SANDIA REPORT 

SAND2007-3442 

Unlimited Release 

Printed June 2007 

Suppression of Pool Fires with HFC-125 

in a Simulated Engine Nacelle 

John C. Hewson and David R. Keyser 

Prepared by 

Sandia National Laboratories 

Albuquerque, New Mexico 87185 and Livermore, California 94550 

Sandia is 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. 

Approved for public release; further dissemination unlimited. 
2

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Suppression of pool fires with HFC-125 in a simulated engine nacelle

J. C. Hewson∗

Fire Science and Technology

Sandia National Laboratories

Albuquerque, NM 87185-1135, USA

D. R. Keyser

INS, Inc.

48015-A Pine Hill Run Rd., Bldg. 1

Lexington Park, MD 20653, USA

Abstract

CFD simulations are conducted to predict the distribution of fire suppressant in an engine nacelle

and to predict the suppression of pool fires by the application of this suppressant. In the baseline configuration, which is based on an installed system, suppressant is injected through four nozzles at a rate fast

enough to suppress all simulated pool fires. Variations that reduce the mass of the suppression system

(reducing the impact of the suppression system on meeting mission needs) are considered, including a

reduction in the rate of suppressant injection, a reduction in the mass of suppressant and a reduction in

the number of nozzles. In general, these variations should work to reduce the effectiveness of the suppression system, but the CFD results point out certain changes that have negligible impact, at least for

the range of phenomena considered here. The results are compared with measurements where available.

Comparisons with suppressant measurements are reasonable. A series of twenty-three fire suppression

tests were conducted to check the predictions. The pre-test predictions were generally successful in

identifying the range of successful suppression tests. In two separate cases, each where one nozzle of the

suppression system was capped, the simulation results did indicate a failure to suppress for a condition

where the tests indicated successful suppression. When the test-suppressant discharge rate was reduced

by roughly 25%, the tests were in agreement with the predictions. That is, the simulations predict a

failure to suppress slightly before observed in these cases.

∗E-mail address: jchewso@sandia.gov

3
4
Executive summary

This report describes simulations of pool-fire suppression and fire suppressant concentrations in a prototype

engine nacelle ground simulator used for testing purposes. These simulations were conducted using the

Vulcan fire-physics simulator. Computational predictions have been compared with available measurements

of suppressant distribution and with actual fire-suppression tests in the engine nacelle simulator. The pretest

predictions were in agreement with the test results in the vast majority of cases. In some pretest simulations

and tests, specific nozzles were removed from the suppressant distribution system. In two unique cases, the

pretest predictions indicated a failure to suppress one step (25% change in suppressant injection rate) before

the actual test did, a conservative result for the predictions.

In general, two criteria need to be met to insure successful suppression. First, the rate of suppressant injection must be high enough that the average concentration in the nacelle exceeds the concentration required

to suppress fires. Second, the suppressant must be distributed with sufficient uniformity that the criteria for

suppression is met in all regions of the nacelle. The rate of injection required to meet the first criteria can be

identified, to a first approximation, with simple zonal models. However, the results presented in the present

study indicate that the first criteria is insufficient to guarantee suppression. The manner in which the suppressant is injected and distributed throughout the nacelle also matters. Poor distribution leads to substantial

inhomogeneities in the suppressant distribution that allow local burning leading to a reemergence of the fire

as the suppressant is dissipated. For the particular nacelle employed here, it is predicted (and the test results

agree) that the average mass fraction in the nacelle should exceed the cup burner value by 30-40% to account

for inhomogeneities. (This number varies according to details of the suppressant injection process.)

Evaluation of the simulation results suggests several general guidelines for fire suppressant distribution.

One of the most significant of these is that suppressant should be injected into or across the primary air inflow

when this inflow is substantial. A failure to inject suppressant into the primary air inflow is associated with

a substantial region with insufficient suppressant to extinguish a flame. This region is also associated with

rapid turbulent mixing that tends to enhance the burning rate if clutter is available to stabilize the flame.

5
6
Contents

1 Introduction 17

2 Terminology 19

I Suppressant distribution predictions 20

3 Key physics parameters in predicting suppression 20

4 An Overview of the Vulcan fire-physics code 23

5 Overall geometry 24

6 Flow field 25

6.1 Flow field without fire or suppressant injection . . . . . . . . . . . . . . . . . . . . . . . . . 26

6.2 Flow field with suppressant injection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26

7 HFC-125 concentration histories at specific locations 57

8 On the sufficient distribution of suppressant throughout the nacelle 59

9 Suppressant distribution without single nozzles 64

II Fire suppression predictions 70

10 Pool fires in the nacelle 73

10.1 Vulcan pool-fire simulation procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73

10.2 Nacelle test pool-fire procedures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

10.3 Pool evaporation rate results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 76

10.4 Nacelle temperatures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 78

10.5 Suppressant bottle test procedure and discharge rate . . . . . . . . . . . . . . . . . . . . . . 85

11 Suppression model 90

12 Mixing times for clutter 94

7
13 Suppression of pool fires: initial predictions 95

13.1 Results with four nozzle system . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95

13.2 Results with individual nozzles capped: inhomogeneities . . . . . . . . . . . . . . . . . . . 97

14 Suppression of pool fires: pre-test predictions and test results 98

14.1 Key results from pretest predictions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99

14.2 Test plan . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100

14.3 Summary of test results and corresponding pretest simulations . . . . . . . . . . . . . . . . 101

15 Summary and Conclusions 107

8
List of Tables

I Nozzle characteristics. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

II The locations of Halonizer probes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57

III Pool characteristics. Here “All” refers to pools B1, B2, B3C3 but not B5C5, the latter being

to difficult to reach in the course of the tests. . . . . . . . . . . . . . . . . . . . . . . . . . . 75

IV Fire suppression tests with an air-flow rate of 1.0 kg/s. . . . . . . . . . . . . . . . . . . . . 95

V Fire suppression tests with an air-flow rate of 1.0 kg/s and individual nozzles capped. . . . . 97

VI Summary of tests conducted. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 102

9
List of Figures

1 The maximum suppressant mass fraction, averaged over the nacelle, attained for the indicated dimensionless suppressant and air mass flux. . . . . . . . . . . . . . . . . . . . . . . 23

2 Velocity vectors (top) near the nacelle center, 0.63 m from the starboard reference point,

(middle) on the starboard side of the engine, 0.20 m from the starboard reference point and

(bottom) on the port side of the engine, 1.06 m from the starboard reference point. Air flow

into the nacelle is 0.66 kg/s. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

3 Contour plots of C2HF5 at 0.5 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 30

4 Contour plots of C2HF5 at 1.0 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 31

5 Contour plots of C2HF5 at 1.5 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 32

6 Contour plots of C2HF5 at 2.0 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 33

7 Contour plots of C2HF5 at 2.5 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 34

8 Contour plots of C2HF5 at 3.0 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 35

9 Contour plots of C2HF5 at 3.5 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 36

10 Contour plots of C2HF5 at 4.0 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 37

11 Contour plots of C2HF5 at 4.5 s after start of suppressant injection near the nacelle centerline, 0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or

1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom). 38

10
12 Contour plots of C2HF5 at 0.5 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39

13 Contour plots of C2HF5 at 1.0 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

14 Contour plots of C2HF5 at 1.5 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41

15 Contour plots of C2HF5 at 2.0 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42

16 Contour plots of C2HF5 at 2.5 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43

17 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44

18 Contour plots of C2HF5 at 3.5 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

19 Contour plots of C2HF5 at 4.0 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

11
20 Contour plots of C2HF5 at 4.5 s after start of suppressant injection on the starboard side of

the engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47

21 Contour plots of C2HF5 at 0.5 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

22 Contour plots of C2HF5 at 1.0 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49

23 Contour plots of C2HF5 at 1.5 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

24 Contour plots of C2HF5 at 2.0 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51

25 Contour plots of C2HF5 at 2.5 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52

26 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

27 Contour plots of C2HF5 at 3.5 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

12
28 Contour plots of C2HF5 at 4.0 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55

29 Contour plots of C2HF5 at 4.5 s after start of suppressant injection on the port side of the

engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s

(top) or 1.0 kg/s (center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4

s (bottom). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56

30 HFC-125 mole fractions at Halonizer sensor locations. Measurements during the qualification tests [25] are compared with simulations with 3.2 kg of HFC-125 injected over 3 s and

4 s. Air flow into the nacelle is 1.0 kg/s and suppressant mass flow is proportional to the

nozzle area for four production nozzles. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 60

31 HFC-125 mole fractions at Halonizer sensor locations. Measurements during the qualification tests [25] are compared with simulations with 3.2 kg of HFC-125 injected over 3 s and

4 s. Air flow into the nacelle is 1.0 kg/s and suppressant mass flow is proportional to the

nozzle area for four production nozzles. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61

32 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with

1.0 kg/s air flow into the nacelle. The mass of HFC-125 injected is 3.2 kg, injected over various periods as indicated with mass flow through each nozzle proportional to the measured

nozzle-throat area. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62

33 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 when

3.2 kg of HFC-125 is injected over 3 s. The air flow rate into the nacelle is 0.5, 0.66 or 1.0

kg/s as indicated. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63

34 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 when

3.2 kg of HFC-125 is injected over 6 s. The air flow rate into the nacelle is 0.5 or 1.0 kg/s

as indicated. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

35 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with

all nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg,

injected over 3 s. Air flow into the nacelle is 1.0 kg/s and suppressant mass flow through

active nozzles is proportional to the measured nozzle-throat area. . . . . . . . . . . . . . . . 65

36 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with

all nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg,

injected over 3 s. Air flow into the nacelle is 0.66 kg/s and suppressant mass flow through

active nozzles is proportional to the measured nozzle-throat area. . . . . . . . . . . . . . . . 66

13
37 Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with

all nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg,

injected over 3 s. Air flow into the nacelle is 0.5 kg/s and suppressant mass flow through

active nozzles is proportional to the measured nozzle-throat area. . . . . . . . . . . . . . . . 67

38 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side

(top), near the nacelle centerline and on the port side (bottom). Nozzle 3 has been capped

and suppressant flows through nozzles 1, 2 and 4 only. Air flow into the nacelle is 0.66 kg/s. 68

39 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side

(top), near the nacelle centerline and on the port side (bottom). Nozzle 4 has been capped

and suppressant flows through nozzles 1, 2 and 3 only. Air flow into the nacelle is 0.66 kg/s. 69

40 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side

(top), near the nacelle centerline and on the port side (bottom). Nozzle 1 has been capped

and suppressant flows through nozzles 2, 3 and 4 only. Air flow into the nacelle is 0.66 kg/s. 71

41 Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side

(top), near the nacelle centerline and on the port side (bottom). Nozzle 2 has been capped

and suppressant flows through nozzles 1, 3 and 4 only. Air flow into the nacelle is 0.66 kg/s. 72

42 Schematic representation of the relative locations of the pool fires and the suppressant injection nozzles showing the ribs and central longerons. Not to scale. . . . . . . . . . . . . . . . 75

43 Evaporation rates predicted in the simulations (closed symbols) and measured in the tests

(open symbols). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77

44 Surface plot showing cells where combustion is taking place for pool fires in locations B1

(top) and B2 (bottom). Surface indicates at least 1 ppm by volume combustion in a computational cell. Air flow into the nacelle is 1.0 kg/s. . . . . . . . . . . . . . . . . . . . . . . . 79

45 Surface plot showing cells where combustion is taking place for pool fires in locations B3C3

(top) and B5C5 (bottom). Surface indicates at least 1 ppm by volume combustion in a

computational cell. Air flow into the nacelle is 1.0 kg/s. . . . . . . . . . . . . . . . . . . . . 80

46 Temperatures at locations in the nacelle corresponding to the locations of channels 1 through

5 of the Halonizer probes are plotted as a function of time for case where air flow is 1 kg/s.

All pools are ignited and suppressant injection occurs at 20 s. . . . . . . . . . . . . . . . . . 81

47 Thermocouple measurements (uncorrected) along the lower nacelle for run 1d(25) are shown

in the upper figure as a function of the test time. The air flow is nominally 1 kg/s and the

lower figure shows the temporal variation in the measured air mass flow rate (in lb/s). All

pools are ignited prior to the start of data acquisition. . . . . . . . . . . . . . . . . . . . . . 83

14
48 Thermocouple measurements (uncorrected) along the lower nacelle for run 5b(25) are shown

in the upper figure as a function of the test time. The air flow is nominally 0.5 kg/s and the

lower figure shows the temporal variation in the measured air mass flow rate (in lb/s). All

pools are ignited prior to the start of data acquisition. . . . . . . . . . . . . . . . . . . . . . 84

49 Temperature contours at locations 15.16 m (top) and 15.55 m (bottom). The air flow rate is

0.5 kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple

temperature from Figs. 47 and 48 (in degrees K) is shown with white lettering. . . . . . . . 86

50 Temperature contours at locations 15.81 m (top) and 16.20 m (bottom). The air flow rate is

0.5 kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple

temperature from Figs. 47 and 48 (in degrees K) is shown with white lettering. . . . . . . . 87

51 Temperature contours at locations 16.59 m (top) and 17.01 m (bottom). The air flow rate is

0.5 kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple

temperature from Figs. 47 and 48 (in degrees K) is shown with white lettering. . . . . . . . 88

52 Temperature contours at locations 17.41 m (top) and 17.83 m (bottom). The air flow rate is

0.5 kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple

temperature from Figs. 47 and 48 (in degrees K) is shown with white lettering. . . . . . . . 89

53 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. The curves are labeled with the test

label to be given below in Table VI. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91

54 Suppressant mole (left) and mass (right) fractions required to extinguish flames for a given

turbulent mixing time scale. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93

55 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. . . . . . . . . . . . . . . . . . . . . . . . . . . 104

56 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. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105

15
16
1 Introduction

Suppression of fires in aircraft engine nacelles and other similar environments is challenging for a number

of reasons. To begin with, these environments are highly cluttered. The clutter both acts as a flame stabilizer

and impedes the transport of suppressant to this flame-stabilization region. High air-flow rates designed to

remove combustible gases from the nacelle also act to rapidly remove suppressant from the nacelle, limiting

the potential mixing time. Limits on the mass of suppression systems that can be carried while still allowing

aircraft mission success are a further constraint. All of these physical realities contribute to a relatively tight

set of requirements for successful fire suppression.

