NIST Technical Note 1904 

Burning Velocity Measurements and 

Simulations for Understanding the 

Performance of Fire Suppressants in 

Aircraft 

Gregory T. Linteris 

John L. Pagliaro 

This publication is available free of charge from: 

http://dx.doi.org/10.6028/NIST.TN.1904 
 

 

NIST Technical Note 1904 

Burning Velocity Measurements and 

Simulations for Understanding the 

Performance of Fire Suppressants in 

Aircraft 

Gregory T. Linteris 

John L. Pagliaro 

Energy and Environment Division 

Engineering Laboratory 

This publication is available free of charge from: 

http://dx.doi.org/10.6028/NIST.TN.1904 

January 2016 

U.S. Department of Commerce 

Penny Pritzker, Secretary 

National Institute of Standards and Technology 

Willie May, Under Secretary of Commerce for Standards and Technology and Director 
Certain commercial entities, equipment, or materials may be identified in this 

 document in order to describe an experimental procedure or concept adequately. 

Such identification is not intended to imply recommendation or endorsement by the 

National Institute of Standards and Technology, nor is it intended to imply that the 

entities, materials, or equipment are necessarily the best available for the purpose. 

National Institute of Standards and Technology Technical Note 1904 

Natl. Inst. Stand. Technol. Tech. Note 1904, 42 pages (January 2016) 

CODEN: NTNOEF 

This publication is available free of charge from: 

http://dx.doi.org/10.6028/NIST.TN.1904 
Burning Velocity Measurements and Simulations for Understanding the 

Performance of Fire Suppressants in Aircraft1
 

Letter Report Prepared for Meggitt 

Gregory T. Linteris
 

Energy and Environment Division 

Engineering Laboratory 

NIST 

John L. Pagliaro 

MSSI Consultant 

November 20, 2015 

1.0 Introduction 

As described by Reinhardt and co-workers [1], overpressure in the Federal Aviation Administration 

(FAA) Aerosol Can Test (FAA-ACT) has been observed for the halon replacements HFC-125, Novec 

1230, and 2-BTP when added at sub-inerting concentrations. Work by Linteris and co-workers [2-5] 

has involved measurements of the flammability limits, explosion pressure and burning velocity, 

together with equilibrium, stirred-reactor, and premixed flame speed simulations to interpret this 

overpressure. Additional investigations were performed to validate the kinetic mechanisms and 

understand the chemical reaction pathways for the pure agents in air [6-11]. 

The goal of the present work is to predict the overpressure behavior of two proprietary blends (each a 

50 % mass fraction of CO2 with either Novec (C6F12O, Novec 1230) or 2-BTP (C3H2F3Br, 2-Bromo-3,3,3­

trifluoropropene), hereafter referred to as Blend A or Blend D, respectively) using the techniques and 

understanding developed in the previous work. To this end, experimental measurements of the 

flammability limits, explosion pressure, and premixed flame speed have been made, under 

conditions believed to provide the most insight into the blends’ behavior in the FAA-ACT. Following 

the approaches developed in prior work, equilibrium, stirred reactor, and premixed flame speed 

simulations have also been performed to interpret and predict the behavior of the agents in the FAAACT. Reviews of the concepts and experimental approaches are also provided to aid in the 

interpretation of the results. 

2.0 Background 

The exothermic reaction of halogenated hydrocarbons in combustion conditions has been described 

(sparsely) in the literature, for select compounds and experiments, as briefly reviewed by Linteris et 

al. [3]. As a result of the findings using the FAA-ACT, however, the potential for fire suppressants 

per se to react exothermically has recently been explored more comprehensively [2-11]. While the 

pure agents in air usually are not flammable themselves, when in the vicinity of a diffusion flame, 

they can react exothermically and contribute to the heat release (and hence, pressure rise if there is 

containment). For example, recent simulations [9] have predicted that HFC-125 added to the air 

stream in co-flow diffusion (e.g., cup-burner) flames (just below the extinguishing concentration) will 

increase the total heat release by a factor of up to three. This is the undesirable behavior leading to 

1

 Official contribution of NIST, not subject to copyright in the United States. 

1
 

the overpressure in the FAA-ACT. Moreover, it has been shown that since the agents (premixed in 

the oxidizer) themselves are reactive, their presence in the oxidizer stream increases the volume of 

oxidizer required for stoichiometric reaction of the fuel and agent. This also contributes to a higher 

overpressure. 

For premixed systems, it has also been found that some fire suppressants (e.g., Novec, HFC-125, etc.), 

when added to a premixed flame below its flammability limit, can actually increase the flame speed 

above the uninhibited condition, bringing the flame into the flammable regime. That is, the mixture 

becomes more reactive with the addition of the fire suppressant. Finally, the explosion pressure of 

lean mixtures can be higher with fire suppressant added to the air stream than without it. 

For example, Table 1 lists the effect of CF3Br, C2HF5, C6F12O, and CH2CBrCF3 on the: 1. calculated 

heat release in a cup burner, 2. stoichiometric air requirement, 3. explosion pressure, and 4. laminar 

burning velocity. Undesirable behavior contributing to overpressure in the FAA-ACT is indicated by 

a red up arrow, while desirable behavior is indicated by the green down arrow. 

Table 1 – Effect of agent addition on various test parameters related to overpressure in the FAA-ACT 

(↓ decreased; ↑ increased; 0: unaffected). 

Test Parameter: 

CF3Br C2HF5 C6F12O CH2CBrCF3 

Diffusion Flames

 Cup-Burner Heat Release ↓ ↑ No data ↑

 Stoichiometric Air Requirement 0 ↑ ↑ ↑

Premixed Flames 

 Explosion Pressure 1.0 

0.6 

↓

↓

↑

↑

↑

↑

↓

↑

 Laminar Burning Velocity 1.0 

0.6 

↓

↓

↓

↑

↓

↑

↓

? 

Our experimental approach for evaluating possible overpressure in the FAA-ACT by Blend A and 

Blend D is these last two tests: 1.) Enhanced explosion pressure of lean premixed flames; and 2.) 

enhanced burning velocity of lean premixed flames. Our numerical approach includes simulations of 

the equilibrium temperature and pressure, the overall reaction rate (as determined with perfectlystirred reactor (PSR) simulations), and premixed flame burning velocity calculations. 

2
In previous work [3-5] equilibrium and perfectly stirred reactor simulations were used to understand 

the experimental behavior of C2HF5, C6F12O, and CH2CBrCF3 in the FAA-ACT. The equilibrium 

simulations allow one to estimate the overpressure that can occur if there are no kinetic limitations. 

This was found to be the case for pure C2HF5, C6F12O, and CH2CBrCF3 at all concentrations except the 

inerting value; i.e., the equilibrium simulations predicted the pressure rise for the pure agents (C2HF5, 

C6F12O, and CH2CBrCF3) at sub-inerting concentrations. For the blends A and B, equilibrium 

calculations will again provide estimates of the fraction of chamber volume involved in the 

combustion and the maximum possible pressure rise. 

For understanding flame extinction, a measure of the overall reaction rate is required. The perfectly 

stirred reactor simulations provided these for a range of inhibitor volume fractions Xa and fueloxidizer mixing ratios. Since the actual FAA-ACT is a two-phase, turbulent explosive process, the 

mixing state of the fuel and oxidizer is unknown. However, by performing the simulations over a 

range of involved chamber oxidizer fraction , the overall reaction rate as a function of mixing state 

can be obtained. This is useful since in previous work [5], the values of the overall reaction rate 

required for flame extinguishment were determined. To extinguish the end-gases of a low-speed, 

post-explosion flame, the stirred-reactor overall rate must be reduced to about 20 s-1; if the end-gases 

are sufficiently non-flammable, however, then the flame to be extinguished is the turbulent jet near 

the impulsive release of the fuel (which is easier to extinguish and requires less agent), so that the 

overall reaction rate must be reduced to about 1700 s-1 [5]. The PSR simulations provide insight into 

the value of psr as a function of mixing state and inhibitor volume fraction, aiding in extrapolation 

from the premixed experiments to the diffusion flames of the FAA-ACT. 

3.0 Experimental Methods 

3.1 Apparatus and procedure 

 The experimental apparatus for the flammability limits and explosion pressure, as well as 

for the burning velocity measurements, is a constant-volume combustion device with an internal 

volume of about 2 L (referred to as the 2-L chamber), which has been described in previous work [10] 

and is outlined below. A schematic diagram is shown in Figure 1, and a photograph in Figure 2. 

Plumbing and electrical schematics are shown in Figure 3 and Figure 4. The stainless steel (316) 

spherical vessel is similar to previous designs [12-15], with an inner diameter of 15.24 cm, volume of 

1.85 L and wall thickness of 2.54 cm, and is equipped with an ignition system, an absolute pressure 

gage, a dynamic pressure sensor, and a thermocouple. The experiment can provide the flammability 

limits, explosion pressure (constant-volume), and rate of pressure rise; further processing of the latter 

can be used to obtain the laminar burning velocity (1-D spherical) as a function of initial pressure and 

temperature (which increase as the unburned gases are compressed). 

3
Figure 1: Schematic diagram of the 2-L experimental apparatus. 

A vacuum pump reduces chamber pressure below 13.3 Pa prior to reactant addition. Test 

mixtures are prepared in the chamber using the partial pressure method, following injection of first 

liquid then gaseous reactants. Component partial pressures are determined with an absolute pressure 

transducer (Omega, PX811; claimed accuracy of 0.1 % of reading) that is periodically calibrated 

against a Baratron 627D (claimed accuracy of 0.12 %) and a Wallace & Tiernan 1500 pressure gage 

(claimed accuracy of 0.066 %). Liquid suppressants (C6F12O and C3H2F3Br) are injected using a 

syringe and a gas-tight septum separated from the chamber by a ball valve (to ensure leak-free 

operation during the experiment). 

4
Figure 2 - 2-L chamber. 

5
   

   

Legend

         

   

   

~
 PCB 101A06

0‐500 psi, 2%

Omega PX811‐030AV

0‐30 psia, 0.1%

ball valve

check valve

needle valve

relief valve

tube

tube with minimized volume

5‐port valve

air

30 psig

vacuum

100 Torr

vent

1 atm

2 L sphere

TC, type K

Bourdon gage,

0‐600 psig

Ignitor

refrigerant

10 psig

methane

10 psig

3‐port valve

N2

170 psig

vent

1 atm

vent

1 atm

Figure 3 - Plumbing schematic diagram of 2-L chamber. 

6
 

   

   

   

   

   

   

   

100 VAC Variac

2 L sphere

PCB 101A06

0‐500 psi, 2%

TC, type K

Omega PX811‐030AV

0‐30 psia, 0.1%

Readout

TC, type K

National Instruments DAQ

Figure 4 - Electrical schematic diagram of 2-L chamber. 

 The sample gases are CH4 (Matheson Tri-Gas, 99.97 % purity), C3H8 (Scott Specialty Gases, 99.0 % 

purity), CF3Br (Great Lakes Chemical Corp., 99.6 % purity), C6F12O (3M, > 99 % purity), C3H2F3Br 

(American Pacific Corp., > 99 % purity), and carbon dioxide (MG Industries). The air is house 

compressed air (filtered and dried) that is additionally conditioned with a 0.01 μm filter, carbon filter, 

and a desiccant bed to remove small aerosols, organic vapors, and water vapor. The relative humidity 

r.h. of the house compressed air, measured with humidity gage (TSI VELOCICALC, 8386), is less 

than 2 %. 

