NISTIR 6196-1

Sprinkler, Smoke & Heat Vent, Draft Curtain

Interaction -- Large Scale Experiments and Model

Development

Kevin B. McGrattan

Anthony Hamins

David Stroup

United States Department of Commerce

National Institute of Standards and Technology
NISTIR 6196-1

Sprinkler, Smoke & Heat Vent, Draft Curtain

Interaction -- Large Scale Experiments and Model

Development

Kevin B. McGrattan

Anthony Hamins

David Stroup

September 1998

UNITED STA

TE

S OF AMERICA

DEPARTMENT O

F

COMME	RCE

U.S. Department of Commerce

William M. Daley, Secretary

Technology Administration

Gary R. Bachula, Acting Under Secretary for Technology

National Institute of Standards and Technology

Raymond G. Kammer, Director
Executive Summary

This study was the result of a coordinated public-private research effort to gain insight into

the interaction of sprinklers, roof vents and draft curtains through fire experiments and numerical

modeling. The work was conducted over a period of two years.

The International Fire Sprinkler, Smoke and Heat Vent, Draft Curtain Fire Test Project organized by the National Fire Protection Research Foundation (NFPRF) brought together a group of

industrial sponsors to support and plan a series of large scale tests to study the interaction of sprinklers, roof vents and draft curtains of the type found in large warehouses, manufacturing facilities,

and warehouse-like retail stores. A Technical Advisory Committee consisting of representatives

from the sponsoring organizations, the National Institute of Standards and Technology (NIST), and

other interested parties planned 39 large scale fire tests that were conducted in the Large Scale Fire

Test Facility at Underwriters Laboratories (UL) in Northbrook, Illinois. The tests were designed

to address relatively large, open-area buildings with flat ceilings, sprinkler systems, and roof venting with and without draft curtains. To simulate these conditions in the 37 m by 37 m by 15 m

high (120 ft by 120 ft by 48 ft high) main test bay, the vents, draft curtains and sprinklers were

installed on a 30 m by 30 m (100 ft by 100 ft) adjustable-height platform, 7.6 m (25 ft) off the

floor. During the tests, smoke and hot gases filled the volume enclosed by the draft curtains, and

the excess smoke flowed around the edges of the platform into a plenum space above. The smoke

in the plenum spaced was continually exhausted through a smoke abatement system.

Funding for the large scale tests, although substantial, permitted only five high rack storage

commodity fire tests to be conducted. In order to best prepare for these tests and to add to the

information available for model development and verification, 34 easily conducted and controlled

heptane spray burner tests were also performed. Recognizing that the resources in terms of funding and time were limited, the NFPRF Technical Advisory Committee reduced the number of

parameters in this study by selecting one commodity, one rack configuration and ignition source, a

single sprinkler system and one sprinkler discharge density, one commercial vent design, and one

vent/draft curtain arrangement.

In parallel with the large scale fire tests, a program was conducted at NIST to develop a numerical field model, Industrial Fire Simulator (IFS), that incorporated the physical phenomena of

the experiments. A series of bench scale experiments was conducted to develop necessary input

data for the model. These experiments generated data describing the burning rate and flame spread

behavior of the cartoned plastic commodity, thermal response parameters and spray pattern of the

sprinkler, and the effect of the water spray on the burning commodity. Simulations were first

performed for the heptane spray burner tests, where they were shown to be in good quantitative

agreement in terms of both predicting sprinkler activation times and near-ceiling gas temperature

rise. The sprinkler activation times were predicted to within about 15% of the experiments for the

first ring, 25% for the second. The gas temperatures near the ceiling were predicted to within about

15%. Next, simulations were performed and compared with unsprinklered calorimetry burns of the

cartoned plastic commodity. The heat release rates of the growing fires were predicted to within

about 20%. Simulations of the 5 cartoned plastic commodity fire tests were then performed. The

goal of these simulations was to be able to differentiate between those experiments that activated a

large number of sprinklers, and those that activated a small number. This goal has been met. The

model was also used to provide valuable insight into what occurred in the experiments, and also

what would have occurred in the event of various changes of test parameters. There are plans to

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continue the development and evaluation of the IFS model beyond publication of this report.

The overall goal of the testing and modeling program was to investigate the effect of roof vents

and draft curtains on the time, number, and location of sprinkler activations; and also the effect

of sprinklers and draft curtains on the activation time, number, and discharge rates of roof vents.

Based on the experiments and model simulations for the chosen test hardware and configurations,

the major findings in this study were:



The tests and model simulations showed that when the fire was not ignited directly under

a roof vent, venting had no significant effect on the sprinkler activation times, the number

of activated sprinklers, the near-ceiling gas temperatures, or the quantity of combustibles

consumed.



The tests and model simulations showed that when the fire was ignited directly under a roof

vent, automatic vent activation usually occurred at about the same time as the first sprinkler

activation, but the average activation time of the first ring of sprinklers was delayed. The

length of the delay depended on the difference in activation times between the vent and the

first sprinkler. 

The tests and model simulations showed that when the fire was ignited directly under a roof

vent that activated either before or at about the same time as the first sprinkler, the number

of sprinkler activations decreased by as much as 50% compared to tests performed with the

vent closed.



The tests and model simulations showed that when draft curtains were installed, up to twice

as many sprinklers activated compared to tests performed without curtains.



In one rack storage test where the ignition of the fire took place near a draft curtain and the

fuel array extended underneath the curtain, disruption of the sprinkler spray and delay in

sprinkler operation caused by the draft curtain led to a fire that consumed more commodity

compared to the other tests where the fires were ignited away from the draft curtains. This

result was demonstrated by the model simulation, as well.



The significant cooling effect of sprinkler sprays on the near-ceiling gas flow often prevented

the automatic operation of vents. This conclusion is based on thermocouple measurements

within the vent cavity, the presence of drips of solder on the fusible links recovered from

unopened vents, and several tests where vents remote from the fire and the sprinkler spray

activated. In one cartoned plastic commodity experiment, a vent did not open when the fire

was ignited directly beneath it. The model simulations could not predict this phenomenon.



Model simulations indicated that the cooling effect of sprinkler sprays reduced the total vent

discharge rate from that assumed in design calculations for unsprinklered buildings.



Model simulations showed how the activation times of the the first and second sprinklers

had a substantial impact on the overall number of activations in the plastic commodity tests.

In the simulation of one test, it was shown that a delay of approximately one minute in the

activation of the second sprinkler led to the activation of four times as many sprinklers as in

a simulation of a test with no delay. It had been suggested that these different outcomes were

due to the presence of draft curtains in the tests with the sprinkler delay, but the simulations

showed that the curtains had no effect because they were over 9 m (30 ft) away from the

ignition point.

ii
Contents

1 Introduction 1

2 Background 3

3 Heptane Spray Burner Tests, Series I 6

3.1 Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

3.1.1 Test Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

3.1.2 Test Layout . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

3.1.3 Sprinkler Protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8

3.1.4 Smoke and Heat Vent . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8

3.1.5 Heptane Spray Burner . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

3.1.6 Instrumentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

3.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

4 Heptane Spray Burner Tests, Series II 21

4.1 Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

4.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

5 Cartoned Plastic Commodity Fire Tests 33

5.1 Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

5.1.1 Fuel Package . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

5.1.2 Storage Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

5.1.3 Sprinkler Protection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

5.1.4 Test Layout . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

5.1.5 Instrumentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36

5.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37

5.2.1 Test P-1 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38

5.2.2 Test P-2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42

5.2.3 Test P-3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

5.2.4 Test P-4 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

5.2.5 Test P-5 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

5.2.6 Obscuration . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58

6 Discussion of Experimental Results 60

6.1 Effect of Vents and Draft Curtains on Sprinkler Activation Times . . . . . . . . . . 60

6.2 Effect of Vents and Draft Curtains on Number of Sprinkler Activations . . . . . . . 62

6.3 Effect of Vents and Draft Curtains on Sprinkler Discharge Pattern . . . . . . . . . 63

6.4 Effect of Sprinklers on the Number and Time of Vent Activations . . . . . . . . . . 64

6.5 Effect of Sprinklers on the Discharge Rate of Vents . . . . . . . . . . . . . . . . . 65

7 Numerical Modeling 66

7.1 Model Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66

7.1.1 Hydrodynamics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66

7.1.2 The Fire . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70

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7.1.3 Sprinkler Activation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 72

7.1.4 Sprinkler Droplet Size Distribution . . . . . . . . . . . . . . . . . . . . . 73

7.1.5 Sprinkler Spray Dynamics . . . . . . . . . . . . . . . . . . . . . . . . . . 74

7.1.6 Sprinkler Spray Interaction with Burning Commodity . . . . . . . . . . . . 76

7.1.7 Extinguishment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77

7.1.8 Numerical Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . 77

7.2 Model Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81

7.2.1 Heptane Spray Burner Test Series I . . . . . . . . . . . . . . . . . . . . . 81

7.2.2 Fire Growth Validation . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83

7.2.3 Cartoned Plastic Commodity Fire Simulations . . . . . . . . . . . . . . . 90

7.2.4 Mass Flow Rates through Vents . . . . . . . . . . . . . . . . . . . . . . . 97

8 Conclusions 101

References 103

Appendices 107

A Heptane Spray Burner Test Results (Series I) 107

A.1 Sprinkler Activation History . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107

A.2 Experimental and Simulated Temperatures . . . . . . . . . . . . . . . . . . . . . . 113

B Heptane Spray Burner Test Results, Series II 137

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List of Figures

1 Plan view of heptane spray burner test configuration, Series I. . . . . . . . . . . . . 7

2 Diagram of the fire vent used in all the tests. . . . . . . . . . . . . . . . . . . . . . 9

3 Plan view of heptane spray burner. . . . . . . . . . . . . . . . . . . . . . . . . . . 11

4 Results of Heptane Series I curtained tests with fires at location A. . . . . . . . . . 13

5 Results of Heptane Series I curtained tests with fires at location B. . . . . . . . . . 14

6 Results of Heptane Series I curtained tests with fires at location C. . . . . . . . . . 15

7 Results of Heptane Series I curtained tests with fires at location D. . . . . . . . . . 16

8 Results of Heptane Series I uncurtained tests with fires at location A. . . . . . . . . 17

9 Results of Heptane Series I uncurtained test with the fire at location B. . . . . . . . 18

10 Results of Heptane Series I uncurtained tests with fires at location C. . . . . . . . . 19

11 Results of Heptane Series I uncurtained test with fire at location D. . . . . . . . . . 20

12 Configuration of the second series of heptane spray burner tests. . . . . . . . . . . 23

13 Sprinkler activation times of Tests II-5, II-1, II-2 and II-6, Heptane Series II. . . . . 24

14 Sprinkler activation times for Tests II-3 and II-7 of Heptane Series II. . . . . . . . 25

15 Sprinkler activation times for Tests II-4 and II-8 of Heptane Series II. . . . . . . . 26

16 Sprinkler activation times for Test II-9 of Heptane Series II. . . . . . . . . . . . . . 27

17 Sprinkler activation times for Test II-10 of Heptane Series II. . . . . . . . . . . . . 28

18 Sprinkler activation times for Tests II-11 and II-12 of Heptane Series II. . . . . . . 29

19 Vent cavity temperatures for Tests II-1, II-2, II-5 and II-6, Heptane Series II. . . . . 30

20 Vent cavity temperatures for Tests II-3, II-4, II-7 and II-8, Heptane Series II. . . . . 31

21 Vent cavity temperatures for Tests II-9–12, Heptane Series II. . . . . . . . . . . . . 32

22 Configuration of the cartoned plastic commodity storage. . . . . . . . . . . . . . . 35

23 Sprinkler activation history for Plastic Test P-1. . . . . . . . . . . . . . . . . . . . 39

24 Temperature profiles from Plastic Test P-1. . . . . . . . . . . . . . . . . . . . . . . 40

25 Fire damage from Plastic Test P-1. . . . . . . . . . . . . . . . . . . . . . . . . . . 41

26 Sprinkler activation times for Plastic Test P-2. . . . . . . . . . . . . . . . . . . . . 43

27 Temperature profiles from Plastic Test P-2. . . . . . . . . . . . . . . . . . . . . . . 44

28 Fire damage from Plastic Test P-2. . . . . . . . . . . . . . . . . . . . . . . . . . . 45

29 Sprinkler activation times for Plastic Test P-3. . . . . . . . . . . . . . . . . . . . . 47

30 Temperature profiles from Plastic Test P-3. . . . . . . . . . . . . . . . . . . . . . . 48

31 Fire damage from Plastic Test P-3. . . . . . . . . . . . . . . . . . . . . . . . . . . 49

32 Sprinkler activation times from Plastic Test P-4. . . . . . . . . . . . . . . . . . . . 51

33 Temperature profiles from Plastic Test P-4. . . . . . . . . . . . . . . . . . . . . . . 52

34 Fire damage from Plastic Test P-4. . . . . . . . . . . . . . . . . . . . . . . . . . . 53

35 Sprinkler activation times from Plastic Test P-5. . . . . . . . . . . . . . . . . . . . 55

36 Temperature profiles from Plastic Test P-5. . . . . . . . . . . . . . . . . . . . . . . 56

37 Fire damage from Plastic Test P-5. . . . . . . . . . . . . . . . . . . . . . . . . . . 57

38 Local optical density measurements for the Plastic tests. . . . . . . . . . . . . . . . 59

39 Temperatures in the vent cavity during Plastic Test P-2. . . . . . . . . . . . . . . . 65

40 Heat release rate per unit area of the cartoned plastic commodity. . . . . . . . . . . 72

41 Droplet size distribution from a typical industrial fire sprinkler. . . . . . . . . . . . 75

42 Photograph of Plastic Test P-5. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79

43 Numerical simulation of Plastic Test P-5. . . . . . . . . . . . . . . . . . . . . . . 80

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44 Total heat release rates for the two, three and four tier cartoned plastic burns (UL). . 83

45 Convective heat release rates for cartoned plastic burns (UL). . . . . . . . . . . . . 84

46 Simulation of three tier calorimetry burn. . . . . . . . . . . . . . . . . . . . . . . 85

47 Heat release rates for  and  cartoned plastic burns (FM). . . . . 86

48 Heat release rates for 	 and 
 cartoned plastic burns (FM). . . . . 87

49 Results of Plastic Test P-1 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 91

50 Results of Plastic Test P-1 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 91

51 Results of Plastic Test P-4 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 92

52 Results of Plastic Test P-5 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 92

53 Heat release rates for numerical simulations of Plastic Tests P-1 and P-4. . . . . . . 93

54 Results of Plastic Test P-2 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 94

55 Results of Plastic Test P-2 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 94

56 Heat release rates for numerical simulations of Plastic Test P-2. . . . . . . . . . . . 95

57 Results of Plastic Test P-3 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 96

58 Results of Plastic Test P-4 simulation. . . . . . . . . . . . . . . . . . . . . . . . . 96

59 Heat release rates for numerical simulations of Plastic Tests P-3 and P-4. . . . . . . 97

60 Snapshots of the simulations of Plastic Tests P-3 and P-4 after 5 min. . . . . . . . . 98

61 Mass flow rate through a single vent. . . . . . . . . . . . . . . . . . . . . . . . . . 99

62 Sprinkler activation history for Test I-1, Heptane Series I. . . . . . . . . . . . . . . 107

63 Sprinkler activation history for Test I-2, Heptane Series I. . . . . . . . . . . . . . . 107

64 Sprinkler activation history for Test I-3, Heptane Series I. . . . . . . . . . . . . . . 108

65 Sprinkler activation history for Test I-4, Heptane Series I. . . . . . . . . . . . . . . 108

66 Sprinkler activation history for Test I-5, Heptane Series I. . . . . . . . . . . . . . . 108

67 Sprinkler activation history for Test I-6, Heptane Series I. . . . . . . . . . . . . . . 108

68 Sprinkler activation history for Test I-7, Heptane Series I. . . . . . . . . . . . . . . 109

69 Sprinkler activation history for Test I-8, Heptane Series I. . . . . . . . . . . . . . . 109

70 Sprinkler activation history for Test I-9, Heptane Series I. . . . . . . . . . . . . . . 109

71 Sprinkler activation history for Test I-10, Heptane Series I. . . . . . . . . . . . . . 109

72 Sprinkler activation history for Test I-11, Heptane Series I. . . . . . . . . . . . . . 110

73 Sprinkler activation history for Test I-12, Heptane Series I. . . . . . . . . . . . . . 110

74 Sprinkler activation history for Test I-13, Heptane Series I. . . . . . . . . . . . . . 110

75 Sprinkler activation history for Test I-14, Heptane Series I. . . . . . . . . . . . . . 110

76 Sprinkler activation history for Test I-15, Heptane Series I. . . . . . . . . . . . . . 111

77 Sprinkler activation history for Test I-16, Heptane Series I. . . . . . . . . . . . . . 111

78 Sprinkler activation history for Test I-17, Heptane Series I. . . . . . . . . . . . . . 111

79 Sprinkler activation history for Test I-18, Heptane Series I. . . . . . . . . . . . . . 111

80 Sprinkler activation history for Test I-19, Heptane Series I. . . . . . . . . . . . . . 112

81 Sprinkler activation history for Test I-20, Heptane Series I. . . . . . . . . . . . . . 112

82 Sprinkler activation history for Test I-21, Heptane Series I. . . . . . . . . . . . . . 112

83 Sprinkler activation history for Test I-22, Heptane Series I. . . . . . . . . . . . . . 112

84 Plan view of heptane spray burner test configuration, Series I. . . . . . . . . . . . . 114

85 Experimental and simulated temperatures for Test I-1, Heptane Series I. . . . . . . 115

86 Experimental and simulated temperatures for Test I-2, Heptane Series I. . . . . . . 116

87 Experimental and simulated temperatures for Test I-3, Heptane Series I. . . . . . . 117

88 Experimental and simulated temperatures for Test I-4, Heptane Series I. . . . . . . 118

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89 Experimental and simulated temperatures for Test I-5, Heptane Series I. . . . . . . 119

90 Experimental and simulated temperatures for Test I-6, Heptane Series I. . . . . . . 120

91 Experimental and simulated temperatures for Test I-7, Heptane Series I. . . . . . . 121

92 Experimental and simulated temperatures for Test I-8, Heptane Series I. . . . . . . 122

93 Experimental and simulated temperatures for Test I-9, Heptane Series I. . . . . . . 123

94 Experimental and simulated temperatures for Test I-10, Heptane Series I. . . . . . 124

95 Experimental and simulated temperatures for Test I-11, Heptane Series I. . . . . . 125

96 Experimental and simulated temperatures for Test I-12, Heptane Series I. . . . . . 126

97 Experimental and simulated temperatures for Test I-13, Heptane Series I. . . . . . 127

98 Experimental and simulated temperatures for Test I-14, Heptane Series I. . . . . . 128

99 Experimental and simulated temperatures for Test I-15, Heptane Series I. . . . . . 129

100 Experimental and simulated temperatures for Test I-16, Heptane Series I. . . . . . 130

101 Experimental and simulated temperatures for Test I-17, Heptane Series I. . . . . . 131

102 Experimental and simulated temperatures for Test I-18, Heptane Series I. . . . . . 132

103 Experimental and simulated temperatures for Test I-19, Heptane Series I. . . . . . 133

104 Experimental and simulated temperatures for Test I-20, Heptane Series I. . . . . . 134

105 Experimental and simulated temperatures for Test I-21, Heptane Series I. . . . . . 135

106 Experimental and simulated temperatures for Test I-22, Heptane Series I. . . . . . 136

107 Temperatures for Test II-1, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 137

108 Temperatures for Test II-2, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 138

109 Temperatures for Test II-3, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 139

110 Temperatures for Test II-4, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 140

111 Temperatures for Test II-5, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 141

112 Temperatures for Test II-6, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 142

113 Temperatures for Test II-7, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 143

114 Temperatures for Test II-8, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 144

115 Temperatures for Test II-9, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 145

116 Temperatures for Test II-10, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 146

117 Temperatures for Test II-11, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 147

118 Temperatures for Test II-12, Heptane Series II. . . . . . . . . . . . . . . . . . . . . 148

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List of Tables

1 Results of the heptane spray burner Series I. . . . . . . . . . . . . . . . . . . . . . 12

2 Results of the heptane spray burner Series II. . . . . . . . . . . . . . . . . . . . . 22

3 Results of the cartoned plastic commodity tests. . . . . . . . . . . . . . . . . . . . 37

4 Summary of the 39 Large Scale Experiments. . . . . . . . . . . . . . . . . . . . . 61

5 Input parameters used for the fire simulations. . . . . . . . . . . . . . . . . . . . . 78

6 Results of the heptane spray burner Series I simulations. . . . . . . . . . . . . . . 82

7 Experimental and numerical flame spread data for two tier calorimetry burn. . . . . 88

8 Experimental and numerical flame spread data for the four tier calorimetry burn. . . 89

List of Abbreviations

FMRC Factory Mutual Research Corporation

gpm (U.S.) gallons per minute (1 gallon = 3.785412 L)

IIT Illinois Institute of Technology

LIFT Lateral Ignition and Flame Spread Test

NFPA National Fire Protection Association

NFPRF National Fire Protection Research Foundation

NIST National Institute of Standards and Technology

RTI Response Time Index

SP Swedish National Testing and Research Institute

UL Underwriters Laboratories

viii
Acknowledgments

The authors would like to thank the following individuals and organizations for their assistance in this project: Frederick Mulhaupt and Douglas Brown at the National Fire Protection Research Foundation for providing us the opportunity to participate in the International Fire Sprinkler, Smoke and Heat Vent, Draft Curtain Fire Test Project; the members of the Technical Advisory

Committee for their advice and expertise in reviewing the test data and the drafts of the final report;

David Sheppard, William Carey, Daniel Steppan and Pravin Gandhi at Underwriters Laboratories

and Joe Hankins at Factory Mutual for sharing their test data and their experience in commodity

fire testing; Linda Blevins and Glenn Forney of NIST for their work in characterizing sprinkler

sprays; Leonard Cooper and Jay McElroy of NIST for their efforts to characterize the thermal

response properties of fusible vent links; Stephen DiGiovanni of the University of Maryland and

Roy McLane of NIST for their work in characterizing the burning rates of the plastic commodity

used in the tests; and David Evans, Daniel Madrzykowski and Kenneth Steckler of NIST for their

assistance in assembling this report.

ix
1 Introduction

There has been a long-standing debate in the fire protection community about the combined use of

roof vents, draft curtains (curtain boards) and sprinklers. Numerous studies have been conducted

over the past few decades, yet many questions remain about the interaction of these devices. As a

result, a coordinated public-private research effort was organized. A group of industrial sponsors

was brought together by the National Fire Protection Research Foundation (NFPRF) to support and

plan a series of large scale experiments using both a heptane spray burner and cartoned polystyrene

cups (Group A plastic) as fire sources. A Technical Advisory Committee (TAC) made up of representatives of the sponsoring organizations, the National Institute of Standards and Technology

(NIST), and other invited participants was created by the NFPRF to guide the studies. The TAC selected a test facility and a sprinkler system, roof vent, draft curtain design for installation in the test

facility in order to simulate fire protection systems found in warehouses, warehouse retail stores

and manufacturing facilities. In all, 39 tests were specified by the TAC. Fire modeling and supporting laboratory experiments provided by NIST aided in the planning of large scale experiments

and in the analysis of the data.

