TECHNICAL MEMORANDUM NUREG/CR-2490

ANL-CT-81-17

(Distribution Codes:

RI and XA)

ARGONNE NATIONAL LABORATORY

9700 South Cass Avenue

Argonne, Illinois 60439

HAZARDS TO NUCLEAR POWER PLANTS

FROM LARGE LIQUEFIED NATURAL GAS (LNG)

SPILLS ON WATER

by

C. A. Kot, T. V. Eichler,*+ A. H. Wiedermann,*+

R. Pape,* and M. G. Srinivasan

Components Technology Division

Manuscript Completed: March 1981

Date Published: November 1981

Prepared for

Division of Risk Analysis

Office of Nuclear Regulatory Research

U. S. Nuclear Regulatory Commission

Washington, D. C. 20555

under Interagency Agreement DOE 40-550-75

NRC FIN No. A2072

*Staff of IIT Research Institute, Chicago, Illinois

+Currently affiliated with ATResearch Associates, Glen Ellyn, Illinois
ABSTRACT

The hazards to nuclear power plants arising from large spills of

liquefied natural gas (LNG) on water transportation routes are treated by

deterministic analytical procedures. Global models, which address the

salient features of the LNG spill phenomena are used in the analysis. A

coupled computational model for the combined LNG spill, spreading, and fire

scenario is developed. To predict the air blast environment in the vicinity

of vapor clouds with "pancake-like" geometries, a scalable procedure using

both analytical methods and hydrocode calculations is synthesized. Simple

response criteria from the fire and weapons effects literature are used to

characterize the susceptibility of safety-related power plant systems. The

vulnerability of these systems Is established either by direct comparison

between the LM threat and the susceptibility criteria or through simple

response calculations.

The analysis and results indicate that the spreading of LNG vapor

clouds up to the lower flammability limit Is dominated by gravitational

effects. Severe fire and blast hazards occur only at locations directly

engulfed by the LNG vapor cloud or in its immediate vicinity. Thermal loads

resulting from an LNG fire are of short duration and can in general be

tolerated by the safety-related power plant systems and components. On the

other hand, blast loads from LNG vapor cloud explosions can cause severe

damage to those systems. The safety standoff distance between the power

plant site and the LNG spill location is primarily dependent on the windInduced LNG vapor cloud drift. Under strong wind conditions (8.96 m/s) it

is estimated, that severe effects on the power plant may be experienced at

distances in excess of 10 km in the down wind direction. To reach a no

damage level under these adverse conditions a standoff distance of

approximately 15 km may be required.

NRC FIN No. Title

A2072 Hazards to Nuclear Power Plants from Large Liquefied Natural

Gas (LNG) Spills on Water Transportation Routes

Ii:
PREFACE

This report presents the results of an investigation conducted for the

U. S. Nuclear Regulatory Commission (NRC), Office of Nuclear Regulatory

Research, Division of Risk Analysis. The work was performed under a

Standard Order for DOE Work (FIN No. A 2072) with Dr. N. A. Eisenberg, NRC

serving as project monitor. His helpful suggestions and reviews are

gratefully acknowledged. The work reported herein was a joint effort of the

Argonne National Laboratory (ANL) and the lIT Research Institute of Chicago,

Illinois.

C. A. Kot, Manager

Structural Systems Analysis Section

Components Technology Division

Argonne National Laboratory

November 1981

i-i
iv
TABLE OF CONTENTS

Page No.

