i

The views and conclusions contained in this document are those of the National Institute of Standards

and Technology and should not be interpreted as representing the official policies, either expressed or

implied, of the Strategic Environmental Research and Development Program.

Transient Application, Recirculating Pool

Fire, Agent Effectiveness Screen:

Final Report, NGP Project 3A/2/890

William Grosshandler, Anthony Hamins, and Kevin McGrattan

Building and Fire Research Laboratory

and

Cary Presser

Chemical Science and Technology Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899

April, 2001

Sponsored by:

The Department of Defense

Strategic Environmental Research and Development Program

NISTIR 6733
ii

EXECUTIVE SUMMARY*

Because of its many positive attributes, halon 1301, or trifluorobromomethane (CF3Br), has been used as a

fire extinguishing agent in many applications including aircraft, ships, and specialized structures. Due to its

high ozone depletion potential, however, world-wide production was halted in 1994. In the search for a

long-range replacement, novel types of extinguishing agents and delivery mechanisms are under

development. To gauge the suitability of a replacement agent, methods are needed to evaluate the

material's suppression effectiveness under conditions that relate to field applications. 

In this study, a laboratory-scale facility has been developed to screen the suppression

effectiveness of agents that are delivered in a transient fashion such as solid propellant gas generators. 

The transient application, recirculating pool fire (TARPF) agent effectiveness screen features a propane

fire stabilized behind an obstruction, which is known to be a highly challenging suppression configuration.

The character of the flame and the impact of the air flow, fuel flow, obstruction geometry, and rate of

agent addition on the amount of material needed for suppression are evaluated for N2, CF3Br, and a solid

propellant gas generator (SPGG). The importance of the injection process on the flow field and the

transport of the agent downstream is examined, and a simple mixing model is used to explain the observed

trend of decreasing suppressant mole fraction with increasing injection duration, even for agents as

different as CF3Br and N2 . Direct numerical simulation of the suppression event is shown to successfully

predict the quantity and rate of N2 required to extinguish the flame based upon a published global reaction

rate for premixed propane/air flame propagation. 

Important Findings: 

• The minimum mole fraction of agent for suppression, normalized by the cup burner value,

correlates with [1 - exp(- Δt/τ)]-1, where Δt is the injection time interval and τ characterizes the

mixing time behind the obstacle in the flow. 

• The general character of the flame and its extinction by a thermal gaseous agent is captured by a

direct numerical simulation of the flow based upon single-step chemistry, and numerical

experiments have corroborated the simple correlation of the experimental data for N2.

• The measured difference between the decrease in agent storage bottle pressure and the arrival of

the agent at the fire highlight the importance of determining the agent concentration locally and the

difficulty in relating changes in bottle pressure to actual mixing conditions.

• For the first time, both compressed and solid-propellant generated gases can be compared side by

side, and the effect on performance of different formulations, particle loadings and burning rates

for various SPGG designs can be unambiguously discriminated.

• When the temperature of a hot surface downstream of the pool is above 800 oC, the flame,

following suppression, will reignite and stabilize on the hot surface. At a temperature below 800

oC, the number of reignitions approaches zero. This result is in contrast to when the hot surface is

located between the stabilizing step and the fuel pool, in which a delayed reignition can be

observed at temperatures as low as 400 oC.

• The effectiveness of suppression with a liquid agent is dependent upon how well the droplets are

entrained into the flame zone downstream of the stabilizing obstacle.

 

*

The content of this report is the same as that found in NISTIR 6733 (see bibliography).
iii

ACKNOWLEDGEMENTS

The worked reported herein was supported by the Department of Defense's Next Generation Fire

Suppression Technology Program (NGP), (co-funded) by the DoD Strategic Environmental Research and

Development Program (SERDP). This report is prepared as part of that program. The authors would like

to thank Michael Selepak for his tireless effort to assemble and maintain the experimental facility, 

Michelle Donnelly for assembling and programming the data acquisition system, S. Rao Charagundla for

supervising the operation of the laboratory, and Erik Johnsson for conducting the infrared absorptance

measurements. Consultations with Jiann Yang and John Widmann of NIST and Jose Torero of the

University of Maryland are also gratefully acknowledged. We would like to acknowledge Michael

DiGiacomo and Jamie Neidert of Atlantic Research Corporation for providing the hybrid gas generators

and for assisting in the SPGG experiments. 

DISCLAIMER

Certain commercial products are identified in this report in order to specify adequately the equipment used.

 Such identification does not imply recommendation by the National Institute of Standards and Technology,

nor does it imply that this equipment is the best available for the purpose.
iv

TABLE OF CONTENTS

page

INTRODUCTION 

...................................................................................................................................1

The Fire Suppression Problem

Previous Studies

Objectives

EXPERIMENTAL FACILITY

............................................................................................................ 5

Wind Tunnel and Burner

Gaseous Agent Discharge

Aerosol Agent Control

SPGG Facility and Operation

NUMERICAL MODELING AND

ANALYSIS................................................................................. 15

EXPERIMENTAL RESULTS

............................................................................................................. 20

Flow Characterization

Effect of Obstruction Geometry

Effect of Type of Gaseous Agent

Aerosol Agents

SPGG Results

Effect of Hot Surface

Data Correlation

CONCLUSIONS AND

RECOMMENDATIONS............................................................................... 38

REFERENCES..................................................................................................................................

..... 39

BIBLIOGRAPHY..............................................................................................................................

.... 42

APPENDICES....................................................................................................................................

.... 43

A. Mechanical Drawings of Facility

B. G. Jomaas, B.T. Roberts, J. DuBois and J.L.Torero, "A Study of the Mechanisms Leading to

Re-ignition in a “Worst Case” Fire Scenario, Final Report, Cooperative Agreement No.

70NANB8H0043
1

INTRODUCTION

The Fire Suppression Problem

Water sprinklers are commonly chosen to automatically protect buildings and their contents against fire

because sprinklers are highly reliable, inexpensive to install and maintain, and water is non-toxic and

friendly to the environment. Alternatives to water suppression are usually required if the situation calls for

a very rapid response, if the primary fuel source is a gas or an evaporating liquid, if the protected space

contains electrically energized or other water sensitive equipment, or if the application is weight and

volume limited. Until recently, one would most likely find halon 1301, or trifluorobromomethane (CF3Br),

as the fire extinguishing agent alternative to water, especially if people had a chance of being exposed at

the time of discharge. Due to its high ozone depletion potential, however, world-wide production of CF3Br

was halted in 1994. 

In spite of almost a decade of significant research activity, no single chemical or suppression

system has been identified that has all the positive attributes of halon 1301. Depending upon the

application, one or another candidate system may appear to hold promise. Full-scale suppression testing is

always essential to demonstrate acceptability, in spite of the fact that the full-scale tests are strongly

influenced by the initial and boundary conditions surrounding the fire and suppression event. Complex,

non-linear relationships among the agent, the flow field, and the fire cannot be unraveled from these fullscale suppression tests because critical parameters cannot be controlled tightly enough and the number of

tests is constrained by cost, time, safety and environmental considerations.

The U.S. Department of Defense is committed to reducing its dependence on halon 1301 and has

made great strides in this direction by eliminating non-essential uses, totally revamping its fire suppression

system certification and recycling procedures, and replacing halon systems with alternative technologies

where possible. There remain some applications where no substitute chemical or system has been judged

satisfactory, and several others where the alternatives identified to date are saddled with serious

deficiencies. One of these applications is for military aircraft, which are particularly vulnerable to fire

during combat and also need in-flight fire protection during routine missions, a need shared by the

commercial fleet.

The amount of a gaseous agent required to extinguish fires in full-scale engine nacelles varies

greatly with the geometry of the test fixture and the manner in which the flame is stabilized. It has been

observed that if the test is designed to allow fuel to collect behind obstacles in the vicinity of a hot surface,

a significantly higher mass of agent is necessary for sustained suppression [1]. The superior performance

of chemically acting agents such as CF3Br and CF3I relative to a hydrofluorocarbon alternative like HFC125 is also accentuated in some of these tests. Full-scale testing carried out by the Navy using two

different fixtures, each meant to simulate fires in the F/A-18 engine nacelle, led to different conclusions

regarding the amount and relative performance of both HFC-125 and solid propellant gas generator

(SPGG) fire suppressants [2]. 

The complexity and unpredictability of full-scale tests can be traced to two factors: flame

stabilization, and agent mixing. Flame stability is governed by local geometry, surface temperature, and

fuel and air flow patterns. Flame extinction will occur if the agent is entrained into the flame zone in

sufficient concentration, if the fuel and air flows are disrupted enough by the agent discharge process, or

by a combination of the two effects. Entrainment and localized flame stretch are, in turn, controlled by the

way the fire suppression system is designed and by the location of the fire relative to the discharge nozzle.