There were three objectives of the study. First, there was a need to understand how sprinklers,

vents and draft curtains interact. Second, there was a need to develop a numerical model capable

of predicting multiple sprinkler activations and the heat release rate from a burning fuel array both

before and after sprinkler activation with sufficient accuracy to reliably differentiate between fire

scenarios that would produce a large versus small number of sprinkler activations. Third, there

was a desire to look beyond current building practices and generate experimental data, along with

a numerical modeling capability, that could be used to evaluate proposed changes to codes and

standards.

All 39 experiments planned by the TAC were conducted in the Large Scale Fire Test Facility at

Underwriters Laboratories in Northbrook, Illinois. The experiments were divided into three series:

an initial set of 22 heptane spray burner tests (Heptane Series I) [1], 5 cartoned plastic commodity

tests (Plastic Series) [2], and 12 additional heptane spray burner tests (Heptane Series II) [2]. Many

of the test parameters did not change throughout the entire project. The ceiling heights were set at

either 7.6 m (25 ft) or 8.2 m (27 ft), the storage height was nominally 6.1 m (20 ft) in the cartoned

plastic tests, the sprinkler system consisted of 0.64 in orifice upright sprinklers spaced 3 m (10 ft)

apart delivering a 0.34 L/(s  m ) (0.50 gpm/ft ) discharge density, and the 1.2 m by 2.4 m (4 ft

by 8 ft) vents used were of the same design, from the same manufacturer, and UL listed1. The

parameters that did change were fire size, fire/ignition position, mode of vent operation, and draft

curtain placement.

A large effort was made to develop numerical techniques that could be used to interpret and

potentially supplement the physical experiments. This work is a major undertaking by NIST. The

Large Eddy Simulation (LES) Fire Model [3, 4] is a computational fluid dynamics (CFD) code that

solves the differential equations that govern the transport of smoke and hot gases from a fire. The

model being developed and applied in this project is referred to as the Industrial Fire Simulator

(IFS). A series of bench scale experiments was conducted by NIST to develop necessary input

data for the model. These experiments generated data describing the burning rate and flame spread

1Throughout this report, measured quantities will be reported with units conforming to the modern metric system

(SI) with traditional units in parentheses. The exception is where the units are essential to the description of the object,

apparatus, etc.

1
behavior of the cartoned plastic commodity, thermal response parameters and spray pattern of the

sprinkler, and the effect of the water spray on the burning commodity. During its development the

performance of the IFS model was assessed by comparing its results to those of the heptane spray

burner experiments and calorimetry burns of the cartoned plastic commodity. The IFS model was

used in planning experiments and also in the analysis to supplement the results of the relatively

small number of cartoned plastic commodity tests that were conducted.

2
2 Background

Roof vents (usually accompanied by draft curtains) and sprinklers are two hardware systems used

in buildings to provide fire protection for rack storage of materials. While sprinklers are intended

to suppress and control the fire, roof vents are intended to aid the fire fighters. According to NFPA

204M, Guide for Smoke and Heat Venting [5], Paragraph 1-1.2:

...large undivided floor areas present extremely difficult fire fighting problems, since

the fire department must enter these areas in order to combat fires in central portions

of the building. If the fire department is unable to enter because of the accumulation

of heat and smoke, fire fighting efforts may be reduced to a futile application of hose

streams to perimeter areas while fire consumes the interior. Windowless buildings also

present similar fire fighting problems. One fire protection tool that may be a valuable

asset for fire fighting operations in such buildings is smoke and heat venting.

The design practice for sprinklers is given in NFPA 231C, Standard for Rack Storage of Materials [6]. The design practice for roof vents and draft curtains is given in NFPA 204M. These are

independent documents. There is no nationally recognized standard for the combined installation

of sprinklers and roof vents. Indeed, NFPA 204M, Paragraph 6-1 states:

A broadly accepted equivalent design basisfor using both sprinklers and ventstogether

for hazard control (e.g. property protection, life safety, water usage, obscuration, etc.)

has not been universally recognized.

Even though the practice of installing vents in sprinklered buildings has been debated for decades,

and in spite of several projects involving large scale tests and numerical modeling, there is still

disagreement about how roof vents and draft curtains affect the time, number and location of

sprinkler activations; and how sprinklers and draft curtains affect the activation time, number and

discharge rates of roof vents. As a result, there is a great disparity among building codes as to the

proper treatment of these fire protection devices. The issue is even more important today because of

the popularity of large warehouse-type retail stores that combine the function of sales and storage.

Typically these stores will contain merchandise stacked up to 6 m (20 ft), with ceilings as high as

9 m (30 ft), and in areas over 10,000 m (100,000 ft ).

As part of the planning for the experiments described in this report, past experiments and

modeling efforts involving vents and draft curtains (with and without sprinklers) were considered:

Roof Ventilation Requirements for Industrial Plants (1954–1955) Armour Research Foundation

(former name of IIT Research). Experimental and modeling study of unsprinklered, vented

facilities [7].

Underwriters Laboratories Roof Vent Testing (1964) Sprinkler/ventstudy in a 60 by 60 by 16 ft

space [8].

Thomas and Hinkley (1964) Experimental and modeling studies at the Fire Research Station,

UK [9].

Portsmouth Fire Test (1966) Colt Industries(UK). Fullscale fire testsin Victoria Barracks, Portsmouth [10].

3
Rack Storage and Rubber Tire Tests (late 1960’s, early 1970’s) Factory Mutual Research Corporation (FMRC). Experimental studies of sprinklered facility with perimeter venting [11].

FMRC Modeling Study (1974) 1:12.5 scale model of large test facility built with vents [12].

IIT Research (1982) Fire Venting of Sprinklered Buildings. Large scale experiments of sprinkler performance in a vented facility. Includes a literature review of work done up to that

date [13].

Ghent Fire Tests (1989) Colt Industries and Fire Research Station (UK). Thirty-three large scale

tests of vents and sprinklers in a 50 m by 20 m by 10 m facility with 40 roof vents [14].

SP Interaction between Sprinklers and Fire Vents (1992) Swedish National Testing and Research

Institute (SP). A series of experiments involving one or two vents and a single sprinkler [15].

A summary of the research in Europe involving smoke and heat vents can be found in a

technical report from SP [16].

International Seminar on Sprinklers and Vents (1992) Fire Research Station (UK). A discussion of past work in the area, particularly the experiments at Ghent. Tentative guidelines are

put forward to address the problem.

Early work by Thomas and Hinkley [9] in the 1960’s on the performance of roof vents led to

the tentative recommendation that sprinklers should be operated before vents because the cooling

of the hot upper layer gases by the opening of vents could delay sprinkler activation. This recommendation found its way into many codes of practice. Since that time, there have been several

large scale experimental and numerical studies performed to explore this issue. Notable among

these is the Ghent fire test project, which consisted of hexane fires in a 50 m by 20 m (164 ft by

66 ft) facility with 40 roof vents, each with an area of 1.64 m (18 ft ). It has been claimed by

proponents of venting that these tests showed that venting does not appreciably affect the activation

of the first sprinkler, but does significantly reduce both the number of sprinklers opening within

60 s of the first as well as the total number of sprinklers operating. However, in all the vented

fires, enough sprinklers operated to wet the fire and its immediate surroundings [17]. Hinkley’s

modeling work [18] reached similar conclusions, and he states that “Venting is unlikely to delay

the operation of sprinklers in a compartment of such a large area that the hot gases would not

have reached the boundaries before sprinklers operate”. He goes on to say that for smaller compartments, slowly growing fires, and/or when draft curtains are used to build up the smoke layer,

the vents can cause a small delay in sprinkler operation, but that this is a “second order” effect

compared to the rate of fire growth and operating characteristics of the sprinkler.

However, a position paper by N.E. Gustafsson of Industrial Mutual, Helsinki, interprets the

results of the Ghent tests completely differently [19]. He argues that for the rapidly growing fires,

a significant delay in sprinkler activation was caused by the presence of vents. Even though the

delay was about 10 to 20 s in most cases, this allowed the fire to grow from 10.2 MW in the

unvented case to as much as 14.2 MW in one of the vented cases. He also cites the inability of the

sprinkler system to surround the fire in the vented cases.

Rack storage fire tests conducted at Factory Mutual in the late 1960’s raised the issue of the

possible detrimental effects of smoke and heat vents in sprinklered buildings [11]. These experiments have been analyzed by a number of researchers, and the conclusions have been criticized

4
based mainly on the fact that ventilation was from doors and eave line windows rather than roof

vents, thus only the potential disadvantage of venting (higher oxygen content) was maintained.

The series of fire tests conducted by IIT Research Institute in 1982 [13, 20] showed no increase in

oxygen in the air supplying the fire. The issue was not addressed in the modeling studies, nor in

the Ghent or SP experiments due to the use of prescribed burning rates.

The experiments performed by IIT Research Institute with both sprinklers and vents rated

at 74
 C (165
 F) showed a slight reduction in water demand for vent areas considered “large”,

but no change in water demand for a more limited vent area. The report also notes measurable

reductionsin near-ceiling gas temperatures away from the fire. The study concluded that roof vents

do not impair the performance of 74
 C (165
 F) standard response sprinklers, but they provide only

marginal benefit. For 141
 C (286
 F) sprinklers in the presence of propane fires, the benefits of

reduced water demand and near-ceiling gas temperatures were more pronounced.

Draft curtains are usually installed to aid roof vents by creating a deeper layer of smoke and

hot gases than would otherwise be formed in large, open area buildings. Recent experiments at

Factory Mutual [21] investigated the effect of draft curtains on a sprinkler system without roof

vents. As part of a test program to evaluate the effectiveness of various sprinkler designs, two

pairs of high rack storage cartoned plastic fire tests were conducted with and without draft curtains

installed. First, two tests were performed without curtains. Then the tests were repeated after draft

curtains had been installed. The fires were ignited near the intersection of two draft curtains. With

the curtains installed, the fire opened more sprinklers and caused more damage to the fuel array,

apparently due to the disruption of the sprinkler discharge pattern and the lack of prewetting of the

commodity (FMRC Standard Plastic test commodity). The report concluded that “...the presence

of draft curtains close to the fire origin will (1) result in the development of a more severe fire and

(2) deleteriously affect sprinkler protection.”

Most of the past work on sprinkler, vent, and draft curtain interaction hasfocussed on the effects

of vents and draft curtains on a sprinkler system. There has been much less work examining the

effects of sprinklers on vents. The work done at SP [15] includes experiments designed to measure

the effects of a single sprinkler on the temperature and velocity of gases flowing through a 1 m

by 2 m (3.3 ft by 6.6 ft) ceiling vent centered in a 7.5 m by 15 m by 6 m high (25 ft by 50 ft

by 20 ft high) test space that was open on two sides. The fire source was a 1.5 MW propane

burner, located 6 m from the center of the vent. A single sprinkler with a flow rate of either 80 or

100 L/min (21 or 26 gpm) was installed at various locations upstream and downstream of the vent.

The conclusions drawn by SP were that when the sprinkler was installed upstream of the vent (i.e.

between the fire and the vent), it had a “significant” impact on the discharge rate of the vent, with

the impact greatest when it was near the vent and decreasing with distance from the vent. When

the sprinkler was installed downstream of the vent, its influence on the vent discharge was regarded

as “negligible”. No attempt was made in the SP study to examine the effects of a sprinkler on the

activation of a thermally-responsive automatic vent.

5
3 Heptane Spray Burner Tests, Series I

In January, 1997, a series of 22 heptane spray burner experiments was conducted at the Large Scale

Fire Test Facility at Underwriters Laboratories (UL) in Northbrook, Illinois [1]. The objective

of the experiments was to characterize the temperature and flow field for fire scenarios with a

controlled heat release rate in the presence of sprinklers, draft curtains and a single vent. The

results of the experiments were used to evaluate the predictive capability of the IFS model, and also

to provide guidance as to the interaction of vents and draft curtains with sprinklers for planning the

cartoned plastic commodity fires.

3.1 Description

A number of components of this first series of tests were carried over into the cartoned plastic

commodity tests and the second series of heptane spray burner tests. The same test space was

used for all the testing, the same vent and sprinkler type, the same sprinkler configuration, and

almost the same ceiling height. Below is a description of the first series of heptane spray burner

tests. Changes to these parameters in the subsequent test series will be discussed in the appropriate

sections. The Report of Test from UL may be found in Ref. [1].

3.1.1 Test Facility

The Large Scale Fire Test Facility at UL contains a 37 m by 37 m (120 ft by 120 ft) main fire test

cell, equipped with a 30.5 m by 30.5 m (100 ft by 100 ft) adjustable height ceiling. The height

of the ceiling may be adjusted by four hydraulic rams up to a maximum height of 14.6 m (48 ft).

A flexible design sprinkler piping system was available at the ceiling to permit any arrangement

of sprinkler spacing with minimum pressure losses. The sprinklers were provided with water

from a 757 kL (200,000 gal) water reservoir using two 190 L/s (3,000 gpm) water pumps rated at

1.03 kPa (150 psi). The water pumps were computer controlled to maintain a constant pressure on

the sprinklers.

The exhaust flow rate in the test facility could be adjusted up to a maximum of 28 m /s (60,000

ft /min). Four 1.5 m (5 ft) diameter inlet ducts provided make up air and were located at the walls

3 m (10 ft) above the test floor to minimize any induced drafts during the tests. The combustion products from the fire test were exhausted through a regenerative smoke abatement system.

During the heptane tests, the exhaust was maintained at a minimum operating rate of 11 m /s

(24,000 ft /min).

3.1.2 Test Layout

The layout of the first series of heptane spray burner tests is shown in Fig. 1. One 1.2 m by 2.4 m

(4 ft by 8 ft) vent was installed among 49 upright sprinklers on a 3 m by 3 m (10 ft by 10 ft)

spacing. The ceiling was raised to a height of 7.6 m (25 ft) and instrumented with thermocouples

and other measurement devices (see section on instrumentation). The ceiling was constructed of

0.6 m by 1.2 m by 1.6 cm (2 ft by 4 ft by 5/8 in) UL fire rated Armstrong Ceramaguard (Item

6
North Wall

Adjustable Height Ceiling













Draft Curtains





     

               







  



  

    !  

A

B

C

D

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

"

98 91 84 81 78 75 72

99 92 85 82 79 76 73

100 93 86 83 80 77 74

101 94 87 62 58 54 50

102 95 88 63 59 55 51

103 96 89 64 60 56 52

104 97 90 65 61 57 53

68

69

70

67 66

71

#

#

#

#

#

#

#

43

44

45

46

47

48

49

Figure 1: Plan view of heptane spray burner configuration, Series I. The sprinklers are indicated by the solid circles and are spaced 3 m (10 ft) apart. The number beside each sprinkler

location indicates the channel number of the nearest thermocouple. The vent dimensions are

4 ft by 8 ft. The boxed letters A, B, C and D indicate burner positions. Corresponding to each

burner position is a vertical array of thermocouples. Thermocouples 1–9 hang 7, 22, 36, 50,

64, 78, 92, 106 and 120 in from the ceiling, respectively, above Position A. Thermocouples 10

and 11 are positioned above and below the ceiling tile directly above Position B, followed by

12–20 that hang at the same levels below the ceiling as 1–9. The same pattern is followed at

Positions C and D, with thermocouples 21–31 at C and 32–42 at D.

7
602B) ceiling tiles2

, suspended from 3.8 cm (1.5 in) wide steel angle brackets. The manufacturer

reported the thermal properties of the material to be: specific heat 753 J/kg  K, thermal diffusivity

%$&')(%* m /s, conductivity 0.0611 W/m  K, and density 313 kg/m .

Draft curtains 1.8 m (6 ft) deep were installed for 16 of the 22 tests, enclosing an area of about

450 m (4,800 ft ). The curtains were constructed of 1.4 m (54 in) wide sheets of 18 gauge sheet

metal. The seams in the draft curtains were connected with aluminum tape.

3.1.3 Sprinkler Protection

The sprinklers used in both series of heptane spray burner tests plus the cartoned plastic commodity tests were Central ELO-231 (Extra Large Orifice) uprights. The orifice diameter of this

sprinkler is reported by the manufacturer to be nominally 0.64 in, the reference actuation temperature is reported by the manufacturer to be 74
 C (165
 F). The RTI (Response Time Index) and

C-factor (Conductivity factor) were reported by UL3to be 148 (m  s)

+,

(268 (ft  s)

+,

) and 0.7 (m/s)

+,

(1.3 (ft/s)

+,

), respectively [1]. When installed, the sprinkler deflector was located 8 cm (3 in) below

the ceiling. The thermal element of the sprinkler was located 11 cm (4.25 in) below the ceiling.

The sprinklers were installed with 3 m by 3 m (10 ft by 10 ft) spacing in a system designed to deliver a constant 0.34 L/(s  m ) (0.50 gpm/ft ) discharge density when supplied by a 131 kPa (19 psi)

discharge pressure 4.

Branch lines were located 3 m (10 ft) apart. The branch lines were 2 in Schedule 40 pipe. The

lines had a 6.032 cm (2.375 in) outer diameter and were mounted with a 22 cm (8.5 in) nominal

clearance from the centerline of the pipe to the ceiling. The manifolds were mounted below the

branch lines and were used to plumb the water to the branch lines. The manifolds were 6.4 cm

(2.5 in) Schedule 40 pipe with an outer diameter of 7.303 cm (2.875 in). They were mounted with

a 1.04 m (40.75 in) nominal clearance from the centerline of the pipe to the ceiling. Although not

shown in Fig. 1, the branch lines ran from north to south. The branch lines and manifolds were

filled with water during all tests.

3.1.4 Smoke and Heat Vent

A single UL listed double leaf fire vent with steel covers and steel curb was installed in the adjustable height ceiling in the position shown in Fig. 1. This vent was selected in collaboration

with the NFPRF Technical Advisory Committee who sponsored the large scale tests. The vent is

designed to open manually or automatically. The vent doors were recessed into the ceiling about

0.3 m (1 ft). A diagram of the vent used in shown in Fig. 2.

2The mention of particular manufacturer’s products does not constitute endorsement by NIST, nor does it indicate

that the products are necessarily those best suited for the intended purpose.

3The RTI and C-factor reported here are based on the equation governing the thermal activation of the fusible link

of the sprinkler -!./ -1032546

RTI

7.8:9;./=< 9?>

RTI

7./ 9;.A@ <

where ./is the link temperature, .8 is the gas temperature in the neighborhood of the link, .@ is the temperature of

the sprinkler mount, and 6 is the gas velocity.

4The flow rate from a sprinkler is proportional to the square root of the discharge pressure. The proportionality

constant is often referred to as the K-factor, usually reported in units of gallons per minute divided by the square root

of pounds per square inch. For the sprinkler used in these tests, the K-factor was approximately 11.4 gpm/(psi) BC

.

8
Figure 2: Diagram of the fire vent used in all the tests. Note that 12

=0.3048 m and 4-

0 =1.2192 m.

