Stirred reactor calculations to understand unwanted combustion enhancement by potential halon replacements

Combustion and Flame 159 (2012) 1016–1025

Contents lists available at SciVerse ScienceDirect

Combustion and Flame j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / c o m b u s t fl a m e

Stirred reactor calculations to understand unwanted combustion enhancement by potential halon replacements q Gregory T. Linteris a,⇑, Donald R. Burgess b, Fumiaki Takahashi c, Viswanath R. Katta d, Harsha K. Chelliah e, Oliver Meier f a

Fire Research Division, National Institute of Standards and Technology, Gaithersburg, MD 20899-8665, USA Chemical and Biochemical Reference Data Division, National Institute of Standards and Technology, Gaithersburg, MD 20899-8320, USA Case Western Reserve University, Cleveland, OH 44106, USA d Innovative Scientific Solutions, Inc., Dayton, OH 45440, USA e University of Virginia, Charlottesville, VA 22904, USA f The Boeing Company, Seattle, WA 98124, USA b c

a r t i c l e

i n f o

Article history: Received 4 March 2011 Received in revised form 20 September 2011 Accepted 22 September 2011 Available online 7 November 2011 Keywords: Fire suppression Flame inhibition CF3Br C2HF5 Halon replacements Cargo-bay fire suppression

a b s t r a c t Several agents are under consideration to replace CF3Br for use in suppressing fires in aircraft cargo bays. In a Federal Aviation Administration performance test simulating the explosion of an aerosol can, however, the replacements, when added at sub-inerting concentrations, have all been found to create higher pressure rise than with no agent, hence failing the test. Thermodynamic equilibrium calculations as well as perfectly-stirred reactor simulations with detailed reaction kinetics, are performed to understand the reasons for the unexpected enhanced combustion rather than suppression. The high pressure rise with added C2HF5 or C3H2F3Br is shown to be dependent upon the amount of added agent, and can only occur if a large fraction of the available oxidizer in the chamber is consumed, corresponding to stoichiometric proportions of fuel, oxygen, and agent. Conversely, due to the unique stoichiometry of CF3Br, this agent is predicted to cause no increase in pressure, even in the absence of chemical inhibition. The stirred-reactor simulations predict that the inhibition effectiveness of CF3Br is highly dependent upon the mixing conditions of the reactants (which affects the local stoichiometry and hence the overall reaction rate). For C2HF5, however, the overall reaction rate was only weakly dependent upon stoichiometry, so the fuel– oxidizer mixing state has less effect on the suppression effectiveness. Published by Elsevier Inc. on behalf of The Combustion Institute.

1. Introduction Because of its destruction of stratospheric ozone, production of the effective fire suppressant CF3Br has been banned by the Montreal Protocol. Although a critical-use exemption has been granted to the aviation industry for use of recycled halon in cargo bay fire suppression, the European Union requires replacement of halon in new design aircraft by 2018, and in existing aircraft by 2040. Several replacements have been proposed, but they have all been found to produce enhanced burning in the FAA Simulated Aerosol Can test [1], and hence they fail FAA’s Minimum Performance Standard [2]. In particular, C2HF5, C3H2F3Br, and C6F12O all produce higher peak pressures in a simulated cargo bay as compared to no added agent, when they are added at concentrations less than that required to completely suppress the explosion. (Names and

q

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

⇑ Corresponding author. Fax: +1 301 975 4052.

E-mail address: [email protected] (G.T. Linteris).

chemical formulas are listed in Table 1.) The agent CF3Br, at subsuppressing concentrations, does not cause the overpressure. The Aerosol Can Test [1] simulates the situation in which a fire in an aircraft cargo bay container heats an aerosol can (e.g., hair spray) causing it to burst, creating an explosion. In the test, a heated container at about 16 bar, releases its contents (propane, ethanol, and water) as a two-phase impulsive spray via a fastacting valve. A continuous DC arc across electrodes located about 1 m downstream of the valve ignites the mixture. The fireball expands into the chamber atmosphere of ambient air and water vapor and premixed suppressant, and the temperature and pressure in the chamber increases (over a time on the order of a second). During each test, instruments record the pressure, temperature, visual images, and concentrations of agent and oxygen. Unfortunately, when added at sub-inerting concentrations, the final pressure rise in the chamber with any of the halon replacement agents is higher than in the absence of agent. The agent C2HF5, added at a volume fraction of 13.5%, suppressed the explosion; however, when added at volume fractions of 6.2%, 8.9%, and 11.0%, the peak pressure rise was about 3.6 bar, or about twice that

0010-2180/$ - see front matter Published by Elsevier Inc. on behalf of The Combustion Institute. doi:10.1016/j.combustflame.2011.09.011

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

1017

Table 1 Chemical names of compounds of interest. Formula

Trade name

Common name

Systematic name

Structural formula

CF3Br C2HF5 C3H2F3Br C6F12O

Halon 1301 HFC-125 2-BTP Novec 1230, FK-5-112

Bromotrifluoromethane Pentafluoroethane Bromotrifluoropropene Heptafluoroisopropyl pentafluoroethyl ketone

Bromotrifluoromethane 1,1,1,2,2-Pentafluoroethane 2-Bromo-3,3,3-trifluoro-1-propene 1,1,1,2,2,4,5,5,5-nonafluoro-4-(trifluoromethyl)-3pentanone

BrCF3 CF3  CHF2 CH2 = CBrCF3 CF3CF2C(=O)CF(CF3)2

with no added suppressant (1.8 bar). With C3H2F3Br, volume fractions of 3% or 4% both gave a pressure rise of 4.3 bar, and volume fractions of 5% or 6% gave a pressure rise of about 6.7 bar. With C6H12O, a volume fraction of 4.2% gave a pressure rise of 4.5 bar, and above that volume fraction (or sometimes at 4.2% itself) the over pressure was suppressed. In contrast, addition of CF3Br at a volume fraction of about 3.8% suppressed the explosion, and 2.5% CF3Br, the pressure rise was only 0.28 bar. (For reference, unconfined tests without suppressants create a 3.4 m diameter fire ball [3]). The goal of the present work is to understand the reasons for the higher pressure rise in the FAA aerosol can test with C2HF5 or C3H2F3Br (at sub-inerting concentrations), and the lack of this effect with CF3Br. 2. Background Enhanced combustion in the presence of fire suppressants has been observed in previous work. A promotion effect has been shown for halogenated hydrocarbons in shock-tube ignition studies for some fuels and conditions. For example, it is known that the H2–O2 mixture can be sensitized by HBr [4–6] and HCl [4], with promotion under some conditions and inhibition under others. Moen et al. [7] performed experiments and modeling on the influence of CF3Br, CF4, and CO2 on the detonability of ethane–air and ethane– and hydrogen–oxygen mixtures, and found that the inhibition ability of CF3Br was greatly reduced as compared to flame inhibition, and that CF3Br was a sensitizer for some situations. Similarly, Babushok et al. [8] computationally studied the ignition delay of hydrocarbons (H2, CH4, CH3OH, and C2H6) and air with various halogenated fire suppressants (Br2, HBr, CH3Br, CF3Br, and CF3H), and also found both promotion or inhibition of the ignition. The varying behavior was complex, likely depending upon subtle changes to the kinetic pathways as the pressure, temperature, or mixture composition change. Wider flammability limits in the presence of halogenated compounds have also been reported. Saito et al. [9] conducted flammability limit experiments in a tubular flame of premixed hydrocarbon (methane or propane) and air, with added CF3Br, CF3H, C3HF7, or C4F10. The traditional fuel–air concentration limits [10] were determined as a function of agent concentration. With addition of CF3Br, both the rich and lean limits narrowed steeply (the mixture became less flammable); whereas with addition of various hydrofluorocarbons (HFCs), the rich limit narrowed less steeply, and the lean limit sometimes widened (the mixture became more flammable). These results confirmed the ignition tests of Moore et al. [11] performed in a spherical combustion chamber, also with premixed reactants. Recently, Kondo et al. [12] measured the flammability limits of C2HF5 for spark-ignited premixed gases in a glass flask [13]; fuels were methane, propane, propylene, methyl formate, and HFC-152a (1,1-difluoroethane, CF2H–CH3), with dry air. The addition of the agent C2HF5 narrowed the rich limit, roughly linearly, up to the inertion point. Conversely, the lean limit was widened for all fuels, with the order of ranking: C3H6 < C3H8 < HCOOCH3 < C2H4F2  CH4 (i.e., largest widening of the lean limit occurred for the fuels with higher H/C ratio). Hence, these experiments also showed an enhance burning of the fuels caused by C2HF5 under fuel-lean conditions.