In the past, successful fire suppression systems have been developed through a combination of an understanding of these various issues and extensive testing. Detailed computational models, at the level of

computational fluid dynamics (CFD), have had limited success prior to this because of the complexity of the

nacelle environment and the relatively high degree of accuracy required. However, new model developments

and increased computational capabilities have moved CFD to a point where some level of prediction is now

possible.

Further, until their recent phase-out, chemicals like Halon 1301 (CF3Br) and 1202 (CF2Br2) were widely

employed to protect vehicles, structures and equipment from fire damage. Because of their high ozonedepleting potential, production of these fire-fighting agents has largely ceased, and research programs have

identified a number of other promising agents [10]. Takahashi et al. [28] and Hamins et al. [10] have

compiled data on the critical suppressant mole fraction required to suppress fires in various configurations

for Halon 1301 and other potential fire suppressants. It is reported that non-ozone depleting agents identified

with the most desirable toxicity, corrosion, stability, etc., characteristics generally require a greater agent

mass and/or volume, relative to Halon 1301, to suppress a fire. For example, typical mole fractions of Halon

1301 required to suppress fires are between 3 and 4% while HFC-125 (C2HF5) requires between 8 and 9%.

The mass of agent required is similarly greater, with roughly double the mass of HFC-125 required relative

to that for Halon 1301. For applications where the fire suppression-system mass is a key design criterion, as

it is in aircraft, a need to optimize the agent delivery system has been identified to minimize the additional

suppression-system mass. In this work, the use of HFC-125 as a fire suppressant in a simulated aircraft

engine nacelle will be studied. The ability of a CFD model to predict the distribution of suppressants will be

assessed (Part 1) along with its ability to predict the suppression of pool fires (Part 2). The results discussed

herein are focused on suppressants with boiling temperatures well below operating temperatures, such as

HFC-125 (C2HF5). The focus here is on pool fires [14, 15]. Spray fires are another significant concern,

and recent reviews can be consulted for these scenarios [9, 11]. In addition to engine nacelles, it is vital to

also suppress fires in aircraft dry bays [2]. Other work has addressed issues related to higher-boiling point

agents [5, 31, 32].

17
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 (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 [8].

It is important to point out the relationship of the present work to other levels of modeling. In the

fire area, so-called zonal models are frequently employed. While zonal models, treating the volume as a

homogeneous mixture, can predict average suppressant concentrations, the suppressant inhomogeneities are

important because the local areas that are deficient in suppressant can stabilize a fire until the suppressant

has been depleted. The present CFD models are fully capable of describing the variations of the local

suppressant concentrations throughout the nacelle. In Sec. 3 a simple zonal model relevant to the present

aircraft nacelle application is presented. This is provided to help identify the parameter space in which the

nacelle is operating and also to help estimate the nacelle-averaged characteristic conditions. In Sec. 6 the

detailed CFD results are presented showing the local flow field. Then in Sec. 8 more global measures of

the CFD results are presented to compare with the zonal models and indicate the level of inhomogeneity

present.

In Part 2 of the report, simulations of pool fires and the attempted suppression of those fires are presented.

These are presented along with the results of a test program that was conducted at the end of this part of the

NGP program. All of the simulations presented in this report were pre-test simulations. While there was

remarkable agreement between the pre-test simulations and the test results, the careful analysis of the results

does indicate some challenges that are faced. Notable among these was the challenge in correctly predicting

the pool evaporation rate. At the same time, certain experimental configurations that were suggested as

challenging by the simulations could not be realized in the tests due to physical limitations of the nacelle

simulator hardware.

Because of the complex, cluttered nature of the aircraft nacelle environment, it is not expected that CFD

models will provide truly predictive simulations like those possible in relatively clean environment with

18
high-end computational submodels. This indicates that a significant experimental test program will be a

core part of the suppression system design and qualification process for the foreseeable future. However, the

level of success described here indicates that CFD-based methods are coming into a period when they can be

successfully used in conjunction with tests. Of particular interest in the design of optimized fire-suppression

systems is the use of simulations to explore the parameter space of both system design and system test. The

CFD model can be used to iterate over the potential design space to identify designs that provide uniform

protection against fires, and yet reduce the system mass or improve upon some other metric. In the present

report, the CFD model was used to suggest what nozzles might be removed with little detrimental effect,

potentially reducing the mass of piping required. The CFD model is also used to identify the insensitivity

of the suppression results to a somewhat reduced total mass of suppressant carried (although the ability to

protect against the relight of fires is a concern that is not addressed here). The CFD model is also used to

identify regions where relatively robust pool fires can be stabilized, and in this manner could identify the

most challenging tests for a particular fire suppression system. Such tests would provide a good measure of

the success of the fire suppression system.

The cost of setting up a modeling and simulation capability versus the cost of setting up an extensive

test program are both nontrivial, and must be taken into account when considering the approach to be taken.

Despite this, there is the potential for improved designs providing better performance (aircraft lifetime cost

savings), significant cost savings in the design process, or potentially both. The cost of setting up a test-bed,

such as the nacelle simulator described in this report is significant. The cost of running a single experiment is

small relative to the cost of the original facility, yet still substantial. Similarly the cost of setting up an initial

simulation is significant, involving detailed geometry specifications and grid-refinement analyses. The cost

of an additional simulation with slightly differing parameters is small with commodity processors available.

As a guide to costs, the initial simulation may require on the order of hundreds of hours to set up while

additional simulation costs are roughly two orders of magnitude less in terms of man-hours. Computational

requirements are also greater for the initial set up because a large number of test-simulations are involved in

checking the appropriateness of the simulation grid, input parameters, etc. Furthermore, the linking of CFD

methods with optimization-based search methods [7] provides a nearly automated manner of searching for

best possible designs. The result of such a search may point the direction toward the best designs to meet a

certain objective.

2 Terminology

The physical device that is being computationally simulated is generally referred to as the “ground-based

simulator,” and this is to be differentiated from computational simulations. Therefore, here we specify

that further references to the “simulator” indicate the physical device located in the laboratory while the

19

“simulation” indicates the computational representation of the simulator.
Part I

Suppressant distribution predictions

The following sections address predictions of the distribution of fire suppressant in the nacelle simulator.

The limited available data for suppressant distribution is compared with predictions where possible.

3 Key physics parameters in predicting suppression

Successful total-flooding-based suppression relies on providing a sufficient mass fraction of suppressant to

all possible fire locations. There are thus two key issues: First, supply a sufficient overall quantity of suppressant. Second, ensure that the suppressant is sufficiently distributed throughout the nacelle. While the

present paper focuses on nacelles, the results in the present section are for relevant to ventilated compartments in general.

In a ventilated compartment, the provision of sufficient suppressant is hampered by the loss of suppressant through whatever ventilation means are available. This section presents a highly simplified model to

describe the mean mass fraction of agent in a ventilated nacelle where the composition of gases in the nacelle

is assumed to be homogeneous. This analysis identifies two key parameters that affect the overall suppressant mass fraction. These parameters are the ratio of the suppressant injection rate to the total injection rate

and the rate of volume displacement by injected fluids per unit nacelle volume.

This simplified inflow-outflow model considers a chamber into which mass fluxes of air and suppressant,

denoted m ̇ a and m ̇ s, respectively, are possible. Other fluid sources, like fuel or combustion products, can be

added to the air mass source as required. In the simulations conducted here, the additional source attributed

to the fuel is at least an order of magnitude smaller than the air. For the purpose of the analysis, the mass

fluxes represent known values. We assume that the compartment is sufficiently vented that the pressure

will remain constant despite the mass influx. This implies that as mass is added to the system, mass must

also leave the system. Because the density may vary with time, the mass that leaves the system will be

determined by the volume displaced by the inflowing mass.

To make the analysis tractable, we assume the system to be completely homogeneous so that the density

and mass fractions throughout are uniform. Given the pressure, temperature and mean molecular weight,

this mean density is ρ = PW /RT. The mean molecular weight is expressible in terms of the mass fractions

of the two components, Yi, as W = 1/(

PYj/Wj ).

20
suppressant. Multiplying the volume displaced by the mean density gives the mass flow rate of the mixture

leaving the chamber

m ̇ out = ρ



m ̇ a

ρa

+

m ̇ s

ρs



(1)

where ρa and ρs are the air and suppressant densities.

Given the above assumptions, conservation equations for the mean suppressant mass fraction in the

chamber is

dYs

dt =

m ̇ s

ρVT

−



m ̇ a

ρa

+

m ̇ s

ρs



Ys

VT

. (2)

where VT is the chamber volume. A differential equation for the air (and all other gases, which are lumped

with air) is not required since it can be found through overall mass conservation, PYi = 1.

It is convenient to normalize the mass flux for each component with the mass if the chamber were filled

with that single component

μs =

m ̇ s

ρsVT

μa =

m ̇ a

ρaVT

. (3)

These have dimensions of rate (one over time) and indicate how fast the air or suppressant would displace a

volume of fluid VT . Using these with

ρi

ρ

=

X

j

WiYj

Wj

=

X

j

Yjrij (4)

where rij is the ratio of the molecular weights of component i to component j, Eq. 2 is written

dYs

dt = μsrsa − (μsrsa + μa)Ys. (5)

This first-order ordinary differential equation can be solved for the relevant scenario where μa = μ

0

a

, and μs

is zero except for the time interval 0 < t < te during which μs = μ

0

s

. Because of the step-function nature

of μs the solution differs for t < te and t > te. For 0 < t < te the suppressant mass fraction starts at zero

and exponentially approaches a steady state value of μ

0

s

rsa/(μ

0

s

rsa + μ

0

a

) as described by

Ys =

μ

0

s

rsa

μ0

s

rsa + μ0

a

h

1 − e

−(μ

0

s

rsa+μ

0

a)t

i

. (6)

The time constant for the approach to the steady-state value is 1/(μ

0

s

rsa + μ

0

a

). Note that this shows that

the steady-state value is reached more rapidly for higher air mass flow rates but the steady-state value is

correspondingly lower. The rate of change (time-derivative) of Ys is, however, directly proportional to the

suppressant mass flow rate and not the air flow rate.

21

The rate at which gases leave the system is determined by the volume displaced by the inflowing air and
in Eq. 6 according to

Ys =

μ

0

s

rsa

μ0

s

rsa + μ0

a

h

1 − e

−(μ

0

s

rsa+μ

0

a)te

i

e

−μ

0

a(t−te)

. (7)

With the suppressant inflow turned off, the time constant for the decay of the suppressant mass fraction

is 1/μ0

a

and the rate of decay is proportional to μ

0

a

and the value of Eq. 6 evaluated at t = te, which is

represented by the terms on the right-hand side of Eq. 7 prior to the last exponential.

This indicates two key parameters to be considered when designing a suppression system. The ratio of

suppressant mass inflow to the total mass inflow gives the steady-state suppressant mass fraction

Ys,st =

μ

0

s

rsa

μ0

s

rsa + μ0

a

, (8)

corresponding to continuous suppressant injection. The product of the time constant appearing in the exponential Eq. 6 and the duration of the injection, τ = (μ

0

s

rsa + μ

0

a

)te indicates how close one comes to

this steady state mass fraction. After an injection period of roughly 1/(μ

0

s

rsa + μ

0

a

), (1 − 1/e) (roughly

two-thirds) of the steady-state value is reached, and roughly eighty-five percent of the steady-state value

is reached after twice that period. Injection for longer periods serves primarily to increase the residence

time for the highest suppressant concentrations. The maximum averaged suppressant mass fraction is thus

estimated to be

Ys,max =

μ

0

s

rsa

μ0

s

rsa + μ0

a

h

1 − e

−(μ

0

s

rsa+μ

0

a)te

i

. (9)

This is presented graphically in Fig. 1 for the dimensionless variables μste, which represents the mass of

suppressant injected over a time period te normalized by the total mass that would be required to fill the

nacelle, and μate, which represents the air inflow during the period of injection normalized by the mass

of air that fills the nacelle. For the 3.2 kg of HFC-125 in the current nacelle at ambient temperatures,

μste ≈ 0.5 and is represented by the center curve in Fig. 1. The upper and lower curves indicate a 50%

increase or decrease in the suppressant mass. For the conditions considered here, the typical values for

μate are 1 to 2. The range of values on the abscissa show the consequences of changing the duration of

the injection. Note that average temperatures above ambient will increase μi proportionally because μiis

normalized by the average density. This will reduce the time constant for filling and flushing the nacelle,

tending to cause Ys,max to more closely approach Ys,st.

This simple model does not account for any spatial fluctuations that may lead to local conditions where

the fire can be sustained. Prediction of spatial fluctuations is only possible at the CFD level, accounting for

the details of the geometry and boundary conditions. The most that can be said is that higher fluid mixing

rates will generally lead to more uniform suppressant distribution. As will be seen in later sections, though,

the collocation of high mixing rates with the suppressant mixing process is not guaranteed.

22

For t > te the suppressant mass fraction decays exponentially from whatever value ita ttained at t = te
μ

0

a

te

Y

s,m

a

x

0 2 4 6

0

0.2

0.4

0.6

0.8

1

μ

0

s

te= 0.25

μ

0

s

te= 0.5

μ

0

s

te= 0.75

Figure 1: The maximum suppressant mass fraction, averaged over the nacelle, attained for the indicated

dimensionless suppressant and air mass flux.

4 An Overview of the Vulcan fire-physics code

Vulcan has been developed over the past eight years at Sandia National Laboratories (SNL) in collaboration

with the SINTEF Foundation and the Norwegian University of Science and Technology. Vulcan is derived

from the Kameleon Fire model [17] and uses an extension of the SIMPLEC method [27] to solve the conservation equations on a structured, staggered, three-dimensional Cartesian, finite-volume grid. The ability

to resolve the geometry of the system is only limited by the ability to construct the appropriate grid with the

Cartesian grid generator available in Vulcan. First- and second-order upwind schemes can be used for the

convective terms. Turbulence is modeled using a standard two equation k- model [18]. The Vulcan code

has the capability of performing fire simulations, and the combustion simulation is based on Magnussen’s

Eddy Dissipation Concept (EDC) assuming mixing-limited combustion [17, 24]. Models of soot generation and oxidation are also included. Radiation transport is conducted using the discrete-transfer model of

Shah [22]. The suppression model employed in Vulcan is discussed in Sec. 11

23
5 Overall geometry

The engine-nacelle simulations have been designed in accordance with test fixture drawings of the nacelle

simulator [19] and in accordance with visual inspections of the simulator, located at Patuxent River Naval

Air Station. The nacelle simulator is based roughly on the United States Navy’s F-18, which has two engines

located side by side in the fuselage. The ground simulator used in the fire tests simulated here is a mock

up of the port engine nacelle. Therefore, the port side corresponds to the outboard side of the engine and

the starboard side corresponds to the inboard side of the engine. The nacelle is roughly 3.18 m (125 in.)

long. All distances in the text of the present report refer to distances from the front of the nacelle, which is

approximately 15 m behind the nose of the aircraft. In the figures, however, the y-axis is measured from the

reference point at the nose of the aircraft in concurrence with the drawing specifications.