For tests with humid air, water vapor is added to the reactants by adding liquid water (house 

deionized) via the syringe, at the mass required to obtain the desired water vapor volume fraction in 

the air. After liquid water addition, the total pressure is monitored to insure that the partial pressure 

of water reached the vapor phase. In some cases (e.g., after tests with C3H2F3Br), the water is 

sometimes absorbed into deposits on the chamber wall surfaces, preventing the total pressure from 

achieving the calculated value. In these cases, somewhat more liquid water (≈50 %) is added until the 

correct value of the water partial pressure is achieved (corresponding to 100 % r.h. at 20 °C, or a water 

volume fraction of 0.023). After addition of all reactants, a 5 min settling time is used to insure 

dissipation of any turbulence, as described in ref. [12]. 

7
 

 

Two ignition systems are used in the present work: an electrical spark and an exploding platinum 

wire. The spark ignition system is used for the burning velocity measurements, and some of the 

flammability limit and explosion pressure tests (as noted below). The electrical spark method is the 

more common ignition method, and provides the minimum disruption of the developing premixed 

flame (which is necessary for extracting the laminar burning velocity from the pressure rise data). 

The platinum igniter, based on the Japanese High Pressure Gas Law (JHPGL, a translation of which is 

available in ref. [16]), somewhat overdrives the ignition process, but is useful since it provides a much 

more conservative estimate of the flammability limits. It is used in some tests in the present work 

since it was found to more accurately represent the flammability limits of mixtures in FAA Aerosol 

Can Tests. 

 The electrical spark ignition system (capacitive) is based on the work of Shepherd et al. [17]). A 

1 kV to 15 kV power supply (Acopian) and custom-made capacitor banks (1 nF to 50 nF) provide 

variable ignition energies, with an estimated operating range of 0.05 mJ to 500 mJ. Two tungsten 

electrodes form a gap in the center of the chamber. Thin electrodes (0.4 mm diameter) minimize heat 

loss from the flame, and the spark gap is adjustable (2 mm, typical). Ignition is attempted several 

times, while gradually increasing the capacitor charging voltage, until ignition occurs. This ensures 

the ignition energy is within an order of magnitude of the minimum value. (Note that for 

stoichiometric iso-octane–air mixtures, Marshall et al. [18] found that the burning velocity was 

insensitive to the ignition energy for values up to 1000 times the minimum ignition energy.) 

 The platinum wire ignition system is based on that in the JHPGL. A variable transformer AC 

power supply supplies 100 VAC, controlled by a manual switch, to a thin platinum wire. The 

platinum wire melts and ruptures violently during each ignition process, and is replaced for each 

test. The igniter configuration is modified slightly from that in the JHPGL. Rather than using igniter 

leads that enter the chamber from two locations at right angles to each other (per the JHPGL), the 

present igniter (shown in Figure 5) uses two parallel copper leads (57 mm long, 1 mm diameter) 

separated by 4 mm, with crimp-on connections (Digi-Key A34501-ND and A2161-ND) between the 

copper and platinum wires instead of welds. Hence, the igniter could be inserted through a single 

0.25 inch fitting, with easily replaceable fusible wire. The fusible portion of wire consists of a 20 mm 

length of Pt wire formed in a loop. In previous work [16], tests were conducted to explore the 

influence of the wire material and ignition energy, and showed that the platinum wire provided 

wider flammability limits, whereas the other materials (tungsten and copper) provided results more 

like those resulting from spark ignition. The difference was attributed to the catalytic action of the 

molten platinum droplets and the turbulence that they create in the explosion process. 

Figure 5 – Platinum igniter. 

8
 

 With either igniter, the explosion pressure is recorded at 4000 Hz. via a dynamic pressure sensor 

(PCB Piezotronics, 101A06; claimed accuracy of 0.1 % of reading). The product gases are immediately 

purged to vacuum via a large flow of N2 (to minimize heating of the chamber and its exposure to acid 

gases), and the chamber is allowed to cool for 20 minutes between experiments. Figure 6 shows a 

typical pressure trace (curve labeled “P”) obtained from the experiment (other curves in the figure are 

described below). The laminar burning velocity is determined from the pressure trace using a 

thermodynamic model, developed by Metghalchi and Keck [12, 19] and further refined by others [13, 

20]. A brief overview of the method follows. 

Figure 6: Experimental pressure trace P, flame radius rf, flame stretch rate ࣄ, and gas temperature 

(unburned Tu and burned Tb) as a function of time. 

3.2 Burning velocity from the pressure trace 

In the analytical analysis of the propagation of the combustion wave in the chamber, the contents 

of the chamber are divided into burned and unburned zones separated by a reaction sheet, assumed 

to be of zero thickness, spherical, and smooth (no instabilities). Initially, the unburned gas is 

considered mixed and at rest. As the unburned gases react, a spatially uniform increase in pressure 

occurs. The burned gas is in chemical equilibrium and both the burned and unburned gases are 

9
 

 

considered as ideal, semi-perfect gases. Both zones are adiabatic, and the unburned gas is 

isentropically compressed as the mixture reacts in the flame sheet.

 With these assumptions, the instantaneous flame radius and burning velocity can be expressed in 

terms of the chamber pressure by applying conservation of mass. The results are given in Eq. 1 and 

Eq. 2 (detailed formulation of the equations can be found in refs. [12, 20]), 

ݎ ௙ൌ ܴൣ /ܲሻଵ/ఊೠ ܲ଴ ሻሺ 1െݔ ௕1 െ ሺ ଵ/ଷ ൧

/݀ݐሻ ݀ݔ ௕ೠሺ ଵ/ఊ /ܲሻ ܲ଴ ሺଶ ܵ௅ ൌ ܴ/3ሺܴ/ݎ௙ሻ

 (1) 

 (2)

in which ݎ ௙is the flame radius, ܵ௨ is the laminar burning velocity, ܴ is the chamber radius, ݔ ௕is the

mass fraction of burned gas, ܲ is the instantaneous pressure, ܲ଴ is the initial pressure, and ߛ௨ is the

unburned gas specific heat ratio. To determine burning velocity, the mass fraction of burned gas ݔ௕

must be related to the chamber pressure (for simplification, early studies applied a linear relationship 

[21]). Alternatively, a thermodynamic closed system analysis on the contents within the chamber can 

be performed. A two-zone approach is used in the present work. The burned gas zone also can be 

modeled using incremental shell volumes [22, 23], allowing burned gas density and temperature 

gradients; nonetheless, since previous studies have shown that inclusion of the gradients has 

negligible effect on the burning velocity [19, 20], we retain the two-zone model. The fraction of mass 

burned ݔ ௕is found from simultaneous solution of the conservation of mass and energy equations 

given in Eq. 3 and Eq. 4, applied for the two zones, 

௏ ௫್ ଵ

್௫ ݔ݀௨ݒ

ெ ׬ ൌ ଴ ׬ ൅ ݔ݀௕ݒ

ா ௫್ ଵ

݁௨݀ݔ ௫್

ெ ׬ ൌ ଴ ׬ ൅ ݔ݀௕݁

 (3) 
 

 (4)

where V is the volume of the chamber, ܧ is the total internal energy of the gases in the chamber, M is

the mass of the initial gas in the chamber, e and v are the internal energy and specific volume of the 

gas, and the subscripts b and u refer to the burned and unburned gas. 

 Thermodynamic data for the unburned and burned gases are required for model implementation. 

Data for hydrocarbon-air species (CH4, O2, N2, CO2, H2O, CO, NO, OH, H2, and O) are taken from 

GRI-mech 3.0 [24], fluorinated species (C6F12O, C2HF5, F, HF, CF4, and CF2O) from the National 

Institute of Standards and Technology (NIST) Hydrofluorocarbon (HFC) mechanism [25, 26], and 

brominated species (CF3Br, C3H2F3Br, Br, HBr, and Br2) from Babushok et al. [27, 28]. Hydrocarbonair product species are included for all mixtures, along with fluorinated products for test containing 

C6F12O. Additional brominated products are included for tests with C3H2F3Br. 

 The unburned gas properties are related to the chamber pressure through isentropic compression: 

௉ ܶ௨ ൌ ܶ௨଴ ቀ

௉బ

where ܶ௨଴ is the initial unburned temperature. Since

ቁ

ሺఊೠିଵሻൗఊೠ (5) 

ߛ௨ is a function of the unburned gas 

temperature ܶ௨ and the initial mixture composition, ܶ௨ and ߛ௨ are solved iteratively at each pressure

increment. The properties of the unburned gas mixture, ݒ௨ and ݁௨, are determined from the mixture

composition and temperature. 

10
 

 

 

 

The remaining unknowns in the conservation equations (ݒ௕, 

proper values of ܶ௕ and ݔ ௕are obtained. Burned gas species concentrations, required for the

determination of ݒ ௕and ݁௕, are estimated for each value of T and P using a constant volume

݁௕, and ݔ (௕are found through

iteration of ܶ௕ (ݒ ௕and ݁௕ are functions of temperature) and ݔ ௕at each pressure increment, until the 

equilibrium calculation performed via the CEA2 routine of Gordon and McBride [29]. The enthalpies 

of formation of C6F12O, C2HF5, and C3H2F3Br (required for CEA2 calculation) are from refs. [3, 4, 25]. 

Once ݔ௕ሺܲሻ is known, the burning velocity ܵ௨ሺܲ, ܶ௨ሻ is calculated over the experimental range of

pressure and temperature using Eq. 2. 

3.3. Data reduction 

As described above, Figure 6 shows a typical pressure trace (solid line) from an experiment (CH4­

air, =1.0). Also shown are Tb, Tu, and rf, which are outputs of the two-zone model calculation. Only 

a portion of the pressure data is used for obtaining burning velocity, as denoted by the thick line on 

the pressure trace. For small ݎ௙, the flame behavior is affected by flame stretch and the ignition

process, and for large ݎ௙, by heat losses to the walls and cellular instabilities; hence, typically only the 

central 75 % of the pressure data are used. Spherically propagating flames are subject to stretch rates 

inversely proportional to the flame radius [30], 

ൌ ߢ

ଶ ୢ௥೑ (6) ௥೑ ୢ௧

where ߢ is the stretch rate and dݎ௙/dݐ is the flame front velocity. Figure 6 also shows the stretch rate 

for this experiment. To reduce stretch effects (as well as the transient caused by the ignition), data are 

neglected for small flame radii, ݎ>௙3.8 cm (i.e., ݎ>௙R/2, as proposed by Elia et al. [31], and adopted by

=0.5, peak stretch rates range between 110 s-1 others [18, 22, 32]). For ݎ௙/ܴ for uninhibited 

stoichiometric methane-air and 20 s-1 for inhibited cases with burning velocities near 6 cm/s. To 

avoid the effects of heat losses to the walls, only data up to dP/dtmax (i.e., the inflection point in the 

P(t) curve in Figure 6) are used, following the recommendations in refs. [33, 34]. 

 A single experiment provides burning velocity data for a range of pressure and temperature of the 

unburned gas, and these data are fit to the equation: 

் ܵ௅ ൌ ܵ௅,଴ ቀ

்బ

where ܵ௅ is the laminar burning velocity, ܶ଴ is the initial temperature, ܵ௅,଴ is the laminar burning

are the fitting parameters. In the experiments performed, velocity at the initial conditions; ߙ and ܵ௅,଴

ܶ଴ is 293 ±2 K.

ቁ

ఈ

 (7) 

Figure 7 shows the burning velocity of stoichiometric methane-air as a function of 

unburned gas temperature and chamber pressure (related through isentropic compression i.e., Eq. 5). 

The diamonds are the experimental data derived from Eq. 2, and the line is the fit to Eq. 7. In 

discussions following, SL is presented at ambient conditions as obtained from Eq. 7 (and shown as 

black dot in Figure 7). 

11
P (bar) 

1.1 2.0 3.0 4.0 5.0 6.0

S L (cm/s) 

290 340 390 440 490
 

45

40

35

30

Tu (K) 

Figure 7: Plot of burning velocity as a function of pressure P and unburned gas temperature Tu. 