Using a constant volume combustion device, Shebeko et al. [14] studied the effect of various halogenated hydrocarbons on the combustion of methane– and hydrogen–air mixtures. The flammability limits generally narrowed with addition of the agents; however, for C4F8 addition to CH4–air flames, the lean limit widened, and for C2HF5 or C2F5Cl addition to CH4–air flames, the lean limit was nearly unchanged. That is, for a CH4–air flame at the lean limit, a C2HF5 mass fraction of 30% had essentially no effect on the flammability. In addition to the flammability limits, the pressure rise, as well as the rate of pressure rise were also used by Shebeko et al. [14] to illustrate the promoting effect of the halogenated fire suppressant for some conditions. For example, when added to lean H2–air flames, C2HF5 increased the maximum pressure rise at any concentration (up to the extinction point). Further, C4F8 (perfluorocyclobutane) addition increased the maximum pressure rise for both lean and rich H2–air flames. This lack of kinetic inhibition was even further substantiated for the agent C4F10 when added to CH4–air flames, for which even the rate of pressure rise was higher with agent added up to a volume fraction of 4%. In experiments with high-speed turbulent flames in a detonation/deflagration tube, Grosshandler and Gmurczyk [15–18] observed more vigorous combustion with CF3I or CF3Br, or various hydrofluorocarbon inhibitors, while using either propane or ethylene as fuels. For some conditions, the premixed addition of the halogenated agent to the air stream increased both the deflagration/ shock propagation rate as well as the pressure ratio across the shock. The results varied with fuel type, stoichiometry, agent type, and the presence or absence of turbulence-inducing spirals. In tests with co-flow diffusion flames, halogenated hydrocarbons added to either the fuel or air stream have been shown to increase total heat release. Holmstedt et al. [19] reported that HFC-227ea (C3HF7) or HFC-134a (C2H2F4) added to the fuel (propane) stream of a turbulent jet burner increased the total heat release, by a factor of 2 and 3.8, respectively, for concentrations just below that required for extinguishment of the flame. Similarly, Katta et al. [20] found that CF3H added to the oxidizer stream in a methane–air cup-burner experiment increased the total heat release. Increased pressure can also make HFC–air mixtures more flammable. In a survey paper, Ural [21] noted that while halogenated hydrocarbons are often considered to have no or low flammability potential, some have been shown to become flammable at elevated pressure. In a recent study, Kondo et al. [22] measured the rich and lean flammability limits of HFC-32 (CH2F2) and HFO-1234yf (C3H2F4) at pressures of 101–2000 kPa. They found that in general, at increased pressure, the lean flammability limit was relatively unchanged, while the rich limit widened significantly. In large-scale tests, Mawhinney et al. [23] found that application of water mist to a fire caused unwanted accelerated burning, which they believed was due to fluid–dynamic enhancement of the burning. Hamins et al. [24] reviewed previous work on enhanced burning with application of fires suppressants, and also concluded that the enhanced combustion was due to more rapid mixing of fuel vapor with air, from the combined effects of enhanced turbulent mixing and more vigorous liquid fuel atomization from agent jet impingement.

1018

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

While the possibility of enhanced flammability in the presence of HFC fire suppressants has been described in the work presented above, no analyses were performed in those investigations to understand why promotion occurred. More typically, HFC agents act as fire suppressants, and based on previous work [25–27], C2HF5 would be expected to extinguish—or at least weaken—the FAA aerosol can test explosions. In previous work [28], an analysis based on equilibrium thermodynamics was used to gain insight into the behavior of the agents C2HF5, C3H2F3Br, and CF3Br in the FAA tests. The analyses were able to predict the experimental pressure rise for addition of C2HF5 or C3H2F3Br at sub-extinguishing volume fractions (0–11.2% and 0–6%, respectively). The enhanced combustion with addition of agent to the oxidizer was shown to be due to the higher heat release, caused by involvement of increasing fractions of the chamber volume of oxidizer. Nonetheless, for CF3Br at all concentrations, and C2HF5 at its suppression volume fraction (13.5%), the pressure rise was less than predicted by equilibrium thermodynamics, possibly due to kinetic limitations in the rate of energy release caused by addition of these chemically acting agents [29–32]. Hence, it is of interest to examine the kinetic behavior of the aerosol can test fuel with air and added agent (C2HF5 or CF3Br). Because of the distributed nature of the explosively dispersed fuel-agent–air mixtures in the FAA tests [1], perfectly-stirred reactor (PSR) calculations were employed, as described below. While the conditions in the FAA test are likely to include more non-premixed or partially premixed structure (due to the fuel-rich core), premixed stirred reactor calculations nonetheless seem to be a reasonable next step in complexity (after the equilibrium calculations), that will provide some insight into the relevant kinetic limitations. 3. Numerical methods 3.1. Initial conditions The volume fraction of agent in the oxidizer Xinh, is a variable. The turbulent mixing during the aerosol release makes specification of the oxidizer–fuel ratio near the flame difficult, so calculations are performed for a range of fuel–oxidizer ratio. However, because the oxidizer consists of both air and agent (which is also a reactant), the oxidizer is partially premixed. Presentation of the results in terms of a fuel–oxidizer stoichiometric ratio is not practical: as the agent concentration in the oxidizer is increased, u for stoichiometric conditions changes due to the agent’s oxygen demand, as well as due to changes in the equilibrium products as the hydrogen–halogen ratio in the flame changes. Hence, the stoichiometry effect is presented in terms of the fraction (g) of the available oxidizer (the chamber volume) involved in the reaction with the fixed quantity of fuel species from the aerosol can contents. The fuel quantity is taken to be the contents of the aerosol can simulator (270 g ethanol, 90 g propane, and 90 g water). The oxidizer consists of ambient air (oxygen, nitrogen, and water vapor) and the added suppressant. Relative humidity of the test air was not reported in the FAA tests; however, analysis of local weather data for the days of the tests indicates that water vapor volume fractions were always in the range of 0.0016–0.01 (corresponding to 7–40% R.H. at 21 °C). Hence, the calculations were performed for a water vapor volume fraction Xwv of 0, 0.0125, and 0.025 in the O2/N2/H2O oxidizer mix, corresponding to 0%, 50%, and 100% R.H at 21 °C. To summarize: the fuel components and their quantities were fixed; the ratio of fuel to total oxidizer (N2, O2, H2O, and agent) was variable, expressed as g, the fraction of chamber volume (11,400 L) involved in the combustion, and the quantity of each oxidizer component in the initial mixture for the calculation was