EXECUTIVE SUMMARY 1

1. INTRODUCTION 4

2. PROBLEM DEFINITION 6

3. LNG SPILL AND POOL DYNAMICS 9

4. LNG VAPOR CLOUD DYNAMICS 20

4.1 Vapor Cloud Model 20

4.2 Result Summary of LNG Vapor Cloud

Spreading 29

5. LNG FIRES 47

5.1 Fire Model Description 47

5.1.1 Pool Fires 47

5.1.2 Diffusion Flames 52

5.1.3 Premixed Vapor Flames 53

5.2 Integrated LNG Dispersion and Fire Model 55

5.3 Result- Summary for LNG Fires 61

6. LNG EXPLOSIONS 79

6.1 Vapor Cloud Detonation Parameters 80

6.2 Two-Dimensional Explosion Models 81

6.2.1 Quasi-steady Detonation Field 85

6.2.2 Air, Blast Environment Computations 86

6.3 Air Blast Result Summary 89

7. LNG SPILL THREATS 98

7.1 LNG Pool 98

7.2 Vapor Cloud 99

7.3 Fire Potential 100

7.4 Fire Threat Magnitudes 101

7.5 Explosion Threat 104

7.6 Threat Summary and Application 106

8.* SURVEY OF NUCLEAR POWER PLANT SAFETY SYSTEMS 110

9. NUCLEAR POWER PLANT SUSCEPTIBILITY 116

9.1 Personnel Exposure Limits 117

9.2 Air Blast Effects on Structures and Components 118

9.3 Thermal Susceptibility of Structures 119

9.4 Ignition of Combustibles 122

9.5 Fracture of Brittle Materials 125

10. VULNERABILITY OF NUCLEAR POWER PLANTS 129

10i.1 Personnel Vulnerability 130

10.2 Ignition Vulnerability 131

10.3 Thermal Load Vulnerability 133

v
TABLE OF CONTENTS (Cont'd)

Page No.

10.3.1 Response of Ceramic Electrical Insulators

10.3.2 Effects on Steel Structures and Components

10.3.3 Thermal Response of Concrete Walls

10.4 Air Blast Vulnerability

10.5 Summary of Vulnerabilities

11. DISCUSSION, CONCLUSIONS, AND RECOMMENDATIONS

REFERENCES

APPENDIX A: Study of Liquid Natural Gas (LNG) Spill

Phenomenology on Water

APPENDIX B: Study of Liquid Natural Gas (LNG) Spill,

Dispersion, and Combustion Phenomenology

APPENDIX C: Study of Explosion and Fast Deflagration

of LNG Vapor

133

135

143

158

160

164

168

Inside

back

Cover

vi
LIST OF FIGURES

No. Title

1 Model of LNG Pool Spread.

2 Gravity Outflow from Tank of Uniform Cross Section.

3 Influence of Release Duration on Pool Spread;

Total Volume of Spill - 25,000 a .

4 Influence of Release Form on Pool Spread;

Total Volume of Spill - 25,000 mi.

5 Pool Vaporization Histories; Variation with Spill Duration;

Vaporization Rate: 0.003 m/s, ---- 0.006 m/s.

6 Pool Dynamics.

7 Basic Cloud Dispersion Model.

8 Drift Effects on Cloud Dispersion Model.

9 Cloud and Pool Boundary Trajectories.

10 Plan View of LFL Contours.

11 Comparison of Maximum Downwind LFL Distances.

12 Influence of Wind Speed on Downwind Transport.

13 Effect of Wind Speed on Vapor Cloud Volumetric

Concentration.

14 Influence of Spill Duration on Cloud Radius and Drift;

Wind Speed - 4.48m/s.

15 Effect of Relative Humidity on Vapor Cloud Radius for

Pure Methane; Wind Speed - 4.48 m/s.

16 Effect of Relative Humidity on Vapor Cloud Radius for

LNG (1.2); Wind Speed - 4.48 m/s.

17 Influence of Relative Humidity and LNG Composition on

Downwind Distance to LFL; Wind Speed - 4.48 m/s.

18 Emitting-Surface Fire Model.

19 Configuration-Factor Models.

20 Model for Flame Propagating Through a Premixed Cloud.

21 Flame Path for Cloud Fire; Downwind Ignition, Ignition

Time - 100 s, Flame Speed -1 m/s, Wind Speed - 4.48 m/s.

Page No.

10

13

15

16

17

19

22

25

27

28

30

35

37

39

42

43

44

49

50

54

57

vii
No. Title (Contd)

22 Flame Path for Cloud Fire; Downwind Ignition, Ignition

Time - 100 s, Flame Speed - 5 m/s, Wind Speed - 4.48 m/s.

23 Flame Path for Cloud Fire; Upwind Ignition, Ignition

Time - 100 s, Flame Speed - 1 m/s, Wind Speed - 4.48 m/s.

24 Flame Path for Cloud Fire; Upwind Ignition, Ignition

Time 100 s, Flame Speed = 5 m/s, Wind Speed - 4.48 m/s.

25 Thermal Radiation Histories for Pool Fire; Spill Duration

- 600 s, Wind Speed - 4.48 m/s, Target Evaluation - 0.

26 Thermal Radiation Histories at Vertical Targets; Downwind

Ignition at 100 s, Spill Duration - 600 s, Wind Speed -

4.48 m/s, Flame Speed - 5 m/s.

