ASSESSMENT OF FIRE-SUPPRESSION SIMULATIONS USING 

FULL-SCALE ENGINE NACELLE TESTS1

D. R. Keyser 

INS, Inc. 

Pine Hill Technology Park, Bldg. 1 

48015-A Pine Hill Run Rd. 

Lexington Park, MD 20653 

J. C. Hewson 

Fire Science and Technology 

Sandia National Laboratories 

Albuquerque, NM 87185-1135, USA 

ABSTRACT 

Results are presented for a series of fire-suppression tests conducted in a full-scale 

nacelle along with the results of pretest predictions carried out using the Vulcan firephysics simulation code. The purpose of these tests was to assess the utility of using a 

CFD fire-physics computer code (Vulcan) to predict the performance of a firesuppression system. The test plans and test matrix were made to correspond to Vulcan 

simulations previously run to predict the test results. Twenty-five fire tests were 

conducted successfully in the NAVAIR ground nacelle simulator to validate this firemodeling code. This simulator is a full-scale ‘iron bird’ mockup of an engine nacelle 

assembly typical of a combat aircraft with a four-nozzle suppressant-distribution 

system. JP-8 pool fires in a variety of locations were considered, and both the 

suppressant and the overall air flow rates through the nacelle were varied to simulate a 

wide range of operating conditions. In addition, the effects of varying the 

inhomogeneities were investigated by capping individual nozzles. The pretest Vulcan 

simulations correctly predicted the success or failure of extinction in nearly all cases. In 

only two cases were there disagreements between the Vulcan predictions and the test 

results. In both of these cases, when either of the two forward nozzles was capped, the 

test results found that the fires were suppressed using a bottle pressure 25% lower than 

the Vulcan simulations had predicted. This indicates that the Vulcan simulations were 

somewhat conservative in their predictions of the mixing in the forward nacelle region. 

Detailed data analyses of the twenty-five fire tests are presented in this paper. 

INTRODUCTION 

Advances in computational power and physics models relevant for fire-suppression 

applications have reached the point where predictive modeling of fire suppression in 

complex geometries is feasible using Reynolds-Averaged Navier-Stokes (RANS) 

 

1

 This work was conducted in part at Sandia National Laboratories, a multiprogram laboratory operated by Sandia Corporation, a 

Lockheed Martin Company, for the United States Department of Energy’s National Nuclear Security Administration under Contract 

DE-AC04-94AL85000. This research is part of the Department of Defense's Next Generation Fire Suppression Technology Program, 

funded by the Department of Defense Strategic Environmental Research and Development Program. 
Computational Fluid Dynamics (CFD). At this level, it is possible to account for 

inhomogeneities as the mean flows are resolved and mixing is treated using wellestablished models [1-3] having a sound physical basis. This is in contrast to zonal 

models where the composition is assumed to be constant over larger regions, with 

empirical mixing rules between zones. 

In fire-suppression applications, the ability to predict inhomogeneities in the suppressant 

concentration is a key requirement for optimizing fire-suppression systems. If regions of 

the flow field exist where the suppressant concentration is locally below the value that 

leads to suppression (referenced to the cup-burner value or the strained flame-extinction 

value), then combustion can continue in those regions. In certain applications where the 

objective is to reduce the fire until fire-fighting crews can attack it, such behavior is 

acceptable. However, this is not acceptable for fires in aircraft nacelles and other 

inaccessible areas. Three factors come into play in aircraft nacelles and similar spaces: 

First, the quantity of fire suppressant that can be carried is severely limited in weight and 

space in order to satisfy the mission requirements. Second, the need to eliminate trace 

combustible vapors requires ventilation that also quickly sweeps suppressant out of the 

nacelle as it sweeps out combustibles. Third, a failure to completely suppress a fire can 

be catastrophic, since small pockets of fire can quickly propagate through the remaining 

premixed gases in the nacelle leading to accelerated burning under certain conditions. 

The present work under the Next Generation Program has focused on developing a set of 

models and tools to predict suppressant distributions in order to meet these challenges 

[4]. The results discussed herein are focused on suppressants with boiling temperatures 

well below operating temperatures, such as HFC-125 (C2HF5). Other work has addressed 

issues related to higher-boiling point agents [5-8]. 

