane18
air flames was conducted using computations and experiments. An axisymmetric flame was

computed using a global 3-step chemical reaction mechanism. Computations of normal 

gravity flames showed good agreement with experimental observations of flame curvature, 

flame location, temperature profile [10] and extinction limits [8]. An axisymmetric 

formulation was shown to be very important in explaining the multi-dimensional structure of

1-g counterflow nonpremixed flames, which cannot be understood using a one-dimensional

model. 

The 1-g flame was classified in two regimes. The specific maximum heat release rate 

was introduced to quantify the local flame strength and to elucidate the extinction mechanism. 

For relatively high strain rates (Regime B), an increased diluent concentration in the reaction

zone causes extinction, and flame extinction initiates near the flame centerline. In this regime, 

the extinction limit is nearly identical to 1D flame results. For relatively low strain rates 

(Regime A), a unique multi-dimensional flame structure and corresponding extinction

mechanism occur. The fuel distribution and the buoyancy effects make the local strain rate at

the outer edge of low strain 1-g flames much higher than at the centerline. The lateral heat

loss effects were identified from an analysis of the overall fractional contribution by each 

term in the energy equation to the heat release. The relative contribution of the radiative heat 

loss from a 1-g flame is small compared with the 0-g 1D flame results due to a thin flame 

thickness caused by a relatively high local strain rate. 

At low strain rates in Regime A, the coupled effects of the heat release and the high local 

strain rate weaken the flame at the outer edge of the 1-g flame. Flame extinction initiates

from the weak outer flame edge even though the region near the centerline is still strong 

enough to sustain combustion. The extinction limit of the 1-g flame is much lower than that 

of the 0-g 1D flame at low strain rates due to the multi-dimensional flame extinction 

mechanism at Regime A. The extinction characteristics of 1-g flames can be explained 

through consideration of the specific local maximum heat release rate. 

Acknowledgements 

This research was supported by the NASA Microgravity Research Division through Contract 

No. NNC04CB07C with Dr. Sandra L. Olsson serving as Technical Monitor. 

References

 [1] Tsuji H 1982 Prog. Energy. Combust. Sci. 8 93-119 

 [2] Puri I K and Seshadri K 1986 Combust. Flame 65 137-150 

 [3] Peters N and Kee R J 1987 Combust. Flame 68 17-29 

[4] Seshadri K and Kee R J 1988 Combust. Flame 73 23-44 

[5] Maruta K, Yoshida M, Guo H, Ju Y and Niioka T 1998 Combust. Flame 112 181-187 

[6] Lutz A E, Kee R J, Grcar J F and Rupley, F M 1997 Sandia Report SAND96-8243 

[7] Sohrab S H, Liñan A and Williams F A 1982 Combust. Sci. Technol. 27 143 

[8] Bundy M, Hamins A, Lee K Y 2003 Combust. Flame 133 299-310 

[9] Han B, Ibarreta A F, Sung C J and T’ien J S 2004 Proc. Combust. Inst. 30 527-535.

[10] Bundy M, Park J, Oh C B, Hamins A, “Extinction and Temperature Measurements in 

Nonpremixed Methane Flames in Microgravity,” in preparation, 2005. 

[11] Seshadri K and Williams F A 1978 Int. J. Heat Mass Transfer 21 251-253 

[12] McMrutry P A, Jou W H, Riley J J and Metcalfe R W 1986 AIAA J. 28 962-970 

19
[13] Kee R J, Dixon-Lewis D, Warnatz J, Coltrin M E and Miller J A 1986 Sandia Report 

 SAND86-8246 

[14] Leonard B P 1979 Comput. Meth. Appl. Mech. Eng. 19 59-98 

[15] Najm H N and Wyckoff P S 1997 Combust. Flame 110 92-112 

[16] Hirt C W and Cook J L 1972 J. Comput. Phys. 10 324-340 

[17] Kee R J, Rupley F M and Miller J A 1989 Sandia Report SAND89-8009B 

[18] Oh C B, Lee C E and Park J 2004 Combust. Flame 138 225-241 

[19] Oh C B, Hwang C H and Lee C E 2006 submitted to J. Comput. Appl. Math

[20] Dryer F L and Glassman I 1972 Proc. Combust. Inst. 14 987-1003 

[21] Ju Y, Guo H, Maruta K and Liu F 1997 J. Fluid Mech. 342 315-334 

[22] Bowman C T, Hanson R K, Davidson D F, Gardiner W C, Lissianski V, Smith G. P, 

 Golden D M, Frenklach M and Goldenburg M 1999 http://www.me.berkeley.edu/ gri_mech/

[23] T’ien J S 1986 Combust. Flame 65 31-34 

[24] Sung C J, Liu J B and Law C K 1995 Combust. Flame 102 481-492 

[25] Chelliah H K, Law C K, Ueda T, Smooke M D and Williams F A 1990 Proc. Combust. 

 Inst. 23 503-511 

[26] Poinsot T and Veynante D 2001 Theoretical and Numerical Combustion (Edwards) 

p.50 

20
List of Tables 

Table 1. Calculated (2D) specific maximum heat release rate as a function of radial location 

in 1-g flames. 

Table 2. Local strain rate (al) and flame thickness of 0-g and 1-g flames at select global strain 

rate conditions. 

List of Figures 

Figure 1. Schematic drawing of the burner geometry and computational domain. 

Figure 2. Photos (upper) and computed flame shapes (lower) of 1-g counterflow 

nonpremixed flame for (a) ag=15 s-1 and Vr=4, (b) ag=20 s-1 and Vr=3, (c) ag=30 s-1 and Vr=3, 

and (d) ag=50 s-1 and Vr=1. Ca is 0.79 for all flames. Temperature is shown on left-hand side 

of the computed results and the heat release rate and stream lines are shown on the right-hand 

side of the computed results. 

Figure 3. Measured and computed temperature profiles for ag=20 s-1 with Ca=0.81 and Vr=1. 

Figure 4. Computed axial velocity profiles at the centerline for ag=20 s-1 with Ca=0.81 and 

Vr=1. 

Figure 5. Measured and computed extinction limits and combustion regimes of 1-g and 0-g 

counterflow nonpremixed flames. 

Figure 6. Computed extinction process of 1-g counterflow nonpremixed flame for ag=60 s-1

and Vr=1 while Ca increases from 0.79 to 0.825 at time equal to 0 s. The flame temperature is 

shown on the left-hand side, and heat release rate and stream lines are shown on the righthand side. (a) 0 s; (b) 0.06 s; (c) 0.09 s. 

Figure 7. Photos of extinction process of a 1-g counterflow nonpremixed flame for ag=60 s-1

and Vr=1, with Ca increasing from 0.79 to 0.825. 

Figure 8. Computed extinction process of 1-g counterflow nonpremixed flame for ag=20 s-1

and Vr=3 while Ca increases from 0.79 to 0.825 at time equal to 0 s. Flame temperature is 

shown on the left-hand side, and heat release rate and stream lines are shown on the righthand side. (a) 0 s; (b) 0.30 s; (c) 0.34 s. 

Figure 9. Photos of extinction process of a 1-g counterflow nonpremixed flame for ag=20 s-1

and Vr=3, with Ca increasing from 0.79 to 0.825. 

Figure 10. Computed fuel (CH4) and oxidizer (O2) volume fraction distributions in a 1-g 

counterflow nonpremixed flame for (a) ag=60 s-1, Vr=1 and Ca=0.79 and (b) ag=20 s-1, Vr=3 

and Ca=0.79. Fuel flows from the lower duct and oxidizer flows from the upper duct. Stream

lines are shown on the right-hand side. 

Figure 11. Computed flame temperature, reactant (CH4 and O2) concentrations, and heat 

release rate in coordinates normal to the flame surface at three locations for a 1-g counterflow 

nonpremixed flame with ag=60 s-1 and Vr=1. (symbol : location A; line : location B; dashed

line : location C; See Fig. 10 for locations). 

Figure 12. Computed normal component of velocity to the flame surface and local strain rate 

in the coordinates normal to the flame surface at three locations of 1-g counterflow 

nonpremixed flame for ag=60 s-1 and Vr=1. (symbol : location A; line : location B; dashed 

line : location C; See Fig. 10 for locations). 

Figure 13. Computed flame temperature, reactants (CH4 and O2) concentrations, and heat 

release rate in the coordinates normal to the flame surface at three locations of 1-g 

counterflow nonpremixed flame for ag=20 s-1 and Vr=3. (symbol : location A; line : location 

B; dashed line : location C; See Fig. 10 for locations). 

Figure 14. Computed normal component of velocity to the flame surface and local strain rate 

in the coordinates normal to the flame surface at three locations of 1-g counterflow 

nonpremixed flame for ag=20 s-1 and Vr=3. (symbol : location A; line : location B; dashed 

line : location C; See Fig. 10 for locations). 

21
Figure 15. Computed fractional contribution by each term in the energy equation (Eq. 16) to 

the heat release of 0-g and 1-g flames with Ca=0.79 as a function of the global strain rate 

along the flow field centerline. 

22
Tables

Table 1. Calculated (2D) specific maximum heat release rate as a function of radial location 

in 1-g flames. 

Location r/R * Maximum HRR

(J/cm3

-s) 

Local strain rate 

(s-1) 

SMHRR 

 (J/cm3) 

ag=20 s-1, Vr=3, Ca=0.79 

A 0 156 60 2.6 

B 0.5 150 61 2.5

C ** 1.1 73 95 0.8 

ag=60 s-1, Vr=1, Ca=0.79 

A 0 229 105 2.2 

B 0.8 228 107 2.1

C ** 1.6 116 40 2.7 

 

 * where R is the burner inner radius and r/R=0 is at the center of the flow field. 

 ** outer flame edge region has about half the value of the centerline heat release rate. 

Table 2. Local strain rate (al) and flame thickness of 0-g and 1-g flames at select global strain 

rate conditions. 

Global strain rate, ag (s-1) 

15 20 40 60

Local strain rate, 

 al (s-1) 27 35 67 97

0-g 

flame Flame thickness, 

wf* (mm) 10.4 9.1 6.7 5.6

Local strain rate, 

 al (s-1) 55 60 73 105

1-g 

flame Flame thickness, 

wf* (mm) 8.0 7.5 6.5 5.3

 * The flame width, wf , is defined as the full width at 5 % of the maximum flame temperature.

23
Figure 1. Schematic drawing of the burner geometry and computational domain. 

Side B

g 

Computational 

domain 

x R=11.7 mm

Side C

0

r 

Fuel 

(CH4+N2)

Air 

L=15 mm

Lp=1.2 mm

W=3.2 mm 

Side A

Flame 

24
 (a) ag=15 s-1 (b) ag=20 s-1 (c) ag=30 s-1 (d) ag=50 s-1

1660

1570

1480

1390

1300

1210

1120

1030

940

850

760

670

580

490

400

LHS : Temp. (K)

-30 -20 -10 0 10 20 30

-10

0

10

20 x (mm)

r (mm) -30 -20 -10 0 10 20 30r (mm) -30 -20 -10 0 10 20 30r (mm) -30 -20 -10 0 10 20 30r (mm)

15mm

23.4mm

144

134

124

114

104

94

84

74

64

54

44

34

24

14

4

RHS :

HRR (W/cm3)

Figure 2. Photos (upper) and computed flame shapes (lower) of 1-g counterflow 

nonpremixed flame for (a) ag=15 s-1 and Vr=4, (b) ag=20 s-1 and Vr=3, (c) ag=30 s

-1

 and Vr=3, 

and (d) ag=50 s-1 and Vr=1. Ca is 0.79 for all flames. Temperature is shown on left-hand side 

of the computed results and the heat release rate and streamlines are shown on the right-hand 

side of the computed results. 

25
0.0 2.0 4.0 6.0 8.0 10.0 12.0 14.0

Distance from Air Duct (mm)

300

600

900

1200

1500

1800

2100

Temperature (

K)

1-g

0-g

2.2s drop tower results [10]

1-g experiment [10]

2D computation

1D computation, GRI-3.0

1D computation, 3-step

Figure 3. Measured and computed temperature profiles for ag=20 s-1 with Ca=0.81 and Vr=1. 

26
0.0 2.0 4.0 6.0 8.0 10.0 12.0 14.0

Distance from Air Duct (mm)

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0.10

0.15

Axial Velocity at Ce

nterline (m/s)

1-g

0-g

2D computation

1D computation, GRI-3.0

1D computation, 3-step

Figure 4. Computed axial velocity profiles at the centerline for ag=20 s-1 with Ca=0.81 and 

Vr=1. 

27
0.70

0.75

0.80

0.85

0.90

0 20 40 60 80

Global Strain Rate (s

-1)

Critic

al Mole Fra

ction at E

xtinction, C

aex

.

100

1-g experiments [8], d=23.4 mm

1-g computations, 2D, d=23.4 mm

0-g computations, 1D, 3-step

0-g computations, 1D, GRI-3.0

Regime A Regime B

Figure 5. Measured and computed extinction limits and combustion regimes of 1-g and 0-g 

counterflow nonpremixed flames.

28
Figure 6. Computed extinction process of 1-g counterflow nonpremixed flame for ag=60 s-1

and Vr=1 while Ca increases from 0.79 to 0.825 at time equal to 0 s. The flame temperature is 

shown on the left-hand side, and heat release rate and stream lines are shown on the righthand side. (a) 0 s; (b) 0.06 s; (c) 0.09 s. 

-20 -10 0 10 20

-10

0

10

20

x (mm)

r (mm)

-20 -10 0 10 20

-10

0

10

20

x (mm)

-20 -10 0 10 20

-10

0

10

20

x (mm)

(a) 

(b) 

(c) 

206

192

178

164

150

136

122

108

94

80

66

52

38

24

10

RHS :

HRR (W/cm3)

1660

1570

1480

1390

1300

1210

1120

1030

940

850

760

670

580

490

400

LHS :

Temp. (K)

29
(a) 

(b) 

(c) 

Figure 7. Photos of extinction process of a 1-g counterflow nonpremixed flame for ag=60 s-1

and Vr=1, with Ca increasing from 0.79 to 0.825. 

30
Figure 8. Computed extinction process of 1-g counterflow nonpremixed flame for ag=20 s-1

and Vr=3 while Ca increases from 0.79 to 0.825 at time equal to 0 s. Flame temperature is 

shown on the left-hand side, and heat release rate and stream lines are shown on the righthand side. (a) 0 s; (b) 0.30 s; (c) 0.34 s. 

-20 -10 0 10 20

-10

0

10

20 x (

m

m)

-20 -10 0 10 20

-10

0

10

20

x (mm)

-20 -10 0 10 20

-10

0

10

20 x (

m

m)

r (mm)

1660

1570

1480

1390

1300

1210

1120

1030

940

850

760

670

580

490

400

LHS :

Temp. (K)

144

134

124

114

104

94

84

74

64

54

44

34

24

14

4

RHS :

HRR (W/cm3

)

(a) 

(b) 

(c) 

31
(a) 

(b) 

(c) 

Figure 9. Photos of extinction process of a 1-g counterflow nonpremixed flame for ag=20 s-1

and Vr=3, with Ca increasing from 0.79 to 0.825. 

32
(a) ag=60 s-1, (b) ag=20 s-1

Figure 10. Computed fuel (CH4) and oxidizer (O2) volume fraction distributions in a 1-g 

counterflow nonpremixed flame for (a) ag=60 s-1, Vr=1 and Ca=0.79 and (b) ag=20 s-1, Vr=3 

and Ca=0.79. Fuel flows from the lower duct and oxidizer flows from the upper duct. 

Streamlines are shown on the right-hand side. 

-30 -20 -10 0 10 20 30

-10

0

10

20

x (mm)

r (mm)

B A

C

0.20

0.18

0.17

0.16

0.14

0.13

0.12

0.10

0.09

0.07

0.06

Xi

-30 -20 -10 0 10 20 30

-10

0

10

20

x (mm)

r (mm)

B A

C

33
-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

0

400

800

1200

1600

2000

Te

m

perature, T (K)

0.00

0.10

0.20

0.30

0.40

Mole Fraction

0

50

100

150

200

250

300

Heat Release Rate, H

R

R (W/cm3)

Fuel Side Air Side HRRmax

CH4

O2

T

HRR

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

0

400

800

1200

1600

2000

Te

m

perature, T (K)

0.00

0.10

0.20

0.30

0.40

Mole Fraction

0

50

100

150

200

250

300

Heat Release Rate, H

R

R (W/cm3)

Fuel Side Air Side HRRmax

CH4

O2

T

HRR

Normal Distance from the Flame Surface (mm)

Figure 11. Computed flame temperature, reactant (CH4 and O2) concentrations, and heat 

release rate in coordinates normal to the flame surface at three locations for a 1-g counterflow 

nonpremixed flame with ag=60 s-1 and Vr=1. ( • : location A; ⎯ : location B; --- : location C;

see Fig. 10 for location of points A, B, and C). 

34
-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

Normal C

omponent of Velocity (m/s)

Fuel Side HRRMax Air Side

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

Normal C

omponent of Velocity (m/s)

Fuel Side HRRMax Air Side

Normal Distance from the Flame Surface (mm)

(a) normal component of velocity to the flame surface 

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-150

-100

-50

0

50

100

Strain 

Rate (1/s)

Fuel Side HRRMax Air Side

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-150

-100

-50

0

50

100

Strain 

Rate (1/s)

Fuel Side HRRMax Air Side

Normal Distance from the Flame Surface (mm)

(b) local strain rate 

Figure 12. Computed normal component of velocity to the flame surface and local strain rate 

in the coordinates normal to the flame surface at three locations of 1-g counterflow 

nonpremixed flame for ag=60 s-1 and Vr=1. (• : location A, ⎯ : location B, --- : location C; 

see Fig. 10 for location of points A, B, and C). 

35
-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

0

400

800

1200

1600

2000

Te

m

perature, T (K)

0.00

0.10

0.20

0.30

0.40

Mole Fraction

0

40

80

120

160

200

Heat Release Rate, H

R

R (W/cm3)

Fuel Side Air Side HRRmax

CH4

O2

T

HRR

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

0

400

800

1200

1600

2000

Te

m

perature, T (K)

0.00

0.10

0.20

0.30

0.40

Mole Fraction

0

40

80

120

160

200

Heat Release Rate, H

R

R (W/cm3)

Fuel Side Air Side HRRmax

CH4

O2

T

HRR

Normal Distance from the Flame Surface (mm)

Figure 13. Computed flame temperature, reactants (CH4 and O2) concentrations, and heat 

release rate in the coordinates normal to the flame surface at three locations of 1-g 

counterflow nonpremixed flame for ag=20 s-1 and Vr=3. (• : location A, ⎯ : location B, --- : 

location C; see Fig. 10 for location of points A, B, and C). 

36
-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-0.20

-0.10

0.00

0.10

0.20

Normal Component of Velocity (m/s)

Fuel Side Air Side HRRMax

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-0.20

-0.10

0.00

0.10

0.20

Normal Component of Velocity (m/s)

Fuel Side Air Side HRRMax

Normal Distance from the Flame Surface (mm)

(a) normal component of velocity to the flame surface 

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-120

-80

-40

0

40

80

Strain 

Rate (1/s)

Fuel Side Air Side HRRMax

Normal Distance from the Flame Surface (mm)

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0

-120

-80

-40

0

40

80

Strain 

Rate (1/s)

Fuel Side Air Side HRRMax

Normal Distance from the Flame Surface (mm)

(b) local strain rate 

Figure 14. Computed normal component of velocity to the flame surface and local strain rate 

in the coordinates normal to the flame surface at three locations of 1-g counterflow 

nonpremixed flame for ag=20 s-1 and Vr=3. (• : location A, ⎯ : location B, --- : location C; 

see Fig. 10 for location of points A, B, and C). 

37
-0.20

-0.10

0.00

0.10

0.20

0 10 20 30 40 50 60 70 80 90 100

Global Strain Rate (s-1)

Energy Fraction .

1-g, Cr, radial convection

1-g, Dr, radial conduction

1-g, Mr, radial inter-diffusion

1-g, Rs, radiation

0-g, Rs, radiation, 1-D, 3-step

Figure 15. Computed fractional contribution by each term in the energy equation (Eq. 16) to 

the heat release of 0-g and 1-g flames with Ca=0.79 as a function of the global strain rate 

along the flow field centerline. 

38
Appendix 2 

Effect of Buoyancy on the Radiative Extinction Limit of Low Strain Rate 

Non-Premixed Methane-Air Flames

ABSTRACT

The structure and extinction of non-premixed flames were investigated through comparison of 

experiments and calculations using a counterflow configuration. Experiments were conducted at 

the NASA Glenn Research Center’s 2.2 s drop tower to attain suppression and temperature 

measurements in low-strain non-premixed methane-air microgravity flames. Suppression 

measurements using nitrogen added to the fuel stream were performed for global strain rates from

7 s-1 to 50 s-1. Judicious hardware selection and an optimized experimental procedure facilitated 

rapid, controllable, and repeatable flame extinction measurements. The minimum nitrogen 

volume fraction in the fuel stream needed to assure suppression for all strain rates in microgravity

was measured to be 0.855 ± 0.016, associated with the turning point, which occurred at a global 

strain rate of 15 s-1. This value was larger than the analogous value in normal gravity. Flame 

temperature measurements were attained in the high temperature region of the flame (T >1200 K) 

using visible emission from a SiC filament positioned axially along the burner centerline. The 

suppression and temperature measurements were used to validate a two-dimensional flame 

simulation developed here, which included buoyancy effects and finite rate kinetics. The 

simulations yielded insight into the differences between microgravity and normal gravity

suppression results, and also explained the inadequacy of the one-dimensional model results to 

explain the microgravity suppression results. 

39
1. INTRODUCTION 

Unwanted fires typically occur due to non-premixed combustion. The agent 

concentration required to achieve the suppression (or extinction) of non-premixed flames is an 

important consideration for design of fire protection systems as engineers rely on suppression 

data to estimate suppressant mass requirements for total flooding applications. Information 

regarding critical suppressant concentrations in the high strain rate regime is less important than 

in the low strain rate regime, because low strain rate conditions typically require higher agent 

concentrations and fire protection design must be based on worst-case conditions. The initial 

stage of an unwanted fire in a microgravity environment, such as an orbiting space platform, will 

likely occur at a low strain rate. The actual value of the strain rate will depend on the combustion 

configuration and local environmental conditions such as air currents generated by the fire itself 

and any forced ventilation. 