27 Extent of Flame Region as a Function of Ignition Delay Time

for Flame Speeds of 1 m/s; Wind Speed - 0.

28 Extent of Flame Region as a Function of Ignition Delay Time

for Flame Speeds of 5 m/s; Wind Speed - 0.

29 Extent of Flame Region as a Function of Ignition Delay Time

for Flame Speeds of 10 m/s; Wind Speed - 0.

30 Time Durations of Pool, Cloud, and Total Fires as a Function

of Ignition Delay Time for Flame Speed of 5 m/s.

Page No.

58

59

60

62

63

65

66

67

68

31 Total

Level

Flame

32 Total

Level

Flame

33 Total

Level

Flame

Thermal Load Received by Vertical Targets at Ground

as a Function of Distance and Ignition Delay Time for

Speeds of 1 m/s.

Thermal Load Received by Vertical Targets at Ground

as a Function of Distance and Ignition Delay Time for

Speeds of 5 m/s.

Thermal Load Received by Vertical Targets at Ground

as a Function of Distance and Ignition Delay Time for

Speeds of 10 m/s.

69

70

71

73

74

75

34 Pool Fire Contribution to Total Thermal Load Received by Vertical Targets at Ground Level with Ignition Delay Time

as Parameter.

35 Comparison of Total Thermal Loads Received by Vertical and

Horizontal Targets at Ground Plane.

36 Peak Thermal Load Flux at Vertical Targets at Ground Level

as a Function of Distance and Ignition Delay Time for Flame

Speeds of 5 m/s.

viii
No. Title (Contd) Page No.

37 Peak-to-Average Radiant Flux 1atio as a Function of Distance

and Ignition Delay Time. 76

38 Typical Cloud Configuration at 10Z Fuel Concentration. 83

39 Assumed Explosion Scenario. 84

40 Wave Diagram for Quasi-Steady Detonation Field. 87

41 Dimensionless Pressure History/Profile Behind Detonation

Front at Ground Level. 88

42 Pressure Histories for LNG Cloud Explosion. 90

43 Pressure Histories for LNG Cloud Explosion, (Cont'd) 91

44 Peak Overpressure Variation for LNG Cloud Explosion. 93

45 Impulse Variation for LNG Cloud Explosion. 94

46 Comparisons of Peak Overpressure Estimates for LNG Cloud

Explosion. 96

47 Comparisons of Impulse Estimates for LNG Cloud Explosion. 97

48 Temperature Rise in Steel Plate vs. Heat Flux;

Flux Duration 600 s, 137

49 Temperature Rise in Steel Plate not Engulfed by Fire;

Constant Radiant Heat Flux. 139

50 Temperature in Steel Plate not Engulfed by Fire as

Function of Heat Flux. 140

51 Temperature Rise in Steel Plate Engulfed by Fire. 142

52 Temperature Histories In Concrete Slab at Various Depths;

Radiant Flux - 50 kW/m • 145

53 Temperature Histories in ioncrete Slab at Various Depths;

Radiant Flux - 201.4 kWha (TR - 1373K). 146

54 Temperature Histories in Concrete Slab at Various Depths. 147

55 Surface Temperature of Concrete Slab as Function of

Incident Radiant Heat Flux. 148

56 Temperature in Concrete Slab as Function of Incident

Radiant Heat Flux; Depth - 50 mm from Heated Surface. 150

57 Temperature in Concrete Slab as Function of Incident

Radiant Heat Flux; Depth - 100 mm From Heated Surface. 151

ix
No. Title (Contd) Page No.