The vent could be operated manually or automatically. In cases where manual operation was

performed, the opening times were chosen so that the vent would either be opened about 25 s prior

to the first sprinkler activation, or 25 s after the first sprinkler activation. Preliminary calculations

indicated that the first sprinkler in most cases would activate 60 to 70 s after ignition. In tests where

automatic operation of the vents was desired, UL listed fusible links rated at either 74
 C (165
 F)

or 100
 C (212
 F) were installed. In most tests, the 74
 C link was used. To determine the thermal

response properties of the fusible link, a plunge tunnel test was performed on a representative

link assembly that consisted of a fusible link rated at 74
 C bolted to a steel tab that was welded

to a steel support bar. The link was spring-loaded with a force of 12.2 N (2.75 lbf). The link

was instrumented with a thermocouple attached to the center of its rear surface. The tunnel air

temperature was 134
 C (273
 F), the tunnel velocity was 2.5 m/s (8.2 ft/s), the ambient temperature

was 25
 C (77
 F). When inserted into the tunnel perpendicularto the flow direction, the link reached

74
 C after 34 s and remained at that temperature until it fused at 68 s. When parallel to the flow,

9
the link reached 74
 C after 35 s and remained at this temperature until it fused at 63 s [22]. The

interval between the time when the link reached its activation temperature and the time when it

fused was significant, suggesting that a one parameter model of link activation may not suffice to

fully characterize the thermal response of the link [23]. However, for the present study, an effective

bulk RTI5. for the link assembly based on the fusing time was calculated to be between 167 and

180 (m  s)

+,

(302 and 326 (ft  s)

+,

).

3.1.5 Heptane Spray Burner

The heptane spray burner consisted of a 1 m by 1 m (40 in by 40 in) square of 1/2 in pipe supported

by four cement blocks 0.6 m (2 ft) off the floor (Fig. 3). Atomizing spray nozzles were used to

provide a free spray of heptane that was then ignited. The number and locations of the atomizing

nozzles used depended on the maximum heptane flow rate expected during a test. The two configurations used were nozzles in Positions A, B, C and G for fires less than or equal to 5 MW, and A,

B, C, D, E, F and G for fires greater than 5 MW.

For all but one of the tests, the total heat release rate from the fire was manually ramped up

following the curve

6

EGFD EIHKJML D NPO 

where EQHF D  MW and N F 
s (N F 
 s was used in Test I-16). The fire growth rate

was intended to approximate the estimated growth rate of the cartoned plastic commodity burns

conducted at FMRC [21]. The fire growth curve was followed until a specified fire size wasreached

or the first sprinkler activated. After either of these events, the fire size was maintained at that level

until conditions reached roughly a steady state, i.e. the temperatures recorded near the ceilings

remained steady and no more sprinkler activations occurred.

The flow of heptane to the burner was manually controlled using two float type flow meters.

The two flow meters were connected in parallel. The first flow meter had a resolution of 0.02 gpm

and a range of 0.68 L/min (0.18 gpm) to 9.1 L/min (2.4 gpm). The second flow meter had a

resolution of 1.1 L/min (0.3 gpm) and a range of 0.91 L/min (0.24 gpm) to 11 L/min (3.0 gpm). The

heat release rate from the burner was confirmed by placing it under the large product calorimeter

at UL, ramping up the flow of heptane in the same manner as in the tests, and measuring the total

and convective heat release rates. It was found that the convective heat release rate was 0.65R0.02

of the total.

3.1.6 Instrumentation

The instrumentation for the tests consisted of thermocouples, gas analysis equipment, and pressure

transducers. The locations of the instrumentation are referenced in the plan view of the facility

5The term “bulk RTI” indicates that the temperature of the object is assumed to be governed by the equation

-!./

-10S2

46

RTI T

7.8U9V./=<

where ./ is the temperature of the object and .8is the temperature of the gas streaming past the object with speed 6.

This definition of RTIT does not explicitly account for conductive losses from the object to surrounding objects, and

should not be confused with the definition of RTI applied to fusible sprinkler links in this report.

6Because the flow rate of fuel was controlled manually, the growth curve for each test differed slightly.

10
Fuel Supply 

 A     W    



 A 



  

  

X X

X

X

X X

X

X

H A

G

F

E D

B

C

Figure 3: Plan view of heptane spray burner.

(Fig. 1).

Temperature measurements were recorded at 104 locations. Type K 0.0625 in diameter Inconel7

sheathed thermocouples were positioned to measure (i) temperatures near the ceiling, (ii)

temperatures of the ceiling jet, and (iii) temperatures near the vent. The thermocouples numbered

50–65 were positioned near the sprinklers, 10 cm (4 in) below the ceiling. These were intended

to measure near-sprinkler gas temperatures as well as to detect sprinkler activation when wetted.

Thermocouples 66–104 were placed 5 cm (2 in) below the ceiling. Thermocouples 43–49 ran

down the centerline of the vent at the level of the ceiling, and were spaced 0.3 m (1 ft) apart.

Thermocouples 1–42 were mounted on arrays hanging above each fire location. The positions are

noted in the caption to Fig. 1.

Oxygen, carbon dioxide and carbon monoxide sampling probes were placed at the ground

(5 cm (3 in) from the floor, 2 m (6 ft) from the burner), and at the vent (15 cm (6 in) below the

ceiling, vent center). The oxygen concentrations were measured with paramagnetic analyzers with

a resolution of 100 ppm. The carbon monoxide and carbon dioxide concentrations were measured

with nondispersive infrared (NDIR) analyzers. The CO meters had a resolution of 10 ppm, the

CO

meters 100 ppm.

One differential pressure was measured across the vent. The static pressure taps were mounted

0.15 m (6 in) above the top of the vent and 0.15 m below the ceiling. Both pressure taps were

mounted at the center of the vent. The differential pressure was measured with a 0 to 133 Pa (0 to

1 Torr) differential pressure transducer.

All thermocouple, pressure transducer and gas species data were collected electronically at a

2 s scan rate.

7

Inconelis a registered tradename of INCO Alloys, Inc., 3800-T Riverside Drive,P.O. Box 1958, Huntington, West

Virginia, 25705-1771.

11
3.2 Y

Results

The test results are summarized in Table 1. Sprinkler activation times for the 22 tests can be found

in Figs. 4–11. Thermocouple temperatures are included in Appendix A, along with a comparison

of the results of the experiments with those of the numerical model.

Heptane Spray Burner Test Series I

Test Burner Vent First No. of Draft Heat Release Rate

No. Pos. Operation Sprinkler Sprinklers Curtains MW @ s

I-1 B Closed 1:05 11 Yes 4.4 @ 50

I-2 B Manual (0:40) 1:06 12 Yes 4.4 @ 50

I-3 B Manual (1:30) 1:04 12 Yes 4.4 @ 50

I-4 C Closed 1:00 10 Yes 4.4 @ 50

I-5 C Manual (0:40) 1:12 9 Yes 4.4 @ 50

I-6 C Manual (1:30) 1:02 8 Yes 4.4 @ 50

I-7 C 74
 C link (DNO) 1:10 10 Yes 4.4 @ 50

I-8 B 74
 C link (9:26) 1:00 11 Yes 4.4 @ 50

I-9 D 74
 C link (DNO) 1:10 12 Yes 4.4 @ 50

I-10 D Manual (0:40) 1:12 13 Yes 4.4 @ 50

I-11 D 74
 C link (4:48) N/A N/A Yes 4.4 @ 50

I-12 A Closed 1:08 14 Yes 4.4 @ 50

I-13 A 74
 C link (1:04) 1:09 5 Yes 6.0 @ 60

I-14 A Manual (0:40) 1:14 7 Yes 5.8 @ 60

I-15 A Manual (1:30) 1:04 5 Yes 5.8 @ 60

I-16 A 74
 C link (1:46) 1:46 4 Yes 5.0 @ 110

I-17 B 100
 C link (DNO) 0:58 4 No 4.6 @ 50

I-18 C 100
 C link (DNO) 0:58 4 No 3.7 @ 50

I-19 A 100
 C link (10:00) 0:56 10 No 4.6 @ 50

I-20 A 74
 C link (1:20) 0:54 4 No 4.2 @ 50

I-21 C 74
 C link (7:00) 0:58 10 No 4.6 @ 50

I-22 D 100
 C link (DNO) 1:00 6 No 4.6 @ 50

Table 1: Results of the heptane spray burner Series I. Note that DNO means “Did Not Open”.

Also note, the fires grew at a rate proportional to the square of the time until a certain flow

rate of fuel was achieved at which time the flow rate was held steady. Thus, the “Heat Release

Rate” was the size of the fire at the time when the fuel supply was levelled off.

12
North Wall

Adjustable Height Ceiling

Draft Curtains

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7:24 2:28 3:14

2:12 1:16 1:18 2:06

2:40 1:10 1:08 1:48

2:14 2:24 1:48

Closed

—- —- —-

—- 1:47 1:27 —-

—- 1:33 1:09 2:53

—- —- —-

1:04

—- —- —-

7:48 2:50 1:40 —-

7:02 2:08 1:14 3:26

—- —- —-

0:40 [

—- —- —-

—- 1:10 1:10 —-

—- 1:06 1:04 2:08

—- —- —-

1:30 [

—- —- —-

—- —- 2:16 —-

—- 2:06 1:46 2:08

—- —- —-

1:46

Figure 4: Plan view of Large Scale Fire Test Facility with the results of Heptane Series I

curtained tests where the fires were at location A. The activations times are from Tests I12, I-13, I-14, I-15 and I-16, respectively. Note that an asterisk beside a vent opening time

indicates manual operation.

13
North Wall

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Draft Curtains

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—- 2:06 1:09 1:10 1:53

2:13 1:09 1:05 1:44

3:29 2:17 2:19 —-

Closed

3:48 2:14 1:14 1:10 2:04

2:20 1:10 1:06 1:58

—- 4:08 2:54 3:42

0:40 [

4:46 2:10 1:10 1:12 2:08

2:10 1:10 1:04 1:52

6:02 3:52 —- 3:48

1:30 [

—- 4:24 1:06 1:00 4:39

4:35 1:16 1:12 1:58

—- 6:01 4:43 5:53

9:26

Figure 5: Plan view of heptane spray burner test (Series I) with the results of curtained tests

where the fires were at location B. The activation times are from Tests I-1, I-2, I-3 and I-8,

respectively. Note that an asterisk beside a vent opening time indicates manual operation.

14
North Wall

Adjustable Height Ceiling

Draft Curtains

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3:44 2:08 1:56

3:28 1:06 1:02

—- 1:10 1:00

—- 2:10 1:58

Closed

6:47 2:20 2:28

—- 1:16 1:14

—- 1:20 1:12

—- 2:43 2:18

0:40 [

—- 2:18 2:28

—- 2:18 1:08

—- 1:08 1:02

—- 2:50 1:54

1:30 [

—- 2:26 2:26

3:00 1:12 1:16

3:52 1:12 1:10

—- 2:30 2:00

DNO

Figure 6: Plan view of heptane spray burner configuration with the results of curtained tests

where the fires were at location C. The activation times are from Tests I-4, I-5, I-6 and I-7,

respectively. Note that an asterisk beside a vent opening time indicates manual operation.

15
North Wall

Adjustable Height Ceiling

Draft Curtains

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—- 2:58

11:15 1:28 1:16 2:14

2:50 1:18 1:10 5:26

3:44 2:26 2:38

DNO

5:48 2:40

4:20 1:28 1:20 2:38

3:58 1:28 1:12 2:52

6:52 2:06 3:30

0:40 [

Figure 7: Plan view of heptane spray burner configuration with the results of curtained

tests where the fires were at location D. The activation times are from Tests I-9 and I-10,

respectively. Note that an asterisk beside a vent opening time indicates manual operation.

16
North Wall

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9:43 1:02 0:56

6:47 8:41

10:00

—-

—- 3:16 1:24 —-

—- 1:46 0:54

—- —-

1:20

Figure 8: Plan view of heptane spray burner configuration with the results of uncurtained

tests where the fires were at location A. The activation times are from Tests I-19 and I-20,

respectively. Note that the vent opening time of 10:00 in Test I-19 is approximate because the

electronic timing circuit malfunctioned.

17
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1:20 1:04 "

1:00 0:58

DNO

Figure 9: Plan view of heptane spray burner configuration with the results of uncurtained

test (Test I-17) where the fire was at location B.

18
North Wall

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—- —-

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—- —-

DNO

3:20 3:44

4:06 1:14 1:08

4:06 0:58 0:58

2:48 3:34

7:00

Figure 10: Plan view of heptane spray burner configuration with the results of uncurtained

tests where the fires were at location C. The activation times are from Tests I-18 and I-21,

respectively.

19
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4:03 1:08 1:04 4:25

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DNO

Figure 11: Plan view of heptane spray burner configuration with the results of uncurtained

test (Test I-22) where the fire was at location D.

20
4 Heptane Spray Burner Tests, Series II

A series of additional heptane spray burner tests was performed during a three day period while the

vents and draft curtains were installed for the cartoned plastic commodity tests. The objective was

to generate larger fires than those of the first series of heptane burns, to provide more data in areas

thought to be lacking in the previous tests, and to perform controlled tests in a vented, curtained

area conforming to current model codes [24].

4.1 Description

Twelve tests were performed. The vent, draft curtain and burner positions are shown in Fig. 12. In

each test the heptane spray burner was ramped up to 10 MW in 75 s following a L-squared curve.

After 75 s the heat release rate was held constant. Draft curtains were installed at the ceiling for all

of the tests. The same sheet metal curtains from the first series of heptane tests were installed to

enclose an area 21 m by 23 m, 490 m (70 ft by 76 ft, 5,320 ft ) under the 8.2 m (27 ft) adjustable

ceiling. The same type of vent used during the first heptane spray burner series was used again.

The vents were designed to operate independently (as opposed to gang operation) at a temperature

of 74
 C (165
 F), or manually. The area of the largest quadrant in Fig. 12 was selected to provide

a larger vent to floor ratio (1:42) than called for by the Uniform Fire Code (1:50 for up to 20 ft

of storage height and less then 6,000 ft of curtained area) [24]. Only the vent in the northwest

corner of the curtained area was instrumented, thus the fires were ignited at various locations in

the northwest quarter of the curtained area.

During each 10 min test, the entire test facility was exhausted from the center of the 15 m

(50 ft) high space (7.0 m (23 ft) above the adjustable ceiling) at the minimum operating rate of

11 m /s (24,000 ft /min).

Thermocouples were used to measure gas temperatures and “structural” steel temperatures.

Two velocity probes were used to monitor the air flows, one in the instrumented vent, the other

about 3 m (10 ft) to the south and 3 m to the east of the vent, 10 cm (4 in) below the ceiling. The

velocities were measured with bi-directional probes connected to pressure transducers with a range

of R133 Pa (1.0 mm Hg). A thermocouple array was placed 3 m (10 ft) to the south and 3 m to

the east of the instrumented vent, as well as at a point 6 m (20 ft) south and 6 m east of the vent

(Position D). A third thermocouple array was placed in front of the vent outside of the curtained

area. Details about the thermocouple arrays may be found in Ref. [2]. A thermocouple was placed

near each sprinkler to record temperatures and activation times. Electronic timing circuits were

installed in the vents to determine when each opened.

Positioned near the fusible link installed within the vent cavity were three calibrated brass disks

with different thermal responses, plus a 0.0625 in Inconel sheathed type K thermocouple. The bulk

RTI8 values of the disks were determined from plunge tests at UL [2]. The values were reported

8The term “bulk RTI” indicates that the temperature of the object is assumed to be governed by the equation

-!./

-10S2

46

RTI T

7.8U9V./=<

where ./ is the temperature of the object and .8is the temperature of the gas streaming past the object with speed 6.

This definition of RTIT does not explicitly account for conductive losses from the object to surrounding objects, and

should not be confused with the definition of RTI applied to fusible sprinkler links in this report.

21
to be 32, 164 and 287 (m  s)

+,

(58, 297 and 519 (ft  s)

+,

) for the “fast”, “medium” and “slow” disks,

respectively [2]. The measurements of each ranged between -10% and +10% of the reported value.

4.2 Results

The results of the second series of heptane spray burner tests are summarized in Table 2. The sprinkler activation times can be found on the following pages, where the times of similar and repeated

tests can be found on the same diagram to emphasize differences. The temperature histories of all

100 near-sprinkler thermocouples can be found in Appendix B. Note that the ambient temperature

within the facility varied between 15
 C and 20
 C (59
 F and 68
 F) for all the tests.

Figures 19–21 show the temperatures recorded by the thermocouple and the calibrated brass

disks within the vent located at the northwest corner of the curtained area. The brass disks were

used as a means to assess the thermal environment of the vent cavity by providing a measure of

temperature more indicative of that of the fusible link than the relatively fast-responding thermocouple. None of the disks could simulate the behavior of the fusible link exactly, however, because

the fusing of the link involves the melting of solder.

Heptane Spray Burner Test Series II (10 MW Fires)

Test Burner Vent Sprinklers First Last Avg. Peak Temp.

No. Position Operation Opened Activation Activation 
 C 
 F

II-1 D 74
 C link (DNO) 27 1:15 6:13 129.4 264.9

II-2 D All Open at Start 28 1:05 5:53 128.8 263.8

II-3 A 74
 C link (1:15) 12 1:08 4:00 101.8 215.2

II-4 B 74
 C link (1:48) 16 1:03 5:54 108.8 227.8

II-5 D 74
 C link (DNO) 28 1:10 7:07 130.0 266.0

II-6 D All Open at Start 27 1:10 5:21 127.5 261.5

II-7 A Closed 18 1:09 4:11 117.2 243.0

II-8 B 74
 C link (1:12) 13 1:10 3:34 107.7 225.9

II-9 E 74
 C link (DNO) 23 1:07 3:28 115.8 240.4

II-10 F 74
 C link (3:20) 19+ 1:14 3:01 108.4 227.1

II-11 C 74
 C link (DNO) 23 1:02 3:56 123.4 254.1

II-12 C All Open at Start 23 0:58 4:55 119.0 246.2

Table 2: Results of the heptane spray burner Series II. Note that all fires were ramped up

to 10 MW in 75 s following a L-squared curve. Also, the plus sign appended to a value in

the “Sprinklers Opened” column indicates that the area of sprinkler activation spread to

the edge of the adjustable height ceiling, thus more activations might have occurred had the

ceiling extended further.

22
North Wall

South Wall 











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Figure 12: Plan view of large scale fire test facility with the burner, vent and draft curtain

positions for the second series of heptane burner tests. The sprinklers are indicated by the

solid circles and are spaced 10 ft apart. The branch lines run north to south. The vents are 4

ft by 8 ft.

23
North Wall

South Wall

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4:16

DNO DNO DNO

DNO DNO

1:05 1:07

1:16 1:20

1:45 1:28 1:28

1:46

1:46

1:50

1:52

1:52

1:55

1:59 2:02

2:19

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3:13 3:14

3:24

4:26

4:48

4:59

5:17

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5:21

5:53

DNO Open Open

Open Open

1:13 1:10

1:25 1:25

1:30 1:30

1:47

1:49

1:51

1:54

1:54 1:54

1:54

2:01

2:12 2:03 2:31

2:50 2:47

2:50

3:18

4:01

4:06

5:19

5:26

5:28

5:35

7:07

DNO DNO DNO

DNO DNO

3:30

3:45

3:06

1:51

1:49

1:57

2:52

5:21

1:54

1:24

1:10

1:33

2:21

1:57

1:21

1:10

1:35

2:10

4:22

2:33

1:51

1:53

1:51

3:21

4:49

4:15

4:35

DNO Open Open

Open Open

Figure 13: Plan view of large scale fire test facility with the sprinkler activation times of

(from top to bottom) Tests II-5, II-1, II-2 and II-6 of the second heptane test series. Tests

II-5 and 1 were performed with the vents controlled by fusible links. Tests II-2 and II-6 were

performed with the vents open from the start.

24
North Wall

South Wall

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DNO 1:15 DNO

DNO DNO

2:26

2:16

2:34

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1:32

4:08

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1:27

1:09

1:30

3:47

1:49

1:24

1:19

1:32

4:11

DNO Closed DNO

DNO DNO

Figure 14: Plan view of large scale fire test facility with the sprinkler activation times for

the second heptane test series where the burner was directly under the northwest vent in the

curtained area (Position A). In Test II-3 (upper times), the vent link fused at 1:15, whereas in

Test II-7 (lower times) the vent was held closed.

25
North Wall

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2:46

1:30

1:05

1:48

4:09

DNO 1:48 DNO

DNO DNO

1:17 1:10

1:24 1:27

1:33

1:43

1:51

2:27

2:29

2:31

3:04

3:34 3:28

DNO 1:12 DNO

DNO DNO

Figure 15: Plan view of large scale fire test facility with the sprinkler activation times for the

second heptane test series where the burner was ten feet north of the northwest vent in the

curtained area (Position B). In Test II-4 (upper times), the vent link fused at 1:48, whereas in

Test II-8 (lower times), the vent link fused at 1:12.

26
North Wall

South Wall

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1:20

1:20

1:42

2:20

1:43

1:39

DNO DNO DNO

DNO DNO

Figure 16: Plan view of large scale fire test facility with the sprinkler activation times for

Test II-9 of the second heptane test series where the burner was under the north-south draft

curtain (Position E).

27
North Wall

South Wall

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DNO 3:20 DNO

DNO DNO

Figure 17: Plan view of large scale fire test facility with the sprinkler activation times for

Test II-10 of the second heptane test series where the burner was under the intersection of

the draft curtains (Position F).

28
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Figure 18: Plan view of large scale fire test facility with the sprinkler activation times for the

second heptane series Tests II-11 and II-12 where the burner was at Position C.