determined based on the value of g and the specified volume fraction of each component. The initial inhibitor volume fraction in the oxidizer gases was varied from zero to 13.5% for C2HF5, 6%, for C3H2F3Br, and 5% for CF3Br, corresponding to approximately the maximum amount added in the FAA tests [1,33]. The fraction of the chamber volume involved in the combustion, g, was varied from about 0.23–1.00. 3.2. Equilibrium thermodynamics The equilibrium conditions of the aerosol can test were calculated using both the STANJAN-III program of Reynolds [34], and CEA2 of Gordon and McBride [35]; the two codes gave results very close to each other. The calculations were performed over the wide range of initial conditions described above. Constant enthalpy, constant pressure solutions were obtained, as described in Ref. [28]. 3.3. Kinetic mechanism A kinetic mechanism to describe the chemical inhibition of the aerosol can test fuel (propane, ethanol, water) with the HFC and bromine-containing species was assembled from sub-mechanisms available in the literature. For the hydrocarbon mechanism, an optimized model for ethylene oxidation proposed by Wang and co-workers was employed [36,37], that included 111 species and 784 elementary reactions. This model has been optimized by considering experimental ignition delay and species profiles data from shock tubes, laminar flame speeds, species profile data from flow reactors, and species profile data from flat flames. To this mechanism, more detailed reactions of ethanol were added (5 species and 36 reactions), as proposed by Dryer and co-workers [38–40]. For the reactions of the hydrofluorocarbons (HFCs) in hydrocarbon flames, the NIST HFC mechanism was used [41,42]. Subsequent updates to that mechanism were made by NIST workers, as noted L’esperance et al. [43]. Other changes to the NIST HFC mechanism were made in the present work based on recent experimental measurements and theoretical calculations [44–50] as listed in Table 2.

Table 2 Reaction rate modifications as suggested by Saso et al. [44], Takahashi and co-workers [45,46], Vetters (for T Ü 330 K) [47], Fernandez and Fontijn [48], and Tsai and McFadden [49,50], that were incorporated into the NIST HFC mechanism. Reaction

A (cm/mol/s)

b

Ea (J/mol)

References

CHF3 + H ) CF3 + H2 CO + F + M ) CFO + M CFO + H ) CO + HF CF3 + O ) CF2O + F CF3 + H ) CF2 + HF CF2 + O ) CFO + F CF2 + H ) CF + HF CF + O2 ) CFO + O CHF3 + O ) CF3 + OH CHF + H ) CH + HF CHF + H ) CF + H2

3.76  1013 3.09  1019 2.5  1013 1.54  1013 5.33x 1013 2.45  1013 3.98  1013 6.62  1012 3.07  1014 0.64  1014 2.30  1014

0 1.4 0 0 0 0 0 0 0 0 0

54,810 2040 0 0 0 0 19,000 7070 79,290 0 0

[44] [44] [44] [45] [45] [46] [46] [45] [48] [50] [50]

Table 3 Additional reactions added to the NIST HFC mechanism to account for C2HF5 reaction with early – forming radicals from propane and ethanol decomposition. Reaction

A (cm/mol/s)

b

Ea (J/mol)

CHF2  CF3 + C2H5 = CF3  CF2 + C2H6 CHF2  CF3 + nC3H7 = CF3  CF2 + C3H8 CHF2  CF3 + iC3H7 = CF3  CF2 + C3H8 CHF2  CF3 + C2H4OH = CF3  CF2 + C2H5OH CHF2  CF3 + CH3CHOH = CF3  CF2 + C2H5OH CHF2  CF3 + CH3CH2O = CF3  CF2 + C2H5OH

5.7  1010 5.7  1010 5.7  1010 5.7  1010 5.7  1010 5.7  1010

0.0 0.0 0.0 0.0 0.0 0.0

49,400 43,100 56,500 44,400 66,500 37,200

1019

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

A list of potentially important reactions of C2HF5 with the radicals from initial fuel (propane or ethanol) decomposition was developed, and the rates were estimated as given in Table 3. The barriers for the reactions were estimated in Evans-Polanyi fashion by analogy to that for the reference reaction CHF2  CF3 + CH3 = CF3 – CF2 + CH4 contained in the NIST HFC mechanism by increasing the barriers in proportion (0.3) to the decrease in the heat of reactions relative to the reference reaction. The HFC sub-mechanism finally adopted contained 51 species and 600 reactions. To describe flame inhibition by CF3Br, the bromine parts of the mechanism of Babushok et al. [8,51] were adopted, adding 10 species and 74 reactions. That sub-mechanism has been validated in premixed flame speed and ignition delay studies [8,29,52]. The final mechanism used for the simulations of the aerosol can test fuel, with C2HF5, and CF3Br, has 177 species and 1494 reactions. The aerosol can test fuel is predominantly ethanol (about 2/3 of the energy release), and other models are available that describe ethanol combustion [53–55] (and even more for propane). Nonetheless, we selected the C1–C4 model of Wang as the base-case (and added the ethanol reactions of Dryer and co-workers) since extensive validation studies have been performed by the developers of these mechanisms. 3.4. Perfectly-stirred reactor calculations Flame extinction caused by suppressants is controlled by the characteristic times for chemical reaction and transport, as described by the Damköhler number Da = sr/sc, in which sr is the flow residence time, and sc is the chemical time [56]. Hence, an important step for understanding flame suppression is to obtain some measure of the overall reaction rate. Given the explosive, twophase, turbulent mixing process occurring during release of the aerosol can test simulator fuel [3], the reaction zone there may be simulated reasonably well by a stirred reactor. Further, the stirred-reactor blow-out residence time has been correlated with both the laminar flame speed [57] and with extinction of laminar diffusion flames with added inert suppressants [58], indicating its utility as a measure of overall reaction rate. The residence time in the _ in which q is the mixture density, reactor s is defined as s ¼ qV=m, _ is the mass flow. Heat losses from V is the reactor volume, and m the reactor to the surroundings can also be considered, but are ne-