 A review of the literature indicates that there has been little work on low (< 30 s-1) 

strain rate flame suppression by an agent, and practically none on the suppression of low strain 

rate adiabatic flames in microgravity. Related studies include Refs. [1-10]. Maruta et al. [1]

conducted the first comprehensive suppression measurements of very low strain non-premixed 

flames in microgravity using the Japan Microgravity Center (JAMIC) 10 s drop tower. In their 

study, the suppression of methane-air diffusion flames with N2 added to the fuel stream was 

measured under microgravity conditions. The minimum methane concentration required to 

sustain combustion was measured to decrease as the strain rate decreased until a critical value 

was observed. As the global strain rate was further reduced, the required methane concentration 

increased. This behavior was denoted as a "turning point" and was attributed to the enhanced 

importance of radiative loss in low strain rate flames. In terms of fire safety, the turning point 

agent concentration represents a fundamental limit for nonpremixed flames, assuring suppression 

for all flow conditions. Bundy et al. [2] observed similar extinction behavior in low strain rate 

normal gravity flames. Zhang et. al. [3] studied extinction of premixed flames in microgravity 

40
using a 2.2 s drop tower. Ronney [4] examined the effect of CF3Br on the burning velocities, 

flammability limits and minimum ignition energies of alkane/air mixtures and determined that 

flame propagation occurred over a wider range of conditions at normal gravity than in 

microgravity. VanDerWege et al. [5] reported on the shape and the visual character of methane

jets (Re=100) flowing into a quiescent agent laden environment. Other investigators have 

examined extinction of diffusion flames through conductive losses to walls [6] or transient twophase diffusion effects in flames burning alcohol fuels [7]. Olson and T’ien [8] reported on the 

structure of low stretch PMMA diffusion flames burning in normal gravity. 

Experimental observations under microgravity conditions show that counterflow flames 

appear to be nearly flat [1]. Under normal gravity conditions, high strain rate flames are also flat 

as buoyancy is overwhelmed by convection. As the strain rate decreases in normal gravity,

buoyancy forces become relatively more important and the flames are clearly curved. Onedimensional flame codes such as OPPDIF [9] are based on a similarity solution that neglects 

buoyancy. These one-dimensional steady state simulations fail to describe the multi-dimensional 

structure of low strain rate normal gravity flames, nor do they capture the dynamics of a flame 

during the transition from normal to microgravity. To gain physical insight, a numerical method 

is needed that is capable of simulating flames under the influence of buoyancy, and to better 

understand the effects of finite burner size on the detailed flame structure. Frouzakis et al. [10] 

developed a two dimensional (2D) direct numerical simulation that considered potential 

problems arising from the use of a 1D similarity solution along the flow field centerline 

to represent the character of 2D counterflow flames [10]. In their study, the structure of 

low strain rate counterflow H2-air flames was investigated.

The objective of this study was to investigate the structure and suppression of low strain 

rate counterflow non-premixed flames in both normal gravity and microgravity through 

measurements and numerical simulations. The suppression effectiveness of a suppressant (N2) 

41
added to the fuel stream of low strain rate methane-air diffusion flames was measured. Flame

temperature measurements were attained in the high temperature region of the flame by radiative 

emission from a thin filament positioned axially along the burner centerline. A two-dimensional 

(2D) flame simulation, including buoyancy effects was developed, and was validated with the 

flame suppression and the temperature profile measurements. 

As mentioned above, microgravity suppression measurements similar to those reported here

were previously conducted [1]. Figure 1 replots the results presented in Fig. 4 of Ref. [1], which 

shows a comparison of the previous microgravity measurements with their calculations based on 

a 1D similarity solution. In this study (and in Figure 1), the global strain rate (ag) was defined 

[11] as: 

 ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = +

a a

f f a

g

V

V

L

V

a

ρ

ρ 1 2

 (1) 

where the parameters V and ρ denote the mean velocity and density of the reactant 

streams at the duct boundaries, L is the duct separation distance, and the subscripts a and 

f represent the air and fuel streams, respectively. The definition of the global strain rate used 

here is larger than the definition used in Ref. [1] by a factor of two. While there are similarities in 

the gross trends between the measurements and the simulation results in Fig. 1, some interesting 

differences are evident , particularly in the very low strain rate regime, where the calculations

overpredict the critical agent concentration in the fuel stream at extinction. The differences

between the model and the experimental results could be associated with a number of issues. 

Reference [1] reports that the location of mesh screens in the exits of the burner ducts was varied 

for each of the experimental conditions. The location of the mesh screens provides a well-defined 

boundary condition on the flow field within the duct, which impacts the flow boundary conditions 

and possibly comparison of the suppression results with the model. In addition, the combustion 

test chamber in Ref. [1] was filled with air, rather than an inert, non-reactive gas such as nitrogen. 

42
Experiments conducted in normal gravity show that use of an air ambient in the combustion 

chamber impacts the extinction results [2]. For these reasons, the work reported here revisits the 

previous microgravity suppression measurements. In addition, this study extends the previous 

experimental work by investigating the detailed flame structure of the near extinction flames 

through measurements of the local temperature. This information was useful for validation of a 

2D numerical code developed here and subsequently used to investigate the mechanisms of flame 

extinction under normal and microgravity conditions.

Like the measurements, there are also a number of questions related to the application of 

the 1D similarity model to the previous experimental measurements [1]. At low strain rates, the 

flames are rather thick, and the flame diameter/flame thickness ratio was not very large. Only

when this ratio is large enough, can the flame be classified as one-dimensional. In this case, 

lateral heat and diffusion losses may be significant, and neglect of these terms through use of a 

1D similarity model may be problematic. The 2D model developed here attempts to address 

these issues. 

Han et al. [12] recently reported on the structure of ultra low-stretch (~2 s-1) 

methane nonpremixed flames. Their study considered radiation effects and extinction in 

curved flames established in normal gravity by flowing diluted fuel in a downward 

direction through a porous spherically symmetric burner with a large radius of curvature. 

Heat losses to the burner were characterized and the computational results were found to 

be consistent with experimental observations. 

In this study, conductive heat losses to the burner were avoided, which impacted 

the range of strain rates investigated, because flame thickness increases as the strain rate 

decreases. Here, the global strain rate (ag) was varied from 7 s-1 to 50 s-1 in microgravity 

flames and from 14 s

-1 to 50 s-1 in normal gravity flames to avoid conductive heat losses to 

43
the burner - as determined through consideration of the profile of temperature near the 

burner boundaries To maximize the range of strain rates considered in the normal 

gravity flames considered here, the approach used by Ref. 2 was applied in which the 

velocity ratio of the air stream to the fuel stream (Va/Vf in Eq. 1) was adjusted such that 

the flame location was forced towards the center of the flowfield, away from the burner 

ducts.

2. EXPERIMENTAL METHOD 

Experimental Apparatus 

Microgravity experiments were conducted at the NASA Glenn Research Center 2.2 s drop tower. 

A 15 mm diameter stainless steel counter-flow burner was enclosed in a 25 L cylindrical 

chamber. The experimental hardware was mounted in a standard NASA drop rig. A schematic 

of the burner and flow system is shown in Fig. 2. To control each gas flow (air, CH4, and N2

added to the fuel stream), a pressure transducer, solenoid valve, pressure regulator, and a fast 

response time (≈50 ms) pressure controller were placed upstream of a critical flow orifice. The 

burner was designed to have minimal dead volume as glass beads and a series of fine mesh steel 

screens were used to impose a near plug-flow velocity profile. The duct separation distance, L, 

was 15 mm (see Fig. 2) and the wire mesh screens were positioned ≈1 mm into each of the ducts. 

This exact geometry was considered in the 2D numerical model developed here. The reactants

were stored in pressurized 500 cm3

 stainless steel cylinders. The oxygen content in the bottled air 

was measured to be (20.93 ± 0.04) % using a paramagnetic analyzer [13]. The fuel was research 

grade methane (99.99% purity). Flow control and data acquisition were performed using a 

PC104 microprocessor with 320 MB of flash memory and a 16 bit data acquisition board 

sampling at 200 Hz. 

44
Suppression Measurements

Attaining accurate suppression measurements in 2.2 s of microgravity was achieved by

judicious hardware selection (50 ms time response pressure transducers, associated critical flow 

orifices, and a small dead-volume flow system) and development of an optimized experimental 

procedure in which microgravity conditions were initiated at near-extinction (normal gravity) 

flame conditions. The combustion chamber was evacuated and filled to 101 kPa with N2 in 

contrast to the previous measurements [1]. The flow configuration used for these experiments

used air from the top duct and methane diluted with nitrogen from the bottom duct. The initial N2

volume fraction was set as close to normal gravity suppression conditions as possible. Ignition 

was accomplished by positioning and energizing a 0.25 mm diameter (6 cm long) coiled Pt + 

30% Rh wire between the ducts after the reactant flows were initiated. Upon ignition, the ignition 

wire was simultaneously de-energized and slowly retracted from the flame at a rate of 2 cm/s to 

minimize convective disturbances. About 5 s later, after the flame had stabilized, microgravity 

conditions were initiated by releasing the drop package. From this point on, the extinction 

experiments were automated. 

Flame suppression was achieved by increasing the agent flow and simultaneously decreasing 

the fuel flow, while maintaining a constant global strain rate, until a critical agent volume fraction 

in the fuel stream was attained. For each experiment, the onboard computer recorded the 

transient strain rate, nitrogen concentration in the fuel stream, the chamber pressure, the start and 

end of the drop period, the flow meter flows, and the signal from a photodetector inside the 

chamber. Acceleration levels during the drop were measured as less than 10-3 m/s

2

. The 

photodetector signal dropped abruptly when the flame was extinguished and the time of 

suppression was independently confirmed from inspection of a digital video recording. The 

45
suppression concentration was inferred by determining the N2 concentration based on the time of 

flame suppression and the delay time of the system. 

The experimental method is illustrated in Figure 3, which shows the measured nitrogen 

volume fraction in the fuel stream of a 15 s-1 methane-air flame as a function of time after drop 

initiation. The photodetector signal is also shown, which indicated the time of flame extinction 

when its value rapidly decreased. The critical N2 concentration at extinction was inferred from

the concentration-time data through consideration of the system delay time, which was the time 

of travel from the fuel/N2 mixing tee (see component number 7 in Fig. 2) to the flame zone, at the 

middle of the flow field. The delay time was calculated through consideration of the system

volume and the gas flows. The calculation was verified by performing extinction experiments for 

different N2 “ramping” rates during the experimental drop for the same flame conditions, with the 

ramping rates varying about 20 % to 30 %. These experiments resulted in determination of the 

values of the critical N2 volume fraction at extinction to within 0.1 %, verifying the estimates of 

the system delay times. Experiments in normal gravity also showed that the critical N2

suppression volume fraction was independent of the N2 ramping rate, for the rates considered 

here. Since the N2 volume fraction at the beginning of the experimental drop was very close to

the extinction concentration (within 2 %, see Fig. 3), uncertainty in the delay time was not a 

significant contributor to the uncertainty in the N2 volume fraction. The expanded relative 

uncertainty in the N2 volume fraction was 2 %, dominated by uncertainty in the flows, based on a 

propagation of error analysis and repeat measurements. For all measurements reported here, the 

expanded uncertainty is presented with a coverage factor of two, i.e., an uncertainty of 2σ

representing a 95 % confidence interval.

The time for the flame to make the transition to microgravity was determined from

observation of video images in experiments in which the N2 concentration was held constant. 

46
The longest transition time, at 7 s-1, was about 0.2 s, which was less than 10 % of the total 

microgravity time available. 

Temperature Measurements

Measurement of the visible emission intensity from a 12.5 (± 0.5) μm β-SiC filament placed 

along the burner centerline allowed determination of flame temperatures for T>1200 K. Below 

this temperature, the signal to noise ratio was inadequate. Radiation emitted by the filament was 

recorded using a digital CCD camera with a close-up lens such that the spatial resolution of the 

image was 0.07 mm/pixel. The camera exposure was adjusted to prevent image saturation (overexposure) at the maximum flame temperature. Flame emission intensity was insignificant 

compared to the filament intensity. 

The total emissive power from a surface can be expressed using Plank’s Law 

integrated over all wavelengths as E = εσT4

. The fraction of the total radiance captured 

by the detector is a function the spectral sensitivity of the detector, the transmissivity of

the optics of the detector, and the solid angle. Because the spectral response curve of the 

detector and the transmissivity of the optics are not constant (and are unknown), a linear 

relationship between the detector signal and T4

 could not be assumed. A calibration 

source was required to find the relationship between signal intensity and filament 

temperature. 

The intensity measurements were calibrated using the one-dimensional OPPDIF flame 

code [9], following Ravikrishna and Laurendeau [14]. In order to equate the fiber 

luminescence to a fiber temperature, the profiles of the flame temperature and the emitted 

intensity along the length of the fiber for a given set of flame conditions was required. 

The correlation was based on a comparison of centerline counterflow flame 

measurements with OPPDIF calculations for a microgravity moderately-strained 

47
methane-air diffusion flame with a N2 concentration of 81 % by volume in the fuel 

stream. The OPPDIF predicted gas temperature profile was corrected to the predicted 

fiber temperature based on heat transfer to the fiber. The fiber temperature was then 

compared to the measured fiber luminosity and a correlation was determined with 

luminosity taken as proportional to the fiber temperature to the fifth power. The expanded 

uncertainty in the temperature measurement was typically 60 K, based on repeat measurements. 

Because the conductivity along the filament is negligible [15], the energy balance 

becomes: 

h P (T – Tf) = σ εP Tf

4 (2) 

where h is the convection coefficient, T is the gas temperature, Tf is the filament temperature, P is 

the perimeter of the filament, σ is the Stefan-Boltzman constant, and ε is the filament emissivity. 

Solving Eq. 2 for T, a radiative correction for the gas temperature results: 

T = Tf + σ ε Tf

4

/ h (3) 

The convection coefficient was found by using a numerical correlation for low velocity 

cross flow over a cylinder [16]. Experiments compared the filament intensity with the filament 

oriented horizontally and vertically in the counterflow flame to determine if there was a 

significant difference caused by filament orientation. The difference between the two 

orientations was 15 K, which was within experimental error, so the cross flow assumption for the 

convection coefficient was considered acceptable. Extinction measurements, performed with and 

without the filament present, showed that the filament had a negligible effect (<0.1 % difference) 

on the agent concentration required for extinction. Using an optical microscope, inspection of

the filament after flame exposure revealed that there was no observable change in the diameter or 

other degradation in these relatively low temperature short-duration experiments.

48
3. NUMERICAL METHOD 

The present computation employed a time-dependent axisymmetric configuration to treat 

counterflow diffusion flames formed between two opposed circular ducts. A coupled set of 

model-free equations with a low Mach number approximation was used following Ref. [17]. 

QUICK [18] and second-order central difference schemes were used to discretize the convective 

and diffusive terms. A predictor-corrector scheme nearly identical to that of Najm et al. [19] was 

used for time integration of the governing equations. Efficient algebraic relaxation for the 

velocity-pressure correction was performed using the HSMAC method [20]. CHEMKIN-II [21]

and TRANFIT [22] were adopted for calculation of thermodynamic and transport properties. The 

computational domain in the axial direction was taken to be 40 mm. In the radial direction, 

domains of 70 mm and 50 mm were used for the zero and normal gravity conditions, respectively. 

In addition, 268 x 70 grids were used for the zero gravity simulation and 268 x 65 grids for 

normal gravity. A uniform grid spacing of 0.15 mm was used in the axial direction. Radially, 

non-uniform grids with a minimum spacing of 0.3 mm were clustered near the centerline. A grid 

sensitivity study confirmed that the calculated extinction limit and the temperature maximum was 

unchanged when a finer grid (0.10 mm in the axial direction) was used, indicating that the heat 

release rate zone was adequately resolved. 

The inflow temperature boundary condition for the reactant streams was 298 K. Uniform

axial velocity was imposed at the mesh screens, which were positioned about 1.2 mm into each of 

the ducts. This condition is appropriate for global strain rates in 0G and 1 G above about 7 s-1

and 15 s-1, respectively, as exemplified by the temperature gradient, which is nominally flat 

under those conditions. Sides A and B shown in Fig. 2 were treated as outflow boundary

conditions for the normal gravity and zero gravity calculations, respectively. A no slip condition 

was applied on the burner walls and the wall temperature was taken as 298 K. A no mass flux 

condition was applied at all boundaries. The oxidizer stream was composed of undiluted air 

composed of 21% oxygen and 79% nitrogen by volume, and the fuel was methane diluted by

49
nitrogen. The ambient gas was set to pure nitrogen, as in the experiment, to prevent secondary 

combustion of fuel. 

Chemical reaction mechanism and radiation model

A 3-step irreversible reaction mechanism [23] for methane oxidation was used. Use of a larger, 

more comprehensive chemical mechanism was precluded due to practical reasons associated with

computational cost. The 3 step chemical model is adequate, as this paper focuses on the

energetics associated with flame suppression, and not detailed reaction pathways. The detailed 

mechanism and reaction rates were taken as [23]: 

CH4 + (3/2)O2 → CO + 2H2O (4)

CO + (1/2)O2 → CO2 (5)

CO2 → CO + (1/2)O2 (6) 

 -d[CH4]/dt = 1011.68 exp(-23500/T)[CH4]

0.7

[O2]

0.8 (7) 

-d[CO]/dt = 1012.35 exp(-19200/T)[CO]1.0

[H2O]0.5[O2]

0.25 (8) 

 -d[CO2]/dt = 1012.50 exp(-20500/T) )[CO]1.0

[H2O]0.5[O2]

0.25 (9) 

with reactions rates in kmole-m-3-s-1. 

The viability of the 3 step global mechanism was demonstrated by performing low strain 

rate flame calculations using OPPDIF [24] and by comparing the calculations to flame 

measurements (see the discussion of Figs. 5 and 7 below). OPPDIF calculations using the GRI3.0 mechanism yielded very similar agent extinction requirements as OPPDIF calculations using 

the 3-step mechanism, while the temperature peak using the 3-step mechanism was less than 100 

K larger than that obtained using GRI-3.0. Although detailed chemical reactions are not included 

in the 3-step mechanism, it was considered appropriate to predict the extinction limits of N2-

diluted methane flames, because Nitrogen is known to extinguish a flame mainly by physical 

influences, such as heat capacity and dilution effects, rather than through chemical effects. An 

optically-thin radiation sub-model [26] was implemented, which considered heat loss from CH4, 

50
CO2, H2O, and CO. Soot was not considered in these thin blue flames (see Fig. 4). The 

volumetric rate of radiative heat loss in the energy equation can be written as: 

4 4 4 ( r p q Kσ T T ) = − − ∞ & (10)

where T and T∞ are the local and the ambient temperature, respectively. Kp is the Plank mean

absorption coefficient of the mixture, which is expressed as: 

K P p = + CH4 4 KCH PCO2 KCO2 + PH2O KH2O + PCO KCO (11) 

where and denote the partial pressure and Plank mean absorption coefficient of species i, 

respectively, with Kp taken from Ref. [26]. 

Pi Ki

4. RESULTS AND DISCUSSION 

Temperature profiles were measured along the burner axis and comparisons were made between 

the microgravity and normal gravity results. As the flame transitioned into microgravity, the 

maximum flame temperature location shifted towards the center of the flowfield and the flame 

width increased. Figure 5 compares the measured axial temperature profiles in normal and 

microgravity for ag = 20 s-1 and an N2 volume fraction in the fuel stream equal to 0.80. The 

measured microgravity peak temperature was essentially the same as in normal gravity. The 

predicted (2D model) peak temperatures in the 0-g and 1-g flames were in agreement with the 

measured peaks to within experimental uncertainty (60 K), while the locations of the peaks were

within about 0.3 mm of the measurements. The computed temperature fields are also shown in 

Fig. 6, which shows that the microgravity flame is thicker along the axis and over the entire flow 

field consistent with the flame images seen in Figs. 4 and the calculated results shown in Fig. 5. 

The simulated flames appear to be similar in shape to the experimentally-observed visible flame 

shapes, but there is a tendency for the model to over-predict the fuel-side temperature. This is 

likely due to the limitations associated with the simplified global chemical model, which does not 

51
address real kinetic effects such as dissociation, recombination, chain branching, etc. Near the 

edge of the burner, even the microgravity flames are not simply one-dimensional in shape. 

Figure 7 shows measurements, and one-dimensional (1D) and two-dimensional (2D) model 

calculations of the critical N2 volume fraction in the fuel stream required to extinguish the 

methane-air diffusion flames as a function of the global strain rate. The figure compares the 

microgravity suppression measurements conducted here with the normal gravity results reported 

by Bundy et al. [2]. For the normal gravity flames, the air/fuel velocity ratio was adjusted, such 

that the flame was positioned away from the burner duct so that heat loss to the burner was 

negligible. The extinction data for normal gravity and microgravity collapse into one curve for 

moderate global strain rates (ag > 40 s-1) as convective forces begin to dominate the effects of 

buoyancy. Figure 7 also shows the microgravity suppression measurements of Maruta et al. [1],

which are in agreement with the microgravity results presented here. The one-dimensional (1D) 

calculations of flame extinction (reported in [1] and confirmed in this study) failed to 

quantitatively reproduce the experimentally determined microgravity extinction results 

determined here and in Ref. [1]. This is attributed to the importance of lateral heat losses 

associated with radial conduction in the energy equation, which are neglected in the 1D similarity 

solution, but are considered in the 2D model calculation. The 2D modeling results shown in

Fig. 7 appear to be adequate, showing agreement with the microgravity experiments to within 

experimental uncertainty, despite the use of simplified chemistry and radiation submodels. For 

the very lowest strain rates (10 s-1), the 2D model somewhat underpredicts the critical agent 

concentration, due perhaps to the use of the optically thin radiation sub-model, which tends to 

overpredict radiative heat loss, especially in the relatively thick lower strain rate flames. 

The N2 volume fraction in the fuel stream at the turning point represents the minimum agent 

needed to assure suppression regardless of strain rate. Its value in microgravity was measured as 

0.855 ± 0.016 at a global strain rate of 15 s-1 ± 5 s-1. This value was larger than the analogous 

value in 1-g. Figure 7 shows that the microgravity suppression measurements and the 2D model 

52
suppression results drop sharply for strain rates lower than the turning point. The ‘turning point’

in normal gravity occurred at a higher global strain rate than in microgravity for both the 

simulations and the measurements. 

At low strain rates, the extinction process is very different in normal gravity and 

microgravity, due to differences in flame structure. In this regard, it is helpful to consider the 

specific maximum heat release rate per unit volume divided by the local strain rate (SMHRR),

which represents the flame strength along a flame surface as introduced by Sung et al. [27]. 

Calculations show that there is no effective difference between using the peak versus integrated 

reaction rates in comparing flame strength. Smaller values of SMHRR imply that the flame is 

weak and easier to extinguish. Table 1 compares the flame strength at three locations for flames 

with global strain rates of 10 s-1 (0-g) and 20 s-1 (0-g and 1-g), and an agent fuel stream volume 

fraction of 0.79. Table 1 lists the calculated (2D model) maximum heat release rate (HRRmax), 

the local strain rate (al), and the specific maximum heat release rate (SMHRR). as a function of 

radial location in the normal gravity and microgravity flames. The table shows that SMHRR at 

the center of the flow field is larger in the 0-g 20 s-1 flame as compared to the 1-g 20 s-1 flame. 

This is consistent with the results shown in Fig. 7 in which the normal gravity flame requires less 

agent to extinguish than the 0-g flame, both at 20 s-1. The calculated value of the SMHRR in 1-g 

is nearly constant near the center of the flow field, but decreases by 1/3 of this value near the edge 

of the burner (at a radial distance, r, equal to the radius of the burner duct, R). This is consistent 

with the observation that the 1-g low strain rate flame extinguishes from the outer edge first. The 

SMHRR of the 0-g flame is also constant near the center, but increases towards the flame edge. 