The diamonds represent the data from an experiment, the line represents the fitted results using 

Eq. 7, and the black dot is the reported SL at ambient (T0=298 K, P0=1 bar) conditions. 

In the present method, spherical flame propagation is a critical condition for accurate 

determination of SL from the pressure trace. Buoyancy can distort the shape of the flame, especially 

for slow burning mixtures (which are of particular interest in the present work). Takizawa et al. [13] 

estimated the minimum burning velocity measurable without error associated with buoyancy. They 

performed experiments in a spherical chamber, using the pressure rise to determine burning velocity. 

Separate experiments were performed in a cylindrical chamber with visual access to provide 

comparison. The shape of the flame front was recorded with high speed video and the burning 

velocity was calculated using the constant-pressure method [35, 36]. To minimize the effects of 

buoyancy, as recommended by Pfahl et al. [37], the rate of change of the flame radius with respect to 

time was traced in the two horizontal directions. Results [13] showed that burning velocities as low 

as 6 cm/s could be measured with the constant-volume method. In the present work, cellular 

instabilities, which also invalidate the spherical flame assumption, are monitored through inspection 

of the SL data of individual test runs. The onset of cellular instabilities is typically detected via a 

distinct increase in SL [18, 22], and these data (if occurring) are omitted during the experimental data 

fitting. 

3.4. Fuel Selection 

The fuel in the FAA_ACT is intended to simulate the contents of an aerosol spray can, and contains 

5.87 moles of ethanol, 2.05 moles of propane, and 5 moles of liquid water. Although it would be most 

desirable to run experiments for the same conditions, it is not possible, since the water content of 

these tests is greater than 100 % r.h. at room temperature. In order to overcome this limitation, yet 

collect data for a representative fuel system, we analyzed various surrogate fuels. The best option 

was to use propane-air mixtures at 100 % r.h. This fuel system best approximated the adiabatic flame 

temperature, and the hydrogen atom content of the total fuel-air mixture in the FAA-ACT, and was 

12
 

4 

adopted for the present experiments. Nonetheless, for the simulations, we used the actual FAA-ACT 

test fuel whenever modeling the FAA-ACT, and of course, for modeling experiments with the moist 

propane-air fuel, we used the propane and water vapor content as measured in the initial conditions 

of the experiment. 

3.5. Uncertainties 

 The uncertainty analysis consists of individual uncertainty components and root-sum-of-squares 

(RSS) components [38]. Uncertainties in the measured parameters are reported as expanded 

uncertainties: kuc, from a combined standard uncertainty (estimated standard deviation) uc, and a 

coverage factor k=2 (level of confidence approximately 95 %). Likewise, when reported, the relative 

uncertainty is kuc/X. Uncertainties in initial temperature and pressure, dynamic and peak pressure 

rise, agent volume fraction, equivalence ratio, and the burning velocity calculation and fitting to Eq. 7 

are considered. The expanded relative uncertainties related to mixture composition are as follows: 

0.5 % for equivalence ratio; and 0.4 %, 0.8 %, and 4 % for the reactant mole fractions of air, CH4, and 

agent. Relative uncertainties of the dynamic and peak pressure rise, initial pressure, and initial 

temperature are 1.3 %, 0.5 %, and 1 %. The expanded relative uncertainty for experimentally 

determined burning velocities is 12 %. 

3.6 Validation of Experimental Techniques 

Flammability limits 

Using the present experimental technique, the flammability limits of CH4-air mixtures and CH2F2­

air mixtures (as a function of stoichiometry) have been determined in previous work [16], and are 

shown in Figure 8. As illustrated for methane (left frame), the copper wire igniter reproduces the 

traditional flammability limits, and the platinum wire igniter provides slightly wider limits. For 

CH2F2-air mixtures, the previously measured flammability limits are well reproduced by the tests 

with the exploding platinum wire. 

LFL UFL 

Cu Igniter

Pt Igniter

LFL UFL 

4 

Chamber Pressure Rise / bar

3 

2 

Chamber Pressure Rise / bar 3 

2 

1 

0 

0 5 10 15 20 25 30 0 5 10 15 20 25 30 

CH4 Initial Partial Pressure / % CF2H2 Initial Partial Pressure / % 

Figure 8 - Chamber pressure rise in the 2-L chamber as a function of fuel partial pressure, using 

either Pt or Cu wire igniter, together with literature values of the flammability limits (vertical 

lines) [39, 40]; left frame: CH4–air, right frame: CH2F2-air. 

13

0 

1 
 

 

Burning velocity 

To validate the accuracy of the experimental facility and the post-processing procedure, the 

burning velocity of methane-air flames was determined over a range of equivalence ratios (0.6 to 1.3). 

Figure 9 compares the present results to published data at standard (298 K, 1 bar; lower curve) and 

compressed (400 K, 3 bar; upper curve) conditions. The solid black squares show the present data, 

other black symbols show data collected using the same experimental technique [13, 14], blue 

symbols show stretch-corrected data from spherical flames [41-43], and red symbols show stretchcorrected data from counter-flow flames [44, 45]. 

 For the initially ambient mixtures, the burning velocities are in excellent agreement with 

previous results using the constant-volume method with a similarly sized chamber and a two-zone 

model [13]. Values are within 1 % at all equivalence ratios except for =0.7 and =1.2 where SL is 5 % 

higher and 3 % lower (the symbols in Figure 9 representing the data of ref. [13] are not visible because 

they are so closely aligned with the larger square symbols showing the present data). Results are 

within 5 % of ref. [14], except at =1.2 where SL is 8.5 % lower. Burning velocities are also in 

satisfactory agreement with stretched-corrected spherically propagating [41-43] and counterflow [44, 

45] flame data, within the scatter of reported values for the entire range of . (For comparison, the 

same experimental P vs. t data were post-processed using the linear relationship between P and xb 

[21]. This technique yielded burning velocities that were roughly 8 % higher than the two-zone model 

approach, with SL=38.9 cm/s at =1.0. Similar findings were reported by refs. [34, 46]; thus, the twozone model approach was selected for post-processing the inhibited flame data.) 

The upper set of symbols in Figure 9 compares burning velocities at the compressed conditions 

with previous results using the same constant-volume method. The present burning velocities are in 

excellent agreement with the results of Stone et al. [14] and in satisfactory (+-9 %) agreement with 

Takizawa et al. [13]. Overall, the methane validation results show that the present experimental 

approach provides burning velocities in agreement with those of other investigations at standard and 

compressed (400 K, 3 bar) conditions. 

Figure 9: Burning velocity of premixed methane-air flames at 298 K and 1 bar (lower) and 400 K 

and 3 bar (upper) as a function of equivalence ratio, together with previously published results. 

14
4. Numerical Methods 

4.1 Overview 

Numerical simulations were performed to provide insight into the potential for pressure rise and the 

chemical reactivity of the system, as a function of fuel-air stoichiometry, as well as agent volume 

fraction. Simulations included chemical equilibrium, premixed steady 1-D planar laminar flame 

speed, and stirred reactor residence time just above blow-out. The National Aeronautics and Space 

Administration (NASA) equilibrium code CEA2 [29] and the SANDIA numerical codes PREMIX [47] 

and PSR [48] were used, along with the chemical kinetics [49] and transport [50] interpreters. 

The equilibrium composition, temperature, and pressure for a mixture of gases at some given initial 

state can be calculated [39, 51]. Automated numerical techniques are available, which use the method 

based on minimizing the Gibbs free energy, for a large number of species typically present in 

combustion systems. The equilibrium conditions of the aerosol can test were calculated using the 

CEA2 of Gordon and McBride [29]. The calculations were performed over a wide range of initial 

conditions including inhibitor volume fractions (Xi) from 0 to 0.30, with a water vapor volume 

fraction of 0.0125, and fractions of chamber air involved () of about 0.15 to 1.00. The equilibrium 

simulations were used to predict the final temperature of the involved reactants, and from this, the 

maximum pressure of the FAA test chamber. The adiabatic combustion temperature and the 

predicted pressure rise were presented in a previous publication for pure C6F12O, C3H2F3Br, and 

CF3Br [10], and these previous results are presented in the present work for comparison. 

PSR simulations were used to understand kinetic limitations associated with the explosion pressure 

predictions of the equilibrium simulations. Flame extinction caused by suppressants is controlled by 

the characteristic times for chemical reaction and transport, as described by the Damköhler number 

Da=r/c, in which r is the flow residence time, and c is the chemical time [52]. Hence, an important 

step for understanding flame suppression is to estimate the overall reaction rate. Given the explosive, 

two-phase, turbulent mixing process occurring during release of the aerosol can test simulator 

fuel [53], the reaction zone might be simulated reasonably well by a PSR. Additionally, the PSR 

blow-out residence time has been correlated with both the laminar flame speed [54] and extinction of 

laminar diffusion flames with added inert suppressants [55], indicating its utility as a measure of 

overall combustion reaction rate. For example, to obtain the characteristic chemical time at extinction 

using a stirred-reactor model [48], the blow-out conditions are determined, as described previously 

[3, 55], and the overall chemical rate psr is the inverse of the chemical time. While any enhanced 

mixing in the FAA-ACT occurring due to turbulence cannot be modeled directly with the PSR 

simulations, the simulations can bracket the range of mixing conditions that might occur, and 

quantify the overall reaction rate therein. As with the equilibrium comparisons, the results of the PSR 

simulations for the pure compounds (Novec and 2-BTP) are compared with the blends. 

As an illustration of the utility of stirred-reactor chemical rates for predicting flammability, Figure 10 

plots, for a premixed methane-air system, the adiabatic flame temperature, stirred reactor overall 

rate, and the laminar burning velocity as a function of methane volume fraction in the mixture. As 

illustrated, good correlation is shown between the traditional flammability limits, and either the 

laminar burning velocity or the stirred reactor overall rate. 

15
2000
 

10000
 

T / K 

Tad psr 

LFL UFL

SL x 250 

0 5 10 15 20
 

1500
 



psr / s-1 or 

SL / cm s-1 

1000 

5000 

500 

0 0 

Methane Volume Fraction 

Figure 10 – Calculated overall chemical rate from stirred reactor simulations psr, laminar burning 

velocity, and adiabatic flame temperature for methane-air mixtures as a function of methane 

volume fraction [16]. Vertical lines show the separately measured flammability limits of methane 

[39]. 

4.2. Kinetic Mechanisms 

The kinetic mechanisms used for the simulations have been developed in separate work [4, 5]. For 

both mechanisms, the hydrocarbon sub-mechanism is the C1–C4 model of Wang et al. [56] (111 

species and 784 reactions), with the addition of 5 species and 36 reactions related to ethanol 

combustion, from Dryer and co-workers [57]. The starting mechanism to describe reactions of the 

hydrofluorocarbons in hydrocarbon flames is the NIST HFC mechanism [25], including modifications 

suggested in more recent work [58, 59], also as summarized in Ref. [3]. In addition, more recent 

modeling of the combustion of pure fluorocarbon agents [60] suggests that relatively large 

concentrations of F atoms are observed in the reaction and post-flame zones. Thus, some formation of 

F2 is expected, and reactions related to its formation and consumption (not previously considered) 

have now been added [7]. 

The kinetic model for flame inhibition by C3H2F3Br is based on a decomposition model developed at 

NIST [6], which describes the reactions down to Br-containing species of C2 species and smaller. 

16
Subsequent reaction of these species are from the kinetic model for CF3Br inhibition of hydrocarbonair flames [27] with more recent modifications. The complete C3H2F3Br sub-model includes 96 

reactions of Br-containing species, as presented in ref. [6]. Thermodynamic data for most of the Brcontaining species are from ref. [27] or Burcat et al. [61], while data for the set of species related to the 

decomposition of 2-BTP are estimated or calculated, as described in ref. [6]. The complete kinetic 

model for the simulations with C3H2F3Br has 1609 reactions and 188 species, and is available in ref. 