The width and height vary significantly along the length, but are contained within a region 1.45 m high

and 1.15 m wide. Coordinates are measured from a reference point on the top starboard side of the simulator

test fixture so that all vertical coordinates are negative while other coordinates are positive. In conjunction

with the rearward taper of the airframe, the nacelle is generally tallest and widest near the forward end.

At various sections, this narrowing is countered by an abrupt increase in distance of the outer nacelle skin

from the engine on the sides and top. The lower nacelle surface slopes smoothly up from the forward to

the aft, crossed by five ribs of varying heights. There are roughly four longerons running along the lower

nacelle surface. The two central longerons end about 0.6 m before the aft end to provide space for the lower

diamond vent.

The engine, a hollow cylinder in the ground simulator, extends from the forward to the aft of the nacelle.

The engine diameter is nominally 0.77 m with a slight narrowing over the front 0.61 m to a minimum

diameter of 0.62 m. From 0.81 m aft to the aft end of the nacelle, the engine is uniformly 0.77 m in

diameter. The outer surface of the engine is generally smooth, although various later-described clutter items

are attached to it.

In addition to the ribs and longerons, the most significant clutter is located near the forward end of the

nacelle. It is dominated by a “gear-box” assembly located from 0.1 m to 0.75 m behind the forward end

and mostly below the engine. The gear-box assembly is a collection of several generally parallelepiped

and cylindrical objects of varying sizes that obstruct the flow in the large space between the engine and the

lower nacelle in the vicinity of the air inlet scoop. Additional large parallelepiped clutter objects are located

along the engine, particularly on the port side as far back as 1.5 m behind the forward end. Smaller clutter,

primarily tubing and wire bundles, exists along the lower half of the engine across the length of the nacelle.

Some of this smaller clutter has been identified as a potential relight surface because of the relatively low

thermal capacity and rapid heating to very high temperatures of smaller clutter [25]. The afterburner control

vanes at the aft end of the nacelle have not been included in the computational model because the setup

24

(initially) focused on fire scenarios near the forward end of the nacelle.
The grid employed in the simulations is coarse, with a resolution of one inch (2.54 cm control volume

edges). The total grid size is then roughly 320,000 control volumes, but the majority of these are not in

relevant regions (outside the nacelle or inside the engine core); a more efficient grid can be obtained using

an unstructured grid as with Sandia’s Fuego code. The coarse grid used here is selected specifically because

the clutter in the nacelle is not to be resolved, but is to be treated with a clutter model [4]. Unfortunately, at

the time the simulations were initiated, the clutter model treatment of near wall clutter (the most significant

clutter in this nacelle simulator) was not correctly implemented and thus the clutter model was turned off.

The clutter model has since been correctly implemented in Vulcan [23]. No rigorous grid resolution studies

were conducted during the course of this study for the nacelle simulator geometry, but it is expected that

finer grid resolution will change at least the details of the results. Because the coarseness of the resolution is

relatively uniform with respect to the flow dynamics, it is expected that the general flow characteristics will

not change dramatically.

6 Flow field

In the absence of fires and suppressant flow, the flow field in the nacelle is dominated by flow through an

inlet air scoop located near the front of the nacelle coming up through the nacelle bottom. The inlet air flows

from the inlet across the engine toward the upper aft of the nacelle. For the purposes of the simulations, air

mass flow rates of 0.33, 0.5, 0.66 and 1.0 kg/s have been considered. These air flow rates cover the range

(0.57, 0.68 and 0.95 kg/s) described in Ref. [6] that were obtained from F-18 C/D flight test data and scaled

up for the E/F aircraft [6]. It is noted that these flow rates can occur over varying altitudes and external air

temperatures so that the corresponding air densities may differ somewhat. For the present simulations, only

expected ambient conditions at the Patuxent River NAS test facility are considered; these are assumed to

be 101000 Pa pressure and 293 K temperature. A sensitivity analysis with respect to air density would be

straightforward with the Vulcan CFD code, but is not performed here.

The inlet air mass flow rates determine the air-exchange times. The air-exchange time is the time over

which the mass of air entering and leaving the nacelle equals the mass of air in the nacelle. For the simulation, the internal volume of the nacelle, neglecting clutter objects, is just over 1.4 m3

. For the range of

air-flow rates (0.33 to 1.0 kg/s) at ambient densities, this gives an air-exchange time of 5.4 to 1.6 s. The air

exchange time is equivalent to 1/μ0

a

in the nomenclature of Sec. 3. Shorter air-exchange times correspond

to higher mass flow rates. Similarly, for a fixed mass flow rate and a lower density, the air exchange time

would be proportionately lower. This provides an indication of the effects of operation at different altitudes

where the ambient density differs. Thus, the change in air-exchange time associated with lower ambient

densities can be simulated on the ground with higher inlet mass flow rates.

25
The inlet cross section is approximately 0.008 m2

and air-flow velocities considered range from 50 to

150 m/s. This results in an inlet Reynolds number, with the length scale defined as the square root of the

inlet cross section area, ranging from 3 × 105to 9 × 105. The inlet air flow is thus strongly turbulent even

without the clutter-enhanced mixing.

6.1 Flow field without fire or suppressant injection

In the absence of suppressant injection, the flow field is dominated by the inlet air scoop located on the

nacelle bottom near the forward end of the nacelle and directed toward the nacelle aft, but also slightly

upward and toward the starboard side, the latter being the centerline of the aircraft. The inlet air flows

upward onto the engine and up around the starboard side of the engine. This creates a large circulating flow

on the starboard side of the engine. The flow above the engine in the nacelle is primarily in the forward

direction, and much of this flow moves down around the engine toward the forward end of the nacelle and

is entrained by the inlet air flow. On the port side of the engine, the flow is observed to be primarily in the

forward direction, similar to the flow along the top of the nacelle. Figure 2 demonstrates some of these flow

patters with velocity vector plots.

Air leaves the nacelle to balance inflow through several vents and openings. The relative flow through

the various paths was measured and found to be independent of air-flow rate [6,20] and the relative flow rates

measured there have been reproduced in the present simulations to within a few percent. The majority of the

air (65%) leaves the nacelle through the upper diamond vent located in the top of the nacelle 2.85 m behind

the nacelle front. Another 10% of the air leaves through the balance piston, centered on the nacelle top 0.75

m from the nacelle front. The remaining 25% of the air leaves through four openings in the nacelle front

simulating connections to the AMAD bay and other parts of the aircraft. In the ground simulator, the flow

out the lower diamond vent, located directly below the upper diamond vent, is effectively zero; this may be

because the fine mesh screen covering the vent was clogged with soot, rust and other particulate [6,20]. The

flow rate out of the various openings was matched to the measured flow rates [6,20] by varying the porosity

of the opening, given the known area of the opening. It would be possible to equate the lower diamond vent

porosity to that of the upper diamond vent to ascertain any changes in the flow patterns that might occur in

fleet aircraft (not subject to the repeated fires in the ground test simulator) with clean diamond vents, but

this has not been done.

6.2 Flow field with suppressant injection

The nacelle simulator was outfitted with a set of four suppressant nozzles that were found to be sufficient to

suppress a fire when 3.2 kg (7 lbs.) of suppressant were discharged [25]. These are denoted by the numbers

1 through 4 with larger numbers corresponding to nozzles that are further toward the rear of the nacelle.

26
y

z

15.5 16 16.5 17 17.5 18

-1

-0.5

0

Vector Magnitude 10 m/s

y

z

15.5 16 16.5 17 17.5 18

-1

-0.5

0

Vector Magnitude 10 m/s

y

z

15.5 16 16.5 17 17.5 18

-1

-0.5

0

Vector Magnitude 10 m/s

Figure 2: Velocity vectors (top) near the nacelle center, 0.63 m from the starboard reference point, (middle)

on the starboard side of the engine, 0.20 m from the starboard reference point and (bottom) on the port side

of the engine, 1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s.

27
Table I: Nozzle characteristics.

Orifice Nozzle location

Nozzle diameter [cm] [25] x [m] y [m] z [m]

1 .396 .09 15.78 -1.02

2 .475 .09 15.84 -1.15

3 .475 .19 16.82 -.41

4 .330 1.03 17.70 -.46

Simulations have been conducted for suppressant injection through these nozzles into the nacelle in the

absence of a fire to obtain information on the suppressant distribution through the nacelle. The suppressant

was assumed to enter the nacelle in the vapor phase or to vaporize fast relative to other time scales. Except

where noted the net suppressant mass flux into the nacelle is 1.06 kg/s so that the injection duration for the

standard 3.2 kg of HFC-125 was 3 s. This fixed mass of suppressant injected corresponds to μ

0

s

te ≈ 0.5 in

the nomenclature of Sec. 3. This 3 s duration for the injection was chosen because the available HFC-125

concentration measurements showed an approximately 3 s rise time (see Figs. 30 and 31). The injection

rate was taken to be constant during the duration of the injection; this is known to be a gross simplification,

but better information was not available at the time of the simulations. The suppressant mass flux for each

nozzle was assumed to be proportional to the nozzle orifice area.

The mass fraction of HFC-125 in vertical planes near the nacelle center and near the starboard and port

sides are shown in Figs. 3 through 29 at half-second intervals. The mass fractions are colored in such a

manner that cool colors indicate likely suppression and hot colors indicate that suppression is unlikely. The

critical mass fraction for suppression in the absence of intense mixing is somewhat less than 0.3 (the cupburner suppressant mass fraction is 0.28 and strained counterflow flames extinguished with a mass fraction

of 0.16 [9], so that this is a conservative estimate), and so the dividing line for likely suppression can be

taken to be the dividing line between yellow-green and green. Depending on various conditions, suppression

may occur for lower mass fractions of HFC-125. Velocity vectors are also provided in Figs. 3 through 29 to

indicate the general flow direction.

Nozzle 1 is located on the starboard side of the nacelle in the forward section and is directed upward.

The suppressant from nozzle 1 is observed to initially fill the upper forward nacelle and is advected forward

and into the lower forward nacelle. The high suppressant mass fractions in the upper forward sections of

the nacelle visible in Figs. 14, 16, 23 and 25 arise primarily from nozzle 1. Note that some suppressant

from nozzle 1 appears to leave the nacelle soon after injection from the upper balance piston vent in Figs. 3

through 7

Nozzle 2 is located near nozzle 1 on the starboard side of the nacelle and directed to the port side

28
across the top of the inlet air flow. Suppressant from nozzle 2 is spread rearward by the inlet air along the

engine. There is some uncertainty in the momentum of the suppressant flow, but with the current uniform

3 s discharge the momentum is sufficient that more suppressant flows along the port side of the inlet jet

and along the lower port side of the engine. The high suppressant mass fraction along the lower port side

forward of the 16 m point in Figs. 5 and 7 comes from nozzle 2. If the suppressant jet momentum has been

grossly overestimated, the suppressant may not cross the jet and would flow along the lower starboard side

of the engine. There is reason to believe that late in the suppressant injection period as the bottle pressure

drops that the suppressant momentum is less than currently described.

Nozzle 3 is located just past the mid-point of the nacelle, in between the second and third ribs, on the

upper starboard side of the engine. Suppressant flow is directed at roughly a 45 degree angle upward across

the engine and into the recessed area at the top of the nacelle. The majority of the suppressant mixes in this

recessed area with flow moving forward and is transported gradually along the upper nacelle in the forward

direction as evidenced by the high suppressant mass fractions along the forward half of the upper nacelle in

Figs. 3 through 7. A significant minority of the suppressant from nozzle 3 leaves the upper recessed region

on the port side of the engine. In this region the flow coming from the inlet air is rearward and this advects

the suppressant into the regions on the port side near the fourth rib. However, the suppressant transported in

this direction is sufficiently diluted that it does not show up on the contour plots.

Nozzle 4 is located on the port side of the engine near the aft of the nacelle just behind the fifth rib.

The suppressant is directed downward along the side of the engine as shown in Figs. 21 through 25. This

flow is counter to the general upward flow toward the upper diamond vent and so the suppressant is split

between suppressant that goes down and around the engine in the lower nacelle area (Figs. 3 through 7) and

suppressant that flows rearward and upward to fill the area near the afterburner control actuators. Forward

transport of suppressant is inhibited by the overall aft-ward flow from the inlet to the upper diamond vent.

The large sixth rib in the lower nacelle just forward of nozzle 4 and also the large fifth rib significantly

restrict forward transport of agent past these planes because they constrict the flow leading to relatively high

aft-ward velocities.

29
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 3: Contour plots of C2HF5 at 0.5 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

30
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 4: Contour plots of C2HF5 at 1.0 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

31
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 5: Contour plots of C2HF5 at 1.5 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

32
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 6: Contour plots of C2HF5 at 2.0 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

33
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 7: Contour plots of C2HF5 at 2.5 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

34
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 8: Contour plots of C2HF5 at 3.0 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

35
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 9: Contour plots of C2HF5 at 3.5 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

36
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 10: Contour plots of C2HF5 at 4.0 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

37
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 11: Contour plots of C2HF5 at 4.5 s after start of suppressant injection near the nacelle centerline,

0.63 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

38
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 12: Contour plots of C2HF5 at 0.5 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

39
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 13: Contour plots of C2HF5 at 1.0 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

40
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 14: Contour plots of C2HF5 at 1.5 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

41
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 15: Contour plots of C2HF5 at 2.0 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

42
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 16: Contour plots of C2HF5 at 2.5 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

43
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 17: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

44
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 18: Contour plots of C2HF5 at 3.5 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

45
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 19: Contour plots of C2HF5 at 4.0 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

46
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 20: Contour plots of C2HF5 at 4.5 s after start of suppressant injection on the starboard side of the

engine, 0.20 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s

(center, bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

47
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 21: Contour plots of C2HF5 at 0.5 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

48
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 22: Contour plots of C2HF5 at 1.0 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

49
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 23: Contour plots of C2HF5 at 1.5 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

50
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 24: Contour plots of C2HF5 at 2.0 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

51
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 25: Contour plots of C2HF5 at 2.5 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

52
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 26: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

53
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 27: Contour plots of C2HF5 at 3.5 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

54
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 28: Contour plots of C2HF5 at 4.0 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

55
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 29: Contour plots of C2HF5 at 4.5 s after start of suppressant injection on the port side of the engine,

1.06 m from the starboard reference point. Air flow into the nacelle is 0.66 kg/s (top) or 1.0 kg/s (center,

bottom). Suppressant flow is 3.2 kg over 3 s (top, center) or over 4 s (bottom).