Figure 1 shows a full-scale mock-up of an F-22 engine nacelle built by the Air Force Research

Laboratory at Wright-Patterson AFB, which was used to test the suppression effectiveness of HFC-125

and SPGG as compared to Halon 1301 [3]. When the engine mock-up is slid in on its rails,
2

Figure 1. Full-scale F-22 Engine Nacelle Simulator at WPAFB [3, Hamins et al, 1997]

3.8 m
3

an annular region is formed about the core where flammable vapors can build and leaking fuel can

accumulate. Air is brought in through a scoop at the lower right of the structure, it mixes with fuel vapor

and/or a spray from a simulated leaking hydraulic, fuel or lubricating oil line, and is ignited by a spark or hot

surface such as the bleed-air line. A fire can be stabilized anywhere that the fuel and air mixture is within

its flammable limits and the local velocity is below the flame speed. The many obstructions in the flow

formed by the nacelle frame, fuel lines, and miscellaneous equipment in the nacelle and on the engine core

act as flame holders. The fire extinguishing agent is released into the nacelle near the air inlet. Depending

upon the amount of agent released, the rate of agent addition, the presence of hot surfaces or electric arcs,

and the location of the fire relative to the injection point, the fire may or may not be fully extinguished. An

engine nacelle simulator such as the one in Fig. 1 is useful for demonstrating that a particular agent or

system can successfully control full-scale fires, but it is not suited for screening the suppression

effectiveness of new compounds, for ranking alternative technologies, or for gaining understanding of fire

suppression physics.

Previous Studies

Agent suppression effectiveness for liquid fuel fires such as those described in NFPA 2001 is determined

in a laminar diffusion flame using a cup burner [4]. Measurements of suppression effectiveness using a

counter-flow burner are preferred for revealing the detailed chemical pathways in the process [5]. In

either of these burners the minimum concentration for suppression is found by increasing the agent flow

slowly until a critical mole fraction is achieved in the oxidizer stream and flame extinction is observed. In

practice, however, agents designed to replace CF3Br are discharged rapidly, not quasi-statically. Solid

propellant gas generators, for example, typically discharge in 10 ms to 500 ms. A robust and repeatable

means to evaluate the effectiveness of different formulations and application rates requires a nonconventional screening device.

Hirst and Sutton [6] developed a wind tunnel to explore the impact of step height, air flow, and

pressure on the blow-out of a jet fuel pool fire stabilized behind a backward facing step. Hirst et al. [7]

studied the suppression of these types of fires using various halons, and concluded that a liquid pool

burning in a flow behind an obstacle is the most difficult fire to extinguish. This was born out in full-scale

tests done later [8]. Experiments by Hamins et al. [9], in cooperation with Walter-Kidde Aerospace, were

conducted in a wind tunnel scaled down from the earlier work by Hirst to examine the performance of

HFC-125 and HFC-227ea. Investigations at the Air Force Research Laboratory [10] as part of the Next

Generation Program (NGP) sought to determine the detailed structure during suppression of a nonpremixed methane/air flame stabilized behind a step. The changing character of the flame with step height

and air velocity was examined, along with the amount of halon 1301 required to suppress the flame as a

function of the flow parameters and injection interval. 

A turbulent spray burner was designed in earlier work at NIST [11] to simulate an engine nacelle

spray fire resulting from a ruptured fuel or hydraulic fluid line. In this apparatus, the agent release and

mixing process were well controlled and the air flow was maintained constant with a sonic orifice. This

arrangement allowed the agent to be discharged without disrupting the incoming air. The turbulent spray

burner was used with both gaseous and powdered agents. Hamins et al. [9] redesigned the burner to

include a heated disk in the center of the flow downstream of the fuel nozzle. They showed that the

concentration of nitrogen necessary to extinguish a turbulent propane flame increased substantially with

surface temperature. The same trend, but to a lesser degree, was observed with hydrofluorocarbons. 

Previous studies have demonstrated the effectiveness of SPGGs and their hybrids in full-scale fire

suppression experiments [12, 3]. Solid propellant gas generators undergo rapid solid-phase reactions

producing inert gases, principally carbon dioxide, water vapor, and nitrogen, as well as particulates

composed of inorganic salts. Each component individually behaves as a fire suppressant. For many

applications, particulate use is unacceptable and the SPGG hardware is typically designed to filter

particulate mass during deployment. Hybrid generators are particularly attractive for some special

situations such as inhabited spaces or cold temperature applications [13]. The hybrids use the hot inert
4

SPGG exhaust to vaporize and expel a secondary suppressant, typically a liquid, such as water or a low

boiling point halocarbon, through a nozzle [12].

Objectives

The research reported here addresses the issues mentioned above for which a full-scale facility is not

suited; i.e., screening the fire suppression effectiveness of new chemical compounds, ranking alternative

technologies, and gaining understanding of fire suppression physics. The specific objectives put forth in

the original proposal were to provide a well-characterized test fixture for screening the effectiveness of

agents to suppress and prevent re-ignition of a turbulent, obstructed flame; to allow evaluation of the

impact of transient agent delivery on flame extinguishment; and to provide a means to screen the effects

of new and currently available agents on solid and liquid fuel surfaces.

Full funding for the three year period was to deliver the following products:

• a well-characterized bench-scale suppression screen for comparing the performance of gaseous

agents and dispersed fluid mists in suppressing and preventing re-ignition of a turbulent, obstructed

flame burning either gaseous or condensed fuels;

• documentation on the screen apparatus including detailed design information and experimental

procedures, with round-robin testing among interested partners to insure utility of the documentation;

• an evaluation of the impact of transient agent delivery on flame extinguishment; and

• development of a means to screen the effects of new agents on condensed (solid and liquid) fuel

surfaces.

The focus of the research deviated a bit from what was originally proposed, in some instances due to

budget reductions and in others due to discovery. 

• Solid fuels were not examined because there was no evidence that they posed a greater threat

than propane in this scenario. 

• A formal round-robin was not established, but work conducted during the same period at the

University of Maryland (see below) and the Air Force Research Laboratory helped to support the

conclusions drawn in the current study. 

• The design for the liquid mist agent delivery system was not finalized due to the complexity of the

process, although the general direction for such a system and its likely capabilities were revealed. 

• More emphasis was placed on the deployment of an SPGG suppression effectiveness screen.

As part of the initial cost-sharing commitment, NIST gave a grant to the University of Maryland to

examine fluid mechanical and heat transfer phenomena that cause an engine nacelle fire to be particularly

difficult to suppress [14]. The objectives of that study were to evaluate the stability of a recirculation

zone behind a backward-facing step under conditions expected in an aircraft engine nacelle, and to

evaluate the effects of the flow structure on a propane diffusion flame established downstream of the step

to characterize a “worst case” fire scenario. The final report from that grant is included as Appendix B

to this document.
5

EXPERIMENTAL FACILITY

A Transient Application, Recirculating Pool Fire (TARPF) suppression facility has been developed in the

current study to screen different agents and prototype systems, and as an indicator of full-scale

performance. The TARPF agent suppression screen is designed to reproduce the most difficult fire

situation and to allow control of critical agent discharge parameters, including agent discharge rate and

duration, air flow, and obstacle geometry. The performance of gaseous agents, aerosols and solidpropellant gas generators can be examined.

Wind Tunnel and Burner

The TARPF is a small wind tunnel consisting of a number of sections, as shown schematically in Fig. 2.

The main portion of the tunnel is 2.5 m long with a square cross section 92 mm on a side. (Refer to

Appendix A for detailed mechanical drawings of key components of the facility.) Air, supplied by a

compressor rated for a maximum flow of 180 g/s at 1.0 MPa, can be delivered to the tunnel at speeds,

averaged over the 92 mm square cross section, up to 17 m/s. Flow is monitored using a calibrated sonic

orifice and a piezoelectric pressure transducer. A heater is available to increase the inlet air temperature

to above 200 oC. A honeycomb flow straightener and mixing screens are located a meter upstream of the

test section. The burner consists of a sintered bronze plate, 92 mm wide by 190 mm long. Propane is the

primary fuel used, which can be supplemented with a JP-8 spray. The expanded relative uncertainty* in

the flows of fuel and air are ± 5 % of the measured value, based upon the manufacturers' specifications

for the metering orifice and mass flow controllers.

Stainless steel baffles between 10 mm and 55 mm high are located upstream of the burner. A

ramp can also be inserted to form a 25 mm high backward-facing step, as seen in Fig. 2. A 25 mm high

vee-shaped obstruction can be located downstream of the propane pool to form a cavity and a second

anchoring position for the flame. An ethane torch located below the vee and external to the tunnel is used

to heat the surface temperature up to 900 oC for reignition studies. The fire is initiated by a spark across

two protruding electrodes located on the side wall of the test section 20 mm above the surface of the

burner and 20 mm downstream of the step. Heat release rates assuming complete combustion are up to

10 kW. The flame is viewed from above and the side through 6 mm thick Pyrex windows by a 30 Hz

video camera to record the suppression process. For some experiments, a high-speed digital camera

(1000 Hz) is used to investigate the flame dynamics. A photograph of the TARPF with a close up of a

typical flame taken through the window is shown in Fig. 3. 

Gaseous Agent Discharge

The fire suppressing agent is injected downstream of the air metering orifice. Since the air flow is choked

by the metering orifice, the introduction of the agent can be accomplished without altering the total air

flow. This is particularly critical when trying to distinguish chemical from physical modes of extinction. 

Mixing of the agent with the air is facilitated by injecting the agent through two opposed radial ports into

the reduced diameter entrance region. (See Appendix A for details on the injector design.)

 

*

The expanded relative uncertainty on all dimensions and independent parameters described in this report

is ± 50 % of the highest significant place reported, with a coverage factor of 2 [15], unless otherwise

stated.
6

Figure 2. Schematic of step-stabilized pool fire apparatus. Dimensions are in millimeters.

spark igniter

92

190

h

2500

reignition

surface

backward step

flow straightener

suppressant gas

air

diffuser

propane

92
7

Figure 3. Photograph of TARPF with close-up of baffle-stabilized propane flame.
8

Figure 4. Schematic diagram of gaseous agent injection system.