Unlike the 1-g flame, the microgravity flame under these conditions abruptly and simultaneously

extinguishes over the entire flame surface. For the same global strain rate (see Eq. 1), the local 

strain rate along the flame axis in the 0-g flame (= 36 s-1) is about 1⁄2 the value of the local strain 

rate in the normal gravity flame (= 62 s-1). In 1-g, buoyancy tends to enhance the local strain rate 

as hot gases are “pulled” above the burner (see Fig. 6). The table also shows that the SMHRR 

53
value at the edge of the 0-g flames is smaller in the 10 s-1 flame than in the 20 s-1 flame (both in 

0-g). This is consistent with the results shown in Fig. 7 in which the 10 s-1 flame requires 

somewhat less agent to extinguish than the 20 s-1 flame (both at 0-g). 

Figure 8 shows the 2D model results of the local fractional energy loss contributed by key

terms in the energy equation (and normalized by the heat release rate) along the flowfield 

centerline of 0-g and 1-g flames with 0.79 N2 volume fraction in the fuel stream as a function of 

the global strain rate. The figure also shows the radiative loss fraction determined in this study 

using the 1D model. The calculations show that axial conduction and diffusion are by far the 

largest heat loss mechanisms. The 2D model results indicate that radiative heat transfer is a 

significant energy loss mechanism in the 0-g flames, becoming larger as the strain rate decreases. 

Radiative heat loss is not negligible even at moderate strain rates in near-extinction 0-g flames, 

nor is it zero in 1-g flames. Radial diffusion is found to be a relatively small contributor to energy

loss for global strain rates above about 20 s-1 in both the 0-g and 1-g flames. As the strain rate 

decreases below 15 s-1. Its magnitude is about 1⁄2 that of the radiative loss term for a global strain 

rate of 10 s-1 in the 0-g flames. The combination of these loss mechanisms becomes larger as the 

strain rate decreases below the turning point at 15 s-1, consistent with the observed decrease in 

agent concentration requirements presented in Fig. 7. Figure 8 shows that the radiative loss terms 

in the2D and 1D models were similar in magnitude. 

5. SUMMARY AND CONCLUSIONS 

A fundamental study of the structure and suppression of low strain nonpremixed flames was 

conducted using experiments and numerical computations. For the first time, temperature 

profiles were determined in near extinction diffusion flames in microgravity. The temperature 

profile measurements, along with measurements of the critical agent suppression requirements, 

were used to validate the 2D flame code. The 2D model calculations also showed agreement with 

experimental observation of flame curvature. 

54
Analysis, using flame simulations, allows insight of the heat transfer processes that 

control flame structure and extinction. At low strain rates, the flames are rather thick and lateral 

heat and diffusion losses are not negligible. For finite size burners, the simulations confirm that 

low strain rate counterflowing nonpremixed flames cannot be assumed to be purely onedimensional in normal gravity or even microgravity. The commonly used one-dimensional flame 

codes do not provide a full picture of the multi-dimensional counterflow structure for experiments 

involving low strain normal gravity or microgravity flames. By providing physical insight, twodimensional simulations are a useful tool for guiding the interpretation of counterflow flame 

experiments. 

ACKNOWLEDGEMENTS 

This research was partially supported by the NASA Microgravity Research Division through 

Contract No. C–32066–T with Dr. Sandra Olson serving as Technical Monitor. The authors are 

grateful to Drs. K. Maruta, K.Y. Lee, J. Logue, W.C. Park, J. Park, and I.K.Puri for helpful 

discussions. 

55
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56
Table 1. The computed (2D) local maximum heat release rate (HRRmax), the local strain rate 

(al), and the specific maximum heat release rate (SMHRR) as a function of radial location in 0-g 

and 1-g flames at ag=10 s

-1

 and 20 s-1 for an agent fuel stream volume fraction of 0.79.

Location a

HRRmax

(J/cm3s) 

al

(s-1) 

SMHRR 

(J/cm3) 

0g; ag = 10 s-1 

r/R=0.0 58.7 19 3.1

r/R=0.5 53.9 17 3.3

r/R=1.0 37.4 10 3.8

0g; ag = 20 s-1

r/R=0.0 110 36 3.1

r/R=0.5 107 33 3.2

r/R=1.0 82 19 4.3

1g; ag = 20 s-1

r/R=0.0 165 62 2.7

r/R=0.5 164 67 2.4

r/R=1.0 147 84 1.8

a. R is the inner duct radius; r/R=0 is on the central axis. 

57
List of Captions for the Figures 

Figure 1. Computed (1D model) and previous measurements of the critical nitrogen volume 

fraction in the fuel stream for suppression of CH4 flames in microgravity [1].

Figure 2. Schematic of counterflow system including (1) gas cylinders, (2) pressure transducers, 

(3) solenoid valves, (4) pressure regulators, (5) pressure controllers, (6) critical flow orifices, (7) 

mixing tee, (8) burner, and (9) glass beads and metal screens. 

Figure 3. Measured nitrogen model fraction in the fuel stream of a 15 s

-1 flame as a function of

time. The photodetector signal is also shown, which indicates the time of flame extinction. 

Figure 4. Flames in normal gravity (left) and microgravity for ag=20 s-1 and the fuel stream N2

volume fraction = 0.80. 

Figure 5. Computed (2D model) and measured temperature profiles in normal gravity and 

microgravity for 20 s-1 methane–air flames with 0.80 N2 volume fraction in the fuel stream. 

Figure 6. Computed 2D temperature distributions of steady ag= 20 s

-1 near suppression methaneair flames in (a) normal gravity and (b) zero gravity.

Figure 7. Computed and measured critical nitrogen volume fraction in the fuel stream for 

suppression of CH4 flames in microgravity and normal gravity.

Figure 8. Results of 2D model calculations of the fractional contribution by axial convection, 

axial diffusion, radial diffusion, and radiation in the energy equation (normalized by the heat 

release rate) in 0-g (lines) and 1-g flames (symbols) along the flow field centerline with 0.79 N2

volume fraction in the fuel stream as a function of the global strain rate. The radiative loss 

fraction determined using the 1D calculation is also shown.

58
0.78

0.79

0.80

0.81

0.82

0.83

0.84

0.85

0.86

0.87

0 10 20 30 40 50 6

Global Strain Rate (1/s)

Critical Volu

me Fractio

n at Extinctio

n

0

μ-g experiments [1]

0-g, 1D model [1]

Figure 1. Previously computed (1D model) and measured critical nitrogen volume fraction 

in the fuel stream for suppression of CH4 flames in microgravity [1]. 

59
1 

 Pressure Chamber 

CH4 

2 

3 

4 

5 

6 

7 

L 

Air N2 

8 

9 

Side B

g 

Side A

Figure 2. Schematic of counterflow system including (1) gas cylinders, (2) pressure transducers, 

(3) solenoid valves, (4) pressure regulators, (5) pressure controllers, (6) critical flow orifices, (7) 

mixing tee, (8) burner, and (9) glass beads and metal screens. 

0.84

0.85

0.86

0.87

0.88

0.89

0 0.5 1 1.5 2

Time from start of drop (s)

N2 Volume Fraction . 

0

0.05

0.1

0.15

0.2

0.25

Photodetector Signal (V)

extinction from

photodetector

delay time

inferred

extinction

concentration

Figure 3. Measured nitrogen volume fraction in the fuel stream of a 15 s-1

 methane-air 

flame as a function of time. The photodetector signal is also shown, indicating the time 

of flame extinction. 

60
Figure 4. Flames in normal gravity (left) and microgravity for ag=20 s-1 and a fuel stream 

N2 volume fraction of 0.80. 

300

500

700

900

1100

1300

1500

1700

1900

0.000 0.003 0.006 0.009 0.012 0.015

Distance from air duct (m)

T

emperature (K)

1-g, exp.

μ-g, exp.

1-g, 2D model

0-g, 2D model

Figure 5. Computed (2D model) and measured temperature profiles in normal gravity and 

microgravity for 20 s-1 methane–air flames with 0.80 N2 volume fraction in the fuel stream. The 

expanded uncertainty in the temperature measurement was estimated as 60 K. 

61
0 5 10 15 20 25 30 35 40 r (mm)

-10

-5

0

5

10

15

20

25

x(mm)

1600

1459

1318

1177

1036

895

753

612

471

(a) normal g 330

0 5 10 15 20 25 30 35 40 r (mm)

-10

-5

0

5

10

15

20

25

x(mm)

1600

1459

1318

1177

1036

895

753

612

471

(b) zero g 330

T (K)

Figure 6. Computed 2D temperature distributions of steady ag= 20 s-1 near suppression 

methane-air flames in (a) normal gravity and (b) zero gravity.

62
0.74

0.76

0.78

0.80

0.82

0.84

0.86

0.88

0 10 20 30 40 50 60 70 8

Global Strain Rate (1/s)

Critical Volume Fraction at Extinction

0

μ-g experiments [1]

Current μg experiments

1-g experiments [2]

0-g, 2D model

1-g, 2D model

0-g, 1D model

Figure 7. Computed and measured critical nitrogen volume fraction in the fuel stream for 

suppression of CH4 flames in microgravity and normal gravity. 

63
Global strain rate [1/s]

Energy loss fr

action

0 20 40 60 80 0.30

0.35

0.40

0.45

0.50

0.55

0.60

0-g, axial convection

1-g, axial convection

0-g, axial diffusion

1-g, axial diffusion

Global strain rate [1/s]

Energy loss fr

action

0 20 40 60 80 -0.05

0.00

0.05

0.10

0.15

0-g, radial diffusion

1-g, radial diffusion

0-g, radiation

1-g, radiation

0-g, radiation, 1D

Figure 8. Results of 2D model calculations of the fractional contribution by axial 

convection, axial diffusion, radial diffusion, and radiation in the energy equation (normalized by

the heat release rate) in 0-g (lines) and 1-g flames (symbols) along the flow field centerline with 

0.79 N2 volume fraction in the fuel stream as a function of the global strain rate. The radiative 

loss fraction determined using the 1D calculation is also shown.

64
Appendix 3 

Numerical Study of the Suppression of Methane-Air Flames with Carbon Dioxide in 

Normal Gravity and Microgravity 

ABSTRACT

The structure and extinction of non-premixed flames was investigated through comparison of 

experiments and calculations in a methane-air counterflow flame with CO2 added to the fuel 

stream. Flames with low global strain rates (20 s-1 to 120 s

-1) are considered. Axisymmetric 

flame simulations yield insight into the superior effectiveness of CO2 as an agent as compared to 

N2. Differences in radiative exchange when CO2 rather than N2 is used as a suppressant are 

highlighted. In addition, a comparison of differences in the structure and extinction of 

microgravity and normal gravity flames are also considered. Measurements and computations of 

the maximum flame temperature near the flow field centerline for near-extinction normal gravity 

methane-air flames diluted by CO2 show behavior analogous to N2 diluted flames in which the

maximum temperature has a minimum value between strain rates of 20 s-1 to 30 s-1. By providing 

physical insight, two-dimensional simulations are a useful tool for guiding the interpretation of

counterflow flame experiments. The analysis, using flame simulations, allows insight of the heat 

transfer processes that control flame structure and extinction. 

65
1. INTRODUCTION 

A series of measurements and calculations have been undertaken in an effort to improve the 

understanding of the suppression of low strain rate flames by an agent in microgravity. The 

current study extends previous work by developing a two-dimensional numerical simulation of

flame structure and extinction in micro-gravity and normal gravity for methane-air non-premixed 

flames with carbon dioxide as a suppressant added to the fuel stream. 

The investigation is relevant from both a fire safety and a fundamental perspective. 

Unwanted fires typically occur due to non-premixed combustion. The initial stage of an unwanted 

fire in a microgravity environment, such as an orbiting space platform, will likely occur at a low 

strain rate. A primary motivation of this work is to characterize the agent concentration required 

to achieve extinction of counterflowing non-premixed (or diffusion) flames at low strain rates. 

The agent concentration required to achieve the suppression of non-premixed flames is an 

important consideration in design of total flooding fire protection systems. Carbon dioxide is of

interest because it is currently used as the suppressant of choice on-board orbiting space 

platforms. 

The counterflow flame geometry is a convenient configuration for control of the flow 

field strain rate as well as the flame position [1]. From an experimental perspective, the 

counterflow configuration has been used to acquire well-defined data on the structure and 

extinction behavior of flames. Furthermore, the counterflow configuration is amenable to 

mathematical and numerical analysis, and has provided useful insight into local “flamelet” 

structure [1-4]. 

It is common engineering practice to develop and apply models that are as simple as 

possible, while maintaining the salient physics. Viewed as a flamelet, the counterflow flame is 

typically modeled as one-dimensional (1D). Experimental observations under microgravity

conditions have confirmed that counterflow flames are nearly flat [5]. Under normal gravity

conditions, high strain rate flames are also flat near the flow field centerline as buoyancy is 

dominated by convection. As the strain rate decreases in normal gravity, however, buoyancy

forces become relatively more important and counterflow flames are clearly curved. Onedimensional flame codes such as OPPDIF [6], which are based on a similarity solution that 

neglects buoyancy fail to describe the multi-dimensional structure of low strain rate normal 

gravity flames, nor do they capture the dynamics of a flame during the transition from normal 

gravity to microgravity. To gain physical insight, a numerical method is needed that is capable of 

simulating realistic flames that are multidimensional due to the influence of buoyancy. The price 

66
of using a multidimensional modeling approach is simplification of the complex flame chemistry,

such that the duration of a computation is practical. Such approximations are acceptable provided 

that the most significant observables are adequately represented. 

There have been many studies on the structure and extinction of moderately strained 

counterflow nonpremixed flames [1-4, 7]. Relatively few studies, however, have considered low 

strain rate flames. Maruta et al. [5] conducted the first comprehensive extinction measurements of 

very low strain nonpremixed flames in microgravity using the 10 s JAMIC drop tower in Japan. 

In that study, the extinction of methane-air diffusion flames with N2 added to the fuel stream was 

measured under microgravity conditions. The minimum methane concentration required to 

sustain combustion was measured to decrease as the strain rate decreased until a critical value 

was observed. As the global strain rate was further reduced, the required methane concentration 

increased. This behavior was denoted as a "turning point" and was attributed to the enhanced 

importance of radiative loss in low strain rate flames. In terms of fire safety, the turning point 

agent concentration represents a fundamental limit for nonpremixed flames, assuring suppression 

for all flow conditions. Bundy et al. [8] observed similar behavior in low strain rate normal 

gravity flames. Recently, the extinction behavior of ultra low-stretch, buoyant counterflowing 

nonpremixed flames established in a large Tsuji-type burner was investigated [9]. In that study, 

information about the extinction limit was not provided, although multi-dimensional effects on 

flame extinction were discussed. 

The objective of this study is to investigate the structure and extinction characteristics of 

low strain rate counterflowing methane-air nonpremixed flames using carbon dioxide as the agent. 

An axisymmetric 2D computational model, including buoyancy effects and radiative heat transfer 

was developed. The model was validated through comparison of the 2D model with experiments 

and 1D computational results for counterflow nonpremixed flames diluted by nitrogen. Once 

validated, the calculation was used in this study to explore the near-extinction flame structure for 

flame suppressed by CO2.

2. EXPERIMENTS 

The calculation results were validated by comparison to experimental measurements with

N2 and CO2 as the suppressant using a burner system that is shown in Figure 1. Experimental

methods were described previously [1]. The inner burner duct diameter (D) was 23.4 ± 0.1 mm

67
and the duct separation distance (L) was set to 25.0 ± 0.1 mm. Four 200 mesh stainless steel

screens were positioned 1.6 ± 0.1 mm from the duct exits and at 1.6 mm intervals to impose a tophat velocity profile. A water-cooled jacket about the top burner duct prevented heating of the 

metal burner and preheating of the supplied reactants. The reactant flows were controlled using 

mass flow controllers that were calibrated using a dry cell primary flow meter with an uncertainty

of 1 %. Flame temperature measurements were conducted using thin filament pyrometry [2]. The 

details of the measurement and calibration methods are given in Refs. [1, 2]. 

Still flame images were taken with a camera using a 1 s shutter opening and an f stop of

2.8 to compare the steady flame shapes with the computations. Images of flame extinction was 

recorded with a digital video camera to clarify the difference between modes of extinction at low

and high strain rates. The calculation results for the extinction limits as a function of a global

strain rate for normal gravity (1-g) counterflow flames was compared with previous

measurements given in Ref [1]. 

Counterflow flame conditions were defined by the global strain rate and the concentration 

of agent (N2 or CO2) in the methane fuel stream. The fractional amount of agent in the fuel 

stream was denoted as Ca. For a given global strain rate, the flame extinguished for a critical 

value of Ca. The global strain rate, ag, is used to characterize the flow field in a counterflow 

flame. Its definition is [3]:

 ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = + ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = +

r O

r F F

O O

O F F

g

V ρ

ρ

L

V V

V ρ

V ρ

L

V 1 2 1 2

a (1) 

where the parameters V denotes the velocity and ρ the density of the reactant streams at the duct 

boundaries, L is the duct separation distance, and the subscripts O and F represent the air and fuel 

streams, respectively. The velocity ratio, Vr, is defined as equal to VO/VF. The global strain rate

concept allows quantification of a characteristic flame residence time or flow time. 

3. NUMERICAL METHOD 

The present study computation employed a time-dependent axisymmetric configuration 

to treat counterflow diffusion flames formed between two opposed circular ducts. A coupled set 

of model-free equations with a low Mach number approximation was used following Ref. [4].

QUICK [5] and second-order central difference schemes were used to discretize the convective 

and diffusive terms. A predictor-corrector scheme nearly identical to that of Najm et al. [6] was

68
used for time integration of the governing equations. Efficient algebraic relaxation for the 

velocity-pressure correction was performed using the HSMAC method [7]. CHEMKIN-II [8]

and TRANFIT [9] were adopted for calculation of thermodynamic and transport properties. 

The computational domain in the axial direction and radial direction was taken to be 40

mm and 70 mm respectively. In addition, 268 x 70 grids were used for the zero gravity 

simulation and 268 x 65 grids for normal gravity. A uniform grid spacing of 0.15 mm was used 

in the axial direction. Radially, non-uniform grids with a minimum spacing of 0.3 mm were 

clustered near the centerline. A grid sensitivity study confirmed that the calculated extinction 

limit and the temperature maximum was unchanged when a finer grid (0.10 mm in the axial 

direction) was used, indicating that the heat release rate zone was adequately resolved. 

The inflow temperature boundary condition for the reactant streams was 298 K. Uniform

axial velocity was imposed at the mesh screens, which were positioned about 1.2 mm into each of 

the ducts. This condition is appropriate for global strain rates in 0G and 1 G above about 7 s-1 and 

15 s-1, respectively, as exemplified by the temperature gradient, which is nominally flat under

those conditions. The computational geometry is shown in Fig. 2. Sides A and B in Fig. 2 were 

treated as outflow boundary conditions. A no slip condition was applied on the burner walls and 

the wall temperature was taken as 298 K. A no mass flux condition was applied at all boundaries. 

The oxidizer stream was composed of undiluted air composed of 21% oxygen and 79% nitrogen

by volume, and the fuel was methane diluted by suppressants. The ambient gas was set to pure

nitrogen, as in the experiment, to prevent secondary combustion of fuel.

3.1 Chemical reaction mechanism and radiation model 

A 3-step irreversible reaction mechanism [10] for methane oxidation was used. Use of a

larger, more comprehensive chemical mechanism was precluded due to practical reasons

associated with computational cost. The 3-step chemical model is adequate, as this paper focuses 

on the energetics associated with flame suppression, and not kinetic minutiae. The detailed

mechanism and reaction rates were taken as [10]: 

CH4 + 3/2 O2 → CO + 2 H2O (4)

CO + 1⁄2 O2 → CO2 (5)

CO2 → CO + 1⁄2 O2 (6) 

and the associated reaction rates were: 

-d[CH4]/dt = 1011.68 exp(-23500/T)[CH4]

0.7

[O2]

0.8 (7) 

69
-d[CO]/dt = 1012.35 exp(-19200/T)[CO]1.0

[H2O]0.5[O2]

0.25 (8) 

-d[CO2]/dt = 1012.50 exp(-20500/T) )[CO]1.0

[H2O]0.5[O2]

0.25 (9) 

with reaction rates in units of kmole-m-3-s-1. 

The viability of the 3 step global mechanism was demonstrated by performing low strain 

rate flame calculations using OPPDIF [11]. The calculations showed that the GRI-3.0 

mechanism [12] and the 3-step mechanism yielded nearly the same extinction and flame 

temperature profile results. An optically-thin radiation sub-model [13] was implemented, which 

considered heat loss from CH4, CO2, H2O, and CO. Soot was not considered in these thin blue 

flames. The volumetric rate of radiative heat loss in the energy equation can be written as: 

4 4 4 ( q K r p σ T T ) = − − ∞ & (10)

where T and T∞ are the local and the ambient temperature, respectively. Kp is the Plank mean

absorption coefficient of the mixture, which is expressed as: 

K P p = + CH4 4 KCH PCO2 KCO2 + PH2O KH2O + PCO KCO (11) 

where, and denote the partial pressure and Plank mean absorption coefficient of species i, 

respectively, with Kp taken from Ref. [13]. 

Pi Ki

(12)

1

∑ ∑ ∑ = = =

+ − ∂

∂

∂

∂

+

∂

∂

∂

∂

+

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

∂

∂ ⎟+

⎠

⎞ ⎜

⎝

⎛

∂

∂

∂

∂

+

∂

∂ − ∂∂ ⎟=− ⎠

⎞ ⎜

⎝

⎛

∂

∂

N

i 1

i

0

r i i

N

i 1

i

pi im

N

i 1

i

pi im

p p p

q W h ω

r

T

r

Y

ρc D

x

T

x

Y

ρc D

r

T

rλ

x r r

T λ

r x

T

ρc v

x

T

ρc u

t

T

ρc

& &

4. Results and discussion

4.1. Validation of the computational results 

The computed flame shape was compared with measurements to validate the 2D model. 

Figure 3 compares photographs of the thin non-luminous flame and computed flame shapes using 

the 2D model for 20 s-1 in both 0-g and 1-g, the flames were near extinction with N2 dilution in 

the fuel stream. 

The high heat release rate region represents the high temperature zone and the peak heat 

release rate can be considered to be near the flame surface. The computed flame shape and the

flame surface location by photographs were similar as show in Figure 3. The photographs in 

Figure 3 show that the flame size, the shape, and the flame surface location of the 1-g flames

70
were reasonably well predicted by the 2D computations for a variety of global strain rate and 

velocity ratio conditions. 

Figure 4 compares the calculated temperature profiles with measurements in a nearextinction 1-g methane-air nonpremixed counterflow flame with ag=20 s-1 with the fuel stream

diluted by N2 (Ca=0.81). Extinction occurs for Ca=0.84. The temperature profile computed by 

the 2D model was in good agreement with previous measurements obtained using thin filament

pyrometry. The flame location was also reasonably predicted by the 2D model. The 2D model 

developed here provides a more complete description of the flame physics as seen in the flame

structure, including the flame size, the shape, the flame surface location and the flame

temperature for a variety of global strain rate and velocity ratio conditions. 