56
Table II: The locations of Halonizer probes

Qualification Report [25] Simulation

Channel y [m] x [m] y [m] z [m]

1 15.27 7:00 .84 15.27 -1.07

2 15.29 12:00 .60 15.29 -.23

3 16.35 3:00 .15 16.35 -.65

4 16.35 12:00 .58 16.35 -.20

5 17.26 2:30 .17 17.26 -.46

6 17.26 6:00 .60 17.26 -1.18

7 17.26 10:30 1.00 17.26 -.46

8 17.98 12:00 .60 17.98 -.20

9 17.98 4:30 .25 17.98 -.97

10 17.98 7:30 .95 17.98 -.97

11 15.63 7:30 .82 15.63 -1.02

12 15.81 7:30 .98 15.81 -1.05

7 HFC-125 concentration histories at specific locations

To evaluate the predictive capability of the simulations in describing the suppressant distribution in the nacelle, we have compared HFC-125 concentration-time traces with those measured using a Halonizer in the

F-18 qualification tests [20, 25]. These comparisons are shown in Figs. 30 and 31. There is an unknown

uncertainty associated with the Halonizer concentration measurements, both in terms of the absolute concentration measured and, significantly, in the locations where the measurements were taken. The indications

from the qualification test report are that the measurements are fairly repeatable so that significant errors are

most likely of the systematic type. The probe locations are not specified in detail in the report; for example,

the probe corresponding to channel 1 is identified as being 604 inches aft of the aircraft nose and at 7:00.

This places it roughly in front of the upper port edge of the gearbox assembly, but there is some latitude

on the exact placement. The stated locations and the coordinates used in the simulations are indicated in

Table II. Each probe location within the simulation represents an average over one cubic inch of volume.

While the agreement described in the following paragraphs is reasonable, it also highlights model sensitivities in the areas of (1) the suppressant spray momentum vector direction and magnitude relative to the

air inflow and (2) the rate of suppressant vaporization. To indicate the effect of the relative momentum,

simulations are conducted with the 3.2 kg of HFC-125 injected in the gas phase over either a 3 or 4 s injection period; this represents a 30% change in the jet momentum. It is evident in Figs. 30 and 31 that this

modest change in the suppressant momentum can lead to as much as a 50% change in the concentration

predicted at a specific point. This substantial concentration variation arises because the high-concentration

suppressant-jet centerlines either hit a measurement location or are shifted by perhaps 10 cm to the side of

57
the measurement location. This highlights both the sensitivity to the spray momentum and the sensitivity

to the Halonizer-probe locations. Because of this enormous sensitivity, we have only moderate confidence

in the significance of agreement or disagreement in these results. More significant are the general trends

discussed below.

In these simulations, the suppressant has been assumed to enter the nacelle in the vapor phase or instantly

vaporize upon entering. Other simulations, not reported here, indicate that there is a finite vaporization time.

This actual vaporization time will depend on the square of the initial particle diameters, assuming a d

2

-

vaporization rate, but reasonable droplet diameters (100 to 500 μm) have estimated evaporation times of 0.1

to 2.5 s. These estimates assume flash vaporization of roughly a third of the suppressant acts to substantially

cool the remaining liquid suppressant to its boiling point. While relatively few large droplets are expected,

this highlights the need to understand the droplet size distribution leaving the suppressant nozzles. The

effects of finite-rate evaporation are clear in the qualification test measurements in Figs. 30 and 31 where

the suppressant concentrations drop off more gradually than simulated after the end of the 3 or 4 s injection.

The concentrations also rise more gradually than predicted.

The various Halonizer measurement channels have been grouped roughly according to their locations

in the nacelle. In the upper pane of Fig. 30 are measurements located in the upper forward quadrant of the

nacelle, channels 2, 3 and 4. Channel 2 is near the front of the nacelle while 3 and 4 are almost half way aft.

Channels 2 and 4 are located at 12:00, along the top of the nacelle, while 3 is located on the starboard side.

Results for channels 3 and 4 are reasonably good, with 3 and 4 s injection periods roughly bounding the

measurements. Predictions for channel 2 are consistently high and this implies that much more suppressant

is predicted to move to the forward nacelle than is measured. This implies that the flow from nozzle 1 (up

the starboard side) should be turned forward less than shown in Figs. 3 through 29. The reason that this

would occur is unclear, but it could imply that much more suppressant goes out the balance piston or that

the dominant flow path is circumferentially around the engine where it narrows at 15.6 m.

In the lower pane of Fig. 30 are the Halonizer channels that are located in the lower forward quadrant of

the nacelle with channel 1 in front of and channels 11 and 12 behind the gearbox assembly and over the air

inlet. Suppressant in this region comes primarily from nozzle 1 (over the top of the engine) and from nozzle

2 (under the engine). Measurements for all three channels here bounded by the 3 and 4 s injections with the

shorter, higher momentum injection giving the higher concentrations since it is better able to cross the air

inlet and reach the probe locations without being dispersed downstream. For channel 12 the concentration

differences for the two suppressant injection rates exceed a factor of two.

Because the turbulence generated by the suppressant injection is also not well known, we have also

varied this parametrically to determine its sensitivity. The concentrations measured by probe channels 11 and

12 do show some sensitivity to turbulence levels generated by the spray nozzles: a two-order of magnitude

increase in the level of turbulence in the initial spray (3 s injection) brings the concentration profiles for

58
channels 11 and 12 into good agreement with the Halonizer measurements. These are the only results

reported that show more than a few percent variation with a substantial change in the suppressant-injection

turbulence levels; we conclude that the sensitivity to the injection momentum and the turbulence generated

as the suppressant mixes is more substantial than the sensitivity to the turbulence imposed as a boundary

condition.

In the upper pane of Fig. 31 are sensors located between nozzles 3 and 4, roughly two-thirds of the

way aft. Channel 6 has been located below the engine and just in front of one large rib while channels 5

and 7 are located on the upper starboard and port sides respectively. Simulated results for channels 5 and

6 are relatively insensitive to the rate of injection and these simulations bound the measurements for those

channels. Results for channel 7, which is located just ahead of nozzle 4, show that as the momentum of

nozzle 4 is reduced there is a competition between the downward momentum from the suppressant injection

and overall upward flow caused by the flow from the air inlet to the upper diamond vent. This creates strong

oscillations in the flow field for the 4 s duration injection that are visible Figs 21 through 29 as well as in

Fig. 31. The relatively high suppressant concentrations measured for channel 7 with the Halonizer imply

that some (steady) forward flow from nozzle 4 is occurring in the simulator that is not picked up for these

conditions in the simulations (or that nozzle 4 is directed in a forward direction by a few degrees).

Certain probes (channels 8, 9, and 10) were listed in the qualification test report [25] with locations

beyond the aft end of the aircraft; we have located these in the afterburner vane region where there were

sensor ports in the fuselage that did not correspond to listed probe locations. The Halonizer measurements

shown in the lower pane of Fig. 31 correspond to these channels. The predictions are substantially lower

because we have neglected to model the afterburner control vanes that are in the vicinity of these probes;

such vanes would act to better mix the flow in this region.

We note that the discrepancies noted above in the measured concentrations do not indicate a change in

state from one where suppression is expected to one where suppression is not expected, or vice versa. When

the simulations or Halonizer probe measurements differ noticeably, one shows a concentration that is sufficient to suppress a fire while the other shows a concentration markedly higher than the mean concentration,

indicating that the suppressant spray is still highly concentrated when it reaches that probe location.

8 On the sufficient distribution of suppressant throughout the nacelle

As an integral measure of the sufficiency of suppressant distribution in the nacelle, we have tracked the

fraction of the nacelle volume for which the suppressant mass fraction exceed 30%. Even with low intensity

mixing, this is a generally sufficient level for suppression and represents a conservative estimate of suppression ability. Figure 32 shows this fraction when the air flow rate is 1.0 kg/s with varying rates of suppressant

injection (same mass over different durations). The mass of suppressant to be injected was fixed at 3.2 kg

59
time [s]

H

F

C

-

1

2

5

M

ole

F

r

a

c

tio

n

10 12 14 16

0

0.1

0.2

0.3

Channel 2: qual test

Channel 3: qual test

Channel 4: qual test

Channel 2: 3sec

Channel 3: 3sec

Channel 4: 3sec

Channel 2: 4sec

Channel 3: 4sec

Channel 4: 4sec

time [s]

H

F

C

-

1

2

5

M

ole

F

r

a

c

tio

n

10 12 14 16

0

0.1

0.2

0.3

Channel 1: qual test

Channel 11: qual test

Channel 12: qual test

Channel 1: 3 sec injec

Channel 11: 3 sec injec

Channel 12: 3 sec injec

Channel 1: 4 sec injec

Channel 11: 4 sec injec

Channel 12: 4 sec injec

Figure 30: HFC-125 mole fractions at Halonizer sensor locations. Measurements during the qualification

tests [25] are compared with simulations with 3.2 kg of HFC-125 injected over 3 s and 4 s. Air flow into the

nacelle is 1.0 kg/s and suppressant mass flow is proportional to the nozzle area for four production nozzles.

60
time [s]

H

F

C

-

1

2

5

M

ole

F

r

a

c

tio

n

10 12 14 16

0

0.1

0.2

0.3

Channel 5: qual test

Channel 6: qual test

Channel 7: qual test

Channel 5: 3sec injec

Channel 6: 3sec injec

Channel 7: 3sec injec

Channel 5: 4sec injec

Channel 6: 4sec injec

Channel 7: 4sec injec

time [s]

H

F

C

-

1

2

5

M

ole

F

r

a

c

tio

n

10 12 14 16

0

0.1

0.2

0.3

Channel 8: qual test

Channel 9: qual test

Channel 10: qual test

Channel 8: 3sec injec

Channel 9: 3sec injec

Channel 10: 3sec injec

Channel 8: 4sec injec

Channel 9: 4sec injec

Channel 10: 4sec injec

Figure 31: HFC-125 mole fractions at Halonizer sensor locations. Measurements during the qualification

tests [25] are compared with simulations with 3.2 kg of HFC-125 injected over 3 s and 4 s. Air flow into the

nacelle is 1.0 kg/s and suppressant mass flow is proportional to the nozzle area for four production nozzles.

61
and the period over which it was released varied from 2 to 6 s. Shorter injection periods would correspond

to higher bottle pressures driving the injection process. To the extent that suppressant vaporization is slow,

the period over which the suppressant is effectively injected will increase. Figure 32 shows that the entire

volume is essentially filled for at least one second for injection durations of 3 s or less. For longer periods

of injection, the volume is not filled with high concentrations since suppressant is continually swept out the

various vents. The failure to completely fill the nacelle does not indicate that suppression will not occur, but

that one is in a less conservative regime where suppression confidence is reduced.

time [s]

v

ol(

H

F

C

>

0.3

)

10 12 14 16 18 20 22

0

0.2

0.4

0.6

0.8

1

1.0 kg/s air; 3.2 kg HFC-125 in 2 s

1.0 kg/s air; 3.2 kg HFC-125 in 3 s

1.0 kg/s air; 3.2 kg HFC-125 in 4 s

1.0 kg/s air; 3.2 kg HFC-125 in 6 s

Figure 32: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with 1.0

kg/s air flow into the nacelle. The mass of HFC-125 injected is 3.2 kg, injected over various periods as

indicated with mass flow through each nozzle proportional to the measured nozzle-throat area.

The peak average suppressant mass fraction estimated with Eq. 9 for the simulations shown in Fig. 32

are 0.59, 0.51, 0.44 and 0.35 with smaller mass fractions corresponding to the longer injection duration.

This indicates that, if the the turbulent mixing in the nacelle were sufficient to create homogeneity, the mass

fraction throughout the nacelle would exceed 0.3 for all of these scenarios. Clearly, inhomogeneities arising

from imperfect mixing are significant. This indicates that, while an analysis as in Sec. 3 does provide an

indication of the trends, the CFD analysis can provide significant information on the degree of inhomogeneity. Specifically the CFD analysis can provide an estimate of the volume that fails to meet a criteria such as

the one indicated here (suppressant mass fractions exceeding 0.3). The inhomogeneities that are observed

62
in the simulations are attributable to finite turbulent mixing rates. That is, the nacelle vapors are swept out

of the nacelle (displaced by incoming air and suppressant) before they have had time to become sufficiently

homogeneous. The rate of turbulent mixing is largely dependent on the turbulent kinetic energy and the

length scales in the nacelle. Within this given nacelle the length scales do not change (changing clutter

would affect this, though), so that the turbulent kinetic energy drives the relative rates of mixing here. When

the suppressant is injected at a slower rate (i.e. over 6 s), the suppressant injection generates less kinetic

energy for mixing. This leads to greater inhomogeneities along with the lower average suppressant mass

fraction, and this is reflected in Fig. 32.

Similar results are shown in Fig. 33 where the air flow rate is varied from 0.5 kg/s to 1.0 kg/s while the

suppressant injection is held fixed at 3.2 kg over 3 s. Except for the fact that the suppressant resides in the

nacelle for somewhat longer after the end of injection, the air flow rate has little effect on the distribution

when all four nozzles are employed (different results are obtained in the following section with with some

nozzles capped). This is not necessarily true if the suppressant is not filling the nacelle volume. In Fig. 34

the volume of the nacelle filled for the slowest rate of suppressant injection shown in Fig. 32 is compared

with the volume filled when the air flow rate is reduced by half. While this is an extreme case, the volume

filled does increase as the rate of dilution by the air is reduced.

time [s]

v

ol(

H

F

C

>

0.3

)

10 12 14 16 18 20 22

0

0.2

0.4

0.6

0.8

1

1.0 kg/s air; 3.2 kg HFC-125 in 3 s

0.66 kg/s air; 3.2 kg HFC-125 in 3 s

0.5 kg/s air; 3.2 kg HFC-125 in 3 s

Figure 33: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 when 3.2

kg of HFC-125 is injected over 3 s. The air flow rate into the nacelle is 0.5, 0.66 or 1.0 kg/s as indicated.