29
Figure 19: Northwest vent cavity temperatures for Tests II-1, II-2, II-5 and II-6 of the second

heptane spray burner test series. The curves on the left display temperatures of the thermocouple near the fusible vent link. The curves on the right display temperatures of the brass

disks. The straight horizontal lines indicate the link’s activation temperature (74
 C, 165
 F).

In each of the tests, the burner was located at Position D. In Test II-1 the vent did not open;

in Test II-2 the vent was open from the start, in Test II-5 the vent did not open; in Test II-6

the vent was open from the start.

30
Figure 20: Northwest vent cavity temperatures for Tests II-3, II-4, II-7 and II-8 of the second

heptane spray burner test series. The curves on the left display temperatures of the thermocouple near the fusible vent link. The curves on the right display temperatures of the brass

disks. The straight horizontal lines indicate the link’s activation temperature (74
 C, 165
 F),

and the vertical lines indicate times of vent openings. In Test II-7 the vent was held closed.

31
Figure 21: Northwest vent cavity temperatures for Tests II-9–12 of the second heptane spray

burner test series. The curves on the left display temperatures of the thermocouple near

the fusible vent link. The curves on the right display temperatures of the brass disks. The

straight horizontal lines indicate the link’s activation temperature (74
 C, 165
 F), and the

vertical lines indicate times of vent openings. In Test II-9 the vent did not open; in Test II-10

the vent outside the draft curtains opened; in Test II-11 the vent did not open; in Test II-12

the vent was open from the start.

32
5 Cartoned Plastic Commodity Fire Tests

Following the analysis of the results of the first series of heptane spray burner tests, a series of

high rack storage cartoned plastic commodity fire experiments was performed at the Large Scale

Fire Test facility at UL, the same space that was used for the heptane burner tests described in

Sections 3 and 4. In planning these experiments, the Technical Advisory Committee considered

a number of factors, including the results of the heptane spray burner tests, numerical modeling,

similar experiments performed by Factory Mutual, and current fire protection practices. In order

to develop test specifications that represented a realistic situation, a number of documents were

reviewed, including the 1994 edition of the Uniform Fire Code [24]; the 1991 edition of NFPA

231C, Standard for Rack Storage of Materials [6]; the 1991 edition of NFPA 204M, Guide for

Smoke and Heat Venting [5]; and selected Factory Mutual data sheets [25, 26, 27]. In addition, an

attempt was made to maintain consistency between these tests and the recent work completed by

FMRC on “Extra Large Orifice” sprinklers for rack storage without vents [21, 28].

5.1 Description

Much of the description for the cartoned plastic commodity tests carries over from the heptane

spray burner tests, including the test facility, sprinkler type, vent type and instrumentation. Differences are described below. The Report of Test from UL may be found in Ref. [2]. The UL report

contains a complete description of the instrumentation and the measurements.

5.1.1 Fuel Package

The Factory Mutual Research Corporation (FMRC) Standard Plastic test commodity, a Cartoned

Group A Unexpanded Plastic, served as the fuel for this test series [21]. This commodity has been

used extensively for testing since 1971 [29]. The complete fuel package consists of a combination

of the cartoned plastic commodity and Class II commodity.

The cartoned plastic commodity consists of rigid crystalline polystyrene cups (empty, 0.47 L

(16 fl oz) size) packaged in compartmented, single-wall, corrugated paper cartons. The cups are

arranged open end down in five layers, 25 per layer for a total of 125 per carton. Each carton, or

box, is a cube 0.53 m (21 in) on a side. Eight boxes comprise a pallet load. Two-way, 1.06 m by

1.06 m by 0.13 m (42 in by 42 in by 5 in) slatted deck hardwood pallets support the loads. A pallet

load weighs approximately 80 kg (170 lb), of which about 36% is plastic, 35% is wood and 29%

is corrugated paper [21].

A Class II commodity was used in the target arrays. Typically, the commodity was located

beyond the expected area of the fire spread. This commodity consists of double tri-wall corrugated

paper cartons with five-sided steel stiffeners inserted for stability. The two cartons plus the liner

form a single 1.06 m (42 in) cube having a combined nominal wall thickness of 2.5 cm (1 in). The

single cube was supported on the same type of pallet supporting the cartoned plastic commodity.

5.1.2 Storage Method

A commodity storage height of 6 m (20 ft) with a ceiling height of 8.2 m (27 ft) represents one of

the most severe arrangements allowable under NFPA 231C without requiring in-rack sprinklers.

33
Under the Uniform Fire Code, storage of the test commodity is required to be protected using

sprinklers, vents and draft curtains when the size of the high piled storage area exceeds 230 m

(2,500 ft ).

The test array (Fig. 22) was designed to create a challenging storage situation. Each storage

array consisted of a main (ignition) double-row rack at the center, flanked on two sides by single

row target racks. The rows were separated by 8 ft wide aisles. Each of the two rows of the main

array consisted of four 2.4 m (8 ft) long bays; a 0.15 m (6 in) flue separated the rows. Longitudinal

flues of 0.2 m (7.5 in) were used to separate the pallets within a row. The overall loaded area of the

double-row rack measured approximately 2.3 m (7.5 ft) wide by 10 m (33 ft) long. The racks were

divided vertically into 4 tiers; the overall loaded height was 5.8 m (19 ft). A similar configuration

was used in a series of FMRC burns documented in Ref. [21].

The fire was ignited with 2 standard igniters which consisted of 8 cm (3 in) long by 8 cm

diameter cylinders of rolled cotton material, each soaked in 120 mL (4 oz) of gasoline and enclosed

in a polyethylene bag. The rolls were placed against the carton surfaces in the first tier, just above

the pallet (Fig. 22). The igniters were lit with a flaming propane torch at the start of each test.

5.1.3 Sprinkler Protection

The sprinklers used in all of the tests were Central ELO-231 (Extra Large Orifice) uprights. The

sprinklers, flow rates and activation temperatures were the same as in the first series of heptane

spray burner tests. Even though UL listing and FM approval of this sprinkler with this type of storage arrangement are based on a minimum density requirement of 0.6 gpm/ft , the lower density of

0.5 gpm/ft was used to allow for more challenging, but still controllable, fires and more sprinkler

activations.

5.1.4 Test Layout

Draft curtains were installed at the ceiling for three of the five tests. The same sheet metal curtains

from the heptane tests were installed to enclose an area 21 m by 23 m, 490 m (70 ft by 76 ft,

5,320 ft ) under the adjustable ceiling (Fig. 23). Five vents were installed in the adjustable ceiling,

following the layout shown in Figure 23. This same type of vent was used during the first heptane

spray burner series. The vents were designed to operate independently (as opposed to gang operation) at a temperature of 74
 C (165
 F), or manually. The area of the largest quadrant in Fig. 29

was selected to provide a larger vent to floor ratio (1:42) than called for by the Uniform Fire Code

(1:50 for up to 20 ft of storage and less then 6,000 ft of curtained area) [24]. Only the vent in

the northwest corner of the curtained area was instrumented, thus the fires were ignited at various

locations in the northwest quarter of the curtained area.

The layout of each of the five tests is shown on the following pages, along with the sprinkler

activation times. In designing the layouts, there were several rules that were followed due to both

the geometry of the test facility and a desire to maintain consistency from test to test: (1) the longer

dimension of the 1.2 m by 2.4 m (4 by 8 ft) vents had to run north-south because of the arrangement

of the structural elements in the adjustable ceiling; (2) the vents had to be placed between branch

lines, which also run north-south; (3) the vents had to be placed between two sprinklers, roughly

at the quarter points of the main test quadrant; (4) the draft curtains had to be placed between rows

of sprinklers so as to maintain a 1.5 m (5 ft) spacing between the curtains and sprinklers; and (5)

34
Plan View

]

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]^

Ignition Source

 		 





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_   	`$a
     $b
  



 

Side Elevation View

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Figure 22: Configuration of the cartoned plastic commodity rack storage. Note that the

pallets without perpendicular lines indicate ClassII targets. The ignition location is indicated

by two small rectangles in the center transverse flue of the Plan View.

35
the racks had to be arranged perpendicularly to the branch lines to eliminate “shadowing” effects.

Three ignition locations were chosen: the first farthest from the vents, the second directly under a

vent, and the third near the intersection of the draft curtains.

During each 30 min test, the entire test facility was exhausted from the center of the 15 m

(50 ft) high space (7.5 m (25 ft) above the adjustable ceiling) at a rate of 28 m /s (60,000 ft /min),

the maximum capacity.

5.1.5 Instrumentation

Visual observations of test events were recorded and assessments made of the amount of commodity consumed by the fire (fire damage). Photographic documentation included video tape and color

slides. In addition, infrared video was used to record each test for analysis, and also to aid the fire

fighters who extinguished the residual fire following each 30 min test.

Thermocouples were used to measure gas temperatures and “structural” steel temperatures.

Two velocity probes were used to monitor the air flows, one in the instrumented vent, the other

about 3 m (10 ft) to the south and 3 m to the east of the vent, 10 cm (4 in) below the ceiling.

The velocities were measured with bi-directional probes connected to pressure transducers with a

range of R133 Pa (1.0 mm Hg). A thermocouple array was placed 3 m (10 ft) to the south and

3 m to the east of the instrumented vent, as well as at a point 6 m (20 ft) south and 6 m east of

the vent, above the ignition point for Test P-1. A third thermocouple array was placed in front

of the vent outside of the curtained area. Details about the thermocouple arrays may be found in

Ref. [2]. A thermocouple was placed near each sprinkler to record temperatures and activation

times. Electronic timing circuits were installed in the vents to determine when each opened. In

addition, a thermocouple and three calibrated brass disks were placed in one of the vent cavities to

measure the air temperature and to simulate the thermal response of the fusible link. These same

disks were installed during the second series of heptane spray burner tests.

Sprinkler water flow and pressure were continuously monitored. Oxygen, carbon monoxide

and carbon dioxide concentrations were recorded near the ignition point. Three optical sensors

consisting of a white light source separated from a photocell by 1 m were positioned 1 m above

the adjustable ceiling, 1.8 m (6 ft) below the ceiling and 1.5 m (5 ft) above the floor at the location

indicated in the plan views shown below.

36
5.2 Results

Five cartoned plastic commodity tests were conducted over the course of a month (October 1997).

The results of the five tests are summarized in Table 3. In four of the tests, the vents were equipped

with a fusible link, rated at 74
 C (165
 F). In one test (Test P-5) the vents were manually operated.

Further details of each test can be found on the following pages.

Cartoned Plastic Commodity Test Series

Test Number P-1 P-2 P-3 P-4 P-5

Test Date Sept. 30 Oct. 2 Oct. 16 Oct. 21 Oct. 28

Indoor Temperature (
 C/
 F) 20/68 18/65 19/67 17/63 18/64

Outdoor Temperature (
 C/
 F) 14/57 11/52 10/50 10/50 8/46

Relative Humidity (%) 51 51 48 30 43

Commodity Moisture Content (%) 9 9 8 6 8

Draft Curtains No No Yes Yes Yes

Fire Pos. rel. to vent 20E,20S Under 10N 20E,20S 20E,20S

No. Vent Openings 0 1 1 0 4 (manual)

Vent Opening Time DNO 6:04 4:11 DNO 1:14

First Sprinkler Activation 1:16 1:40 1:07 1:33 1:14

Last Sprinkler Activation 13:41 5:37 16:06 2:40 5:04

Activations in the 1st 6 min 4 23+ 8 5 7

Total Sprinklers Activated 20 23+ 19+ 5 7

Estimated Boxes Consumed 117 127 184 103 81

Table 3: Results of the cartoned plastic commodity test series. Note that position of fire is

in relation to the vent in the northwest corner of the curtained area. The vent that operated

in Test P-2 was not the vent directly above the ignition source. The plus sign appended to a

value in the “Activations in the 1st 6 min” or the “Total Sprinklers Activated” rows indicates

that the area of sprinkler activation spread to the edge of the adjustable height ceiling, thus

more activations might have occurred had the ceiling extended further.

37
5.2.1 Test P-1

The fire was ignited in a transverse flue, at the point where two boxes on the same pallet abut. The

two ignitors were situated about 12 cm (5 in) off the floor. The flames reached the top of the third

tier about 54 s after ignition; the fourth tier about 60 s. At this time, the flames could just be seen

emerging from the array on the south face and entering the longitudinal flue separating the rows

within the array. The first sprinkler activated 76 s after ignition, knocking the flames below the top

of the fourth tier. The flames re-emerged from the longitudinal flue at about 1:53, followed by the

second sprinkler activation at 2:14 on the opposite side of the ignition flue. The 58 s delay between

the first and second sprinkler activation allowed the fire to spread more rapidly towards the east

end of the array, ultimately opening up more sprinklers on the eastern side of the north-south line

of symmetry. Although 20 sprinklers opened by about 15 min, only 4 had opened after 8 min.

Examining the thermocouple data near each unactivated sprinkler, it appears that the near-ceiling

temperatures outside of the first ring of sprinklers steadily grew following the first few activations,

but eventually decreased following the 11 activations that occurred between 8 and 13 min.

By about 10 min after ignition the smoke layer throughout the test facility was at the height of

the third tier. Up to this point in time, much more smoke was observed in the north aisle, reducing

visibility to only a few meters. The south aisle remained relatively clear of smoke during the first

10 min.

The fire damage was limited to the bays surrounding the ignition point, both on the south and

north side of the main array (Fig. 25). Most of the fire damage was found in the bay east of the

ignition flue on tiers 1 through 3 due to the asymmetrical sprinkler activation pattern. The fourth

tier was protected from fire damage by the sprinkler spray.

38
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DNO DNO DNO

DNO DNO

Figure 23: Plan view of Large Scale Fire Test Facility with the sprinkler activation times for

Plastic Test P-1. The circled “S” denotes the location of the optical density meters or “smoke

eyes”.

39
Figure 24: Temperature profiles from all 100 near-sprinkler thermocouples from Plastic Test

P-1. The arrangement of the plots mimics that of Fig. 23.

40
View of Main Array from the South

View of Main Array from the North

Figure 25: Fire damage from Plastic Test P-1.

41
5.2.2 Test P-2

Test P-2 was intended to present an extreme situation as far as venting is concerned. The ignition

point was put directly under a vent. The objective was to see how vent activation soon after ignition

could affect sprinkler response. In the experiment, flames reached the top of the central array at

about 65 s and the vent cavity at about 70 s. The first sprinkler activated at 100 s, followed 8 s later

by the sprinkler on the opposite side of the vent. The vent above the ignition point did not open

at any time during the 30 min test. However, the vent 6 m (20 ft) to the west of the ignition point

did open at 6:04. Unlike Test P-1, in Test P-2 all 23 sprinkler activations occurred within 6 min of

ignition.

By about 10 min after ignition the smoke layer throughout the test facility was at the height of

the third tier. Up to this point in time, there was no observable difference in the amount of smoke

in the north and south aisles. The fire damage in Test P-2 was similar to that of Test P-1, with more

damage observed in the fourth tier (Fig. 28).

42
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DNO

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DNO DNO

Figure 26: Plan view of Large Scale Fire Test Facility with the sprinkler activation times for

Plastic Test P-2.

43
Figure 27: Temperature profiles from all 100 near-sprinkler thermocouples from Plastic Test

P-2. The arrangement of the plots mimics that of Fig. 26.

44
View of Main Array from the South

View of Main Array from the North

Figure 28: Fire damage from Plastic Test P-2.

45
5.2.3 Test P-3

Test P-3 was designed to be similar to a test performed during the high rack storage cartoned

plastics series at Factory Mutual where the fire was ignited near the intersection of two draft

curtains [21]. However, the test performed at Factory Mutual involved slatted wood shelving, a

slightly different rack configuration, and different sprinkler spacing and flow rate.

The first sprinkler activation occurred after 67 s at the position closest to the curtain intersection, followed closely by the second at 72 s. The draft curtains channeled the smoke and hot

gases towards the east and south, opening more sprinklers in these rows than would be expected if

the curtains were not there. The vent did automatically activate at 4:11, but by that time the two

sprinklers on each side of it had already activated.

The draft curtains did delay by 3 to 4 min the activation of the sprinklers to the north of the first

two that activated, and blocked the water from wetting the north face of the center array. The fire

was able to spread both to the north and to the west underneath the draft curtains, but activations

within the adjacent quadrants were able to control the spread. The sprinkler activations went to the

north edge of the adjustable height ceiling.

By about 10 min after ignition the smoke layer throughout the test facility was at the height of

the third tier. Up to this point in time the north aisle remained relatively clear of smoke, the south

aisle filled with smoke within the first 5 min. The fire damage from this fire was more extensive

than that of Tests P-1 and P-2, going beyond the central two bays in tiers 2, 3 and 4 (Fig. 31).

46
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Figure 29: Plan view of Large Scale Fire Test Facility with the sprinkler activation times for

Plastic Test P-3.

47
Figure 30: Temperature profiles from all 100 near-sprinkler thermocouples from Plastic Test

P-3. The arrangement of the plots mimics that of Fig. 29.

48
View of Main Array from the South

View of Main Array from the North

Figure 31: Fire damage from Plastic Test P-3.

49
5.2.4 Test P-4

Test P-4 was intended to be a replicate of Test P-1 with the only exception being the presence of the

draft curtains. As in all the tests prior to this one, the flames emerged from the top tier after about

60 s. The first sprinkler opened after 93 s followed 1 s later by the second. This activation time was

about 10 to 15 s longer than in previous tests under a flat ceiling, but it appeared to those observing

the test that the quick succession of activations contained the fire to a volume roughly two pallet

loads wide, three pallets high, and one pallet deep. The temperatures at all of the unactivated

sprinklers decreased relatively quickly following the initial activations. There was no build-up of

the fire following the first activation that was observed in Test P-1.

The smoke layer descended to the height of the third tier after about 15 min. Up to this point

in time the south aisle was more heavily smoke logged than the north. The fire damage from this

test was less extensive than that from the first three tests (Fig. 34).

50
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Figure 32: Plan view of Large Scale Fire Test Facility with the sprinkler activation times for

Plastic Test P-4.

51
Figure 33: Temperature profiles from all 100 near-sprinkler thermocouples from Plastic Test

P-4. The arrangement of the plots mimics that of Fig. 32.

52
View of Main Array from the South

View of Main Array from the North

Figure 34: Fire damage from Plastic Test P-4.

53
5.2.5 Test P-5

Given the difference in sprinkler activity between Test P-1 (20 activations) and Test P-4 (5 activations), it was decided to repeat Test P-4, only now the vents were to be manually opened following

the first sprinkler activation. It had become clear by this time in the project that the vents were

unlikely to open when the fire was ignited more than about 4.6 m (15 ft) away.

The results of Test P-5 were similar to those of Test P-4. Again, the first two sprinklers opened

within 1 s of each other, but about 20 s sooner than those in Test P-4. No change in test conditions

could explain why the sprinklers activated sooner, thus it was attributed to the variability in the

initial rate of heat release. Seven sprinklers opened in all, compared with 5 in Test P-4. The fire

damage was slightly less extensive in this test than in Test P-4, possibly due to the earlier sprinkler

activation.

The smoke layer descended to the height of the third tier after about 15 min. Up to this point

in time the north aisle was more heavily smoke logged than the south. The fire damage from this

test was similar to that of Test P-4.

54
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1:15 1:14

2:26

2:27

3:21

3:33

5:04 DNO 1:14 1:14

1:14 1:14

Figure 35: Plan view of Large Scale Fire Test Facility with the sprinkler activation times for

Plastic Test P-5. The vents in the curtained area were opened manually following the first

sprinkler activation.

55
Figure 36: Temperature profiles from all 100 near-sprinkler thermocouples from Plastic Test

P-5. The arrangement of the plots mimics that of Fig. 35.

56
View of Main Array from the South

View of Main Array from the North

Figure 37: Fire damage from Plastic Test P-5.

57
5.2.6 Obscuration

Three optical sensors consisting of a white light source separated from a photocell by 1 m were

positioned 1 m above the adjustable ceiling, 1.8 m (6 ft) below the ceiling and 1.5 m (5 ft) above

the floor at the location indicated in the plan views of each test. The smoke eyes were not moved

from test to test, even though the racks were. Figure 38 presents the data from all five tests.

Because during every test, the exhaust fan above the plenum space was drawing smoke from the

test space at a constant rate of 28 m /s (60,000 ft /min), the optical density data (averaged over

time to eliminate local smoke logging effects) can be used as an indirect indicator of the relative

size of the fire. Tests P-1, P-2 and P-3 were bigger fires than Tests P-4 and P-5 as indicated by the

increased obscuration of the test facility and the amount of commodity consumed. Unfortunately,

what the optical density data does not reveal is the effect of opening the smoke and heat vents

because there is no way to separately account for the smoke flowing through the vents and the

smoke flowing around the adjustable ceiling and up through the 3 m (10 ft) wide gap between the

ceiling and the wall of the facility. In effect, all the tests were vented at the perimeter.

58
Figure 38: Local optical density measurements for the Plastic tests, averaged over a minute.

59
6 Discussion of Experimental Results

As in studies preceding this one, a series of large scale tests was used to gather information about

the interaction of sprinklers, vents and draft curtains. In this study, engineering analysis using

fire modeling techniques will provide additional insight into the behavior of these systems. To

emphasize the contrast provided, it is instructive to pause here and discuss what can be determined

from the large scale tests alone. Later, additional information from fire modeling will help bolster

the confidence in these conclusions and provide additional insight.