Nitrogen (0.99995 volume fraction pure), CF3Br (commercial grade), and HFC-125 (commercial grade)

are stored as gases in one and two liter stainless steel vessels with the pressure monitored by a high-speed

 (1 ms response) piezoelectric transducer, and the temperature measured with a chromel-alumel (76 mm

diameter) thermocouple. An electronic timer controls the interval (10 ms to 1000 ms) that a solenoid

valve on the agent vessel remains open. The agent passes through a 6 mm diameter orifice before it is

injected through two opposed radial ports into the air passage upstream of the diffuser. A computer

monitors the flow controllers, pressure transducers, and thermocouples, and sends a signal to the electronic

timer to open and close the solenoid valve while releasing the flow of suppressant. Figure 4 shows

schematically the elements of the gaseous injection system.

The mass of the gaseous agent released is determined from the change in pressure and

temperature in the storage vessels. The expanded uncertainty in the calculated mass is ± 2 %, with a

minimum absolute uncertainty of 0.12 g attributable to the resolution of the pressure transducer. The

discharge rate and duration are controlled by the initial agent pressure and an electronically actuated

solenoid valve. Figure 5 is a pressure trace taken during a typical nitrogen discharge. The initial pressure

is set to 0.96 MPa with an electronically controlled metering valve located between a standard high

pressure nitrogen gas bottle and the stainless steel agent bottle. The computer acquires background data

T P

P

Computer control

and data acquisition

Gaseous agent

pulse to TARPF

injection point

1 L agent

vessel

2 L agent

vessel

Solenoid

control valve

Fixed area

metering orifice

High pressure agent

storage bottle

Remotely

controlled needle

valve

Vessel

isolation

valve
9

on the initial state for one second, at which point the electric solenoid valve is opened to allow nitrogen to

enter the wind tunnel. An electronic timer closes the valve after the desired interval. Pressure and

temperature in the agent bottle are measured at a frequency of 1000 Hz during the discharge process. 

The piezoelectric pressure transducer is able to follow the change within about 5 kPa, but the

thermocouple is too slow. To determine the instantaneous mass discharge (dm/dt in Fig. 5), the nitrogen is

assumed to be an ideal gas with the expansion inside the bottle occurring isentropically. The temperature

of the thermocouple is recorded to within 1 oC. However, the actual gas temperature could differ by tens

of degrees since the thermocouple bead adjusts much more slowly due to its thermal inertia. The gas

temperature is more appropriately estimated by the theoretical adiabatic value, Tad. Assuming errors

propagate in a normal fashion, the uncertainty of dm/dt is estimated to be ± 2 g/s.

Three observations can be made from the shape of the discharge curve in Fig. 5: (i) the discharge

data are much noisier than the pressure data, because they rely on the gradient of pressure; (ii) the

measured interval (about 180 ms) is longer than the 150 ms interval set in the electronic timer, due to the

inertia of the solenoid valve; and (iii) the discharge rate at the start of the process is markedly higher than

the average rate.

Aerosol Agent Control

Aerosol agents require a totally different injection system than what has been described above or used

previously. Yang et al. [16, 1999] developed a means to accurately meter alternative liquid agents using a

nebulizer, but their screening device operated in a continuous rather than impulsive mode. Commercial 

Time, ms

-100 0 100 200 300 400 500

Temperature, K

240

250

260

270

280

290

300

Pressure, kPa

400

500

600

700

800

900

1000

Mass Flow, g/s

0

10

20

30

40

50

T

Tad

P

dm/dt

1 liter vessel, N2

set time=150 ms

Liquid reservoir

to drain

Ramp

Hot surface for re-ignition

Solenoids

from compressed air cylinder

70 cm

Porous burner

(Propane 33 mL/s)

65.5 cm

Mist nozzle

Air 

Wind tunnel

Figure 5. Storage vessel pressure trace and temperature during discharge of nitrogen into TARPF.
10

Figure 6. Liquid aerosol agent injection system.

fuel injectors were considered but the flow rates and injection duration were not adjustable over the range

and with the precision necessary for the current application. 

Figure 6 shows the aerosol injection system developed for the TARPF. Several kinds of

atomizers were used to introduce and fully disperse misted agents into the upstream air passage under

various conditions, including the introduction of aerosol at different locations along the tunnel air passage,

different atomizer operating conditions such as different spray angles, liquid flow rates, and transient

duration times. The liquid dispensing system consists of a liquid reservoir, a compressed air cylinder, two

computer-controlled solenoids, and a hollow-cone atomizer. The atomizer was positioned equi-distanced

vertically between the ramp and the upper wall of the wind tunnel, and at discrete locations along the air

passage from 65.5 cm upstream of the propane porous burner to a location downstream near the ramp. 

Several different atomizers were used while trying to obtain a well-dispersed spray of droplets in the air

passage. The atomizers were operated at line pressures that varied between 0.27 MPa and 0.69 MPa in

an attempt to affect the quality of atomization. The atomizers were commercially manufactured to the

following: (1) a nominal spray angle of 70° when operated at 1.03 MPa with a flow rate of 1.9 mL/s, and

(2) a nominal spray angle of 60° when operated at 0.69 MPa with a flow rate of 1.1 mL/s. Operating

these atomizers at nominal pressures normally results in a fully developed spray, and thus an increase

above the nominal condition has a small effect on further reducing droplet size. When positioned at

upstream locations, the spray appeared to fill the entire air passage cross section. 

To initiate a mist for a fixed duration, the data acquisition system activated the solenoid that was

connected to the reservoir. At the end of the discharge, the solenoid is deactivated to terminate the flow

to the atomizer. The second solenoid leading to a drain is simultaneously activated to prevent any residual

flow to the nozzle and dripping from the nozzle.

Liquid Fuel Injector/Hot Surface Arrangement

To investigate different re-ignition scenarios, a hot surface obstruction was set up immediately

downstream of the propane porous burner (see Fig. 6). The vee-shaped obstruction was 25 mm in height,

the base was 25 mm wide, and was centered across 50 mm of the wind tunnel passage (see Appendix A.

for detailed drawings). An S-type thermocouple was place on the external upstream surface of the

obstruction. A high-resistance ceramic adhesive (max: 1500 oC) was used to adhere the thermocouple to

the Inconel surface, 15.5 mm thick. An ethane torch was supported underneath the obstruction to heat the

inner surface to over 1100 oC. A mass flow controller was used to regulate the obstruction surface

temperature. A stream of JP-8 droplets was used to provide a source of re-ignition for the propane flame

by directing the droplets directly onto the heated obstruction after flame suppression. The droplet array

was generated from a 0.14 mm sapphire orifice that was pressure fit into a 3.1 mm tube. The tube was

inserted into the wind tunnel passage from the top wall at a position 25 mm upstream of the flame

stabilizing obstacle. The end of the tube (3.8 mm) was bent 90o into the direction of the air stream, and

was centered between the passage top wall and face of the ramp. The JP-8 was forced through the

injector with nitrogen back pressure at 170 kPa. The low back pressure and several sintered filters (with

a pore size of 7 m and 2 m) were used to help prevent clogging of the injector. Initially, piezo-electric

crystals were used to initiate droplet breakup, but impingement of the fuel on the heated obstruction was

difficult due to entrainment of the individual droplets into the high-velocity air flow. Impingement was

achieved successfully by directing the non-atomized fuel stream on the heated obstruction. The JP-8

injection time (controlled by a solenoid valve), and flow rate were varied in order to optimize the

impingement process. The impingement of the fuel onto the obstruction surface resulted in a significant

decrease in the surface temperature and thus care was taken to compensate for this effect during

experiments.
11

SPGG Facility and Operation

Gas generators are manufactured in discrete units using specific chemical formulations and orifice sizes

that are designed based on the particular application. Unlike a compressed gas or aerosol discharge, the

TARPF operator can control neither the total mass discharged nor the injection time interval of the SPGG.

 To accommodate this limitation, a custom discharge chamber was designed to allow the operator to select

the fraction of the SPGG discharge that is injected into the flame zone. The hardware was

designed to allow standard size gas generators (which contain significantly more material than is required

for suppression in the TARPF) to be evaluated by repeating the test sequence with identical gas

generators and incrementally increasing the fraction of SPGG effluent allowed to flow into the TARPF air

stream. The SPGG effluent (gas and particulate) passes through a metering orifice and is injected into the

air stream using the same manifold as for the compressed gaseous agents. (Refer to Fig. 8a.) The flow

through the bypass port is discarded into a laboratory exhaust hood. The pressure in the discharge

chamber is monitored by a 1 ms response piezoelectric transducer and the temperature is measured with a

76 μm type K thermocouple. A computer monitors the pressure transducer and thermocouple.

Figure 7 shows a photograph of the TARPF and the location of the SPGG injection system; Figure

8a is a schematic diagram of the SPGG injection system; and Fig. 8b is a close-up photograph. The

discharge chamber is made of stainless steel with an internal volume of approximately 200 mL. There are

four main ports on the discharge chamber as seen in Fig. 8: (1) a port with a 3⁄4 NPT female thread

connecting to the gas generator cartridge holder, (2) a variable area metering orifice (1.6 mm to 6.4 mm

diameter) to limit the flow into the TARPF, (3) a bypass port tapped for a 2 NPT nipple, and (4) a 19 mm

port for mounting a pressure relief blow-out diaphragm. A housing made of 6 mm thick steel (seen in Fig.