[6]. 

The kinetic model for C6F12O flame inhibition is an approximate model, outlined in ref. [4]. The 

initial C6F12O steps were modelled (which leads to C2HF5 and C3HF7), and the rest of the fluorinatedspecies mechanism was the same as in the C3H2F3Br mechanism described above. 

4.3 Kinetic Model Validation 

As initial validation of the kinetic mechanisms for flame inhibition by pure C6F12O and C3H2F3Br, the 

burning velocity of premixed CH4- and C3H8-air flames inhibited by these compounds has been 

measured and predicted in separate work. For addition of pure C6F12O, Figure 11 shows the 

measured (points) and predicted (lines) unstretched laminar burning velocity for C3H8–air flames (of 

various equivalence ratios) with added C6F12O [62]. As indicated, the model predictions are very 

close to the experimental results, except for the very lean flames (=0.6), for which the uncertainty in 

the measured burning velocity is large (± 2 cm/s) relative to the measured burning velocity (3 cm/s 

to 10 cm/s) (likely due to radiative heat losses, buoyancy, and stretch effects, which become more 

important at low burning velocity). For C3H2F3Br addition, Figure 12 shows equivalent results for 

CH4- and C3H8–air flames [11]. Again, the agreement is generally excellent, except for the very lean 

flames (=0.6), which have higher uncertainty in the measured burning velocity. These results 

provide confidence in the predictive ability of the premixed flames simulations for addition of C6F12O 

and C3H2F3Br to hydrocarbon flames, since similar simulations are presented below for the blends of 

these agents with CO2. Similarly, the stirred reactor simulations are expected to be accurate, since 

they use the same kinetic mechanism. 

17
Su

0 (cm/s) Su

0 (cm/s) 

40

5

10

15

20

25

30

35 Φ=1.0 

Φ=1.2 

Φ=0.8 Φ=0.6 

5

10

15

20

25

30

35 Φ=1.0 

Φ=1.2 

Φ=0.8 

Φ=0.6 

0 0

0 0.01 0.02 0.03 0.04 0 0.01 0.02 0.03 0.04

C6F12O volume fraction C6F12O volume fraction 

Figure 11 - Comparison of experimental (symbols) and numerical (lines) Su0 for C6F12O inhibited 

CH4-air (left frame) and C3H8-air (right frame) flames at =0.6, 0.8, 1.0, and 1.2. 

Su

0 (cm/s) CH4-air Su

0 (cm/s) C3H8-air 

40

35 Φ=1.0 

35

30

30

25

25

20

20

15

15

10 10

5 5

0 0

0 0.01 0.02 0.03 0.04 0 0.01 0.02 0.03 0.04

C3H2F3Br volume fraction C3H2F3Br Volume Fraction 

Φ=1.2 Φ=0.8 

Φ=0.6 

Φ=1.0 

Φ=1.2 Φ=0.8 

Φ=0.6 

Figure 12 - Comparison of experimental (symbols) and numerical (lines) Su0 for C3H2F3Br-inhibited 

CH4-air (left) and C3H8-air (right) flames at  = 0.6, 0.8, 1.0, and 1.2; open symbols are non-linearly 

extracted, bars are linearly extracted, and crosses are from ref. [10]. 

18
5. Results and Discussion 

5.1. Experimental Results 

5.1.1. Flammability of Agent-Air Mixtures 

Since the end-gases in the FAA-ACT may be nearly all air and suppressant (i.e., containing little of 

the fuel from the aerosol can contents), it is useful to explore the flammability of the air-agent 

mixtures (both for pure agents and their blends with CO2). Using the 2-L chamber, the flammability 

of both the pure agents and their blends were tested, using the platinum wire ignition system 

described above. For pure C6F12O or C3H2F3Br addition to humid air at stoichiometric conditions 

(agent volume fractions of 0.065 and 0.0775 and water vapor volume fraction of 0.023), the pressure 

did not rise above that caused by the igniter itself (about 0.07 bar); i.e., they were non-flammable, 

even with the energetic platinum igniter. The blends of the agents with CO2 were added at blend 

volume fractions of 0.30; for both blends, there was no significant pressure rise, again indicating no 

ignition. This is the highest concentration of agent expected, and is believed to be the most 

flammable concentration. Hence, like the pure agents themselves, at ambient conditions, the blends 

are non-flammable. This is not surprising since the inerting mole fraction of CO2 for hydrocarbon-air 

mixtures is around 0.30, and lean agent-air mixtures would be much less flammable. Nonetheless, 

under heated conditions the pure agents become flammable (as discussed below in Section 5.3.2. Pure 

Agent-Air Premixed Flames). Consequently, the potential flammability of the CO2/agents blends at 

elevated temperatures is discussed in that section (we do not currently have capabilities to measure 

flammability limits at elevated temperatures). 

5.1.2. Explosion Pressures and Inerting Concentrations 

5.1.2.1. Experimental Data 

As discussed in previous work [10], the explosion pressure of an agent/air mixture can describe the 

potential exacerbation of overpressure caused by agent addition, particularly if the experiments are 

conducted over a range of equivalence ratio . For example, for the pure compounds C6F12O, 

C3H2F3Br, and CF3Br, Figure 13 (right frame), Figure 14, (right frame) and Figure 15 [10] show the 

experimentally measured explosion pressure in the 2-L chamber for premixed mixtures of methaneor propane-air, for =1.0 and 0.6. As indicated in Figure 15, CF3Br lowers the explosion pressure at 

all loadings and all values of  that were tested. For C6F12O (Figure 13, right frame), however, agent 

addition increases the explosion pressure somewhat (up to about 10 %) for stoichiometric flames 

(=1.0), and rather strongly (about 220 %) for lean flames (=0.6). As has been discussed, this is due 

to the competition between the fuel effect of the agent (which increases the temperature and hence 

reaction rate with added C6F12O to lean flames) and the slower kinetics of the fluorinated system 

(compared to the hydrocarbon system) [10]. For pure C3H2F3Br addition (Figure 14, right frame), the 

explosion pressure of stoichiometric flames increases insignificantly (about 1 % at Xa=0.01) and 

decreases for Xa>0.01. The lean flames (=0.6) are inhibited at low C3H2F3Br volume fraction (around 

Xa=0.01) (due to the strong catalytic radical recombination by Br-species at lower temperatures [7]), 

but above this value of Xa, the explosion pressure rises (at Xa=0.03) to about 175 % of the uninhibited 

value, and then drops for Xa>0.03. 

The results for the Blend A are shown in Figure 13 (left frame). As indicated, the behavior 

(surprisingly) is similar to that of pure C6F12O: leaner flames show enhancement of the explosion 

pressure with agent addition. At =1.0, the explosion pressure is monotonically reduced for 

19
20

increasing Xa, but at lower values of , the explosion pressure is increased with Blend A addition (up 

to about Xa=0.10), and then reduced. This figure can also be used to estimate the inerting 

concentration (the concentration of agent that will prevent explosion for all values of ). Slight 

extrapolation of the curves in Figure 13 indicates that a volume fraction of Blend A Xa of 0.27 is 

sufficient to inert these mixtures. 

ΔPmax (bar) ΔPmax (bar) C6F12O 

8 8 Phi=1.0 

7 =1.0 

6 6

Phi=0.6 

5

4

Phi=0.55 4 =0.6 

3

2

Phi=0.5 

2
 

1
 

0

Phi=0.3 Phi=0.4 

0 ‐1

012345 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35

Blend A Volume Fraction Agent volume fraction (%) 

Figure 13 – Measured explosion pressure in the 2-L chamber for moist propane air flames of 

various , as a function of the volume fraction of Blend A in the initial mixture (left frame); or pure 

C6F12O in dry methane-air flames (right frame) . 

ΔPmax (bar) ΔPmax (bar) C3H2F3Br 

8 8

Phi=1.0 

7

=1.0 

6 6

Phi=0.6 5

4

4

Phi=0.55 

3

=0.6 2

2 1

Phi=0.5 

0

Phi=0.4 Phi=0.3 

0 ‐1

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 012345

Blend D Volume Fraction Agent volume fraction (%) 

Figure 14 – Measured explosion pressure in the 2-L chamber for moist propane air flames of 

various , as a function of the volume fraction of Blend D in the initial mixture(left frame); or pure 

C3H2F3Br in dry methane-air flames (right frame). 
4

The results for Blend D are shown in Figure 14 (left frame). For stoichiometric flames, the explosion 

pressure with added Blend D is monotonically reduced as Xa increases. For the lean flames, while the 

shape of the curves mimics that of pure C3H2F3Br (i.e., decreasing at very low agent loading, 

increasing for intermediate, and then decreasing again), the explosion pressure is never higher with 

added Blend D. That is, the explosion pressure behavior of Blend D is unlike C3H2F3Br, C6F12O, or 

Blend A, for which enhanced explosion pressure was observed. Apparently, the reduction in 

temperature caused by the CO2 addition sufficiently moderates temperature increase from reaction of 

the C3H2F3Br. Moreover, as discussed in ref. [7], the catalytic cycle for radical recombination from Brspecies is more effective at lower temperature, and then becomes less effective as the temperature is 

increased with more agent. To explore these features in more detail, the experimental data in Figure 

13 and Figure 14 are plotted in alternative form below, together with the results of equilibrium 

simulations for the conditions of these experiments. As with Blend A, the inerting concentration of 

Blend D for moist propane-air flames can be determined from the present explosion sphere 

experiments (Figure 14, left frame), and is found to be Xa=0.18. 

ΔPmax (bar) CF3Br 

8

6 =1.0 

2

0

=0.6 

0 1 2 3 4 5

Agent volume fraction (%) 

Figure 15 – Measured explosion pressure in the 2-L chamber (Pt-wire igniter) for methane-air 

flames (=1.0, and 0.6) as a function of the volume fraction of CF3Br in the initial mixture [10]. 

5.1.2.2. Equilibrium Simulations and Experimental Data 

Figure 16 re-plots the measured explosion pressure data for methane-air flames from ref. [10] (=1.0 

left frame, 0.6 right frame) with added CF3Br, C6F12O, and C3H2F3Br. The top curves (right axes) 

show the calculated equilibrium adiabatic flame temperatures, while the middle curves (left axes) 

show the calculated equilibrium explosion pressure (lines) and the bottom curves, the measured 

values (points with connecting lines). As discussed previously [10], the potential for overpressure 

with added suppressant is illustrated by the equilibrium curves, and exists even for addition to 

21
 

500

0

500

1000

1500

12

14

500

0

500

1000

1500

12

14

 0 1 2 3 4 5

‐

C6F12O 

C3H2F3Br 

CF3Br 

Equil 

Exp 

Tad 

stoichiometric flames (=1.0). Due to heat losses, kinetic limitations, and buoyancy-induced 

quenching, however, the experimental pressure rise is always lower than the equilibrium values. 

Most noteworthy is that the results for agent addition to the lean (=0.6) methane-air flames (right 

frame of Figure 16) indicate that a large enhancement of the explosion pressure is predicted (and 

experienced) with addition of any of the pure agents (except CF3Br). 

Φ=1.0 Φ=0.6 

ΔPmax (bar) Tad (K) ΔPmax (bar) Tad (K) 

2500 2500

2000 2000

10 10

8 8

6 6

4 4

2 2

0 ‐ 0

0 1 2 3 4 5

C6F12O 

CF3Br 

C3H2F3Br 

Tad 

Equil 

Exp 

Agent volume fraction (%) Agent volume fraction (%) 

Figure 16 – Measured explosion pressure (Pt-wire igniter) in the 2-L chamber for methane-air 

flames (=1.0 left frame; =0.6, right frame) as a function of the volume fraction of CF3Br, C6F12O, 

and C3H2F3Br volume fraction in the initial mixture [10]. Top curves show the calculated 

equilibrium temperature, and the bottom curves show the equilibrium and measured explosion 

pressure. 