58 Temperature Rise with Time in Concrete Wall 0.3 m Thick;

Constant Surface Temperatures = 1373 K, Initial Temperature - 293 K. 152

59 Temperature Variation Across 0.3 m Thick Concrete Wall at

Time t = 1800 s; Constant Surface Temperature - 1372 K,

Initial Temperature = 293 K. 153

60 Temperature Rise with Time in Concrete Wall 0.3 m Thick;

Constant Surface Temperature = 2148 K, Initial Temperature - 293 K. 155

61 Temperature Variation across Concrete Wall 0.3 m Thick at

Time t - 1800 s; Constant Surface Temperature - 2148 K,

Initial Temperature -293 K. 156

62 Temperature Rise with Time in Concrete Wall 0.3 m Thick;

Initial Temperature - 293 K. 157

x
LIST OF TABLES

No. Title

1 List of Variables for Cloud Dynamics Model Including

Code Details

2 Reference Conditions

3 Influence of Wind and Entrainment Coefficient on Critical

Cloud Dimensions

4 Influence of Spill Duration on Critical Cloud Dimensions

5 Influence of Humidity and Molecular Weight on Critical

Cloud Dimensions

6 Influence of External Heating on Critical Cloud

Dimensions

7 Total Thermal Loads Received by Vertical Surface Targets

600 s Spill Duration

8 Summary of LNG Threats

9 Nuclear Power Plants Along Waterways

10 Systems Susceptible to Off-Site LNG Threats

11 Systems Susceptible to On-Site LNG Threats

12 Air Blast Susceptibility Criteria; Structures and

Components

13 Summary of Nuclear Power Plant Vulnerability

Page No.

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33

34

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45

77

107

i1

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114

120

162

xi

1.

Executive Summary

The prime objective of this investigation is to define the hazards to

nuclear power plants arising from large LNG spills on water transportation

routes. The emphasis is on how the spill environment could affect safetyrelated systems which are necessary for a safe shutdown of the plant and for

subsequent heat removal. Motivation for such a study exists because a

sizable number of nuclear power plants are located along navigable

waterways. At the same time, some movement of LNG already exists in the

costal waters, and this traffic may be expected to increase. Further the

shipments of LNG are very large (of the order of 105 m 3) and contain immense

amount of energy. Hence, the consequences of a massive spill could indeed

be catastrophic. A secondary aim of the effort is to provide technical

information which will assist the U.S. NRC in formulating guidelines

concerning the siting and operation of nuclear power plants as they relate

to the threats posed by possible LNG spills on water transportation

routes. To achieve these objectives, the scope of this program consists of

two major tasks: (1) definition of the LNG spill phenomenology and

quantification of the arising threats and (2) assessment of the nuclear

power plant susceptibility and vulnerability to the possible threats.

The study was not intended as a major research effort. Rather a synthesis

and critical evaluation of existing data, methodologies, and predictive

techniques primarily used to obtain estimates of the LNG spill environment

and to define were the vulnerabilities of nuclear power plant systems.

Since the existing information defining the LNG spill and spreading

phenomenology on water was

too disparate and conflicting some developmental work had to be undertaken

in this area and the major portion of the program effort was directed

towards accomplishing this task.

A global approach, which defines only the most salient features of the

phenomenology, but still permits reasonable estimates of the expected

threats, has been used in this study. The study is completely deterministic

and probabilistic estimates of the hazards are not made. In the

vulnerability analysis, simple susceptibility criteria are used for the

various plant systems and components. Most of the blast-response

vulnerabilities are established by comparing overpressure criteria for

various plant systems with the blast environment. Simple calculations are

carried out to define the thermal response. In the accident scenario it is

assumed.that the. entire contents of one LNG shipping container, i.e., 25 x

103 i 3 , are spilled. Ignition is assumed to be possible during all stages of

spill development.
2

The analysis outlined in this report permits an overall description of the

phenomena associated with a large spill of LNG on water. A major

achievement is the development of a combined LNG spill, spreading, and fire

model. This approach permits the simultaneous and coupled computation of

all the major aspects of an LNG fire. A tool is thus available to carry out

detailed analyses of the fire threat for any postulated scenario and

specific power-plant arrangements and sites. Another important result of

the analysis effort is the development of an approach for the prediction of

air blast in the vicinity of an exploding -pancake-like" vapor cloud. While

the method is approximate, the results are scalable with cloud height and

therefore are applicable to a broad class of problems. It is believed that

those predictions are a better approximation of the air blast field in the

vicinity of the vapor cloud than those obtained from TNT equivalency

considerations.

Based on the analyses, it is found that the thermal loads resulting from a

large LNG fire on water can in general be tolerated by the important safetyrelated plant systems. This is primarily due to the short durations of the

expected fires. On the other hand, it is shown that nuclear power plant

systems are very vulnerable to blast effects from an exploding LNG vapor

cloud. It is also concluded that all serious effects are limited to the

immediate vicinity of the fire/explosion source, i.e., the LNG vapor

cloud. Because of this, the standoff distance, i.e., the distance from the

LNG spill location, which is required to ensure the safety of the important

nuclear power plant systems, is primarily dependent on the wind-induced

vapor cloud drift. Under certain atmospheric conditions, i.e., high wind

velocities and air humidities, the cloud may drift a long distance in the

down wind direction before the LFL is reached. Thus to avoid severe

hazards, stand off distances in excess of 10 km are required.