8) encloses the entire injection system as a precaution against a premature or explosive discharge

Appendix A contains detailed fabrication diagrams of the discharge chamber assembly. The main

chamber body (76 mm x 76 mm x 100 mm) is sandwiched between a cartridge inlet plate and a metering

orifice outlet plate that limits the flow of effluent into the TARPF. Copper gaskets are used to create a

seal between the plates and the chamber. Two 1/8 NPT holes on the side of the main chamber body

allow access for pressure and temperature transducers. The SPGG cartridge holder (see Fig. 8) is

mounted vertically facing upward to Port 1 with a 3/4 NPT nipple. The holder is made of steel pipe 229

mm long and 50 mm in diameter. The lower cap was designed to accommodate the SPGG hardware and

ignition wiring. (Refer to the appendix for details.)
12

Figure 7. Photograph of TARPF showing location of SPGG injection system.

The fraction of the SPGG effluent injected into the TARPF was varied by selecting the size of

Ports 3 and 4 (the pressure relief port). The combined area is referred to here as the bypass area. In an

arrangement similar to that of Port 1, two plates could be attached to Port 3, allowing variation of the

effective opening area (51 mm, 38 mm, or 25.4 mm). In many of the experiments, Port 4 was left open to

maximize the bypass area.

The mass of the SPGG injected into the TARPF is determined from the total mass discharged and

the bypass area ratio, equal to the ratio of the area of the metering orifice to the combined area of the

metering orifice and the bypass area. The expanded uncertainty in the calculated mass is ± 2 %, with a

minimum absolute uncertainty of 0.12 g attributable to the resolution of the pressure transducer. From the

temperature and pressure measurements, the rate of suppressant addition to the incoming air, dm/dt, can

be estimated within an expanded uncertainty of ± 2 g/s. The concentration of the SPGG in the air flow is

determined from the rate of suppressant addition (dm/dt) and the mass flow rate of air. As in the gaseous

agent experiments, the rate of air flow is invariant during the agent discharge through the use of a sonic

orifice positioned upstream of the agent discharge location, in the air duct.

SPGG Injection

System
13

(a.) Schematic diagram of SPGG injection system

(b) Photograph of injection system with housing cover removed, showing the SPGG cartridge holder,

discharge chamber, and ports: (1) SPGG outlet, (2) metering orifice, (3) bypass, and (4) blow-out.

Figure 8. SPGG injection system chamber. 

Burner

Bypass

Exhaust

Air

Agent Injector Section

Discharge

Chamber

SPGG Cartridge

Holder

Explosion

Protection

Housing

Blow-Out Disk

SPGG Ignition Wire

 2

Discharge Chamber

1

 Cartridge Holder

3

4

Protective

Housing
14

In the current study, identical commercial air-bag hybrid gas generators were used, one in each of

fifty experiments. Each generator released 20.7 g ± 0.1 g. The discharged mass was specified by the

manufacturer and confirmed experimentally by weighing the generators before and after each discharge.

The agent is composed of twenty grams of compressed argon gas and 0.7 g of a solid propellant, which at

equilibrium converts to KCl (s) , H2O, N2 and a small amount of gaseous CO2 [17]. 

The gas generator was discharged after steady-state fire conditions were achieved in the TARPF.

The discharge was controlled by engaging an electronic switch on a control box that completed a circuit

leading from a 12 V battery to the electrical connector located on the gas generator. One ampere was

required to fire the 40 mg squib, which is an intrinsic part of the gas generator. The squib ignites the solid

propellant, which rapidly discharges. The combustion products of the solid propellant propel the gaseous

argon from the generator casing, located within the cartridge holder, into the discharge chamber. SPGG

cartridges were changed and prepared for the next run in less than four minutes.
15

NUMERICAL MODELING AND ANALYSIS

To better understand the fluid dynamics of the suppression event, a computational fluid dynamics (CFD)

model was used to simulate the baffle or step-stabilized flame in the TARPF facility. Similar numerical

studies have been performed in the past few years. Among these, Liou and Hwang [18] used a twodimensional CFD model to study the residence time of tracer particles within the recirculation zone of a

backward-facing step. Weller et al. [19] applied a large eddy simulation (LES) model to study a premixed

turbulent flame stabilized by a backward-facing step. Here, a low Mach number CFD model is applied to

study the suppression event in the TARPF facility. The model is called the Fire Dynamics Simulator

(FDS), and the simulations described here were performed with the first publicly released version [20].

The model is typically used to simulate large-scale fires, in which case turbulence is handled with a simple

LES model. As applied here, however, the sub-grid scale turbulence model is not used, but rather the

coefficients of viscosity, thermal conductivity and mass diffusivity are derived from kinetic theory and

empirical extrapolation [21]. Thus, the calculations directly simulate the fluid motion (although not the

combustion).

A brief description of the model equations is given below. The full numerical method used to solve

the equations is given in Ref. [20]. First, consider the following conservation equations of mass,

momentum and energy of a mixture of perfect gases in the low Mach number limit:

Conservation of Mass

(1)

Conservation of Species

(2)

Conservation of Momentum

(3)

Conservation of Energy

( ) + ∇ • ( u) − = − Χ −∇ • q

∂

∂

•

(1 )q' ' '

Dt

Dp h h

t

r r r

(4)

Here, r is the density, Yi the mass fraction, Wi

&

 the production rate of the ith component of the mixture,

u=(u,v,w) is the velocity vector, p the pressure, g the gravity vector, h the enthalpy, q&''' the rate of heat

release per unit volume, and Χr the radiative heat loss fraction, taken as 0.20 for methane [22]. The

components of the viscous stress tensor t are
16

















∂

∂

−

∂

∂

+

∂

∂

=

k

k

ij

i

j

j

i

ij

x

u

x

u

x

u

3

2

t m d (5)

where m is the dynamic viscosity. The energy flux vector q is given by

i i

i

q = −k∇T − ∑rDih ∇Y (6)

where k is the thermal conductivity of the mixture, T is the temperature, and hi is the enthalpy of the ith

component.

The conservation equations are supplemented by an equation of state relating the thermodynamic

quantities r, p and h. The pressure is first decomposed into three components, a ``background'' pressure

po, a hydrostatic contribution, − rgz, and a perturbation to the hydrostatic, .

~p The coordinate z is the

vertical spatial component. Using this definition of pressure, the equation of state can be written in a form

appropriate for a perfect gas in the low Mach number regime [23]:

(7)

Here, Mi is the molecular mass of the ith species, M is the average molecular weight of the mixture and

R is the universal gas constant. The pressure p in the state and energy equations is replaced by the

background pressure po to filter out sound waves that travel at speeds that are much faster than typical

flow speeds expected in fire applications.

A further assumption about the thermodynamic variables is that the constant-pressure specific

heat of the ith species cp,i is assumed to be independent of temperature. Under this assumption, the

enthalpy can be written as:

(8)

The specific heat for each species can be expressed in terms of the number of internal degrees of

freedom active in that molecule.

(9)

If the ratio of specific heats for all the species is assumed to be that of a diatomic molecule

(n = 5,g = 7 / 5) , the equation of state can be rewritten in the form:

(10)
17

The basis of this approximation is that nitrogen will be the dominant species in the simulations.

The coefficients of viscosity, thermal conductivity and material diffusivity are approximated from kinetic

theory. The viscosity of the lth species is given by

(11)

where sl is the Lennard-Jones hard-sphere diameter (Å) and Ωnis the collision integral, an empirical

function of the temperature T. The thermal conductivity of the lth species is given by

(12)

where the Prandtl number Pr is 0.7. The viscosity and thermal conductivity of a gas mixture are given by

(13)

The binary diffusion coefficient of the lth species diffusing into the mth species is given by

(14)

where Mlm=2(1/Ml+1/Mm)

-1

, slm sl = ( +sm) / 2 , and ΩD is the diffusion collision integral, an empirical

function of the temperature T [21]. It is assumed that nitrogen is the dominant species in any combustion

scenario considered here, thus the diffusion coefficient in the species mass conservation equations is that

of the given species diffusing into nitrogen:

(15)

where species 0 is nitrogen.

A simple one-step, finite-rate reaction is used to model the combustion of propane:

(16)

The fuel depletion rate (unit mass / unit time / unit volume) is given by the expression

(17)

The heat release rate term is
18

(18)

where ΔH is the heat of combustion. Westbrook and Dreyer [24] suggest values for propane of

11 B = 8.6×10 cm3

 mol-1 s

-1

, E=126.6 kJ/mol, a=0.1 and b=1.65 for propane. The heat of combustion is

assumed to be 46,400 J/g.

Both two and three-dimensional simulations have been performed. The advantage of the two

dimensional calculations is that greater spatial and temporal resolution can be exploited. The disadvantage

is that the much of the complex structure of the turbulent flame cannot be simulated. Three-dimensional

calculations are much costlier to perform, but yield a great deal of information about the flame structure.

The solution of the conservation equations governing the flow in the tunnel is computed on either a

two-dimensional, uniformly-spaced grid spanning a plane 552 mm long and 92 mm high, or a three

dimensional, uniformly spaced grid spanning a volume 552 mm long, 92 mm wide and 92 mm high. The

number of the grids are 576 by 96 and 256 by 48 by 60, respectively. The backward-facing step is

approximated with masked grid cells. The reactant flows consist of oxygen, nitrogen and propane.