The purpose of the tests described here is to assess the utility of CFD models like Vulcan 

to predict the outcomes of a fire-suppression system’s performance. The tests and 

simulations are conducted in the context of a full-scale ‘iron-bird’ aircraft nacelle groundtest simulator located at the Patuxent River Naval Air Station. This nacelle simulator is 

representative of actual nacelles incorporating realistic arrangements of structural 

support, piping, machinery and other items collectively referred to as ‘clutter.’ This 

scenario tests both the fluid flow and the fire-suppression aspects of Vulcan in an 

application where the flow is under-resolved, and some of the physics is captured in 

subgrid-scale clutter models. The test sequence was developed based on Vulcan 

simulations described in [9,10] with the conditions designed to cross from successful to 

failed suppression. While simulations and tests were conducted in a simulated aircraft 

engine nacelle as determined by the focus of the NGP plan, the capabilities demonstrated 

herein are expected to be applicable to a wider range of scenarios including aircraft dry 

bays, land-based vehicles, and ships. 

In this paper, we describe in brief the Vulcan simulations and the test procedures 

employed. The results of the tests are then compared to the prior Vulcan predictions, and 

the performance of the CFD model is assessed. 
DESCRIPTION OF THE NACELLE GROUND-TEST SIMULATOR 

Figure 1 shows the fire test simulator. The air inlet source is seen in the lower left 

coming up into the lower, forward end of the nacelle. The air is supplied from an 

electric-motor-driven centrifugal compressor and measured with a calibrated turbine 

meter. There is adequate straight pipe, according to ASME Standards [11] and a flow 

straightener between the compressor and the turbine meter to provide uniform flow 

profiles for flow measurement. Likewise there is adequate straight pipe downstream of 

the turbine meter, and downstream of the 45 degree elbow there is an Etoile swirlremoving conditioner in the straight pipe leading to the nacelle providing uniform 

velocity profiles similar to those in flight. 

Suppressant is introduced into the nacelle through as many as four nozzles each on 

different streamlines and fuselage stations. The nozzles orifices have diameters ranging 

from 3 to 5 mm and are indicated with identifying numbers in the text where the numbers 

increase from nozzle 1 to nozzle 4 moving from the forward to the aft end of the nacelle. 

The suppressant flows to the nozzles from a bottle located on the inboard side of the 

nacelle through a piping structure. The bottle is filled with suppressant as described in 

the Test Method section. The nozzle characteristics and the suppressant concentrations 

generated by these nozzles are detailed in [10]. 

There are two vents in the top: a diamond-shaped vent all the way aft, and a balance 

piston round vent at the forward third. There are also four small holes (2 to 3 in. dia.) in 

the front face simulating vents and connections to the AMAD bay. Air, products of 

combustion, and suppressant can leave the nacelle through these vents and holes as 

described in [10]. 

An array of 21 type ‘K’ thermocouples was installed over the bottom of the nacelle in 

three rows, 8 each on the sides and 5 along the center, spaced approximately equally 

streamwise about 2 inches above the ribs. These were used to measure the temperatures 

above the pool fires and to detect the presence of the flames. These also provide some 

indication of the flow patterns that can be compared to the Vulcan model predictions. 
Figure 1. The F/A-18 Nacelle Ground Test Simulator, Outboard Side. 

Instrumentation employed in the tests and their uncertainties:

1.Supply Air Flow Measurement: 6-inch Turbine Meter, Sponsler Co. Inc., Model SP6-

CB-PH7-C-4X. S/N 130619 Calibrated in water 3 Sep 04: + 0.75% systematic, + 0.06% 

random. 

2.Temperature Measurements: 21 Type K thermocouples, calibrated at ambient: +1.1 

deg. F systematic and +1.8 deg. F random.

3.Digital Stopwatches (2): Calibrated error less than 0.1 sec over 15 minute period 

4. Graduated Cylinders for measuring tares and final fuel quantities from the pools, + 5 

ml.

5. Video Cameras (4) to record the existence of fire and its extinguishment inside the 

nacelle. 