63
time [s]

v

ol(

H

F

C

>

0.3

)

10 12 14 16 18 20 22

0

0.2

0.4

0.6

0.8

1

1.0 kg/s air; 3.2 kg HFC-125 in 6 s

0.5 kg/s air; 3.2 kg HFC-125 in 6 s

Figure 34: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 when 3.2

kg of HFC-125 is injected over 6 s. The air flow rate into the nacelle is 0.5 or 1.0 kg/s as indicated.

9 Suppressant distribution without single nozzles

A series of simulations have been conducted to ascertain the effect of removing a single suppressant nozzle

while keeping the overall mass of suppressant injected constant. In conducting simulations with specific

nozzles removed, the mass flow rate out of the remaining nozzles was kept at the values used in other

simulations (i.e. those in Figs. 32 to 34). This corresponds to an assumption of choked flow (driving

pressure and nozzle area determine flow rate) where the pressure driving the suppressant is not affected by

the removal of a nozzle, and that the individual nozzle areas have not been changed. The injection duration

is thus increased somewhat to allow the full 3.2 kg of HFC-125 to flow with a somewhat reduced overall

flow rate.

The nacelle volume fraction where the suppressant mass fraction exceeds 0.3 is indicated in Figs. 35,

36 and 37 for air flow rates of 1.0, 0.66 and 0.5 kg/s, respectively. In all of these scenarios, it appears that

removing nozzle 3 has no detrimental effect since the complete volume is indicated as filled, and the duration

for which it is filled is somewhat increased because of the slightly slower rate of injection with only three

nozzles employed. Figure 38 shows the distribution of the suppressant mass fraction with nozzle 3 assumed

64
time [s]

f

r

a

c

tio

n

(

H

F

C

>

0.3

)

10 12 14 16 18 20

0

0.2

0.4

0.6

0.8

1

1.0 kg/s

1.0 kg/s w/o nozzle 1

1.0 kg/s w/o nozzle 2

1.0 kg/s w/o nozzle 3

1.0 kg/s w/o nozzle 4

Figure 35: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with all

nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg, injected over 3 s.

Air flow into the nacelle is 1.0 kg/s and suppressant mass flow through active nozzles is proportional to the

measured nozzle-throat area.

to be capped. Nozzle 3 is located on the starboard side of the nacelle, about halfway back, and it discharges

into the upper recessed region visible in the centerline plot. The flow from nozzle 3 serves the purpose of

filling the recessed area above the engine. However, it appears that this area would be filled even in the

absence of nozzle 3 with flow convected upwards around the engine from nozzle 2 visible in on the port

and starboard side in Fig. 38. Also, the suppressant from nozzle 3 flowing out of the upper recessed areas

is convected into two regions where suppressant is provided for by other nozzles. Suppressant convected

towards the forward portions fills a volume that is also filled by nozzle 1. Suppressant transported aft towards

the port side fills a region that is also filled by nozzle 4. It appears that the coverage provided by the other

nozzles is sufficient. This assumption will be revisited in the later-described fire suppression simulations.

Based on the criteria shown in Figs. 35 through 37, removing nozzle 4 also has little apparent effect on

the sufficient distribution of the suppressant. The distribution of the suppressant mass fraction with nozzle

4 assumed to be capped is shown in Fig. 39. Nozzle 4 is located in the aft section of the nacelle, on the

port side directed downwards just behind the sixth and last rib. In the absence of nozzle 4, the aft section

of the nacelle fills with suppressant more slowly because it relies on the general aft-ward flow to convect

65
time [s]

f

r

a

c

tio

n

(

H

F

C

>

0.3

)

10 12 14 16 18 20

0

0.2

0.4

0.6

0.8

1

0.66 kg/s

0.66 kg/s w/o nozzle 1

0.66 kg/s w/o nozzle 2

0.66 kg/s w/o nozzle 3

0.66 kg/s w/o nozzle 4

Figure 36: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with all

nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg, injected over 3 s. Air

flow into the nacelle is 0.66 kg/s and suppressant mass flow through active nozzles is proportional to the

measured nozzle-throat area.

suppressant from other nozzles. The mass fraction contour plots indicate that, except for a small region at

the very aft of the nacelle where the afterburner control vanes would be, the suppressant is well distributed

even in the absence of nozzle 4. We note that the lower diamond vent in the nacelle simulator is modeled

here as essentially blocked based on observations of the test simulator; assuming that this vent is not blocked

on fleet aircraft, the removal of suppressant in the aft region may be substantially faster making the retention

of nozzle 4 highly desirable. Also, the additional vent in the aft region would mean that an even more

substantial portion of the suppressant injected through nozzle 4 would rapidly leave the nacelle. Additional

simulations to consider the effect of an open diamond vent could help to address this question.

The removal of nozzle 1 or, in particular, nozzle 2 appears to have a significant effect on the distribution

of suppressant throughout the nacelle. Based on the criteria shown in Figs. 35 through 37, removing nozzle

1 can leave 10-20% of the nacelle with insufficient suppressant converage while removing nozzle 2 can

leave as much as half of the nacelle with insufficient coverage.

The distribution of suppressant mass fraction in the nacelle with nozzle 1 assumed to be capped is shown

for vertical planes near the nacelle center and near the starboard and port sides in Figs. 40. As indicated

66
time [s]

f

r

a

c

tio

n

(

H

F

C

>

0.3

)

10 12 14 16 18 20

0

0.2

0.4

0.6

0.8

1

0.5 kg/s

0.5 kg w/o nozzzle 1

0.5 kg w/o nozzzle 2

0.5 kg w/o nozzzle 3

0.5 kg w/o nozzzle 4

Figure 37: Fraction of the nacelle internal volume for which HFC-125 mass fractions exceeds 0.3 with all

nozzles and with indicated nozzles blocked. The mass of HFC-125 injected is 3.2 kg, injected over 3 s.

Air flow into the nacelle is 0.5 kg/s and suppressant mass flow through active nozzles is proportional to the

measured nozzle-throat area.

earlier, nozzle 1 is located in the forward starboard area of the nacelle and directed upward. The most

significant deficiencies in the suppressant distribution with nozzle 1 capped occur on the starboard side,

especially in the vicinity of nozzles 1 and 2. The mass fraction in the lower nacelle, both forward and aft

of the inlet jet, is also reduced; this is most critical in the region aft of the inlet jet where the mass fractions

were closer to the critical levels for suppression. These lower mass fractions arise because less suppressant

from nozzle 1 is present in the forward areas to be entrained by the inlet jet.

It is noteworthy that the volume fraction over which the suppressant mass fraction exceeds 0.3 with

nozzle 1 capped is less when the air flow rate is less (c.f. Figs. 35 through 37). That is, the coverage is less

sufficient when the air flow rate is less. If one considers the analysis in Sec. 3, this results is counterintuitive.

In Sec. 3 it was shown that the average suppressant mass fraction is greater when the air flow rate is less.

This is true here, where the nacelle-averaged suppressant mass fraction (not shown) is increasing as the air

flow rate is reduced while the volume fraction covered is decreasing. This indicates that there are more

substantial inhomogeneities at lower air flow rates, which is something that the global analysis of Sec. 3

cannot identify. Only a series of measurements or a CFD analysis can identify this phenomena. The reason

67
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 38: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side (top),

near the nacelle centerline and on the port side (bottom). Nozzle 3 has been capped and suppressant flows

through nozzles 1, 2 and 4 only. Air flow into the nacelle is 0.66 kg/s.

68
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 39: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side (top),

near the nacelle centerline and on the port side (bottom). Nozzle 4 has been capped and suppressant flows

through nozzles 1, 2 and 3 only. Air flow into the nacelle is 0.66 kg/s.

69
for this behavior is related to the initial momentum of the inlet air and the role this plays in the suppressant

mixing within the nacelle. The momentum associated with the inlet air flow that drives overall nacelle

circulation and the flow past the injection nozzles. This affects the volume of air into which the remaining

nozzles are injecting suppressant. Doubling the inlet air velocity moves roughly twice the volume of air

past the other nozzles, and thus initially distributes the suppressant over twice the volume of nacelle vapors.

The most relevant nozzles here are nozzle 2 and 3, since suppressant from these nozzles circulates into the

forward nacelle and does so at a greater rate when the air flow is greater.

The distribution of the suppressant mass fraction is shown in Fig. 41 with nozzle 2 assumed to be capped.

Nozzle 2 is located in the forward starboard area of the nacelle and directed across the lower nacelle in

the port or outboard direction. Without nozzle 2 present, the inlet jet entrains relatively little suppressant

and this results in little suppressant along the lower nacelle where pool fires might exist. The suppressant

injected through nozzles 1 and 3 is observed to be largely concentrated in the upper forward nacelle while

that injected through nozzle 4 is observed to be primarily restricted to the region aft of the sixth rib. It

is only after the suppressant injection is complete that the agent becomes well distributed, and the mass

fractions in the lower nacelle at these times are low enough, due to continuous ventilation, that suppression

is questionable or doubtful. It is clear that nozzle 2 plays a key role in distributing the suppressant throughout

the nacelle through its interaction with the air inlet. The fact that the flow rates in front of nozzle 2, which

feeds directly into the inlet air, are so great leads to the great importance of this nozzle in creating a sufficient

suppressant distribution.

To summarize, the sufficiency of the suppressant distribution depends on the locations of the nozzles

relative to the air flow, the rate of air flow past the nozzles and the rate of turbulent mixing. The nozzle

of greatest significance for providing a uniform distribution is the one that feeds into the highest velocity

air flows (nozzle 2 here) since this suppressant is distributed most uniformly by the air flow. Nozzle 1 is

also significant in this respect since it fills the forward region where the air inlet entrains additional vapors.

The kinetic energy of the flow is important for mixing the suppressant, and reducing the kinetic energy

by reducing the suppressant injection rate (or the air flow rate in other scenarios) tends to increase the

inhomogeneities because of reduced mixing rates. The momentum associated with the initial air flow is

also significant since it determines the overall mass of vapors passing the various suppressant nozzles; for

this reason, when nozzle 1 was capped the suppressant coverage was increased with greater air flow even

while the average mass fraction is decreasing. We note here that later fire-suppression evidence indicates

that the Vulcan simulations could be under-predicting the mixing rates to some degree and therefore the

actual uniformity of coverage may be somewhat greater than indicated here. The overall conclusions appear

to remain valid, despite this minor correction.

70
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 40: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side (top),

near the nacelle centerline and on the port side (bottom). Nozzle 1 has been capped and suppressant flows

through nozzles 2, 3 and 4 only. Air flow into the nacelle is 0.66 kg/s.

71
y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

y

z

16 17 18

-1

-0.5

0

C2HF5

0.5

0.45

0.4

0.35

0.3

0.25

0.2

10 m/s

Figure 41: Contour plots of C2HF5 at 3.0 s after start of suppressant injection on the starboard side (top),

near the nacelle centerline and on the port side (bottom). Nozzle 2 has been capped and suppressant flows

through nozzles 1, 3 and 4 only. Air flow into the nacelle is 0.66 kg/s.

72
Part II

Fire suppression predictions

The following sections address predictions of fire suppression in the nacelle simulator. Following the predictions described in this section, a series of fire-suppression tests were conducted at Patuxent River Naval

Air Station. The results of these tests are compared with the Vulcan predictions throughout these sections.

It is noted that the test facility capabilities and required procedures prevented a direct match with the simulated conditions, so that most comparisons are subject to some additional uncertainty associated with known

differences in the test conditions. These differences are identified throughout this section.

10 Pool fires in the nacelle

Pool fires stabilized behind obstructions have been identified as among the most challenging fires to suppress [14–16]. Obstructions such as structural ribs provide a region of recirculating flow where hot products

help stabilize the flame that suppressant is relatively slow to penetrate. In certain scenarios, such as those

described in [9,13,28], the suppressant concentration must be maintained at an elevated level in the flow past

the stabilization region for a substantial period of time to ensure that suppressant penetrates the stabilization

region. In the current simulator the mixing times in the flame stabilization regions are estimated to be short

compared to the suppressant injection times (see Sec. 12), but the stabilization region is still important in

stabilizing the fire.

10.1 Vulcan pool-fire simulation procedures

Because the previous tests with the nacelle simulator [25] focused on spray fires and because obstructionstabilized pool fires are known to be challenging to suppress, the present study focuses on pool fire suppression. We consider pools located in four sections of the nacelle as indicated in the schematic in Fig. 42.

These sections are delineated from the forward to the aft end with increasing numbers as each rib is passed.

The letter ’B’ is applied as a prefix to the pool location to indicate that the pool is located between the two

central longerons. These are the only locations where substantial quantities of fuel can collect. To augment

the quantity of fuel, the basin just to the port of that indicated by ’B’ is sometimes assumed to hold some

fuel also; this basin is denoted with the prefix ’C.’ The pool located at B1 is between the forward end of

the nacelle and before the first rib. That denoted B2 is located just aft of the first rib; the inlet air flows

partially over this pool. The pool denoted B3 is located behind the second rib. The pool denoted B5C5 is

located between the fourth and fifth ribs; these are the two larger ribs. The simulations indicated that pool

fires at locations that would be denoted B4 and B6 are more sensitive to heat losses and therefore difficult

73
to stabilize because of the particular predicted flow circulation. Also, these pools were determined to hold

less fuel than the others. Pools B4 and B6 are not considered further here. The approximate surface area

of the largest possible pools was measured in the F-18 simulator. Similar pool areas have been used in the

simulations; the pool areas are indicated in Table III. The pools are assumed to be filled with two liters of

JP-8 and the net evaporation rate based on heat feedback to the pool is also indicated in Table III. The JP-8

is modeled as a single component liquid with a boiling temperature of 423 K, a enthalpy of vaporization of

280 kJ/kg, an enthalpy of combustion of 43.2 MJ/kg, a heat capacity of 2.47 kJ/kg/K and a density of 808

kg/m3. The pool is assumed to start at a temperature of 293 K and evaporation occurs in the simulations as a

consequence of thermal feedback from the fire to the pools. The heat flux to the pools is partitioned between

pool heating and evaporation. Initially roughly 70% of the heat flux goes to pool heating and 30% of the

heat flux going to pool evaporation for the current simulations, but by the time the suppression occurs the

pools have heated substantially towards the boiling temperature. These pool evaporation and heating rates

presume that the pool is not losing heat through conduction to the nacelle wall during the course of the fire;

as is noted later, it is possible to include heat loss if warranted. In general, it is difficult to accurately predict

the thermal feedback to a pool from a fire. Because these fires are partially blown beyond the pools by the

advective flows in the nacelle, it is particularly hard to determine whether the evaporation rates are reasonable without suitable measurements. The evaporation rate is viewed as one of the significant uncertainties

in the simulation with uncertainties as high as 50% not being unreasonable.