4.2. Extinction limits and suppression characteristics of CO2 addition in fuel stream 

Figure 5 shows a comparison between the measured and computed extinction limits of 

the critical CO2 volume fraction in the fuel stream to extinguish the methane-air diffusion flame 

as a function of the global strain rate for equal fuel and oxidizer flow velocities (Vr=1). The flame 

extinction curve has a turning point in which the critical CO2 volume fraction increased as the 

strain rate decreased, until the volume fraction reached a maximum value near ag=30 s-1 and then 

decreased as the global strain rate increased. The measured critical agent volume fraction of CO2

in the methane fuel stream assuring suppression for all global strain rates was 0.773±0.009, and 

that of 2D model was 0.76 which is raged in the error bar. The 2D model under-predicted for 

smaller ag of 30, and over-predicted for higher ag of 30. 

Figure 6 shows the computed flame shamp of 1-g counterflow for ag=20 with different dilution 

volume fractions of CO2. For the case of Ca=60, the edge of the computed flame reached to the 

outer lip of the burner, but increasing of Ca made to decrease the flame size. Near the extinction 

limit (Ca=66), the flame zone was located near the center axis of the burner. 

Figure 7 show the temperature profile along the center axis of the counterflow flame for 

different CO2 volume fraction in the fuel stream. The maximum temperature decreased and the 

flame position moved up to air duct as the volume fraction of CO2 increased. But, the flame 

thickness and overall flow structure was not significantly varied with increasing of Ca even if the 

flame size and temperature field was changed. 

Table 1 represent the computed fractional contribution to the heat release rate by key terms in the 

energy equation along the flow field centerline for the ag=20. As the volume fraction of CO2

increased, the contribution of radiation and axial convection term increased while that of axial 

71
conduction decreased. The results indicate that axial convection is the largest contribution of the 

heat loss mechanism near the extinction limit. 

Table 1. The computed fractional contribution by each term in the energy equation to the

heat release rate of 1-g, ag=20. 

Ca axial convection axial conduction radiation 

60 0.4532 0.4502 0.0928

65 0.4627 0.4296 0.1042

66 0.4640 0.4225 0.1071

Figure 8 shows the computed flame shape using 2D model for the fuel stream diluted 

with CO2. The flame was wider with increasing the global strain rate, and the flame apart from

the burner tube of air incoming. The flame thickness and curvature decreased as the global strain 

rate increased. As an increase in global strain rate tend to increase the distance between flow 

stagnation point and flame surface at the centerline of the burner. 

For the case of low strain rate, some of the mixture of fuel and CO2 did not rise up to 

flame surface and dispersed out near lower burner tube. This was probably because the inertia of 

the fuel and CO2 mixture is low. But for the case of ag=40, most of incoming flow rose toward the 

flame surface. 

Figure 9 shows the computed local strain rate and the position of maximum heat release 

rate at the centerline of the burner for CO2 volume fraction of 65. The negative maximum value 

of local strain rate is further from the air duct as the global strain rate increased. The flame 

surface location defined by the maximum heat release rate was also moved down as shown in 

Figure 8. The position of the maximum heat release rate was just above the maximum negative 

local strain rate. 

Figure 10 represents the computed fractional contribution of axial convection and 

diffusion term in the energy equation to the heat release for 0.65 CO2 as function of the global 

strain rate along the centerline. The contribution of axial convection fraction increased with 

increasing the global strain rate. The contribution of the axial diffusion fraction is linearly

72
increased with increasing the global strain rate until ag=35. But it was nearly constant in the 

higher global strain rate of ag=80. 

Figure 11 shows the computed fractional contribution of radiative term in the energy

equation to the heat release rate as function of global strain rate near the extinction limit of N2 and 

CO2 suppression. The conditions are Ca=65 for CO2 suppression and Ca=79 for N2 suppression. 

The results show that the radiative energy loss mechanism becoming larger as the global strain 

rate decreases. 

As expected, the fractional contribution of radiative energy loss of CO2 suppression is 

almost twice times higher than that of the N2 suppression. N2 is radiatively inactive gas while 

CO2 is a strong absorber. Some radiation energy emitted near the flame surface can not be 

absorbed by the reactant when the fuel is diluted with CO2. Therefore, the radiative heat loss may

affect on the difference of extinction limit between CO2 and N2 suppression. 

Figure 12 indicate the maximum temperature at the centerline of the burner as function of 

Ca for global strain rate of 20, 30, and 60. The maximum temperature decreased with increasing 

Ca, it has the minimum value at the extinction limits. The maximum temperature near the 

extinction limit has the minimum value around the global strain rate of 30. 

Figure 13 shows the maximum flame temperature near the flame extinction with 2D 

computation and measurements. 2D computation in microgravity shows similar trends with 

experiments in the maximum temperature along the centerline of the burner. For the nitrogen 

agent in microgravity, the maximum centerline temperature had a minimum value near the global 

strain rate of 20 s-1. For the CO2 agent in normal gravity, the trend was similar with N2 in 

microgravity except the turning point slightly moved to global strain rate of 30 s-1. 

SMHRR(Specific Maximum Heat Release Rate) considered to investigate the order of 

flame strength with the volume fraction of CO2. Figure 14 shows the computed maximum heat 

release rate and local strain rate as function of Ca. The maximum heat release rate decreased with 

increasing volume fraction of CO2 while the local strain rate in which the maximum heat release 

rate occurred was almost constant. Accordingly, the SMHRR decreased as the volume fraction of 

CO2 increased. The value of SMHRR for ag=60 was about 2.0 near the extinction limit. 

Figure 15 shows the computed SMHRR with maximum heat release rate and local strain 

rate as function of the global strain rate near the extinction limits. The maximum heat release rate 

and local strain rate had the minimum value around the global strain rate of 30 near the extinction 

limit of CO2 volume fraction. For the higher global strain rate than 30, the maximum heat release 

rate and local strain rate linearly increased as the global strain rate increased. Except the case of 

ag=20, the flame was extinguished under 2.0 of SMHRR. 

73
Figure 16 shows the maximum temperature and heat release rate at the centerline of the 

burner near the extinction limit at ag=20 for the N2 and CO2 , respectively. The maximum heat 

release rate, the maximum temperature and flame location for the CO2 flame suppression were 

similar with that of the N2, but the flame thickness of the CO2 suppression is thinner than that of 

the N2 suppression. 

Figure 17 shows the computed results of the temporal evolution of 1-g flame extinction at 

ag = 30 and Vr=1. The CO2 volume fraction in the fuel stream of 1-g flame was increased at time

equal zero from an initial steady value of 0.74 to 0.78, its value at the extinction limit. The right

hand side of the computational results shows the computed temperature field, while the left hand 

side shows the computed heat release rate and streamlines. Figure 6 (b) shows that the extinction 

of the 1-g flame initiate near the centerline as the increased CO2 volume fraction in the fuel 

stream reaches the flame surface. The flame zone around the centerline becomes weak as the heat 

release rate decreases and the entire flame abruptly extinguishes. 

5. Conclusions 

The structure and extinction of non-premixed flames was investigated through comparison of 

experiments and calculations in a methane-air counterflow flame with CO2 added to the fuel 

stream. The simulations yielded insight into the effectiveness of CO2 as an agent, highlighting 

differences in the radiative loss when CO2 rather than N2 is used as a suppressant. The 

differences between microgravity and normal gravity suppression results were modeled. 

Measurements and computations of the maximum flame temperature near the flow field 

centerline for near-extinction methane-air flames diluted by CO2 show behavior analogous to N2

diluted flames in which the maximum temperature has a minimum value between strain rates of 

20 s-1 and 30 s

-1. 

The analysis, using flame simulations, allows insight of the heat transfer processes that 

control flame structure and extinction. At low strain rates, the flames are rather thick and lateral 

heat and diffusion losses are not negligible. By providing physical insight, two-dimensional 

simulations are a useful tool for guiding the interpretation of counterflow flame experiments. 

74
6. References 

[1] Bundy M, Hamins A, Lee K Y 2003 Combust. Flame 133 299-310 

[2] Bundy M, Park J, Oh C B, Hamins A, “Extinction and Temperature Measurements in 

Nonpremixed Methane Flames in Microgravity,” submitted to Combustion Theory and Modeling, 

2006. 

[3] Seshadri K and Williams F A 1978 Int. J. Heat Mass Transfer 21 251-253 

[4] C.B. Oh, C.E. Lee, J. Park, Combust. Flame, 138 (2004) 225-241. 

[5] B.P. Leonard, Comput. Meth. Appl. Mech. Eng.,, 19 (1979) 59-98. 

[6] H.N. Najm, P.S. Wyckoff, Combust. Flame, 110 (1997) 92-112.

[7] C.W. Hirt and J.L. Cook, J. Comput. Phys., 10 (1972) 324-340.

[8] R.J. Kee, F.M. Rupley, J.A. Miller, CHEMKIN-II: A Fortran Chemical Kinetics Package for 

the Analysis of Gas Phase Chemical Kinetics, Report No. SAND89-8009B, Sandia National 

Laboratories, 1989. 

[9] R.J. Kee, G. Dixon-Lewis, J. Warnatz, M.E. Coltrin, J.A. Miller, A Fortran Computer Code 

Package for the Evaluation of Gas-phase Multicomponent Transport Properties, Report No. 

SAND86-8246, Sandia National Laboratories, 1986. 

[10] F.L. Dryer, I. Glassman, Proc. Combust. Inst., 14 (1972) 987-1003. 

[11] A. Lutz, R.J. Kee, J. Grcar, J., F.M. Rupley, “A Fortran Program Computing Opposed Flow 

Diffusion Flames”, SAND96-8243, Sandia National Labs, Livermore (1997). 

[12] Smith, G. P., Golden, D. M., Frenklach, M., Moriatry, N. W., Eiteneer, B., Goldenberg, M, 

Bowman, C. T., Hanson, R. K., Song, S., Gardiner, W. C., Lissianski, V. and Qin, Z., 

http://www.me.berkeley.edu/gri_mech/. 

[13] Y. Ju, H. Guo, K, Maruta, F, Liu, J. Fluid Mech., 342 (1994) 315-334. 

75
Figure 1. Schematic drawing of counterflow burner. 

Figure 2. Schematic the computational domain. 

Side B

g 

Computational domain 

x R=11.7 mm

Side C

0

r 

 Fuel 

(CH4+agent)

Air 

L=25 mm

Lp=1.2 mm

W=3.2 mm 

Side A

Flame 

76
 (a) 1-g (b) 0-g 

Figure 3. Comparison of flame shape in normal gravity and microgravity for methane-air nonpremixed 

flames at ag=20 s-1 and a fuel stream N2 volume fraction of 0.80. The left hand sides are photographs,

while the right hand sides are simulation results. 

Figure 4. Measured and computed temperature profiles for ag=20 s-1 with 0.81 volume fraction of CO2 in

the fuel stream and Vr=1. 

77
Figure 5. Computed and measured critical CO2 volume fraction in the fuel stream (Ca) as a function of the 

global strain rate for suppression of CH4 flames in normal gravity. 

78
 (a) Ca= 0.60

 (b) Ca= 0.65

 (c) Ca=0.66 

Figure 6. Computations of a normal gravity counterflow nonpremixed flame diluted with CO2 in the fuel 

stream for the global strain rate of 20 s-1. Calculated temperatures are shown on the left-hand side of the 

figures and the heat release rate and streamlines are shown on the right-hand side. 

79
Figure 7. The computed local heat release rate for different CO2 volume fractions in the fuel stream (0.60, 

0.65, and 0.66) at a global strain rate of ag=20 s-1. 

80
a) ag= 20 s

-1

 (b) ag = 30 s

-1

 (c) ag = 40 s

-1

Figure 8. The calculated structure of a normal gravity counterflow nonpremixed flame with a CO2 volume 

fraction of 0.65. The temperature is shown on left-hand side of the figure; the local heat release rate and 

streamlines are shown on the right-hand side. 

81
Figure 9. Computed local strain rate along the centerline for a CO2 volume fraction of 0.65 in the fuel

stream with Vr=1 and global strain rates of 20 s-1, 30 s

-1, and 40 s-1. 

Figure 10. Computed fractional contribution by each term in the energy equation to the heat release of a 1-

g flame with Ca=0.65 (agent = CO2) as a function of the global strain rate along the flow field centerline. 

82
Figure 11. Comparisons of computed radiative fraction for N2 and CO2 for near-extinction 1-g flames as a 

function of the global strain rate along the flow field centerline. (Ca=0.65 for CO2 and Ca=0.79 for N2) 

Figure 12. Maximum temperature along the flow field centerline as a function of the CO2 volume fraction

in the fuel stream. 

83
Figure 13. Measured and computed maximum flame temperature near the flow field centerline as a 

function of the global strain rate for near-extinction methane-air flames diluted by N2 and CO2. 

Figure 14. Computed maximum heat release rate, the local strain rate, and the specific maximum heat 

release rate (SMHRR) as a function of volume fraction of CO2 in the fuel stream for ag=60. 

84
Figure 15. Computed maximum heat release rate, the local strain rate, and the specific maximum heat 

release rate (SMHRR) as function of global strain rate near the extinction limits (ag=20, Ca=0.66, ag =30,

Ca=0.75, ag =40, Ca=0.75, ag =60, Ca=0.74, ag =80 Ca=0.74, ag =100, Ca=0.73, ag =120, Ca=0.72) 

Figure 16. Comparison of centerline temperature and heat release rate between N2 and CO2 suppression 

near the extinction limits at ag=20 s-1. 

85
(a) t = 0.0 s 

 (b) t = 0.14 s 

 (c) t = 0.15 s 

 (d) t = 0.16 s 

Figure 17. Computed transient extinction of a normal gravity counterflow nonpremixed flame diluted by

CO2 in the fuel stream for ag=30 s-1 and Vr=1 with Ca increasing from 0.74 to 0.78. The flame temperature 

is shown on the right-hand side and the heat release rate and streamlines are shown on the left-hand side. 

 

86
Appendix 4

Suppression Limits of Low Strain Rate 1G Non-Premixed Propane Flames 

ABSTRACT 

The suppression of low strain rate non-premixed normal gravity propane flames was investigated 

experimentally in a counterflow configuration. Conductive heat losses were minimized by

varying the velocity ratio of fuel to oxidizer to adjust the flame position such that conductive 

losses to the burner were reduced. This was confirmed by temperature measurements using 

thermocouples (0.05 mm wire diameter) near the reactant ducts. The critical concentration of N2 , 

CO2, and CF3Br added to the fuel and the critical concentration of N2 added to the oxidizer stream

required to obtain extinction of propane-air non-premixed flames was measured as a function of 

the global strain rate. In agreement with previous measurements performed with methane-air 

diffusion flames, limiting non-premixed flame extinction behavior in which the agent 

concentration obtained a value that insures suppression for all global strain rates was observed. 

A series of extinction measurements varying the air: fuel velocity ratio showed that the critical N2 

concentration was invariant with this ratio, unless conductive losses were present. In terms of fire 

safety, the measurements confirm the existence of a fundamental limit for suppressant 

requirements in normal gravity flames as was previously seen in methane-air flames. This can be 

thought of as analogous to agent flammability limits in premixed flames. The critical agent 

volume fraction in the propane fuel stream assuring suppression for all global strain rates was 

measured to be .923±0.003 for N2, 0.879±0.004 for CO2, and 0.517±0.01 for CF3Br. These 

results were distinct from the methane results. The critical agent volume fraction in the oxidizer 

stream assuring suppression for all global strain rates in the propane flames was measured as 

0.354±0.003 for N2.. 

 

Measurements using thin filament pyrometry measures flame temperatures for diluted propane-air 

flame near extinction for several different strain rates. Insight was gained into the near extinction 

flame temperatures as a function of strain rate. 

87
INTRODUCTION 

The agent concentration required to achieve the suppression of low strain rate 

non-premixed flames is an important consideration for fire protection applications as 

engineers rely on suppression data to estimate suppressant mass requirements for total 

flooding applications. Information regarding critical suppressant concentrations in the 

high strain rate regime is less important than in the low strain rate regime, because low 

strain rate conditions typically require higher agent concentrations and fire protection 

design must be based on worst-case conditions. Currently, there is a scarcity of 

experimental data on the effectiveness of suppressants in low strain rate non-premixed 

flames. 

 Low strain rate agent suppression requirements in counterflow flames are also of 

interest because they correspond to agent suppression requirements in buoyancydominated flames, such as cup burner flames [1]. The burner of choice for testing agent 

suppression effectiveness has traditionally been the cup burner. In that burner, the flame 

strain rate is not controlled and preheating of the fuel cup can lead to flame stabilization. 

A study comparing inert and halocarbon suppressant effectiveness in the counterflow and 

cup burner geometries was performed by Saso et al. [2]. Their results confirmed the 

results of Hamins et al. [1], showing a correspondence in flame extinguishment 

concentration between the cup burner and the counterflow burner at a global strain rate of 

30 s-1. 

The counterflow geometry is a convenient configuration for direct control of the 

strain rate and flame position [3]. The effect of flame position on suppressant 

requirements has been considered in a number of counterflow flame suppression studies. 

88
In some counterflow studies, the fuel and oxidizer reactant streams were adjusted such 

that the flame position is maintained [4]. Flame position has been shown to effect 

suppressant requirements if the flame is too close to the reactant ducts, which was 

attributed to conductive rather than convective heat losses. 

The effectiveness of chemically acting inhibitors on methane-air counterflow 

diffusion flames was investigated by Vora et al. [5] and Siow et al. [6]. In each of these 

studies the suppressant effectiveness was related to the measured OH concentration in 

flames with a global strain rate of 20 s-1. 

The first comprehensive extinction measurements of very low strain nonpremixed flames were reported by Maruta et al. [7], who performed the experiments 

under micro-gravity conditions. In that study, the extinction of methane-air flames with 

N2 added to the fuel stream was investigated using the JAMIC 10 s drop tower. The fuel 

stream nitrogen concentration required to achieve extinction was measured to increase as 

the strain rate decreased, until a critical value was obtained. As the global strain rate was 

further reduced, the required N2 concentration decreased. This behavior was denoted as a 

"turning point" and was attributed to the enhanced importance of radiative losses at low 

strain rates. Numerical calculations were performed to quantify the radiative heat loss

effects. The fraction of radiative heat loss to total flame enthalpy was found to be greater 

than 10% for near extinction flames with a global strain rate less than 4 s-1. Several other 

works to date have established the importance of radiative heat loss on the stability of 

weakly strained diffusion flames [8-10]. 

Recent work at the National Institute of Standards and Technology studied the 

extinction characteristics of methane-air diffusion flames at low strain rates in 1-g. The 

89
turning point phenomenon was observed for 1-g methane flames as well as the μg 

methane flames. With the exception of the previous work performed by Bundy et al [11], 

the studies that have considered suppression by an agent in normal gravity counterflow 

flames have not carefully investigated the low strain rate region (see Ref. [1] for 

example). There are experimental difficulties associated with establishing steady flames 

at low strain rates in the laboratory including flame instabilities associated with ambient 

currents and forced exhaust. 

The objective of this work was to study the suppression of low strain rate nonpremixed propane flames in normal gravity through the establishment of steady flames 

free from significant conductive losses to the burner ducts. In the experiments described 

here, gaseous N2 was used as a suppressant added to both the oxidizer and fuel streams to 

investigate whether turning point behavior occurs in suppression of normal gravity 

propane flames, as previously observed in normal gravity methane flames [11]. In 

addition, the suppression effectiveness of CO2 and CF3Br are also reported for flames 

with agent added to the fuel streams of propane-air diffusion flames. 

EXPERIMENTAL METHOD 

Experiments were conducted using the counterflow burner shown in Figure 1. 

The inner diameter of the reactant ducts was (23.4 ± 0.1) mm and the duct separation 

distance (L) was set to (25.0 ± 0.1) mm. Four 200 mesh stainless steel screens were 

positioned (1.6 ± 0.1) mm from the duct exits and at 1.6 mm intervals to impose a tophat velocity profile. A water-cooled jacket about the top reactant duct prevented heating

of the metal burner and preheating of the reactants. The reactants were air and research 

90
grade propane (99 % purity), which were stored in pressurized cylinders. The oxygen 

content in the bottled air was measured to be (20.93 ± 0.04) % using a paramagnetic 

analyzer [12]. The reactant flows were controlled using mass flow controllers that were 

calibrated using a dry cell primary flow meter with an uncertainty of 1%. The central

axis of the burner was aligned vertically with gravity. 

The burner was fitted with a concentric tube (64 mm o.d.) about the bottom duct

for flow of an inert gas to shield the reactants from ambient air. Experiments were 

performed using the “curtain” flow of nitrogen to isolate the reactants from the ambient 

environment. The average velocity of the curtain was tuned to approximately match the 

fuel flow from the lower duct; however the magnitude of the curtain flow rate had a

negligible effect on the extinction results for curtain flows in the range of 20 cm3

/s to 80 

cm3

/s. Several studies have explored the impact of the composition of the ambient 

environment on flame stability as exemplified by the concentration of suppressant 

required to achieve extinction on flame stability as exemplified by the concentration of 

suppressant required to achieve extinction. Some studies have used an air ambient [7], 

whereas others have used an inert environment [12]. In previous work done on 1-g 

methane extinction [11] it was shown that as the curtain composition was varied from 

pure N2 to pure air, the N2 concentration in the fuel stream required to achieve extinction 

increased significantly. This suggested that the enhanced stability associated with the air 

curtain was due to excess fuel reacting with the oxygen in the curtain flow; therefore tall 

measurements were performed using a pure N2 curtain. 

A number of previous studies investigating the structure and extinction of gaseous 

diffusion flames applied a flow of nitrogen [13,14] about one of the reactant streams (i.e., 

91
a “curtain” of nitrogen), while applying suction through an exhaust section about the 

opposing reactant stream. Use of a forced exhaust was found to create turbulence in low 

strain rate flames. Instead, the burner was isolated from ambient disturbances by placing 

it in a nearly closed compartment (0.5 m × 0.5 m × 0.6 m) with a round 10 cm diameter 

exhaust port at its top. The compartment was placed inside a chemical hood, allowing 

controlled removal of combustion byproducts through buoyancy-driven ventilation. The 

annular top exhaust section of the burner was not used for these experiments. 

In the experiments described here, the global strain rate (ag) was varied from 12 s1

 to 70 s-1. The global strain rate concept allows quantification of a characteristic flame 

residence time or flow time. Suppression measurements were performed by 

incrementally increasing the agent flow, while maintaining a constant global strain rate. 

This was accomplished by simultaneously reducing the air or fuel flow. The global strain

rate is defined as: 

, (1) 

which is a factor of two larger than the definition used by Maruta et al. [7] or Puri and 

Seshadri [14]. The parameters V and ρ denote the velocity and density of the reactant

streams at the duct boundaries, L is the duct separation distance, and the subscripts a and 

F represent the oxidizer and fuel streams, respectively. The parameter Vr is defined as 

equal to Va/VF. The expression given by Eq. (1) was derived by Seshadri and Williams

[15] to describe global strain rate for counterflow flames. Chelliah et al. [16] used Eq. 

(1) to compare global and local oxidizer strain rates for undiluted propane-air flames near 

extinction. Pogliani et al. [17] performed a numerical investigation of flame stretch and 

92
radiation on flame extinction at low strain rates. Their calculations showed that there is a 

proportional relationship between the global and oxidizer side strain rates. Zegers et al. 