6. Pressure Gauge, NoShok, 0-1000, + 6 psi. (0-71 + 0.4 bar) 

7. Dead Weight Tester, Druck Co., 0-10,000 psi. + 2.5 psi.(0-710 +0.2 bar) 

8. Pressure Transducer, Patriot Gage Co. 0-3000 psi. + 15 psi.(0-207 + 1 bar) 

9. Digital Data Acquisition System integral with the ground test nacelle. 

10. Calibrated Weigh Scale for the HFC-125. 

TEST METHOD 

This test series focused on extinguishing JP-8 pool fires contained by structural ribs and 

longerons along the lower nacelle surface. Prior to the testing, the bottom of the nacelle

was lowered and the spaces between the ribs were sealed from their normal drain holes so 

that each could become a container for fuel. In the test series, three pools were used, 

beginning at the front of the nacelle and progressing in order to the middle of the nacelle.

These pools are designated B1, B2 and B3 (or B3C3 indicating that the pool crossed one

longeron). Prior to the tests, measured quantities of water were added to each pool, and 

the dimensions of the resulting pools were recorded to document the surface area of each
pool as a function of liquid contained in the pool. The areas measured are tabulated in 

[9]. 

In most of the tests, all three pools were ignited because this provided the strongest fire 

source. In several tests only the first or third pool was used in conjunction with a reduced 

number of agent-distribution nozzles in order to assess the spatial distribution of the 

agent in both the Vulcan model and the real nacelle. The fuel burning rate and the nacelle 

temperatures were observed to be substantially lower when a single pool was burned as 

predicted in the Vulcan pretest simulations [9]. 

The mass of agent to be used in each test was measured by the difference over the tare of 

the empty ‘bottle’ on a calibrated scale. HFC-125 was the agent used in all tests. The 

rate of injection was controlled by the nitrogen pressure level in this bottle, and three 

rates of injection were targeted by setting the bottle pressure to one of three values. 

These values targeted injection rates that were expected to roughly correspond to a 

designed system (515 psig, 35.5 bar), a rate roughly 75% of the designed system (275 

psig, 19 bar) and a rate roughly 50% of the designed system (122 psig, 8.5 bar); at these 

ambient temperatures the vapor pressure of HFC-125 was about 122 psig (8.5 bar) so that 

no nitrogen charge was added for the lowest pressure. Nitrogen is somewhat soluble in 

this agent, and consequently the fire-suppression agent in these tests was a mixture of 

both HFC-125 and nitrogen; the latter is ignored in the Vulcan simulations. The 

suppressant bottle discharges downward so that the initial discharge is driven by the 

nitrogen pressure, which drops as liquid agent is displaced from the bottle. At an 

intermediate pressure the rate of pressure change slows; this is associated with agent 

discharge being driven largely by the boiling of the agent or the dissolved nitrogen. Note 

that for the lowest bottle pressure there is no initial nitrogen-driven stage. The last stage 

of the discharge follows, in which the gaseous agent and nitrogen in the bottle flow from 

the bottle. The pressure profiles from several tests, where discharge was through all of 

the nozzles, are shown in Fig. 2. The discharges from the two tests denoted 1a(21) and 

1a(25) are essentially the same conditions, as are the four tests denoted 1b(21), 5b(25), 

8a(26) and 1b(26). The spread in these pressure profiles is indicative of uncertainties 

attributable to, for example, ambient temperature changes that may heat the bottle during 

the few minutes that the test is being set up. Such uncertainties are thought to be the 

most significant and are estimated to be roughly +10%. 
0

100

200

300

400

500

600

0 2 4 6 8 10 12

Time after start of discharge [s]

B

ottle pressure [psig]

1a(25)

1a(21)

1b(21)

1d(25)

5b(25)

8a(26)

1b(26)

 Figure 2. The suppressant discharge pressure, measured at the bottle, 

for various initial pressures with all nozzles discharging and a total of 

3.2 kg of agent. 

Attempts to model this discharge based on the pressure data with simple pipe flow 

network models has proven challenging, and the results indicate that the flow at the 

nozzle where most of the pressure drop occurs is most likely multiphase flow. We

estimate somewhat more gas than liquid phase at the nozzles, with our best estimate 

being 60% gas and 40% liquid by volume. This volume ratio is expected to vary during 

the evolution of each discharge, moving towards 100% vapor at the end of the discharge.