Nitrogen is used to represent the products of combustion simply to reduce the computational time. Propane

is introduced into the flow domain through a 190 mm by 92 mm vent in the floor of the tunnel at a uniform

rate of 33 mL/s. Air and agent are introduced into the domain 100 mm upstream of the step with a top hat

velocity profile. At the outflow boundary a constant ambient pressure is assumed. Ignition is achieved by

momentarily heating up a small patch on the floor of the tunnel just downstream of the step. At all other

wall locations, the temperature is maintained at 200 oC.

Presented in Fig. 9 are sequences of images separated in time by 0.01 s that are taken from

simulations that illustrate the dynamics of the suppression event. Figure 9a is a 2-D simulation of flow

(moving left to right) over a 25 mm step. The darkness of the image reflects the extent of local heat

release. Figure 9b is a 3-D simulation of a flame stabilized on a 25 mm baffle that is undergoing a

successful suppression. In both Figs. 9a and 9b the top images show the flame just prior to discharge of 

nitrogen into the air stream. Upon injection, the flame is disturbed by a large vortex generated by the

pressure pulse. Due to the low Mach number approximation, the gas upstream of the step is essentially

incompressible, and the velocity jump from 2.1 m/s to 5.7 m/s is conveyed to all points in the flow domain

in 0.02 s, the time of the ramp-up from the base velocity to the injection velocity. Thus, even before the

agent arrives at the step, the flame has already been dramatically transformed from its original state. The

generation of the large vortex at the step produces a pathway by which the agent can penetrate the region

just behind the step, mixing with the gases, cooling and diluting the fuel and oxygen.
19

Figure 9. The disruption of a stabilized flame by the injection of nitrogen upstream of 25 mm high obstacle:

(a).2-D simulation of flow over step; (b) 3-D simulation of flow over baffle (frames are separated by 0.01

s).

(a)

(b)
20

EXPERIMENTAL RESULTS

Flow Characterization

The facility was operated over a range of propane and air flows to examine the flame behavior. Blow out

can be achieved either by increasing the air flow or decreasing the propane flow. At low air velocities, a

fluctuating laminar flame is anchored on the top downstream edge of the step or baffle and extends well

downstream of the porous plate. As the velocity increases, the flame becomes turbulent and less luminous.

 Near blow-out, the orange color disappears and the visible blue flame shrinks. With the backward-facing

step installed, an average air velocity above the step of over 23 m/s is necessary to blow out the flame if

the propane flow is greater than 33 mL/s (corresponding to a transverse velocity of 1.9 mm/s). 

Two air flows were chosen to evaluate the ability of the agents to suppress the propane pool fire

stabilized by the backward-facing step. The low and high mean air velocities (just above the step) were

2.1 m/s ± 0.2 m/s and 5.4 m/s ± 0.2 m/s, respectively. Corresponding propane flows of 33 mL/s ± 2

mL/s and 85 mL/s ± 2 mL/s, respectively, were utilized. The low flow condition corresponds to what

Takahashi et al. [10] describe as regime I suppression (rim-stabilized flame), and the high flow is

transitional between regimes I and II (intermittent turbulent flame).

The velocity distribution of the air 76 mm upstream of the burner was measured with a 3 mm

diameter pitot tube at seven locations across the duct. Figure 10 compares the results with and without

the flame present for the high flow condition. The velocity profiles are seen to be flat within 5 % over the

central three-fourths of the duct. The boundary layer above the step appears to be less than 7 mm thick. 

 The presence of the flame tends to increase the pitot probe signal, which is likely due to a combination of

preheating the air upstream by the flame, acceleration in the flow due to partial blockage of the duct

caused by the expanding combustion gas, and a possible shift in electrical output due to heating of the pitot

probe and transducer.

The facility is designed to introduce suppressant impulsively without altering the air flow. This is

achieved by maintaining a choked condition for the air independent of modest changes in downstream

pressure generated by the injection process. The pitot tube was used to measure the instantaneous flow

76 mm ahead of and 5 mm above the backward step during the discharge of C2HF5 into the air stream for

the two different air flow conditions (without fuel flowing). The dashed lines in Fig. 11 show the combined

effect of the velocity (V) and density (ρ) change, V(ρ/ρο)

1/2, created by the injection process, where ρο is

the initial density of the air stream. Figure 11a is for the low air flow condition and a high rate of agent

discharge; Fig. 11b represents a high air flow with a low rate of agent addition. The pressure, P, in the

agent storage vessel is also plotted in Fig. 11, from which the rate of C2HF5 mass added, dm/dt, is

calculated. The injection interval is 130 ms ± 5 ms for both cases, but the amount of agent added in Fig.

11b is 1/3 the amount added in Fig. 11a because the agent storage volume was 1 L and 3 L, respectively. 

The sizable increase in V(ρ/ρο)

1/2 seen in Fig. 11a within 0.10 s of the passage of the acoustic

wave results from the slug of air between the injector and the pitot tube being shifted downstream by the

addition of agent. High-speed video images of the flame during the discharge corroborate this description.

 The shift is barely discernable in Fig. 11b since the amount of agent added is small relative to the flow of

air. The time that the agent itself arrives at the pitot tube is limited by the bulk convection and the distance

the probe is downstream of the point of injection (1.1 m). The times of arrival of the agent at the pitot tube

can be estimated to be 0.58 s and 0.24 s for the conditions in Figs. 11a and 11b, respectively. Interpretation

of the pitot signal is complicated by the much higher density of C2HF5 as compared to air. 

An instrument developed by Pitts et al. [25] was used to measure the infrared absorption by

C2HF5 at a wavelength of 8.7 μm ± 0.1 μm. Figure 11 also contains a plot of the absorptance,
21

Figure 10. Velocity (normalized by nominal mean) measured upstream of burner, with and without flame

lit.

integrated across the width of the duct, during the discharge. The viewing port is 2.3 m downstream of

the agent injector and 1.1 m downstream of the pitot tube. The absorptance pulse in Fig. 11b is similar in

shape to the rate of discharge, but about 0.1 s wider. Nonuniform mixing of the agent with air as it flows

over the step and into the recirculation zone contributes to the long absorptance tail that is evident for over

a second in Fig. 11a, although the agent injection duration was only 0.13 s.

The amount of suppressant necessary to extinguish a fire in the TARPF depends upon the fuel and air

flows chosen to challenge the suppressant. Figures 12 and 13 show how the amount of nitrogen

necessary to extinguish a 25 mm baffle-stabilized flame varies with the flow of air and fuel. The

uncertainty in any given value is estimated to be ± 0.2 g. The filled circles in the figures indicate

extinction and the crosses represent no extinction. When the air speed is less than 5 m/s and the fuel flow

is fixed at 45 mL/s ± 2 mL/s, decreasing the speed (see Fig. 12) reduces the amount of nitrogen

necessary to extinguish the flame. No flame extinction occurred between 5 m/s and 15 m/s because the

amount of nitrogen necessary exceeds the maximum amount contained in the storage vessel. Above 16

m/s, the strain on the flame is sufficient at times to extinguish the flame without the need for any nitrogen.

 (The dashed lines are included in the figure to assist the eye in identifying the extinction boundaries.) 

The propane flow does not have much affect on the amount of nitrogen needed to extinguish the flame if

the injection interval and air flow are fixed, as shown in Fig. 13. There is a lower limit for the propane

(<12 mL/s) that leads to extinction due to heat loss to the burner, even with no nitrogen dilution. The upper

limit of propane (120 mL/s) is dictated by the maximum safe operating temperature of the burner;

however, since the mass of nitrogen needed to suppress the flame does not

Normalized Velocity

0.0 0.5 1.0 1.5

Height Above Step, mm

0

20

40

60

flame off

flame on
22

Figure 11. Bottle pressure, agent mass flow, velocity above step, and IR absorptance 1 m downstream of

step during 0.125 s discharge of C2HF5; (a) low air flow, high agent flow, (b) high air flow, low agent flow.

Time, s

0.0 0.5 1.0 1.5 2.0

Pressure, kPa

0

100

200

300

400

500

600

Absorptance [unitless], Mass Flow [g/s], or V(ρ/ρo

)1/2 [m/s]

0

1

2

3

4

5

6

Time, s

0.0 0.5 1.0 1.5 2.0

Pressure, kPa

0

100

200

300

400

500

600

Absorptance [unitless], Mass Flow [g/s] or V(ρ/ρo

)1/2 [m/s]

0

1

2

3

4

5

6

α x 1 0

dm/dt

P

α x 10

dm/dt

P

V (ρ/ ρo)

1/2

V (ρ/ ρο)

1/2

(a)

(b)
23

Nominal Air Speed, m/s

0 5 10 15 20

Mass of Nitrogen, g

0

2

4

6

8

10

12

Extinction

No Extinction

Extinction

Figure 12. Impact of air speed on extinction of 25 mm baffle-stabilized flame; propane flow is 45 mL/s, 

nitrogen injection time is 312 ms: circles imply extinction, crosses imply no extinction.

Propane Flow, mL/s

20 40 60 80 100 120

Mass of Nitrogen, g

0

1

2

3

4

5

6

7

Extinction

No Extinction

Figure 13. Impact of propane flow on N2 required for suppression of 25 mm baffle-stabilized flame;

air flow is 3.88 m/s, injection time is 185 ms; circles imply extinction, crosses imply no extinction.
24

appear to increase, there is no need to operate the burner at higher fuel flows. (Note: the dip in mass

required for propane flows around 75 mL/s is attributed to the low number of experiments conducted in

this region.)