glected in the present analyses. The governing equations of conservation of mass, species, and energy form a system of coupled nonlinear algebraic equations, which can be solved numerically. In the present work, we employ the SANDIA PSR code [59]. To obtain the characteristic chemical time at extinction using a stirred-reactor model, one must determine the blow-out condition. The process is illustrated in Fig. 1, which shows the reactor temperature as a function of residence time, for three values of the volume fraction of C2HF5 in the oxidizer. At a very low reactor mass flow rate, the residence time in the reactor is long, essentially yielding the equilibrium conditions. As the mass flow in the reactor is increased, the temperature decreases slightly due to incomplete reaction, and there eventually becomes a point at which there is insufficient time to achieve substantial reaction in the vessel; because of the exponential dependence of reaction rate on temperature, this point is a very abrupt change, where the mixture ‘‘blows-out,’’ without reacting, yielding a blow-out time schem. Near blow-out, a criterion of <0.5% change in the mass flow rate was used to determine schem. Figure 1 shows the blow-out condition for a mixture of the aerosol can test fuel, with 65% of the oxidizer in the test chamber (g = 0.65) at a relative humidity of 50%, for C2HF5 added to the chamber air at 0%, 7.7%, and 14.4% volume fraction. As indicated, the characteristic chemical times, schem, obtained for these inhibitor concentrations are 0.43 ms, 2.4 ms, and 32 ms, corresponding to overall chemical rates in the stirred-reactor (xpsr = 1/schem) of 2330 s1, 416 s1, and 31 s1. As the concentration of added C2HF5 increases in this range, the characteristic chemical time increases (i.e., the overall characteristic reaction rate decreases).

4. Results 4.1. Equilibrium calculations The adiabatic flame temperature (of the involved reactants) Tad was calculated for the FAA aerosol can test in the presence of each of the three suppressants (C2HF5, C3H2F3Br, CF3Br) premixed in the chamber air. Calculations were performed over a range of values of

C2HF5 X inh

2500

η ≈ 0.65 50 % r.h.

2000

0

7.7 X C2HF5 = 0 %

14.4

1500

0.072

Tad / K

Reactor Temperature / K

2000

1500

0.135

1000

500

0.43 0 0.0001

0.001

τ = 32 ms

2.4 0.01

0.1

1000

1

10

Residence Time / s Fig. 1. Stirred reactor temperature as a function of residence time for the FAA aerosol can test fuel, g  0.65, 50% R.H., and C2HF5 added at 0%, 7.7%, and 14.4% volume fraction in the oxidizer stream.

0.0

0.2

0.4

η

0.6

0.8

1.0

Fig. 2. Calculated adiabatic equilibrium temperature (Tad) as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of the suppressant C2HF5 in the chamber air.

1020

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

g and Xi, as described in more detail in Refs. [28,60], which provide more details on specification of each of the mole fractions (propane, ethanol, and water in the fuel mix, and N2, O2, ambient water vapor and agent in the oxidizer stream) for the initial conditions. The pressure was 1.01 bar. The results for C2HF5 are shown in Fig. 2. With no agent, the shape of the curve mimics the variation in Tad with fuel–air equivalence ratio [61], and the peak adiabatic flame temperature Tad,peak is reached when about one third of the chamber volume of oxidizer reacts with the aerosol can contents (g = 0.33). As Xinh increases, Tad,peak decreases slightly, but still remains near 2200 K, and the peak value is reached at larger values of g. For inhibitor loadings greater than Xinh = 0.072, Tad,peak decreases abruptly, and thereafter decreases more steeply as Xinh increases. This occurs because above Xi = 0.072, the fluorine content of the system is greater than the hydrogen content, so HF cannot be formed as the most stable final product; rather, COF2 is also formed, with the attendant lower total chemical enthalpy release. The most notable feature of these flames is that as Xinh increases to the value which suppressed the aerosol can test explosion, 0.135, the peak Tad is still around 1900 K, and all of the chamber volume (of oxidizer–inhibitor mix) is required to achieve the peak temperature. Further, at high values of Xinh, the behavior near the peak Tad is more of a plateau than a peak, such that a wide range of values of g will produce nearly the same Tad. This is different from the behavior at low Xinh, for which Tad drops off on either side of the peak, similar to the uninhibited case. Since reaction rates are exponential in temperature, this behavior is likely to affect the overall chemical rates at low and high Xinh. The results for C3H2F3Br are shown in Fig. 3. The result is similar to that of C2HF5; however, the system reaches unity hydrogen to fluorine ratio at Xi = 0.06, so values above that condition are not displayed in the figure. As a result, the Tad,peak is maintained nearly constant at about 2200 K for all values of Xinh up to 0.06. As with C2HF5, increasing amounts of C3H2F3Br require increasing amount of chamber volume of oxidizer to achieve Tad,peak, such that with Xi = 0.06, nearly all of the chamber volume is involved at the fuel–oxidizer ratio corresponding to Tad,peak. Nonetheless, note that in the FAA aerosol can test experiment, even with C3H2F3Br added

at Xinh = 6%, the explosion from the aerosol can fire ball was not suppressed (despite all the oxidizer estimated to be consumed). As with C2HF5, as Xinh increases with added C3H2F3Br, the curves for Tad are very flat near the peak—that is, Tad is not very sensitive to the value of g for high Xinh. The results for CF3Br are different. As shown in Fig. 4, addition of CF3Br up to suppression concentration (around 4%) again decreases the peak Tad only slightly (100 K), but the shape of the curves at high Xinh are essentially the same as at low Xinh. That is, for peak Tad, the fraction of chamber volume required is constant (g = 0.33), not increasing as Xinh increases, as occurs with addition of the other two agents. The effects of the inhibitor volume fraction Xinh on the peak adiabatic flame temperature are summarized for the three agents (C2HF5, CF3Br, and C3H2F3Br) in Fig. 5. The left axis shows the peak Tad, while the right axis shows the value of g required to achieve the peak Tad (gmax). As indicated, all the agents have a minor (lowering) effect on the peak Tad for Xinh < 7%, with CF3Br lowering the temperature the most. At higher Xinh with the agent C2HF5, the peak Tad is lower, due to the inability of the system to form HF (when there are more F atoms in the system than H atoms). The effect of the different agents on the fraction of chamber volume required to achieve peak Tad is dramatic. As agent is added, gmax increases rapidly for C2HF5, and even more rapidly for C3H2F3Br— but not at all for CF3Br. For the halon alternatives, this is a result of their fuel-like nature, (i.e., carbon and hydrogen content), so that the larger molecules have a larger effect on gmax. That is, adding the agent is like adding fuel to the air stream, so that to oxidize the aerosol can test fuel, more oxidizer volume is required. As Xinh goes up, the value of the oxidizer for this purpose is decreased more, so a higher volume fraction of the chamber is required. The reason that CF3Br does not show this effect is due to the unique stoichiometry of this agent in hydrocarbon systems. Since [F]/[H] is always <1 for CF3Br in the FAA aerosol can test, and since there is sufficient water in the products, there are always sufficient H and O molecules left over from the hydrocarbon oxidation (ordinarily in the form of H2O), to supply the H and O necessary to oxidize the CF3Br. A global reaction (exothermic) representing this is: CF3Br + 2H2O ) CO2 + 3HF + HBr (note that HF is a more stable product for H atoms than H2O).