Note that the majority of the agent mass remains almost entirely liquid, but the presence 

of the gas phase dramatically reduces the rate at which the agent flows into the nacelle. 

Once the bottle was filled, it was installed into the distribution piping on the nacelle, and 

its manual safety valve was opened. Predetermined quantities of fuel were poured from 

graduated cylinders into the pools chosen for each test. At the end of each test, the 

remaining fuel was drained from each pool so that the amount consumed could be

accounted; these data, along with the elapsed time of the test, gave an estimate of the 

heat-release rate during the test. 

Igniting the pools of JP-8 fuel at ambient temperatures is not trivial. The bottom of the 

nacelle was warmed with hot air to enhance the volatility of the fuel and to reduce heat 

losses. A minimum air flow from the blower was established to provide sufficient air for

combustion. An electric igniter was inserted over each pool in turn while a more volatile 

fuel, pure ethyl alcohol, was sprayed onto the pool. Once the pool ignited, it required 
about 30 seconds to become vigorously inflamed. After all pools had established a 

durable flame, the air flow was ramped up to the predetermined rate—either1 kg/s or 0.5 

kg/s according to the turbine meter. This steady flow was maintained for about 15 to 30 

seconds before the agent was released. The elapsed period of burning was recorded with 

stopwatches as well as the digital data acquisition system. Four video cameras recorded 

the views inside and outside the nacelle by which elapsed times for extinguishment were 

recorded. This data acquisition system also recorded the signals from the array of 21 

thermocouples in the nacelle. The temporal response of these thermocouples was not 

sufficient to determine the temperatures in the interior of the nacelle; the indicated 

temperatures had not attained steady state over the duration of the fires. 

After each release of agent, a reduced air flow was maintained to cool the nacelle and 

remove potentially inflammable vapors. Then the remaining fuel was collected and 

measured in the graduated cylinders. 

KEY RESULTS FROM THE PRETEST SIMULATIONS 

Certain results from the Vulcan pretest simulations and simplified analyses helped guide 

the design of the experimental test plan and matrix. These results estimated the overall 

concentration of suppressant in the nacelle, the duration of the transients as the 

concentrations rise and fall, and the magnitude of the distribution of suppressant. To 

begin with, a simplified analysis of the relative masses of air and suppressant flowing 

into the nacelle can provide some guidance on the overall suppressant concentration and 

the transients. The suppressant mass-injection rate relative to the total injection rate 

provides a characteristic mean mass fraction within the nacelle, Yss. The injection must 

proceed for a long duration for the mean mass fraction to reach this characteristic steadystate value, but the mean mass fraction approaches this value in an exponential manner 

with an exponential time constant that is proportional to the total nacelle volume divided 

by the total volumetric influx. This time constant indicates the time scale for transients. 

In the tests conducted here, it was found that suppression is less sensitive to the total 

injected mass than it is to the rate of injection. 

The Vulcan predictions indicate, and the test results confirm, that the overall or average 

mass fraction of suppressant resulting in fire suppression is substantially greater than the 

cup burner value, which is Ycb=0.28 [12]. In general, the estimated average fraction 

required for suppression is on the order of 30% to 40% greater than the cup-burner value, 

according to the Vulcan simulations. A similar excess was required for the tests. The 

fact that a greater overall mass fraction is required indicates that inhomogeneities are 

significant. In other words, the suppressant’s mass fraction in certain regions of the flow 

is substantially less than the mean. It is noted that the high rates of mixing present in the 

nacelle tend to reduce the mass fraction required to suppress the fire. It is found for HFC125 that strained laminar flames are indicated to be suppressed at mass fractions as low 

as 0.16 [12]. 

One of the primary objectives of the present study is to ascertain the predictive 

capabilities regarding the degree of inhomogeneity in the nacelle. To this end it is noted 

that predictions with the Vulcan models were very successful at reproducing extinction in 
a geometry where the suppressant was introduced in a uniform manner and the only 

mixing processes were related to a fire-stabilizing recirculation zone [3]. The present 

geometry is appreciably more complicated, and the leading challenge was expected to be 

the transport of the suppressant rather than the well-established suppression model itself. 