Thus far the results of 39 full-scale fire experiments have been presented. In this section, these

results will be discussed with an emphasis on how roof vents and draft curtains affect the time,

number and location of sprinkler activations; and how sprinklers and draft curtains affect the time,

number and discharge rates of roof vents. To facilitate the discussion, a summary of the 39 large

scale experiments is presented in Table 4.

6.1 Effect of Vents and Draft Curtains on Sprinkler Activation Times

For all three series of tests, when the fire was not ignited directly underneath a vent, the activation

times of the nearest sprinklers to the fire were not affected by the opening of vents either prior to

or after the first sprinkler activation. When the fire was ignited 3 m (10 ft) from the vent center,

the only discernible affect of the vent opening on sprinkler activation was for those sprinklers

immediately downstream of the vent. For example, consider the sprinkler activation times for

Tests I-4, I-5, I-6 and I-7 (Fig. 6). The two sprinklers to the west of the vent did not activate in

Tests I-5 and I-6 when the vent was manually operated at 40 and 90 s. However, in Test I-7 when

the vent did not open, both sprinklers activated, and in Test I-4 when the vent was held closed, one

of the two sprinklers activated. Consider the peak gas temperatures near the two sprinklers in Tests

I-4 and I-7 compared with those in Tests I-5 and I-6 (Figs. 88-91). Compared to the unopened

tests, the temperatures were 25
 C (45
 F) lower when the vent was opened at 40 s, and 5
 C (9
 F)

lower when the vent was opened at 90 s. When examining the temperatures and activation times

for Tests I-1, I-2, I-3 and I-8 where the fire was in the B position (3 m (10 ft) to the north of the

vent), the effect of opening the vent is not as clear because the shorter dimension of the vent is

facing the fire.

In tests where the fire was ignited directly beneath a vent, vent openings prior to the activation

of the nearest sprinklers had an effect on the sprinkler activation times. The earlier the vent opening, the more noticeable the effect. For example, in Tests I-12, I-13, I-14 and I-15 where the fire

was positioned directly under the vent and the draft curtains were installed, the average sprinkler

activation time of the nearest four sprinklers was 1:13 when the vent was held closed (Test I-12),

1:29 when the vent opened automatically at 1:04 (Test I-13), 1:58 when the vent was opened manually at 0:40 (Test I-14), and 1:08 when the vent was opened manually at 1:30 (Test I-15). This

data suggests that the earlier the vent activation, the longer the delay in activation of the first ring

of sprinklers.

Test I-16 was performed with a different fire growth curve, and cannot be directly compared

with any other test. In that test, the first sprinkler activated at the same time that the vent opened

(1:46), followed by the next two sprinklers at 2:06 and 2:08. One of the four sprinklers nearest the

fire did not activate at all. The temperature near this sprinkler was 140
 C (284
 F) at the time of the

vent opening, but it decreased to about 80
 C (176
 F) over the next few minutes. A similar effect

60
Summary of Sprinkler, Smoke&Heat Vent, Draft Curtain Project Experiments

Test Fire Location Fire Size Draft Vent Sprinklers Avg. Peak Temp. # Boxes

No. Pos. Rel. to Vent MW @ s Curtains Operation First Last Tot. f C f F Consumed

I-12 A 0 4.4 @ 50 Yes Closed 1:08 7:24 14 111.0 231.8I-13 A 0 6.0 @ 60 Yes 74 f C link (1:04) 1:09 2:53 5 96.2 205.2I-14 A 0 5.8 @ 60 Yes Manual (0:40) 1:14 7:48 7 95.2 203.4I-15 A 0 5.8 @ 60 Yes Manual (1:30) 1:04 2:08 5 113.8 236.9I-16 A 0 5.0 @ 110 Yes 74 f C link (1:46) 1:46 2:16 4 87.7 189.9I-19 A 0 4.6 @ 50 No 100 f C link (10:00) 0:56 9:43 10 84.4 183.8I-20 A 0 4.2 @ 50 No 74 f C link (1:20) 0:54 3:16 4 75.8 168.5

II-3 A 0 10 @ 75 Yes 74 f C link (1:15) 1:08 4:00 12 101.8 215.2II-7 A 0 10 @ 75 Yes Closed 1:09 4:11 18 117.2 243.0

P-2 A 0 Plastic No 74 f C link (DNO) 1:40 5:37 23 85.8 185.9 127

I-1 B 10 g N 4.4 @ 50 Yes Closed 1:05 3:29 11 110.9 231.6I-8 B 10 g N 4.4 @ 50 Yes 74 f C link (9:26) 1:00 6:01 11 99.7 211.4I-2 B 10 g N 4.4 @ 50 Yes Manual (0:40) 1:06 4:08 12 105.9 222.7I-3 B 10 g N 4.4 @ 50 Yes Manual (1:30) 1:04 6:02 12 109.2 228.5

I-17 B 10 g N 4.6 @ 50 No 100 f C link (DNO) 0:58 1:20 4 81.2 178.1

II-4 B 10 g N 10 @ 75 Yes 74 f C link (1:48) 1:03 5:54 16 108.8 227.8II-8 B 10 g N 10 @ 75 Yes 74 f C link (1:12) 1:10 3:34 13 107.7 225.9

P-3 B 10 g N Plastic Yes 74 f C link (4:11) 1:07 16:06 19 90.5 194.9 184

I-4 C 10 g E 4.4 @ 50 Yes Closed 1:00 3:44 10 109.6 229.3I-7 C 10 g E 4.4 @ 50 Yes 74 f C link (DNO) 1:10 3:52 10 107.3 225.1I-5 C 10 g E 4.4 @ 50 Yes Manual (0:40) 1:12 6:47 9 98.9 210.0I-6 C 10 g E 4.4 @ 50 Yes Manual (1:30) 1:02 2:50 8 105.2 221.4

I-18 C 10 g E 3.7 @ 50 No 100 f C link (DNO) 0:58 1:18 4 70.9 159.7I-21 C 10 g E 4.6 @ 50 No 74 f C link (7:00) 0:58 4:06 10 85.6 186.1

II-11 C 10 g E, 10 g S 10 @ 75 Yes 74 f C link (DNO) 1:02 3:56 23 123.4 254.1II-12 C 10 g E, 10 g S 10 @ 75 Yes Manual (0:00) 0:58 4:55 23 119.0 246.2

I-9 D 20 g W, 20 g S 4.4 @ 50 Yes 74 f C link (DNO) 1:10 11:15 12 108.7 227.7

I-10 D 20 g W, 20 g S 4.4 @ 50 Yes Manual (0:40) 1:12 6:52 13 107.5 225.5I-11 D 20 g W, 20 g S 4.4 @ 50 Yes 74 f C link (4:48) N/A N/A N/A 184.9 364.9I-22 D 20 g W, 20 g S 4.6 @ 50 No 100 f C link (DNO) 1:00 4:25 6 88.7 191.7

II-1 D 20 g E, 20 g S 10 @ 75 Yes 74 f C link (DNO) 1:15 6:13 27 129.4 264.9II-5 D 20 g E, 20 g S 10 @ 75 Yes 74 f C link (DNO) 1:10 7:07 28 130.0 266.0II-2 D 20 g E, 20 g S 10 @ 75 Yes Manual (0:00) 1:05 5:53 28 128.8 263.8II-6 D 20 g E, 20 g S 10 @ 75 Yes Manual (0:00) 1:10 5:21 27 127.5 261.5

P-1 D 20 g E, 20 g S Plastic No 74 f C link (DNO) 1:16 13:41 20 93.8 200.8 117P-4 D 20 g E, 20 g S Plastic Yes 74 f C link (DNO) 1:33 2:40 5 82.6 180.7 103P-5 D 20 g E, 20 g S Plastic Yes Manual (1:14) 1:14 5:04 7 83.3 181.9 81

II-9 E 10 g W, 15 g S 10 @ 75 Yes 74 f C link (DNO) 1:07 3:28 23 115.8 240.4

II-10 F 10 g W, 15 g N 10 @ 75 Yes 74 f C link (3:20) 1:14 3:01 19 108.4 227.1

Table 4: Summary of the 39 Large Scale Experiments. Note that DNO means “Did Not Open”. Test No. I- hrefers to the first

series of heptane spray burner tests, II- hthe second series, and P- hthe Plastic tests. The “Avg. Peak Temp.” refers to the average

of the maximum temperatures recorded at all the near-sprinkler thermocouples in the test area.

61
was seen in Test I-20 (Fig. 8). Following a vent opening at 1:20, one of the four sprinklers nearest

the fire did not open until 3:16. The first sprinkler activation was at 0:54.

During the second series of heptane spray burner tests, two tests were performed with the

burner directly under a vent (Fig. 14). In Test II-7, where the vent was held closed, the average

activation times of the nearest two sprinklers was 1:14 and the nearest six 1:24. In Test II-3, where

the vent opened automatically at 1:15, the average of the nearest two sprinklers was 1:17 and the

nearest six 1:32.

Draft curtains also had an effect on sprinkler activation times. Given two sprinklers equidistant

from a fire, the one nearer to a curtain showed a tendency to activate sooner. For example, in tests

where the fire was placed at Position B in the first series of heptane tests, activation at sprinkler 52

(1 m from east curtain, 6 m from north curtain) occurred about 20 s on average before activation

at sprinkler 64 (10 m from east curtain, 6 m from north curtain). The Position B experiments with

curtains (Tests I-1, I-2, I-3, I-8) also showed activation at sprinkler 90 in Tests I-2 and I-3. This

sprinkler was over 7.6 m (25 ft) from the burner, the farthest distance of any activation in the test

series. Based on an examination of thermocouple data, activation at this sprinkler was believed to

be due to the channeling of the smoke and hot gases by the north draft curtain.

6.2 Effect of Vents and Draft Curtains on Number of Sprinkler Activations

In general, draft curtains increased the number of sprinkler activations. Inspection of Table 4 indicates that in Tests I-1 and I-8 there were 11 activations when the draft curtains were installed and

the vent was closed, and in Test I-17 there were 4 activations when the curtains were not installed

and the vent was closed. Tests I-4 and I-7 both had 10 activations with the curtains installed and

the vent closed, Tests I-18 and I-21 had 4 and 10 activations with the curtains removed. Tests

I-9 and I-10 had 12 and 13 activations with curtains installed, Test I-22 had 6 activations with the

curtains removed. This data indicates that in tests performed with draft curtains where the fire was

not directly beneath a vent, there were up to twice as many sprinkler activations compared to tests

performed without draft curtains.

The reason for the increased number of activations is that draft curtains lead to an increase

of the near-ceiling gas temperatures. Consider, for example, the peak gas temperatures near the

second ring sprinklers in Test I-1 (Fig. 85) compared to those of Test I-17 (Fig. 101). The temperatures were between 20
 C and 30
 C (36
 F and 54
 F) lower in Test I-17. Similar differences can be

seen when comparing temperatures in Tests I-1 through I-16 with those in Tests I-17 through I-22.

The difference between temperatures in the curtained and uncurtained tests can be explained by

considering a fire plume impinging on a well-developed, 1.8 m (6 ft) deep smoke layer as opposed

to a thinner layer. In the latter case, the plume can entrain more cool air before it reachesthe ceiling

layer, and therefore the smoke is cooler by the time it reaches the ceiling. Plus, the deeper smoke

layer formed by the draft curtains insulates the sprinklers from cooler air below the layer, leading

to more activations.

What effect did the vents have on the number of activations? When the fire was ignited directly

under a vent (Position A), the number of activations was reduced. Consider Test I-12 versus Tests

I-13, I-14, I-15 and I-16. The number of activations was roughly halved due to the opening of the

vent directly above the fire. Tests II-3 and II-7 show the number of activations reduced from 18

to 12. However, when the fire was not ignited under a vent, there was either a small decrease or

no decrease at all in the number of sprinkler activations. Tests I-1 and I-8 compared with Tests

62
I-2 and I-3 showed no reduction in the number of activations when the fire was ignited 3 m (10 ft)

north of the vent. Tests I-4 and I-7 compared to Tests I-5 and I-6 showed a reduction of 1 and 2

sprinklers from 10. Tests II-11 and II-12 showed no reduction at all. Tests I-9 and I-10, as well

as Tests II-1, II-5, II-2 and II-6 showed no reduction either. Thus, unless the ignition took place

under or very near a vent, there was no evidence in this data set that venting reduced the number

of sprinkler activations.

To see why vents had little effect on the number of sprinkler activations, consider the average

peak temperatures in the curtained area in Tests II-1, II-2, II-5, II-6, II-11 and II-12. In Tests

II-1 and II-5 where the fire was located at Position D and no vents operated, the average peak

temperatures were 129.4
 C and 130.0
 C, respectively. In Tests II-2 and II-6 where the fires were

at Position D but all the vents were opened at the start of the tests, the average peak temperatures

were 128.8
 C and 127.5
 C, respectively. Similarly, in Test II-11 where the fire was at Position C

and the vent did not operate, the average peak temperature was 123.4
 C, whereas in Test II-12,

where all the vents were opened at the start, the temperature was 119.0
 C.

6.3 Effect of Vents and Draft Curtains on Sprinkler Discharge Pattern

In the cartoned plastic commodity Test P-3, the draft curtain to the north of the ignition point

delayed the operation of sprinklers further north and blocked the spray of sprinklers on either side

of it. In this test, the fuel array extended beneath the north and west curtains. The fire spread

to the north side of the main array because the commodity there was unwetted due to a delay in

sprinkler activation on the north side of the curtain and blockage of the sprinkler spray from the

south side. The results of Test P-3 reinforced evidence provided by two similar9

tests performed by

Factory Mutual [21]. In the FMRC tests, the fires spread underneath the curtains, resulting in the

development of a more severe fire, a greater number of sprinkler operations, an atypical sprinkler

opening pattern, distorted sprinkler discharge patterns which affected prewetting of commodity,

and more smoke production. Although the fire damage and number of sprinkler activations in Test

P-3 were not as great as that seen in the tests performed at FMRC, the fire damage was substantially

higher in this test than in any other test performed in the series, even though the first two sprinkler

activations were relatively early (67 and 72 s). This early activation was most likely due to the

close proximity of the fire to the intersection of the draft curtains. However, the early jump on the

fire did not lead to a rapid decrease in temperatures or sprinkler activations as was the case in Tests

P-4 and P-5, the other two tests performed with draft curtains installed. Instead, the fire spread

to the unprotected north face of the central array; and even though it was eventually controlled

by sprinklers on the north side of the east-west curtain, it ultimately consumed approximately 184

boxes, nearly twice as much as Tests P-4 and P-5. A vent did automatically activate at 4:11, but

by that time the two sprinklers on each side of it had already activated. Based on an examination

of the the sprinkler activation pattern and the thermocouple data, the opening of the vent had no

influence on the test results.

9The tests performed at FMRC involved slatted wood shelving, a slightly different rack configuration, and different

sprinkler spacing and flow rate.

63
6.4 i

Effect of Sprinklers on the Number and Time of Vent Activations

Based on the test data collected in this study, it is difficult to assess how, in general, sprinklers

affect the activation of vents because (1) there is little information about how the vents would have

operated in an unsprinklered facility because only one test was performed without sprinklers, and

(2) only one vent design was used in the test program. However, it appears from the data below

that the sprinkler spray influenced the thermal response characteristics of this particular vent, and

it is believed that sprinklers could have a similar influence on similar vent designs.

In the one unsprinklered test of the study (Test I-11), the vent opened at 4:48. The heptane

spray burner was 8.6 m (28 ft) from the vent center. Six other tests were performed with the fire

at this distance from the vent when the vent was equipped with a fusible link, and in none of these

tests did the vent open. In the unsprinklered Test I-11, the temperature near the vent was about

170
 C (338
 F), whereas in Test I-10, with the fire at the same location, the temperature near the

vent was about 90
 C (194
 F) after the sprinklers had activated around the fire (Figs. 94 and 95).

Examination of the near-ceiling temperatures from all the tests indicates that sprinklers of this type

have a significant cooling effect, and this will certainly have an effect on thermally-responsive,

independently-controlled vents.

To better understand the thermal environment in the vicinity of the vent’s fusible link, a thermocouple and three calibrated brass disks were placed near the link of the vent located at the

northwest corner of the curtained area during the cartoned plastic and the second series of heptane

spray burner tests (see Section 4.1 for a description of the instrumentation). Figures 19–21 show

the temperatures recorded by the thermocouple and the brass disks within the vent cavity during

the heptane tests. In Tests II-3 and II-4, when the vent opened automatically, the temperature of the

“medium” and “slow” disks rose above the rated temperature of the link (74
 C, 165
 F) at about the

same time that the vent opened. In Test II-8, the vent opened about 10 s before the “medium” disk

temperature reached 74
 C, and 30 s before the “slow” disk temperature reached 74
 C. In Tests II-9

and II-11, where the vent did not open, the temperatures recorded by the “medium” and “slow”

disks were comparable to those recorded by the disks in Tests II-3, II-4 and II-8.

In Plastic Test P-2, the fire was ignited directly under a vent. In the experiment, flames reached

the top of the central array at about 65 s and the vent cavity at about 70 s. The first sprinkler

activated at 100 s. The vent did not open at any time during the 30 min test even though another

vent 6 m (20 ft) to the west of the unopened vent opened at 6:04. The temperature history of the

brass disks within the cavity of the unopened vent is given by Fig. 39. After the test, the fusible

link was examined, and it was observed that the solder holding the two strips of metal together

had begun to melt. This observation had been made when examining the links after several of the

heptane spray burner tests, as well.

This data, along with the plunge tunnel measurements reported in Section 3.1.4, suggests that

the fusible link reached its activation temperature before or at about the same time as the the first

sprinkler activated, but the link did not fuse. It is not clear whether the link did not fuse because

it was cooled directly by water drawn upwards into the vent cavity, or whether the sprinkler spray

simply cooled the rising smoke plume enough to prevent the link from fusing. In any event, this

phenomenon deserves further study.

64
Figure 39: Temperatures inside the instrumented vent cavity during Plastic Test P-2. The

curve on the left displays the temperature of the thermocouple near the fusible link. The

curve on the right displays the temperature of the brass disks. The horizontal line on the left

plot indicates the rated temperature of the fusible link.

6.5 Effect of Sprinklers on the Discharge Rate of Vents

The cooling of the near-ceiling gases due to the operation of sprinklers will affect the rate of

discharge through a vent. To measure the flow of gases through a vent, a velocity probe and

thermocouples were positioned in the vent nearest the fire location in the second series of heptane

burner tests and the cartoned plastic commodity tests. Unfortunately, the velocity data was deemed

unreliable, thus there was no means to directly measure the discharge rate. The numerical model

to be described in the next section will be used to examine the effect of sprinklers on the discharge

rate of vents.

An indirect effect of sprinklers on vent performance is that sprinkler sprays entrain smoke and

hot gases, cool them, and transport them towardsthe floor. No measurements were made during the

tests to quantify this phenomenon, but visual observations were made to determine what areas of

the test space filled with smoke during the first 5 or 10 minutes of the cartoned plastic commodity

tests. In Test P-1, earlier and more frequentsprinkler activation occurred to the north of the ignition

point, leading to heavier observed smoke logging in the north aisle. In Test P-2 it was less obvious

which aisle was more heavily smoke logged. In Test P-3, the south aisle was more smoke logged

because sprinklers to the north of the ignition point were delayed by the draft curtains. The curtains

also blocked the smoke from the north aisle, at least initially. In Test P-4, the south aisle was more

heavily smoke logged; in Test P-5, the north aisle. The sprinkler activation pattern in Tests P-4 and

P-5 was consistent with these observations.

65
7 Numerical Modeling

The numerical model used to simulate the physical experiments is based on techniques common

to the computational fluid dynamics (CFD) community. Often these types of models are referred

to as “field models” by the fire research community, to distinguish them from “zone models”. The

basic idea behind most CFD models is to divide the space of interest into small control volumes

or computational cells, and in each cell compute the density, velocity, temperature, pressure and

species concentration based on conservation laws of mass, momentum and energy. The accuracy

of the results often depends on the number of cells used to discretize the volume of interest. The

technique being applied in this project is referred to as large eddy simulation (LES). The idea

behind this approach is to divide the test space into as many cells as possible (in this case, hundreds

of thousands to over a million) to resolve as much of the convective motion of the gases (air,

smoke) as possible. In this way, much of the mixing of the hot gases from the fire with the cool

surrounding air can be captured directly, reducing the dependence on empirical entrainment or

turbulence parameters that are often subject to much debate and uncertainty.

7.1 Model Description

The heart of the numerical model is an algorithm that solves the set of partial differential equations

describing the transport of smoke and hot gases from the fire and its subsequent mixing with the

surrounding air. This is often referred to as the hydrodynamic model. The driving forces, like

the fire and the sprinkler spray, are represented by source terms in the governing equations. The

physical boundaries and their properties provide the boundary conditions for the equations. In the

sections to follow, the major equations and the assumptions behind them are put forward.

7.1.1 Hydrodynamics

Consider a thermally expandable ideal gas driven by a prescribed heat source. The equations of

motion governing the fluid flow are written in a form suitable for low Mach number applications

[30]. Sometimes, this form of the equations is referred to as “weakly compressible”. The most

important feature of these equations is that in the energy conservation equation and the equation

of state the spatially and temporally varying pressure is replaced by an average pressure j

H

that

depends only on time. This is done to filter out acoustic waves. The efficiency of the numerical

solution of the equations is dramatically increased by this approximation.