[18] used LDV measurements to determine local strain rates in non-premixed methane/air 

and methane/air flames. Their results also show a linear relationship between the 

maximum velocity gradient on the oxidizer side of the flame and the global strain rate 

defined by Eq. 1. 

The flame extinction measurements conducted here were repeated at least four 

times. Experiments were preformed with velocity ratios of Vr = 1⁄2, Vr = 1, Vr = 2 and Vr = 

3. The Vr = 1⁄2 and Vr = 1 tests where performed with the fuel jet from the top burner duct 

and the air jet from the bottom burner duct and the Vr = 2 and 3 cases were tested with the 

opposite orientation. The mean value of the relative expanded standard uncertainty in the 

agent extinction concentration was typically 0.8 % based on repeat measurements and a 

propagation of uncertainty analysis. All uncertainties reported here are for a coverage 

factor of one, which is the combined standard uncertainty [19]. Measurement 

repeatability typically accounted for approximately one-quarter of the uncertainty. 

Changing the velocity of the reactant streams led to changes in flame position and 

flame shape. Flame shape was characterized in terms of flame curvature through the use 

of digital photography and image analysis. The curvature was defined as the reciprocal 

of the average radius of a circle associated with the luminous flame in the cylindrical 

zone between the ducts. The analysis of curvature was based on three points defined by 

the intersection of the lower edge of the luminous flame with (1) the central axis and (2) 

the segments connecting the inner walls of the top and bottom burner ducts. 

93
Temperature measurements were conducted along the centerline near the burner 

ducts using thermocouples. Heat transfer from the flame to the burner is indicated by the 

temperature gradient near the burner ducts. The temperature measurements were 

conducted using small thermocouples with the wires aligned horizontally along an 

isotherm to minimize conductive losses. Measurements were made using 0.05 mm (0.002 

in) diameter (Pt/Pt + 10 % Rh) S-type thermocouples. Strain rates at which significant 

conductive losses were found were removed from the final results. 

Temperature measurements were performed for T ≥ 1250 K by measuring the 

luminous intensity of a thin filament positioned in the flame. This technique was 

originally developed by Vilimpoc et al. [20] and is detailed in several publications on 

laminar diffusion flames [21,22]. A (12.5 ± 2) micron SiC filament was aligned axially 

between the burner ducts at various radial locations. The radiative energy emitted from 

the filament was digitally recorded using a CCD camera and close-up lens. The spatial 

resolution of the image was 0.047 mm. The image luminous intensity was digitized to 8 

bit resolution, this represents a 2°K resolution. The exposure setting on the camera was 

adjusted so that the image was not saturated (over-exposed) at the maximum flame 

temperature. Image processing software was used to acquire the fiber intensity data 

from the digital image. The emission intensity at each axial location was found by 

subtracting the background intensity from all pixels and integrating the radial intensities 

across the fiber. 

For a specific set of flame conditions the measured fiber intensity profile was 

compared to the calculated fiber temperature profile to determine the relationship

between fiber emissions and fiber temperature following the method of Vilimpoc et al. 

94
[20] OPPDIF provided the calculated gas temperature profile for a given set of 

conditions. The OPPDIF code uses multi-step chemistry and a narrowband radiation 

model and has been validated for predicting temperatures in counterflow flames. The 

radiation model is described in detail by Pogliani et al. [17]. The OPPDIF gas 

temperature profile was used to calculate the heat transfer to the fiber and the fiber 

temperature profile given the material properties of the fiber. The fiber temperature was 

then compared to the measured fiber luminosity and a correlation was determined with 

luminosity being proportional to the fiber temperature to the fifth power. The 

relationship between fiber temperature and illuminosity was recalibrated for each set of 

tests to take into account any changes in the camera settings or fiber position. 

The SiC fiber was used because it has a known emissivity of 0.88 over a wide 

range of wavelengths [20] and a low conductivity allowing point measurements along the 

entire length of the fiber. [23] Since the conductivity is low, it can be neglected and the 

energy balance becomes 

h P (Tg – Tf) = σεP Tf

4 (2) 

where h is the convection coefficient, Tg is the gas temperature, Tf is the fiber 

temperature, P is the perimeter of the fiber, σ is the Stefan-Boltzman constant, and ε is 

the fiber emissivity. Solving for Tg a simple radiative correction for the fiber results: 

Tg = Tf + σε Tf

4

/ h (3) 

The convection coefficient was found by using a numerical correlation for low velocity 

cross flow over a cylinder. [17] Experiments compared the filament intensity with the 

fiber oriented horizontally and vertically at several locations in the flow field to 

95
determine if there was a significant difference caused by fiber orientation [24]. The 

differences found between the two orientations were within experimental error and 

showed that the crossflow assumption for the convection coefficient was adequate. 

Results and Discussion 

Observations 

The flames were laminar, steady and the suppression measurements were highly 

repeatable for strain rates as low as 15 s-1 depending on the agent type and configuration. 

Below that value, the conductive losses from the flame to the burner become significant. 

Undiluted, low strain rate counterflow propane-air flames are yellowish in color. As N2

and CO2 were added to either the fuel or oxidizer streams, flame luminosity decreased 

until flame emission was entirely blue in color near extinction. This was not the case 

when CF3Br was added to the flames. In that case, the flames became increasingly

luminous as more agent was added. These flames were highly luminous near extinction 

and large amounts of soot were observed to escape from the flame. This is similar to 

observations made in methane/air flames [11]. 

Figure 2 shows the measured position of the central axis of the luminous flame 

zone as a function of the global strain rate for several values of the velocity ratio (Vr) 

when the air stream was flowing from the top duct. In these experiments, the fuel side jet 

contained 98% of the extinction value of N2 in the flow for a given strain. The results in 

96
Fig. 2 show that the flames were lifted towards the top duct as the global strain rate was 

decreased due to the effects of buoyancy. When the propane was injected form the top 

jet, Vr = 1⁄2 and Vr = 1, the curvature appears to reach a maximum in the ag= 20 –25 s-1

range at a curvature of roughly 0.4 cm-1. When the fuel was injected from the bottom jet 

the curvature appeared to increase linearly with decreasing strain rate. At higher strains 

the curvature decreases with increasing strain rate. The magnitude of the difference in 

the curvature between velocity ratios seems to increase with increasing strain rate. The 

flame position also increased linearly with increasing strain rate. Increasing the velocity 

of the top jet compared to the bottom jet pushed the flame further from the top duct. The 

results also showed that as Vr was increased from 1 to 3, the luminous flame shifted away 

from the top duct and the flame curvature generally decreased. Through adjustment of 

the value of Vr, the suppression measurements reported here are for flames that were 

sufficiently far from the ducts such that conductive heat losses were negligible. 

Conductive Heat Losses to the Burner 

Thermocouple measurements were conducted to characterize the temperature 

field near the top reactant duct as Vr was varied. The most problematic conditions in 

terms of conductive losses are the lowest strain rate flames, which have the broadest 

temperature profiles and are located closest to the top duct as seen in Fig. 2. To 

determine the losses to the burner for the 1-g tests for different velocity ratios and global 

strain rates, the temperature was measured along the centerline near the burner ducts 

using thermocouples as conductive heat transfer from the flame to the burner is indicated 

by the temperature gradient near the burner ducts. The results are shown in figures 4 and 

97
5. Measurements were made using 0.050 mm (0.002 in) diameter (Pt/Pt + 10 % Rh) Stype thermocouples. The distance from the thermocouple to the orifice was determined 

by taking an image of the thermocouple at each measurement location and finding the

pixel distance from the thermocouple to the duct and the distance per pixel for the 

camera. The rate of energy lost to the duct ring due to conductivity was estimated. In 

order to decide if the losses were significant, the energy loss per unit time due to 

conduction to the rig was equated to the uncertainties in the flow rate. The uncertainty in 

the flow rate of the diluent corresponds to an uncertainty in the energy loss from the 

flame. The uncertainty in the energy loss from the flame is equal to the energy it would 

take to raise the volume of diluent that represents the possible error in the flow to the 

flame temperature. This value was calculated and compared to the conductive losses for

each strain rate. If the conductive losses were less then the energy loss rate uncertainties 

due to the uncertainties in the flow rates, then it was assumed that the conductive losses 

were negligible, if not then the data point was removed from this analysis. The 

conductive losses were considered significant for ag ≤ 17 s-1 when Vr = 1 and for ag < 13 

s

-1 when Vr = 2, 0.5, 3, and 0.3. 

Flame Temperature Measurements 

Flame temperatures were attained by measuring the luminous emission from a 

12.5 micron SiC filament place axially between the burner ducts. Low strain rate flames

near extinction were established to observe the effects of curvature and air-fuel velocity 

98
ratio on flame structure. Previous work showed that the flame position, but not the flame 

temperatures, changed with the fiber distance from the centerline. [11] This suggests that 

even though the flames are curved by buoyancy, the local agent concentration required 

for suppression will not differ significantly along the flame sheet. Further evidence that 

the flame curvature does not significantly affect the stability of these flames is given by 

the extinction results described below. Previous work also indicated that the flame 

structure is not significantly impacted as the velocity ratio is varied over the small range 

of velocity ratios used in this study. [11] 

Temperature measurements were taken for a certain percentage of the extinction 

diluent concentration present in the flow. The temperature data presented here is for the 

nitrogen-diluted flames only. Temperature data was taken for Va/Vf = 1 at 98% and 

99.5% of extinction and for Va/Vf = 2 at 98% of extinction. The maximum temperature 

for each strain rate and flow condition is shown in figure 5. The same trend can be seen 

for all conditions. At higher strain rates the near extinction temperature decreases with 

decreasing strain rate until a minimum extinction temperature is reached, then as strain 

rate decreases, near extinction temperatures increase. When comparing the graph to the 

extinction data, it is seen that the minimum temperature correlates with the maximum 

diluent concentration. The flame is most stable at the turning point and this is when 

flames can be stabilized at the lowest temperatures. As the percentage of extinction 

diluent in the flow is increased, the flame temperature decreases but the shape of the 

curve remains the same. When comparing the different velocity ratios it can be seen that 

shapes of the curves are similar but the temperatures at 98% of extinction for the Va/Vf = 

2 case are slightly lower then the temperatures for the Va/Vf = 1 case. The Va/Vf = 2 

99
case has a fuel jet exit velocity that is less then the Va/Vf = 1 case and the air jet velocity 

is two times larger than the fuel velocity. The changes in these parameters affect the 

reaction rate and the thermal and mass diffusivities at the flame surface. When the 

velocity ratio is equal to one it appears that there is slightly more heat generation then for 

a velocity ratio of two. The differences in the extinction measurements taken for the two 

velocity ratios are insignificant, but the different flow configurations do apparently have 

an effect on the energy production due to combustion. 

Flame Extinction Measurements 

Previous work has shown the Vr can be treated as an independent variable that 

does not have a large effect on the amount of diluent needed for extinction in a given 

strain rate methane-air non-premixed flame. [11] Extinction measurements done in this 

study also confirm that Vr can be treated as an independent variable for low strain rate 

propane-air non-premixed flames. Figure 7 shows the measured N2 volume fraction in 

the fuel stream required for extinction as a function of the global strain rate for a number 

of values of Vr and two different burner orientations for flames without significant 

conductive heat losses. Conditions below the data points represent stable flames, 

whereas conditions on and above the data represent flames that have been extinguished. 

Error bars representing the expanded combined standard uncertainty are shown for all 

points. Turning point behavior was exhibited, in which the critical suppressant 

concentration increased as the strain rate decreased, until the concentrations reached a 

maximum value near ag = 30 s-1 and then decreased as the global strain rate further 

100
decreased. For a particular value of the global strain rate, the agent suppression 

concentration did not significantly change as the velocity ratio was varied. 

Flame curvature can affect flame stability, although in counterflow diffusion 

flames the radius of curvature must be extremely small before there is a noticeable effect. 

Finke and Grunefeld [25] found measurable differences in the extinction strain rate of 

hydrogen diffusion flames with a radius of curvature less than 10 mm. The smallest 

radius of curvature of the flames in this study was 18 mm. Figure 2 shows that the 

largest absolute difference in flame curvature as Vr was varied from 1/2 to 3 was 

observed for a global strain rate of approximately 40 s-1. The measurements in Fig. 7 

show negligible differences in N2 suppression requirements at ag = 40 s-1 as the velocity 

ratio was varied from Vr = 1/2 to Vr = 3, suggesting that differences in flame curvature 

had a negligible effect on agent extinction requirements and flame stability for the flames 

considered in this study. 

Figure 8 shows the comparison of the 1-g methane extinction curve [11] and the 

1-g propane extinction curve. In the 1-g tests the turning point occurred at lower strain

rates for the propane then for the methane. The curve of the turning point behavior of the 

nitrogen extinction of propane was also much shallower then that of the nitrogen 

extinction of methane as can be seen in figure 8. The reason for the differences in turning 

point locations for the two fuels is most likely due to differences in flame structure and 

radiative emissions. The main cause for heat loss from the flame in 1-g is convective and 

radiative losses. Different fuel masses cause different exit momentums from the jets 

affecting the convective heat transfer and the relative importance of the losses due to 

buoyancy. The conductivity of the gases passing through the flame front also affects the 

101
convective heat transfer and the flame temperature based on the fuel chemistry affects the 

buoyancy forces and the radiative losses. These effects all play a role in determining the 

location of the turning point. For the propane, the conductive losses appear to be more 

significant then the radiative losses at lower global strain rates then for the methane. 

Figure 9 shows suppression measurements of the critical CO2, and CF3Br 

concentration in the fuel stream of non-premixed C3H8-air flames as a function of the 

global strain rate. The averaged results for N2 from Figure 7 are included in this figure. 

As expected, the most effective suppressant was CF3Br, followed by CO2 and then N2. 

The critical suppressant concentration increased as the strain rate decreased, with its 

value leveling off near 30 s-1, except for CF3Br, which flattened near 40 s-1. The results 

demonstrate the existence of turning point behavior for both inert and chemically active 

agents in normal gravity diffusion flames when agent is added to the fuel stream. The 

turning point occurs at lower values of ag than in the 1-g methane measurements [11].

Figure 10 shows measurements of the N2 volume fraction in the oxidizer stream

required to achieve extinction as a function of the global strain rate. Similar to the results 

for agent added to the fuel stream (Fig. 8), these measurements also show turning point 

behavior. The turning point occurs within the same range of global strain rates, at values 

between 30 s-1 to 35 s-1. 

Masri et al. [26] suggest that the relevant parameter when comparing agent 

effectiveness in fuel-side and air-side dilution should be the effective mass fraction. The 

effective mass fraction is defined as the agent mass fraction that would exist at the 

theoretical stoichiometric flame sheet location in cold flow, neglecting preferential 

diffusion. 

102
For a propane-air flame, the effective mass fraction is given by, Yeff,st = Ya,f + 

(Ya,o - Ya,f)ξst , where ξst = [1 + 3.636 YC3H8,f / YO2,o]

-1. Yeff,st at extinction for both air 

and fuel stream dilution is shown in Figure 10. The fuel-side dilution extinction 

experiments from Figure 9 are represented by the closed symbols in Figure 11 and the 

air-side dilution extinction experiments from Figure 10 correspond to the open symbols 

in Figure 10. For each of the three different agents, the air-stream agent addition required 

a greater amount of agent by mass for extinction than for fuel-stream agent addition. 

Trees et al. [27] attribute differences in air-stream and fuel-stream agent effectiveness to 

flame structure and preferential diffusion effects. 

SUMMARY AND CONCLUSIONS 

In summary, the range of global strain rates typically investigated in normal 

gravity was extended by isolating the burner in order to reduce disturbances by ambient 

currents and by varying the velocity ratio of fuel and oxidizer to reduce conductive losses 

to the burner. By treating Vr (=Va/Vf) in Equation 1 as an independent experimental 

variable, it is possible to adjust the flame location and thereby minimize conductive heat 

losses from the flame to the burner. A flame free of conductive losses is relatively more 

stable than a flame with conductive losses. Such a flame will require larger agent 

suppression requirements. Whereas flame position and the location of the stagnation 

point may change as the parameter Vr is varied in these flames, the suppression and 

maximum flame temperature measurement results suggest that flame stability is 

essentially unchanged. 

103
The critical concentration of N2, CO2, and CF3Br added to the fuel stream and the

critical concentration of N2 added to the oxidizer stream required to obtain extinction of 

propane-air non-premixed flames was measured as a function of the global strain rate. 

In terms of fire safety, the measurements demonstrate the existence of a fundamental 

limit in the suppressant requirements in normal gravity diffusion flames, analogous to 

agent flammability limits in premixed flames. A fundamental distinction between 

premixed flammability limits and non-premixed flame suppression limits is that the value 

of the limiting concentration in a diffusion flame will obtain different values depending 

on whether the suppressant is added to the oxidizer stream or the fuel stream. The

consideration of configuration does not arise in a premixed flame. The critical agent 

volume fraction in the propane fuel stream assuring suppression for all global strain rates 

was measured to be 0.923±0.003 for N2, 0.879±0.004 for CO2, and 0.517±0.001 for

CF3Br. The critical agent volume fraction in the oxidizer stream assuring suppression for 

all global strain rates was measured as 0.354±0.003 for N2. 

The structure of normal and micro gravity flames at moderate strain rates can be 

expected to be similar, as buoyancy effects are relatively unimportant for momentum

dominated flames. At low strain rates, however, significant differences in flame structure 

are anticipated. The ability to work with low strain rate flames in normal gravity enables 

experimentation at reduced cost as compared to microgravity flames. Further work is 

ongoing to determine the differences in the structure of low strain rate non-premixed 

flames under normal and microgravity conditions. 

104
ACKNOWLEDGEMENTS 

This research was partially supported by the NASA Microgravity Research 

Division through Contract No. C–32066–T (NIST) with Dr. Sandra L. Olson serving as 

Technical Monitor. The authors are grateful to Professor I.K. Puri of the University of 

Chicago at Illinois for helpful discussions. 

105
REFERENCES 

1. Hamins, A., Trees, D., Seshadri, K., & Chelliah, H., Combust. Flame, 99:221 (1994). 

2. Saso, Y., Saito, N., Chihong, L., and Ogawa, Y., Fire Safety Journal, 26:303 (1996). 

3. Tsuji, H., Prog. Ener. Combust. Sci. 8:93-119(1982). 

4. Macdonald, M.A., Jayaweera, T.M., Fisher, E.M, and Gouldin, F.C., Combust. Flame

116:166-176 (1999). 

5. Vora, N., Siow, J.E., and Laurendeau, N.M., Combust. Flame 126:1393 (2001). 

6. Siow, J.E., and Laurendeau, N.M., Combust. Sci. Technol. 174:91(2002). 

7. Maruta, K., Yoshida, M., Guo, H., Ju, Y., & Niioka, T., Combust. Flame, 112:181 

(1998). 

8. Liu, F., Smallwood, G.J., Gülder, Ö.L., and Ju, Y., Combust. Flame, 121:275 (2000). 

9. T’ien, J.S., Combust. Flame, 65:31 (1986). 

10. Olsen, S.L., T’ien, J.S., Combust. Flame, 121:439 (2000). 

11. Bundy, M.,Hamins, A., Lee, K.Y., Combustion and Flame, 103:299 (2003) 

12. Kelly, M., NIST, Report of paramagnetic analyzer test results, personal

communication, March 20, 2000. 

13. Yang, M.H., Hamins, A. & Puri, I.K., Combust. Flame 98:107-122 (1994). 

14. Puri, I.K., & Seshadri, K., Combust. Flame, 65:137-150 (1986). 

15. Seshadri, K., & Williams, F.A., Int. J. Heat Mass Transfer, 21:251 (1978). 

16. Chelliah, H.K., Law, C.K., Ueda, T., Smooke, M.D., and Williams, F.A., Proceedings 

of the Twenty-Third Symposium (International) on Combustion, The Combustion 

Institute, Pittsburgh, 1990, 503-511.

17. Pogliani, B., Bundy, M., Hamins, A., Puri, I.K. Proceedings of the Second Joint 

Meeting of the US Sections of the Combustion Institute, Oakland, 2001. 

18. Zegers, E.J.P., Williams, B.A., Fisher, E. M., Fleming, J.W., and Sheinson, R.S., 

Combust. Flame, 121:471 (2000). 

19. International Organization of Standardization, Guide to the Expression of Uncertainty

in Measurement, Geneve, Switzerland, 1993.

20. Vilimpoc, V., Goss, L.P., and Sarka, B., Opt. Lett. 13:93-95 (1988). 

21. Ravikrishna, R.V., and Laurendeau, N.M., Combust. Flame., 120:372-382 (2000). 

22. Pitts, W.M., Proceedings of the Twenty-Sixth Symposium (International) on 

Combustion, The Combustion Institute, Pittsburgh, 1996, 1171-1179.

23. Gross, Larry P., Weaver, William, L., and Trump, Darryl D. Thin Filament 

Velocimetry and Pyrometry Innovative Scientific Solutions Inc. 

106
24. Logue, J.: “Characteristics of Low Strain Near Extinciton Flames in mg and 1-g”; 

Master’s Thesis, University of Maryland, Department of Fire Protection Engineering 

(2003) 

25. Finke, H., & Grunefeld, G., Proceedings of the Combustion Institute, 28:2133 (2000).

26. Masri, A.R., Combust. Sci. Technol. 96:189(1994).

27. Trees, D., Grudno, A., Seshadri, K., Combust. Sci. Technol. 124:311(1997).

28. Pogliani, B., Puri, I.K., and Hamins, A., “Effect of Thermal Radiation on the 

Extinction of Diluted Counterflow Nonpremixed Flames,” presented at the Western 

States Section Technical Meeting of the Combustion Institute, Salt Lake City, Utah,

October 15-16, 2001.

29. Pandya, T.P., and Srivastava, N.K., Combust. Sci. Technol. 11:165(1975).

30. Sheinson, R.S., Penner-Hahn, J.E., and Indritz, D., "The Physical and Chemical 

Action of Fire Suppressants," Fire Safety Journal 15, 43 (1989).

107
Reactants 

Exhaust 

Figure 1 Schematic drawing of counterflow burner. 

Water

Water

Reactants 

N N

Screens

N2 Curtain 

Baffle

L

Flame

108
0

2

4

6

8

10

12

14

16

18

20

10 20 30 40 50 60

Global Strain Rate(1/s)

Distance From Top Duct (mm)

0

0.2

0.4

0.6

0.8

1

1.2

1.4

Curvature(1/cm)

Va/Vf=1/2 Position

Va/Vf=1 Position

Va/Vf=2 Position

Va/Vf=3 Position

Va/Vf=1/2 Curvature

Va/Vf=1 Curvature

Va/Vf=2 Curvature

Va/Vf=3 Curvature

Figure 2 The position of the luminous flame zone and the flame curvature as a function 

of the global strain rate for various ratios of the velocity of the air stream to the velocity

of the fuel stream (Vr). 

0

20

40

60

80

100

120

140

23.8 24 24.2 24.4 24.6 24.8 25 25.2

Distance from Bottom Air Duct (mm)

Teemperature (c)

ag= 13 1/s, Va/Vf=1

ag= 20 1/s, Va/Vf=1

ag= 13 1/s, Va/Vf=2

ag= 20 1/s, Va/Vf=2

Figure 3 The measured temperature as a function of location near the fuel (upper) duct 

for four near-extinction propane-air flames with N2 added to the fuel stream wit two 

different velocity ratios. 