In Figure 17, data for Blend A and Blend D are presented (together) in the same format as in Figure 

16. In Figure 17, different frames refer to different values of , as indicated, and the blue solid lines 

refer to Blend A, while red, dotted lines refer to Blend D. As the figure shows, agent addition can 

either increase or decrease the equilibrium explosion pressure depending upon Xa and . Moreover, 

the predicted equilibrium explosion pressure (and adiabatic combustion temperature) of Blend A and 

Blend D are very close, with each sometimes higher than the other depending upon the Xa and . 

For the experimental explosion pressure, however, addition of Blend D always lowers the explosion 

pressure, and Blend D is always more effective than Blend A. 

As noted, both the equilibrium adiabatic combustion temperature and the calculated equilibrium 

explosion pressure as a function of Xa are similar for the two agents. Since the actual explosion 

pressure for addition of Blend D is always lower than for addition of Blend A, Blend D has greater 

kinetic inhibition effect on the flames than does Blend A. In the 2-L spherical flame experiments, it is 

difficult to separate the effects of stretch-induced quenching (due to buoyancy-induced flow, which 

lowers the temperature) and purely kinetically-induced quenching (i.e., lowered extent of reaction 

22
due to slower chemistry of the halogen-containing system). A lowered overall reaction rate reduces 

the flame speed, allowing buoyancy more time to act, which then causes vortex formation and flowinduced quenching, further lowering the reaction rate. Nonetheless, the experimental results in 

Figure 17 can be used as rough estimates of the kinetic effects of agent addition; increased deviation 

of the experimental pressure rise from the equilibrium value (when agent is added) is due to kinetic 

effects of the agent. For example, for Blend A in lean flames, at Xa=0.10 the fuel effects dominates 

over any kinetic effects, and at Xa=0.20, the kinetic effects dominate. For Blend D, kinetic effects 

always dominate (adding agent to mixtures of any stoichiometry always increases the deviation of 

the actual pressure rise from the equilibrium results). 

5.2 Numerical Results 

The predicted performance of the blends in the FAA-ACT is explored via thermodynamic 

equilibrium simulations (for temperature and hence pressure rise), and perfectly-stirred reactor (PSR) 

simulations (to provide a measure of the overall reaction rate, which is related to the extinction 

condition). In the figures below (Figures 18 to 23), the left frames refer to the blends, and the right 

frames refer to the pure agents (for which the data and figures are from previous work for C6F12O [4] 

and C3H2F3Br [5]). 

5.2.1. Equilibrium Simulations 

5.2.1.1 Blend A 

Figure 18 shows the calculated equilibrium adiabatic combustion temperature Tad for the FAA-ACT 

as a function of the fraction of the chamber volume  that is assumed to react with the fuel of the 

simulated aerosol can contents (ethanol, propane, water). For example, in the figure, the curve 

marked Xa=0 refers to the uninhibited oxidizer mixture (i.e., air at 50 % r.h.), and shows that about a 

third of the chamber volume =0.29 is required for peak temperature. As agent is added (Xa 

increases), the curves change as follows: 1. shift down (due to the high fraction of CO2 in the blend, 

which lowers the temperature), 2. to the right (due to the oxygen demand of the C6F12O in the blend 

and O2 displacement by the added agent, and 3. have less sharp of a peak (i.e., the mixture more 

resembles a premixed flame of air and agent, for which the curves would be flat, then diffusion flame 

of aerosol can fuel with air). For comparison, the pure C6F12O flames (right frame) maintain their 

high Tad with agent addition, until Xa ≈ 0.027, above which a steep drop occurs (due to insufficient 

hydrogen in the system to produce the stable product HF, as has been discussed [4]). For Xa > 0.075, 

temperature again drops because the system (air-agent-fuel) is overall fuel rich. These features are 

captured in Figure 19, which shows the peak Tad, and the value of  at which the peak occurs. As 

indicated, the value of  for peak Tad increases to about 1.0 as Xa approaches the inertion value (0.27 

for Blend A, and around 0.07 to 0.08 for C6F12O). 

The increase in  with agent addition has ramifications for the potential pressure rise. Figure 20 

shows the predicted pressure rise with addition of Blend A or C6F12O [4]. For increasing Xa, as 

increases (due to the increase of oxygen demand of the agent itself and displaced O2), the mass of 

reactants also increases (the chamber oxidizer volume contains both air and agent), increasing the 

potential pressure rise. Thus, at values of Xa just below that required for inertion, there can be 

pressure rise of about 3.5 bar for Blend A addition, or 7 bar for C6F12O addition. 

23
500

0

500

1000

12

14

500

0

500

1000

12

14

500

0

500

1000

12

14

500

0

500

1000

12

14

500

0

500

1000

12

14

500

0

500

1000

12

14

Φ=1.0 Φ=0.5
 

‐

1500

2000

2500

0

2

4

6

8

10

Tad (K) ΔPmax (bar) 

Tad 

Equil 

Exp 

Blend D

Blend A

‐

1500

2000

2500

0

2

4

6

8

10

Tad (K) ΔPmax (bar) 

Tad 

Equil 

Exp 

0 5 10 15 20 25 30 0 5 10 15 20 25 30
 

Agent volume fraction (%) Agent volume fraction (%) 

Φ=0.6 Φ=0.4 

ΔPmax (bar) Tad (K) ΔPmax (bar) Tad (K) 

0 5 10 15 20 25 30 0 5 10 15 20 25 30
 

Agent volume fraction (%) Agent volume fraction (%) 

Φ=0.55 Φ=0.3 

ΔPmax (bar) Tad (K) ΔPmax (bar) Tad (K) 

‐

1500

2000

2500

0

2

4

6

8

10

Tad 

Equil 

Exp 

‐

1500

2000

2500

0

2

4

6

8

10

Tad 

Equil 

Exp 

‐

1500

2000

2500

0

2

4

6

8

10

Tad 

Equil 

Exp 

‐

1500

2000

2500

0

2

4

6

8

10

Tad 

Equil 

Exp 

0 5 10 15 20 25 30 0 5 10 15 20 25 30
 

Agent volume fraction (%) Agent volume fraction (%) 

Figure 17 – Measured (bottom curve) and calculated equilibrium (middle curves) explosion 

pressure in the 2-L chamber for (Pt-wire igniter) for moist propane-air flames (various ) with 

24
                       

Tad,max / K

 

1500 

2000 

added agent; calculated equilibrium adiabatic combustion temperature is shown by the top curves. 

Blue lines are Blend A; red (dotted) lines Blend D. 

1000 

1500 

2000 

Tad / K 

Xa 

0

0.29

0.11

0.20

Blend A 

1000 

1500 

2000 

Tad / K Xa 

0

0.100

0.027

0.053

C6F12O

0.0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0 

Fraction of Chamber Volume Involved in Combustion,  Fraction of Chamber Volume Involved in Combustion,  

Figure 18 – Equilibrium adiabatic combustion temperature as a function of , the fraction of 

chamber volume involved in combustion. Each curve refers to a different value of the agent 

volume fraction Xa in the chamber (left frame: Blend A; right frame, C6F12O [4]). 

T Blend A

ad / K 1.0 Tad,max / K C6F12O

2000 

0.5 



Tad,max ) 



Tad,max ) 

0.5 

1500 

1000 0.0 1000 0.0 

0 10 20 30 0 1 2 3 4 5 6 

Inhibitor Volume Fraction in Oxidizer, Xa (%) Inhibitor Volume Fraction in Oxidizer, Xa (%) 

Figure 19 – Peak temperature (from Figure 18) and the value of  at which the peak temperature 

occurs (left frame: Blend A; right frame, C6F12O [4]). 

25

1.0 
 

 

6 6

 

     

5 

P at Tmax / bar Blend A 

4 

3 

0 

1 

2 

3 

4 

5 

0 5 10 

P at Tmax / bar 

RH = 0 %

100 %

FAA Data

50 %

C6F12O

2 

1 

0
 

0 10 20 30 
 

Inhibitor Volume Fraction in Oxidizer Xa (%) Inhibitor Volume Fraction in Oxidizer Xa (%) 

Figure 20 – Predicted pressure rise in the FAA-ACT as a function of the volume fraction of added 

inhibitor (left frame: Blend A; right frame, C6F12O [4]). 

5.2.1.1. Blend D 

For Blend D (or C3H2F3Br) addition to the FAA-ACT, Figure 21 shows in similar format Tad as a 

function of  and Xa, and Figure 22 shows the peak values of Tad, and the values of  at which they 

occur. As the figures show, Tad decreases steadily with Blend D addition, but is constant (around 

2150 K) with C3H2F3Br addition. Based on these equilibrium simulations, the maximum pressure rise 

in the FAA-ACT for Blend D and C3H2F3Br are shown in Figure 23. The predicted pressure rise at 

the inertion point is about 4 bar with Blend D and about 8 bar with C3H2F3Br. Thus, even without 

kinetic limitations, the pressure rise just below the inertion point with the blends is expected to be 

about one half that of the pure agents just below their inertion point. The key question is whether 

with blend addition to the FAA-ACT, the reaction rate is slowed faster than the pressure rise 

increases. To examine this we turn to stirred-reactor simulations to determine the overall chemical 

rate as a function of mixing state and additive loading. 

26
 

2000 

T Blend D ad / K 

2000 

Xa 

1500 0

0.10

1000 

0.29 0.20

T C3H2F3Br ad / K 

2000 

Xa 

0 1500 

0.03 

0.06 

1000 

0.0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0 

 

Figure 21 – Equilibrium adiabatic combustion temperature as a function of , the fraction of 

chamber volume involved in combustion. Each curve refers to a different value of the agent 

volume fraction Xa in the chamber (left frame: Blend D; right frame, C3H2F3Br [5]) T ad,max ) 1000 

1500 

2000 

T ad,max ) 

Taft,max T / K / K C3H2F3Br ad,max Blend D 1.0 

0 10 20 30 

0.5 0.5 

1500 

1000 0.0 

0 2 4 6 

Inhibitor Volume Fraction in Oxidizer, Xa (%) Inhibitor Volume Fraction in Oxidizer, Xa (%) 

Figure 22 – Peak temperature (from Figure 21) and the value of  at which the peak temperature 

occurs (left frame: Blend D; right frame, C3H2F3Br [5]). 

27

0.0 

1.0 
 

6 6

5 5 

4 4 

3 3 

2 2 

1 1 

0 0 

0 

Inhibitor Volume Fraction in Oxidizer Xa 

10 20 

Inhibitor Volume Fraction in Air, Xa (%) 

Figure 23 – Predicted pressure rise in the FAA-ACT as a function of the volume fraction Xa of 

added inhibitor (left frame: Blend D; right frame, C3H2F3Br [5]). 

5.2.2. Perfectly Stirred Reactor Simulations 

Note that the assumptions that went into the calculation of the pressure rise as shown in Figure 20 

and Figure 23 have been shown to reproduce the actual pressure rise in the FAA-ACT when there are 

no kinetic limitations. While the deviation of the experimental pressure rise in the 2-L experiments 

from the calculated equilibrium pressure rise illustrated the kinetic effects of the blends, it is of value 

to explore them in more detail for the FAA-ACT using perfectly stirred reactor simulations. In 

particular, the PSR simulations can: 1. determine the overall reaction rate of the mixtures for a wide 

range of  and Xa; 2. explore reaction rates with the actual FAA-ACT fuel (instead of the moist 

propane-air mixtures of the 2-L experiments), and 3. compare expected reaction rates for the blends 

with those of the pure agents, for which the extinction behavior in the FAA-ACT experiments is 

available for comparison. 

