Anthony Hamins

Kevin McGrattan

Glenn Forney

Unwanted Accelerated Burning After 

Suppressant Delivery

Sponsored by: 

The Department of Defense Strategic Environmental Research and Development 

Program

The views and conclusions contained in this document are those of the authors and should not 

be interpreted as representing the official policies, either expressed or implied, of the Strategic 

Environmental Research and Development Program, NIST, or any other part of the U.S. 

Government.

i
NIST Special Publication 1004 

Anthony Hamins

Kevin McGrattan

Glenn Forney

Unwanted Accelerated Burning After 

Suppressant Delivery

ii
NIST Special Publication 1004 

Anthony Hamins

Kevin McGrattan

Glenn Forney

Building and Fire Research Laboratory

Unwanted Accelerated Burning After 

Suppressant Delivery

September 2003

iii

National Institute of Standards and Technology

Arden L. Bement, Jr., Director

Technology Administration

Phillip J. Bond, Under Secretary for Technology

U.S. Department of Commerce

Donald L. Evans, Secretary
Certain commercial entities, equipment, or materials may be identified in this 

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

identification is not intended to imply recommendation or endorsement by the 

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

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

National Institute of Standards and Technology Special Publication 1004

Natl. Inst. Stand. Technol. Spec. Publ. 1004, 48 pages (September 2003)

CODEN: NSPUE2

For sale by the Superintendent of Documents, U.S. Government Printing Office

Internet: bookstore.gpo.gov — Phone: (202) 512-1800 — Fax: (202) 512-2250

Mail: Stop SSOP, Washington, DC 20402-0001

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 2003

_________________________________________

iv
Executive Summary 

In a number of real-scale fire tests, accelerated or enhanced unwanted burning has 

been observed immediately following the delivery of a fire suppressant. The conditions 

under which this phenomena occurs, its potential for harm, and its effect on the ability of 

an agent to extinguish a fire have not been comprehensively documented. The objective 

of this work is to identify the causes of enhanced burning with an emphasis on aircraft 

dry bay and engine nacelle applications and to provide recommendations on how it can 

be avoided when it may become a significant safety risk. 

A survey of the fire and combustion literature was undertaken to document cases 

of enhanced burning after suppressant delivery. Informal interviews were conducted with 

a sample of scientists and engineers involved with full and reduced-scale fire suppression 

experiments. Analysis of the information suggests that there is no common terminology 

or unique definition used to describe unwanted enhanced burning. Several instances of 

enhanced burning have been documented and it is apparent that unwanted enhanced 

burning is not uncommon. Manifestation of the phenomena has been documented through 

measurements of increased pressure, temperature, and heat release rate, as well as 

changes in the appearance (intensity, color, volume) of a fire for a number of 

experimental configurations. Although many events could be described as enhanced 

unwanted burning after suppressant delivery, the main hazard associated with this 

phenomena appears to be system over-pressurization due to enhanced combustion from a 

vaporizing liquid fuel that mixes with air after suppressant delivery. In many instances, 

the phenomena do not represent a significant safety risk. 

To further understand the phenomena, three experiments in which system overpressurization was suspected were analyzed in some detail. The cases involved overpressurization of full-scale suppression experiments for an aircraft dry bay and two cases 

of over-pressurization of full-scale enclosure fire experiments. The three cases had many 

differences in terms of experimental conditions, time scales of the suppression events, 

and the geometric configuration of the experiments, yet in all three cases volatile liquid 

fuels were present. When wood, a solid-phase fuel, was substituted for liquid fuels in the 

suppression experiments in an enclosure (Case 2), large pressure fluctuations were not 

v
observed. The case studies were analyzed using differing approaches, depending on the 

details of the scenario. 

The first case-study examined was that of enclosure fires conducted at the Naval 

Research Laboratory (NRL) in which it was documented that suppressant discharge can 

cause significant fire flare-up and pressure increases for unconfined volatile liquid fuels 

in the presence of a hot ignition source. Deflagration was observed for HFC agents, 

water mist, and hybrid agents involving HFC/powder mixtures. The results showed that 

suppressant delivery caused an intensification of the fire as defined by an increase in the 

flame volume, suggesting that (premixed) fuel-air mixtures can be created in a postsuppression environment under certain conditions. Interestingly, all deflagrations 

observed were in the presence of a hot ignition source (at ≈550 oC). An analysis of the 

detailed experimental data provided by Sheinson and coworkers suggest that unwanted 

burning effects occurred when there was insufficient suppressant to assure suppression 

and prevention of re-ignition. In this sense, unwanted accelerated burning effects are 

already embedded in the design equations resulting from full-scale test results. 

The second case-study involved the unanticipated catastrophic over-pressurization 

that destroyed a C-130 Wing Leading Edge Dry Bay Test Facility on August 28, 1998 

during real-scale fire suppression testing. Six tests previous to the catastrophic event 

using N2-pressurized HFC-125 had led to routine suppression results. The seventh test 

was designed to test the impact of the N2 used for pressurization on the fire. An Air Force 

investigation of the event concluded that “increased level of fluid ignition (combustion)” 

was the source of over-pressure. Experimental peculiarities were noted including the use 

of a suppressant delivery line that was air-filled. 

The fire scenario was extremely complex due to the presence of a ballistic round, 

accompanying shock waves, fuel splashing (two-phase flow), multiple ignition sources 

(hot shrapnel), and so on. It should be noted that according to measurements made during 

the experiment, failure of the C130 dry bay fixture occurred when the pressure inside of 

the apparatus reached 54 kPa (7.8 psig). The apparatus had survived earlier experiments 

when the pressure reached 30 kPa (4.3 psig). A plausible explanation for the overpressurization is that delivery of the agent in this case caused enhanced mixing of unburnt 

fuel and air. The experiment used an inert suppressant (N2), but at a volume fraction that 

vi
was 0.4 % of the cup burner value, an amount that would have little impact on the 

stability of the fire and certainly would not extinguish the fire. 

To better understand the conditions in the dry bay during fire suppression, an 

analytic approach was taken to simulate the pressure in the enclosure. The time 

associated with equilibration of a pressure wave within the dry bay was estimated to be 

less than 1 ms, whereas the timeline associated with the spray fire was on the order of 

≈100 ms, which allowed application of the incompressible form of the conservation 

equations. 

 The third case-study examined pressure fluctuations observed by Kashiwagi et al. 

[2000] during suppression experiments on moderate-sized (≈ 400 kW) heptane pool-fires 

in an enclosure. The experiments were simulated using the NIST Fire Dynamics 

Simulator [McGrattan et al., 2000] with a one-step chemical reaction. Numerical 

prediction of the time dependent enclosure pressure and the amount of agent required to 

suppress the fire yielded predictions that were very sensitive to the Arrhenius rate 

parameters. Using Arrhenius rate parameters that correspond to measurements in smallscale suppression experiments failed to yield pressure fluctuations similar to those 

observed by Kashiwagi et al. [2000]. 

In summary, the main hazard associated with unwanted burning after suppressant 

delivery appears to be over-pressurization likely due to enhanced mixing of vapor with 

air. From the investigation of three key case studies, the most important overall finding is 

that suppressant delivery does not appear to pose undue safety risk unless insufficient 

amounts of suppressant are applied to a fire. It is concluded that these effects are already 

embedded in full-scale test results and associated design equations. This report concludes 

with a discussion of safety issues associated with unwanted enhanced burning after 

suppressant delivery. 

.

vii
Table of Contents

Page 

Executive Summary.............................................................................. v 

1. Technical Problem .............................................................................. 1 

2. Objectives ........................................................................................ 1 

3. General Methodology ........................................................................... 2 

4. Background ...................................................................................... 4 

4.1 Interview Findings: Full-Scale Suppression Testing .................................. 4 

4.1.1 Large Pressure Fluctuations & Over-Pressurization ........................... 6 

4.1.2 Flare-ups ............................................................................ 8 

4.2 Literature Review & Interview Findings: Bench-Top Experimentation .......... 11 

4.2.1 Splashing ........................................................................... 11 

4.2.2 Flame Straining .................................................................... 12 

4.2.3 Mixing .............................................................................. 16 

4.2.4 Heat Transfer and Chemistry-Related Effects ................................ 17 

4.3 Assessment of Background Information ............................................... 17 

5. Case Studies ..................................................................................... 18 

5.1 Case 1: Analysis of Over-Pressurization during Suppression in the 

 NRL Moderate-Scale Enclosure .............................................. 18 

5.2 Case 2: Modeling of Global Properties during Fire Suppression in a Dry bay ... 24 

 5.2.1 Background ........................................................................ 24

 5.2.2 Analytic Approach ................................................................ 27 

 5.2.3 Summary ........................................................................... 37 

 5.3 Case 3: Pressure Fluctuations During Suppression in an Enclosure .............. 38 

 5.4 Accelerated Burning in an Engine Nacelle ........................................... 41 

6. Summary, Conclusions, and Recommendations for Future Research .................... 42 

7. Acknowledgements .............................................................................. 44 

8. References ........................................................................................ 45 

viii
1. Technical Problem

Over the last decade, there has been a world-wide research effort to find long and 

short term alternatives to ozone depleting halon fire suppressants. This has led to testing 

of candidate agents in reduced and full-scale fire suppression testing. A key tenet of fullscale testing is to maximize fire suppression efficiency while minimizing the amount of 

stored agent necessary for complete suppression. This optimization demands a better 

understanding of the behavior of fires during the suppression process. One part of the 

suppression process that is not well understood is the “enhanced burning”, which has 

been observed just after an extinguishing agent has been applied to a fire. 

In a number of real-scale fire tests, accelerated or enhanced unwanted burning has 

been observed immediately following the delivery of a fire suppressant. The phenomena 

may manifest itself in a variety of forms including a fire flare-up or a sudden enclosure 

over-pressure. The phenomena is associated with the delivery of a suppressant and would 

not occur without a suppressant discharge. 

A literature review reveals that there have been few published descriptions of 

enhanced burning during suppressant deployment. Reports of the phenomena have 

largely been anecdotal rather than scientifically rigorous. Little analysis has been 

conducted considering the implications of the observations. 