109
0

50

100

150

200

250

300

350

400

24 24.05 24.1 24.15 24.2 24.25 24.3 24.35 24.4 24.45 24.5

Distance From Bottom Air Duct(mm)

Tem

p

erat

ure (

C)

ag=11 1/s, Va/Vf=1

ag=15 1/s, Va/Vf=1

ag=17 1/s, Va/Vf=1

ag=25 1/s, Va/Vf=1

ag=11 1/s,

dT/dy(y= 25 mm) =-721.3

K/mm

ag=15 1/s,

dT/dy(y= 25 mm)

=-49.7 K/mm

ag=25 1/s,

dT/dy(y= 25 mm)

=0 K/mm

ag=17 1/s,

dT/dy(y= 25 mm)

=-21.3 K/mm

Figure 4 The measured temperature as a function of location near the fuel (upper) duct 

for four near-extinction propane air flames with N2 added to the fuel stream under 

conditions of a constant velocity ratio for various values of the global strain rate. 

110
1400

1450

1500

1550

1600

1650

1700

1750

10 15 20 25 30 35 40 45 50 55

Global Strain Rate (1/s)

Maximum Flame Temperature(K)

Vr=2, propane from the bottom

(98%)

Vr=1, propane from the bottom

(98%)

Figure 5 Maximum temperatures as a function of strain rate for different velocity ratios 

and percentages of extinction. 

1300

1350

1400

1450

1500

1550

1600

1650

1700

0.006 0.007 0.008 0.009 0.01 0.011 0.012 0.013 0.014 0.015

Distance from the Top Duct (m)

Temperature (k)

ag=20 1/s, Vr=1

ag=35 1/s, Vr=1

ag=20 1/s, Vr=2

ag=35 1/s, Vr=2

Figure 6 Temperature profiles for ag = 20 s-1 and ag = 35 s-1 for diluent concentrations of 

98% and 99.5% of the volumetric extinction concentration and Vr/= 1 and 2. 

111
0.895

0.9

0.905

0.91

0.915

0.92

0.925

0.93

0 20 40 60 80

Global Strain Rate(1/s)

Critical N2 Volume Fraction

Vr=3, Propane from the bottom

Vr=2, Propane from the bottom

Vr=1, Propane from the top

Vr=1/2, Propane from the top

Vr=1/3, Propane from the top

Figure 7 The measured N2 volume fraction in the fuel stream required for suppression 

in flames free of conductive losses as a function of the global strain rate (ag) for a number 

of air:fuel velocity ratios (Vr) and two different rig orientations, with the fuel injected

from the bottom and the top jets, in the presence of an N2 curtain. 

112
0.78

0.8

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0 10 20 30 40 50 60 70 80 90

Global Strain Rate (1/s)

Critical N2 Volume Fraction

Propane extinguished with N2

Methane extinguished with N2

Figure 8 Comparison of extinction curves for propane and methane air diffusion flames

diluted with N2 from the fuel side. 

0.8

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0 20 40 60 80 100

Global Strain Rate(1/s)

Critical N2 and CO2 Volume 

Fractions

0.4

0.45

0.5

0.55

0.6

0.65

0.7

Critical CF3Br Volume Fraction

N2

CO2

Halon

Figure 9 The critical agent concentration in the fuel stream required for suppression of 

propane-air diffusion flames as a function of global strain rate.

113
0.32

0.325

0.33

0.335

0.34

0.345

0.35

0.355

0.36

15 20 25 30 35 40 45 50 55

Global Strain Rate (1/s)

Critical N2 Volume Fraction

Test one 

Test two

Test three

Average

Figure 10 The critical concentration of N2 in the oxidizer stream required for 

suppression of propane-air diffusion flames as a function of global strain rate. 

114
0.1

0.15

0.2

0.25

0.3

0.35

0.4

0 10 20 30 40 50 60 70 80 90

Global Strain Rate(1/s)

Effective Mass Fraction, Yeff,s

fuel + N2

air + N2

fuel + CO2

fuel + CF3Br

Figure 11 The critical effective mass fraction required for extinction of propane-air 

diffusion flames as a function of global strain rate. 

115
Appendix 5

A Technique for Extrapolating Absorption Coefficient Measurements to High 

Temperatures 

Abstract

An extrapolation technique that provides semi-quantitative estimates for the infrared 

absorption coefficients of gaseous fuels at temperatures beyond those for which 

measurements are generally practical (> 700 K) is presented. The new method is based on 

a simplified expression for molecular line intensities consisting of three fitting parameters 

and two variables (temperature and frequency). The accuracy of the extrapolations was

tested first by comparing predictions of absorption coefficients for CO, CO2

, and H2O 

vapor to the corresponding values obtained directly from the HITEMP molecular 

database. Finally, in order to establish the practical utility of the method, the spectrum of 

propane at 1000 K obtained from the extrapolation technique was compared to actual 

experimental measurements. 

116
Introduction 

The transfer of thermal energy between the flames and fuel in a combusting system is 

mediated by the absorption and emission of infrared radiation (IR). Thus, to accurately 

model fires or incinerators, the spectroscopic properties of the gas phase combustion 

intermediates must be known. This information can be obtained from non-intrusive 

measurements of the attenuation in the intensity of IR radiation from an external source, 

such as from a Fourier Transform Infrared (FTIR) spectrometer. Whether the objective is 

combustion modeling or diagnostics, calibration spectra corresponding to the gases of 

interest at elevated temperatures must be available. This poses a problem because it is 

difficult to measure calibration quality spectra at temperatures above 700 K. Furthermore, 

since the IR spectra of gas phase molecules undergo dramatic changes with increasing 

temperature, the common strategy of using spectra measured at room temperature to 

approximate high temperature absorption coefficients [1] yields unacceptable errors. 

Much attention has been focused on radiation from combustion products (H2O, CO2, CO, 

and soot) in flames. However, relatively little is known about absorption properties of 

large hydrocarbon molecules near the fuel source. DeRis [2] performed experiments 

measured radiative feedback to the fuel surface in PMMA pool fires. His work has 

motivated several recent studies on fuel absorption at elevated temperatures [3-8]. The 

focus of this investigation was to develop a semi-quantitative method to extrapolate 

absorption coefficients, obtained for temperatures up to 700 K, to the higher temperatures 

(1000 K) that are required to describe the transfer of radiation between a flame and its 

fuel-rich core. 

Extrapolation Technique 

According to the Beer Lambert law, the absorption coefficient, κν, at any wavenumber, ν, 

can be expressed in terms in terms of measurable quantities as

where τν is the spectral transmittance, p is the partial pressure of the absorbing specie (Pa),

and L is the measurement path length (m). From a more fundamental perspective [9], the 

absorption coefficient is the product of line intensity (S), line shape (g(ν-ν0

)) and the

number of absorbing molecules per unit volume and pressure: 

where νo is the center of an absorbing line in wavenumbers and NL is Loschmidts’ number 

(NL = 2.447 ·1017 molecules/cm3

/kPa at 296 K). The line intensity (S) corresponds to the 

product of the quantum mechanical probability for the transition and the population 

difference between the initial (absorption) and final (emission) states [11]: 

117
Here, h and k are the Planck and Boltzmann constants, respectively, c is the speed of light, 

E′ is the energy of the lower state, r is the transition dipole moment, and hcν is the 

difference in energy between the upper and lower states. The second bracketed term is 

the quantum mechanical transition probability, where gl is the degeneracy of the lower 

state. Finally, Ia is the isotope fraction, gi is the nuclear spin degeneracy, and Q is the 

rotational partition sum. Substituting S from eqn. (3) and the (pressure broadened) 

Lorentzian line shape for g(ν-ν0

) [9] in eqn. (2) gives: 

where Pt is the total pressure and g is a broadening parameter (cm-1/kPa). 

The approach taken in this study was to simplify this expression to a form suitable for 

fitting the absorption coefficients measured over a range of temperatures. This was 

achieved by reducing eqn. (4) to an expression containing three fit parameters (S0

, E′, and 

p) and two variables (ν and Τ). The values of the absorption coefficients at higher 

temperatures, which are difficult to measure, can then be estimated by extrapolation. 

The terms with negligible temperature or frequency dependence were consolidated into 

S0 as indicated in eqn. (5). This includes everything in the numerator with the exception 

of the wavenumber, ν, the exponentials, and explicit temperatures. In the denominator, it 

includes all but Tn and the temperature dependence of the partition sum Q. 

The rotational partition function is presumed to be proportional to T (m = 2 rotational 

degrees of freedom) in a linear molecule and T

3/2 

(m = 3) in a nonlinear molecule. With 

this in mind, all powers of T that were not contained in the exponential terms were 

included in the temperature exponent, p. As a first approximation, we also ignored the 

118
explicit contributions from overlapping rotational lines (i.e., we set ν = ν0

) so that the 

effects of the Lorentzian line shape functions are effectively subsumed into the S0 and T

p 

terms. Since the temperature dependence of each rotational transition is unique, this 

approximation is likely to result in larger errors when the resolution of the measurements 

is not sufficient to differentiate between the individual lines. After substituting E′ = hcνr = 

1.439 νr for the energy of the lower state, eqn. (4) becomes: 

In this expression temperature, T (K), and wavenumber, ν (cm-1), are independent 

variables and the parameters S0

, νr, and p were fit over a range of temperatures at each

wavenumber. 

Extrapolation data processing 

1. HITEMP Database (Carbon monoxide, Carbon Dioxide, and Water Vapor) 

Absorption data for CO, CO2 and H2O were obtained from the HITEMP database [10]. 

HITEMP is the high temperature analog of the HITRAN database, which is an extensive 

compilation of molecular line absorption data for 36 species [11]. The accuracy of the 

spectra derived from the HITEMP database for applications to combusting systems at 

temperatures exceeding 1000 K has already been examined by others [12]. Parameters 

(So

, νr, p) were fit to eqn. (6) for absorption coefficients obtained from the database at 

300, 400, 450, 500 and 600 K. Using these parameters, an extrapolation was performed to 

1000 K and comparisons were made to absorption coefficients calculated from spectra 

generated directly from the HITEMP database at 1000 K. 

2. Experiment 

A Mattson Galaxy 7020 FTIR spectrometer was modified so that the IR beam from a 

globar source passes through the interferometer but is then diverted outside the 

spectrometer, bypassing the standard sample compartment and detector. The IR beam

travels along the axis of a quartz flow cell, which is located inside a tube furnace, and

onto an external Mercury Cadmium Telluride (MCT) detector (MI0465, Graseby 

Infrared). A diagram of our experimental setup is shown in Figure 1. The entire optical 

path, including the spectrometer and external detector chamber was purged with N2 

(99.5% Super Dry) to eliminate absorption from ambient water vapor and CO2. The 

interferometer is equipped with corner-cube optics. This arrangement minimizes the 

effect of sample emission on the transmission measurements [13]. This was confirmed by 

measuring the emissions (with the source blocked) from the heated (1000 K) cell 

containing propane gas, which was found to be negligible. The furnace (Lindberg/Blue M 

119
HTF53347C with CC58434C-1 controller) has three heating zones designed to maintain a 

uniform temperature over the full length (31.75 ± 0.01 cm) of the cell (inner diameter = 

2.54 cm). The gas temperatures were measured with K-type thermocouples at the inlet 

and outlet of the cell. The temperature difference between two thermocouples was 

typically about 5K, which was considered acceptable for the purposes of this 

investigation. The ends of the quartz cell were sealed with Zinc Selenide (ZnSe) laser 

grade windows (0.3cm thickness). The temperature of the windows was kept below 700 

K to prevent oxidation of the ZnSe. This was accomplished by directing a jet of N2 on the 

outside of the windows as seen in Figure 1. The pressure inside of the cell was monitored 

and regulated to maintain 101.3 kPa during the measurements. Mixtures containing 1000 

± 20 ppm, 2500 ± 50 ppm, and 4000 ± 80 ppm of propane (CP Grade, Matheson) in N2 

(99.5% Super Dry) were made by regulating the flow of these gases into the cell using 

mass flow controllers (20 sccm for propane and 200 sccm for N2

). 

Comparisons of our extrapolation technique to experimental measurements were 

performed for propane at 800 K and 1000 K based on fits determined from measurements 

made at temperatures of 300, 400, 450, 500 and 600 K. The absorption spectra were 

measured at 1 cm-1 resolution and signal averaged over 128 scans.

 

Accuracy assessment 

In assessing the accuracy of our extrapolations, integrated absorption coefficients 

from the extrapolated spectra are compared to the corresponding values from the 

experimental (or HITEMP) spectra. In this equation, ν1 and ν2 are the frequencies 

corresponding to the band limits. In the case of propane, an average integrated absorption 

coefficient from spectra of the three (1000 ppm, 2500 ppm, and 4000 ppm) gas mixtures 

was obtained. The accuracies of the extrapolations are assessed on the basis of relative 

errors computed by taking the ratios of the differences between the integrated absorption 

coefficients obtained from the extrapolations and the “actual” values obtained either from

the HITEMP (for CO, H2

O, and CO2) or experimental spectra (for C3H8) and dividing by 

the “actual” values. 

Extrapolation Results 

As a first attempt, the accuracy of the extrapolation procedure was examined without 

introducing complications due to experimental errors. Extrapolations of the spectra for 

CO, CO2 and H2

O (based on parameters determined by fitting eqn. (6) to HITEMP 

spectra over temperatures 300, 400, 450, 500 and 600 K) were compared to the 

corresponding spectra at 1000 K generated directly from the HITEMP database. This 

exercise provides an estimate of the magnitude of the errors resulting from the 

simplifications introduced in representing eqn. (4) by eqn. (6). The practical utility of 

120
using our methodology in conjunction with experimental measurements is demonstrated 

by comparing predictions (based on fits of spectra measured at lower temperatures) for 

the absorption of IR radiation by the C-H stretching region in propane (extending from 

2700 cm-1 to 3200 cm-1) at 800 K and 1000 K. 

Carbon Monoxide 

Parameters were fit for the R branch of the 4.7 μm CO band, over the range from

2150 cm-1 to 2270 cm-1. The data were generated at a resolution of 1 cm

-1, resulting in an 

average spacing of 0.482 cm-1. This was sufficiently high that the individual lines in the 

band were resolved. Figure 2 (a) is a comparison of extrapolated absorption coefficient 

from eqn. (6) to the one from HITEMP at 300 K. The difference in the integrated 

absorption between the original and fit data, obtained from eqn. (6), is less than –0.1 %. 

The small error is expected because the reference spectrum was included in the training

set. In Figure 2 (b), however, the extrapolated absorption coefficients are compared to the 

HITEMP data at 1000 K, which is 400 K above the highest temperature spectrum used in 

the fitting procedure. The difference in the integrated absorbance at this temperature is

still only 3.6 %. While this agreement may be a little misleading because there is clearly a 

cancellation of errors, the magnitudes of the residuals (Figure 3) are on average at least 

an order of magnitude smaller than the absorbance values. 

The spectra of heavier molecules are not as well resolved at elevated temperatures as is 

the CO spectrum shown in Figure 2. Closely spaced broadened lines will, in effect, smear 

out distinguishable line structure. Since this is the case for many fuels, it is important to 

investigate the effect of reducing the resolution on the accuracy of the extrapolations. 

Figure 4 (a) and (b) compare the extrapolated absorption coefficients and HITEMP data 

at 1000K and 4 cm-1. At 4 cm-1 resolution, the CO line structure is gone and the band 

appears as a continuous absorption. The fit procedure was carried out using these data 

with similar results as the higher resolution case. The integrated absorbance at 1000 K

using the fit parameters differed from the HITEMP data by only –0.28 %. The errors in 

the temperature extrapolations for CO appear to remain tractable upon de-resolving the 

spectrum even though the contributions from overlapping rotational lines were not 

explicitly accounted for in the derivation of eqn. (6).

Water Vapor 

Parameters were fit to lines in the 2.7 μm water band, over the range between 3500 cm-1 – 

4000 cm-1. The temperatures used were the same as those in the CO case. The water 

absorption data was generated from the HITEMP database at a resolution of 4 cm-1, 

resulting in an average data spacing of 1.93 cm-1. At this resolution, most of the line 

structure was still visible in this band. Figure 5 (a) is a comparison of extrapolated 

absorption coefficient from eqn. (6) to the one from HITEMP at 300 K, using the 

parameters generated in the fit. The results of this comparison were comparable to those 

from the CO fit at 300 K. The difference in the integrated absorbance was -0.02 %. Again, 

we anticipated a small error because the parameters were fit to data at this temperature. 

However, the comparison is still favorable at 1000 K, where the difference in the 

integrated absorbance is only 3.6 %. This result is shown in Figure 5 (b). The residuals 

are shown as a function of wavenumber in Figure 6.

121
Carbon Dioxide 

The accuracy of the extrapolation method for the 4.3 μm band in CO2 over the spectral

range 2250 cm-1 to 2350 cm-1 (P branch) was also examined. The procedure was carried 

out using data at two resolutions: 0.5 cm-1 and 4 cm-1, resulting in average data spacing of 

0.241 cm-1 and 1.93 cm-1, respectively. The HITEMP and fit data at T = 300 K and 0.5 

cm-1 resolution are compared in Figure 7a. In this case the difference in the integrated 

absorption coefficient was –2.1 %, compared with –1.1 % at 4 cm-1 resolution. At 1000 K,

the difference at 0.5 cm-1 resolution was only –0.7 %. The HITEMP and fitted spectra are 

compared in shown in Figure 7 (b). While the error in the integrated absorbance is small, 

a frequency resolved comparison, as indicated by the residual spectrum in Figure 8, does 

reveal a small, but noticeable systematic error between 2270 cm-1 and 2300 cm-1. At 4 

cm-1 resolution (and 1000 K) the difference in the integrated absorption coefficient at 

1000 K is 1.3 %. The discrepancies are apparent in the comparison shown in Figure 9. A 

summary of the integrated and residual RMS errors for the extrapolations of the HITEMP 

data is presented in Table 1. Chu et al. [14] report uncertainties on the order of 2 % - 3 % 

for the measurement of absorption coefficients in the NIST Quantitative Infrared 

Database. Additionally, their study reports that line intensity variations on the order of ± 

10 % can frequently be found in comparisons of quantitative reference spectra. Although 

the cancellation of positive and negative errors makes it difficult to discern a pattern in 

the discrepancies between the integrated absorption coefficients obtained from direct 

calculation and the extrapolation, it does appear that the RMS errors grow more rapidly 

with temperature in the extrapolations from the lower resolution spectra. However, in 

either case, the errors in the extrapolations are comparable to the experimental 

uncertainties reported by Chu.

Propane (C3H8) 

Figure 10 shows how the absorption coefficients for propane vary with temperature. The 

band broadens with increasing temperature as higher energy rotational states become

populated. Furthermore, the peak absorption drops by more than an order of magnitude 

from 300 K to 1000 K. These changes in the shape and intensity of the absorption band 

are captured in the extrapolations. 

The extrapolated absorption coefficients at 800 K (a) and 1000 K (b) are compared to the 

corresponding experimental values in Figure 11. Although there is at least qualitative 

agreement, a more detailed comparison reveals that the extrapolations exhibit more 

structure and underestimate the absorbance at the wings of the bands. The fine structure 

apparent in the extrapolations is presumably an artifact resulting from the retention of the 

higher resolution inherent at lower temperatures. The discrepancy at the band wings is 

probably due to the inability of the extrapolations to capture the effects of “hot bands”, 

corresponding to vibrational transitions that are not populated at the lower temperatures 

used in determining the fitting parameters. Despite these deficiencies, the errors in the 

integrated absorption coefficients are only about 10% to 20% at 800 K and 1000 K. The 

accuracy of the extrapolations is better for frequencies in the vicinity of the band center 

and become less accurate at the wings (Table 2). The residuals are shown as a function of 

wavenumber in Figure 12. 

122
Summary and Conclusions 

The variations in the IR absorption coefficients for three combustion products (CO, CO2, 

and H2O) and one fuel (C3H8) were examined as functions of temperature from 300 K to 

1000 K. As expected, the intensities and shapes of the absorption bands of these 

molecules undergo dramatic changes over this temperature range, becoming broader and 

less intense (by as much as an order of magnitude) with the increasing population of 

higher energy rotational states. On this basis, it is concluded that this temperature 

dependence must be considered to accurately model the transfer of IR radiation and to 

measure the concentrations of hot gases in fires and incinerators by transmission 

spectroscopy. With this in mind, an extrapolation technique was developed to extend the 

range of molecular absorption coefficients to temperatures beyond the range for which 

measurements are routinely available. Experimental measurements of propane indicate 

that errors in the extrapolation are about 10 - 20 % over the temperature range of interest. 

Thus, by employing this extrapolation procedure, the accuracy of absorption coefficients 

used in models of radiative transfer and in combustion diagnostics can be improved by as 

much as an order of magnitude over what would otherwise be obtained using values 

measured at lower temperatures.

Acknowledgements 

This work was partially supported by the National Research Council Postdoctoral 

Associate Program, the NASA Microgravity Program (Contract #C32066T; Dr. Sandra 

Olson, scientific officer), and the Fire Research Grants Program (Contract # 

60NANB1D0075). The authors acknowledge helpful discussions with Professors Jungho 

Kim and Greg Jackson of the Department of Mechanical Engineering of the University of 

Maryland. 

123
References 

1. Modest, M.F., Bharadwaj, S. P., JQSRT 73 (2002) 329-338. 

2. DeRis, J., Proc. Combust. Inst. 17 (1979) 1003-1015. 

3. Fuss, S.P., Ezekoye, O.A., Hall, M.J., ASME J. Heat Transfer, 118 (1996) 918-923. 

4. Fuss, S.P., Hall, M. J., Ezekoye, O. A., Applied Optics, 38 (1999) 2895-2904. 

5. Brosmer, M.A., and Tien, C. L., JQSRT, 33 (1985) 521-532. 

6. Brosmer, M.A., and Tien, C. L., ASME J. Heat Transfer, 107 (1985) 943-948. 

7. Brosmer, M.A., and Tien, C. L., Combust. Sci. Tech., 48 (1986) 163-175. 

8. Park, S., H., Stretton, A. J., and Tien, C. L., Combust. Sci. Tech., 62 (1988) 257-271. 

9. Gross, L.A., Griffiths, P.R., and Sun, J.N.-P., in: J. Wormhoudt (Ed.), Infrared 

Measurements for Gaseous Measurements, Marcel Decker: New York. 1985, p. 81-132. 

10. Rothman, L.S., Camy-Peyret, C., Flaud, J-M., Gamache, R., Goldman, A., Goorvitch, 

D., Hawkins, R., Schroeder, J., Selby, J., Wattson, R., JQSRT, being prepared. 

11. Rothman, L.S., Rinsland, C.P., Goldman, A., Massie, S.T., Edwards, D.P., Flaud, 

J.M., Perrin, A., Camy-Peyret, C., Dana, V., Mandin, J.Y., Schroeder, J., McCann, A., 

Gamache, R.R., Wattson, R.B., Yoshino, K., Chance, K.V., Jucks, K.W., Brown, L.R., 

Nemtchinov, V., and Varanasi, P., JQSRT, 60 (1998) 665-710. 