In the equations to follow, all symbols have their usual fluid dynamical meaning: k is the

density, l the velocity vector, m

F5n l the vorticity, j the pressure, o the gravity vector, prq the

constant-pressure specific heat, s the temperature, t the thermal conductivity, L the time, D u  

the

prescribed volumetric heat release rate, v the gas constant equal to the difference of the specific

heats v

F

prqxwyp{z , | the external force term (in this case the sprinkler spray), and } the standard

stress tensor for compressible fluids.

Conservation of Mass

~

k~

LV n

kl F 

(1)

66
Conservation of Momentum

k ~l

~

Ll3 n l  n

jwkAo F |

 n

} (2)

Conservation of Energy

kpq ~s

~

L 

l ns'wGj

H



L

F D u    n

tns (3)

Equation of State

j

H

 L

F kvs (4)

The divergence of the flow n

!l is a very important quantity in the analysis to follow, and it is

readily found by combining Eqs. (1) and (3), using the equation of state (4)

j

Hn l 





j

H



L

F  w







D u  n

tns



(5)

where  F

pqpz . Integrating Eq. (5) over the entire domain  yields a consistency condition for

the background pressure j

H

 L



j

H{ l ` 

j

H



L

F  w





J D u 

 { tn

s ` O (6)

where 

is the volume of the enclosure. The background pressure can be expressed in terms of a

background temperature s

H

 L



and density k

H

 L



j

HF vkHsH

(7)

These spatially averaged quantities play the same role that ambient conditions do in the Boussinesq

approximation. Perturbations to each are represented by the relations

s

F sH

 L



   s

 ¡ k F

k

H

 L



   k



(8)

The perturbation values are thus simply related

  ¢s



   k

£F 

(9)

Defining s

H

and k

H

through the adiabatic process

k

H

k¤

F

 j

H

j¥¤ ¦ ̈§ ̈© (10)

allows the energy equation to be expressed in terms of the perturbation temperature s and the

divergence ~

s

~

L 

l n

s

F

  ?s

«a¬n l 





j

H

j

H



L®­ (11)

67
The background pressure is found from Eq. (6).

Walls and other solid surfaces that form the boundary of the flow domain are heated by the

surrounding gases. Heat is conducted into the solid according to the equation

k ̄%p! ̄

~s° ̄ ~

L

F

~

~2± tA ̄ ~s3 ̄ ~2± (12)

where k ̄ is the density of the solid, p! ̄ the specific heat, s° ̄ the temperature, tA ̄ the conductivity,

and ±

the normal direction coordinate. Heat is transferred to solids by convection and thermal

radiation. More detail on these processes is included below.

The pressure is composed of three components, the background j

H

 L



, the hydrostatic, and a

perturbation to the hydrostatic j

j

μ ́·¶ L

F jH

 L

 wkH

 L

1 ̧%o  j

¬ ́·¶ L



(13)

where  ́

is the position vector μ»1⁄4¶3⁄41⁄2¶ oA and ois the vertical coordinate. Subtracting off the hydrostatic pressure gradient~ from the momentum equation (2), and then dividing by the density yields

l

~

L



 n¿ l

¿  wl

m





k

n

jÀw

kIwkH

k

o F



k

n

} (14)

Note the use of the vector identity: lS nÀ l FÁ¿ l

¿  

 wlm. To simplify this equation further,

the density in the pressure term is assumed ambient, and then the term ¿ l

¿  



is combined with the

perturbation pressure jÂkHand written as a total pressure, Ã. This approximation assumes that the

buoyancy generated vorticity dominatesthe baroclinic contribution caused by the non-alignment of

perturbation pressure and density gradients. This is not a good approximation at the small scales

where the actual combustion heat release takes place. However, at the resolvable scales this is

nothing more than the assumption that these are buoyancy dominated flows.

There are numerous ways of handling the viscous terms in the momentum equation – the

simplest is to simply prescribe a constant viscous coefficient in cases where the grid resolution

is fine enough to capture the mixing processes at all relevant scales. More often one is forced

due to lack of adequate grid resolution to use a sub-grid scale model. For the problem of smoke

movement in a warehouse, the latter condition applies. The treatment of sub-grid scale mixing

follows very closely the analysis of Smagorinsky [31]. The stress tensor } in Eq. (14) is replaced

by the Reynolds stress tensor Ä whose components are written in the form

N1ÅÇÆ F  k

 ̈È ̄)É 

¿ËÊÌ¿ÍÊÅÇÆ ¡ ÊÅÇÆ F



  ~2Î ~ Å

»

Æ 

~2Î~ Æ

»

Å  (15)

where È

 ̄ is taken as 0.21, É

F

 ̈Ï»Ï1⁄2ÐÏoA +Ñ

, Ï», Ï1⁄2 and Ï oare the grid cell dimensions, and ¿ÇÊÌ¿F

ÒÊÅÇÆ ÊÅÇÆ . There have been numerous refinements of the original Smagorinsky model [32, 33, 34],

but it is difficult to assess the improvements offered by these newer schemes for the problem of

smoke movement. There are two reasons for this. First, the structure of the fire plume is so dominated by the large scale, resolvable eddies that even a constant eddy viscosity gives results almost

identical with those obtained with the Smagorinsky scheme [4]. Second, the lack of precision in

most large scale fire data makes it difficult to sort out the subtleties associated with these models.

For the time being, the Smagorinsky model with the given È

 ̄ producessatisfactory results for most

large scale applications where boundary layers are not important or are not resolvable.

68
To obtain the pressure perturbation,~

we take the divergence of the momentum equation

l

~

L ÔÓÕ nÃ

F 

(16)

where all the convective and diffusive terms have been incorporated in the term Ó. The resulting

equation for the total pressure Ã is an elliptic partial differential equation

n

Ã

F

w

~

n

l ~ 

L w n

 Ó

(17)

The linear algebraic system arising from the discretization of Eq. (17) has constant coefficients

and can be solved to machine accuracy by a fast, direct (i.e. non-iterative) method that utilizes fast

Fourier transforms. No-flux boundary conditions

~

are specified by asserting that

Ã

~2± F

w«Ö:× (18)

at solid walls, where Ö1⁄4× is the normal component of Óat the wall. This equation asserts that

the normal component of velocity at the wall does not change with time, and indeed remains

zero assuming the flow velocity is initially zero. At open external boundaries it is assumed that

the perturbation pressure is zero, a reasonable assumption as long as the fire is well within the

numerical domain.

Direct Poisson solvers are most efficient if the domain is a rectangular region, although other

geometriessuch as cylinders and spheres can be handled almost as easily. For these solvers, the noflux condition (18) issimple to prescribe at external boundaries. However, many practical problems

involve more complicated geometries. For building fires, doors and windows within multi-room

enclosures are very important features of the simulations. These elements may be included in the

overall domain as masked grid cells, but the no-flux condition (18) cannot be directly prescribed at

the boundaries of these blocked cells due to consistency issues. However, it is possible to exploit

the relatively small changes in the pressure from one time step to the next to enforce the no-flux

condition. At the start of a time step, the components of the convection/diffusion term Óare

computed at all cell faces that do not correspond to walls. Then, at those cell faces that do, set

Ö1⁄4× Fw

~Ã

~2±Ø ÚÙ

Î

× (19)

where Ö:× is the normal component of Ó

at the wall, and Ùis a relaxation factor empirically determined to be about 0.8 divided by the time step. The asterisk indicates the most recent value of

the pressure. Obviously, the pressure at this particular time step is not known until the Poisson

equation is solved. Equation (19) asserts that following the solution of the Poisson equation for the

pressure, the normal component of velocity Î ~2Î × will be driven closer to zero according to × ~

LÜÛ wÙ

Î

× (20)

This is approximate because the true value of the velocity time derivative depends on the solution

of the pressure equation, but since the most recent estimate of pressure is used, the approximation

is very good. Also, even though there are small errors in normal velocity at solid surfaces, the

69
divergence of each blocked cell remains exactly zero for the duration of the calculation, and the

consistency condition (6) ensures global mass conservation. In other words, the total flux into a

given obstacle is always identically zero, and the error in normal velocity is usually at least 3 or 4

orders of magnitude smaller than the characteristic flow velocity. When implemented as part of a

predictor-corrector updating scheme, the no-flux condition at solid surfaces is well maintained.

7.1.2 The Fire

The fire is represented by a large number of Lagrangian elements (particles that move with the

flow) that release heat as they are transported by the thermally-induced motion. Since the fluid

motion determines where the heat is actually released, and the heat release determines the motion,

the large scale features of the coupling between the fire and the smoke transport are retained.

The heat release rate per unit mass of burning fuel is determined from experiment. The spread

of the fire through the fuel array is predicted by the calculation. Smoke is simulated by tracking

the convected elements after the fuel burnout is completed. A specified percentage of the fuel

consumed is assumed to be converted to smoke particulate. Thus, a knowledge of the spatial

distribution of the Lagrangian elements is equivalent to a specification of the smoke particulate

density at any instant of time. This “thermal element” model, which represents the combustion

heat release as a large number of point sources convected by the resolvable flow field, is in fact a

simple combustion model in its own right. It is consistent with more detailed combustion theories

currently in use, and it permits the use of experimental data from fire experiments in a way that

does not violate the consequences of those theories. A more detailed discussion of this model can

be found in Ref. [35].

The volumetric heat release rate term in the energy conservation equation (3) is essentially

the sum of the heat released from the thermal elements that represent the fire. Thermal elements

originate at solid (burning) surfaces. In the case of boxed polystyrene cups, the cardboard box

heats up due to both convective and radiative heat transfer from the surrounding gas. When the

surface heats up to its measured ignition temperature, thermal elements are ejected from the surface

and burned at a prescribed rate (an input heat release rate per unit area). The surface is assumed

to be thermally-thin. This assumption is technically not correct, but a more detailed treatment of

the heat transfer properties of cardboard would complicate an already complicated situation. As

with all the assumptions made in the model, future refinements will be made if warranted. The

assumed thermally-thin cardboard heats up according to its given density k ̄ , specific heat p1 ̄ and

characteristic thickness Ï s° ̄



L

F

D u  Ý D u  Þ wD u  ß k ̄àp1 ̄ Ï

(21)

where D u  Ý ,

D u  Þ and D u  ß represent heat fluxes due to convection, radiative absorption, and radiative

emission. The individual values of the parameters k ̄ , p1 ̄ and Ï

are not relevant here; rather it is their

product that is determined experimentally, and this product is referred to as the lumped thermal

capacitance of the fuel. For a vertical cardboard sample of the cartoned plastic commodity used in

the tests, k ̄Íp1 ̄ Ïwas determined from the LIFT (Lateral Ignition and Flame spread Test) apparatus

at NIST to be 1.5R0.4 kJ/m  K [36]. Note that the data analysis was based on a thermally-thin

assumption, following the method outlined by Ohlemiller [37]. The ignition temperature of the

sample was determined to be (370
 C, 700
 F).

70
The convective heat flux to the surface

D u  Ý is estimated from a correlation of the form

D u Ý F È És

+Ñ És W/m (22)

where És is the difference between the wall and gas temperature, and Èis an empirical constant

of value 1.43 for a horizontal surface and 0.95 for a vertical surface [38].

The radiative heat flux to the surface,

D u  Þ , is calculated based on the assumption that a prescribed

fraction of the heat released from a thermal element is radiated away, and this energy is absorbed

by the surrounding surfaces without attenuation by the surrounding gas. Usually, the fraction of

energy radiated away from the fire is obtained by comparing the total heat release rate with the

convective heat release rate. For a given point  ́

 ̄ on a wall or surface, the radiative flux is given as

D u  Þ Fá

Åãâ ¦;äWåæ  ̈çÅ  èÞ

D uÅ

é ¿  ́

Å w

 ́

 ̄

¿ 

(23)

where  ́

Å

is the position of the ê th thermal element, èÞis the fraction of the heat release rate of the

ê th element, D uÅ, converted into radiative energy, and ç

Å

is the angle formed by the normal to the

surface and the vector  ́

qVw  ́ ̄ . Since there are hundreds of thousands of thermal elements in a

typical calculation, the above summation is made over a sampling of the elements.

As the cardboard heats up, energy is lost to radiation according to

D u  ß Fìë3í sî ̄ wÕsî¤



(24)

where ëis the Stefan-Boltzmann constant ëÚF 
)$a  Ô)()ï W/m /Kî , and íis the emissivity of

the surface. In this case, the emissivity was assumed to be 0.8.

Once the cardboard reaches its ignition temperature, Lagrangian particles representing pyrolyzed fuel are ejected at a rate of D ±

 

particles per unit time per unit area (about 1,000 particles/m  s)

with a small normal velocity (Û0.20 m/s) into the flow domain. The heat release rate for the fire at

a given time L is then expressed by summing over all of the thermal elements in the flow field

D u

 L

F á D u 

D ±

  

N

 L w L

H«ð N

(25)

where the D u is a prescribed rate of heat release per unit area, Nis a characteristic element burn-out

time, and L

H

is the time that a given particle leaves a solid surface. The heat release rate per unit area D u 

is determined from small scale cone calorimeter measurements [39]. Figure 40 displays the heat

release rate per unit area for a small sample compartment consisting of a single polystyrene cup

surrounded by cardboard [40]. The burn-out time is obtained from the plume correlation of Baum

and McCaffrey [41]. In the flame region, the vertical centerline velocity is given as ñ F %$=_)ò ̧%o

for  ðóoôð $&	A·õØ, where õØ F

ED pqk¤Ìs°¤ ò ̧` 

§÷ö is the characteristic diameter of a fire

whose total heat release rate is ED

. The burn-out time is the time it takes a thermal element to reach

the top of the flame region

N Û 

¦1ø  ù1⁄4ú

H



o

%$=_ò ̧%o

F



%$=_:û $&	A·õØ  ̧ (26)

and is usually a few tenths of a second.

71
Figure 40: Heat release rate per unit area of the FMRC Standard Plastic test commodity

obtained from a cone calorimeter measurement. The cone was in a vertical configuration,

and the peak radiative flux on the object was 75 kW/m .

Oxygen transport and consumption is included in the calculation, and the burn-out time of

any thermal element will vary based on the concentration of oxygen in the gas surrounding it. It is

assumed that 1 kg of oxygen is consumed for every 13,100 kJ of energy released. When the oxygen

mass fraction üÂý, falls to a certain prescribed lower limit (about 12%), combustion is assumed to

stop, and the unburned fuel associated with the thermal elements remains unburned until more

oxygen can be found.

Another useful input parameter associated with the given solid fuel is the total amount of

potential energy contained within a unit volume of the unburned fuel. The computational cell

representing a piece of the fuel will disappear once the energy contained within that volume is

consumed. This parameter is referred to as the Energy Per Unit Volume, and is expressed in

units of kJ/m . If prescribed, this parameter permits fire spread through the solid fuel, as in the

combustion of a boxed commodity or a crib.

7.1.3 Sprinkler Activation

The temperature of the sensing element of an automatic sprinkler is estimated from the differential

equation presented by Heskestad and Bill [42, 43]

s°þ



L

F ò

Î

RTI sÂÿÐw s°þ  w ÈRTI sþ`w s



(27)

where s°þ is the link temperature, sÂÿ is the gas temperature in the neighborhood of the link, s

is the temperature of the sprinkler mount, and Îis the gas velocity. The thermal sensitivity of

72
the detector is indicated by the value of RTI10

. The heat lost to the mount due to conduction is

characterized by the “C-factor”, È. A heated wind tunnel (plunge test) is used to determine both

of these parameters by creating an environment in which the air velocity and temperature, plus

the mount temperature, are held at constant values. The C-factor is measured first. There are two

methods of performing the tests [44], both of which are designed to pinpoint combinations of air

temperature and velocity at which an energy balance is established for the heat gained and lost by

the sensing element. For these combinations of air temperature and velocity, the right hand side of

Eq. (27) is zero, and consequently

È

F

 sÂÿwÕs sþÝ w's w

  ò

Î

(28)

Following the determination of the C-factor, the RTI is determined from the solution of Eq. (27)

(assuming the values of sÿ , s, and Î

are constant)

s°þ  L

F s  sÂÿÐw's

 ÔÈò

Î

a  w
	 w

  ÔÈò

Î

 ò

Î

RTI

L

­

(29)

The formula for the RTI is given by

RTI F wL Ý   ÚÈò

Î

 ò

Î


  w  (!  ¦#"%$`§'&(

 
) ( *

(30)

where s°þ+Ý, is the mean liquid bath operating temperature of the sprinkler, and L Ý is the activation

time following the introduction of the sprinkler into the heated wind tunnel.

7.1.4 Sprinkler Droplet Size Distribution

Once a sprinkler has activated, the sizes, temperatures and trajectories of a representative sample of

the water droplets are computed. The initial conditions are deduced from measurements of droplet

sizes and density patterns of sprays not subjected to a fire plume. Note that tracking every droplet

is prohibitively expensive, and unnecessary. The sampling of droplets has been referred to as the

“superdrop” concept [45]. It is directly analogous to the thermal element concept discussed above,

and indeed tracking water droplets does not introduce any new machinery into the numerical code.

In the calculations that will be discussed below, typically five to ten thousand droplets from each

active sprinkler interact with the gas at any given time. This number of droplets ensures that a

sufficient distribution of the water is obtained.

In order to compute the droplet trajectories, the initial size and velocity of each droplet must

be prescribed. This is done in terms of random distributions. The initial droplet size distribution of

the sprinkler spray is expressed in terms of its Cumulative Volume Fraction (CVF), a function that

relates the fraction of the water volume transported by droplets less than a given diameter. Typically, this function is represented very well by a combination of log-normal and Rosin-Rammler

10Often the thermal response of a sprinkler is characterized by a single parameter, also called the Response Time

Index or RTI. This parameter lumps together the RTI as used in Eq. 27 and the C-factor. When reporting an RTI, it

should be made clear what form of the governing equation is being solved. In this report, when only an RTI value is

used to characterize the thermal response of an object, it is referred to an a “bulk” RTI.

73
distributions [46]

Ö

 

·F-. /

.0

¦ &

21 3

H

¦ 4%576 (98 : ;=<?>A@BDCE ,

,GF , Ï  IH 



 w 6 (

H

ø J'K 

 B

B

'L

  ð





(31)

where  is the median droplet diameter (i.e. half the mass is carried by droplets  or smaller

in diameter), and and ëare empirical constants equal to about 2.4 and 0.58, respectively. The

median diameter varies from sprinkler to sprinkler, and it also varies with the pressure applied. It

appears that is less sensitive to changes in sprinkler design and operating conditions.

In the numerical algorithm, the size of the sprinkler droplets are chosen to mimic the RosinRammler/log-normal distribution. A Probability Density Function (PDF) for the droplet diameter

is defined

M

 

·F Ö



 





 N

¤

H

Ö



 ̈Ï 

Ï

 Ï

(32)

Note that Ö



 



denotes the first derivative of the function Ö

 



. Droplet diameters are randomly

selected by equating the Cumulative Number Fraction of the droplet distribution with a uniformly

distributed random variable O

O

 

·F53

H M

 ̈Ï  Ï

(33)

Figure 41 displays typical Cumulative Volume Fraction and Cumulative Number Fraction functions. Of course, every droplet from a given sprinkler is not tracked. Instead, a sampled set of

the droplets is tracked. Typically, 1,000 particles per sprinkler per second are released. The total

number of droplets represented by each computed droplet is given by PRQD 



D ± P

3



, where PRQDis

the mass flow rate of water from a single sprinkler,

D ±

is the number of droplets tracked per sprinkler per second, and P

3

is the average mass of a droplet. The average mass of a droplet can be

expressed in terms of the PDF

P

3 F



	

é

k

Q ¤

H M

 Ï

  Ï

 



Ï

(34)

where k

Q

is the density of water.

7.1.5 Sprinkler Spray Dynamics

The influence of the water spray is introduced into the governing equations of fluid motion through

the force term | in Eq. (2). Thisterm representsthe momentum transferred from the water droplets

to the gas, and it is computed by summing the force from each droplet in a control volume

|

F



TS kp3VUW3 l

3 wyl1⁄4ÿ ¿ l3 wlÿ ¿



Ýz

(35)

where p

3

is a drag coefficient, UW3 is an effective cross sectional area of the particle, lÿ is the

velocity of the surrounding gas, l

3

is the velocity of the droplet, and 

Ýz is the volume of the

control volume. The trajectory of an individual droplet is governed by the equation





L P

3 l3

·F P

3 ow



 k p

3XUW3 lÿwyl3

¿ l1⁄4ÿ wyl3¿

(36)

74
Figure 41: Cumulative Volume Fraction and Cumulative Number Fraction functions of the

droplet size distribution from a typical industrial fire sprinkler. The median diameter  is

1 mm, ë3F `$a
_ and  F %$&	.

where P

3

is the mass of the droplet. The initial droplet velocity and direction are estimated from

measurements of the water flux near the floor through a trial and error sequence of calculations.

Typically, the initial direction is chosen randomly from a range of angles from the vertical through

which the droplet exits the sprinkler deflector.