The conditions under which accelerated burning occurs, its potential for harm, 

and its effect on the ability of an agent to extinguish a fire are currently unknown. While 

this behavior could pose a significant safety risk, there has been no systematic study 

investigating the conditions necessary for the existence of this phenomenon. Enhanced 

burning could potentially pose a safety risk if it causes structural damage, strengthens a 

fire enough such that it survives extinguishment, or increases the likelihood of unwanted 

re-ignition or fire spread. 

2. Objectives 

The objective of this work was to identify the causes of enhanced burning after 

extinguishing agents are applied to fires with an emphasis on aircraft dry bay and engine 

nacelle applications, and to provide recommendations on how enhanced burning can be 

avoided when it becomes a significant safety risk. 

1
3. General Methodology

The technical approach and methodology used in this research are shown 

schematically in Figure 1. The approach consisted of a literature review, an informal 

survey of fire suppression practitioners, and detailed analysis of the phenomena when it 

appeared to pose a significant safety risk. Because the number of documented cases 

where the phenomenon was identified or suspected was small, the study focused on an 

analysis and investigation of case studies identified by practitioners of fire suppression 

testing and research. 

 To further understand the phenomena, calculations were undertaken when 

appropriate. In one case, these focused on simple bracketing estimates of pressure 

increase in an enclosure due to suppressant release and enhanced combustion. A second 

case was suited for computational fluid dynamic (CFD) fire modeling. An attempt to do 

this was unsuccessful due to the complexity of the scenario and the current limitations of 

CFD fire models. This report concludes with a discussion of safety issues associated with 

unwanted enhanced burning after suppressant delivery. 

The approach was developed based on the fact that little is known about the 

phenomena. The approach consciously does not consider characterization of the 

splashing of a fuel by a high momentum agent stream, which was the approach initially 

proposed as part of this project. This does not imply that splashing is not particularly 

dangerous. A splashing fuel spray can enlarge the surface area of a burning material, 

leading to a condition that is potentially difficult to extinguish. Subsequent analysis 

suggested that while a study of splashing may lead to interesting new information, such 

information would not aid in the analysis of unwanted accelerated burning. In addition, it 

is a given that splashing of a fuel pool will occur when impinged by a high momentum 

agent stream. It was decided that detailed investigation of the splashing phenomena 

would not lead to insight into how to prevent unwanted burning after suppressant 

delivery. The solution to the splashing problem is self-evident and involves directing the 

high momentum agent discharge away from possible fuel sources. 

 The effort discussed in this report instead focused on other issues associated with 

the unwanted accelerated burning phenomena, namely, over-pressurization of the 

2
enclosure by ignition of a combustible fuel vapor-air mixture. The worst case would be 

the case of a uniform stoichiometric mixture throughout the enclosure. Ignited at one end 

of the enclosure, it would propagate at its turbulent flame speed (O[10 m/s]), heating and 

expanding until it reached the other end of the nacelle. 

The remainder of Section 1 of this report contains a description of the phenomena, 

a literature review, and the results of an informal survey of a number of scientists and 

engineers conducting full-scale fire suppression research and testing. Section 2 contains 

an analysis of three case studies including two enclosure fires and a fire in a dry bay. 

Section 3 contains summaries of the findings and makes conclusions regarding avoidance 

of enhanced burning after suppressant delivery. 

identify phenomena 

safety 

risk?

low high 

postulate mechanism(s)

detailed analysis: 

case studies 

recommendations

evaluate data

survey & literature review

Prioritize

Figure 1. Flowchart outlining the research approach taken in this study. 

3
4. Background 

Several instances of enhanced burning have been documented in the literature and 

it is apparent that unwanted enhanced burning after suppressant delivery is not 

uncommon. Yet, analysis of information available in the literature indicates that there is

no unique definition to describe unwanted enhanced burning. In general, the phenomena 

are not well understood or recognized. Because of this situation, there is no common 

terminology to describe the phenomena, which hampers progress in this area. 

A review of the literature and discussion with other researchers indicates that 

evidence of the phenomena of unwanted burning after suppressant delivery is typically 

anecdotal, rather than scientific in nature. Manifestation of the phenomena has been 

documented through measurements of increased pressure, temperature, or heat release 

rate after suppressant delivery, as well as observed changes in fire appearance (intensity, 

color, volume). The phenomena have not been limited to one particular suppressant type 

or configuration. Observations of the phenomena have been made in a number of very 

different configurations including full and reduced-scale experiments on aircraft dry bays 

and enclosure fires. To date, there has been no comprehensive study of this phenomenon 

or its possible ramifications in terms of fire suppression and life safety. 

A survey of the fire and combustion literature was undertaken to document cases 

of enhanced burning after suppressant delivery. In addition, a number of interviews were 

conducted with researchers involved in full and reduced-scale fire suppression 

experiments. 

4.1 Interview Findings: Full-Scale Suppression Testing

 Interviews were conducted with those directly involved with full-scale 

suppression testing on engine nacelle, dry bay aircraft fire safety, and vented and 

unventilated enclosure fires. Interviews were conducted during the years 2001 and 2002 

with personnel performing experiments at Wright Patterson Air Force Base (WPAFB) 

[Bennett, 2002; Tucker, 2002; Cyphers, 2002], the Naval Research Laboratory (NRL) 

[Sheinson, 2002; Maringhides, 2002], NAVAIR at Lakehurst [Tedesci, 2002; Wolf, 

2002], the Research Institute of Fire and Disaster of Japan (RIFD) [Saso, 2002], the 

Federal Aviation Administration (FAA) [Ingerson, 2002; Rhinehart, 2002], and the 

4
Aircraft Survivability Division, U.S. Navy Facility at China Lake, CA (China Lake) 

[Manchor, 2002] as listed in Table 1. 

 The interviews and literature search provided information on the multiplicity of 

scenarios, conditions, configurations, and suppressant types that are relevant to the 

discussion of unwanted burning. A number of observations are noted below that appear 

to be related to unwanted enhanced burning or combustion after suppressant delivery. 

Almost all investigators involved in full and reduced-scale suppression experiments have 

observed “flare-ups” of a fire after suppressant delivery. In a “flare-up”, the luminosity 

and volume of the fire appear to rapidly increase immediately after suppressant 

deployment as additiont of the suppressant causes displacement of flames, fuel vapors or 

hot combustion products. Several investigators involved in full and reduced-scale 

suppression experiments observed over-pressurization of the test enclosure. 

Table 1. Interviews: full-scale suppression testing 

Configuration 

Nacelle Dry Bay Enclosure 

Bennett, Tucker (WPAFB)1 Cyphers (Skyward)2 b Sheinson, Maranghides 

(NRL) 3 b

Ingerson (FAA) 4 Manchor (China Lake) 5 Saso (RIFD) 6 b

Tedeschi (NAVAIR 7

Lakehurst) 

Wolf (NAVAIR Lakehurst)7 Rhinehart (FAA) 4

b Indicates that over-pressurization of enclosure after suppressant delivery observed. 

1. Wright Patterson Air Force Base, Dayton, Ohio. 

2. Skyward, Ltd., Dayton, Ohio. 

3. Naval Research Lab, Washington, D.C. 

4. Federal Aviation Administration, William J. Hughes Technical Center, Atlantic City, 

New Jersey. 

5. Naval Ordnance Test Station, China Lake, California. 

6. Research Institute of Fire and Disaster, Mitaka, Japan.

7. Naval Air Systems Command, Lakehurst, New Jersey. 

5
4.1.1 Observation of Large Pressure Fluctuations and Over-Pressurization

Of the several observations detailed by those performing full-scale fire 

suppression testing, one particular type of event appeared to be particularly hazardous. 

Several investigators involved in full and reduced-scale suppression experiments 

observed over-pressurization of the test enclosure (see notes in Table 1). This situation 

was observed in unventilated and slowly ventilated enclosure fires, as well as a Dry bay 

Test Facility. The phenomena are described in more detail below. Whereas there were a 

number of commonalities between these three cases, there were many more differences, 

for example, in ventilation and clutter conditions, suppression timescales (Δt), and 

configuration type, as seen in Table 2. The most hazardous pressure excursions occurred 

for the configurations with a re-ignition source present. The one commonality of note 

was that volatile liquid fuels were present in all cases. When wood replaced heptane in 

Configuration 3, pressure fluctuations were not observed. 

Table 2. Configurations in which post-suppression over-pressurization or pressure 

fluctuations were observed. 

Configuration Re-ignition 

Source 

Ventilation Clutter Δt (s) 

suppression 

Fuel 

1. Enclosure 

(NRL) 

Yes None High ≈10 MethanolHeptane 

2. Dry bay 

(C130 WPAFB) 

Yes Projectile 

damage 

Low ≈0.1 JP8 

3. Enclosure 

(RIFD) 

No Mechanical None ≈10 Heptane* 

* large pressure fluctuations not observed when wood cribs were used as fuel. 

Over-pressurization was observed and reported in the literature by 

experimentalists investigating fire suppression in ventilated and unventilated enclosure 

fires. To further understand the most hazardous phenomena, analyses were carried out on 

three experiments in which system over-pressurization was suspected. The cases 

6
involved over-pressurization of full-scale suppression experiments for an aircraft dry bay 

and two cases of over-pressurization of full-scale enclosure fire experiments. The three 

cases had many differences in terms of experimental conditions, time scales of the 

suppression events, and geometric configuration of the experiments, yet in all three cases 

volatile liquid fuels were present. When wood, a solid-phase fuel, was substituted for the 

liquid fuel in the suppression experiments in an enclosure (Case 3), the amplitude of the 

pressure fluctuations were reduced. The case studies were analyzed using approaches 

tailored to each incident, depending on the details of the scenario. A brief synopsis of the 

three cases follows with a more detailed analysis given below. 