Since nuclear power plants are only affected by the hazards from an LNG

spill when the vapor cloud is directly on the plant site or in its close

vicinity, safety measures may be taken to mitigate at least some of the

hazardous effects. The most obvious is a warning system that will make the

operating personnel aware of the impending danger. This can be followed by

plant isolation, i.e., closing of all dampers, doors, etc., and elimination

of possible ignition sources. The latter requires plant shutdown and the

deenergizing of all unnecessary power sources and electrical equipment.

Isolation of the plant from the exterior should be quite effective against

fire hazards, but will not provide significant protection against air blast

overpressure. Such protection can only be achieved by hardening the vital

plant systems. The most fragile safety-related system at a typical plant
3

site appears to be the off-site power supply. It therefore is a prime

candidate for blast hardening.

The conclusions of the current study are based on a purely deterministic

analysis. Thus, worst-case assumptions are often made, and no consideration

is given to the likelihood of the occurrence of an event. In applications

to actual power plants, probabilistic considerations must be included, in

particular when a preliminary deterministic study indicates significant

hazard levels.

The methodology and results of the current study represent a significant

improvement in the treatment of the hazard problem arising from large LNG

spills on water. The method can be readily used in the analysis of hazards

in specific cases. It is particularly valuable as a rapid and inexpensive

screening procedure when many possible spill scenarios and environmental

conditions must be considered. Coupled with probabilistic analysis, it can

become a valuable tool in assessing the risks to nuclear power plants

arising from large LNG spills on water transportation routes. It should

also be noted that the developed methods can be directly adapted to

investigate the spreading and effects of other hazardous materials.
4

1. INTRODUCTION

In recent years, the effects of off-site hazards have become an important

factor In the siting and design of nuclear power plants. There exists

particular concern over the shipment of large bulk quantities of hazardous

materials, i.e., explosive, toxic, flammable, and corrosive substances, in

the vicinity of the plants. The current investigation concerns specifically

the hazards that may arise from a large spill of liquified natural gas (LNG)

on water-transportation routes. A strong motivation for such a study exists

because a sizable number of nuclear power plants are located along navigable

waterways. At the same time, considerable movement of LNG already exists in

the coastal waters, and this traffic is expected to increase. In at least

one case, the separation between an existing plant and an LNG terminal is

but a few miles. Further shipments of LNG are very large (of the order of

105 mi3 ) and contain immense amounts of energy. Therefore, the consequences

of a massive spill could indeed be catastrophic.

The major objective of the current study is to define the hazards to nuclear

power plants arising from large LNG spills on water transportation routes.

The emphasis is on how the spill environment could affect safety-related

systems which are necessary for a safe shutdown of the plant and for

subsequent heat removal. A secondary aim of the effort is to provide

technical information which will assist the U.S. NRC in formulating

guidelines concerning the siting and operation of nuclear power plants as

they relate to the threats posed by possible LNG spills on water

transportation routes.

To achieve these objectives, the scope of this program consists of two major

tasks: (1) definition of the LNG spill phenomenology and quantification of

the arising threats and (2) assessment of the nuclear power plant

susceptibility and vulnerability to the possible threats.

A number of research investigations, which deal with the phenomena of LNG

spills on water and related areas, have been conducted in the past [1,2] and

are now in progress [3]. Although these studies have produced much

information, great disparities and uncertainties remain in the obtained

results. The purpose of the current program is to reduce some of these

uncertainties. However, originally, the study was not intended as a major

research effort. Rather a synthesis and critical evaluation of existing

data, methodologies, and predictive techniques were planned in order to

obtain reasonable estimates of the LNG spill envitonment. Similarily, in

determining the effects of the threats on the power plant, existing

information concerning the susceptibility of various systems, components,
5

and structures is to be used, modified, or adapted to arrive at plant

vulnerabilities.

Early in the study, however, it was recognized that existing information

defining the LNG spill environment on water is too disparate and conflicting

to permit a synthesis into a meaningful predictive tool. Therefore some

developmental work had to be undertaken in the area of LNG spill

phenomenology, and a large portion of the program effort was directed

towards accomplishing this task.