The sprinkler spray droplet temperature s

3

and radius Y

3

are governed by the following equations

s

3



L

F[Z ]\B

Ý,^Þ B`_=B sÿ wÕs3

 s

3 ð sÂz

 s

3 F sÂz (37)



Y

3



L

F[Z  s

3 ð sÂz w \B \Va _bB sÂÿ wÕs3

 s

3 F sÂz (38)

where pQis the specific heat of water, k

3

is the density of the droplet, sÂÿ is the gas temperature, sz

is the vaporization temperature of water (100
 C), c`z is the heat of vaporization of water, c

3

is the

heat transfer coefficient between air and the water droplet, given by

c

3 F

Nu td 

Y

3

Nu is the Nusselt number given by

Nu F  

`$a Re

+,

Pr

+Ñ

75
The Reynolds number is based on the velocity of the droplet, its diameter and the viscosity of air

at ambient temperature. The Prandtl number is about 0.7, and te is the thermal conductivity of air.

The change in gas temperature due to the presence of the water droplets is given by

sÂÿ



L

F á J PRQ

P

 O 

pQ

pq 

	 c

3

pQY

3 k3

s

3 wÕsÿ 

(39)

where PRQ and P

 denote the mass of water and air in the control volume over which the summation

is being performed.

7.1.6 Sprinkler Spray Interaction with Burning Commodity

When a water droplet hits a solid horizontal surface, it is assigned a random horizontal direction

and moves at a fixed velocity until it reaches the edge, at which point it drops straight down at a

constant speed. This terminal velocity was determined by injecting a colored dye into a stream

of water that cascaded down the side of a carton. It was roughly 0.55 m/s. Both the heating of

unburned surfaces and the heat release rates from burning surfaces are affected by the droplets.

The heat transfer coefficient between the surface and the water film is calculated based on an

empirical correlation for forced flow past a flat plate [47]

Nu F cgfih

tQ

F `$a Re

+,

Pr

+Ñ Û


 for water flowing at 0.55 m/s (40)

The characteristic length h is assumed to be the size of the fuel package. For computational convenience, the water continues to be tracked in the form of droplets, but the heat transfer coefficient

between the water droplets and their surroundings is written in terms of an equivalent heat transfer

coefficient between the thin film of water and the surface. By doing this, the same thermodynamic

formulae may be applied to the droplet whether it is airborne or not. This is a numerical convenience because the transport of water is much easier to describe in terms of Lagrangian droplets

rather than sheets of liquid. If the water formed a thin sheet, its temperature sQ

would be governed

by

sQ



L

F cgf s3 ̄Uw'sQ 

pQjP Q

(41)

On the other hand, as a droplet the water temperature s

3

would be governed by

s

3



L

F

	 c

3

pQY

3 k3

sÂÿwÕs3



(42)

Equating sQ

 

L with s

3  

L and taking sÿ F s° ̄ , an effective heat transfer coefficient is derived

c

3 F cgfkY3 k

3

PQ

(43)

Here, Y

3

is an arbitrarily chosen droplet radius, and P Qis the mass of water per unit area.

76
7.1.7 Extinguishment

Extinguishment of the fire is the single most difficult component of the numerical model. To date,

most of the work in this area has been performed at Factory Mutual. An important paper on the

subject is by Yu et al. [48]. Their analysis yields an expression for the total heat release rate from

a rack storage fire after sprinkler activation

EGFD EQH

D 6 (ml   (no (44)

where EIH

D

is the total heat release rate at the time of application L

H

, and t is a fuel-dependent

constant that for the FMRC Standard Plastic test commodity is given as

t F `$=  PD  Q w

`$a%	` s(

¦

(45)

The quantity PD Qis the flow rate of water impinging on the box tops, divided by the area of exposed

surface (top and sides). It is expressed in units of kg/m /s.

Unfortunately, this analysis is based on global water fluxes and burning rates. The numerical

model requires more detail about the burning rate as a function of the local water flux. Until

better models can be developed, the present extinguishment model consists of an empirical rule

that decreases the local heat release rate as more water is applied

D u  Fqpr  w  PQ

P Q

o 

tsu D uH 

(46)

The critical water density P Q

o is estimated from small scale calorimeter burns of the commodity.

7.1.8 Numerical Methodology

In summary, the equations that are solved numerically are the energy equation (11), the momentum equation (14), and a Poisson equation for the total pressure (17). The background pressure,

temperature and density are found from Eqs. (6), (7) and (10). Each of the conservation equations

emphasize the importance of the divergence and vorticity fields, as well as the close relationship

between the thermally expandable fluid equations [30] and the Boussinesq equations for which

the authors have developed highly efficient solution procedures [49, 3]. These are applied directly

to the equations presented here with minor modifications and no loss in performance. The only

changes from earlier methodology are a return to a uniform rectangular grid with blocks of cells

masked to simulate internal boundaries; and the use of a second order Runge-Kutta scheme to advance the velocity and temperature fields in time. The speed and accuracy of this technique enable

calculations on current generation workstations that involve over a million computational cells,

yielding the spatial range of two orders of magnitude for a three-dimensional calculation.

Table 5 provides a summary of the input parameters required to perform the simulations of

the experiments. As an example, Figs. 42 and 43 show a photograph of Plastic Test P-5 and the

corresponding numerical simulation. The geometry of the simulation consists of the individual

pallet loads of the cartoned plastic commodity, the vents and draft curtains. The wooden pallets

are accounted for by modifying the burning characteristics of the bottom of the simulated boxes.

The supporting racks are not included in the calculation.

77
Input Parameters for an Industrial Fire Simulation

Category Parameter Value

Building Facility UL Large Scale Fire Test Facility

Region of Simulation Below adjustable ceiling + 1 m above

Simulation Dimensions 30.5 m v 30.5 m v 8.2 m (100 ft v 100 ft v 27 ft)

Numerical Grid Dimensions 144 v 144 v 40

Cell Size 0.21 m v 0.21 m v 0.21 m (8 in v 8 in v 8 in)

Lateral Boundary Conditions Ambient Pressure

Ceiling Boundary Conditions Thermally Thick

Floor Boundary Conditions Adiabatic

Ceiling Product Name Armstrong Ceramaguard (Item 602B)

Specific Heat, wyx (Eq. 12) 753 J/kg z K

Conductivity, {x (Eq. 12) 0.0611 W/m z K

Density, |}x (Eq. 12) 313 kg/m~

Fuel Type Racked cartoned plastic commodity

Boxes + Pallet Dimensions 1.1 m v 1.1 m v 1.2 m (42 in v 42 in v 47 in)

Ignition Temperature 370  C

Lumped Thermal Capacitance, |x wx' (Eq. 21) 1.5 kJ/(mz K)

Heat Release Rate Per Unit Area,   (Eq. 25) 500 kW/m

Energy Content Per Unit Volume 500 MJ/m~

Radiative Fraction,  (Eq. 23) 0.50

Emissivity,  (Eq. 24) 0.8

Critical Water Density,  ` (Eq. 46) 0.125 kg/m

Ignitor Peak Heat Release Rate 25 kW

Sprinkler Response Time Index, RTI (Eq. 27) 148 (m z s) BC(268 (ft z s) BC)

C-factor (Eq. (27)) 0.7 (m/s) BC(2.3 (ft/s) BC)

Activation Temperature, ./+ =G (Eq. 28) 74  C (165  F)

Flow Rate 189 l/min (50 gpm)

Spray Angle, (Eq. 36) 5 – 80

Initial Droplet Velocity, (Eq. 36) 8 m/s (26 ft/s)

Median Droplet Diameter, -W@ (Eq. 31) 1 mm (0.04 in)

Sprinkler Spacing 3 m v 3 m (10 ft v 10 ft)

Distance below ceiling 8 cm (3 in)

Vent Dimensions 1.2 m v 2.4 m (4 ft v 8 ft)

Bulk Response Time Index, RTI 175 (m z s) BC(317 (ft z s) BC)

Activation Temperature 74  C (165  F)

Table 5: Input parameters used for the cartoned plastic fire simulations. Note that the grid

dimensions varied from case to case, but the cell size did not. The dimensions listed above

represent the largest calculations performed.

78
Figure 42: Photograph of Plastic Test P-5.

79
Figure 43: Numerical simulation of Plastic Test P-5.

80
7.2 Model Results

The first series of heptane spray burner tests provided a rich set of data with which to validate the

hydrodynamics and sprinkler spray algorithm of the model. Then, the flame spread and fire growth

algorithms were used to predict the results of the calorimetry burns conducted at UL. The empirical extinction algorithm (Eq. 46) was checked against commodity classification data reported

in Ref. [48]. Finally, the model, as developed and verified, was used in the analysis of the cartoned plastic commodity fires to verify that certain phenomena observed during the tests were not

anomalies, and also to explore various “what ifs” that could not be studied experimentally.

7.2.1 Heptane Spray Burner Test Series I

Numerical simulations of the first series of heptane spray burner experiments were performed to

check the accuracy of the hydrodynamic and sprinkler spray models. For Tests I-1–16 in which

the draft curtains were installed, the computational domain extended from the floor to a height

2 m (7 ft) above the ceiling, and encompassed the area enclosed by the draft curtains. For Tests

I-17–22, the domain was extended beyond where the draft curtains were in the previous tests to

account for the extra ceiling area. The numerical grid in all cases consisted of 186,624 cells, with

dimensions 72 by 72 by 36, with the grid cells stretched in the vertical direction to cluster more

cells near the ceiling. The width of each cell was about 0.3 m (1 ft), the height varied from about

0.15 m to 0.45 m (0.5 ft to 1.5 ft). Calculations performed with denser grids did not change the

results significantly.

Heat losses to the ceiling were computed by solving a one-dimensional heat conduction equation for the internal temperature of the insulating tile. Because this is an insulating material, the

heat losses to the ceiling were not significant compared with the energy absorbed by the sprinkler

sprays. Over the fire the energy lost to the ceiling was on the order of 2 to 4 kW/m , and it dropped

off rapidly in the radial direction.

The convective heat release rate for the heptane spray burner was measured using the large hood

at UL. For the sizes of fires used in the tests, about 65% of the total heat release rate was convective.

This was the number used by the model to define the heat output of the fire. The radiative energy

did not play a role in the numerical simulation of a single gas burner. The calculations of the

cartoned plastic commodity fires discussed in the following sections did not neglect radiation heat

transfer because it was an important component of the flame spread algorithm.

The most obvious check of the numerical model is how well it predicted sprinkler activation

times. Examination of the activation times of nearby sprinklers was also important, especially in

cases where several sprinklers were at equal distances from the fire and, in theory, should have

activated at the same time. Table 6 presents a summary of the numerical simulations, comparing

the number and time of sprinkler activations and the average peak temperatures near the fire. Appendix A.1 contains the complete sprinkler activation history for all 22 tests, both experimental

and numerical. Appendix A.2 contains thermocouple data at sprinklers near the fire, along with

the numerical predictions.

In most of the tests, the numerical model predicted the activation of the first four sprinklers

surrounding the fire to within about 5 or 10 s. For the next ring of sprinklers, the model underpredicted the activation times by 15 to 30 s, on average. This difference probably was due to the

uncertainty in the droplet size distribution. Work is still underway to more accurately measure the

81
Heptane Spray Burner Test Series I Simulations

Test Burner Vent Avg. Act. Time No. of Avg. Peak Draft Heat Release

No. Pos. Operation First Row Sprinklers Temp. (C) Curtains Rate

Exp. Sim. Exp. Sim Exp. Sim. MW @ s

I-1 B Closed 1:09 1:08 11 12 137 148 Yes 4.4 @ 50I-2 B Manual (0:40) 1:10 1:12 12 9 134 137 Yes 4.4 @ 50I-3 B Manual (1:30) 1:09 1:12 12 10 138 141 Yes 4.4 @ 50I-4 C Closed 1:05 1:09 10 10 136 142 Yes 4.4 @ 50I-5 C Manual (0:40) 1:16 1:10 9 8 125 132 Yes 4.4 @ 50I-6 C Manual (1:30) 1:24 1:13 8 8 137 134 Yes 4.4 @ 50I-7 C 74  C link (DNO) 1:13 1:09 10 10 133 142 Yes 4.4 @ 50I-8 B 74  C link (9:26) 1:09 1:09 11 12 125 148 Yes 4.4 @ 50I-9 D 74  C link (DNO) 1:18 1:12 12 12 134 143 Yes 4.4 @ 50

I-10 D Manual (0:40) 1:22 1:16 13 12 133 136 Yes 4.4 @ 50I-11 D 74  C link (4:48) N/A N/A N/A N/A 196 152 Yes 4.4 @ 50I-12 A Closed 1:13 1:22 14 12 138 148 Yes 4.4 @ 50I-13 A 74  C link (1:04) 1:29 1:48 5 4 119 114 Yes 6.0 @ 60I-14 A Manual (0:40) 1:58 2:14 7 4 116 108 Yes 5.8 @ 60I-15 A Manual (1:30) 1:08 1:19 5 4 143 153 Yes 5.8 @ 60I-16 A 74  C link (1:46) 2:03 2:56 4 4 110 106 Yes 5.0 @ 110I-17 B 100  C link (DNO) 1:06 1:01 4 10 105 122 No 4.6 @ 50I-18 C 100  C link (DNO) 1:06 1:05 4 9 98 113 No 3.7 @ 50I-19 A 100  C link (10:00) 1:06 1:15 10 12 110 127 No 4.6 @ 50I-20 A 74  C link (1:20) 1:50 1:25 4 4 104 117 No 4.2 @ 50I-21 C 74  C link (7:00) 1:05 1:03 10 10 119 125 No 4.6 @ 50I-22 D 100  C link (DNO) 1:04 1:00 6 12 115 131 No 4.6 @ 50

Table 6: Results of the heptane spray burner Series I simulations. Note that “Avg. Peak Temp.” in this table refers to the average

peak temperature of the nearest 16 sprinkler-mounted thermocouples. The “Avg. Act. Time First Row” is the average activation

time of the nearest 4 sprinklers.

82
droplet size distribution. A sensitivity analysis was performed to determine what parameters had

the most impact on the results of the calculations, and droplet size was shown to be one of the more

important. The reason for this is because the heat transfer between the hot gases and the droplet

is directly proportional to the surface area of the droplet. Thus doubling the size of the droplets

reduces the number of droplets by a factor of 8 and reduces the heat transferred from the gas by a

factor of 4.

7.2.2 Fire Growth Validation

Experimental burns of the cartoned plastic commodity were performed at UL. Two, three and four

tier configurations were tested. The ignition method was the same as the large scale commodity

burns. Figure 44 plots the total11 heat release rates versus time for the three experiments up to the

point where water was applied. Figure 45 shows the convective

12 heat release rates for the same

experiments compared with those computed in the numerical model. Figure 46 shows what one of

these calculations looks like. For the period of time before water application, the simulation heat

release rate is within 20% of the experiment.

Figure 44: Total heat release ratesfor the two, three and four tier cartoned plastic commodity

calorimetry burns. Water was applied to the fire at 130 s in the two tier test, and at 95 s in

the three tier test.

11Based on oxygen consumption calorimetry.

12Based on measurement of the temperature and velocity of the exhaust gases through the hood.

83
Figure 45: Comparison of experimental and simulated convective heat release rates for the

two, three and four tier cartoned plastic calorimetry burns. Water was applied to the fire at

130 s in the two tier test, and 95 s in the three tier test. The simulations were performed with

no water application.

Similar calorimetry experiments have been performed at Factory Mutual. There are two differences between the tests performed at FM and those at UL. The first is that the first pallet in the FM

tests is raised 23 cm (9 in) off the floor whereas the first tier at UL is on the floor. Second, the FM

burns are centrally ignited with four half-ignitors arranged in a pinwheel pattern at the intersection

of the four pallets, whereas the UL burns are ignited in the center of one of the lateral flues with

two half-ignitors, one on each side of the flue. As a result, the fire growth with the FM configuration is more rapid because of the fire has access to more fresh air from all four flues and from

beneath. Figures 47 and 48 present the results of the FM burns, compared with the corresponding

simulations.

The fire growth of the 2, 3 and 4 tier UL burns was also measured by taping bare-bead thermocouples to the sides of the boxes in order to determine at what time the fire had reached that

particular place. Small scale experiments in the LIFT apparatus at NIST indicated a piloted ignition temperature for the cardboard to be about 370 C (700 F) [36, 37]. Using this temperature as

an indicator of ignition, the simulation and experiment could be compared. Tables 7 and 8 give

the locations and ignition times for the 2 and 4 tier experiments and simulations. The comparison

between the two sets of data shows some large differences in ignition times, especially for locations outside of the ignition flue and far from the ignition point. This is to be expected, given the

84
Figure 46: Simulation of three tier calorimetry burn at 30, 60 and 90 s following ignition.

85
Figure 47: Comparison of experimental and simulated heat release rates for the 

and  cartoned plastic calorimetry burns conducted at Factory Mutual. The ignition

point was centrally located and the first tier pallet was raised 23 cm (9 in) off the ground.

The actual experiments were extinguished with a large amount of water. The simulations did

not include a water spray.

simplistic combustion sub-models. Nevertheless, the agreement with the heat release rate data is

encouraging. It provides validation for the methodology for at least a few minutes of simulation,

during which time most of the important interactions are taking place. The model in its present

form can be used to analyze the cartoned plastic commodity experiments for the first few minutes.

For longer times, a better characterization of the burning and extinguishment processes need to be

developed and incorporated into the model.

86
Figure 48: Comparison of experimental and simulated heat release rates for the 

and  cartoned plastic calorimetry burns conducted at Factory Mutual. The ignition

point was centrally located and the first tier pallet was raised 23 cm (9 in) off the ground.

The actual experiments were extinguished with a large amount of water. The simulations did

not include a water spray.

87
Two Tier Cartoned Plastic Calorimetry Burn

TC coordinates (m) Ignition Times (s)

   ¡

Exp. Sim. Diff.

-0.69 0.08 2.27 69 69 0

-0.76 0.08 1.19 99 80 -19

-0.69 0.08 1.12 93 71 -22

-0.76 0.00 2.64 82 78 -4

-0.76 0.00 1.12 100 86 -14

-0.46 0.08 1.19 72 68 -4

-0.28 0.08 1.19 60 55 -5

-0.08 0.08 1.19 53 50 -3

0.08 0.08 1.19 60 50 -10

0.28 0.08 1.19 72 68 -4

0.46 0.08 1.19 84 75 -9

0.27 0.00 1.12 58 57 -1

-0.46 0.00 2.11 50 56 +6

-0.08 0.00 2.11 38 45 +7

0.08 0.00 2.11 47 45 -2

0.46 0.00 2.11 67 67 0

-0.46 0.00 2.64 48 59 +11

-0.28 0.00 2.64 39 49 +10

-0.08 0.00 2.64 38 49 +11

0.08 0.00 2.64 48 49 +1

0.28 0.00 2.64 55 58 +3

0.46 0.00 2.64 62 69 +7

Table 7: Location of thermocouples relative to the ignition point and the ignition times of the

experiment and the numerical simulation for the two tier high cartoned plastic commodity

burn. The 

coordinate indicates the distance along the ignition flue, with the negative direction being towards the center of the array, the  

coordinate indicates the distance down the

non-ignition flue, and the ¡coordinate indicates the height off the floor. All dimensions are

in meters.

88
Three Tier Cartoned Plastic Calorimetry Burn

TC coordinates (m) Ignition Times (s)

   ¡

Exp. Sim. Diff.

-1.50 0.00 4.95 86 89 +3

-0.95 0.00 4.95 72 74 +2

-0.69 0.80 4.95 78 90 +12

-0.69 0.27 4.95 69 69 0

-0.27 0.00 4.95 42 56 +14

0.00 0.00 5.25 42 60 +18

-1.50 0.00 3.50 98 95 -3

-0.95 0.00 3.50 83 75 -8

-0.69 0.27 3.50 71 67 -4

-0.69 0.80 3.50 81 92 +11

-0.27 0.00 3.50 41 44 +3

0.00 0.00 3.75 38 48 +10

-1.50 0.00 1.95 93 104 +11

-0.95 0.00 1.95 93 73 -20

-0.69 0.27 1.95 71 68 -3

-0.69 0.80 1.95 81 90 +9

-0.27 0.00 1.95 38 29 -9

0.00 0.00 2.20 31 31 0

0.27 0.00 1.95 33 35 +2

-0.27 1.07 1.95 81 – –

Table 8: Location of thermocouples relative to the ignition point and the ignition times of the

experiment and the numerical simulation for the four tier high cartoned plastic commodity

burn. The coordinate indicates the distance along the ignition flue, with the negative direction being towards the center of the array, the  

coordinate indicates the distance down the

non-ignition flue, and the ¡coordinate indicates the height off the floor. All dimensions are

in meters.

89
7.2.3 Cartoned Plastic Commodity Fire Simulations

A very useful application of the numerical model has been to simulate the five cartoned plastic

commodity tests discussed in Section 5 in order to explain and evaluate the results of the experiments. The benefit of the numerical model in this application is that it provides a consistent means

of varying test parameters. In other words, if two calculations are performed in which only one

input parameter is different, then the difference in the results of the two calculations can safely be

attributed to the difference in the input parameter. A drawback of large scale testing is that this

type of sensitivity analysis usually requires more tests than can be afforded. If a sufficient number

of replicates cannot be performed, then the outcomes of the experiments are often subject to debate

as to whether differences in test results were due to changes in test parameters or due to random

variations.