The simulations are conducted by simulating the ignition of the pools with quiescent flow followed by a

six-second ramp up in the air-flow rate to the appropriate flow rate for the test. After this, the simulation was

allowed to run for ten seconds to insure steady state had been reached. In Sec. 10.4, it will be shown that

the simulated flow reached steady state in roughly two seconds. This procedure was followed because the

test procedure was expected to be similar. One of the things that the simulations identified was the difficulty

of igniting the pools in the strong air flow that occurred in the nacelle at operating conditions. In the later

tests, it was determined that it was impossible to ignite the pool fires except at low air-flow rates, and the

test procedure did follow the simulation procedure, although the time to light the pool fires was appreciably

longer, and in a few tests the air flow was held at moderate air-flow rates in an attempt to heat up the nacelle,

ostensibly making it more difficult to extinguish or easier for relight to occur.

Following approximately ten seconds of steady-state simulation, the suppressant was injected. For the

purposes of these simulations, the suppressant, HFC-125 was injected as in its vapor phase. That is, the

suppressant was presumed to flash vaporize upon leaving the nozzle at ambient temperatures. At the time

of the simulations, data was not available regarding the variation of the mass flow rate during the injection

process, and a fixed injection rate over a specified duration was employed as described in Part I. The two

parameters that were used to specify the injection rate were the total mass to be injected and the duration

over which that mass was to be injected. Over that duration, the mass flow rate of the suppressant was

74
1

2

3

4

FRONT VIEW

  



  







  



  







			 			 









   





 





 



 











   









4

B2

B1

B3C3

B5C5

1

2

3

STARBOARD PORT

TOP VIEW

AFT

FRONT

Figure 42: Schematic representation of the relative locations of the pool fires and the suppressant injection

nozzles showing the ribs and central longerons. Not to scale.

Table III: Pool characteristics. Here “All” refers to pools B1, B2, B3C3 but not B5C5, the latter being to

difficult to reach in the course of the tests.

Area Air Flow Pred. Evaporation Volume of Volume where

Pool [kg/s] [m2] rate [kg/s] combustion T > 700 K

B1 0.084 1.0 0.0089 5.9% 4.6%

B2 0.100 1.0 0.011 4.6% 22%

B3C3 0.094 1.0 0.0082 5.5% 37%

B5C5 0.043 1.0 0.0043 1.5% 5.2%

All 0.321 1.0 0.057 16.6% 98.4%

All 0.321 0.5 0.050 7.6% 94.5%

fixed except for 0.1 s ramp-up and ramp-down periods at the start and end of the injection. It is now known

that the injection rate over the first half second is probably substantially larger than the fixed rate, while the

injection rate tapers off after around two seconds (see Sec. 10.2).

10.2 Nacelle test pool-fire procedures

The basic test procedure followed involved (1) the filling of the suppressant bottle with the desired mass

of suppressant, (2) the capping of any suppressant nozzles that were to be turned off for a particular test,

75
(3) the installation of the suppressant bottle on the test fixture, (4) the filling of the each pool (Table III)

with 2 l of JP-8 fuel, (5) the establishment of a minimum air flow and the ignition of each pool fire, (6) a

stabilization period where the pool fires became fully engulfed, (7) a gradual increase of the inlet air-flow

rate to its target value, (8) another stabilization period at the target air-flow rate and (9) the discharge of

the suppressant. Following suppression a reduced air flow was maintained to cool the nacelle and remove

potentially flammable vapors, and then the remaining fuel was collected and measured in the graduated

cylinders. For fires that were not successfully suppressed, the air flow was reduced and a flight-line fire

extinguisher was employed to suppress the fire. Air-flow rates were measured using a 6 inch turbine meter.

Details regarding the measurement equipment employed as well as uncertainties are available in [20].

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–either 1 kg/s or 0.5 kg/s–as measured with a 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 an 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 (see Sec. 10.4).

10.3 Pool evaporation rate results

Evaporation rates in the tests are estimated and these can be compared with those predicted in the Vulcan

simulations. The evaporation rates in the tests are estimated by measuring the difference between the volume

of fuel put into the nacelle and the volume of fuel drained after each test. The volume difference is divided

by the fuel density, estimated at 808 kg/m3, and the estimated duration of the burn. Unfortunately, it is

difficult to assign a duration to the burn since the pools were ignited in sequence and in quiescent air. Once

fires were established the air-flow rate was gradually ramped up to the specified flow rate of 0.5 to 1.0 kg/s.

As a consequence, it is also difficult to assign an air-flow rate to the duration of the burn, and only peak airflow rates are used for comparison. The air-flow rates from test data are compared with those from Vulcan

predictions in Fig. 43 where it is evident that the measured evaporation rates tend to be lower than those

predicted by Vulcan. The tests shown in Fig. 43 correspond to those described in Sec. 14. In light of the

76
Nominal Air Flow Rate [kg/s]

M

e

a

n

E

v

a

p

o

r

a

tio

n

R

a

t

e

[

k

g

/

s

]

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1

0

0.01

0.02

0.03

0.04

0.05

0.06

Pools B1, B2, C2 B3, C3

Pools B3, C3

Pool B1

Vulcan Predictions Pools B1, B2, C2 B3, C3

Vulcan Predictions Pools B3, C3

Vulcan Predictions Pool B1

Figure 43: Evaporation rates predicted in the simulations (closed symbols) and measured in the tests (open

symbols).

77
uncertainties in the burn duration and the air-flow rate, it is expected that the indicated measurements from

the test data in Fig 43 are lower than the actual evaporation rates during the most intense burning rates at

high air-flow rates. It is estimated, based on the time that it took to light the fires and ramp up the air-flow

rate, that the actual burning rates are between 30% and 80% higher than indicated. This suggests that the

evaporation rates predicted in Vulcan using a simple pool model with are still too high. Since the Vulcan

simulations were conducted with no specific heat losses (heat flux to the pool is partitioned between pool

heating and evaporation) it is expected that the Vulcan-predicted evaporation rates would represent an upper

bound. It is possible to employ a heat loss model in conjunction with the pool model but this was not done

because there was not adequate information regarding heat losses when the simulations were conducted

(prior to the tests).

10.4 Nacelle temperatures

During the course of the tests, thermocouple measurements were conducted. An array of 21 thermocouples

was distributed in the nacelle. The temporal response of these thermocouples was not sufficient to determine

the temperatures in the interior of the nacelle, however. The indicated temperatures did not attain steady state

over the duration of the fires. Here the predicted temperatures in the nacelle are compared with the thermocouple measurements. Note that the Vulcan thermocouple model was not employed so that corrections for

convection and radiation balances are not incorporated here. Rather this provides a qualitative indication of

the temperature field.

As a measure of the degree to which fires in different sections are predicted to spread throughout the

nacelle, Figs. 44 and 45 show the volume over which combustion is occurring. For pools denoted B1, B2

and B3C3 combustion (heat release) is occurring in 4 to 6% of the volume as indicated in Table III; for pool

B5C5 the volume over which combustion occurs is only about 1.5% of the nacelle internal volume because

the fire is contained between the larger structural ribs surrounding this pool. Also indicated in Table III is

the nacelle volume fraction over which hot product gases are spread; the criteria for hot product gases is

that the control volume averaged temperatures exceed 700 K (427 C or 800 F). Heat release from pool fires

B2 and B3C3 is substantially spread throughout the nacelle by the main air inlet that passes over the fires in

these regions. Since the pool evaporation rates are likely over-predicted, the volume over which combustion

is occurring should be correspondingly over-predicted.

Temperature-time profiles at locations corresponding to several of the Halonizer probes are presented

in Fig. 46. These temperature profiles are indicative of the variations in the nacelle during the course of

the simulation. A fire is ignited with minimal air flow near t = 0 s. While the air flow is minimal there

are substantial fluctuations in the temperature since the hot gases are convected in a chaotic manner around

the nacelle. As the air-inflow rate increases linearly from t = 2 s to t = 8 s the fire approaches a more

78
x

0

0.5

1

y

15

16

17

18

z

-1.5

-1

-0.5

0

Y

X

Z

x

0

0.5

1

y

15

16

17

18

z

-1.5

-1

-0.5

0

Y

X

Z

Figure 44: Surface plot showing cells where combustion is taking place for pool fires in locations B1 (top)

and B2 (bottom). Surface indicates at least 1 ppm by volume combustion in a computational cell. Air flow

into the nacelle is 1.0 kg/s.

79
x

0

0.5

1

y

15

16

17

18

z

-1.5

-1

-0.5

0

Y

X

Z

x

0

0.5

1

y

15

16

17

18

z

-1.5

-1

-0.5

0

Y

X

Z

Figure 45: Surface plot showing cells where combustion is taking place for pool fires in locations B3C3

(top) and B5C5 (bottom). Surface indicates at least 1 ppm by volume combustion in a computational cell.

Air flow into the nacelle is 1.0 kg/s.

80
t [s]

T

e

m

p

e

r

a

t

u

r

e

[

K

]

0 5 10 15 20 25 30

0

400

800

1200

1600

2000

1

2

3

4

5

Channel

Figure 46: Temperatures at locations in the nacelle corresponding to the locations of channels 1 through 5 of

the Halonizer probes are plotted as a function of time for case where air flow is 1 kg/s. All pools are ignited

and suppressant injection occurs at 20 s.

81
determinate state, being driven by the air inflow. The flow field stabilizes before t = 10 s and is completely

stationary for ten seconds of simulated time. At t = 20 s, the suppressant is injected followed quickly by

suppression, and the temperature drops rapidly.

The predicted evolution of the temperature can be compared with thermocouple measurements conducted during the course of the tests. An array of twenty-one thermocouples was placed along the lower

nacelle in an effort to measure the temperature field and provide an indication of the fire location. The thermocouples were placed in three rows with up to eight thermocouples in a row. The rows extended from the

forward to the aft end of the nacelle with thermocouples spaced every eight inches (20 cm). One row was

approximately located along the center of the nacelle while the other rows were roughly twelve inches (30

cm) to the inboard and outboard edges of the nacelle. The middle row contained only five thermocouples

in order to not interfere directly with the pools. An example of the thermocouple temperature profiles is

provided in Figs. 47 and 48 where notable differences with Fig. 46 are evident (note that the test measurements are reported in Fahrenheit in Figs. 47 and 48). Most significant is the fact that the thermocouple time

constants are comparable to the duration of the burn, which was generally limited by the time it would take

to consume the fuel in pool B2. That is, under the high air-flow rates used in the tests, the fuel evaporation

was significant over periods on the order of twenty seconds, which is short compared to the time constant

of the shielded thermocouples. In order to avoid uncertainties related to fuel depletion, we were forced to

stabilize the flow for only 10-20 seconds at high flow rates before discharging the suppressant bottles. This

implies that the thermocouple data does not represent the actual nacelle temperatures. It is noted that the

gas temperatures are expected to equilibrate in roughly two seconds, which is the air turnover time scale

for the nacelle. The thermocouple data, therefore, is only representative of the combustion temperatures in

the nacelle and is not considered accurate because it continues to be influenced by the ambient temperature

when the fires were initially ignited. It is also noted that, based on cool-down times, the thermocouple

time constants are observed to vary substantially. These differences in the time constant are attributable to

the different air-flow rates past the thermocouples; high convection coefficients reduce thermocouple time

constants. This suggests that even relative temperatures, between thermocouples in the same test, are not

good indicators of even relative temperatures. The test number selected for Fig. 48 was chosen because, for

this test, a substantially longer burn duration (on the order of 90 s after the pools were burning at low airflow rates) was employed before suppression, allowing the thermocouples to rise nearer to their steady-state

values. It is clear from Fig. 48 that the thermocouple temperatures are still rising after this long duration

burn. The burn time corresponding to the test in Fig. 47 was approximately 15 s so that the thermocouple

temperatures can be expected to rise substantially further if a steady state is reached.

Thermocouple measurements conducted along the lower portion of the nacelle during the course of the

tests are presented in Figs. 47 and 48. The simulations also predict the full temperature field throughout

the nacelle. With the understanding that the higher predicted fuel-vaporization rates lead to higher overall

82
Figure 47: Thermocouple measurements (uncorrected) along the lower nacelle for run 1d(25) are shown in

the upper figure as a function of the test time. The air flow is nominally 1 kg/s and the lower figure shows

the temporal variation in the measured air mass flow rate (in lb/s). All pools are ignited prior to the start of

data acquisition.

83
Figure 48: Thermocouple measurements (uncorrected) along the lower nacelle for run 5b(25) are shown in

the upper figure as a function of the test time. The air flow is nominally 0.5 kg/s and the lower figure shows

the temporal variation in the measured air mass flow rate (in lb/s). All pools are ignited prior to the start of

data acquisition.

84
heat release and correspondingly higher temperatures, Vulcan-predicted temperature contours are provided

at roughly the axial locations where temperature measurements along the lower nacelle are also available.

These are presented in Figs. 49 through 52 for air-flow rates of 0.5 kg/s on the left pane and 1.0 kg/s on

the right pane. In these figures, the peak (not yet steady-state) temperatures taken from Figs. 47 and 48 are

imprinted on the contour plots in white lettering. Note that the indicated positions of the temperatures are

only approximate and because the temperatures do not reach a steady state the actual gas temperatures are

expected to be higher. In general the agreement between the Vulcan-predicted temperatures and the thermocouple measurements is only moderate. The disagreement is attributable to the long thermocouple time

constants, differing time constants among different thermocouples, the lower-than-predicted pool burning

rates, and the uncertainty in the thermocouple positions. The thermocouple measurements in the forward

section of the nacelle are generally the lowest, in agreement with the temperature contours in Fig. 49. However, the thermocouple temperatures rise relatively little compared to the predicted temperatures. It is noted

that the higher thermocouple temperatures in the left column are indicative of the longer test run time for the

0.5 kg/s air-flow case (Fig. 48) and these do not necessarily indicate that the temperatures of the low-flowrate tests are higher. In Fig. 50 the low and high air-flow rate simulations show a different trend regarding

the fire around the air inlet: at low flow rates the fire is concentrated around the inflow jet while at high

flow rates the inflow velocity is high enough that there is little heat release evident around the jet. In the

tests, it was observed that at the highest air-flow rates, the fire over pool B2 was largely blown back towards

pool B3. This could be evidence of local extinction just from high rates of mixing, a phenomena referred

to as blow out. The thermocouple measurements indicate the highest temperatures in those planes shown in

Fig. 51. The thermocouple temperatures in the last two planes indicate cooling that is not evident in Fig. 52;

this is likely a consequence of the lower observed evaporation rates in the tests relative to the simulations.