The relationship between the total mass of nitrogen required for suppression and the injection time

interval is plotted in Fig. 14 for a 3.9 m/s ± 0.2 m/s air flow, 45 mL/s ± 2 mL/s propane flow flame

stabilized on the 25 mm baffle. The open symbols are experiments that did not extinguish the flame, and

the closed symbols are experiments that led to extinction. (The total number of experiments conducted

exceeds by a factor of ten the number of data points plotted in this and the following curves; for clarity,

only those conditions close to the extinction boundary are included.) As the injection interval increases

from 100 ms to 500 ms, the minimum mass required increases over three-fold. The rate of mass addition

(calculated by dividing the total mass by the estimated injection interval) decreases with increasing

injection interval, as shown in the right-hand figure.

Effect of Obstruction Geometry

The data plotted with squares in Fig. 15 were taken with the ramp placed in front of the 25 mm baffle to

form a backward step-stabilized flame, rather than the simple baffle-stabilized flame represented in Fig.

14. The air flow is the same in these two cases but the propane flow is higher, 85 mL/s, in Fig. 15. The

addition of the ramp and increase in propane flow do not have much influence on the mass of nitrogen

required for suppression. For both the baffle and backward step, just under 6 g of N2 are required when

the injection interval is 200 ms ± 10 ms. The data plotted as circles in Fig. 15 were taken with the nominal

air speed reduced to about 1.5 m/s and the propane flow reduced a proportionate amount to near 33 mL/s.

 The squares in Fig. 15 represent experiments conducted at the high air and propane flows, and circles

represent experiments conducted at the lower flows. Open symbols indicate that the flame was not

extinguished, and filled symbols indicate flame extinction. Less than 4 g ± 0.2 g of N2 are needed to

extinguish this flame if injected over a 200 ms ± 10 ms interval. The differences in rates of mass addition

to suppress the high flow and low flow flames can also be seen at the right in Fig. 15. 

Figures 16 and 17 show what happens to the required nitrogen mass and addition rate if the baffle

height is decreased to 10 mm or increased to 55 mm (blockage from 11 % to 60 %), respectively. The

symbols have the same meaning as in Fig. 15. The short baffle produces a fire which is the easiest to

extinguish, and the high baffle the most difficult in terms of the amount and rate of N2 addition.

The effect of baffle height is not large if the injection interval is at least 150 ms, as can be seen

more clearly in Fig. 18. (Note that 6 mm has been added to the height of each obstacle to account for the

distance between the floor of the tunnel and the recessed top surface of the burner.) The bottom curve

delineates the minimum amount of nitrogen for suppression when the air flow is fixed at its high value and

the agent injection interval is maintained at 175 ms ± 10 ms. The open circles represent the largest mass

of N2 that did not result in extinction for flames stabilized on the different sized baffles; the filled circles

are the minimum mass of agent that successfully extinguished the flames. The diamonds are the results

for the 25 mm baffle with the ramp in place (backward-facing step). Experiments were also conducted

with and without the reignition obstruction shown in Fig. 2. The amount of N2 necessary for suppression

was unchanged.

The rate of mass addition is plotted in the upper curve of Fig. 18 (the triangles are the backwardfacing step, and the squares are for the baffles). The data are plotted two ways: the higher value is the

rate of nitrogen addition computed during the first 50 ms that the solenoid valve is open (refer to the shape

of the dm/dt curve in Fig. 5); the lower value is the average over the entire open interval measured from

the pressure trace. The estimated rate of mass addition varies substantially, especially for the 55 mm

baffle, depending upon whether the averaging period is the first 50 ms or the entire time that the solenoid

remains open.
25

Injection Time, ms

100 200 300 400 500

Mass, g

0

2

4

6

8

10

Extinction

Injection Time, ms

100 200 300 400 500

Rate of Mass Addition, g/s

0

10

20

30

40

50

Extinction

No Extinction

Injection Time, ms

100 200 300 400

Mass, g

0

2

4

6

8

10

Injection Time, ms

100 200 300 400

Rate of Mass Addition, g/s

0

10

20

30

40

50

No Extinction

Figure 14. Mass and rate of nitrogen addition required to extinguish 3.88 m/s air flow, 

45 mL/s propane flame: filled diamonds, extinction; open diamonds, no extinction.

Figure 15. Mass and rate of nitrogen addition required to extinguish high flow (squares) and

low flow (circles) air/propane flames: filled symbols, extinction; open symbols, no extinction.

25 mm Baffle 25 mm Baffle

25 mm Step 25 mm Step
26

Injection Time, ms

100 200 300 400

Mass, g

0

2

4

6

8

10

Injection Time, ms

100 200 300 400

Rate of Mass Addition, g/s

0

10

20

30

40

50

Injection Time, ms

100 200 300 400

Mass, g

0

2

4

6

8

10

Injection Time, ms

100 200 300 400

Rate of Mass Addition, g/s

0

10

20

30

40

50

Figure 17. Mass and rate of nitrogen addition required to extinguish high flow (squares) and

low flow (circles) air/propane flames: filled symbols, extinction; open symbols, no extinction.

10 mm Baffle 10 mm Baffle

55 mm Baffle

55 mm Baffle

Figure 16. Mass and rate of nitrogen addition required to extinguish high flow (squares) and

low flow (circles) air/propane flames: filled symbols, extinction; open symbols, no extinction.
27

Obstacle Height, mm

0 10 20 30 40 50 60 70

Mass, g

0

2

4

6

8

10

12

Rate of Mass Addition, g/s

0

10

20

30

40

injection time = 175 ms

air flow = 3.88 m/s

Figure 13. Impact of obstacle height on the total mass and rate of addition 

of nitrogen required to suppress obstacle stabilized pool fire.

Figure 18. Impact of obstacle height on the total mass and rate of addition of nitrogen required to suppress

obstacle stabilized pool fire.

Effect of Type of Gaseous Agent

Figure 19 is a plot of the minimum agent mole fraction that extinguished the fires (Χ) as a function of the

agent injection time interval (Δt) for both N2 and CF3Br. The parameter Χ is defined as the average

volume flow of agent during the injection interval divided by the sum of the agent and bulk air flows. The

data represent experiments conducted over a range of conditions including air velocities (defined above the

obstacles) that varied from 2 m/s to 9 m/s, propane flows from 33 mL/s to 85 mL/s, and baffle heights

between 10 mm and 55 mm, in addition to the 25 mm backward step. The open and closed symbols

represent the low and high air flow conditions, respectively. Figure 19 shows that Χ decreases with

increasing injection time interval for all obstacle types and both agents. The highest mole fraction

requirements were consistently for the low air flow conditions. For some experiments, the value of Χ was

nearly 0.8 for short injection intervals. The most challenging geometric configuration was the 55 mm

baffle, followed by the 25 mm obstacles, and the 10 mm baffles. There was little difference in Χ between

the 25 mm step, 25 mm cavity and the 25 mm baffle and those data are presented as one group in Fig. 19.

The effectiveness of CF3Br was compared to that of N2 using the 25 mm high backward step. The 1 L

storage vessel was used to accentuate the pressure change associated with the small quantities of CF3Br

required for suppression. Only 1.6 g ± 0.2 g of CF3Br injected for 100 ms was needed to extinguish the

flame under the high air flow conditions (corresponding to Χ ≅ 0.075), as compared to 3.9 g ± 0.2 g

for N2 as agent (Χ ≅ 0.5) under similar conditions. These results are consistent with numerous studies

that show that CF3Br is a more effective suppressant than N2 for both free standing and baffle stabilized

flames.
28

Figure 19. Mole fraction of agents (N2 and CF3Br) added to air at extinction boundary for high and low

flow conditions, as a function of the injection time interval and obstacle geometry.

Aerosol Agents

The aerosol injection system described earlier was used by Pitts et al. [26] to compare the suppression

performance of a water mist and methoxynonafluorobutane in the TARPF to that measured in the steadystate Tsuji burner [16]. The air flow in the tunnel was fixed at 6.67 x 10-3

 m

3

/s and the porous burner was

operated at a propane flow rate of 33 mL/s. The mist discharge duration was varied from 1 s to 10 s. 

The mass flow rate of the liquid agent, magent

& , can be calculated using the density of the liquid and the

calibrated volume flow rate. The mass fraction of the liquid agent, bagent, in the air stream is then

 

agent air

agent

agent m m

m

& &

&

+

b = (19)

where mair & is the calculated mass flow rate of air. Note that in writing Eq. (19), it is implicitly assumed

that the mist droplets are homogeneously dispersed in the carrier phase (air). 

Injection Time Interval, Δt, ms

0 100 200 300 400 500 600

Mole Fraction of Agent, 

Χ

0.0

0.2

0.4

0.6

0.8

1.0

circles - 55 mm baffle

triangles - 25 mm step, baffle, cavity

squares - 10 mm baffle

open symbols - low flow

filled symbols - high flow

CF3Br

N2
29

A number of suppression experiments were run with the mass fraction of the water calculated to

be 0.11. The pool fire stabilized behind the backward-facing step of the ramp could not be completely

suppressed under these conditions, although useful observations could be made. Before the mist injection,

the flame was luminous and yellow. During the mist injection, the yellow flame was nearly extinguished

with pockets of localized blue flame regions. The mist was also observed to have high momentum, which

resulted in the major portion of the mist bypassing the flame completely. The duration of the small blue

flames coincided with that of the mist application. Once the mist application period was over, the flame

regained its original burning intensity and yellow luminosity.