The stirred-reactor overall chemical rate psr as a function of  is shown in Figure 24 for Blend A (left 

frame) and C6F12O (right frame) and in Figure 25 for Blend B and C3H2F3Br. The different curves 

refer to different values of Xa, as indicated. The green dots indicate the value of  that provided the 

peak Tad, and hence are expected to be the mixing condition at which the mixture in the FAA-ACT 

will burn (if possible). The black squares are the critical values of psr for extinguishment for the 

high-strain flame near the fuel release, or in the low-strain regions at the end-gases of the chamber. 

That is, near the explosive release of the aerosol can simulator and high-voltage igniter, the reaction 

rate must only be lowered (from about psr = 9000 s-1 for the uninhibited mixture) to about psr = 

1700 s-1, whereas in the low-strain regions of the chamber, if flammable mixtures exist, their reaction 

rate much be lowered to about psr = 20 s-1. The mixtures of pure C6F12O or C3H2F3Br create 

overpressures in the FAA-ACT because the green dots for  > 0.7 (shown in the right frames of Figure 

24 and Figure 25) have higher reactivity than both the value necessary for inertion (psr = 20 s-1), and 

the value of the uninhibited flames (the blue curve marked Xa=0). That is, addition of this agent to 

the lean mixtures enhances flammability. Unfortunately, for Blend A, there is a region near =0.7 for 

which the green dots also meet these criteria (the reactivity of the system at its desired stoichiometry 

is higher than for the uninhibited reactants). That higher reactivity, coupled with the increase in  as 

P at Tmax / bar Blend D 

0 1 2 3 4 5 6 7 

P at Tmax / bar C3H2F3Br 

28
Xa increases, is likely to lead to overpressure at sub-inerting concentrations. In contrast, in Figure 25 

for Blend D, the reactivity at the conditions of the thermodynamically favorable reactant mixtures 

(i.e., the green dots) is always lower than that of the uninhibited system (curve for Xa=0). That is, 

adding the suppressant always lowers the reactivity for any initial fuel-air stoichiometry. 

10000 

psr / s-1 Blend A 

10000 

psr / s-1 C6F12O 

1000 1000 

100 100 

10 10 

1 1 

Xinh 

0 

0.20 

0.11 

0.29 

0.058 

0.0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0 

Xinh 

0 

0.100 

0.011 

0.037 

0.042 

0.047 

0.052 

  

Figure 24 – Calculated stirred-reactor overall reaction rate psr as a function of , the fraction of 

chamber volume involved in combustion. Each curve refers to a different value of the agent 

volume fraction Xa in the chamber (left frame: Blend A; right frame, C6F12O [4]). 

10000 

psr / s-1 Blend D 

10000 

psr / s-1 C3H2F3Br 

1000 1000 

100 100 

10 10 

1 

 

1 

 

0.0 0.2 0.4 0.6 0.8 1.0 

Xinh 

0 

0.10 

0.047 

0.20 

uninhibited

1.8 1.3 1.0 0.8 0.7 0.6 0.5 0.4 0.3

0.0 0.2 0.4 0.6 0.8 1.0 

Xinh 

0 

0.04 

0.02 

0.06 

uninhibited

1.8 1.3 1.0 0.8 0.7 0.6 0.5 0.4 0.3 

Figure 25 – Calculated stirred-reactor overall reaction rate psr as a function of , the fraction of 

chamber volume involved in combustion. Each curve refers to a different value of the agent 

volume fraction Xa in the chamber (left frame: Blend D; right frame, C3H2F3Br [5]). 

29
 

 

The effect of agent addition on overall reactivity can be shown more clearly by replotting the data of 

Figure 24 and Figure 25 in alternative format. Figure 26 shows psr as a function of Xa (top frames: 

left C6F12O; right C3H2F3Br; bottom frames: left Blend A, right Blend D). In these frames, each curve 

refers to a specific value of  (or equivalently, the equivalence ratio of the uninhibited mixtures 

uninhibited=0.29/, as listed on the labeled curves). The horizontal line at psr=20 s-1 indicates the value 

of psr at which no reaction is expected anywhere in the FAA-ACT chamber (i.e., this is the overall 

rate necessary for inertion of all reaction, even at very low strain). For C6F12O (top left frame) with 

little fuel (i.e., high eta), the curve for / = 0.90/0.32 indicates that at low agent loadings, adding 

agent increases the reactivity, while for Xa > 0.025, it drops, so that at Xa=0.06, the mixture is 

predicted to be inert, even for low-strain flames in the FAA-ACT. The reactivity is relatively constant 

(and close to the inerting value) for 0.06 ≤ Xa ≤ 0.10 (note that the FAA-ACT experiments were only 

conducted at Xa = 0.042 and 0.081, with no inertion and inertion, respectively, consistent with the 

prediction from Figure 26). For C3H2F3Br (top right frame in Figure 26), the curve for / = 1.06/0.27 

indicates that the reactivity of the mixture is also increased with agent addition at low loading, such 

that the mixture is predicted to burn for low-strain conditions of 0.028 ≤ Xa ≤ 0.06 (i.e., mixtures with 

Xa=0.06 will not be inerted, as was experienced in the FAA-ACT experiments). 

For Blend A (bottom left frame in Figure 26) the curve for fuel-lean condtions (=1.0/=0.28) 

indicates that increasing Xa makes the system more flammable, and brings it above the threshold for 

reactivity in the FAA-ACT (psr=20 s-1); indeed, this is generally true for uninhibited=>0.65: adding 

Blend A increases the reactivity of the system, and brings it above the overall reaction rate for 

inertion (psr=20 s-1). In contrast, adding Blend D at the most dilute condition simulated (with regard 

to fuel from the aerosol can, =0.87/=0.48), does not bring psr, above the overall reactivity 

necessary for burning of the low-strain flames (psr=20 s-1). Hence, Blend D is not expected to cause 

overpressure in the FAA-ACT. 

Using Figure 24 and Figure 25, the expected inertion values for Blend A and Blend D in the FAA-ACT 

can be estimated. For example, in the figures, the curves (i.e., the value of Xa) for which psr is lower 

than that for inertion (psr=<=20 s-1) at the expected mixing condition (i.e., the green dots) provides 

the inertion value. For Blend A and Blend D, these are Xa=0.27 and Xa=0.19. These values are very 

close to the estimates from the 2-L explosion vessel (0.27 and 0.18), which is remarkable since: 1. the 

overall PSR rate for inertion (psr=20 s-1) is a rough estimate based on previous calculations [3-5] and 

experimental results in the FAA-ACT [1, 63], and 2. the inerting volume fraction of Blend A and 

Blend D determined from the 2-L chamber are not resolved very well in the experiments, but rather, 

read from extrapolations of the curves in Figure 13 and Figure 14. 

It is also possible that Blend D will inert the FAA-ACT at a lower volume fraction than Xa=0.19. In 

Figure 25, the volume fraction of Blend D estimated to extinguish the turbulent flame near the 

location of the fuel release and igniter (the upper black square in the figure) is estimated to be 

Xa=0.11. Physical locations away from this region have two differences: more diluted reactants (i.e., 

less fuel from the aerosol can), and lower strain (lower velocity gases). Further from the region of 

peak reactivity near ACT fuel release, the ACT fuel is diluted by the air/agent mixture, lowering the 

reactivity of the mixture; however, further from that region of explosive fuel release, the strain rate is 

lower, so that more agent is required to extinguish flames at that flow condition. See, for example, 

Figure 27 [64] which show the calculated agent volume fraction for extinction of a counterflow 

diffusion flame of methane-air, as a function of the strain rate. As indicated, the low-strain flames 

30
 

 

require more agent for extinction. It is not possible to know the complex way in which these two 

effects act, but if reaction rate is lowered (due to slower kinetics) faster than the agent requirement 

increases (because the lower strain flames require more agent), then the flames will extinguish at the 

lower value of Xa=0.11. 

1 

10 

100 

1,000 

10,000 

0 2 4 6 

Xa, Inhibitor Volume Fraction in Air (%) 

 

psr / s-1 C3H2F3Br 

1.06/0.27 

0.54/0.54 

0.17/1.70 

0.28/1.03 

psr / s-1 Blend A psr / s-1 Blend D 

10,000.0 

1,000.0 

100.0 

10.0 

1.0 

0.1 

0 10 20 30 0 10 20 

Xa, Inhibitor Volume Fraction in Air Xa, Inhibitor Volume Fraction in Air (%) 

 

1.0/0.28

0.74/0.39

0.86/0.33

0.65/0.45

0.1 

1.0 

10.0 

100.0 

1,000.0 

10,000.0 

 

0.87/0.33

0.75/0.39

0.60/0.48

Figure 26 – Stirred reactor overall reaction rate as a function of inhibitor volume fraction in the 

FAA-ACT. Each curve refers to a different value of the fraction of chamber volume allowed to 

react with the fixed quantity of aerosol can test fuel (i.e., a different initial equivalence ratio of the 

uninhibited mixture) (top frames: left, C6F12O, right, C3H2F3Br; bottom frames: left, Blend A, right 

Blend D). 

31
Figure 27 – Agent volume fraction at extinction vs. strain rate for a methane-air flame [64]. 

The potential for combustion enhancement by Blend A and Blend D has been explored above via 

experimental measurements of the explosion pressure in the 2-L combustion chamber, and via 

simulations of the equilibrium flame temperature and pressure and the stirred reactor overall 

reaction rate. In the section below, premixed burning velocity calculations and measurements are 

used to obtain additional insight, and further validate the kinetic mechanisms used for the 

simulations. 

5.3. Premixed Flame Burning Velocities 

5.3.1. Inhibited Hydrocarbon-air Flames 

Figure 28 shows the calculated 1-D, laminar, unstretched burning velocity of premixed C3H8-air 

flames (initial conditions: 1 bar, 298 K, 100 % r.h.) with added Blend A (left fame) or Blend D (right 

frame). Results are shown for equivalence ratios of the mixture (before addition of agent) of 1.0, 0.6, 

0.55, and 0.5. As described above, the accurate performance of the kinetic mechanisms for C6F12O 

and C3H2F3Br flame inhibition has been demonstrated in previous work [11, 62]. As additional 

illustration of the accuracy in the present work, their effect on the burning velocity of premixed, 

moist, C3H8-air flames was determined using the 2-L chamber (as described above). Figure 29 shows 

the calculated (lines) and measured (points) burning velocity for addition of Blend A (left frame) and 

Blend D (right frame). As indicated, for the stoichiometric flames, the agreement is excellent, while 

for the lean (=0.6) flames, the experiments provide a slower burning velocity (by about 3 cm/s), for 

both neat mixtures and with a Blend A loading of Xa=0.10. As has been discussed [11, 62], this is 

likely due to radiative heat losses, which are a large fraction of the total heat release for slow burning 

flames. Nonetheless, the relative reduction in burning velocity caused by addition of Blend A is 

accurately predicted, as shown in Figure 30, which shows the data of Figure 29 as the normalized 

32
 

burning velocity (in which the burning velocity is normalized by the burning velocity of the 

uninhibited mixture at the same stoichiometry). As indicted, the effect of adding either agent is well 

captured in the simulations performed. Note that for Blend D in the lean flame (=0.6), the available 

spark ignition energy (estimated to be 500 MJ) was not powerful enough to initiate reaction (as 

observed when pure C3H2F3Br was added to lean propane-air flames [11]). 

As indicted in Figure 28, Figure 29, and Figure 30, Blend A is about half as effective on a molar basis 

as Blend D when added to stoichiometric flames; moreover, Blend A is predicted to increase the 

burning velocity when added to flames of <=0.6. This may lead to overpressure in the FAA-ACT. 