The thermocouple results do corroborate other observations. The thermocouple temperatures are appreciably lower than the Vulcan simulations, even if attempts are made (not presented here) to correct for

radiative losses and to predict the steady-state temperatures. The highest five temperatures in Fig. 47 at an

air-flow rate of 1 kg/s range from approximately 700 K to 950 K (800 F to 1250 F). The highest six temperatures in Fig. 47 at an air-flow rate of 0.5 kg/s range from approximately 800 K to 1150 K (1000 F to 1600

F). The lower-than-predicted temperatures corroborate the observed reduced fuel-evaporation rates that are

measured relative to those predicted by Vulcan. Reduced evaporation rates will lead to less heat release per

unit air inflow and lower averaged gas temperatures.

10.5 Suppressant bottle test procedure and discharge rate

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 (C2HF5) was the agent used in all tests. The rate of injection was controlled by the

85
x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

640

820

480

400

650 400

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

540 410 390 410 Figure 49: Temperature contours at locations 15.16 m (top) and 15.55 m (bottom). The air flow rate is 0.5

kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple temperature from

Figs. 47 and 48 (in degrees K) is shown with white lettering.

86
x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

570

800

440 570

540

480

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

550 820 600

630 Figure 50: Temperature contours at locations 15.81 m (top) and 16.20 m (bottom). The air flow rate is 0.5

kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple temperature from

Figs. 47 and 48 (in degrees K) is shown with white lettering.

87
x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

580 1130 770

720

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

820

1080

700 940

800 550

Figure 51: Temperature contours at locations 16.59 m (top) and 17.01 m (bottom). The air flow rate is 0.5

kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple temperature from

Figs. 47 and 48 (in degrees K) is shown with white lettering.

88
x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

700

620

480

770 600

440

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 0.5 kg/s

x

z

0 0.5 1

-1

-0.5

T

1800

1600

1400

1200

1000

800

600

400

Air flow 1.0 kg/s

600

540

450 560

500

400

Figure 52: Temperature contours at locations 17.41 m (top) and 17.83 m (bottom). The air flow rate is 0.5

kg/s in the left panes and 1.0 kg/s in the right panes. The peak indicated thermocouple temperature from

Figs. 47 and 48 (in degrees K) is shown with white lettering.

89
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 pressure

measurements suggest a three-stage discharge process. The suppressant bottle discharges downward so that

the initial stage of the 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. That is, once the bottle

pressure approaches the agent vapor pressure, the agent boiling keeps the bottle pressure approximately

constant. 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. 53.

Test labels will be described below in Sec. 14.2 and given in Table VI. 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 affect the discharge rate

by roughly ±10%. Note that the discharge rate is proportional to the square root of the pressure, to leading

order.

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

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 (by volume) 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 liquid (volume fractions are given in the previous sentence), but

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

11 Suppression model

The fire suppression model contained within Vulcan [13] is key to its ability to predict the nacelle physics.

Because the full range of chemical and fluid dynamic length scales cannot be resolved in engineeringscale turbulent flows, the effects of chemical kinetics are captured within subgrid-scale models. To predict

90
Figure 53: 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. The curves are labeled with the test label to be given

below in Table VI.

91
fire suppression requires a prediction of local flame extinction. An extinction criterion can be expressed in

terms of a Damkohler number,  ̈ Da = τη/τext, the ratio between fluid mixing and chemical times. A suitable

chemical time relevant to flame extinction, τext, can be defined as a perfectly stirred reactor (PSR) mixing

time corresponding to a switch from combustion occurring to combustion being impossible. In other words,

τext is the chemical time scale representing the fastest possible chemical-reaction rates. For Damkohler  ̈

numbers smaller than a critical value, Dacrit , the flame will be extinguished, so that for

τη/τext > Dacrit (10)

combustion may occur, while for

τη/τext < Dacrit (11)

combustion may not occur. Combustion in Vulcan is modeled using the EDC model that can be thought of

as a distribution of PSR’s that relate the fuel-consumption rate to the fluid-mixing rate, the latter obtained

from turbulent time scales. The submodels within Vulcan assume that the fuel-air mixing rates scale with

the inverse of the Kolmogorov time scale, τη,

τη = 0.41(ν/)

1/2

(12)

where ν is the molecular viscosity and  is the turbulent kinetic-energy dissipation rate. The actual mass

rate of reactant mixing is proportional to τη and also a function of the local reactant mass fractions. It is

noteworthy that in high-fidelity modeling of turbulent diffusion flames, the fluid time scales corresponding

to extinction are found to be very close to the Kolmogorov time scale [12]. For mixing rates that exceed a

critical value, the chemical kinetics and the associated heat release cannot keep up with the rate of reactant

influx and the rate of product out flux. In this case, the flame will be extinguished. This critical mixing rate

is true of both PSR’s and turbulent flows, but as indicated below, τη is not exactly the same as the critical

mixing rate in PSR’s.

The chemical time scale corresponding to extinction, τext, is a function of the reactant composition and

enthalpy. This time scale is essentially the inverse of the maximum possible rate of reaction within the

flame. Fire suppressants typically increase the chemical time scale by slowing down the overall chemical

kinetics in at least one of the following manners: (1) dilution reduces the frequency of collisions between

reactive molecules, (2) chemical inhibition causes a reduction in the levels of active radicals through an

enhancement of chain-termination reactions and (3) cooling of the flame reduces the overall kinetic rates

through a reduced probability of a molecular collision crossing an energetic barrier. The chemical time

scale is obtained through detailed chemical-kinetic calculations using the hydrocarbon mechanisms from [3]

and suppressant mechanisms from [1, 21]. Using these chemical-kinetic mechanisms in the Chemkin PSR

92
τη

X

C

2

H

F

f

o

r

s

u

p

p

r

e

s

sio

n

10

›5 10›4 10›3 10›2

0

0.02

0.04

0.06

0.08

0.1

τη

Y

C

2

H

F

f

o

r

s

u

p

p

r

e

s

sio

n

10

›5 10›4 10›3 10›2

0

0.1

0.2

0.3

Figure 54: Suppressant mole (left) and mass (right) fractions required to extinguish flames for a given

turbulent mixing time scale.

code [26], τext was determined for PSR’s with varying levels of suppressant mass fraction and varying

degrees of cooling.

The critical Damkohler number that determines the ratio between  ̈ τη and τext at flame extinction is not

known a priori because the PSR mixing rate is not identical to the turbulent mixing rates, but is related by

an order-unity constant. The critical Damkohler number is determined empirically by matching blow-off  ̈

criteria for an ethane jet flame as described in [30]. Because this test involves no suppressant, this measure

is independent of knowledge of the suppressant chemistry. A value of Dacrit = 0.563 is obtained in this

manner indicating that the PSR and turbulent mixing rates are indeed of the same order. Combining this

extinction criteria with the various chemical times leading to PSR extinction, as described in the previous

paragraph, yields an extinction criteria with fire suppressants. For HFC-125 employed here, the extinction

criteria are plotted in terms of the suppressant required for a given fluid mixing rate, τη, in Fig. 54. It is

seen that as the suppressant mass fraction increases, the mixing time required for complete combustion,

and thus flame stabilization, increases. This corresponds to the reduction in the chemical rates associated

with the introduction of suppressants as described in the previous paragraphs. It is presumed that there is a

maximum mixing time in relevant systems that arises because of the expected range of physical scales and

the buoyancy generated mixing rates. This time scale is 0.6 ms, and this leads to the maximum required

suppressant mass fraction of 0.28 evident in Fig. 54. Such a limit is in agreement with cup-burner test

results [9].

93
12 Mixing times for clutter

There are two key challenges in suppressing clutter-stabilized pool fires. First, clutter creates a region in

the recirculation zone where fluid-mixing time scales may increase. As indicated in Fig. 54, this leads to an

increase in the required suppressant. Second, there is a delay in the mixing of suppressant from the flow past

the obstruction into the recirculation zone. This delay can be approximated using a perfectly-stirred reactor

mixing model. Under such an assumption, the mean suppressant mass fraction in the flame stabilization

zone, Ystab, can be approximated as

Ystab = Y∞[1 − exp(−t/τmix)] (13)

where Y∞ is the mass fraction in the fluid flowing past the recirculation zone and τmix is the mixing time

constant. For simple flow past a single step or rib with no other obstructions, Takahashi et al [29] measured

the mixing time constant to be

τmix = 34.7(h/u∗) (14)

where h is the obstruction height and u

∗

is the average of the mean flow velocity over the step and the mean

flow velocity without the step. Based on the CFD simulation predicted velocities, it is possible to estimate

mixing times using Eq. 14 although these estimates are complicated by the facts that the velocity varies

laterally across the ribs and that the flow is not exactly normal to the ribs. For flow past the first rib, which

stabilizes flames in pool B2, the mixing times range from very small where the inlet flow crosses the rib to

0.9 s where velocities are lower. For flow past the second rib, which stabilizes fires in pool B3, mixing times

are in the vicinity of 0.1 to 0.2 s; mixing times for the third rib are similar. For flow past the large fourth and

fifth ribs, which stabilize fires in pool B5C5, the mixing times range from 0.6 to 1.5 s. Except for the last set

of ribs, the mixing times are generally small relative to the residence time for the suppressant in the nacelle

so the finite mixing time is not expected to play a significant role.

Hewson et al [13] showed that additional obstructions can affect the mixing time constant. In general,

when obstruction wakes are coincident with the stabilization zone, the associated wake low-pressure region

tends to pull the flow away from the stabilization zone and increase the time required for suppressant to mix

into the flame stabilization zone. If the obstruction is located so that the high-pressure stagnation region in

front of the obstruction is coincident with the wake, the opposite behavior is observed, resulting in reduced

suppressant mixing times.

94
Table IV: Fire suppression tests with an air-flow rate of 1.0 kg/s.

Pool mass HFC-125 [kg] injection rate [kg/s] Nozzles Suppression Time [s]

B1 3.2 1.07 1,2,3,4 0.8

B1 2.2 1.07 1,2,3,4 0.8

B2 3.2 1.07 1,2,3,4 0.9

B2 2.2 1.07 1,2,3,4 0.9

B3C3 3.2 1.07 1,2,3,4 1.2

B3C3 2.2 1.07 1,2,3,4 1.2

B3C3 3.2 0.53 1,2,3,4 none

B3C3 3.2 0.64 1,2,3,4 none

B3C3 3.2 0.71 1,2,3,4 none

B3C3 2.2 0.73 1,2,3,4 none

B5C5 3.2 1.07 1,2,3,4 1.8

B5C5 2.2 1.07 1,2,3,4 1.8

13 Suppression of pool fires: initial predictions

Suppression simulations were conducted in two phases followed by tests conducted on the physical nacelle

simulator. The first series of simulations identified some test-design issues that led to a rethinking of the test

procedure. Specifically, the challenge of igniting and stabilizing flames in specific areas of the nacelle was

identified in the first stage of simulations. Also, the sensitivity of the suppression to pool heat losses was

identified. To overcome concerns raised by these simulation results, the test plan was revised and a second

set of simulations was carried out to conform to a set of tests that were expected to be easier to conduct

and less sensitive to uncontrollable parameters. In this section, the results of the first series of simulations

are presented. Results are first presented using the four-nozzle suppression system for various pool fire

configurations presented in Sec. 10, and then for a three-nozzle suppression system in which one of the

nozzles was capped. In the following section the results of the second series of simulations are presented

along with the test results.

13.1 Results with four nozzle system

In general, suppression is predicted for pool fires at all locations in the nacelle using the four-nozzle configuration with 3.2 kg of suppressant when this mass is injected in three to four seconds. A three second

injection duration (1.07 kg/s of suppressant injection) corresponds to the nominal suppression system design

identified in previous testing, so the present results indicate that the production system should be sufficient

to suppress pool fires. In previous sections (see Sec. 3 and 8), it was observed that reducing the mass injection rate of the suppressant reduced the average mass fraction in the nacelle, and also resulted in apparently

95
insufficient coverage to suppress certain fires. To examine the effect of varying the mass of suppressant

injected and the rate of suppressant injection, a series of additional simulations for fires in pool B3C3 were

conducted. Simulations with other pools were conducted, but other pools proved easier to extinguish, and

these are not reported. When the rate of suppressant injection is reduced to just below 0.75 kg/s, no suppression was predicted in the simulations, although the combustion would certainly be described as tenuous for

these conditions since the fire was restricted to a corner of the pool C3. As the rate of suppressant injection

is reduced, the fire stability increases, and at 0.53 kg/s the fire is evidently stabilized, at least near the corner

of pool C3. The fire tends to stabilize in this location because the mixing rates are lower.

It is noteworthy that the mass of suppressant injected plays little role in determining the occurrence of

suppression here. For the two cases with suppressant injection rates of 0.71 and 0.73 kg/s, the resulting

tenuous combustion region is very similar. There will certainly be a lower bound in terms of suppressant

mass injected over a short period of time, although this has not been identified in the simulations. More

significant in terms of the mass of suppressant employed is the duration over which suppressant will act

to inhibit reignition and allow cooling of heated surfaces. Thus the suppressant requirements are likely to

be dictated by two factors. First, to insure suppression by flooding the compartment with a high enough

concentration, a sufficiently high rate of injection is required. Second, to maintain that concentration for

long enough to prevent reignition a minimum mass is required. Only the former factor is considered here.

In general, the suppression effectiveness is found to correlate most strongly with the mass injection rate.

Though the results are not presented here, suppression is sensitive to heat flux from the fire to the

pool. If the heat flux and, hence, the fuel evaporation are reduced by 50% by imposing a 50% heat loss

on the pools, suppression is substantially easier for the cases considered in this section. Specifically, the

reduced suppressant injection rates shown in Table IV that fail to suppress the fire in pool B3C3 succeed in

suppressing the fire if the heat flux is reduced by 50%. The reason for this is that the fire can more readily

be completely separated from the fuel source with slower evaporation rates. Similar results are expected if

the fuel is cold or if there are substantial heat losses through the nacelle under the pool. While relatively

high engine temperatures are expected during operations, temperatures during ground tests may be such that

reduced evaporation is experienced.

In light of the lower evaporation rates observed in Sec. 10, it is likely that if tests were carried out

for some or all of these simulations, then fires might have been extinguished more readily. However, the

sensitivity to heat losses that was identified with a single pool fire led to a shift in the test plan towards tests

with several pools simultaneously burning.