The above observations indicate partial localized flame suppression due to sufficient droplet

entrainment into some portions of the mixing layer, but an insufficient number of droplets entrained over

time to prevent the flame from flashing back into the previously extinguished zones. In flow experiments

conducted with no flame present, the droplets were observed to traverse the burner very rapidly with little

entrainment of the mist into the recirculation zone. With a flame, gas expansion effects are enhanced and

are expected to further carry the droplets axially past the flame. Although the droplet statistics were not

measured, one would expect that the size distribution from these atomizers would cover a range between 1

μm and 200 μm. For droplet entrainment into the recirculation zone behind the ramp to occur, a high

concentration of micrometer-sized droplets needs to be dispersed throughout the passage cross section

and entrained in the air flow. Apparently, a large enough concentration of small droplets (with a small

relative velocity) was unavailable to suppress the flame completely. 

In the Tsuji burner experiments, Yang et al. [16] used a droplet delivery system that assured

efficient transport of fine liquid droplets to the flame. They found that a calculated water mist mass

fraction of 0.03 in the air stream was sufficient to extinguish their laminar counter-flow flame. This is less

than 1/3 the mass fraction of water injected into the TARPF without achieving suppression. 

Thermodynamic estimates of the cooling effect of water suggest that a mass fraction of almost 0.13 is

required to bring the equilibrium temperature of a propane/air flame down to 1500 K, which is the

extinction temperature found by Pitts et al. [25] for purely thermal agents. Thus, one could argue that a

0.11 mass fraction should not be expected to suppress the TARPF flame. 

More experiments and computational modeling are necessary to sort out the discrepancy between

the TARPF and Tsuji results to identify the relative importance of droplet entrainment, size distribution,

and transient vs. steady-state effects on the mass fraction of water needed to ensure flame extinction. 

This may best be achieved using an ultrasonic humidifier (producing all droplets less than 2 μm), or twinfluid atomizer, to inundate the recirculation zone with submicron droplets, as long as the droplets do not

pre-vaporize before reaching the flame. Note that entrainment of droplets into the recirculation zone will

be dependent on determination of the Stokes number, St = rD

2

DU/18md, where r is the droplet density,

DU is the relative velocity between the droplet and surrounding air, m is the air viscosity, and d is a

characteristic size of a vortex structure formed by the air stream. For values of

St < 1, droplets will be dispersed with the surrounding stream while for large values of St insufficient time

will be available to influence the droplet motion. Thus, entrainment will be optimized for smaller droplets

that have little relative velocity to the airflow.

SPGG Results

For the SPGG experiments, the nominal velocity of the air above the backward-facing step was

maintained constant at 5.4 m/s; the propane flow was 85 mL/s (at standard temperature and pressure). 

The pressure build-up in the discharge chamber was measured for a range of bypass areas, as seen in Fig.

20. The high pressures produced in the SPGG discharge chamber and the known area of the metering

orifice allow the mass flow of agent added to the air stream of the TARPF to be estimated by assuming

that the flow through the orifice is choked. Figure 21 shows the repeatability of the agent injection process

with five overlaying mass flow and thermocouple temperature traces. (The thermocouple does not reflect

the true gas temperature, which is expected to be hundreds of degrees Celsius hotter than recorded in the

figure.)
30

The SPGG discharge time was consistently 20 ms ± 1 ms, which is over three times faster than

the shortest N2 or CF3Br injection interval, and not much affected by the bypass port area. The bypass

ratio (Ain/Atot) is the area of the inlet metering orifice (port 2) divided by the total open area available for

flow to exit the discharge chamber (ports 2, 3, and 4) The time interval, Dt, is shown in Fig. 22 as a

function of the estimated discharge mass. The total mass delivered to the air stream during the discharge

process is found by integrating dm/dt over Dt. Excluding the highest and lowest area ratios, the estimated

mass delivered can be seen in Fig. 22 to be linearly proportional to the area ratio; however, almost 50 %

more mass is estimated than one would expect. The dashed line in Fig. 22 indicates that a 1-to-1 relation

would exist if the mass were directly proportional to the area ratio.

There are several factors that contribute to an uncertainty in the estimate of the absolute mass of

agent. First, uncertainties in the gas composition and temperature upstream of the metering orifice affect

the estimate since the mass is proportional to the square root of the molecular weight divided by the

temperature. A factor of two under-estimate in this parameter would cause a 40 % over-estimate in

mass, which, if corrected for, would cause the data plotted in Fig. 22 to more closely align with the dotted

line. A second source of uncertainty is the complexity of the flow within the discharge chamber created

by the jet emanating from the SPGG. The calculation assumes that the upstream flow is steady and

parallel to the axis of the metering orifice, but the flow is highly transient and more perpendicular.

Suppression of the propane pool fire with the hybrid gas-generators was found to be successful if

at least 1.5 g of agent was injected into the fire; conversely, extinction never occurred when less than 0.7

g was added. The percent of the fires suppressed varied when the agent mass was between these limits,

as shown in Fig. 23. The suppression statistics were generated by lumping the mass from thirty-three

discharges into bins 0.2 g wide, centered about the data plotted. The solid line is a fit to the data assuming

that the shape is sigmoidal. It is apparent from the curve that there is a 50 % chance that suppression will

be successful if the amount of agent is 0.9 g ± 0.1 g, and there is a 90 % success rate for 1.3 g ± 0.1 g of

agent.

Figure 24 is a plot of the mass of agent required for suppression versus the injection time interval.

 All of the SPGG data are lumped around 20 ms since the injection time was fixed. The N2 and CF3Br

results extend to much longer injection time intervals. Linear fits to the data yield intercepts of 1.6 g for

nitrogen and 1.2 g for halon 1301. The SPGG data fall close to the halon results. The significance of the

linear shape and the value of the intercept is unclear; however, the superior performance of the gas

generator is undeniable.

The mass fraction of agent, β, is defined as the total mass injected divided by the injection

duration, Dt, over the sum of the mass flow of air plus the mass flow of agent. The percentage of

occasions that the flame was extinguished is plotted in Fig. 25 as a function of β. All of the fires were

extinguished when β was greater than 0.62; none for a mass fraction below 0.49. The definition of β for

the SPGG differs somewhat from the definition as applied to the gaseous agents because the mass

discharge profiles for SPGG and the gaseous agents are different. Whereas the mass injection rate of the

gaseous agents is controllable through selection of the injection hardware, the SPGG mass injection rate is

practically dependent on the propellant effluent generation rate (i.e., the propellant burning rate). The

character of the mass injection rate, dm/dt, is qualitatively different than that of the gaseous agents. Its

value for the SPGG is best represented by the pressure profile shown in Fig. 20, as compared to dm/dt

seen in Fig. 5 for gaseous N2. For the same mass, variation of the discharge profile will lead to variation in

agent effectiveness. Optimization of the rate of agent discharge is an area that would benefit from further

study, which could be approached from analytic, numerical, and experimental perspectives.
31

TIME, s

-0.02 0.00 0.02 0.04 0.06 0.08 0.10

PRESSURE, kPa-a

0

500

1000

1500

2000

2500

3000

1.27x10-3 m2

0.51x10-3 m2

0.79x10-3 m2

1.55x10-3 m2

Abypass=1.42x10-3 m2

TIME, s

-0.02 0.00 0.02 0.04 0.06 0.08 0.10

TEMPERATURE, K

280

290

300

310

320

330

340

350

360

dm/dt, kg/s

0.00

0.05

0.10

0.15

T

dm/dt

Figure 21. Discharge chamber thermocouple temperature and estimated

mass flow of SPGG efluent added to air stream for multiple identical runs.

Figure 20. Chamber pressure created by SPGG discharge as a function of bypass area.
32

(Ain/Atotal)mtot, g

0.00 0.25 0.50 0.75 1.00 1.25

MASS DELIVERED, g

0.0

0.5

1.0

1.5

2.0

INJECTION INTERVAL, ms

0

20

40

60

80

100

0.303+1.368(Ain/Atot)mtot

MASS, g

0.50 0.75 1.00 1.25 1.50 1.75 2.00

% SUPPRESSED

0

20

40

60

80

100

Figure 6. Injection interval (triangles) and calculated mass delivered to flame (circles) 

as a function of area ratio times total mass of gas generated (20.7 g).

Figure 7. Percentage of flames extinguished as a function of mass delivered to flame by

SPGG.

Figure 23. Percent of flames extinguished as a function of mass delivered to flame by SPGG.

Figure 22. Injection interval (triangles) and calculated mass delivered to flame (circles) as function of area

ratio times total mass of gas generated (20.7 g).
33

INJECTION TIME, ms

0 100 200 300 400

MASS, g

0

2

4

6

8

10

CF3Br

N2

SPGG

Figure 24. Impact of injection time interval on the mass of agent (CF3Br, N2, or SPGG)

required to suppress step-stabilized propane pool fire.

Effect of Hot Surface

Figure 2 shows the location of the inverted-vee obstacle placed downstream of the propane pool. An

ethane-fueled torch was used to preheat the upstream surface of the obstacle to produce temperatures in

excess of 1000 oC as measured by a type-S thermocouple mounted on the exterior face. Pyrometer

readings of the interior (i.e., fire-side) surface indicated temperatures were about 50 oC cooler. 

Experiments were conducted to determine the minimum temperature needed to ignite the propane pool

when the air and propane were set to the low flow conditions. Ignition occurred when the hot surface

reached temperatures close to 1000 oC. 

A spray nozzle was located upstream of the stabilizing step that allowed liquid JP-8 to be sprayed

over the propane pool and onto the heated inverted-vee surface. The fuel acted to cool the surface

approximately 150 oC below the temperature that had been obtained when only the gaseous

propane fuel was flowing. The JP-8 was ignited and formed a flame that stabilized behind the vee when

the temperature of the surface exceeded 900 oC.