Case 1: Analysis of Over-Pressurization during Suppression in the NRL Moderate-Scale 

Enclosure

Sheinson and coworkers of NRL have shown that a suppressant discharge in a 

moderate-scale enclosure (28 m3) can cause significant fire flare-up and lead to rapid 

pressure increases for unconfined volatile liquid fuels in the presence of a hot ignition 

source [Maranghides and Sheinson, 2000]. Some aspects of the experiments are 

described in Table 2 (above). In their experiments, repeatable rapid over-pressurization 

was often observed after suppressant delivery, which was attributed to a 

combustion/deflagration wave. A variety of suppressant types were used including 

gaseous, compressed liquid, and hybrid suppressants composed of liquid-phase plus dry 

powder. Through the analysis of video records, the experiments documented that 

suppressant delivery could cause intensification of the fire and an increase in the burning 

area. Because an ignition source was present throughout the experiments, it was unclear 

whether the deflagration initiated due to a failure to fully suppress the initial fire or as a 

re-ignition after complete suppression. 

Case 2: Over-Pressurization during Fire Suppression Testing in the C130 Dry Bay 

Facility at WPAFB 

In a replica of a C130H WLE Dry bay test article, unanticipated catastrophic 

failure occurred after suppressant delivery failed to extinguish an experimental fire. An 

incident investigation suggested that the release of agent (N2 in this case) may have acted 

7
to mechanically agitate the air, contributing to a more vigorous and rapid combustion 

event. In the post-suppression environment of an enclosure fire, deflagration and overpressurization was attributed to enhanced mixing of unburned vapor/air mixtures. 

Case 3: Total flooding suppression of a pool fire in an enclosure at RIFD, Japan

A series of total flooding suppression experiments were performed using heptane 

pools fire in an enclosure at RIFD, Japan [Kashiwagi, 2001]. The suppressants tested 

included gaseous inert mixtures as well as compressed halogenated liquids. A door 

(closed before suppressant delivery) and a small vent near the ceiling provided 

ventilation. Pressure fluctuations were observed in the enclosure after suppressant 

delivery. No controlled ignition source was present in the enclosure, although the hot 

metal burner or fire-heated thermocouples could have acted as an ignition source. The 

reasons for the pressure fluctuations are not clear. It is possible, however, that the 

fluctuations were due to dynamics of the suppressant flow shielding air from fuel. 

Kashiwagi et al. [2001] highlight the importance of the pressure fluctuations, stating that, 

“agent discharge enhances flame intensity.... over-pressurization during suppression is 

caused not only by the agent discharge but also by the expansion of gases due to the 

combustion... application of the total-flooding systems to volatile liquid fuel fires has 

potential risk of compartment destruction due to the rapid mixing of unburned vapor with 

air.” 

4.1.2 Flare-ups

While a high momentum suppressant discharge in an engine nacelle or dry bay 

may or may not extinguish a fire, the discharge can significantly alter the flowfield. In 

such a case, the suppressant flow may cause displacement of flames, fuel vapors or hot 

combustion products associated with the fire. This may include smoke, CO2, and water 

vapor. A dramatic example of such a situation is shown in Fig. 2, which illustrates a 

typical flare-up observed during full-scale V-22 fire suppression experiments at the China 

Lake test facility. Flare-ups were also been observed immediately after suppressant 

delivery in reduced-scale experiments such as a turbulent spray burner in a wind tunnel 

8
[Hamins et al., 1994] and the Wright Patterson Engine Nacelle Test Facility [Tucker, 

2002]. 

In the China Lake experiments, a solid propellant gas generator was used to 

extinguish a fire in the wheel bay area, which is to the right in the images shown in 

Fig. 2. The images in Fig. 2 are taken approximately 0.4 s after generator deployment in 

the adjacent volume on the right. Flames appear to be pushed into adjacent volumes that 

are in communication (have an opening) with the primary protected volume. This is 

likely due to the rapid volume displacement caused by the effluent of the solid propellant 

gas generators that occurs within the protected volume, which tends to push the hot gases, 

fuel vapors, and flames through the opening into the adjacent volume [Manchor, 2002]. 

Similar observations were made in many other experimental facilities with other 

suppressant types. For example, in the case of the NIST turbulent jet spray burner, 

gaseous suppressant release caused sudden changes in the appearance of the fire [Hamins 

et al., 1996]. The danger posed by a flare-up relates to the possible ignition of a fuel 

puddle or leak that may be present in the vicinity of the flare-up. The duration of a flareup can be expected to be on the order of the duration of the suppressant release. The 

intensity of a flare-up is related to the relative location of the fire to the suppressant 

release and its momentum. If suppression occurs, then the extended flame which is 

downstream of the flame anchoring region can be expected to disappear only after flame 

extinguishment occurs near the flame anchoring position. If suppression does not occur 

then the flare-up will disappear and the flame will return to its previous appearance. 

Heat transfer from a flare to a nearby object could lead to ignition depending on 

the critical ignition temperature and the duration of flame impingement. An increase in a 

material’s temperature depends on the thermal inertia of the object, and the amount and 

duration of the heat flux. The duration of the heat flux due to flame impingement from a 

flare could be expected to be less than or equal to the time of the suppressant delivery 

that typically varies from 1⁄4 s from a gas generator to several seconds for a compressed 

liquid discharge. Direct flame impingement on a surface would cause ignition of many 

liquid fuels with flash points above ambient temperature. 

9
Figure 2. Flare-up observed during full-scale V-22 fire suppression experiments 

at the China Lake test facility when gas generators are used to extinguish fire in a 

dry bay area. Top and bottom images were taken approximately 0.6 s apart. The 

top image was taken just after release of a solid propellant gas generator in the 

wheel bay, which is a volume adjacent to the right of the images. 

10
4.2 Literature Review and Interview Findings: Bench-Top Experimentation

Unwanted enhanced burning after suppressant delivery may occur through flowrelated effects that involve fuel splashing or flame straining. Although these effects are 

due to the consequences of the discharge of the suppressant, they are each distinct in their 

form. Other mechanisms may also contribute to accelerated burning. For example, it was 

shown that halogenated agents can enhance soot formation and heat release rate in benchscale flame experiments [Smyth and Everest, 1996; Homstedt et al., 1993]. 

4.2.1 Splashing 

 In engine nacelle and dry bay applications, suppressant delivery is typically a high 

momentum two-phase (liquid and gas) jet. Fast, high momentum suppressant application 

was shown to be advantageous in terms of suppressant mass requirements [Hamins et al., 

1994]. In aircraft applications, agent is often delivered through a dispersing nozzle or 

simply through a hollow tube. The suppressant discharge may or may not be directed at 

the fuel source of the fire. If the suppressant is aimed directly at a liquid fuel source, the 

momentum associated with suppressant delivery can cause splashing of the liquid fuel, 

impacting the character of the burning fire. Given sufficient momentum, the suppressant 

discharge can cause the ejection of liquid droplets from the burning fuel surface. 

 Yang and coworkers demonstrated the impact of a high momentum jet on the 

burning of a liquid hydrocarbon diffusion flame [Yang, 2002; Pitts et al., 1994]. In their 

experiments performed at NIST, gaseous nitrogen or HCFC-22 was released from a 

pressurized cylinder (4.1 MPa or 600 psig), which was located 1.6 m directly above a 

10 cm diameter liquid pool of burning heptane. A fast opening solenoid valve controlled 

the suppressant discharge. Figure 3 shows a series of frames from high speed filming 

(250 frames/s) of gaseous N2 delivery onto a burning heptane pool fire. The total time 

elapsed between frames 3(a) and 3(f) was approximately 0.5 s. Image analysis showed 

that the velocity of the N2 stream was approximately 90 m/s. 

 Figure 3a depicts the fire just before the N2 release. It is a typical pool fire with 

the flames forming a characteristic plume shape above the fuel surface. Figure 3b shows 

that within a few milliseconds after agent release, the gas-phase flame is disrupted and 

appears to be squashed or compressed. After 20 ms, Figure 3c shows that the flame 

11
appears to be extinguished. Small pockets of flame appear to remain about the periphery 

of the burner. Figure 2d shows the liquid fuel splashing about 80 ms after the release. A 

significant amount of fuel atomization was evident. The atomized fuel mixed with the 

surrounding air and burning of the fuel/air mixture occurred despite the presence of the 

agent. The occurrence of combustion in this scenario is governed by the mixing dynamics 

of the fuel, the air and the suppressant. Figures 3e and 3f, photographed about 300 ms 

and 500 ms after the release, respectively, show the formation of a large volume of flame. 

The visible flame area can be thought of as representative of the degree of fire 

enhancement. As the flame area increases, the fire heat release rate increases. In this 

manner, the images clearly demonstrate enhanced burning. Similar behavior was 

observed by Yang and Pitts [1994] during release of the compressed liquid CHF2Cl 

(HCFC-22) above the 10 cm heptane pool fire, despite the fact that HCFC-22 is almost a 

factor of three times more effective a suppressant than N2, on a volumetric basis [Hamins 

et al., 1994]. 

 Yang and Pitts observed that the delivery of a high momentum jet enhances the 

burning rate of the pool fire as the liquid fuel is forced from the burner. As this 

happenned, the flame luminosity increased and the effective burning area appeared to 

increase. The fact that the jet was composed of a suppressant did not assure that burning 

was suppressed throughout the flowfield. Fire suppression was assured only at locations 

with sufficient suppressant volume fraction. Thus, the dynamics of the agent flow control 

combustion and suppression of the pool flame. While there may be many effective 

strategies for extinguishing a 10 cm pool fire using a gaseous agent, a high momentum 

agent jet is not effective, due to fuel splashing and non-uniform agent mixing. 

 Although the splashing phenomena were documented through the suppression 

experiments described above, there is little understanding of the splashing behavior of a 

liquid fuel puddle subjected to a high momentum fluid jet. Trabold and Obot [1997] 

investigated the onset of splashing from free liquid surfaces exposed to impinging gas 

jets. Predictive equations for the critical mean nozzle exit velocity at the onset of liquid 

splashing were developed in terms of the relevant flow, geometric, and physical 

parameters. At a critical jet momentum, splashing occurs whereby droplets are ejected 

from the liquid surface. Yet there is no known available data on the dynamic spatial 

12
distribution and momentum of a splashing liquid and its interaction with an impinging jet 

or spray. 