In order to evaluate the mixing process for several scenarios, simulations and tests were 

conducted with different rates of injection and with different nozzle configurations. As 

an added variable, the air inflow rate was varied to simulate varying flight conditions. As 

indicated above, the ratio of the suppressant’s injection rate to the combined, total mass 

rates is indicative of the mean of the suppressant mass fractions in the nacelle. 

TEST PLAN 

From the Vulcan model simulations of various test conditions, it was predicted that the 

suppression would be more sensitive to the rate of injection of agent than the amount of 

suppressant, or equivalently the duration of injection. Further, suppression is indicated to 

be sensitive to the distribution of agent via the distribution of nozzles about the nacelle. 

With these results, a test plan was generated in the form of a rule-based sequence of tests. 

The test sequence is initiated with an approximation of the production suppression 

system, which is expected to result in successful suppression. From this point subsequent 

tests reduce the effectiveness of that system by either: 

• reducing the suppressant injection rate by reducing the suppressant bottle pressure so 

that the suppressant’s mass fraction is likely to be reduced or 

• removing a nozzle so that the distribution of suppressant is likely to be less uniform, 

or 

• reducing the mass of suppressant injected for a given bottle pressure so that the 

fraction is reduced and held for a shorter duration. 

As indicated above, tests were run also for reduced air flow rates. However, because of 

limitations on the rate of injection imposed by the HFC-125 vapor pressure, it was not 

possible to reach a condition where suppression failed at lower air flows. That is in 

agreement with predictions for lower air flows: the overall mass fraction of the agent for 

the physically attainable injection rates is always sufficiently high to extinguish the fires. 

For the tests in which a nozzle was removed (by capping the end), the sequence was 

repeated also by lowering the bottle pressure until suppression failed. The general 

sequence for changing the injection rate was to reduce the injection rate by 50% if the 

previous attempt succeeded and to increase the injection rate back to 75% of the original 

rate if the previous attempt failed to extinguish. The lowest injection rate that is 

attainable is 50% of the nominal design injection rate, so this process results in 

bracketing the suppression in proximity to 100%, 75%, and 50% of normal injection 

rates. The nominal pressures corresponding to these tests were 122, 275, and 515 psig 

(8.5, 19.0, and 35.5 bar) though these pressures vary somewhat with the mass of agent in 

the bottle. While these are relatively wide margins, the available resources did not allow 

for additional tests to narrow these bands. The tests actually conducted are summarized 

in Table 1. 
Table 1. Summary of Tests Conducted. 

Variations from the baseline conditions, apart from pressure, are indicated in bold.

Test 

indicator 

air 

inflow 

rate pools nozzles

supp.

mass

bottle 

press.

targ 

supp 

rate 

[lb/s] [lb] [psig] [lb/s] 

1a(25) 2 all all 7 515 2.33 

1a(21) 2 all all 7 515 2.33 

1b(21) 2 all all 7 122 1.16 

1d(25) 2 all all 7 265 1.75 

5d(25) 1 all all 7 265 1.75 

5b(25) 1 all all 7 122 1.16 

7a(26) 2 all all 4.84 458 2.33 

7d(26) 2 all all 4.84 270 1.75 

8a(26) 1.5 all all 7 121 1.16 

1b(26) 2 all all 7 131 1.16 

2a(27) 2 3 not 2 7 515 1.6 

2d(27) 2 all not 2 7 275 1.2 

2b(27) 2 all not 2 7 515 1.6 

2c(27) 2 3 not 2 7 275 1.2 

3d(27) 2 all not 1 7 275 1.35 

3b(27) 2 all not 1 7 515 1.8 

3e(28) 1.5 all not 1 7 275 1.35 

3f(28) 2 1 not 1 7 275 1.35 

3g(28) 2 3 not 1 7 275 1.35 

4a(28) 2 all not 3 7 275 1.2 

4b(28) 2 all not 3 7 122 0.8 

7e(28) 2 all all 3 515 2.33 

7f(28) 2 all all 2.25 600 2.33 

SUMMARY OF MEASUREMENTS AND OBSERVATIONS 

A total of twenty-five tests were conducted following the plan to explore the edges of the 

extinguishment envelope for this ground test nacelle simulator. Two of these tests were 

conducted to determine whether or not individual pools (B1 and B3) could stabilize a 

fire. It was demonstrated that these pools could sustain a fire. This is in agreement with 