C3H2F3Br CF3Br X inh

2000

0

Tad / K

2000

Tad / K

Xinh 0

1500

0.046

1500

0.03

0.06

1000

1000 0.0

0.2

0.4

0.6

0.8

1.0

η Fig. 3. Calculated adiabatic equilibrium temperature (Tad) as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of the suppressant C3H2F3Br in the chamber air.

0.0

0.2

0.4

0.6

0.8

1.0

Fraction of Chamber Volume Involved in Combustion, η Fig. 4. Calculated adiabatic equilibrium temperature (Tad) as a function of chamber volume fraction involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of the suppressant CF3Br in the chamber air.

1021

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

C3H2F3Br

CF3Br 1.0

Tad,max / K

2000

C2HF5

0.5

max

C3H2F3Br

CF3Br

1500

0

2

4

6

8

10

12

14

0.0

Xinh (%) Fig. 5. Calculated maximum adiabatic equilibrium temperature (left scale) as a function of initial volume fractions of the suppressant in the chamber air, for C2HF5 (4), CF3Br (h), C3H2F3Br (s). The fraction of chamber volume (g) involved in the combustion which gives the peak temperature is shown on the right scale for each suppressant (same symbols); error bars: effects of water vapor.

8

6

max

/ bar

C3H2F3Br

Pressure Rise at

The relative humidity (R.H.) of the ambient air is known to have a significant effect on the behavior of HFCs in hydrocarbon flames (because of the sensitivity of the system to the fluorine–hydrogen atomic ratio [62–64]). Hence, the thermodynamic equilibrium calculations (as well as the kinetic calculations described below) were performed for water vapor in the air corresponding to 0%, 50% and 100% R.H., at 294 K; i.e., a water vapor volume fraction in the ambient air of 0, 0.0125, and 0.025. For the aerosol can test chamber with no inhibitor, saturating the air with water vapor (100% R.H.) as opposed to dry air (0% R.H.) lowered Tad by about 6% for stoichiometric and lean conditions (0.34 6 g 6 1.0), while for rich conditions (0.25 6 g 6 0.20), Tad was raised by 1–2%. With addition of C2HF5, and g > 0.30, Tad was 1–5% higher with saturated air. For CF3Br addition, Tad was at most 1% lower with saturated air. Note than in the FAA tests, the variation of ambient water vapor content was much smaller (6–40% R.H.) than the range examined here, (0–100%). Based on flame equilibrium calculations, and using the value of g which produces the peak adiabatic flame temperature Tad (gmax), it is possible to estimate the pressure rise in the chamber [28,66]. In the calculation, the value of g is selected, and determines the fraction of chamber volume oxidizer (O2–N2–agent) allowed to react with the fuel mix in the aerosol can simulator. An equilibrium calculation (constant P, constant H) gives the volume and temperature of the products, and these are allowed to mix adiabatically with the fraction of chamber volume (1  g) which is treated as inert. The ideal gas law gives the pressure which this volume of products would have attained. (Note that equilibrium calculations assuming constant U, and constant V for the reacting fuel and chamber fraction, essentially the Adiabatic Mixed Explosion Model [65], gave equivalent results.) The results of the calculations of pressure rise are shown in Fig. 6, together with the experimental data from the FAA tests [1,33]. The error bars (in Figs. 5 and 6) show the effects of water vapor (the error limits showing the results for 0% R.H., and 100% R.H.). As indicated, for C2HF5 and C3H2F3Br, the calculations predict the pressure rise reasonably well, except for the case of XC2HF5 = 0.135, for which the explosion was suppressed. Similar

4

C2HF5

2

CF3Br

0

0

5

10

15

Xinh (%) Fig. 6. Calculated pressure rise based on equilibrium thermodynamics evaluated at gmax, as a function of inhibitor volume fraction in the chamber air, for addition of C2HF5, C3H2F3Br, or CF3Br (small symbols and lines: calculation with polynomial curve fits; large circles: FAA experimental data; error bars: effects of water vapor).

to the case of C2HF5 at 13.5%, the pressure rise is not predicted well with added CF3Br added at around 4%, for which the explosion is again completely suppressed. Apparently, kinetic effects limit the extent of reaction with CF3Br, or C2HF5 at the suppression point (or with CF3Br at lower values of Xinh). Below, we employ stirredreactor simulations to estimate the overall rate of reaction of the aerosol can test fuel with air and suppressant; as in the equilibrium calculations above, PSR simulations were performed for a range of values of g and Xinh.

1022

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

4.2. Perfectly stirred reactor calculations The initial conditions for the stirred-reactor calculations are identical to those of the equilibrium calculations, and the results can be presented in a similar format. The stirred-reactor temperature Tpsr is evaluated at the reactor residence time just above the blow-out condition. It is shown as a function of g and Xinh for C2HF5 and CF3Br in Figs. 7 and 8. For each of these suppressants, the peak Tpsr is lower than the peak Tad, by 150–400 K, varying with g, Xinh, and inhibitor type. For C2HF5, Figure 7 shows that as Xinh increases, Tpsr decreases for rich conditions (low g), but increases then decreases for lean conditions (large g), illustrating the fuellike nature of C2HF5 in the stirred reactor. Also, as with the adiabatic flame temperature, for a given value of Xinh, Tpsr is sensitive to g for low Xinh, but relatively insensitive to g for high g; i.e., the curves in Fig. 7 flatten out for high Xinh and g. For CF3Br, Figure 8 shows that for rich flames (g < 0.33), Tpsr is unaffected by addition of agent, while for lean flames (g > 0.33), Tpsr increases somewhat.

The effect of addition of the agents on the overall chemical rate, over a range of g and Xinh, is shown in Figs. 9 and 10. For C2HF5, Figure 9 shows that the effect of C2HF5 on the overall rate depends upon the value of g. At low g (rich conditions), adding agent reduces xpsr drastically. At intermediate values of g (around 0.55), addition of C2HF5 has no effect on xpsr for Xinh up to about 0.06; while at high g (lean conditions), adding C2HF5 first increases the chemical rate, then decreases it. That is, for large fractions of the chamber air involved, adding suppressant enhances the reactivity of the system. Conversely, for a given amount of agent, the effect of g is different at low Xinh versus high Xinh. For example, with no agent, xpsr is highly dependent upon g, whereas for Xinh = 0.135, the overall chemical rate is more mildly dependent upon g (i.e., the curve in Fig. 9 is relatively flat). Clearly the amount of oxidizer (air–agent–water vapor) involved in the reaction with the aerosol can contents has a large effect on the behavior of the system.

C2HF5

10000

X inh

C 2HF5

0 1000

X inh 0

ω psr / s

Tpsr / K

-1

1700

100

0.072

0.072

1500

10

1300

0.133

0.133 1

1100

0.0

0.2

0.4

η

0.6

0.8

1.0

Fig. 7. Calculated Tpsr as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of the suppressant, C2HF5, in the chamber air.

0.0

0.2

0.4

η

0.6

0.8

1.0

Fig. 9. Overall chemical rate xpsr, for C2HF5 as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of C2HF5 in the chamber air.

CF3 Br 10000

CF3Br X inh = 0.046

1000

ω psr / s

Tpsr / K

-1

1700

1500

X inh 0

100

0.024 10

0

1300

0.046

1 1100

0.0 0.0

0.2

0.4

η

0.6

0.8

1.0

Fig. 8. Calculated Tpsr as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of the suppressant, CF3Br, in the chamber air.