 The ejection of burning liquid fuel drops may subsequently be transported to a 

location where ignition of non-combusting fuel could occur. Such a scenario is possible 

if the suppressant is a high momentum gaseous jet, but it is also possible if the 

suppressant is a low momentum liquid spray, emanating, for example, from a water 

sprinkler. A number of studies have been conducted on the effect of a water sprinkler 

spray on pool fires [Kokkala, 1992; Han et al., 1997; Kim et al., 1997]. In their flame 

suppression studies of liquid pool fires using overhead water sprays, flame-enhancing 

behavior similar to the high momentum suppressant jet was also observed [Kokkala, 

1992; Kim et al., 1997]. In these experiments, fire enhancement was attributed to 

increased mixing of fuel and air as a result of water interacting with the pool surface. 

Bench-top experiments with liquid water droplets falling onto the surface of a boiling 

hydrocarbon have also shown that splashing of the target liquid can occur depending on 

the impact Weber number of the incoming drop [Manzello and Yang, 2002]. 

13
 

 

 

 

(a) 

 

(b) 

Figure 3. Response of a 10 cm heptane pool fire to a high pressure, overhead nitrogen release.

(c) (d) 

 

(e) 

 

(f) 

14
4.2.2 Flame Straining

Rapid changes in the appearance of a flame such as its size or luminosity after 

suppressant discharge could be interpreted as enhanced or accelerated burning. Flame 

appearance including its shape and color is a natural consequence of the conditions that 

define the flow. Changes in the appearance of a flame may occur due to any number of 

factors including changes in the local instantaneous flows of fuel or air, or changes in the 

local pressure or temperature. Upon suppressant release, the flow field in a nacelle or dry 

bay may significantly change and the appearance of a fire may change as the local flows 

of fuel or air change. When suppressant delivery is complete, flowfield conditions may 

return to the previous state or could differ. If a flame experiences increased air flow, then 

the flame may undergo straining. 

An unstrained hydrocarbon diffusion flame is typically luminous in appearance. 

As the flow into such a flame increases, the flame weakens, becoming shorter and blue in 

appearance. Subsequent decrease in the flow would alter the flame appearance again, 

elongating the flame length and changing its color to one that is luminous. An example 

of this phenomenon is exemplified in Fig. 4, which shows JP-8 spray flames [Hamins et 

al., 1996] under conditions of high and low air flow, all other conditions being the same. 

The high air flow case represents near-extinction conditions and the flames are shorter 

and bluish. The low air flow case is much longer and the flames are more luminous, 

which would give rise to a relatively higher radiative heat flux. 

 10 cm 

 10 cm 

 12 cm 

Figure 4. Baffle stabilized JP-8 spray flames in a wind tunnel under conditions of 

approximately 10 m/s air flow (left) and 2 m/s air flow (right). 

15
4.2.3 Mixing 

The nature and rate of reactant mixing plays a role in the burning rate of a fire. 

This is true for both premixed and non-premixed flames. The discussion here is focused 

on non-premixed flames, since most actual fires are non-premixed by nature. A nonpremixed flame is reactant-limited by definition. A ventilation-limited fire is one type of 

reactant-limited fire, where the transport of air to the reaction zone limits the burning 

rate. A fuel-limited fire is another type of reactant-limited fire. It is defined by overventilated conditions and the transport of fuel to the reaction zone limits the burning rate. 

Transport of reactants to a fire can certainly be affected by agent delivery. This can 

happen in one of several ways. Air can be entrained into the flow of a suppressant or 

pushed ahead of a suppressant jet. Alternatively, a fuel leak in the form of a spray, for 

example, could be displaced or disrupted by a suppressant jet. Each of these scenarios is 

possible. Reactant delivery can also impact turbulence levels, which can impact heat 

transfer and burning rates. The impact of agent induced mixing on a fire will be 

dependent on the details of each of these scenarios and will depend on the specifics of the 

geometric configuration, the flowfield, and agent effectiveness. Air entrainment into the 

centerline of axisymmetric jets was previously studied for steady state flows [Pitts, 1991]. 

There is no information on entrainment in unsteady jets, or on the transient flow of air 

pushed just ahead of an instantaneous flow of a jet. In addition, these issues were not 

investigated in the context of a suppressant flow and the effect of entrainment on fire 

suppression. 

 7 cm 

 

4.2.4 Heat transfer and Chemistry related effects

 Unwanted enhanced burning after suppressant delivery may occur through heat 

transfer related effects. Heat transfer effects may include re-ignition and/or flame 

stabilization that can occur due to preheating of a solid object that acts like a flame 

anchor or ignition source. There are no known studies on the effect of a suppressant 

discharge on fire heat transfer. Yet, it is conceivable that the suppressant discharge could 

tilt a fire and promote heat transfer to a solid object such as a metal bulkhead or flange. 

Such a situation could enhance flame stability. Subsequent heating of such an item could 

promote re-ignition and flame stabilization, enhancing flame stability.

16
 Addition of the some types of suppressants can alter the flame chemistry and 

impact soot levels. For example, the chemically active suppressant CF3Br can enhance 

the concentration of soot in a hydrocarbon diffusion flame [Smyth and Everest, 1996]. 

This was demonstrated in both laminar and turbulent diffusion flames. Enhanced soot 

levels lead to increased radiative heat transfer, which is a key mechanism of fire spread 

and growth. 

 Some suppressants can be considered to be fuels as they are exothermic. For 

example, halogenated fire suppressants release enthalpy, albeit relatively small, when 

introduced into a hydrocarbon flame. This can lead to larger visible flame volumes, for 

example a 6 % volume fraction of CF3Br in an ambient temperature oxidizer stream has 

an adiabatic equilibrium temperature of 600 K. There are some studies in the literature on 

the chemical effects of a suppressant. Holmstedt and Anderson [1993] and Holmstedt et 

al. [1994] measured the heat release rate using oxygen consumption calorimetry for 

various suppressants added to propane/air diffusion flames. Some agents were found to 

increase the heat release rate by almost a factor of four. As the heat release rate of a fire 

increases, the flame volume may be expected to increase and the appearance of a fire may 

suddenly become very large with application of such a suppressant. 

4.5 Assessment of Background Information

The background information attained through interviews and the literature review 

indicates that a multiplicity of mechanisms may be relevant to an understanding of 

unwanted burning after suppressant delivery. These include flowfield-related effects, 

heat transfer effects, chemical effects, and enhanced mixing effects. It is often difficult to 

categorize these mechanisms in a simple manner because the physics are complex and the 

mechanisms are often coupled. Each of these effects may have an impact on flame 

volume or the appearance of a fire and on fire suppression requirements. Consideration of 

the background information, however, suggests that the main hazard associated with 

unwanted accelerated burning after suppressant delivery is system over-pressurization. 

Whereas changes in flame appearance or flame volume are not necessarily hazardous, 

system over-pressurization, on the other hand, could lead to structural damage of an 

aircraft. 

17
In an effort to better understand the hazards associated with the phenomena of 

unwanted accelerated burning, the next portion of this report analyzes three cases that 

exhibit system over-pressurization or pressure fluctuations after suppressant delivery. 

5. Case Studies

Section 4.1 of this report identified a number of full-scale experimental facilities 

in which unwanted accelerated burning after suppressant release appears to be hazardous. 

In those cases, over-pressure led to the destruction or the potential for destruction of the 

experimental facilities. In this section, the most hazardous cases are further analyzed in 

an effort to understand the basis for the unwanted accelerated burning after suppressant 

release. 

5.1 Case 1: Analysis of Over-Pressurization during Suppression in the NRL ModerateScale Enclosure 

Through a series of carefully crafted experiments, Maranghides and Sheinson 

[2000] at NRL documented that suppressant discharge in a moderate-scale enclosure 

(28 m3

) can cause significant fire flare-up and pressure increases for unconfined volatile 

liquid fuels in the presence of a hot ignition source. The experiments were conducted to 

evaluate the performance of various suppression technologies for use in flammable liquid 

storage rooms on-board Navy vessels. This section of the report summarizes their 

experimental findings, which were reported previously [Maranghides and Sheinson, 

2000] and then analyzes the experimental results. 

The suppression experiments are first described. Two types of fires burned 

simultaneously. The first consisted of a burning liquid fuel that flowed in a cascading 

manner over shelving filled with 20 L buckets simulating storage containers to be found 

shipboard. To control fuel delivery, the buckets were sealed and empty, and fuel 

associated with the cascading fire was metered. The fire was three-dimensional in nature 

and was obstructed from direct suppressant impingement. A burning fuel-filled 

rectangular pan (0.3 m x 0.3 m) was also used. The fuel was composed of 80 % 

methanol and 20 % heptane (by volume). Methanol was selected because it is a 

18
shipboard material that presents a difficult extinguishment challenge and the heptane was 

included to make the flames visible for safety purposes. 

The fire was allowed to burn for 20 s before the suppressant was delivered. The 

experiments used agent discharge cylinders filled with HFC-227ea (1,1,1,2,3,3,3-

Heptafluoropropane or FM-200) pressurized with N2 to 4.1 MPa, which required 

approximately 1 kg of nitrogen. The nozzle used for the tests had a 1 cm diameter and 

was located overhead in the center of the compartment and oriented downwards with its 

four outlets oriented in the cardinal directions (0o, 90o, 180o, 270o). Suppressant 

discharge times were roughly 10 s. 

 Figures 5a-c show the NRL fire suppression test facility from exterior and interior 

views. The exterior view (Fig. 5a) shows that the enclosure wall is fitted with four large 

(0.36 m2

) explosion relief panels. The first interior view (Fig. 5b) was taken through the 

door shown in (Fig. 5a). The photo shows the three-dimensional pan fire with cascading, 

burning liquid fuel. The fuel in this case was heptane, a volatile liquid. Figure 5c is an 

image captured from a video camera viewing the enclosure interior. The reddish glow is 

identified as a deflagration and appears as a radiating volume of combusting gas. This 

image was taken just before the explosion panels in Fig. 5a ruptured due to the large 

rapid pressure increase within the enclosure. 

(a) (b) (c) 

Figure 5. (a) Exterior view of the moderate-scale (28 m3) NRL enclosure; (b) Interior view of 

the NRL enclosure with an obstructed, cascading three-dimensional pan fire; (c) deflagration 

wave propagating within the enclosure. 