Vulcan simulations, although the Vulcan simulations indicated that the stability of fires in 

individual pools was sensitive to heat losses. Specifically, if the heat losses associated 

with conduction through the pool to the nacelle are as great as 50 % of the heat flux to the 

pool (essentially reducing the vaporization rate by 50 %), then certain pools not 

employed in the present series of tests could fail to sustain themselves. In the tests it was 

necessary to apply heat sources (heat lamps) to the nacelle under the pools to minimize 
heat losses just to get the fires stabilized, and this lends support to the Vulcan 

observations. 

Of the remaining tests, two tests were replications of the first two tests in order to gain 

confidence in the results; in each replication, the results of repeated tests were identical. 

Because of the physical limitations, namely the vapor pressure of the HFC-125 and the 

inability to light or stabilize fires in certain pools, certain Vulcan simulations could not be 

reproduced. Consequently certain tests were run without pretest simulation results, but 

by using the arguments in the section on ‘Key Results from the Pretest Simulations’ a 

similar range of parameter space was identified, and these results were in agreement with 

the Vulcan simulation trends. The results of the testing are described now for: 

(1) tests with all nozzles in which the ratio of the suppressant mass injection rate 

to the total mass injection rate was varied, 

(2) tests where the total mass of agent was reduced, and 

(3) tests where one of the nozzles was capped. 

The results will be presented in terms of the target ratio of the suppressant mass injection 

rate to the total mass injection rate. Ideally, the pressure data indicated in Fig. 2 would be 

used to determine the actual suppressant discharge rate to a level of accuracy similar to 

the accuracy of the inlet flow meter. However, the unknowns associated with the phase 

transitions occurring both in the bottle and in the distribution piping prevent such a 

determination. Instead, the discharge rate is estimated. For this purpose, discharge rates 

identical to those in the Vulcan simulations are employed. Specifically, with all nozzles 

discharging, the 3.2 kg of HFC-125 was presumed to discharge uniformly over 3, 4.5 and 

6 s for bottle pressures of 515 (35.5 bar), 275 (19 bar), and 122 psig (8.5 bar). The 

discharge rate was presumed to be reduced in accordance with the reduction in the total 

nozzle area when nozzles were capped. Clearly, the assumption of constant discharge 

rates in the Vulcan model will be a source of test uncertainty. This uncertainty varies 

over the duration of the suppressant injection period, and the errors in the estimates 

provided here are likely to be greatest in the earliest (fraction of the first second) and 

latest periods of the injection. If the rate of suppressant injection is considered averaged 

over the significant couple of seconds, say the first 2-3 s of the injection process, then 

based on the results of our analysis, we estimate uncertainties on the order of +15 %. 
0.0 0.2 0.4 0.6 0.8

1a(25)

1a(21)

1b(21)

1d(25)

5d(25)

5b(25)

7a(26)

7d(26)

8a(26)

1b(26)

7e(28)

7f(28)

ratio of mass injection rates

7f(28)

7e(28)

1b(26)

8a(26)

7d(26)

7a(26)

5b(25)

5d(25)

1d(25)

1b(21)

1a(21)

1a(25)

Figure 3. A summary of results for tests with all nozzles. Blue bars 

represent successful extinction, while red bars represent a failure to 

suppress the fire. In all of these cases, the Vulcan predictions were in 

agreement with the tests. 