In contrast, addition of Blend D is predicted to lower the reaction rate at all initial mixture 

stoichiometries (i.e., it never increases the reaction rate of the mixtures). These results reinforce the 

results found above in the experimental tests for explosion pressure and flammability limits, as well 

as the simulations for equilibrium temperature and pressure, and the stirred-reactor overall rates. 

SL (cm/s) SL (cm/s) 

0

0 0.1 0.2 0.3 0 0.1 0.2 0.3

Blend A Volume Fraction Blend D Volume Fraction 

Figure 28 – Calculated unstretched laminar burning velocity of premixed, moist C3H8-air flames 

(of initial equivalence ratio  of 1.0, 0.6, 0.55, and 0.5) with added suppressant (left frame: Blend A; 

right frame, Blend D). 

0

5

10

15

20

25

30

35

Φ=0.6 

Φ=0.55 

Φ=0.5 

Φ=1.0 

5

10

15

20

25

30

35

Φ=1.0 

0.55 

0.5 

0.6 

33
 

 

SL (cm/s) SL (cm/s) 

0

0 0.1 0.2 0.3 0 0.1 0.2 0.3

Blend A Volume Fraction Blend D Volume Fraction 

Figure 29 – Calculated (lines) unstretched and measured (points: □, =1.0; Δ, =0.6) laminar 

burning velocity of premixed, moist C3H8-air flames (of initial equivalence ratio  of 1.0 and 0.6) 

with added suppressant (left frame: Blend A; right frame, Blend D). 

SL/SLo (cm/s) SL (cm/s) 

0.0

0 0.1 0.2 0.3 0 0.1 0.2 0.3

Blend A Volume Fraction Blend D Volume Fraction 

5

10

15

20

25

30

35

0.6 

Φ=1.0 

0

5

10

15

20

25

30

35

Φ=1.0 

0.0

0.2

0.4

0.6

0.8

1.0

Φ=0.6 

1.0 

0.2

0.4

0.6

0.8

1.0

Φ=1.0 

Figure 30 – Calculated (lines) unstretched and measured (points: □, =1.0; Δ, =0.6) laminar 

burning velocity, normalized by the uninhibited value, for premixed, moist C3H8-air flames (of 

initial equivalence ratio  of 1.0 and 0.6) with added suppressant (left frame: Blend A; right frame, 

Blend D). 

34
5.3.2. Pure Agent-Air Premixed Flames 

The pure compounds C6F12O and C3H2F3Br and their blends (Blend A and Blend D) were found to be 

non-flammable in the 2-L tests described above. The present simulations predict a peak burning 

velocity of 0.14 cm/s for pure C6F12O-air at 308 K and 100 % r.h. and 1.4 cm/s for pure and C3H2F3Brair at 300 K and 100 % r.h. (note that the high values of r.h. are selected because these are the most 

flammable conditions for pure agent-air mixtures). These low calculated burning velocities for pure 

C6F12O- and C3H2F3Br-air flames are consistent with the non-ignition in the 2-L chamber tests with 

the platinum igniter. It is not necessary to calculate the burning velocities of Blend A-air and Blend 

D-air mixtures at near-room temperature conditions: if the pure agents had very low values of SL

o

 at 

these conditions, the agents diluted with CO2 will have even lower burning velocities. Again, the 

predicted burning velocities are consistent with the lack of ignition of the mixtures in the 2-L 

chamber at 295 K. Nonetheless, the flammability of the pure agent-air mixtures must be considered at 

higher temperature. 

In the FAA-ACT, the end-gases are compressed by combustion of the earlier-burning reactants, thus 

increasing the temperature of the end-gases. For example, Figure 31 shows the calculated 

temperature from an adiabatic compression of the gases in the FAA-ACT (air or air and fuel) to the 

indicated final pressure [64]. The arrows illustrate the final pressure which would be experienced in 

the chamber at a value of Xa just below the measured inertion point for each agent. As shown, for 

addition of HFC-125, Novec, or 2-BTP, the compressed end-gases can have initial temperatures of 

around 455 K, 480 K, or 515 K, respectively. 

Figure 31 – Temperature of gases in the FAA-ACT after adiabatic compression to the indicated 

final pressure, for air or FAA-ACT fuel with air (as noted). 

In previous work, Babushok et al. [60] calculated the burning velocity for pure C6F12O-air and 

C3H2F3Br-air flames at elevated temperatures and with added water vapor, as shown in Figure 32. 

35
   

     

   

      

All of these mixtures are well above 1 cm/s, and would be expected to burn in the moist, 

compressively heated end-gasses of the FAA-ACT. The C6F12O-air burning velocity has a strong 

temperature dependence. In previous work, Babushok et al. [60] calculated the burning velocity for 

pure dry C6F12O-air flames, and the peak calculated burning velocity was found to be 0.367 cm/s at 

400 K and 1.6 cm/s at 500 K with dry air [60]. With water vapor, the burning velocity is higher, for 

example (from Figure 32) at 500 K, increasing to about 2 cm/s at 7 % water vapor. For both of these 

pure agents with air, the predicted burning velocity behavior is consistent with the overpressures 

obtained in the FAA-ACT at high agent loadings, for which both C6F12O and C3H2F3Br in the 

endgasses clearly burned. Burning Velocity

/ cm

s‐1

6

5

4

3

2

1

0

BTP‐air, Tinit= 500 K

BTP‐air, 400 K

Novec 1230‐air, Tinit= 500 K

0 0.02 0.04 0.06 0.08 0.1 0.12
 

H2O Volume Fraction

Figure 32 – Calculated burning velocity of pure C3H2F3Br-air and C6F12O-air flames as a function of 

initial temperature and water vapor content. 

For Blend A and Blend D, however, the maximum overpressure (if there are no kinetic limitations) 

will only be 3.5 bar to 4 bar, leading to initial temperatures of about 430 K or 460 K, respectively. It is 

useful to know the burning velocities of pure flames of Blend A or Blend D with air at the elevated 

initial temperatures corresponding to their potential overpressure. This will provide an assessment 

of the combustion propensity of the mixtures at the potential end-gas condition. Although these 

burning velocities could be measured (if they are high enough), our devices are not equipped for 

measurements at elevated temperature. Since the numerical predictions of burning velocity 

accurately reproduce measured values, the simulations are used to estimate the burning velocities at 

the heated conditions. 

For the blends of these agents with CO2, the reactivities (i.e., burning velocities) in the end gases will 

be much lower than for the pure agents. The CO2 in the agent has two effects: lowering the 

temperature and reactivity at any initial temperature, and providing dilution, which reduces the 

compressive heating in the FAA-ACT—again lowering the reactivity. 

36
   6.05E 06x2 3.83E 03x + 6.19E

   

Figure 33 shows the calculated laminar premixed 1-D planar unstretched burning velocity SL

o

 of 

Blend A and Blend D as a function of initial reactant temperature (at 100 % r.h.). (In the figure the 

data points are the result of the calculations at the indicated initial temperature, while the lines are 

smooth curve fits to the data.) For Blend A at 430 K, the calculated value of SL

o

 is interpolated to be 

approximately 0.09 cm/s while for Blend D at 460 K, it is 0.82 cm/s. Both of these values are below 

the 1 cm/s value which appeared to be the lower limit of flame propagation in the endgasses of the 

FAA-ACT [65] 

y = ‐ ‐ ‐ ‐01

300 350 400 450 500

SL (cm/s)

Blend A

Blend D 1.0

0.8

0.6

0.4

0.2

0.0

Initial Temperature / K

Figure 33 - Calculated laminar premixed 1-D planar unstretched burning velocity SL

o

 of Blend D as 

a function of initial gas temperature. 

37
Conclusions 

In order to understand the potential for overpressure in the FAA-ACT, experimental measurements 

and numerical simulations have been made. The following main points have been elucidated, and 

the tests or calculations leading to them are explained. 

The endgases of pure Blend A or Blend D in the FAA-ACT are not likely to combust. 

Explosion pressure tests indicate that pure C6F12O or C3H2F3Br, and Blend A and Blend D in moist air 

are not flammable (at their most reactive stoichiometry), at 295 K. Numerically calculated burning 

velocities indicate that the pure agents are flammable at temperatures expected in the compressed 

endgases of the FAA-ACT, and this behavior is consistent with their large overpressures in the FAAACT. For the Blends A and B, however, predicted burning velocities indicate that the blends are not 

flammable at the possible compressed conditions of the FAA-ACT for these agent. Hence, while the 

combustion of the endgases containing nearly pure agent and air will occur for the FAA-ACT, this 

will not happen with the blends. This characteristic will help reduce the overpressure tendency of 

the agents. 

The inerting concentrations of Blend A and Blend D have been determined. 

Blend A is predicted to inert the FAA-ACT at a volume fraction Xa of 0.27, and Blend D at Xa=0.18. 

This is based on the inerting concentrations in the 2-L chamber tests with the conservative Platinum 

igniter with moist propane-air mixtures at a range of initial fuel-air stoichiometries. Also, stirred 

reactor simulations independently predict inerting values in the FAA-ACT of 0.27 and 0.19 (i.e., 

nearly the same values). There is some chance that Blend D will inert the FAA-ACT at Xa=0.11, but it 

is difficult to say for certain. 

Blend A is much more likely to cause an overpressure in the FAA-ACT than Blend D (which is 

predicted to be effective in the FAA-ACT). 

There are many reasons for this result. As they are added to the FAA-ACT, both agents increase 

(the fraction of chamber volume involved in combustion), so the mass of combustible fuel increases 

as Xa goes up. This exacerbates the overpressure potential of both agents, although the magnitude of 

the potential pressure rise is only half of the neat agents C6F12O and C3H2F3Br. Moreover, for Blend 

A, the reactivity of the lean mixtures is enhanced with agent addition. This was shown in the 

explosion pressure tests, the overall reaction rates calculated by the PSR simulations, and in the 

premixed flame speed calculations. Hence, with Blend A, the potential for overpressure from the 

increase in , together with the enhanced reactivity of many of its mixtures with fuel are likely to 

result in overpressure (although much lower than that with added C6F12O). In contrast, Blend D 

always reduces the reactivity of the fuel-air mixture (at any value of ), so it has a much better chance 

to not cause the overpressure. For this to happen, Blend D must reduce the combustion efficiency in 

the chamber, on average, more effectively than its addition causes  to go up (and the pressure rise to 

go up). Based on the tests results and calculations described above, it is believed that Blend D will 

not cause an overpressure in the FAA-ACT. 

38
References: 

1.	 J. W. Reinhardt, Aircraft cargo MPS test of FK-5-1-12, in: International Aircraft Systems Fire 

Protection Working Group, October 25-26, 2006, Atlantic City, NJ, 2006, Federal Aviation 

Administration, Atlantic City, NJ, 

2.	 G. T. Linteris, F. Takahashi, V. R. Katta, H. K. Chelliah, O. Meier, Thermodynamic analysis of 

suppressant-enhanced overpressure in the FAA aerosol can simulator, in: Fire Safety Science -­

Proceedings of the Tenth International Symposium, M. Spearpoint, (Ed.) College Park, MD, 

6/19/2011, International Association for Fire Safety Science, Boston, MA, 2011, 307-320. 

3.	 G. T. Linteris, D. R. Burgess, V. R. Katta, F. Takahashi, H. K. Chelliah, O. Meier, Combustion 

and Flame 159 (2012) 1016-1025. 

4.	 G. T. Linteris, V. I. Babushok, P. B. Sunderland, F. Takahashi, V. R. Katta, O. Meier, 

Proceedings of the Combustion Institute 34 (2013) 2683-2690. 