19
Typical measurements of the transient pressure within the NRL enclosure during 

the suppression experiments are shown in Figs. 6 and 7 [Sheinson and Maranghides, 

2002]. The suppressant mass discharged during the experiments was designed such that 

the suppressant volume fraction in the enclosure was equal to 9.1 % and 11.5 % for the 

conditions reported in Figs. 6 and 7, respectively. These values are equal to the cup 

burner volume fractions times a factor of 1.1 and 1.4, respectively. (The critical cup 

burner volume fraction for the methanol/heptane mixture was measured to be 8.3 % 

[Maranghides and Sheinson, 2000].) 

The experimental compartment was sealed during the experiments. As the fire 

heated the enclosure, the pressure was observed to increase. Figures 6 and 7 show that 

the pressure increased steadily after ignition. Experiments proceeded with 20 s of burning 

before initiation of the discharge of the HFC-227ea suppressant. The pressure in the 

enclosure (see Figs. 6 and 7) decreased rapidly as the compressed liquid agent vaporized, 

absorbing heat from within the compartment. This led to rapid cooling of the 

compartment and a consequent immediate decrease in the pressure. The cooling effect is 

quite large considering that the pressure drop occurs despite a significant amount of 

suppressant mass being added to the compartment. The temperature data (not shown) 

from these experiments are consistent with this interpretation. 

Figure 6 shows a rapid pressure rise shortly after the suppressant discharge. After 

the initial cooling effect, the suppressant in the compartment quickly expands, creating a 

pressure increase in the compartment. At time zero in Fig. 6, the suppressant was 

deployed and the fire was not extinguished by the suppressant discharge. During the 

suppressant discharge, cascading fuel continued to flow into the compartment. A portion 

of this fuel was consumed by the fire. A fraction of the fuel, however, was unburned and 

vaporized during the suppressant discharge process. The vaporized fuel mixed with air, 

leading to rapid combustion in the form of a deflagration wave. These factors led to a 

pressure increase of approximately 13 kPa, causing rupture of the explosion panels. 

Through analysis of the video records, the experimental results indicated that suppressant 

delivery caused intensification of the fire and an increase in the burning volume. 

Maranghides and Sheinson [2000] suggest that the suppressant discharge increased fuel 

evaporation, leading to increases in the fuel burning area associated with the cascading 

20
fire. The pressure measurements in concert with the video record suggest that a 

combustion/deflagration wave developed in the enclosure. It should be noted that a hot 

ignition source (≈ 550  ̊C) was present throughout the experiments. The deflagration 

could have been initiated by the ignition source or by an unsuppressed kernel of the fire. 

Other experiments at NRL showed that the pressure increase from the suppressant 

discharge alone (without the presence of the fire) did not cause the explosion panels to 

rupture. Experiments with a fire alone (no suppressant discharge) did not cause the 

explosion panels to rupture. In that case, the fire leveled-off at approximately 5.3 kPa 

almost 30 s after fire initiation and the fire self-extinguished in the closed compartment 

after 75 s. 

Figure 7 represents a fire that was identical to that corresponding to the fire 

associated with Fig. 6 except that the suppressant mass was approximately 40 % greater. 

The fire in Fig. 7 was extinguished by the suppressant discharge. The pressure trace in 

Fig. 7 was similar to Fig. 6 until the agent was released. At that time, the pressure rapidly 

decreased until its value reached the baseline. At that time, the video record shows that 

the fire was extinguished and no deflagration wave was observed. 

In previous tests, rapid over-pressurization was observed to occur for a variety of 

suppressant types including gaseous, compressed liquid, and hybrid suppressants 

composed of liquid-phase plus dry powder. The results suggest that a combustible 

(premixed) fuel-air mixture can propagate when complete extinguishment does not occur 

and suppressant inertion volume fractions are not reached. The inertion volume fraction 

of HFC-227ea in a fuel mixture of 80 % methanol plus 20 % heptane was never 

measured. However, the inertion volume fraction of HFC-227ea has been reported as 

11 % to 11.5 % by Holmstedt et al. [1994] in a propane/air mixture. Experiments were 

conducted in an 8 L vacuum tight cube fitted with an ignition electrode. Using the same 

device, Holmstedt et al. [1994] measured the inertion volume fraction of Halon 1301 to 

be 6.5 % by volume in agreement with Hamins et al. [1994]. The effects of ignition 

energy and enclosure size were shown to impact the results. 

Flames burning methanol are expected to be more difficult to extinguish. The fuel 

of interest, however, is a 20 % methanol plus 80 % heptane mixture (by volume). Macek 

[1974] shows that as alcohols become less oxygenated, the temperature limits resemble 

21
pure hydrocarbons. Calculations using the NASA CEC program [Gordon and McBride, 

1976] of the limiting volume fraction in a heptane/air mixture based on a critical 

temperature criteria for flammability limits [Macek, 1974] yields the same inertion 

volume fraction as a propane/air mixture. Therefore, the 11 % value is considered a rough 

estimate of the inertion volume fraction for HFC-227ea in the methanol/heptane fuel 

mixture. 

In Figs. 6 and 7, the suppressant volume fractions of 9.1 % and 11.5 % of HFC227ea were insufficient and sufficient, respectively, to extinguish the fires and prevent reignition. These results are consistent with the 11 % inertion requirements discussed 

above. The situation is complicated by a number of factors. First, it is probably safe to 

assume that suppressant mixing was not homogeneous throughout the enclosure. Second, 

the HFC-227ea flammability data in this case (and typically) is for reactant mixtures 

initially at ambient temperature. Flammability limits, however, are temperature 

dependent and regions of the fires associated with Figs. 6 and 7 were preheated by the 

fires. These facts suggest that more information is needed to make definitive conclusions 

regarding the NRL experiments. 

A separate series of experiments were conducted with a 2 min preburn duration 

and compartment ventilation turned off 30 s preceding agent discharge. The O2 depletion 

and CO2 generation during these experiments weakened the fires, significantly enhancing 

the performance of the fire suppressant and resulting in very rapid fire extinguishment. 

No over-pressurization or deflagrations were observed under these conditions.

22
23

Alternatives Test B4

FM200 - 45 Second Preburn

Com partm ent Pressure

-0.5

0

0.5

1

1.5

2

-60.0 -30.0 0.0 30.0 60.0 90.0 120.0

Tim e (sec)

Pressure (psi)

fire 

suppressant 

Conc = 11.5 % = 1.8·CB

Figure 7. NRL fire suppression test results for FM-200 delivery an obstructed, cascading 3D/pan fire 

in a moderate scale enclosure (28 m3

) with a design volume fraction of 11.5 %, which is equal to the 

cup burner volume fraction (CB) + 40 %. The fire was extinguished and the pressure panels did not 

blow-out. 

Alternative Test B-2

FM-200 - 20 Second Preburn

Compartment Pressure

-0.5

0

0.5

1

1.5

2

-30.0 0.0 30.0 60.0 90.0 120.0

Time (sec)

Pressure (psi)

fire 

Panel ruptures Conc = 9.1 % = 1.4·CB 

Suppressant 

applied 

Time (s) 

Figure 6. NRL fire suppression test results for FM-200 delivery onto the obstructed, cascading 

3D/pan fire in the moderate scale enclosure (28 m3

) with a design volume fraction of 9.1 %, 

which is equal to the cup burner volume fraction (CB) + 10 %.
5. 2 Case 2: Modeling of Global Properties during Fire Suppression in a Dry Bay 

5.2.1. Background

An unanticipated catastrophic event occurred during real-scale fire suppression testing on 

August 28, 1998 at the C-130 VRP Phase IA-Wing Leading Edge Dry Bay Fire 

Extinguishing Agent Evaluation Baseline test (WLE-R-B-07), which was conducted in 

the Aircraft Survivability Research Facility Upper Range 3. This was the seventh and 

final baseline test on a replica C130H WLE dry bay test article containing ≈1000 L of 

JP8. In that incident, over-pressurization destroyed the full-scale test fixture. Six tests 

previous to the catastrophic event using N2-pressurized HFC-125 had led to routine 

suppression results. The seventh test was designed to investigate the impact of the N2

used for pressurization on the fire. It should also be noted that many hundreds, if not 

thousands, of suppression experiments were conducted in the Wright Patterson Dry Bay 

facilities without the occurrence of catastrophic over-pressurization. This included 

experiments in which suppressant delivery either extinguished or did not extinguish the 

test fire. 

 Dry bays are structural elements that are typically located in landing gear wheel 

wells, wing leading edges, munition bays, avionics equipment bays, and the fuselage of 

military aircraft. Dry bays vary considerably in their geometry, size, and ventilation, and 

have been described as void volumes within the mold line of an airplane, excluding air 

inlets, engine compartments, and exhaust nozzles [Bennett, 2002]. Dry bays normally 

contain electrical and mechanical components such as electrical cables, flight control 

actuators, hydraulic accumulators, and liquid oxygen containers. They are often adjacent 

to fuel tanks and may have fuel, hydraulic, coolant, and/or lubricating fluid lines passing 

through them. During normal operating conditions, dry bays are free of flammable liquids 

and vapors. During combat or equipment failure, however, flammable liquids may be 

released, leading to the possibility of a fire. 

24
Bleed Air Duct 

CO2 Nozzle Agent nozzle

Fixed WLE 

Figure 8. Photograph showing the inside of the Fixed Wing Leading Edge 

(WLE) C-130 Dry Bay Test Fixture with agent nozzle. 

Fuel Tank 

Engine Area

Fixed WLE 

Figure 9. Drawing of the Fixed Wing Leading Edge (WLE) C-130 

Dry Bay Test Fixture including the engine area and fuel tank. 

 Dry bays are particularly vulnerable to fire when an anti-aircraft projectile 

ruptures a fuel tank or fluid line, necessitating rapid automatic detection and deployment 

of a suppressant in order to avoid aircraft loss. Full-scale fire suppression testing of dry 

bay fixtures has been conducted at a number of facilities in the U.S including the Wright 

Patterson Air Force Base, where an extensive experimental program has been conducted 

25
for many years. Testing was based on simulation of realistic operating conditions and 

“live” incendiary projectiles varying in size from 12.7 mm to 23 mm. Figures 8 and 9 

show the Fixed Wing Leading Edge C-130 Dry Bay Test Fixture including the agent 

nozzle, fuel tank and engine area. 