The results of the tests employing all nozzles are shown in Fig. 3. We first discuss the 

results of tests with labels starting in ‘1’ ‘5’ or ‘8’ because they correspond to baseline 

cases using approximately seven pounds (3.2 kg) of suppressant at varying bottle

pressures and air flows. This series can be viewed as varying the ratio of the

suppressant’s mass injection rate to the total mass influx rate. Suppression is observed to 

fail when this ratio is below (approximately) 0.4 (+ 10% uncertainty in this value, but the

total uncertainty is estimated at +19%). This corresponds to those cases where the air 

flow rate is greatest (2 lb/sec, 0.9kg/s) and the bottle pressure is lowest (122 psig, 8.5 

bar). For reduced air flow rates (1.5 lb/sec, 0.68 kg/s) or increased bottle pressures (265 

psig, 18.2 bar) suppression is successful. The Vulcan predictions are in agreement with 

the test results for all of these tests. The fact that the ratio of injection rate exceeds the 

cup burner mass fraction by a factor of 1.4 (0.4/0.28) is indicative of the degree of 

inhomogeneity in the system. 

The results in Fig. 3 with labels starting ‘7’ are those tests where the mass of suppressant

was reduced from 7 lbs, to 4.84 lbs, 3 lbs and 2.2 lbs (from 3.2 kg to 2.2, 1.4 and 1 kg). 

All of these tests resulted in successful suppression of the fire. In all of these cases, the 

rate of injection was close to that for the designed conditions; in other words, the 

suppressant was injected just as fast, but for a shorter duration, in these tests. It is 

estimated that the bulk of the suppressant is injected in 3 s for 7 lbs, in 2 s for 4.84 lbs, 

and in 1 s for 2.2 to 3 lbs. These results reinforce the importance of the ratio of mass

injection rates. The Vulcan simulations were conducted with 7 and 4.84 lbs (3.2 and 2.2 
kg) of agent, and the results were in agreement with the tests. No Vulcan simulations 

were run for agent masses less than 4.84 lbs (2.2 kg). 

 

0.000 0.200 0.400 0.600 0.800

2a(27)

2d(27)

2b(27)

2c(27)

3d(27)

3b(27)

3e(28)

3f(28)

3g(28)

4a(28)

4b(28)

Ratio of mass injection rates

4b(28)

4a(28)

3g(28)

3f(28)

3e(28)

3b(27)

3d(27)

2c(27)

2b(27)

2d(27)

2a(27)

without nozzle 3

without nozzle 1

without nozzle 2

Figure 4. A summary of test results with specified nozzles capped. 

Those cases shown with the blue bars represent successful extinction, 

and those shown with red bars represent a failure to suppress the fire. 

In all of these cases, the Vulcan predictions were in agreement with 

the test results. Green bars indicate cases where suppression was 

successful in the tests, but Vulcan indicated failure to extinguish. 

Figure 4 summarizes results for those tests in which one of the four nozzles is capped. 

Capping a nozzle has two effects: First of all, the total nozzle area is reduced so that, for

the same nominal bottle pressures employed, the suppressant’s injection rate is reduced.

The reduction depends on each specific nozzle orifice diameter, but the reduction is on

the order of 25%. This effect is captured by the indicated ratio of the mass injection 

rates. The second effect of capping a nozzle is the increased inhomogeneity inside the 

nacelle. More inhomogeneity would require a greater ratio of injection rates to define the

boundary between successful and failed suppression. In the Vulcan simulations, it was 

predicted that higher inhomogeneity would be associated with capping either nozzles

numbered 1 or 2,--but not 3. 

The results of tests when nozzle 3 is capped are indicated by the labels with the prefix ‘4’ 

at the top of Fig. 4. It is seen that the mass ratio where suppression fails is moving 

towards lower values than observed when all nozzles are employed. It is not certain 

whether or not the degree of inhomogeneity actually decreases and performance of the 

system increases by capping nozzle 3, and this conclusion should not be drawn based on 
the present results even if this is implied in the test results. Nozzle 3 injects suppressant 

into the upper nacelle, far from any of the pool fires considered here. For this reason, 

reduced suppressant in the upper nacelle may not impact suppressant levels in the lower 

nacelle where the fires are. The uncertainties in the suppressant’s injection rate and the 

lack of replicate tests prohibit drawing firm conclusions for these cases. 

Tests where nozzle 1 is capped are indicated with the prefix ‘3’ in Fig. 4. For these tests, 

it is necessary to refer to Table 1, because in some tests only specific pools were ignited 

to find the locations where fires failed to extinguish. The tests where all pools are ignited 

are discussed first (3b, 3d, 3e). In cases 3b and 3e the Vulcan simulations and the tests 

showed different outcomes. Specifically, the Vulcan predictions indicated a failure to 

suppress the fire in test 3b (the Vulcan prediction for 3e was not actually run, but 

extrapolated) whereas the fire was extinguished in the tests. Reducing the injection rate 

so that the ratio of mass rates is 20 % less leads to a successful prediction as in case 3d. 