96
Table V: Fire suppression tests with an air-flow rate of 1.0 kg/s and individual nozzles capped.

Pool mass HFC-125 [kg] injection rate [kg/s] Nozzles Suppression Time [s]

B1 3.2 0.83 2,3,4 1.3

B1 2.2 0.83 2,3,4 1.3

B1 3.2 0.74 1,3,4 1.5

B1 2.2 0.74 1,3,4 1.5

B2 3.2 0.83 2,3,4 none

B2 2.2 0.83 2,3,4 none

B2 3.2 0.74 1,3,4 4.6

B2 2.2 0.74 1,3,4 3.2

B3C3 3.2 0.83 2,3,4 1.6

B3C3 2.2 0.83 2,3,4 1.6

B3C3 3.2 0.74 1,3,4 none

B3C3 2.2 0.74 1,3,4 none

B5C5 3.2 0.74 1,3,4 2.4

B5C5 3.2 0.90 1,2,3 1.7

13.2 Results with individual nozzles capped: inhomogeneities

In previous sections, the capping of various nozzles was shown to result in suppressant inhomogeneities that

left certain regions with little suppressant as indicated in Figs. 35, through 37. Of significant interest was

the flow over pools B2 and B3 when nozzles 1 or 2 were capped as indicated in Fig. 40 and 41. To examine

the effects of these inhomogeneities on the predicted suppression, a series of simulations were conducted

where one nozzle was assumed to remain capped. For the majority of these simulations, the flow rate out of

the other individual nozzles was maintained at the flow rate that corresponds to 1.07 kg/s if all the nozzles

were opened. This results in a reduced flow rate for the sum of the three remaining nozzles relative to the

sum of the four nozzles and an increased injection duration. This was done to facilitate the experimental

validation process since the bottle pressure can presumably remain fixed with and without nozzles capped

with the per nozzle flow rate primarily related to the bottle pressure driving the flow. Results of these tests

are summarized in Table V.

For fires in pools B1 and B5C5, the removal of various nozzles is not found to alter the prediction of

suppression. For pool B1 the removal of either nozzle 1 or 2 has little effect beyond delaying the suppression

for a short time. This delay in suppression appears to correspond to the overall slower suppressant injection

since longer suppression times correspond to slower injection.

For pool B5C5, the removal of nozzle 2 results in an increase in the suppression time because suppressant

from nozzle 2 is entrained in the air inflow and carried back across the ribs surrounding the pool. Without

nozzle 2, the suppression is delayed until suppressant from other nozzles works its way around the nacelle

97
to be entrained by the air inflow. The removal of nozzle 4, which is closest to pool B5C5 and injects

suppressant around the aft section of the nacelle, has minimal effect on suppression of this pool since the

mean flow along the nacelle bottom is aft-ward.

For pool B3C3, the removal of nozzle 1 has little effect other than to slightly delay the suppression.

Removing nozzle 2, however, results in a failure to suppress fires in pool B3C3. As indicated in Fig. 41, the

removal of nozzle 2 leaves a significant volume with little suppressant. This volume is the volume directly

over the pool B3C3, and thus nozzle 2 is necessary to suppress fires in pool B3C3.

For pool B2, eliminating nozzle 1 leads to a predicted failure to suppress because the suppressant from

nozzle 1 fills the forward nacelle and is entrained by the inlet air flow. In the absence of the high suppressant

concentrations in the forward nacelle, the pool fire over B2 is predicted to be maintained upstream of the

suppressant from nozzle 2 which passes over the downstream portion of that pool. Capping nozzle 2 is found

to noticeably lengthen the suppression time because the suppressant from nozzle 2 creates strong flows in

the vicinity of pool B2. Without the enhanced flow around pool B2 the entrainment of suppressant from the

forward nacelle area is substantially slower.

14 Suppression of pool fires: pre-test predictions and test results

In the course of conducting the nacelle fire-suppression simulations and designing the tests, a number of

factors were identified that led to a change in the test plan. The resulting test plan is described in this section

along with the results of the pretest simulations and the actual tests. In general, the tests indicated that the

pretest simulation results were largely correct.

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 B3C3. It was determined that pool B5C5 used in

the simulations in the previous section would be too difficult to fill with fuel and ignite given its obstruction

by the larger fourth and fifth structural ribs. In most of the tests, all three pools were ignited because this

provided the strongest fire source; this was done largely to impose a sufficiently strong heat flux to the pools

since the simulations described in the previous section indicated a sensitivity to heat losses. This sensitivity

to heat losses is supported by observations of several tests conducted with a single pool. 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 and indicated in Fig. 43.

98
14.1 Key results from pretest predictions

Results from the pretest predictions described throughout this report and also the simplified analysis in

Sec. 3 helped guide the design of the experimental test plan and test matrix. These results provide estimates

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

rise and fall, and the uniformity of the suppressant distribution. To begin with, the 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

inflow rate provides a characteristic mean steady-state mass fraction within the nacelle, Ys,st (Eq. 8). The

injection must proceed for a long duration for the mean mass fraction to reach this characteristic steady-state

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. This is a consequence of the short

time scale for the transients (τ in Sec. 3), on the order of one to two seconds for the flows considered here,

so that the characteristic steady-state mass fraction, Ys,st, is approached.

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 [9]. 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 mass fraction in certain regions of the flow is substantially less than the

mean. Based on the combined test and simulation results, it is expected that the inhomogeneities in the

suppressant mass fraction are on the order of 30-40%. Accounting for regions of high strain rate in the

flow (near the air inlet) it is likely that the inhomogeneities are substantially larger. High rates of mixing

present in the nacelle tend to reduce the mass fraction required to suppress the fire. It is found for HFC-125

that strained laminar flames are indicated to be suppressed at mass fractions as low as 0.16 [9]. This would

correspond to τη = 0.2 ms (Fig. 54), a typical value observed along the lower section of the nacelle in the

simulations.

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 [13]. 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

99
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 injection rate to the combined, total mass

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

14.2 Test plan

From the Vulcan 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.

Each of these paths was followed until at least one case was identified with a failure to suppress in the

physical tests or until physical limitations on the test equipment were encountered. 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. Because the

suppressant injection rate was identified throughout the pretest simulations as the most significant variable

(among those that could physically be varied), for every other system variable tested, the suppressant injection rate was varied, attempting to reduce it until the suppression failed or until the minimum attainable

injection rate had been tested.

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 second attempt succeeded in suppressing the fire. The lowest injection rate that is attainable is 50% of the nominal design injection rate, so this process results in bracketing

100
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 VI. The tests in Table VI are labeled in the manner in which they were coded in executing

the actual tests. The first digit, a number, is unique if one of the following is varied with respect to the

baseline test configuration: the air flow rate is varied, a nozzle is capped, the suppressant mass is reduced.

The second digit, a letter, varies in the sequence that a subset of tests were conducted. Typically the second

digit varies according to the varying of the suppressant bottle pressure, and thereby the suppressant mass

injection rate, although in some instances the number of pools burning or the mass of suppressant was varied

within a subset of tests to address questions that arose during the test sequence. The final four digits, a day

of the month in parenthesis, indicates the day of October, 2004, on which a test was conducted. A pair of

tests, 1a and 1b were conducted in preparation during the prior week and duplicated on the week on which

the remainder of the tests were conducted.

14.3 Summary of test results and corresponding pretest simulations

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

suppression 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 is 50% of the heat flux to the pool (essentially reducing

the vaporization rate by 50%), then pools in sections B4 and B6 not employed in the present series of tests

could fail to sustain themselves. Since it was not feasible to reach these locations to fill them with fuel and

ignite them, this could not be confirmed. 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 under quiescent conditions,

and this lends support to the Vulcan observations.

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

confidence in the results; in each repeated test the results, in terms of whether suppression was successful or not, 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 Sec. 14.1 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:

101
Table VI: Summary of tests conducted.

Test label air inflow rate pools nozzles supp. mass bottle press. targ supp rate

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

1a(21) 2 all all 7 515 2.33

1a(25) 2 all all 7 515 2.33

1b(21) 2 all all 7 122 1.16

1b(26) 2 all all 7 131 1.16

1d(25) 2 all all 7 265 1.75

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

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

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

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

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

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

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

5b(25) 1 all all 7 122 1.16

5d(25) 1 all all 7 265 1.75

7a(26) 2 all all 4.84 458 2.33

7d(26) 2 all all 4.84 270 1.75

7e(28) 2 all all 3 515 2.33

7f(28) 2 all all 2.25 600 2.33

8a(26) 1.5 all all 7 121 1.16

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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. 53 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, as discussed in Sec. 10.5. Instead, the discharge rate is estimated. For this

purpose, target 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,

especially in light of the nonlinearity of the curves in Fig. 53. 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%.

The results of the tests employing all nozzles are shown in Fig. 55. 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 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. 55 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.

103
Figure 55: 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.

104
Figure 56: 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.

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).

Figure 56 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 injection rate is reduced. The reduction depends on each specific nozzle orifice

diameter (Table I), 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

105
of Fig. 56. 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 improves by capping nozzle 3, and this conclusion should

not be drawn based on the present results even though 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.

Tests where nozzle 1 is capped are indicated with the prefix ‘3’ in Fig. 56. For these tests, it is necessary

to refer to Table VI, 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; based on the results of this overall program,

cases 3b and 3e are considered to be the same parameter space and 3e was run specifically to confirm the

result of test 3b. Specifically, the Vulcan predictions indicated a failure to suppress the fire in test 3b (the

Vulcan prediction for 3e was not actually run, but assumed to provide the same result as 3b) 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 clutter model, as implemented for the pretest simulations, did 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 as 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. Such information is particularly useful in designing tests where one might

want to locate fires in particularly challenging-to-suppress areas.

The results of tests for which nozzle 2 is capped are indicated by the labels with the prefix ‘2’ in Fig. 56.

As with the previous examples, the Vulcan simulations indicated a failure to suppress at the higher injection

rates for tests ‘2a’ and ‘2b’ 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 roughly 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

106
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. Test ‘2a’ was a duplication of ‘2b’ except for the fact that only the third

set of pools (B3C3) is filled in test ‘2a.’ Similarly, test ‘2c’ is a duplicate of test ‘2d’ except that only pool

B3C3 was filled. These results show that extinction of pool B3C3 is difficult when nozzle 2 is capped. This

is in agreement with the Vulcan predictions, except for the above mentioned 20% difference in the injection

rate required in Vulcan.

15 Summary and Conclusions

In this report a detailed CFD analysis of fire suppression predictions within an aircraft engine nacelle is made

using Vulcan. The geometry is that of a nacelle simulator used for ground-based fire tests. The distribution of

the suppressant within the nacelle is described based on predictions and compared with available point-wise

measurements. Simulations are conducted in which pool fires are ignited in the nacelle and then the ability

of various suppressant discharges to suppress these fires is predicted. Following this series of simulations,

tests were conducted in the simulator to evaluate the Vulcan predictions. In most of the cases considered the

tests indicated agreement with the suppression predictions.

The Vulcan suppression model that makes these predictions possible is based on the concept of a critical

Damkohler number for flame extinction. Such a model relates the available fluid time for heat release to the  ̈

maximum chemical heat-release rates. This model captures the full range of dilution, cooling and chemical

inhibition modes of extinction through a detailed chemical-kinetic model of the flame-extinction process.

A general understanding of the suppression-system effectiveness can be obtained by considering two

factors: first, an overall average mass fraction of suppressant provided to the nacelle and, second, a level of

inhomogeneity in the suppressant mass fraction throughout the nacelle. The first can be understood using

zonal type models while the second can only be understood using a well-instrumented test fixture or a CFD

approach like the present.

To leading order, the average mass fraction of the suppressant in the nozzle is determined by the ratio

of the mass injection rate of the suppressant to the total mass injection rate (mostly suppressant plus air).

In addition to the average mass fraction it is necessary to understand how uniformly the suppressant is

distributed within the nacelle. Regions where the suppressant mass fraction is too low provide locations

where a fire can be stabilized until the suppressant has been dissipated. The degree of inhomogeneity is

determined by the location of the suppressant nozzles and the rate of turbulent mixing around these nozzles

and throughout the nacelle. The cluttered nature of nacelles tends to lead to relatively rapid mixing, but the

CFD results suggest, and the tests confirm, that the level of inhomogeneity is still high enough that 30-40%

107
excess average mass fraction is required to insure suppression throughout the nacelle.

Vulcan predictions and tests were conducted using a four nozzle system similar to that installed on a

production aircraft. The predictions and test results indicate that this suppression system is also effective

against pool fires as studied here. In order to cause a suppression failure, it was necessary to reduce the

suppressant bottle pressure by roughly a factor of four so that the estimated suppressant discharge rate was

half of the design discharge rate. In this case, both the predictions and test results indicate a failure to

suppress.

In predictions, the mass of the suppressant injected was reduced by approximately 30%, and this was

found to have no detrimental effect in the suppression. The test results confirmed this, and tests with as

much as a 70% reduction in suppressant mass also resulted in suppression. (The potential for hot-surface

relight may be sensitive to the suppressant mass and this was not tested here.)

To go beyond the installed suppression system design, certain nozzles were capped for certain tests. The

pretest predictions indicated that capping the nozzles referred to as 1 or 2 (Table I) should leave regions

where the suppressant was insufficient to suppress the fires. The test results indicate that for both of these

cases, it was necessary to reduce the suppressant injection rate by 25% before this occurred. In other words,

the Vulcan predictions indicated slightly more more inhomogeneities than the tests suggest when nozzles

1 or 2 are capped. This leads to slightly conservative results for the Vulcan predictions. When nozzle 3 is

capped, the Vulcan results were in agreement with the test results.

Within the report, the CFD predictions are used to indicate the fraction of the volume over which the suppressant is predicted to be sufficient for suppression. These integral measures obtained using CFD provide

information on the inhomogeneities that goes beyond information provided by zonal models.

The results also highlight the challenges of comparing measurements in complex environments with

simulations. Challenges are reported in (1) the sensitivity of point-wise concentration measurements both to

position uncertainties in complex geometries and to the details of the suppressant injection momentum, (2)

the temporal response of thermocouples in dynamic fire environments, and (3) the relating bottle pressures

to mass-discharge rates.

The simulation results also suggest several guidelines regarding the arrangement of the suppressant

injection nozzles. Specifically, the predictions suggest that providing the suppressant into the vicinity of the

primary air inlet is most effective in providing a uniform distribution. The test results suggest that this effect

may be stronger in the simulations than in reality, but it is not possible to make the detailed assessment

within the physical nacelle simulator.

108
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