0.2

0.4

η

0.6

0.8

1.0

Fig. 10. Overall chemical rate xpsr, for CF3Br as a function of chamber volume fraction (g) involved in combustion with the contents of the aerosol can simulator. Different curves refer to different initial volume fractions of CF3Br in the chamber air.

1023

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025

Variation in the water vapor content of the air (from Xwv = 0 to Xwv = 0.0125) has a varying effect on the overall chemical rate in the PSR. When small amounts C2HF5 are added, the effects are large. For example, with Xinh = 0.02, added water vapor increases xpsr by 12% at g = 0.2, and decrease it by 50% at g = 0.94. With Xinh = 0.06, the effect of water vapor is smaller, decreasing xpsr by 0–13% for the same range of g. With Xinh = 0.14, the decrease in xpsr with water vapor addition is 26% at g = 0.4, and 8% at g = 1.0. While these effects are significant, the variations in xpsr in Fig. 9 with g or Xinh are much larger. For CF3Br addition, Figure 10 shows that the overall chemical rate is always reduced, and the effect is much stronger for lean conditions (large g). This is consistent with the known high temperature sensitivity of the effectiveness of the HBr catalytic radical recombination cycle [66]. For example, at g = 0.6, CF3Br addition at Xinh = 0.045 lowers the overall reaction rate by three orders of magnitude (as compared to the uninhibited case), in contrast to C2HF5 (with g = 0.6), for which the inhibitor was relatively ineffective when added at concentrations up to Xinh = 0.08. For CF3Br, water vapor has a significant effect on xpsr at g for stoichiometric combustion (0.33), lowering xpsr 12% at Xinh = 0% and 90% at Xinh = 0.05; whereas at low (0.16) or high (0.64) g, the effect is smaller, about 12–40%. The results for C2HF5 and CF3Br discussed above are plotted together in Fig. 11. Curves for no agent, and for C2HF5 and CF3Br at their suppressing concentration (Xinh,sup = 0.135 and 0.038) and roughly one half that value ½ Xinh,sup (Xinh = 0.075 and 0.022) are shown. The importance of the actual value of g in the aerosol can tests is illustrated by evaluating the effect of agent addition on xpsr at g = 0.33 and 0.47 (as denoted by the vertical dotted lines in the figure and listed in Table 4). For the case of g = 0.33, xpsr = 7600 s1 for the uninhibited flame, and 625 s1 and 2700 s1 for CF3Br and C2HF5 at ½ Xinh,sup, respectively. That is, C2HF5 reduces the reaction rate more than CF3Br with each added at half its suppression concentration. Similarly, when added at Xinh,sup, xpsr = 40 s1 and 980 s1 for CF3Br and C2HF5, so that again, C2HF5 is expected to slow the reaction rate more. On the other hand, referring to the dotted line in Fig. 1 for g = 0.47, xpsr = 2900 s1 for uninhibited flames, and 1100 s1 and 90 s1 for C2HF5 and CF3Br at ½ Xinh, ext, and 55 s1 and 14 s1 at Xinh,sup, and so that for g = 0.47, the inhibition by CF3Br is stronger. At higher values of g, the superior inhibition by CF3Br is even greater. It is clear from the above discussion that the actual fraction of the chamber gases involved in the reaction with the aerosol can 10000

X inh = 0 %

X CF3Br = 2.4 % 1000

X CF3Br = 4.6 %

ω psr / s

-1

X C2HF5 = 7.2 %

100

contents has a large effect on the expected suppression efficiency of the two agents. As discussed above, for the agent C2HF5 it is possible to estimate the values of g from the pressure rise observed in the experiments (and shown in Fig. 6). That is, if fairly large values of g are not employed, it is impossible to have enough energy release to get the pressure rise observed in the FAA tests. Using Figure 5, we obtain the value of gmax (i.e., g necessary to give Tad, which was required to yield the observed pressure rise for C2HF5 and C3H2F3Br). These values of g are shown as the solid circles in Fig. 11. As shown, for C2HF5 at Xinh = 0.075, g must be about 0.47 to get the observed pressure rise, while for Xinh = 0.135, g  1.05 is required. As indicated, at the higher values of Xinh, values of g close to unity are required to give the observed pressure rise. For CF3Br, the specification of g is more difficult, since, as discussed above, the predicted pressure rise with CF3Br is insensitive to the value of g used in the calculation, so the pressure rise cannot be used to estimate g. (Of course, the actual pressure rise for CF3Br is even less than that predicted based on thermodynamics, due to kinetic influences.) If the actual g in the FAA tests were 0.33 for CF3Br, then the PSR calculations suggest that it should not extinguish the aerosol can test as effectively as C2HF5 (which is contrary to the experimental result). Hence, it is more likely that the actual g for the aerosol can test with CF3Br addition is higher than 0.33. If g were about 0.47 for either agent, then the explanation shown by Fig. 11 is consistent with the FAA experimental results for the agents, since CF3Br is more effective than C2HF5 at g = 0.47, as shown in Table 4. The sensitivity of the inhibitor effectiveness to the value of g is consistent with the findings in some of the FAA tests [67] that a delay in the energy application to the igniter had a large effect on the explosion properties of the system. That is, delaying the spark increases the mixing time prior to ignition, changing the effective value of g in the system (as well as changing the local gas composition near the ignitor). Similarly, tests with CF3Br with added N2 in the oxidizer stream found the combination to work better than either agent alone. This is also likely due to the enhanced effectiveness of CF3Br at lower temperature, as illustrated in the simulations above for which higher g leads to lower overall reaction rate (i.e., higher dilution with the relatively inert air stream lowers the peak temperature, increasing the effectiveness of the HBr catalytic cycle). To highlight the importance of the value of g on the effectiveness of the agents in reducing xpsr, Figure 12 shows, xpsr, as a function of the fraction of the suppressing concentration. The different curves in the figure are for different assumed values of g (0.33, 0.47, and gmax). The following conclusions can be drawn from the figure: 1. More agent generally reduces xpsr, for the assumed values of g. 2. For the case g = 0.47, there is little change in xpsr for the curve for C2HF5 up to 30% of the suppression value. 3. For C2HF5 (blue curves), the reduction in xpsr with addition of agent is similar regardless of the value of g; i.e., for g = 0.33, g = 0.47, or g(Tad,peak).

X C2HF5 = 13.3 %

Table 4 Overall chemical rate predicted by the stirred reactor calculation for C2HF5, CF3Br, and 2-BTP simulant at values of constant g = 0.33 and g = 0.47.