 The incident that occurred during real-scale fire suppression testing on August 28, 

1998, which led to destruction of the Wright Patterson C-130 VRP Phase IA-Wing 

Leading Edge Dry Bay was unique. Because the reasons for the failure were unclear, an 

investigation was commissioned. The investigation reconstructed the experimental 

timeline and examined possible causes of the incident. The salient information in the 

investigation report included the following incident timeline: 

1. a projectile penetrated the outer wall of the Dry bay (time = 0). 

2. The detection system was activated and the suppressant (N2) deployed at 20 ms. 

3. At approximately 40 ms, the agent was delivered into the fixture. 

4. Approximately 0.1 L of nitrogen was discharged into the replica dry bay volume of 

1500 L. 

5. The line leading from the agent bottle to the delivery location within the Dry bay was 

filled with air at ambient conditions. 

6. The Dry bay failed at 110 ms. 

7. Fire/flames were observed at 166 ms. 

The rate of the pressure rise for the test was similar to previous tests. There was 

an initial short duration pressure rise of 30 kPa due to the threat impact, a quick drop-off, 

then a ramping to the peak at 95 ms. A rapid drop-off after the peak pressure occurred 

due to failure of the test article and subsequent venting. The investigation report 

concluded that the release of agent (N2) in this case likely acted to mechanically agitate 

the air, which contributed to a vigorous and rapid combustion event that led to overpressurization and mechanical damage of the test apparatus. 

This was a unique incident in terms of aircraft fire suppression testing. According 

to those who perform dry bay suppression testing, there are no other documented 

incidents of catastrophic failure of a real-scale suppression test facility. Conversations 

with those in the aircraft industry suggest that such failures are extremely rare or non26
existent [Rosenfeld, 2002]. For these reasons, the C130 WLE Dry Bay incident provides 

a unique data point on the phenomena of unwanted accelerated burning after suppressant 

delivery. Probable cause of the incident merits attention. 

The investigation report is classified, which precluded obtaining the report in its 

entirety [Haas, 2000; Cyphers, 1998]. From the sections of the report that were finally 

made available to NIST, no modeling of the incident was reported. Quantitative analysis 

was limited to use of the ideal gas law. In this study, modeling is used to gain insight 

into the nature of the phenomena. Detailed computational fluid dynamics (CFD) 

modeling of the C130 WLE Dry Bay suppression experiments, however, would be 

extremely complex due to the presence of a ballistic round, accompanying shock waves, 

fuel splashing (two phase flow), hydraulic ram, multiple ignition sources (hot shrapnel), 

and so on. Instead, a simplified analytic approach was taken to simulate the essential 

dynamics of the incident. That model is outlined below. 

It should be noted that over-pressurization of the test apparatus was not unusual. 

According to measurements made during the experiment, failure of the C130 dry bay 

fixture occurred when the pressure inside of the apparatus reached 54 kPa. The apparatus 

had survived earlier experiments when the pressure had reached 30 kPa. 

The experiment used an inert suppressant (N2), but at a volume fraction that was 

0.4 % of the cup burner value, an amount that would have little impact on the stability of 

the fire and certainly would not extinguish the fire. To better understand the conditions 

in the dry bay during fire suppression, an analytic approach was taken to simulate the 

pressure in the enclosure. A numerical approach was not used due to the complexities 

mentioned above. 

An Air Force investigation of the event concluded that “increased level of fluid 

ignition (combustion)” was the source of overpressure. Experimental peculiarities were 

noted including the use of a suppressant delivery line that was air-filled. 

5.2.2. Analytic Approach

A simplified analytic approach is described below, which was used to simulate the 

essential dynamics of the global properties during fire suppression in the incident that 

27
occurred at the Dry Bay Test Facility at WPAFB on August 28, 1998. Background 

information on the incident was briefly described above. 

 

Projectile 

induced 

hole 

Applied 

external 

airflow Hot upper 

layer gases

Fire 

Dry Bay 

Fuel spray/flow 

Fuel tank 

Figure 10. Fuel spray and fire resulting from projectile impact through 

the dry bay into the fuel tank. 

Suppressant spray 

 In the experiments, a projectile penetrated the outer wall of the Dry bay (time= 0) 

and then penetrated the fuel tank (160 000 L), causing a flow and spray of jet fuel. The 

volume of the dry bay was ≈ 3 m3

. Hot shrapnel rapidly ignited the fuel. Figure 10 

schematically shows the fuel spray and fire resulting from impact of the projectile. 

Within a few milliseconds, an automatic detection system activated the suppression 

system, which delivered agent into the compartment at approximately 40 ms. The static 

pressure transducer in the Dry bay peaked at 54 kPa at 95 ms. Almost immediately (10 

ms later) the pressure in the Dry bay fell to ambient as the test compartment failed. The 

maximum pressure in other similar tests typically obtained 30 kPa and compartment 

failure did not occur. 

Using the formalism developed by Rehm and Forney (1992), the pressure in the 

dry bay during the fire suppression event is derived based on mass and energy 

conservation. Hydrodynamic ram associated with the projectile impact is not considered 

in this analysis. The basic premise of this simple type of fire model is that the enclosure 

can be divided into a small number of regions or zones with approximately uniform 

28
conditions. Zone models are widely used to understand the conditions in an enclosure 

during a fire. Zone models typically divide an enclosure into two control volumes, an 

upper layer of hot gas and smoke, and a lower layer of air. 

The dry bay compartment in question was on the order of 1 m in length. In the 

experiments, a projectile strikes the external dry bay wall. At room temperature, the 

speed of sound in air is approximately 300 m/s, implying that the pressure in the dry bay 

equilibrates in a time on the order of 3 ms. The timeline associated with the spray fire is 

thought to be on the order of 100 ms, much longer than that required for equilibration of 

the compartment pressure, and allows application of the low Mach number form of the 

conservation equations. The equations that govern conservation of mass and energy are 

applicable even for the relatively rapid processes that occur in the situation of interest. 

The gas in each layer has attributes of mass, internal energy, density, temperature 

and volume denoted respectively by mi , Ei, ρi, Ti, and Vi where i = L in the lower layer 

and i = U in the upper layer (see Fig. 10). The compartment as a whole has the attribute 

of a single pressure P. These eleven variables are related by the following constraints: 

ρi = mi / Vi (1) 

Ei = cv mi Ti (2) 

P = R ρi Ti (3) 

V = VL + VU (4) 

Equations 1-4 represent the definitions of density, internal energy of a gas, the ideal gas 

law, and the total volume. The specific heats at constant volume and constant pressure are 

related to the universal gas constant R by: 

 R = cp - cv (5) 

The differential equations for mass in each layer are: 

mL = & dt

dmL (6)

mU = & dt

dmU (7)

29
The first law of thermodynamics states that the rate of increase of internal energy plus the 

rate at which the layer does work by expansion is equal to the rate at which enthalpy is 

added to the gas. The first law can be represented in differential form as: 

i

i i q dt

dV P

dt

dE

=+ & (8)

A differential equation for the pressure can be determined by adding the upper and lower 

layer versions of Eq. 8 and noting that the change in volume of the upper and lower 

layers are related as: 

dt

dV

dt

dVU L −= (9)

Equations 1-3 yield: 

dt

PVd

R

cv i)(

V

)Tmd(c dt

dEi iiv = = (10)

This leads to an equation for the time varying pressure P: 

)q( V

1)-( dt

dP

L qU = & + & γ (11)

Equation 11 can be rewritten in terms of the non-dimensional pressure ( p

~ ) in the 

dry bay enclosure (relative to the ambient external pressure, = penc

~ - pamb~ ) and nondimensional time ( t

~ ) considering enthalpy sources and sinks. The parameter γ represents 

the ratio of the specific heats (Cp/Cv), which is taken as a constant equal to 1.4, the value 

of γ for air. This formulation includes the global effects of forced ventilation and 

suppressant addition and is similar to that derived by Rehm and Forney [1992b] for the 

time-dependent pressure: 

(12) )q-qq( V

1)-( t

~

d

p

~d

f vs = & & ++ && qfv

γ

where V is the enclosure volume and is the dimensionless sensible enthalpy addition to 

the compartment due to several source terms. The subscript “f” represents the fire, “fv” 

qi &

30
is the enthalpy addition to the compartment due to forced ventilation, “s” represents 

enthalpy addition due to suppressant deployment, and ”v” is the enthalpy that is lost to 

the exterior by flow through a hole or vent caused by a projectile. A portion (≈ 10 %) of 

the pressure differential is due to the Venturi effect caused by airflow on the exterior of 

the dry bay, which tends to diminish the pressure in the enclosure. 

The calculation yields two interesting results, namely, that only a small difference 

in could have caused failure of the fixture and that paradoxically, the pressure rise is 

larger for smaller vent sizes (small diameter projectiles). The results are consistent with 

the idea that the N

f q&

2 release caused an increased amount of mixing and a relatively small 

amount of enhanced burning. The small size of the hole caused by the projectile 

exacerbated the situation. 

In the C130 WLE dry bay suppression experiments, flight conditions were 

simulated by blowing air on the outside of the fixture at high speeds, typically 130 m/s 

(300 mph), to simulate in-flight conditions. In this case, the pressure differential must 

account for the Venturi effect, which increases the fluid flow through the vent, although 

this effect is relatively small. Solution of Eq. 12 allows calculation of the pressure in the 

dry bay for various vent sizes. For simplicity, the fire source and the suppressant 

enthalpy addition are taken as constants. The size of the effective “vent” is determined 

by the incoming projectile and will be at least as large as the diameter of the incoming 

projectile. In a real aircraft, the Venturi-related pressure difference across a projectile 

induced vent can vary significantly depending on flight elevation, air speed, wind 

conditions, and the orifice coefficient, which is related to the hole character. 