12. Fleckl, T., Jager, H., and Obernberger, I., J. Physics. D: Applied Phys., 35 (2002) 

3138-3144. 

13. Tripp, C.P., and McFarlane, R. A., Applied Spectroscopy, 48 (1994) 1138-1142. 

14. Chu, P.M., Guenther, F. M., Rhoderick, G., C., and Lafferty, W. J., J. Research 

National Inst. St. Tech., 104 (1999) 59-81. 

124
Table 1 

Summary of errors in the HITEMP extrapolations. 

Error in integrated values (%) RMS of the residual 

Temperature (K) 

Molecule 

(Res, cm-1) Integration Range (cm-1) 

300 1000 300 1000

CO (1) 2150–2270 -0.1 3.6 1.33 E-3 

1.11 E

-3

CO (4) 2150–2270 0.3 -0.3 2.04 E-4 

1.10 E

-3

H2O (4 ) 3500–4000 -0.02 3.6 2.09 E-3 

4.85 E

-3

CO2 (0.5) 

2250–2350 -2.1 -0.7 3.06 E-2 

1.45 E

-2

CO2 (4 ) 

2250–2350 -1.1 1.3 1.36 E-2 

1.18 E

-

Table 2 

Errors in the integrated absorption coefficients for propane at 1 cm-1 resolution. 

Integrated Absorption 

Coefficient (Pa

-1

m

-1

) 

Upper: Extrapolation 

Lower: Experiment 

Error in integrated values 

(%) 

Temperature (K) 

Integration 

Range 

(cm-1) 

300 800 1000 300 800 1000 

0.030 0.022 0.020 

2700 –2850 

Right Wing

0.029 0.037 0.033 

3.5 -41.1 -40.8 

1.064 0.324 0.244 

2850 – 3050

Band Center 

1.084 0.385 0.255 

-1.9 -15.9 -4.3 

0.014 0.013 0.011 

3050 – 3200

Left Wing 

0.012 0.027 0.026 

16.7 -53.7 -59.0 

125
List of Captions 

Figure 1: Diagram of experiment setup 

Figure 2: Comparison of CO spectral absorption coefficient between HITEMP and data 

calculated using eqn. (7) with fit parameters at 1cm-1 resolution: (a) 300K, (b) 1000K. 

The CO partial pressure is 0.101 kPa and the path length is 0.32 m. 

Figure 3: Residual (κFit – κHitemp

) for CO at 1000 K. This represents the difference 

between the data sets shown in Figure 2. 

Figure 4: Comparison of CO spectral absorption coefficient between HITEMP and data 

calculated using eqn. (7) with fit parameter at 4 cm-1 resolution: (a) 300K, (b) 1000K. The 

CO partial pressure is 0.101 kPa and the path length is 0.32 m. 

Figure 5: Comparison of water vapor spectral absorption coefficient between HITEMP 

and data calculated using eqn. (7) with fit parameters at 1 cm-1 resolution: (a) 300K, (b) 

1000K. The H2

O partial pressure is 1.01 kPa and the path length is 0.32 m. 

Figure 6: Residual (κFit – κHitemp

) for H2O at 1000 K. This represents the difference

between the data sets shown in Figure 5. 

Figure 7: Comparison of CO2 spectral absorption coefficient at 300 K between HITEMP 

and data calculated using eqn. (7) with fit parameters at 0.5 cm-1 resolution: (a) 300K, (b) 

1000K. The CO2 partial pressure is 0.341 kPa and the path length is 0.32 m. 

Figure 8: Residual (κFit – κHitemp) for CO2 at 1000 K. This represents the difference 

between the data sets shown in Figure 7. 

Figure 9: Comparison of CO2 spectral absorption coefficient at 1000 K between HITEMP 

and data calculated using eqn. (7) with fit parameters at 4 cm-1 resolution: (a) 300K, (b) 

1000K. The CO2 partial pressure is 0.341 kPa and the path length is 0.32 m. 

Figure 10: Comparison of C3

H8 mean spectral absorption coefficient at 1000 K between 

Experiment and data calculated using eqn. (7) with fit parameters at 1 cm-1 resolution: (a)

Extrapolation, (b) Experiment. The temperature is from 300K to 1000K, and the path 

length is 0.32 m. 

Figure 11: Comparison of C3

H8 mean spectral absorption coefficient between Experiment 

and data calculated using eqn. (7) with fit parameters at 1 cm-1 resolution: (a) 800K, 

(b) 1000K. The path length is 0.32 m. 

Figure 12: Residual (κextrapolation – κexperiment) for C3

H8 at 1000 K. This represents the 

difference between the data sets shown in Figure 11. 

126
127
128
129
130
131
132
133
134
135
136
137
138
Appendix 6 

Determination of Planck Mean Absorption Coefficients for CF3Br 

Abstract 

The Planck mean absorption coefficient, ap, has been calculated for CF3Br over a 

temperature range from 295 K – 675 K using data obtained through measurements. A 

plot of ap vs. temperature showed monotonically decreasing behavior over this 

temperature range. The magnitude of the ap values suggests that CF3Br can significantly 

impact calculations of radiative transfer in flames containing this species. Consideration 

of conditions in an inhibited heptane flame are consistent with this conclusion. 

Nomenclature 

ap Planck mean absorption coefficient (atm-cm)

-1

α band integrated intensity (atm-1

-cm

-2) 

βω extinction coefficient (atm-cm)

-1

σω scattering coefficient (atm-cm)

-1

P pressure (atm) 

L absorption path length (cm)

T temperature (K) 

Eωb blackbody emissive power (W/m

2

/μm)

kω absorption coefficient (atm-cm)

-1

ω wave number (cm-1) 

σ Stefan-Boltzmann constant (W/m

2

/K4) 

 139
Introduction

In large-scale fires, radiation is the dominant mode of heat transfer [1]. In nonsooting and moderately sooting fires, the participation of heated gases and soot, through 

absorption and emission of infrared radiation, plays a significant role in radiative 

exchange. The Planck mean absorption coefficient, ap, characterizes radiative emission

in the source term of the generalized energy equation for an absorbing/emitting (nonscattering) medium [2]. Because it is a function only of the local temperature, ap can be 

conveniently tabulated and efficiently used in complex radiative heat transfer 

calculations, avoiding consideration of wavelength dependence. The Planck mean 

absorption coefficient, ap, is defined as: 

T

E

a 4

,b

p

d

σ

ω

=

∫ κ

∆ω

ω ω

 (1) 

To calculate ap from Eq. (1), the spectral absorption coefficient, κω, for a gas or gas 

mixture, is needed as a function of temperature. In the more general case of a scattering 

medium the absorption coefficient is replaced by the extinction coefficient, βω: 

βω κω σω = + (2)

where σω is the scattering coefficient. However, for a gas, scattering is typically

neglected and the extinction coefficient is equivalent to the absorption coefficient. The 

radiative properties, including κω, of molecular combustion species such as H2O, CO2, 

and CO have been relatively well characterized. A narrow band model (RADCAL) was 

developed which evaluates Eq. (1) for relevant combustion species including CO2, H2O, 

CO, CH4 and soot [3]. 

 140
Alternatively, Abu-Romia and Tien [4] presented a method for calculating ap that is based 

on a simplified narrow band approach. This approach utilizes the integrated intensity of 

individual gas bands that can be determined independently and tabulated. According to 

this method ap can be calculated from: 

∑

α ∫ κ

σ = σ

ω

=

∞

ω ω

j 4

j b, j

4

0

,b

p T

E

T

E

a

d

. (3) 

In this expression, the summation is over any bands that make a significant contribution 

to the absorption coefficient. The approximation is made that the blackbody emissive 

power varies linearly across each absorption band so that the value evaluated at the band 

center, (Eb, j), can be used. The only required parameter is the temperature dependent 

integrated intensity, αj, which characterizes absorption for each band. This method has 

the advantage that it is easily implemented for cases in which the integrated intensity is a

known function of temperature. Tien [5] presented results based on this model for CO2, 

H2O, N2O, NH3, SO2, CH4, CO and NO. Grosshandler and Thurlow [6] extended this 

method to derive an expression for ap based on the equivalence ratio for flames burning 

H2, CO, CH3OH and CH4 in air, and lean regions of other generic flames. 

Although the production of CF3Br (Halon 1301) has been halted due to its adverse 

effects on stratospheric ozone, it is still widely used as a fire suppressant. This agent 

inhibits flames primarily through chemical effects. Heat loss to the surroundings by 

radiative emission, however, can also influence extinction. Radiation can be particularly 

significant in weakly strained flames [7]. Therefore, it is important to characterize the 

 141
radiative properties of HF and HBr so that flame emission can be modeled more 

accurately. 

Radiative transfer involving species such as CF3Br has not been studied 

extensively, although results for HBr and HF are available [8]. HBr and HF are prevalent 

in flames and fires inhibited by CF3Br (and related halogenated compounds). They may 

also be present in incinerators. These are difficult species to study experimentally due to 

their corrosive nature. Halogenated compounds such as CF3Br are widely used as fire 

suppression agents. Measurements have shown that the local CF3Br volume fraction in 

near-extinction CF3Br-inhibited heptane/air counterflow flames can be significant, even

in the high temperature reaction zone [9]. Seshadri, for example, measured volume

fractions larger than 0.02 in such flames. 

The Planck mean absorption coefficient can be used to characterize radiative 

emission from gases of any optical thickness [2] and is thus a useful parameter, 

particularly since it can be determined a priori. The aim of this study was to calculate ap

for CF3Br for use in simplified flame models [e.g., 10] using spectral data determined at 

NIST. 

Experiment 

A FTIR spectrometer was modified so that the IR beam from a globar source passes 

through the interferometer but is then diverted outside the spectrometer, bypassing the 

standard sample compartment and detector. The IR beam travels along the axis of a 

quartz flow cell, which is located inside a tube furnace, and onto an external Mercury

Cadmium Telluride (MCT) detector. A diagram of the experimental setup is shown in 

 142
Figure 1. The entire optical path, including the spectrometer and external detector 

chamber was purged with N2 (99.5% Super Dry) to eliminate absorption from ambient 

water vapor and CO2

. The interferometer is equipped with corner-cube optics. This 

arrangement minimizes the effect of sample emission on the transmission measurements. 

This was confirmed by measuring negligible emission (with the source blocked) from the 

cell heated to 1000 K and filled with propane gas. The furnace has three heating zones

designed to maintain a uniform temperature over the full length (31.75 ± 0.01 cm) of the 

cell (inner diameter = 2.54 cm). The gas temperatures were measured with K-type 

thermocouples at the inlet and outlet of the cell. The temperature difference between two 

thermocouples was typically about 5 K, which was considered acceptable for the 

purposes of this investigation. The ends of the quartz cell were sealed with Zinc Selenide 

(ZnSe) laser grade windows (0.3 cm thickness). The temperature of the windows was 

kept below 700 K to prevent oxidation of the ZnSe. The pressure inside of the cell was 

monitored and regulated to maintain 101.3 kPa during the measurements. Mixtures of 

CF3Br (CP Grade, Matheson) in N2 (99.5% Super Dry) were made by regulating the flow 

of these gases into the cell using mass flow controllers (20 sccm for CF3Br and 

1000 sccm for N2

). Experimental measurements were performed at ambient temperature 

(295 K), 375 K, 475 K, 575 K and 675 K. The measurements were limited to 

temperatures below 700K because of thermal restrictions on the windows. The absorption 

spectra were measured at 4 cm-1 and 16 cm-1 resolution and signal averaged over 128 or 

256 scans. 

 143
Methodology 

Planck mean absorption coefficients were calculated according to Eqn. (3). The 

integrated intensity, α, was determined by integrating the spectral absorption coefficient 

for each band and taking value in the optically thin limit according to: 

. (4) ∫ → ∆

=

ω

ω α κ dω

pL 0

lim

This procedure was carried out for measurements at 295 K, 375 K, 475 K, 575 K, and 

675 K. 

Results 

Table 1 shows the Planck mean absorption coefficient for CF3Br calculated over a 

temperature range from 295 K – 675 K. The values of ap are significantly larger than

those of HF, HBr, and HCl over this temperature range [8], although the general trend of 

monotonic decrease with increasing temperature is similar to these species. At 295 K, the 

Planck mean absorption coefficient for CF3Br is 2.64 (atm-cm)

-1. The values for HF, 

HCl and HBr are, respectively, 0.46, 0.05, and 0.02 (atm-cm)

-1. The reason for this 

difference is the relatively large absorption coefficient possessed by CF3Br as compared

to the other molecules. Figure 2 shows a plot of the spectral transmittance for CF3Br for 

the conditions T = 295 K, pL = 0.031 atm-cm. 

In CF3Br inhibited flames, Seshadri [9] and Masri [11] report that the volume 

fraction of CF3Br is comparable to that of CO in the high temperature reaction zone. For 

a flame with the CF3Br suppressant added to the air stream of a diffusion flame, the 

 144
concentrations of CO2, H2O, and CO will fall as a function of distance from the high 

temperature reaction zone, and CF3Br will be larger than any other species that

participates in radiative exchange [9]. In cool flame regions, CF3Br may absorb flame 

radiation, and block its emission to the surroundings. The inclusion of CF3Br in radiative 

emission calculations of CF3Br inhibited hydrocarbon flames may be significant and 

ought to be considered, depending on the combustion configuration. 

To estimate the impact of CF3Br in the cool flame region, a CF3Br inhibited 

heptane/air counterflow flame [9] was considered. The results showed that, for optically 

thin conditions, CF3Br was found to account for approximately 80 % of the total emission 

in the region approximately 1.5 mm from the reaction zone where the temperature is 

675 K. In Ref. [9], data are not available for CF3Br closer to the reaction zone 

(corresponding to a temperature of 1300oC). Yet, ap values for all relevant species tend 

to decrease with increasing temperature. While the data in Ref. [9] show that CF3Br has 

a relatively small concentration in the reaction zone, given the magnitude of its Planck 

coefficient relative to other flame species considered here (CH4, CO2, H2O, and CO), it is

expected to plays a significant role in the total radiative emission even at higher 

temperatures. 

 

Conclusions 

Values of the Planck mean absorption coefficient have been calculated for CF3Br 

based on FTIR data over a temperature range from ambient to 700 K. The temperature 

dependent values of ap fall off with temperature not unlike HBr and HF. The inclusion of 

CF3Br, HBr and HF, can significantly contribute to emission calculations for fire 

suppression or incinerator scenarios where halogenated compounds are present. 

 145
References

[1] DeRis, J., 1979, “Fire Radiation-A Review,” Proc. 17th Sym (Int) Combust, The 

Combustion Institute, pp. 1003-1015. 

 

[2] Siegel, R., and Howell, J.R., 1992, Thermal Radiation Heat Transfer, Hemisphere 

Publishing, Washington. 

[3] Grosshandler, W.L., 1993, “RADCAL: A Narrow-Band Model for Radiation 

Calculations in a Combustion Environment,” NIST Technical Note 1402, National 

Institute of Standards and Technology, Gaithersburg MD. 

[4] Abu-Romia, M.M., and Tien, C.L., 1967, “Appropriate Mean Absorption 

Coefficients for Infrared Radiation of Gases,” ASME Journal of Heat Transfer, 89,

pp. 321-327. 

[5] Tien, C.L., 1968, “Thermal Radiation Properties of Gases,” Advances in Heat 

Transfer, Irvine, T.F., and Hartnett, J.P., eds., Academic Press, New York, 5, pp. 

253-324. 

[6] Grosshandler, W.L., and Thurlow, E.M., 1992, “Generalized State-Property 

Relations for Nonluminous Flame Absorption Coefficients,” ASME Journal of 

Heat Transfer, 114, pp. 243-249. 

[7] Maruta, K., Masaharu, Y., Hongsheng, G., Yiguang, J., and Niioka, T., 1998, 

“Extinction of Low-Stretched Diffusion Flame in Microgravity,” Combust Flame, 

112, pp. 181-187. 

[8] Fuss, S.P., and Hamins, A., Determination of Planck Mean Absorption Coefficients 

for HBr, HCl, and HF, ASME J. Heat Transfer, 124, 26-30 (2002). 

[9] Seshadri, K. and Williams, F.A., Halogenated Fire Suppressants. ACS Symposium 

Series 16. April 23-24, 1975, San Antonio, TX, American Chemical Society,

Washington DC, (Ed.: R.G. Gann) pp. 149-182, 1975. 

[10] Choi, M.Y., Hamins, A., Rushmeier, H., and Kashiwagi, T., “Simultaneous Optical 

Measurement of Soot Volume Fraction, Temperature, and CO2 in Heptane Pool 

Fire,” Proc. 25th Sym (Int) Combust, The Combustion Institute, pp. 1471-1480, 

1994. 

[11] Masri, A.R., 1994, “Chemical Inhibition of Nonpremixed Flames of Hydrocarbon 

Fuels with CF3Br,” Combust Sc Tech, 96, pp. 189-212. 

 146
Table 1. Calculated values of Planck mean absorption coefficient for CF3Br 

T (K)

Planck mean

absorption coefficient 

(atm-cm)

-1

295 2.64

375 2.59

475 2.18

575 1.80

675 1.43

 

 147
Figure 2: Transmittance spectrum of CF3Br at T = 295 K, pL = 0.031 atm-cm. 

 148
Appendix 7

Determination of Planck Mean Absorption Coefficients for Hydrocarbon Fuels 

ABSTRACT

Infrared absorption of propane, n-heptane and propylene were measured using Fourier 

transform infrared spectroscopy (FTIR) for temperatures ≤ 1000 K to facilitate calculation of 

absorption coefficients of fuel molecules in non-premixed flames. Spectrally resolved fits of the

absorption coefficient data using a semi-empirical quantum-based expression provide a basis for 

calculating infrared spectra at any temperature. Planck mean absorption coefficients of these 

fuels and of methane calculated from the HITRAN database are compared since methane 

absorption has been used to model fuel species absorption in fires. Propane and heptane have 

similar characteristics over the entire temperature range whereas methane and propylene with

their higher proportion of absorption in low frequency bands have peak mean absorption 

coefficients at lower temperatures where blackbody radiation peaks near the spectral range of 

these bands. Propylene with low frequency absorption bands associated with the C=C bond has

the highest Planck mean absorption coefficient at temperatures < 600 K. N-heptane has the 

largest Planck mean absorption at temperatures ≥ 800 K where blackbody emissions peak near or 

above the spectral range of the C-H stretching bands. Implications of these results on fuel 

absorption of radiative heat transfer in large flames are discussed. 

149
INTRODUCTION

The energy required to volatilize solid and liquid fuels in fires is transported by conduction, 

convection and radiation. With small fires less than 0.1 m in base diameter, conduction 

dominates the heat transfer to the fuel source, while convection tends to dominate in fires with 

diameters between 0.1 and 0.3 m. Above about 0.3 m, radiation heat transfer is typically the 

largest source of heat feedback to the fuel source (Burgess and Hertzberg 1974). Because of the 

dominance of radiation for large fires, an accurate assessment of radiation absorption within the 

fuel rich core of the fire becomes imperative for assessing heat feedback to the fuel source. 

For large fires, radiation dominates external heat transfer to the surrounding as well as 

internal heat transfer to the fuel source. Hamins et al. reported that radiation heat transfer 

provides 96%, 80%, and 55% of the total heat feedback to the fuel source for 0.3 m diameter 

pool fires of toluene, n-heptane and methanol respectively (Hamins et al. 1994). Furthermore, 

because external heat transfer from large fires is typically dominated by radiation, assessment of 

fire damage depends on accurate models of radiative heat transfer emitted from fires, which in 

turn depends on the flame temperature. Flame temperature will be impacted by radiative 

absorption by both soot and the hydrocarbon-rich core between the fuel surface and the flame. 

Since the incident radiative heat flux controls fuel volatilization in fires, accurate assessment of 

the effects of gaseous fuel and fuel-derived molecules on radiative exchange in the fuel-rich core 

is needed. 

In most previous studies, combustion products such as carbon dioxide (CO2), water vapor 

(H2O), carbon monoxide (CO) and soot have been the focus of radiative transport (Buckius and

Tien 1977; Lee and Tien 1980; Ludwig et al. 1973). Studies on the radiative absorption of these

molecules have provided absorption coefficients over a broad range of temperatures, and one can 

150
obtain this data through HITRAN (Rothman et al. 2005b). To include contributions of fuels and 

other fuel decomposition products to radiative transport in fires, absorption coefficients for these 

molecules must be measured over a broad range of temperatures characteristic of the fuel rich 

core, from 300 to 1000 K (Brosmer and Tien 1987; Klassen and Gore 1994). 

Many studies incorporating radiative transport in fires or combustion utilize a gray gas

assumption to calculate radiation attenuation. The gray gas emissivity approximation is often 

based on empirical fits to external radiation intensity measurements, e.g., in fires of volatilized 

plastics (Markstein 1979). The effective emissivity and corresponding absorptivity will vary 

significantly with flame temperature particularly for sooty flames and will further depend on 

how close the peak of the broadband flame emissions falls relative to the strong absorption bands 

of the vaporized fuel molecules (De Ris 1979). To date, only a few studies have attempted to 

provide data for assessing the temperature dependent absorption coefficients of fuel molecules at 

temperatures as high as 1000 K. Notable studies in this area include the work of Tien and 

coworkers on hydrocarbon fuels and methyl-methacrylate (Brosmer and Tien 1985a; Brosmer 

and Tien 1985b; Brosmer and Tien 1986; Park et al. 1988), and Fuss et al. on paraffin 

hydrocarbons and acid gases (Fuss et al. 1996; Fuss et al. 1999; Fuss and Hamins 2002).

Difficulties in measuring spectrally resolved infrared absorption coefficients at high temperature

arise because of the temperature limitations of most infrared window materials and issues 

associated with high temperature vacuum seals. 

This study has been undertaken to build a database for absorption coefficients for several 

hydrocarbon molecules for temperatures up to 1000 K. Infrared absorption coefficients of the 

hydrocarbons propane (C3H8), n-heptane (C7H16), and propylene (C3H6) were measured by 

Fourier transform infrared spectrometer (FTIR) in a unique experimental facility for 

151
measurements to 1000 K. These measurements are analyzed and compared to existing data for 

methane (CH4) obtained from the HITRAN database (Rothman et al. 2005a). 

The intensity of thermal radiation from flames is diminished by absorption and augmented by 

emission as it travels through the atmosphere of a fire. The emissivity of the participating 

medium is given in terms of the Planck mean absorption coefficient κP as a function of 

temperature T. κP(T) is expressed as an integral of the product of the spectral absorption 

coefficient κν(T) and the blackbody Planck function Ebν(T) over all wavenumbers ν (Siegel and 

Howell 1992): 

( ) ∫

∞

⎟ −

⎠

⎞ ⎜

⎝

⎛ =

0 2

3

1

4

exp 1

1

ν

ν

ν

κ

σ

κ ν d

T

C

C

T P T (1) 

where C1, C2 and σ are the first and second radiation constants and the Stefan-Boltzmann 

constant, respectively. Since κP is a function only of T, it can be easily presented in tabular form

(Zhang and Modest 2002). Approximate values are obtained by numerical integration and the

accuracy increases with the resolution of the spectral measurements. Accurate Planck mean 

absorption coefficients are important in combustion and fire modeling for solution of radiative 

transport equation without employing computationally-prohibitive, spectrally-resolved 

calculations. However, use of the Planck mean absorption coefficient approach is limited by the 

scarcity of temperature dependent data. In this paper, κP for n-heptane, propane, propylene, and 

methane are calculated for a range of conditions and the impact of hydrocarbon radiative 

absorption in large-scale fires is discussed. 