Consider, for example, the different outcomes of Tests P-1 and P-4 in which the only difference

in test parameters was that draft curtains were installed for Test P-4, but not for Test P-1. Twenty

sprinklers activated in Test P-1, five in Test P-4. Was this result due to the draft curtains? One

would expect that the draft curtains would have increased the number of activations, not decreased

them, based on the results of the first heptane spray burner series. To gain some insight, the

numerical model was run to simulate the experiments. Test P-1 was run two ways. First, the

simulation was run with no manipulation of the sprinkler activation times (Fig. 49). Second, the

calculation was repeated, but with the second sprinkler intentionally delayed a minute to mimic

what actually occurred during the experiment (Fig. 50). The reason for modifying the sprinkler

activation pattern was to see how the delay could have led to the different outcomes observed in

the tests. Tests P-4 and P-5 were simulated without any modification of the sprinkler activation

times (Figs. 51 and 52). In the simulations of Test P-4 and Test P-1 without any imposed sprinkler

delay, 4 sprinklers activated. In the simulation of Test P-1 with the sprinkler delay, 17 sprinklers

activated within the first 5 min. In the actual Test P-1, 20 sprinklers activated, but over about

15 min. The model demonstrated the effect of delaying the second sprinkler. The difference in

outcomes of Tests P-1 and P-4 of the cartoned plastic test series was not due to the draft curtains,

but rather to the sprinkler delay in Test P-1.

Figure 53 presents the heat release rates for the simulations of Test P-1 (with and without

sprinkler delay) and Test P-4. The activation times for the first sprinkler in all cases was between

69 and 71 s. In the case where the second sprinkler is delayed from opening for 58 s (Test P-1,

delay), the heat release rate is substantially greater than the cases where both sprinklers activate

nearly at the same time (Test P-1, no delay and Test P-4).

If the increased number of sprinkler activations of Test P-1 over Test P-4 can be attributed to

the delay of the second nearest sprinkler, did the draft curtains have any effect at all? Comparing

the simulation of Test P-4 with the ‘no delay’ simulation of Test P-1, there is not much difference

in heat release rates. The overall temperatures near the ceiling are slightly higher in Test P-4, most

likely due to the containment of the hot gases by the curtains. This was seen in the analysis of the

first series of heptane spray burner tests.

Another interesting case to examine with the numerical model is Test P-2, where the ignition

point was directly under the vent, but the vent did not open. What if it had? A calculation was

performed where the vent was set to open automatically with a fusible link, and another calculation

was done with the vent held closed. In the first calculation, the vent opened after 83 s, followed a

half second later by the first sprinkler and 3 s later by the second sprinkler. In the calculation with

90
Figure 49: Results of Plastic Test P-1 simulation with no imposed delay in the second sprinkler activation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

Figure 50: Results of Plastic Test P-1 simulation with the

same delay imposed on the second sprinkler activation as was

seen in the actual experiment. “S” denotes simulation sprinkler activation time, “E” experiment. The cross-hatched area

indicates temperatures between 100 and 120  C.

91
Figure 51: Results of Plastic Test P-4 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

Figure 52: Results of Plastic Test P-5 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

92
Figure 53: Heat release rates from the numerical simulations of Tests P-1 and P-4 of the

Plastic series. The term ‘delay’ means that the second sprinkler was intentionally opened

58 s after the first in order to mimic the actual experiment.

the vent closed, the first sprinkler opened after 83 s, followed 1.4 s later by the second. Figures 54

and 55 present the results of the simulations. In the simulation where the vent opened, 14 sprinklers

activated in the first 7 min following ignition. In the simulation where the vent was not allowed

to open, 19 sprinklers activated in the first 7 min. This latter simulation was comparable to the

actual experiment where 23 sprinklers opened in the first 6 min. The opening of the vent in the first

simulation reduced the number of sprinkler activations by exhausting heat from the fire directly

underneath. There is certainly plenty of evidence from the heptane spray burner tests indicating

that when the fire was placed directly beneath an opened vent, the number of activations was

significantly reduced.

Less clear, however, is why so many sprinklers activated in Test P-2. The simulation of Test

P-1 with no manipulation of the sprinkler activation times produced only 4 activations. What was

the difference between the simulation of Test P-1 and Test P-2? Only the presence of a 1.2 m by

2.4 m by 0.3 m deep (4 ft by 8 ft by 1 ft deep) cavity in the ceiling formed by the vent in Test

P-2. This cavity led to a 14 s delay in the first sprinkler activation in the simulation of Test P-2.

The significance of this delay is shown in Fig. 56, in which the heat release rate histories for the

simulations of the two versions of Test P-2 are plotted on the same graph as the heat release rate

curves for the 2 by 2 by 4 tier cartoned plastic calorimetry experiment and the simulation of Test

P-1 with no second sprinkler delay. The growth of the fire during the time period 60 to 100 s after

93
Figure 54: Results of Plastic Test P-2 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

Figure 55: Results of Plastic Test P-2 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

94
ignition was very fast, and it was demonstrated that even a 14 s delay in sprinkler activation could

significantly alter the number of sprinkler activations.

Figure 56: Heat release rates from the numerical simulations of Test P-2 of the Plastic series,

plus the 2 by 2 by 4 tier calorimetry burn of cartoned plastic commodity and the simulation

of Test P-1 with no second sprinkler delay.

The cartoned plastic burn that caused the most fire damage was Test P-3. In this test, the

ignition point was placed close to the intersection of two draft curtains (Fig. 29). An estimated

184 boxes were consumed during the test, compared to 103 and 81 in Tests P-4 and P-5, the other

two Plastic tests conducted with draft curtains in place. Figures 57 and 58 present the results

of the simulations of Tests P-3 and P-4. The vent nearest the ignition point in Test P-3 opened

at 4:11, and no vents opened in Test P-4, thus venting did not have any impact on the results

in either case, at least for the first 4 min. Figure 59 shows the convective heat release rates for

the two simulations. Clearly, the draft curtains had an effect on the performance of the sprinkler

system. The draft curtains delayed the opening of the two sprinklers directly north of the first two

sprinklers to activate. Less obvious, the draft curtains changed the near-ceiling flow pattern of both

the sprinkler spray and the fire plume. Regardless of the sub-model used to simulate the burning

of the cartoned plastic commodity, the calculation showed that less water reached the north side

of the central array when the draft curtains were installed. Figure 60 shows a snapshot of the two

simulations after 5 min. In Test P-3, the fire has spread to the north face of the array, whereas in

Test P-4 the sprinklers in the north aisle prevent the spread to the north face.

95
Figure 57: Results of Plastic Test P-3 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

Figure 58: Results of Plastic Test P-4 simulation. “S” denotes

simulation sprinkler activation time, “E” experiment. The

cross-hatched area indicates temperatures between 100 and

120  C.

96
Figure 59: Heat release rates from the numerical simulations of Tests P-3 and P-4 of the

cartoned plastic series.

7.2.4 Mass Flow Rates through Vents

One final use of the model was to simulate certain experiments in which important measurements

could not be made. Attempts were made to measure the velocity through an opened vent to determine the mass flow rate, but the bi-directional probes produced erroneous results. These measurements of mass flow rate would have been important because the opened vents did not have a

significant effect on the number of activated sprinklers or the near-ceiling gas temperatures during

the second series of heptane spray burner tests (except in cases where the fire was directly underneath a vent). The heptane test data seems at odds with some experiments performed in the past,

notably the Ghent experiments of 1989 [14]. The discrepancy is most likely due to different vent to

floor ratios and sprinkler flow rates. These parameters would directly affect the near-ceiling temperatures and the mass flow rates through the vents. Hinkley points out in the SFPE Handbook [50]

that for temperature rises less than about 75 C (167 F) there is a serious decrease in the mass flow

rate through a vent. This assessment is based on the expression for mass flow rate through a vent

due to buoyancy ¢ £¥¤§¦ ̈a©D«n¬­ ̄®±°n233¬ 3¬μ ́ 23 (47)

where ¦ is an orifice coefficient equal to about 0.6,  ̈a© is the area of the vent, «n¬ is the density

of air, ® is the acceleration of gravity, ° is the depth of the layer of hot gases, 23 is the average

temperature rise in the layer, and 3¬ is the (absolute) ambient temperature. Figure 61 displays

97
LES Simulation of NFPRF Sprinkler/Vent/Draft Curtain

 Group A Plastic Tests 4 (top) and 3 (bottom)

Ignition point near center of curtained area

Ignition point near intersection of draft curtains

Figure 60: Snapshots of the simulations of Plastic Tests P-3 and P-4 after 5 min.

98
Figure 61: Mass flow rate through a single vent as a function of ceiling layer temperature.

The solid line represents the mass flow predicted by Eq. 47 for a 1.2 m by 2.4 m (4 ft by

8 ft) vent and a ceiling layer depth of 1.8 m (6 ft). The points represent the average mass

flow through the vent nearest the fire over the time period between 2 and 5 min for three

numerical simulations.

the idealized mass flow rate as a function of temperature rise (solid curve). In Appendix A of the

1991 edition of NFPA 204M, Guide for Smoke and Heat Venting, the derivation of the venting

relationships contained in that document is explained [5]. Equation 47 is used to estimate the

mass flow through a vent, and the vent to floor ratios recommended in the document are based

on equating the mass flow of the air entrained into the fire plume with that exhausted through the

vents. An important assumption made is that the temperature rise 23 in Eq. 47 is equal to the

99
ambient temperature on the Kelvin (Rankine) scale. Thus, if the ambient temperature were 20 C

(68 F), then the temperature rise 2¶3 near the vent would be assumed to be 293 C (559 F). At

this temperature, the mass flow through the vent would be very near its theoretical maximum. Of

course, this analysis does not take into account the effect of sprinkler sprays because it is stated in

Chapter 6-1 of the 1991 edition of 204M that the document “...representsthe state of technology of

vent design in the absence of sprinklers.” Indeed, the test data and the model predictions reported

here indicate that the temperature increase over ambient in the vicinity of an opened vent in a

sprinklered facility would be far less than 293 C.

The numerical model was used to estimate the mass flow rates through the vent nearest the

fire. Figure 61 presents the rates of three typical calculations compared to the rate predicted by

Eq. 47. The mass flow rates for Test I-10 and P-5 are relatively low compared with the theoretical

maximum because the near-ceiling gas temperatures are greatly reduced by the sprinklers. The

flow rate for Test II-2 is much higher because the ceiling layer temperatures are significantly higher

for the 10 MW fire. The simulation of Test II-2 was rerun with the draft curtains removed. The

computed mass flow from the numerical simulation dropped into a range of 1.5 kg/s to 2.0 kg/s.

In terms of Eq. (47), this reduction in mass flow rate is due to the decrease in the smoke layer

depth, ·, but another contribution is the change in ceiling jet dynamics caused by the draft curtain

removal. This latter effect is not accounted for in Eq. (47), but it is in the numerical model.

100
8 Conclusions

Thirty-nine large scale fire tests were conducted at the Underwriters Laboratories Large Scale Fire

Test Facility in Northbrook, Illinois, to investigate what effect roof vents and draft curtains have on

the time, number and location of sprinkler activations; and what effect sprinklers and draft curtains

have on the time, number and discharge rates of roof vents in a warehouse or warehouse-like retail

store. In addition, a computational fluid dynamics model was used as a planning and analysis tool.

The test site and experimental test parameters were chosen by an industry-led Technical Advisory

Committee to address relatively large, open-area buildings with smooth, unobstructed (except for

draft curtains) horizontal ceilings, adequate sprinkler systems and independently-controlled (i.e.

not grouped) automatic roof vents. Because the smoke was vented into a large plenum space

and not the atmosphere, wind effects were not considered and the effect of venting on smoke

obscuration could not be quantified.

The major findings relative to the interaction of sprinklers, draft curtains and vents based on

the experiments and model simulations in this study were:

 ̧

The tests and model simulations showed that when the fire was not ignited directly under

a roof vent, venting had no significant effect on the sprinkler activation times, the number

of activated sprinklers, the near-ceiling gas temperatures, or the quantity of combustibles

consumed.

 ̧

The tests and model simulations showed that when the fire was ignited directly under a roof

vent, automatic vent activation usually occurred at about the same time as the first sprinkler

activation, but the average activation time of the first ring of sprinklers was delayed. The

length of the delay depended on the difference in activation times between the vent and the

first sprinkler.  ̧

The tests and model simulations showed that when the fire was ignited directly under a roof

vent that activated either before or at about the same time as the first sprinkler, the number

of sprinkler activations decreased by as much as 50% compared to tests performed with the

vent closed.

 ̧

The tests and model simulations showed that when draft curtains were installed, up to twice

as many sprinklers activated compared to tests performed without curtains.

 ̧

In one rack storage test where the ignition of the fire took place near a draft curtain and the

fuel array extended underneath the curtain, disruption of the sprinkler spray and delay in

sprinkler operation caused by the draft curtain led to a fire that consumed more commodity

compared to the other tests where the fires were ignited away from the draft curtains. This

result was demonstrated by the model simulation, as well.

 ̧

The significant cooling effect of sprinkler sprays on the near-ceiling gas flow often prevented

the automatic operation of vents. This conclusion is based on thermocouple measurements

within the vent cavity, the presence of drips of solder on the fusible links recovered from

unopened vents, and several tests where vents remote from the fire and the sprinkler spray

activated. In one cartoned plastic commodity experiment, a vent did not open when the fire

was ignited directly beneath it. The model simulations could not predict this phenomenon.

101
 ̧

Model simulations indicated that the cooling effect of sprinkler sprays reduced the total vent

discharge rate from that assumed in design calculations for unsprinklered buildings.

 ̧

Model simulations showed how the activation times of the the first and second sprinklers

had a substantial impact on the overall number of activations in the plastic commodity tests.

In the simulation of one test, it was shown that a delay of approximately one minute in the

activation of the second sprinkler led to the activation of four times as many sprinklers as in

a simulation of a test with no delay. It had been suggested that these different outcomes were

due to the presence of draft curtains in the tests with the sprinkler delay, but the simulations

showed that the curtains had no effect because they were over 9 m (30 ft) away from the

ignition point.

The Industrial Fire Simulator (IFS) developed in conjunction with the test program was shown

to be in good quantitative agreement with the heptane spray burner tests in terms of both predicting

sprinkler activation times and near-ceiling gas temperatures. The sprinkler activation times were

predicted to within about 15% of the experiments for the first ring, 25% for the second. The gas

temperatures near the ceiling were predicted to within about 15%. Simulations were performed

and compared with unsprinklered calorimetry burns of the cartoned plastic commodity. The heat

release rates of the growing fires were predicted to within about 20%.Simulations of the 5 cartoned

plastic commodity fire tests were then performed. The goal of these simulations was to be able

to differentiate between those experiments that activated a large number of sprinklers, and those

that activated a small number. This goal has been met. The model also provided valuable insight

into what occurred in the experiments, and also what would have occurred in the event of various

changes of test parameters. There are plans to continue the development of the IFS model beyond

publication of this report.

102
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106
A Heptane Spray Burner Test Results (Series I)

The experimental data and numerical model results presented in this section represent a subset of

both the actual data collected by UL and the results of the model.

A.1 Sprinkler Activation History

The figures on the following pages present the sprinkler activation history of the 22 heptane burner

tests (Series I), plus the model predictions. In each figure, the sprinkler activation times from

the simulation (“S”) and the experiment (“E”) are listed at the location of the activated sprinkler.

The temperature contours are instantaneous snapshots from the simulation, showing the predicted

temperatures near the ceiling at some arbitrary time during the simulation. Highlighted with a

cross-hatched filling is the contour interval from 75 C to 100 C. Contours are spaced at intervals

of 25 C. The activation temperature for the Central ELO-231 upright sprinklers used in the tests

is 74 C. The highlighted contour interval separates the area of almost certain activation from the

area of almost no activation. It was observed during the tests that following the activation of the

4 sprinklers nearest the fire after about 65 s, the sprinklers in the next ring would activate if the

surrounding gas temperature was at least 100 C. Thus, the 75–100 C interval represents an area

of less than likely activation. The contours also show the effect of opening a vent. Thermocouple

data along with that of the simulations is included in Appendix A.2.

Figure 62: Sprinkler activations for Test I-1. Figure 63: Sprinkler activations for Test I-2.

107
Figure 64: Sprinkler activations for Test I-3. Figure 65: Sprinkler activations for Test I-4.

Figure 66: Sprinkler activations for Test I-5. Figure 67: Sprinkler activations for Test I-6.

108
Figure 68: Sprinkler activations for Test I-7. Figure 69: Sprinkler activations for Test I-8.

Figure 70: Sprinkler activations for Test I-9. Figure 71: Sprinkler activations for Test I-10.

109
Figure 72: Near-ceiling gas temperatures

from the simulation of Test I-11. There were

no sprinklers deployed in the experiment.

Figure 73: Sprinkler activations for Test I-12.

Figure 74: Sprinkler activations for Test I-13. Figure 75: Sprinkler activations for Test I-14.

110
Figure 76: Sprinkler activations for Test I-15. Figure 77: Sprinkler activations for Test I-16.

Figure 78: Sprinkler activations for Test I-17. Figure 79: Sprinkler activations for Test I-18.

111
Figure 80: Sprinkler activations for Test I-19. Figure 81: Sprinkler activations for Test I-20.

Figure 82: Sprinkler activations for Test I-21. Figure 83: Sprinkler activations for Test I-22.

112
A.2 1

Experimental and Simulated Temperatures

The plots in Figs. 85–106 on the following pages compare the experimental thermocouple temperatures with those of the model. For each of the 22 tests, 104 thermocouples were placed near the

ceiling. Their locations are shown in Figure 84. Only the 16 thermocouples nearest the fire are

shown for each test on the following pages. For fire locations A, B and C, these thermocouples

(channels 50–65) were positioned very close to the sprinkler. For fire location D, 4 of the 16 thermocouples (58, 59, 62, 63) were positioned next to the sprinkler, the rest were placed 2 in below

the ceiling directly above the sprinkler deflector. Following sprinkler activation, the temperature of

the corresponding thermocouple decreases sharply. The experimental and simulated temperatures

are not expected to agree following activation because the actual thermocouple is wet, whereas the

simulated thermocouple continues to record the gas temperature.

113
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Figure 84: Plan view of heptane spray burner configuration, Series I. The sprinklers are

indicated by the solid circles and are spaced 10 ft apart. The number beside each sprinkler

location indicates the channel number of the nearest thermocouple. The vent dimensions are

4 ft by 8 ft. The boxed letters A, B, C and D indicate burner positions. Corresponding to each

burner position is a vertical array of thermocouples. Thermocouples 1–9 hang 7, 22, 36, 50,

64, 78, 92, 106 and 120 in from the ceiling, respectively, above Position A. Thermocouples 10

and 11 are positioned above and below the tile directly above Position B, followed by 12–20

that hang at the same levels below the ceiling as 1–9. The same pattern is followed at Positions

C and D, with thermocouples 21–31 at C and 32–42 at D.

114
Figure 85: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-1 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

115
Figure 86: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-2 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

116
Figure 87: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-3 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

117
Figure 88: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-4 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

118
Figure 89: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-5 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

119
Figure 90: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-6 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

120
Figure 91: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-7 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

121
Figure 92: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-8 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

122
Figure 93: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-9 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

123
Figure 94: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-10 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

124
Figure 95: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-11 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

125
Figure 96: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-12 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

126
Figure 97: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-13 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

127
Figure 98: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-14 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

128
Figure 99: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-15 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

129
Figure 100: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-16 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

130
Figure 101: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-17 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

131
Figure 102: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-18 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

132
Figure 103: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-19 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

133
Figure 104: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-20 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

134
Figure 105: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-21 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

135
Figure 106: Comparison of experimental (dashed curve) and simulated (solid curve) temperatures of the near-sprinkler thermocouples for Test I-22 of the first series of heptane spray

burner tests. The activation temperature of the sprinkler (74 C, 165 F) is denoted with a

short dashed line. See Fig. 1 for the location of the sprinklers.

136
B Heptane Spray Burner Test Results, Series II

Figures 107–118 on the following pages are the temperatures for each of the 12 tests of the second

heptane spray burner series.

Figure 107: Temperature histories for the 100 near-sprinkler thermocouples for Test II-1,

Heptane Series II.

137
Figure 108: Temperature histories for the 100 near-sprinkler thermocouples for Test II-2,

Heptane Series II.

138
Figure 109: Temperature histories for the 100 near-sprinkler thermocouples for Test II-3,

Heptane Series II.

139
Figure 110: Temperature histories for the 100 near-sprinkler thermocouples for Test II-4,

Heptane Series II.

140
Figure 111: Temperature histories for the 100 near-sprinkler thermocouples for Test II-5,

Heptane Series II.

141
Figure 112: Temperature histories for the 100 near-sprinkler thermocouples for Test II-6,

Heptane Series II.

142
Figure 113: Temperature histories for the 100 near-sprinkler thermocouples for Test II-7,

Heptane Series II.

143
Figure 114: Temperature histories for the 100 near-sprinkler thermocouples for Test II-8,

Heptane Series II.

144
Figure 115: Temperature histories for the 100 near-sprinkler thermocouples for Test II-9,

Heptane Series II.

145
Figure 116: Temperature histories for the 100 near-sprinkler thermocouples for Test II-10,

Heptane Series II.

146
Figure 117: Temperature histories for the 100 near-sprinkler thermocouples for Test II-11,

Heptane Series II.

147
Figure 118: Temperature histories for the 100 near-sprinkler thermocouples for Test II-12,

Heptane Series II.

148