The procedure used to test for re-ignition of the propane fire was as follows. The propane fire

was initiated under the operating conditions of an air flow velocity of 1.5 m/s and a propane flow rate of

33 mL/s, and the obstruction was heated to the temperature desired. The flame was observed to extend

past the obstruction. The agent was introduced to the system for three different injection times of 75 ms,

100 ms, and 150 ms, and the occurrence of flame suppression was observed. The cylinder pressure was

increased if suppression did not occur, and then the process repeated. The agents tested were nitrogen,

HFC125, and CF3Br. The agent injection time was varied to observed the effect on re-ignition of the

propane fuel. The JP-8 stream was then directed against the heated obstruction while the unignited

propane still was flowing into the wind tunnel passage. If re-ignition did not occur after 10 s, then the

propane flow was terminated. Experiments were repeated with and without the JP-8 stream in order to

determine the efficacy of the stream in the re-ignition process. A variety of re-ignition results were
34

observed which depended on the injection time interval of the agent and JP-8 stream. Re-ignition was

found to occur when the obstruction was above 900 oC (bright red), and a function of the cooling effect by

the agent and JP-8 liquid stream.

The influence of the surface temperature was found to be bi-modal: when the temperature of the

hot surface (with the JP-8 and propane flowing) was below about 880 oC ± 10 oC, the amount of agent

necessary for suppression was the same as when the surface was unheated; and when the temperature

was above 890 oC ± 10 oC, reignition always occurred within about 10 s independent of higher

temperatures or the amount of N2 added. A similar finding resulted when nitrogen was replaced by

CF3Br, although the measured dividing temperature was about 50 oC higher. This difference may have

been associated with a build-up of carbon on the hot surface that insulated the fire-side surface from the

ethane torch, making an accurate determination of the surface temperature more difficult.

A few experiments were conducted with the temperature reduced to the Leidenfrost point (290

oC for JP-8) to determine if an increase in contact time would compensate for the lower temperature. 

Ignition of the JP-8 and propane did not occur. Satcunanathan and Zaczek [27] measured the ignition

delay of kerosene and Diesel fuels on a hot metal surface and found that ignition could occur in less than

one second for a 2 mm diameter droplet, but that the ignition time increased for temperatures between

about 500 oC and 550 oC due to surface boiling. No ignition of the JP-8 droplets was observed in the

TARPF when the vee-surface temperatures was maintained in the range between 360 oC and 440 oC. 

However, Jomaas et al. [14] found that ignition of a JP-8 pool fire would occur in their step-stabilized

burner for temperatures in the same range. (See Appendix B.)

Data Correlation

The effectiveness of the gaseous agents is compared to that of the SPGG through use of a simple agent

mixing model to describe the suppression phenomena. A more detailed description of the model can be

found in Grosshandler et al. [28] and Hamins et al. [9]. A characteristic time, τ, for mixing of the agent

into the flame zone can be defined in terms of the air and agent volume flows, (V'air+ V'agent), and the step

height, h, as

τ = γh /{(V'air+ V'agent)/[(L-h)L]} (20)

where L is the width of the tunnel and γ is an empirical non-dimensional parameter that relates the ratio of

the distance that a fluid element travels within the recirculation zone to the obstacle height. Takahashi et

al. [10] measured the characteristic mixing time in a similar facility and found γ to be around 40. 

Evaluating Eq. (20) for the range of flows and baffle heights examined in the current study and using a

value of 40 for γ, τ is found to vary between 0.04 s and 0.40 s.

Hamins et al. [9] found that for a specified injection duration it is possible to relate the mole

fraction of agent required to achieve extinction, C, to the characteristic mixing time, τ, according to the

following relation:

C/C* = [1-exp(-Dt/τ)]-1 (21)

where C* can be found experimentally by flowing agent continuously into the air stream at increasing

rates until extinction occurs. If the air flow is low enough, the value of C* is expected to be similar to

the cup burner extinction requirements [9]. For propane in a cup burner, Trees et al. [5] found the value

of X* to be 0.32 for N2, 0.41 for Ar, and 0.039 for CF3Br. Others have found similar results [29]. Figure

26 compares the suppression results for N2, CF3Br, and SPGG through use of Eq. (21), where the

normalized mole fraction is plotted as a function of the non-dimensional injection interval. For the SPGG

results, C

* is assumed to equal the C* value for argon. While the data do not fall exactly on the model, the

trend of the results are well represented by the single curve when one considers run-to-run variations due

to the statistics associated with suppression of a turbulent flame. The results show that the
35

Figure 25. Percent of flames extinguished as a function of the estimated mass fraction of agent.

Figure 26. Normalized mole fraction as a function of non-dimensional injection interval, comparing N2,

CF3Br, and SPGG

b

0.4 0.5 0.6 0.7 0.8

% SUPPRESSED

0

20

40

60

80

100

Δt/τ

0 2 4 6 8

X/X

*

0

1

2

3

4

SPGG

N2

CF3Br

X/X∗= [1-exp(-Δt/τ)]-1
36

Figure 27. Suppression mole fraction of agent (N2 or CF3Br) normalized by cup burner values (Χ*) as

function of injection time interval normalized by characteristic residence time (τ). Gray symbols:

experimental results keyed to Fig. 19; black diamonds: direct numerical simulations with N2.

effectiveness of the SPGG hybrid significantly exceeds the model predictions.

For laminar diffusion flames strained at intermediate rates, Trees et al. [5] showed that the

minimum extinction mole fraction of agent in a counterflow flame decreases from the cup burner value

when the strain rate is 50 s-1 to much smaller values for a strain rate of 400 s-1. Although the flow in the

recirculating region behind a step is much more complicated than in a counterflow flame, the strain rate in

the current study should scale with 1/τ. When the flow of air is increased sufficiently, the flame becomes

strained to the point that agent is not needed for extinguishment (i.e., C

* → 0) and the flame blows out. 

In other words, as the air flow in the TARPF increases, the model represented by Eq. (21) in Fig. 26 must

be adjusted not only for changes in τ, but also for changes in the value of C*.

A number of direct numerical simulations of N2 suppressing the step-stabilized flame were

conducted assuming a finite reaction rate with one step chemistry. The injection interval and mole fraction

Δt/τ

0 1 2 3 4 5

X/X*

0

1

2

3

4

5

X/X*= [1-exp(-Δt/τ)]-1

simulation, suppression

simulation, no suppression
37

of nitrogen were varied, and the flame structure was followed through the transient period. Either the

flame was extinguished, or it was reestablished following the passage of the N2 pulse. Figure 27

summarizes the results. Open diamonds indicate that no suppression occurred; solid diamonds imply that

suppression was successful. The gray symbols refer to the experimental N2 and CF3Br data for all

geometries examined. The dotted line is Eq. (21). 

The computations are able to distinguish regimes of extinction and non-extinction in the case of

nitrogen as an agent. The numerical model cannot at present predict what will happen when a chemically

active agent like CF3Br is introduced into the flame. The cooling and dilution of the flame by the agent can

be predicted, and hopefully, simplified combustion mechanisms for various chemically-active agents can be

developed that will lead to a better understanding of the dynamics of fire suppression.
38

CONCLUSIONS AND RECOMMENDATIONS

A transient application, recirculating pool fire (TARPF) facility has been built for screening the

suppression effectiveness of halon 1301 replacements. The ability to measure the relative effectiveness of

alternative agents is key to the development of new fire suppression systems. The physical and chemical

properties, and the manner of storage and release of the next generation suppression systems may be quite

unlike CF3Br, but their effectiveness must still be bench-marked against it. The TARPF facility provides

the means to screen gaseous agents and solid propellant gas generator concepts in the laboratory for

applications in protected spaces involving baffle-stabilized pool fires.

This report represents a comprehensive study, utilizing experimentation, advanced numerical

modeling, and analysis to develop and characterize a well-controlled fire suppression facility. Sample

experimental results are shown that demonstrate the utility of the facility. Detailed fabrication drawings of

the facility are provided in the Appendix of this report. In addition, the experimental procedures are

described in detail in the text. Nominal air velocities between 2 m/s and 23 m/s flowing over a backwardfacing step were examined. Because the air is metered with a sonic orifice, the injection of agent does not

modulate the air flow. The minimum amount of agent for flame extinguishment is substantially and directly

affected by the air velocity and the interval of injection. A simple mixing model is useful to explain the

observed trend of decreasing suppressant mole fraction with increasing injection duration. A detailed

numerical model of the suppression of baffle stabilized fires was developed. The model was successful in

simulating the suppression event including a quantitative determination of agent mass requirements.

The facility provides the capability to test solid-propellant gas generators allowing direct

comparison of compressed and solid-propellant generated gases for the first time. The capability to test

SPGGs under well-controlled laboratory conditions allows evaluation among different propellant

formulations, particulate yields, and burning rates for various SPGG designs. Results showed that the

effectiveness of the hybrid SPGG exceeded that of CF3Br and that its effectiveness was significantly

greater than expected from the model prediction. Further research could lead to significant improvement

of the SPGG as well as traditional agent through enhanced suppressant system design.

Several other research areas require further investigation. These include the effect of the air flow

on the steady-state extinction mole fraction of agent, the relationship between agent injection and its

concentration history at the flame, and especially the observed differences in the normalized mole fractions

of CF3Br, N2, and the SPGG for very short injection time intervals.
39

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