This implies that Vulcan tends to predict failure to suppress before it actually occurs. 

This may be because the mixing rate within the Vulcan simulation errs on the low side, 

thereby tending to underestimate mixing in highly cluttered areas bounded by walls. This 

is thought to be the case because the current clutter model does not account for enhanced 

mixing properly near the wall. Improvements to the clutter model may improve these 

predictive capabilities, but the fact that Vulcan predictions tend to be conservative, with a 

safety factor on the order of 20% as indicated here, is viewed to be preferable and 

acceptable. In terms of the ratio of the mass injection rates, the tests indicate a critical 

value between 0.4 and 0.45, a statistically insignificant increase over the 0.4 (+/-15 %) 

estimated for tests employing all nozzles. The Vulcan predictions indicated that the 

critical ratio of mass injection rates was between 0.45 and 0.55. In tests 3f and 3g, only 

pools B1 and B3 were filled, respectively. This was done to ascertain the location of the 

inhomogeneity that led to failed suppression. Here it was found that low concentrations 

of agent occur in the vicinity of pool B1 and not pool B3. This agrees with the Vulcan 

simulations. We view the ability of Vulcan to identify the region where the fire could not 

be suppressed as favorable. 

The results of tests for which nozzle 2 is capped are indicated by the labels with the 

prefix ‘2’ in Fig. 4. As with the previous examples, the Vulcan simulations indicated a 

failure to suppress at the higher injection rates due to increased inhomogeneities. The 

test results show that the failure to suppress occurs at the next step down in the injection 

rate (or the suppressant bottle pressure), the same as when nozzle 1 was capped. This 

again indicates a 20 % safety factor relative to the Vulcan simulations. It is noted that the 

failure to suppress when either nozzle is capped occurs at a ratio of injection rates that is 

again spanning 0.4 which is the same value indicated in the tests employing all nozzles. 

This implies that the degree of inhomogeneity in the tests where nozzles were capped is 

not significantly greater than those where all nozzles were open; in the Vulcan 

simulations capping nozzle 2 did indicate greater inhomogeneities in the fire region. 

CONCLUSIONS 

The results of a series of fire suppression tests in the complex environment of a simulated 

aircraft engine nacelle have been compared quite favorably with CFD simulations using 
the Vulcan code. The test plan was developed based on the pretest Vulcan predictions, 

and increments of roughly 25% in the air flow rate and the suppressant injection rate 

were investigated. In addition, the effect of varying the arrangement of suppressantdistribution nozzles was investigated by removing nozzles one at a time. The Vulcan 

predictions were generally successful in predicting the results of the tests (in all but 2 

cases). In general, success or failure in extinguishing the fires is largely correlated with 

the ratio of the rate of injection of suppressant to the total inflow rate (air plus 

suppressant). In agreement with this finding, significant reductions in the total mass of 

suppressant used still resulted in successful extinguishment, in which the ‘standard’ mass 

of 3.2 kg was reduced to just over 1 kg. 

For those cases where all four of the nozzles were employed and for those cases where 

the nozzle towards the middle of the nacelle was capped (#3), all of the test results agreed 

with the Vulcan predictions. For those cases where either of the nozzles in the forward 

section of the nacelle was capped, the test results indicated successful suppression using a 

bottle pressure commensurate with a 25% reduction in the suppressant’s injection rate 

relative to that predicted in the Vulcan simulations—that is, it was easier to suppress than 

predicted. This suggests that the Vulcan simulations are somewhat conservative in 

predicting the mixing of the suppressant into fire regions towards the forward end of the 

nacelle. Thus, Vulcan presents a factor of safety of about 20% relative to the tests 

conducted here. The largest uncertainty in these results resides in the variance between 

the assumed uniform injection rate of suppressant in Vulcan and the difficulties in 

determining the actual two-phase injection rates during the tests. 

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