In the incident of interest, forced ventilation was zero and the initial pressure, 

p

~ ( t

~ = 0), in Eq. 12 was taken as equal to 0.1 to account for the pressure difference 

associated with a venturi induced pressure for a 130 m/s wind (300 mph) blowing past 

the dry bay. The venturi effect tends to decrease the final pressure in the dry bay by a 

relatively small amount (< 10 % ), depending on the flow rate. Figure 11 shows the 

solution to Eq. 12 for p

~ ( t

~ = 0) = 0.1. The pressure obtains 90 % of its steady value 

( p

~ = 0.9) when t

~ = 4 and asymptotes to p

~ = 1 when t

~ = 10. 

31
In the case of the dry bay incident in question, little suppressant was deployed and 

 << and forced ventilation ( ) was absent [Cyphers, 2002]. The pressure 

increase due to the suppressant delivery was less than 300 Pa (0.04 psig). Solution of Eq. 

12 for these conditions is shown in Fig. 12, which is plotted in dimensional units with 

equal to the dimensional fire heat release rate. Figure 12 shows that the calculated steady 

state enclosure pressure is a function of the vent size (projectile hole diameter) and the 

fire heat release rate ( ). For a larger fire size, the calculation results show that the 

pressure will be larger. The figure also shows that for a larger vent size, the final pressure 

in the enclosure 

qs & q f & qfv &

q f &

q f &

2 4 6 8 10

0.2

0.4

0.6

0.8

1

p

~

t

~

Figure 11. Solution of Eq. 12 for the pressure as a function of time in dimensionless 

units for initial conditions defined by p

~ ( t

~ = 0) = 0.1. 

32
0

10

20

30

40

50

60

70

0 100 200 300

Fire Heat Release Rate (kW)

Pressure (kPa)

is smaller. The enclosure pressure for small vent sizes is particularly sensitive to the fire 

size. The hole (assumed to be round) in the incident of interest was approximately 13 cm 

in diameter [Cyphers, 2002] and the critical pressure just before failure of the enclosure 

was 54 KPa (7.8 psig). The enclosure in previous experiments reached pressures as high 

as 30 kPa (4.3 psig) without catastrophic failure. Considering these parameters, solution 

of Eq. 12 leads to the results presented in Fig. 13, which shows the relationship between 

fire size and vent diameter for final enclosure pressures of 54 kPa (7.8 psig) and 30 kPa 

(4.3 psig). The figure shows that the difference in fire size ( ) between these two 

pressures is roughly 25 %, the difference between 600 kW and 800 kW or 200 kW. 

Thus, only a moderate difference in fire size could have caused pressurization sufficient 

to bring about failure of the test fixture. All other conditions being equal, Fig. 13 shows 

that paradoxically, a dangerous situation can arise in the instance of a smaller projectile 

hole size, which may lead to a smaller vent size. Of course, all other conditions may not 

be equal if the hole size differs. 

q f &

The above calculations assumed that the time to reach steady pressure in the 

enclosure depends on the vent and fire sizes. This is not necessarily the case. Figure 14 

Figure 12. The calculated steady enclosure pressure as a function of the (dimensional) 

fire heat release rate ( q f & ) for various vent sizes. 

D=2.5 cm

D=5.0 cm

D=7.5 cm

vent diameter

33
shows the calculated time to obtain a steady pressure in the enclosure ( t

~ = 4 in Fig. 11) 

as a function of hole size for fire with a 50 kW, 100 kW, 200 kW, and 400 kW heat 

release rate. For a 100 kW fire and a 12.7 cm hole, steady pressure is achieved in 

122 ms. The time increases as the fire size increases and as the hole size decreases. In the 

incident in question, the pressure was measured to obtain a maximum 95 ms after 

projectile impact. The fire is assumed to have initiated at the time of projectile impact. 

This suggests that a steady state pressure was likely not obtained in the incident in 

question. The calculated fire sizes would therefore need to be larger than those shown in 

Fig. 12 to achieve the observed failure pressure after 95 ms, but the difference between 

the fire size resulting in pressure failure and the fire size for the largest observed pressure 

when failure did not occur is approximately 25 %. This indicates that there is a fine line 

between safety and failure in terms of the critical fire size. The dry bay problem is 

complicated by the actual geometry in the dry bay, which was composed of two similar 

size enclosures connected by a relatively small duct (13 cm in diameter). A smaller 

effective dry bay volume would require a correspondingly smaller heat release rate to 

obtain the same enclosure pressures given the same heat release rate. Similarly, steady 

burning would take less time to achieve in a smaller enclosure volume. 

Figure 15 shows the solution to Eq. 12, 95 ms after initiation of the fire for a vent 

size of 12.7 cm (5 in). The critical failure pressure of 54 kPa (7.8 psig) is obtained for a 

steady 2.1 MW fire, whereas a pressure of 30 kPa (4.3 psig) is obtained for a fire size of 

1.5 MW, which is a difference of almost 25 %. 

34
0

100

200

300

400

500

600

700

800

900

0 510 15

vent diameter (cm) 

Qf (kW)

Pfinal = 54 kPa

Pfinal = 30 kPa

Figure 13. The fire heat release rate as a function of the vent hole diameter for 

steady pressures equal to 54 kPa (7.8 psig) and 30 kPa (4.3 psig). 

Figure 14. The calculated time to obtain a steady enclosure pressure ( t

~ = 4 in 

Fig. 11) as a function of vent hole diameter for various fire heat release rates. 

10

100

1000

10000

0 5 10 15 20

Hole Size (cm)

Time to Steady State (ms)

Qf = 50 kW

Qf = 100 kW

Qf = 200 kW

Qf = 400 kW

35
0

20000

40000

60000

80000

100000

120000

0 1000 2000 3000

Qf (kW)

Enclosure Pressure (Pa) Figure 15. The calculated enclosure pressure as a function of the fire heat 

release rate 95 ms after fire initiation for a 12.7 cm (5 in) vent hole size. 

If only the dry bay volume is considered, the required fire heat release rate to 

obtain the same enclosure pressures resulting in failure would be approximately 1 MW, 

whereas a pressure of 30 kPa (4.3 psig) would be obtained for a fire size of 800 kW. A 

smaller vent size would require a correspondingly smaller fire heat release rate to obtain 

the same enclosure pressure. A 1 MW fire would require burning an average of 

approximately 20 kg of jet fuel per second. This may not be an unreasonable value 

considering the quantity of fuel available (160 000 L) and the volume of the dry bay 

(≈ 3 m3

). For turbulent hydrocarbon pool fires, heat release rates per volume are typically 

1200 kW/m3 to 2500 kW/m3 depending on the measurement technique used [de Ris, 

1979; Orloff and de Ris, 1983]. It is not unreasonable to expect comparable values for a 

turbulent spray fire (composed of many small non-premixed flames associated with 

burning droplets) burning in a 3 m3

 volume. 

36
The internal investigation of the dry bay incident suggested that the release of 

agent (N2 in this case) acted to mechanically agitate the air contributing to a more 

vigorous and rapid combustion event. A detailed understanding of the exact nature of 

this event for the purposes of this report has been hampered by the availability of detailed 

specific information on the conditions and geometry of the incident in question. This is 

largely due to the availability of documents associated with the test program and the 

investigation itself. The documents that were made available are consistent with the 

conclusions of the internal investigation at Wright Patterson AFB described above. It 

should be emphasized, however, that the lack of an effective amount of agent precipitated 

the over-pressurization event. Contributing to the situation is the fact that N2 is not an 

effective suppressant, requiring almost 4 times more agent by volume than HFC-125, for 

example, to extinguish a heptane cup burner flame [Hamins et al., 1994] and at the 

volume fractions used in this experiments, suppression was extremely unlikely. It should 

also be noted, that over-pressurization was typical during these experiments. The 

difference in the catastrophic incident was that the over-pressurization was more intense 

than over-pressurization due to the agent alone or due to the projectile impact alone. The 

pressure data from 12 experiments showed that over-pressures greater than 48 kPa 

occurred during two experiments. The catastrophic C130 test was unique in that the 

agent amount was too small to be effective, but the lack of suppressant was not the lone 

contributor to the pressure increase in this incident. An analysis of a series of baseline 

experiments associated with this program in the C130 dry bay Simulator suggests that 

there was a non-negligible probability of large (> 54 kPa) static pressure increases just 

from the projectile impact. Unfortunately, these data are not available for general release. 

Clearance issues such as this hampered progress and communication of results associated 

with this study. 

5.2.3 Summary

In summary, the simplified global analysis yields two interesting results, namely 

that a moderate difference in could have caused failure of the fixture and that 

paradoxically, the pressure rise is larger for smaller vent sizes (small diameter 

projectiles). The results suggest that a pressure rise strong enough to cause the overf q&

37
pressurization and apparatus failure was likely due to suppressant (N2) induced mixing 

leading to enhanced combustion. A more complete analysis would consider 

compressibility and dynamic effects. 

5.3 Case 3: Pressure Fluctuations During Suppression in an Enclosure 

The experiments performed by Kashiwagi et al. [2000] on suppression of 

moderate-sized (≈ 200 kW) hydrocarbon pool-fires burning in an enclosure are examined 

below. A schematic drawing of the experimental apparatus taken from Kashiwagi et al.

[2000] is shown in Fig. 16. Twenty experiments were conducted using five types of 

suppressant including HFC-227ea, HFC-23, and the inert mixtures, IG-541 (52 % N2, 

40 % Ar, 8 % CO2), IG-55 (50 % N2, 50 % Ar), and IG-100 (100 % N2). Each suppressant 

used a different type of nozzle. The mass of suppressant discharged was approximately 

20 % above cup burner values. The design volume fraction varied from 38 % to 41 % for 

the inert suppressants, 8 % for HFC-227ea, and 16 % for HFC-23. In the experiments, 

30 kg to 45 kg of suppressant was discharged from a pressurized cylinder over a period 

varying from 10 s to 105 s depending on the suppressant type with the inert 

Pressure transducers

Gas analyzers

FT-IR

Temperature recorder

Heptane pan fire

Windows

Door

2.6 m

4.8 m

Nozzle

Sampling

Door tubes

Suppressant

4.8 m

Vent

To a scrubber

Figure 16. Schematic diagram of the enclosure suppression experiment taken 

from Kashiwagi et al. [2000]. 