EXPERIMENTAL APPARATUS

152
A Mattson Galaxy 7020*

 FTIR spectrometer was modified such that the IR beam from a 

globar source passed through the interferometer and was diverted to bypass the standard sample

compartment and detector and to travel along the axis of a quartz flow cell. The flow cell was 

located inside a three-zone tube furnace, and the IR beam passed through the cell into an external 

Mercury Cadmium Telluride (MCT) detector (MI0465, Graseby Infrared). Absorption spectra

were measured at 1 cm

-1 resolution with the signal averaged over 128 scans. A diagram of the 

experimental setup is shown in Figure 1a. 

The entire optical path, including the spectrometer and external detector chamber was purged 

with N2 (99.995 % min. purity) to eliminate absorption from ambient water vapor and CO2. The 

interferometer was equipped with corner-cube optics which minimized the effect of sample

emission on the transmission measurements (Tripp and McFarlane 1994). This was confirmed 

by measuring the emission with the source blocked from the heated (1000 K) cell containing 

propane gas, which was found to be negligible. The furnace (Lindberg/Blue M HTF53347C)

had three heating zones designed to maintain a uniform temperature over the full length (31.75 ±

0.01 cm) of the cell (inner diameter = 2.54 cm). Gas temperatures were measured with K-type 

thermocouples at the inlet and outlet of the cell. The temperature difference between two 

thermocouples was typically about 5 K, which was considered acceptable for the purposes of this

investigation. The ends of the quartz cell were sealed with zinc selenide (ZnSe) laser grade

windows (0.3 cm thickness). To prevent oxidation of the ZnSe, the temperature of the windows

was kept below 700 K by directing a jet of N2 on the outside of the windows as indicated in 

Figure 1a. This cooling permitted high-resolution infrared measurement at high temperature

 

*

 Certain commercial equipment, materials, or software are identified in this manuscript to specify adequately the 

nature of the research. Such identification does not imply recommendation or endorsement by the National Institute 

of Standards and Technology, nor does it imply that the materials or equipment are necessarily the best available for 

the purpose. 

153
with many scans. The pressure inside of the cell was monitored and regulated to maintain 101 

kPa ± 1 kPa during the measurements. Mixtures containing volume fraction of 1000, 2500, and 

4000 ppmv of propane (99% min. purity) in N2 (99.995% min. purity), and 5000, 10000, and 

15000 ppmv of propylene (99% min. purity) in N2 (99.995% min. purity) were passed through 

the cell using mass flow controllers (MKS 1479A), which had a ± 0.5% uncertainty for the range 

of flows in the current study. 

For n-heptane (99.0% min. purity), a mass flow controller (F.S. 20 cm3

/min N2) was used to 

bubble N2 through three liquid heptane baths, assuring a continuous supply of saturated heptane 

vapor (Figure 1b) at room temperature. The n-heptane concentration was calculated based on the 

liquid temperature, which was measured by a K-type thermocouple within the last heptane bath. 

The n-heptane vapor pressure of the saturated stream was obtained using the Antoine equation 

with parameters obtained from a previous study (Carruth and Kobayashi 1973). Since 

uncertainty of the thermocouple is ± 0.5 K from our calibration, the uncertainty of the fuel 

concentration is ± 2.7%. A second mass flow controller (F.S. 1000 cm3

/min N2) was used to 

dilute the heptane-saturated flow with pure N2 before introduction into the gas cell.

Measurements were conducted for three heptane concentrations, which varied as shown in Table 

1 due to small variations in the fuel bubbler temperature over time. 

METHODOLOGY 

The simplified extrapolation and interpolation technique of absorption coefficient of 

Wakatsuki et al. (Wakatsuki et al. 2005) was used to determine κν as a function of T. This 

approach involves taking the full quantum expression for κν and simplifying it to the following 

expression 

154
n

r

T

T T

S ⎟

⎠

⎞ ⎜

⎝

⎛ ⋅ − ⎥

⎦

⎤ ⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛ ⋅ − −

=

ν ν

ν

κ ν

1.439

exp

1.439 1 exp 0

 (2) 

Equation 2 ignores some of the minor temperature dependencies of some terms in the full 

quantum expression by lumping them into a constant S0, but provides a reasonable 

approximation to the full equation with respect to T. The three parameters S0, νr, and n are 

optimized using a least squares approach to fit measured spectral absorption data such that κν can

be extrapolated and interpolated for any range of T. The fits do very well at capturing the falloff

with T of the various absorption band peaks, but have limited accuracy for the wings of strong 

bands where slight increases in κν with T can arise due to line broadening. However, the wings 

typically account for very small amounts of the total integrated κP and thus the fitted equations 

provide sufficiently accurate values for κν, and for finding κP as a function of T.

In the current study, κP was determined at 50 K temperature increments using the fits for S0, 

νr, and n at 1.0 cm-1 intervals to calculate κν for integration in equation 1. Infrared spectra were 

calculated for propane, n-heptane and propylene from 300 to 1400 K whereas the fits for S0, νr, 

and n were derived from absorption measurements at 300, 400, 450, 500, and 600 K. Earlier 

studies verified that the error in the integrated absorption coefficient using the extrapolation of 

the fits for propane to 800 and 1000 K for propane were < 20% in comparison to measured 

values (Wakatsuki et al. 2005). For lower T, within the range of measurements, interpolated κP

were < 10% of the measured values. The errors in the extrapolations are comparable to the

experimental uncertainties reported elsewhere (Chu et al. 1999). κP for methane was derived

using the HITRAN database to obtain κν for the same range of T. The same temperatures for the 

experimental fuels were used for methane as well. 

155
RESULTS

Figure 2 shows the absorption coefficient of propane at 296 K, which illustrates typical 

absorption spectra for a paraffin hydrocarbon. The temperature dependent normalized blackbody 

spectral emissive power is also plotted as a function of wavenumber to show how characteristic 

bands of such fuels impact absorption of blackbody radiation at varying temperatures. Propane 

contains a CH3- stretching peak centered at 2960 cm-1 and bending peaks at 1460 cm-1 and 1380 

cm

-1 as well as -CH2- stretching peaks at 2930 cm-1 and 2850 cm-1, and a bending peak at 1470 

cm

-1. Figures 3a) and 3b) show the spectral absorption coefficient of the propane stretching and 

bending band regions as a function of T. The plots show the characteristic decline with 

increasing T in absorption coefficients near the band centers and the broadening causing a rise 

with increasing T in spectral absorption near the edges of the band. 

The location of the peak blackbody emission shifts to larger wavenumber (shorter 

wavelength) as T increases as indicated in Figure 3. The low wavenumber absorption bending 

bands thus have a larger impact on emission and absorption of blackbody radiation at T < 800 K. 

On the other hand for T ≥ 800 K, more characteristic of fire emissions, the higher frequency C-H 

stretching peaks fall nearer to the peaks of blackbody emission and thus play the dominant role 

in fuel-based radiative absorption. 

Figure 4 shows the absorption coefficient of n-heptane at 296 K and the temperaturedependent normalized blackbody spectral emissive power as a function of wavenumber. As 

normal alkanes, propane and heptane have similar characteristic band groups (CH3- and -CH2-). 

Figure 5a) and 5b) show the spectral absorption coefficient of the C-H bending and stretching 

peaks for heptane as a function of T. As expected, comparison of Figures 3 and 5 shows that 

156
heptane absorbs more radiant energy per unit mole than propane, due to the increased number of 

C-H bonds. Furthermore, heptane, with its increased chain length, has a reduced sharpness in 

the peaks compared to propane. 

Figure 6 shows the absorption coefficient of propylene at 296 K and the temperaturedependent normalized blackbody spectral emissive power as a function of wavenumber. 

Propylene has several characteristic band groups such as a bending group (1420 cm-1), stretching

of =CH2 (3080 cm

-1), C-H out of plane bending (1860 cm-1), C=C stretching (1645 cm-1), and 

=CH2 out of plane bending (990 cm-1 and 910 cm-1), in addition to a C-H stretching group (2960 

cm

-1 for CH3-). Propylene’s =CH2 out of plane bending enhances radiation absorption and 

emission at lower T. Figures 7a), 7b), and 7c) show the spectral absorption coefficient of 

propylene =CH2 and C-H out of plane bending peaks, and C=C stretching, and C-H in and out of

plane bending and =CH2 stretching peaks as a function of T. 

Figure 8 shows the Planck mean absorption coefficient for methane (obtained from 

HITRAN), and for propane, heptane and propylene (interpolated and extrapolated from the 

experimental data), as a function of temperature up to 1400 K in 50 K increments. Table 2 

summarizes the fourth order polynomial fits given by Eq. 3 to the processed data. 

4

4

3

3

2 κ P = a0 + a1

T + a2T + a T + a T (3) 

Figure 8 shows that propane and heptane have similar characteristics. The number of -CH2- 

groups impacts the magnitude of the absorption coefficient, but it does not influence the general 

shape of κP as a function of T. The peak of κP for methane occurs at much lower T than for the

two larger alkanes due to an increased proportion of its absorption coming in the low-frequency 

bands. Figure 9 shows the absorption coefficient of C-H bending peaks for methane, propane, 

and heptane at 296 K. The bending motion for a methane molecule produces sharp well-defined 

157
peaks at 1300 cm-1, and κP for methane is larger than propane up to about 400 K and similar to 

heptane at 300 K in spite of the fewer number of bonds. 

Propylene has the largest κP for T ≤ 500 K due to the additional contributions of =CH2

bending, but the effect of the bending band on κP is diminished for T > 800 K where blackbody 

emissions peak well above the frequency of those bands. For T > 800 K, propane with the two 

extra C-H bonds, has a slightly higher κP than propylene since t he extra low frequency bands for 

propylene contribute very little to the blackbody radiation at these higher T. The results for 

propylene from the experiments in this paper compare very favorably with previous studies 

(Brosmer and Tien 1986). Since transmittance of the ZnSe window used in the gas cell 

experiments in this study rapidly falls off below 700 cm-1 (above 14.3 μm) as T increases, the 

absorption spectrum of the low frequency 578 cm-1 band was not measurable in the current

study. Brosmer and Tien calculated κP for propylene with and without the 578 cm-1 band by 

approximation from the result of Lord (Lord and Venkateswarlu 1953) and Silvia (Silvia et al. 

1973). They estimated the inclusion of the band gave 15, 10, and 5% of the total absorption for 

blackbody emissions at 300, 400, and 550 K respectively and rapidly dropped below 1% for 

higher T. κP for propylene was corrected to include the 578 cm

-1 band and results showed that 

the current Planck means absorption coefficient was typically 15% higher than those of the

previous study for the temperatures measured. 

Because the current study explored very high temperatures both experimentally and with fit 

extrapolations, it is important to consider how fuel pyrolysis influenced the measurements and 

subsequent calculations for the Planck mean absorption coefficient. Pyrolysis of the fuels of 

interest were calculated using the Aurora program in CHEMKIN (Kee et al. 2003) for 

temperatures up to 1800 K. The detailed chemical kinetics mechanisms used for these 

158
calculations were taken from work by Petrova and Williams for propane and propylene (Petrova 

and Williams 2006) and by Seiser et al. for heptane (Seiser et al. 2000). The experimental cell 

conditions with the calculated residence times (ranging from 15 s at 300 K down to

approximately 3 s at 1000 K) were input into Aurora and pyrolysis of each species within the cell 

was calculated. Figure 10 shows results for each fuel of the volume (mole) fraction passing 

through the cell unreacted. Heptane is predicted to begin to have significant pyrolysis at about

900 K and have about 77 % fuel conversion at 1000 K. Figure 11 compares the heptane infrared 

spectrum obtained by experiment at 1000 K and by extrapolation of fits from lower temperature 

data to 1000 K. The extrapolated data was obtained by using the simplified extrapolation 

technique (Wakatsuki et al. 2005) under the assumption that no fuel pyrolysis occurred at any 

temperature. Figure clearly shows that significant heptane pyrolysis has occurred at 1000 K.

For the other species, pyrolysis of methane, propane, and propylene starts at 1200, 900 and 950 

K respectively. An alternative Planck mean absorption coefficient incorporating the effects of 

fuel pyrolysis temperature might be considered for T > 800 K. These results suggest that the 

significance of various fuel species absorption in a fire will depend on the temperature of the 

radiative flame as well as on the gases themselves in the fuel rich core. Figure 8 indicates that 

absorption by the larger alkanes increases by 4 to 5 times as T rises from 300 K to 800 K. 

Another important question regarding radiation feedback involves the impact of fuel

pyrolysis on overall absorption. In this regard, propylene and methane are interesting molecules 

as both are prominent pyrolysis products during the breakdown of alkanes such as n-heptane. In 

fact two C3H6 molecules plus one CH4 molecule has the same number of C’s and H’s as C7H16, 

and for T > 800 K), the sum of the Planck mean absorption coefficients for the smaller molecules 

is only slightly higher than that of the larger heptane molecule. At lower temperature (T < 600 

159
K), the smaller molecules have much higher Planck mean absorption coefficients whereas the

larger n-heptane does not. This suggests that if pyrolysis products find their way either through 

diffusion or convection into cool regions of the flame core, the impact of pyrolysis on 

hydrocarbon radiation absorption may be very significant. This is an area for further study, both 

with additional data collection and with detailed modeling of large-scale pool fires in which 

chemically resolved fuel rich cores and detailed radiation transport calculations are developed. 

CONCLUSIONS 

Absorption coefficients of hydrocarbon fuels (propane, heptane and propylene) were 

obtained experimentally and the Planck mean absorption coefficients were calculated using a 

simplified fitting and extrapolation equation up to 1400 K and the HITRAN database (methane).

Fourth order polynomial fitting equations for each fuel are provided as a function of temperature.

Large differences in the absolute values of the coefficients were obtained for the different 

hydrocarbon molecules. Based on this observation, it is clear that an accurate description of 

radiative transfer in fires will require temperature dependent absorption coefficients for all fuels 

and decomposition products present in significant concentrations. As the carbon number 

increases, similar trends in the Planck mean absorption coefficients as a function of frequency 

were observed. However, at lower temperatures, a representative alkene, i.e., propylene, shows a 

large contribution for absorption and emission arising from the bending motion of the C=C bond 

and the associated =C-H bond. This suggests that other fuels such as toluene, which contain 

characteristic band groups that absorb at low wavenumber, will significantly impact radiation 

transfer in fuel rich cores at low temperature. The importance of fuel pyrolysis and conversion 

of large alkanes to smaller alkenes and methane may or may not significantly impact radiative

160
absorption as it depends on how much of the high temperature pyrolysis products migrate to 

cooler zones where their mean absorption coefficients increase due to their low wavenumber 

absorption bands. These results suggest the importance of further study on the effects of fuel

absorption in large-scale fires. 

ACKNOWLEDGEMENT 

This work was supported by the Fire Research Grants Program (Contract # 

60NANB1D0075) of the Building and Fire Research Laboratory at the National Institute of 

Standards and Technology (NIST), and by the NASA Microgravity Program (Contract 

#C32066T; Dr. Sandra Olson, scientific officer). 

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165
TABLES 

Table 1 - Volume fraction of heptane in the heptane/nitrogen mixture used in the FTIR 

measurements. 

Temperature (K)

293 510 1060 1671

400 492 1023 1614

450 516 1056 1684

490 510 1037 1680

593 516 1000 1593

794 497 1009 1475

1000 590 1045 1509

Volume Fraction (x 10-6)

Table 2: Values for 4th order polynomial fits with equation 4 to Planck mean absorption 

coefficient data. 

Methane Propane Heptane Propene

a0 4.00E+00 -1.53E+01 -3.59E+01 3.88E+01

a1 6.57E-03 8.47E-02 1.88E-01 -6.46E-02

a2 -1.37E-05 -9.75E-05 -2.08E-04 4.67E-05

a3 7.34E-09 4.29E-08 8.87E-08 -1.63E-08

a4 -1.27E-12 -6.58E-12 -1.33E-11 2.20E-12

R2 9.92E-01 9.96E-01 9.94E-01 9.99E-01

300 K - 1400 K

166
FIGURES 

N2 flow in

TC (fuel) for 

Pvap calcs.

TC (water)

3-stage fuel bubbler

dilution flow

Water Bath

MFC

MFC

liquid vapor 

sample

Sample

in

Sample

Out

N2 Cooling Quartz Cell N2 Cooling

ZnSe

windows

Sample Stream

Three Zone Furnace

IR

S

o

urc

e

Detector

Thermocouples

31.8 cm

(a)

ZnSe

windows

(b)

N2 flow in

TC (fuel) for 

Pvap calcs.

TC (water)

3-stage fuel bubbler

dilution flow

Water Bath

MFC

MFC

liquid vapor 

sample

N2 flow in

TC (fuel) for 

Pvap calcs.

TC (water)

3-stage fuel bubbler

dilution flow

Water Bath

MFC

MFC

liquid vapor 

sample

Sample

in

Sample

Out

N2 Cooling Quartz Cell N2 Cooling

ZnSe

windows

Sample Stream

Three Zone Furnace

IR

S

o

urc

e

Detector

Thermocouples

31.8 cm

(a)

ZnSe

windows

Sample

in

Sample

Out

N2 Cooling Quartz Cell N2 Cooling

ZnSe

windows

Sample Stream

Three Zone Furnace

IR

S

o

urc

e

Detector

Thermocouples

31.8 cm

(a)

ZnSe

windows

(b)

Figure 1 – Diagram a) of high temperature FTIR test facility with N2 cooling of ZnSe windows 

and b) continuous liquid fuel vapor supply for heptane vapor controlled by saturation and 

downstream dilution.

167
4000 3500 3000 2500 2000 1500 1000

0

400

800

1200

1600

2000

2400

0.0000

0.0002

0.0004

0.0006

0.0008

Absorption Coefficient (atm-1m-1

)

Wavenumber (cm

-1)

1200 K

800 K

Normalized Blackbody 

S

pectral Emissiv

e P

ower (1/cm-1)

400 K

Figure 2 – The measured spectral absorption coefficient of propane (C3H8) at 296 K and the 

temperature dependent normalized blackbody spectral emissive power as a function of 

wavenumber. 

168
Wavenumber (cm-1)

1600 1500 1400 1300

Absorption Coefficient 

κ

ν(atm-1m-1)

0

50

100

150

200

250

295 K

396 K

578 K

790 K

1009 K

295 K

396 K

578 K

790 K

1009 K

(a) 

Wavenumber (cm-1)

3200 3100 3000 2900 2800 2700

Absorption Coefficient 

κ

ν(atm-1m-1)

0

500

1000

1500

2000

2500

295 K

396 K

578 K

790 K

1009 K

(b) 

Figure 3 – Measured temperature-dependent spectral absorption coefficient of C3H8; (a) C-H

bending and (b) C-H stretching region. 

169
4000 3500 3000 2500 2000 1500 1000

0

1000

2000

3000

4000

0.0000

0.0002

0.0004

0.0006

0.0008

Absorption Coefficient (at

m

-1

m

-1

)

Wavenumber (cm-1)

1200 K

800 K

Normalized Blackbody

Spectral E

missive Power (1/cm-1)

400 K

Figure 4 – Measured spectral absorption coefficient of n-C7H16 at 296 K and the temperature 

dependent normalized blackbody spectral emissive power as a function of wavenumber.

170
Wavenumber (cm-1)

1500 1400 1300

Absorption Coefficient 

κν(atm-1

m-1)

0

100

200

300

400 293 K

393 K 

594 K

801 K

1002 K 293 K

393 K

594 K

801 K

1002 K

Wavenumber (cm-1)

3200 3100 3000 2900 2800 2700

Absorption Coefficient 

κ

ν(atm-1m-1)

0

1000

2000

3000

4000

293 K

393 K

594 K

801 K 

1002 K

(a) 

(b) 

Figure 5 – The measured temperature dependent spectral absorption coefficient of n-heptane; (a)

C-H bending and (b) C-H stretching. 

171
4000 3500 3000 2500 2000 1500 1000

0

200

400

600

800

0.0000

0.0002

0.0004

0.0006

0.0008

Absorption Coefficient (atm-1

m

-1)

Wavenumber (cm-1)

1200 K

800 K

Normalized Blackbody 

Spectral Emissive Power (1/cm-1)

400 K

Figure 6 – The measured spectral absorption coefficient of C3H6 at 296 K and the temperaturedependent normalized blackbody spectral emissive power as a function of wavenumber.

172
Wavenumber (cm-1)

1100 1000 900 800 A

bsorption Coefficient 

κ

ν(at

m-1

m-1)

0

200

400

600

800

1000

296 K 

390 K 

594 K 

793 K 

1003 K

296 K

594 K

793 K

390 K

1003 K

Wavenumber (cm-1)

1900 1800 1700 1600 1500 1400 1300

Absorption Coefficient 

κ

ν(atm-1

m-1)

0

50

100

150

200

296 K

390 K

594 K

793 K

1003 K

(a) 

(b) 

Wavenumber (cm-1)

3200 3100 3000 2900 2800 2700 A

bsorption Coefficient κ

ν(atm-1

m-1)

0

100

200

300

400

500

296 K 

390 K 

594 K 

793 K 

1003 K

(c) 

Figure 7 – The measured temperature dependent spectral absorption coefficient of Propene 

(C3H6); (a) =CH2 out of plane bending (b) C-H in plane and out of plane bending, and C=C 

stretching and (c) CH3- and =CH2 stretching. 

173
300 600 900 1200 1500

0

5

10

15

20

25

Planck mean absorption

coefficient (atm-1m-1)

Temperature (K)

 Methane

 Propane

 Heptane

 Propene

Figure 8 – The Planck mean absorption coefficient of CH4 (from HITRAN), C3H8, n-C7H16, and 

C3H6 (from fitting and extrapolation of measurements). 

174
Wavenumber (cm-1)

1600 1500 1400 1300 1200

Absorption coefficient (atm-1

m-1)

0

200

400

600

800

1000

1200

Methane

Propane 

Heptane 

Figure 9 – Spectral absorption coefficients for C-H bending peaks for methane (HITRAN), and 

propane and heptane (experimental) at 296 K.

175
Temperature (K)

400 800 1200 1600 2000

Nor

m

aliz

e

d Volum

e Fracti

on 0.0

0.2

0.4

0.6

0.8

1.0

1.2

Methane

Propane

Heptane

Propylene

Figure 10 – Calculated normalized fuel volume fraction of hydrocarbon fuels remaining after 

residence in gas cell as a function of temperature (Kee et al. 2003).

176
Wavenumber (cm-1)

3200 3100 3000 2900 2800 2700

Absorption coefficient (atm-1

m

-1)

0

100

200

300

400

500

600

700

Extrapolation

Experiment

Figure 11 – Comparison of the extrapolated and experimentally measured spectral absorption 

coefficient of heptane (C7H16) for the C-H stretching band at 1000 K. The difference between 

extrapolated and measured spectrum is due to pyrolysis.

177