5.	 G. T. Linteris, V. Babushok, J. L. Pagliaro, D. R. Burgess Jr, J. A. Manion, F. Takahashi, V. R. 

Katta, P. T. Baker, Combustion and Flame, Understanding Overpressure in the FAA Aerosol 

Can Test by C3H2F3Br (2-BTP), accepted (2015). 

6.	 D. R. Burgess Jr, V. I. Babushok, G. T. Linteris, J. A. Manion, International Journal of Chemical 

Kinetics A chemical kinetic mechanism for 2-bromo-3,3,3,-trifluoropropene (2-BTP) flame 

inhibition, accepted (2015). 

7.	 V. Babushok, G. T. Linteris, D. R. Burgess Jr, P. T. Baker, Combustion and Flame 162 (2015) 

1104-1112. 

8.	 V. I. Babushok, G. T. Linteris, P. T. Baker, Combustion and Flame 162 (5) (2015) 2307-2310. 

9.	 F. Takahashi, V. R. Katta, G. T. Linteris, V. I. Babushok, Proceedings of the Combustion 

Institute 35 (2015) 2741-2748. 

10.	 J. L. Pagliaro, G. T. Linteris, P. B. Sunderland, P. T. Baker, Combustion and Flame 162 (1) (2015) 

41-49. 

11.	 J. L. Pagliaro, N. Bouvet, G. T. Linteris, Combustion and Flame, Premixed flame inhibition by 

CF3Br and C3H2F3Br (2-BTP), in preparation (2015). 

12.	 M. Metghalchi, J. C. Keck, Combustion and Flame 38 (1980) 143-154. 

13.	 K. Takizawa, A. Takahashi, K. Tokuhashi, S. Kondo, A. Sekiya, Combustion and Flame 141 (3) 

(2005) 298-307. 

14.	 R. Stone, A. Clarke, P. Beckwith, Combust. Flame 114 (1998) 546-555. 

15.	 T. Iijima, T. Takeno, Combustion and Flame 65 (1986) 35-43. 

39
16.	 G. T. Linteris, J. L. Pagliaro, Test results prepared for Honeywell: Flammability of refrigerants 

in the Japanese high pressure gas law test, NISTIR 7178, National Institute of Standards and 

Technology, Gaithersburg MD, 2011, p 48 p. 

17.	 J. E. Shephard, J. C. Krok, J. J. Lee, Spark Ignition Energy Measurements in Jet A, Explos. Dyn. 

Lab. Rep. FM 99e7, California Institute of Technology, Pasedena, CA, 1999. 

18.	 S. P. Marshall, S. Taylor, C. R. Stone, T. J. Davies, R. F. Cracknell, Combust. Flame 158 (10) 

(2011) 1920-1932. 

19.	 M. Metghalchi, J. C. Keck, Combustion and Flame 48 (1982) 191-210. 

20.	 P. G. Hill, J. Hung, Combust. Sci. Technol. 60 (1988) 7-30. 

21.	 B. Lewis, G. von Elbe, Combustion, Flames and Explosions of Gases, Academic, New York, 

1961. 

22.	 K. Saeed, C. R. Stone, Combust. Flame 139 (1-2) (2004) 152-166. 

23.	 K. Eisazadeh-Far, A. Moghaddas, J. Al-Mulki, H. Metghalchi, Proceedings of the Combustion 

Institute 33 (1) (2011) 1021-1027. 

24.	 G. P. Smith, D. M. Golden, M. Frenklach, N. W. Moriarty, B. Eiteneer, M. Goldenberg, C. T. 

Bowman, R. K. Hanson, S. Song, J. W. C. Gardiner, V. V. Lissianski, Z. Qin, GRI Mech 3.0. 

<http://www.me.berkeley.edu/gri_mech> (Sept. 9, 2015). 

25.	 D. R. Burgess, M. R. Zachariah, W. Tsang, P. R. Westmoreland, Progress in Energy and 

Combustion Science 21 (6) (1995) 453-529. 

26.	 D. Burgess, M. R. Zachariah, W. Tsang, P. R. Westmoreland, Thermochemical and Chemical 

Kinetic Data for Fluorinated Hydrocarbons, NIST Technical Note 1412, Gaithersburg, MD, 

1995. 

27.	 V. Babushok, T. Noto, D. R. F. Burgess, A. Hamins, W. Tsang, Combustion and Flame 107 (4) 

(1996) 351-367. 

28.	 V. I. Babushok, D. R. F. Burgess, W. Tsang, A. W. Miziolek, Simulation studies on the effects of 

flame retardants on combustion processes in a plug reactor, in: Halon Replacements, W. 

Tsang; A. W. Miziolek, (Eds.), ACS Symposium Series 611, American Chemical Society, 

Washington, D.C., 1995, pp. 275-288. 

29.	 S. Gordon, McBride, B.J., Computer program for calculation of complex chemical equilibrium 

compositions and applications, NASA Reference Publication 1311, Cleveland, OH, 1996. 

30.	 F. A. Williams, A review of some theoretical considerations of turbulent flame structure, in: AGARD 

Conference Proceeding, AGARD-CP-164, NATO Science and Technology Organization, 1975, 

31.	 M. Elia, M. Ulinski, M. Metghalchi, Journal of Engineering for Gas Turbines and Power 123 (1) 

(2001) 190. 

40
32. A. Moghaddas, K. Eisazadeh-Far, H. Metghalchi, Combustion and Flame 159 (4) (2012) 1437­

1443. 

33. D. Razus, V. Brinzea, M. Mitu, C. Movileanu, D. Oancea, Energy & Fuels 26 (2) (2012) 901-909. 

34. A. S. Huzayyin, H. A. Moneib, M. S. Shehatta, A. M. A. Attia, Fuel 87 (1) (2008) 39-57. 

35. D. R. Dowdy, D. B. Smith, S. C. Taylor, A. Williams, Proceedings of the Combustion Institute 

23 (1990) 325-332. 

36. L. K. Tseng, M. A. Ismail, G. M. Faeth, Combustion and Flame 95 (4) (1993) 410-426. 

37. U. J. Pfahl, M. C. Ross, J. E. Shepherd, K. O. Pasamehmetoglu, C. Unal, Combustion and Flame 

123 (1-2) (2000) 140-158. 

38. B. N. Taylor, C. E. Kuyatt, Guidelines for Evaluating and Expressing the Uncertainty of NIST 

Measurement Results, NIST Technical Note 1297, National Institute of Standards and 

Technology, Gaithersburg, MD, 1994. 

39. I. Glassman, Combustion, Academic Press, San Diego, CA, 1996. 

40. S. Kondo, Y. Urano, K. Takizawa, A. Takahashi, K. Tokuhashi, A. Sekiya, Fire Safety Journal 41 

(1) (2006) 46-56. 

41. M. I. Hassan, K. T. Aung, G. M. Faeth, Combustion and Flame 115 (4) (1998) 539-550. 

42. G. Rozenchan, D. L. Zhu, C. K. Law, S. D. Tse, Proc. Combust. Inst. 20 (2002) 1461-1469. 

43. C. H. Osorio, A. J. Vissotski, E. L. Petersen, M. S. Mannan, Combustion and Flame 160 (6) 

(2013) 1044-1059. 

44. O. Park, P. S. Veloo, N. Liu, F. N. Egolfopoulos, Proc. Combust. Inst. 33 (1) (2011) 887-894. 

45. C. M. Vagelopoulos, F. N. Egolfopoulos, Proceedings of the Combustion Institute 27 (1998) 

513-519. 

46. C. C. M. Luijten, E. Doosje, L. P. H. de Goey, International Journal of Thermal Sciences 48 (6) 

(2009) 1213-1222. 

47. R. J. Kee, J. F. Grcar, M. D. Smooke, J. A. Miller, A fortran computer program for modeling 

steady laminar one-dimensional premixed flames, SAND85-8240, Sandia National 

Laboratories, Livermore, CA, USA, 1991. 

48. P. Glarborg, R. J. Kee, J. F. Grcar, J. A. Miller, PSR: A FORTRAN program for modeling wellstirred reactors, SAND86-8209, Sandia National Laboratories, Albuquerque, NM, 1986, pp. 1­

54. 

49. R. J. Kee, F. M. Rupley, J. A. Miller, CHEMKIN-II: A fortran chemical kinetics package for the 

analysis of gas phase chemical kinetics, SAND89-8009B, Sandia National Laboratories, 

Livermore, CA, 1989. 

41
50.	 R. J. Kee, G. Dixon-Lewis, J. Warnatz, R. E. Coltrin, J. A. Miller, A fortran computer package 

for the evaluation of gas-phase, multicomponent transport properties, SAND86-8246, Sandia 

National Laboratories, Livermore, CA, 1986. 

51.	 S. R. Turns, An introduction to combustion : concepts and applications, WCB/McGraw-Hill, 

Boston, 2000. 

52.	 F. A. Williams, Journal of Fire and Flammability 5 (1974) 54-63. 

53.	 T. Marker, Initial development of an exploding aerosol can simulator, DOT/FAA/ARTN97/103, Federal Aviation Administration, Washington, D.C., 1998, pp. 1-16. 

54.	 R. B. Barat, Chemical Engineering Science 56 (2001) 2761-2766. 

55.	 S. Liu, M. C. Soteriou, M. B. Colket, J. A. Senecal, Fire Safety Journal 43 (8) (2008) 589-597. 

56.	 H. Wang, X. You, K. W. Jucks, S. G. Davis, A. Laskin, F. Egolfopoulos, C. K. Law, USC Mech 

Version II. High-temperature combustion reaction model of H2/CO/C1-C4 compounds. 

<http://ignis.usc.edu/USC_Mech_II.htm> (2015). 

57.	 J. Li, A. Kazakov, F. L. Dryer, Journal of Physical Chemistry A 108 (38) (2004) 7671-7680. 

58.	 D. L'Esperance, B. A. Williams, J. W. Fleming, Combustion and Flame 117 (4) (1999) 709-731. 

59.	 Y. Saso, D. L. Zhu, H. Wang, C. K. Law, N. Saito, Combustion and Flame 114 (3-4) (1998) 457­

468. 

60.	 V. I. Babushok, G. T. Linteris, O. Meier, Combust Flame 159 (2012) 3569-3575. 

61.	 E. Goos, A. Burcat, B. Ruscic, Extended third millennium thermodynamic database for 

combustion and air-pollution use with updates from active thermochemical tables. 

<Ftp://ftp.technion.ac.il/pub/supported/aetdd/thermodynamics/BURCAT.THR > (August 

2012). 

62.	 J. L. Pagliaro, G. T. linteris, Fire Safety Journal, "Hydrocarbon flame inhibition by C6F12O 

(Novec 1230): Kinetic model validation and influence of water vapor," (2015) In Preparation. 

63.	 J. W. Reinhardt, Behavior of bromotrifluoropropene and pentafluoroethane when subjected to 

a simulated aerosol can explosion, DOT/FAA/AR-TN04/4, Federal Aviation Administration, 

Washington, D.C., 2004, pp. 1-20. 

64.	 G. T. Linteris, J. A. Manion, W. Tsang, D. R. Burgess Jr, V. Babushok, V. R. Katta, F. Takahashi, 

P. B. Sunderland, J. L. Pagliaro, Overpressure in the FAA aerosol can test with halon replacements, 

in: The Seventh Triennial International Fire & Cabin Safety Research Conference, Dec. 3-5, 

2013, Philadelphia, PA, 2013, Federal Aviation Administration, Atlantic City, NJ, 2013, 

65.	 G. T. Linteris, J. A. Manion, D. R. Burgess, I. A. Awan, P. B. Sunderland, J. L. Pagliaro, 

Unwanted combustion enhancement by potential halon replacements: 2011 annual report, 

NIST Technical Note xxx, N. I. o. S. a. Technology, Gaithersburg, MD, 2011. 

42