38
suppressant taking 55 s or longer. Suppressant effectiveness was tested for heptane pool 

fires varying from approximately 200 kW to 400 kW. A number of experiments were 

also conducted using wood cribs. The fire was ignited with the door open and the vent 

closed. The vent was spring activated, responding to the enclosure pressure. The exact 

response function was not characterized. This is unfortunate as the enclosure pressure is 

highly sensitive to the vent opening. 

Approximately 10 s before suppressant discharge, the door was closed. 

Simultaneously, the vent began to open. No controlled ignition source was present in the 

enclosure, although the hot metal of the burner or fire-heated thermocouples could have 

acted as an ignition source. A complete description of the procedure and the 

experimental results is given in Kashiwagi et al. [2000]. Pressure fluctuations as large as 

1 kPa to 3 kPa in the enclosure were measured during the experiments [Kashiwagi et al., 

2000]. A typical pressure trace is shown in Fig. 17. The maximum size of the pressure 

fluctuations in these experiments was an order of magnitude smaller than in the two other 

cases considered in this report (Sections 5.1 and 5.2). The reasons for the pressure 

fluctuations are not clear. It is possible, however, that the fluctuations were due to 

dynamics of the suppressant flow shielding air from fuel. Of course, the pressure traces 

reported by Kashiwagi et al. [2000] are for successful suppression tests only. 

Unsuccessful suppression may have led to over-pressures not unlike those observed in the 

first two cases of this report. In this sense, the Kashiwagi’s results are consistent with the 

first two cases examined in this report – large over-pressures did not occur when 

suppression was successful. 

Pressure excursions were observed when both halogenated or inert agents were 

applied. The pressure fluctuations are attributed to the dynamic nature of the suppression 

process in which a portion of the enclosure volume is suppressed successfully. As fuel 

vapor flows from the liquid puddle, reacting with available air, excursions in pressure 

occur as local combustion occurs. 

Attempts to recreate aspects of the experiments including the pressure fluctuations 

using the NIST Fire Dynamics Simulator (FDS) software were largely unsuccessful. FDS 

is a large-domain CFD fire model that predicts and visualizes the spread, growth, and 

suppression of a fire for practical engineering applications based on the underlying 

39
scientific principles governing fluid motion. The model solves numerically the 

conservation equations of mass, momentum, and energy that govern low-speed, 

thermally-driven flows with an emphasis on smoke and heat transport from fires. The 

code addresses combustion and hydrodynamic phenomena over a range of physical 

length scales varying six orders of magnitude. In common usage, FDS employs either a 

mixture fraction model (infinite rate chemistry) or simplified finite rate chemistry. Here, 

single step chemistry was employed. Depending on the kinetic rates selected, almost any 

type of dynamic enclosure pressure result was possible. Using Arrhenius rate parameters 

that correspond to measurements in small-scale experiments failed to yield pressure 

fluctuations similar to those observed by Kashiwagi et al. [2000]. Flame suppression 

phenomena take place on the physical length scale of a flame sheet, which is on the order 

of 1 mm in length and is over an order of magnitude smaller than the grid size that is 

practical for CFD fire models such as FDS. Further work is needed to develop robust fire 

suppression models for large-scale fire phenomena. 

-20

0

20

40

60

80

-20 0 20 40 60 80 100

Room Pressure (mmH 2O)

Time after Agent Discharge (sec)

(a) Agent: IG-55

Fire source: n-heptane

0.3m 2 (420 kW) 

Door closed.

Figure 17. The room pressure measured experimentally for the case of a heptane 

pan (0.3 m2

 area) fire (420 kW) suppressed by an inert suppressant mixture. The 

figure is taken directly from Kashiwagi et al. [2000]. 

40
5.4 Accelerated Burning in Engine Nacelles

Because of the serious hazard posed by over-pressurization of an enclosure as 

exemplified through the dry bay incident described in the previous section, this section 

first considers this phenomenon. Interviews with those performing full-scale experiments 

in engine nacelles and those working in the aircraft industry suggest that there has never 

been a case of over pressure damage associated with agent release in an engine nacelle. 

According to a Boeing representative, nacelle failure due to suppressant deployment has 

never been recorded [Rosenfelds, 2002]. Nacelles are designed to handle catastrophic 

failure of the bleed air duct, which connects the nacelle to the engine itself. For this 

reason, actual nacelles include blow-out panels and ports that are designed to rapidly 

open in response to significant pressure increases due to rapid additions of large flows 

into the nacelle [Rosenfelds, 2002]. 

Structural damage to a nacelle may occur due to weather, foreign object impact 

(birds, projectile) or catastrophic engine failure (core components shred), but structural 

damage due to over-pressurization from suppressant deployment and fire is improbable 

[Ingerson, 2002]. Finally, nacelle fires are not expected to be as large as dry bay fires. 

Agent quantity is designed to produce a specified or designed volume fraction range. For 

these reasons, structural failure in a nacelle due to agent deployment is unlikely. During 

full-scale suppression testing at the Wright Patterson Engine Nacelle Facility, nacelle 

over-pressure during the F-22 program was not observed. Experiments during the F-22 

experiments used a replica of the F-22 nacelle materials, geometry, and flow conditions. 

A typical pressure trace is shown in Fig. 18. The experiment involved a jet fuel spray 

flame extinguished by a solid propellant gas generator (SPGG) released at time zero in 

the figure. In these experiments, the pressure increase never exceeded 3 kPa to 7 kPa. 

Other suppression experiments at the Wright Patterson facility over the last few years 

have not led to system over-pressurization, structural damage to the facility, or blow-out 

of rupture panels [Tucker, 2002]. These results are consistent with Boeing’s contention 

that nacelle failure due to suppressant deployment is highly unlikely [Rosenfelds, 2002]. 

41
F22 nacelle

Run pac031_b

98

100

102

104

106

108

-25 -20 -15 -10 -5 0 5 10

Time (s)

Static Pressure (kPa)

Figure 18. Pressure trace taken during the 1997 F22 engine nacelle 

test program at Wright Patterson Air Force Base.

6. Summary and Conclusions 

A literature review and survey of practitioners suggested that there are several 

mechanisms associated with accelerated burning after suppressant delivery. An analysis 

of the information suggests that the main hazard of accelerated burning is system overpressurization due to enhanced combustion from vaporizing fuel that mixes with air after 

suppressant delivery. Small amounts of over-pressurization are not problematic, but if 

the over-pressurization is large, then damage of a structural element may occur. 

Depending on the location and function of the structural element, over-pressurization 

may impact aircraft function. 

The phenomenon associated with enhanced burning after suppressant delivery is 

of considerable importance and certainly can impact strategies and mass requirements 

associated with fire suppression in aircraft applications. It is important to note that each 

of the possible mechanisms of unwanted enhanced burning after suppressant delivery that 

are described in Section 1 may occur not only in actual engine nacelles and dry bays, but 

can also be expected to occur during full-scale suppression testing. In this sense, it is 

expected that accelerated burning effects are already embedded in the full-scale test 

42
results and therefore in the design equations that are based on those results. With this 

recognition, a central question becomes, how well do the current generation of full-scale 

fire suppression tests represent the range of conditions experienced in the field? If a fire 

test program considers a broad range of conditions that incorporates all conditions 

experienced in the field, then that aspect of the test program should be considered 

adequate. On the other hand, if a fire test program does not consider a broad range of 

conditions, then such a program should be considered lacking. Fire conditions in fullscale suppression tests should be carefully designed to include a worst case or else 

unwanted burning effects may occur under extreme conditions that were unanticipated 

during a full-scale test program. 

Based on this conclusion, it is recommended that full-scale fire suppression 

testing be founded on a comprehensive scientifically-based design process rather than on 

an ad-hoc basis, which has historically been the case. The design of full-scale fire 

suppression test scenarios ought to use the best analytic and computational tools 

available, which may require a significant portion of the resources of any test program. 

Currently, full-scale fire suppression testing is an art. It needs to integrate scientific 

analysis in the design of experiments and interpretation and analysis of experimental 

results. 

A high momentum suppressant discharge was shown to be advantageous in terms 

of reducing the total mass required to suppress a fire. The splashing of fuel associated 

with suppressant discharge could present a mechanism of fire spread and enhanced 

combustion. There is no evidence, however, that shows that suppression is not possible 

using a high momentum discharge towards a fuel puddle in a dry bay or nacelle, given 

sufficient suppressant. Appropriate safety factors should be determined and applied in 

fire suppression systems design to avoid unwanted enhanced burning effects. Still, it may 

be prudent to direct a high momentum suppressant discharge away from a fuel source 

provided that adequate spatial coverage by the suppressant is not compromised. This 

should prove to be feasible in both nacelles and dry bays. 

In summary, the main hazard associated with enhanced burning after suppressant 

delivery is system over-pressurization likely due to rapid mixing of fuel vapor with air. 

Insufficient suppressant delivery and enhanced mixing can lead to enhanced burning, 

43
depending on the circumstances. Unwanted burning effects did not appear to occur when 

sufficient suppressant was applied to assure suppression and prevention of re-ignition. 

Assuring that sufficient suppressant is delivered to the fire can minimize enhanced 

burning hazards. 

To reiterate the most important conclusion of this work, unwanted accelerated 

burning effects are already embedded in full-scale test results and the resulting design 

equations. Further research is needed to develop accurate general-purpose CFD models 

of fire suppression. This study demonstrates the need for data to better predict 

flammability limits, especially at elevated temperature. 

7. Acknowledgements 

The authors are indebted to Donald Bein (NAVAIR), Mike Bennett (WPAFB), 

Linda Blevins (SANDIA), Dan Cyphers (Skyward Ltd ), Doug Ingerson (FAA), 

Alexander Maranghides (NIST), Joe Manchor (China Lake), Bill Pitts (NIST ), Kuldeep 

Prasad (NIST), Ron Sheinson (NRL), Yuko Saso (FRI), and Jiann Yang (NIST) for many 

helpful discussions and suggestions. 

44
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48