10

Overall chemical rate, xpsr (s1) 1

0.0

0.2

0.4

η

0.6

0.8

1.0

Fig. 11. Overall chemical rate xpsr, for C2HF5 and CF3Br as function of g; curves are given for each agent at the extinction volume fraction, and approximately one half of that value.

g:

0.33

Agent:

C2HF5

0.47 CF3Br

Fraction of suppression volume fraction (%) 0 7600 7600 50 625 2700 100 40 980

C2HF5

CF3Br

2900 1100 90

2900 55 14

1024

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025 10000

CF3 Br, η = 0.33

ω psr / s -1

1000

CF3Br, the reduction in the overall reaction rate with addition of CF3Br is highly dependent upon g. Hence, the relative effectiveness of these agents for reducing the overall reaction rate in the stirred reactor simulations is highly dependent upon the premixed state of the reactants—the assumed composition of which depends upon the mixing state in the FAA test. That is, the relative effectiveness of these agents for inerting the explosion of the FAA test may be dependent upon the particular test conditions employed. Acknowledgments

C2HF5, η= 0.47

100

C2HF5, η= 0.33 C2HF5, η=ηmax CF3 Br, η = 0.47 10

0

50

100

Xinh / Xinh,supp (%) Fig. 12. Overall chemical rate xpsr, for all C2HF5 (D) and CF3Br (C) as function of the fraction of suppression concentration, for different values of g.

4. The effectiveness of CF3Br is very sensitive to the value of g. 5. For CF3Br to be more effective than C2HF5, g must be greater than about 0.4.

5. Conclusions Chemical equilibrium and perfectly-stirred reactor calculations have been performed for the purpose of understanding the unexpected enhanced combustion in the FAA Aerosol Can Explosion simulation test. The equilibrium calculations were used to predict the overpressure in the FAA tests with C2HF5 or C3H2F3Br added at sub-inertion concentrations. With either of these agents, the observed pressure rise is a strong function of both the fraction of chamber volume involved in the combustion (g), and the amount of added agent. At increasing agent volume fractions, the observed large pressure rise is predicted only if a large fraction of the chamber is involved, corresponding to near stoichiometric conditions (including oxygen demand from reaction of the agent), with reaction to equilibrium products. On the contrary, the pressure rise with added CF3Br is shown to be nearly independent of both the fraction of chamber volume involved, and the amount of added CF3Br. This result is shown to be due to the atom balance in the CF3Br–hydrocarbon–air system, for which added CF3Br is predicted to never show an enhanced pressure rise (even if the typical strong chemical inhibition with CF3Br were not to occur). Nevertheless, the high reactivity for the C2HF5 and C3H2F3Br at the concentrations of the FAA tests are somewhat unexpected. To understand the role of kinetics, the inhibition effects of the agents on the overall reaction rate of the system were examined through stirred reactor calculations. Calculations were performed for C2HF5, and CF3Br, over a range of agent concentrations and fraction of chamber oxidizer involved in the combustion. The amount of oxidizer (air + agent) in the test chamber which participates in the high-temperature reaction with the fuel (i.e., the amount of chamber test volume which mixes with and reacts with the fire ball of the fuel reaction), is found to have a major influence on the relative effectiveness of each agent. For the premixed system modeled by the PSR, the role of excess oxidizer (air plus agent) entrainment with the reaction zone varies with the agent. For C2HF5, the reduction in the overall reaction rate with addition of agent is found to be relatively insensitive to the fraction of chamber volume involved (g). On the other hand, for

The authors thank Wing Tsang of NIST, and John Reinhardt at the FAA Technical Center for essential input to the work. Ken Smith and Med Colket at the United Technologies Research Center performed ab initio calculations for C3H2F3Br and graciously provided the thermodynamic data for the C3H2F3Br molecule, allowing us to perform the thermodynamic calculations. Mike Zehe at NASA Glenn Research Center assisted with implementing the C3H2F3Br data in to the NASA CEA program. Brad Williams at NRL kindly provided his changes to the NIST HFC mechanism. The work was supported by The Boeing Company. References [1] J.W. Reinhardt, Behavior of Bromotrifluoropropene and Pentafluoroethane When Subjected to a Simulated Aerosol Can Explosion, DOT/FAA/AR-TN04/4, Federal Aviation Administration, 2004. [2] J.W. Reinhardt, Minimum Performance Standard for Aircraft Cargo Compartment Halon Replacement Fire Suppression Systems (2nd Update), DOT/FAA/AR-TN05/20, Federal Aviation Administration, 2005. [3] T. Marker, Initial Development of an Exploding Aerosol Can Simulator, DOT/ FAA/AR-TN97/103, Federal Aviation Administration, 1998. [4] D.R. Blackmore, G. O’Donnell, R.F. Simmons, Proc Combust. Inst. 10 (1965) 303–310. [5] L.A. Lovachev, V.T. Gontkovskaya, N.I. Ozerkovskaya, Combust. Sci. Technol. 17 (1977) 143–151. [6] L.A. Lovachev, L.N. Lovachev, Combust. Sci. Technol. 18 (1978) 191–198. [7] I.O. Moen, P.A. Thibault, J.H. Lee, R. Knystautas, T. Dean, C.K. Westbrook, Proc. Combust. Inst. 20 (2008) 1717–1725. [8] V.I. Babushok, D.R.F. Burgess, W. Tsang, A.W. Miziolek, in: Halon Replacements, 1995. [9] N. Saito, Y. Saso, C.H. Liao, Y. Ogawa, Y. Inoue, in: Halon Replacements, 1995. [10] H.F. Coward, G.W. Jones, Limits of Flammability of Gases and Vapors, AD0701575, US bureau of Mines, 1952. [11] T.A. Moore, D.S. Diedorf, S.R. Skaggs, in: Proceedings of the 1993 CFC & Halon Alternatives Conference, 1993. [12] S. Kondo, K. Takizawa, A. Takahashi, K. Tokuhashi, A. Sekiya, Fire Saf. J. 44 (2009) 192–197. [13] Number Designation and Safety Classification of Refrigerants, ANSI/ASHRAE Standard 34-2007, American Society of Heating, Refrigerating and AirConditioning Engineers, 2007. [14] Y.N. Shebeko, V.V. Azatyan, I.A. Bolodian, V.Y. Navzenya, S.N. Kopyov, D.Y. Shebeko, E.D. Zamishevski, Combust. Flame 121 (2000) 542–547. [15] G.W. Gmurczyk, W.L. Grosshandler, Fire Safety Science – Proceedings of the Fourth International Symposium, vol. 4, International Association for Fire Safety Science, 1994, pp. 925–936. [16] G. Gmurczyk, W. Grosshandler, Proc. Combust. Inst. 25 (1994) 1497–1503. [17] G.W. Gmurczyk, W.L. Grosshandler, in: W. Grosshandler, R.G. Gann, W.M. Pitts (Eds.), Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, National Institute of Standards and Technology, Gaithersburg, MD, 1995. [18] A. Hamins, G. Gmurczyk, W.L. Grosshandler, C. Presser, K. Seshadri, in: W.L. Grosshandler, R.G. Gann, W.M. Pitts (Eds.), Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST SP 861, National Institute of Standards and Technology, Gaithersburg, MD, 1994. [19] G. Holmstedt, P. Andersson, J. Andersson, in: Fire Safety Science – Proceedings of the Fourth International Symposium, International Association of Fire Safety Science, vol. 4, 1994, pp. 853–864. [20] V.R. Katta, F. Takahashi, G.T. Linteris, Combust. Flame 144 (2006) 645–661. [21] E.A. Ural, Process Saf. Prog. 22 (2003) 65–73. [22] S. Kondo, A. Takahashi, K. Takizawa, K. Tokuhashi, Fire Saf. J. (2010). [23] J.R. Mawhinney, B.Z. Dlugogorski, A.K. Kim, in: Fire Safety Science – Proceedings of the Fourth International Symposium, International Association for Fire Safety Science, vol. 4, 1994, pp. 47–60. [24] A. Hamins, K. McGrattan, G.P. Forney, Unwanted Accelerated Burning After Suppressant Delivery, SP-1004, National Institute of Standards and Technology, 2003.

G.T. Linteris et al. / Combustion and Flame 159 (2012) 1016–1025 [25] D. Trees, K. Seshadri, A. Hamins, in: A.W. Miziolek, W. Tsang (Eds.), Halon Replacements: Technology and Science, ACS Symposium Series 611, American Chemical Society, Washington, DC, 1995. [26] E.J.P. Zegers, B.A. Williams, E.M. Fisher, J.W. Fleming, R.S. Sheinson, Combust. Flame 121 (2000) 471–487. [27] M.R. Nyden, G.T. Linteris, D.R.F. Burgess, P.R. Westmoreland, W. Tsang, M.R. Zachariah, Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST SP 861, National Institute of Standards and Technology, 1994. [28] G.T. Linteris, F. Takahashi, V.R. Katta, H.K. Chelliah, O. Meier, in: Fire Safety Science – Proceedings of the Tenth International Symposium, International Association for Fire Safety Science, 2012, in press. [29] T. Noto, V. Babushok, A. Hamins, W. Tsang, Combust. Flame 112 (1998) 147– 160. [30] G.T. Linteris, D.R. Burgess, V. Babushok, M. Zachariah, W. Tsang, P. Westmoreland, Combust. Flame 113 (1998) 164–180. [31] G.T. Linteris, F. Takahashi, V.R. Katta, Combust. Flame 149 (2007) 91–103. [32] F. Takahashi, G.T. Linteris, V.R. Katta, Proc. Combust. Inst. 31 (2006) 2721– 2729. [33] J.W. Reinhardt, Prevention of a Simulated Aerosol Can Explosion With a Mixture of Halon 1301 and Nitrogen, DOT/FAA/AR-TN08/49, Federal Aviation Administration, 2008. [34] W.C. Reynolds, The Element Potential Method for Chemical Equilibrium Analysis: Implementation in the Interactive Program STANJAN, Version 3, Stanford University, 1986. [35] S. Gordon, B.J. McBride, Computer Program for Calculation of Complex Chemical Equilibrium Compositions and Applications, NASA Reference Publication 1311, NASA Glenn Research Center, 1996. [36] D.A. Sheen, X.Q. You, H. Wang, T. Lovas, Proc. Combust. Inst. 32 (2009) 535– 542. [37] 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, University of Southern California, 2007. . [38] J. Li, A. Kazakov, M. Chaos, F.L. Dryer, in: Proceedings of the Fifth Joint Meeting of the US Sections of the Combustion Institute, Combustion Institute, 2007, p. Paper 26. [39] J. Li, A. Kazakov, F.L. Dryer, Int. J. Chem. Kinet. 33 (2001) 859–867. [40] J. Li, A. Kazakov, F.L. Dryer, J. Phys. Chem. A 108 (2004) 7671–7680. [41] D.R. Burgess, M.R. Zachariah, W. Tsang, P.R. Westmoreland, Prog. Energy Combust. Sci. 21 (1995) 453–529. [42] D. Burgess, M.R. Zachariah, W. Tsang, P.R. Westmoreland, Thermochemical and Chemical Kinetic Data for Fluorinated Hydrocarbons, NIST Technical Note 1412, National Institute of Standards and Technology, 1995.

1025

[43] D. L’esperance, B.A. Williams, J.W. Fleming, Combust. Flame 117 (1999) 709– 731. [44] Y. Saso, D.L. Zhu, H. Wang, C.K. Law, N. Saito, Combust. Flame 114 (1998) 457– 468. [45] K. Takahashi, Y. Sekiuji, Y. Yamamori, T. Inomata, K. Yokoyama, J. Phys. Chem. A 102 (1998) 8339–8348. [46] Y. Yamamori, K. Takahashi, T. Inomata, J. Phys. Chem. A 103 (1999) 8803– 8811. [47] B. Vetters, B. Dils, T.L. Nguyen, L. Vereecken, S.A. Carl, J. Peeters, Phys. Chem. Chem. Phys. 11 (2009) 4319–4325. [48] A. Fernandez, A. Fontijn, J. Phys. Chem. A 105 (2001) 8196–8199. [49] C.P. Tsai, D.L. Mcfadden, J. Phys. Chem. 94 (1990) 3298–3300. [50] J.S. Francisco, J. Chem. Phys. 111 (1999) 3457–3463. [51] V. Babushok, T. Noto, D.R.F. Burgess, A. Hamins, W. Tsang, Combust. Flame 107 (1996) 351–367. [52] G.T. Linteris, M.D. Rumminger, V.I. Babushok, W. Tsang, Proc. Combust. Inst. 28 (2000) 2965–2972. [53] N.M. Marinov, Int. J. Chem. Kinet. 31 (1999) 183–220. [54] P. Saxena, F.A. Williams, Proc. Combust. Inst. 31 (2007) 1149–1156. [55] A. Frassoldati, A. Cuoci, T. Faravelli, E. Ranzi, Combust. Sci. Technol. 182 (2010) 653–667. [56] F.A. Williams, J. Fire Flamma. 5 (1974) 54–63. [57] R.B. Barat, Chem. Eng. Sci. 56 (2001) 2761–2766. [58] S. Liu, M.C. Soteriou, M.B. Colket, J.A. Senecal, Fire Saf. J. 43 (2008) 589–597. [59] P. Glarborg, R.J. Kee, J.F. Grcar, J.A. Miller, PSR: A FORTRAN Program for Modeling Well-Stirred Reactors, SAND86-8209, Sandia National Laboratories, 1986. [60] G.T. Linteris, O. Meier, Unwanted Combustion Enhancement by Potential Halon Replacements: Thermodynamic Considerations, NIST Technical Note 1693, National Institute of Standards and Technology, 2011. [61] I. Glassman, Combustion, Academic Press, San Diego, CA, 1996. [62] G.T. Linteris, in: A.W. Miziolek, W. Tsang (Eds.), Halon Replacements, ACS Symposium Series 611, American Chemical Society, Washington, DC, 1995. [63] K. Takizawa, A. Takahashi, K. Tokuhashi, S. Kondo, A. Sekiya, Combust. Flame 141 (2005) 298–307. [64] G.T. Linteris, G.W. Gmurczyk, in: HOTWC-1995, National Institute of Standards and Technology, 2006, pp. 227–238. [65] R.A. Ogle, Process Saf. Prog. 18 (1999) 170–177. [66] G. Dixon-Lewis, R.J. Simpson, Proc. Combust. Inst. 16 (1977) 1111–1119. [67] J.W. Reinhardt, in: Aircraft Cargo Compartment ‘‘Testing Update’’, International Aircraft Systems Fire Protection Working Group Meeting, Federal Aviation Administration, 2007.