Because of the limited size of the total effort, it was never the intention

to develop detailed analyses and computational methods which would require

the implementation of large computer codes. Therefore, a global approach,

which defines the most salient features of the phenomenology, but still

permits reasonable estimates of the expected threats, has been used in this

study. Another constraint imposed on the program effort is that a

completely deterministic approach has been taken. This program is not

intended as a risk analysis, and probabilistic estimates of the hazards and

threats are not part of this study. A definition of the problem treated in

this study and the approach taken are given in the next section. This is

followed by sections that summarize the various phenomenological models for

the LNG spill environment as well as the magnitude of threats. A survey of

the nuclear power plant safety systems that may be affected by the spill

environment is then given, and susceptibilities to the various threats are

established. Finally, estimates of plant vulnerabilities are provided, and

the limitations of the results are analyzed and discussed. Additional

details of the various LNG spill phenomenological models and threat

definitions are given in the appendices.
6

2. PROBLEM DEFINITION

The problem considered here is a large spill of LNG from a tanker ship on an

essentially open body of water such as an estuary or wide river. The

quantities of LNG carried by a single ship can be in excess of 105 m3 .

However, regardless of the ship design, the LNG cargo is in general held in

compartments not exceeding 2.5 x 10 4 m 3 in capacity [4]. Since the current

investigation relies on a completely deterministic approach and since the

rupture of a single compartment appears as the most credible shipping

accident, the contents of such a compartment, i.e., 2.5 x 104 m3 , are taken

as the maximum credible spill in this study. Although this may appear to be

an arbitrary limitation, it was felt that no a priori justification exists

for considering the involvement of the remaining tanker compartments and

that such justification could only be developed through detailed

investigations of possible accident scenarios. However, such an effort is

clearly beyond the scope of the current study. Note also that even the

contents of a single compartment contain an enormous amount of energy.

Since LNG has a density of about 450 kg/m3 (28 ibm/ft 3 ), there are

approximately 107 kg or 10 kilotons of fuel in a compartment. Considering

that the energy content of hydrocarbon fuels is approximately tenfold that

of TNT, this indeed represents a large hazard. For completeness, the case

of instantaneous release of the entire cargo of one compartment is included

in this study. However, the primary emphasis is on more credible spill

scenarios in which the cargo Is released over a finite period of time.

Much of the uncertainty in the current estimates of the hazards associated

with a massive LNG spill on water arises from the complex phenomenology of

such a spill. This complexity in turn is largely due to the physical

characteristics of LNG. Being a cryogenic liquid at a very low temperature

of -160*C (-256*F) at a density of about 450 kg/m 3 (28 lbm/ft 3 ) [51, LNG upon

spilling will float on water and spread in an ever increasing pool. This

spreading will be accompanied by violent boiling and vaporization of the LNG

with simultaneous ice crystal formation in the water. Since the vaporized

LNG remains extremely cold and even after considerable expansion and mixing

with air may still be at -73 0 C (-100°F), moisture in the air will condense

and freeze. The density of cold LNG vapor is considerably larger than that

of air. Hence the formed vapor cloud will rapidly spread by gravity. At

the same time, wind forces will cause a global drift of the cloud. Shear

forces and air turbulence, primarily at the upper surface of the essentially

pancake-shaped cloud, will induce air entrainment into the cloud,

simultaneously heating the mixture and changing its density. The cloud may

ultimately reach neutral buoyancy, and further dispersion is then primarily

due to atmospheric turbulence and wind-induced drift.
7

An LNG fuel-air mixture is combustible between the upper flammability limit

(UFL) (15% fuel by volume) and the lower flammability limit (LFL) (5% fuel

by volume). At the boundary of the vapor cloud flammability conditions will

persist at all times until the bulk of the vapor is dispersed below the

LFL. Therefore, ignition of the vapors and/or spreading LNG pool is

possible at any stage of development and various fire scenarios are

possible. These include pool fires, diffusion flame fires in which fuelrich vapors burn around their periphery, and fires of premixed clouds which

may burn very rapidly. In this latter stage of fire development, the

possibilities of cloud explosion or rapid deflagration with the accompanying

production of significant overpressure or blast waves cannot be excluded.

Although no evidence exists that unconfined pure methane-air mixtures will

detonate, it must be remembered that LNG is "spiked" with significant

fractions of higher hydrocarbons [5]. Some experimental evidence that such

mixtures can explode even when unconfined [6]. Partial confinement and/or

turbulence induced by obstacles will only enhance this process [7].

The preceding indicates that all phases of the LNG spill phenomenology are

interrelated; i.e., fuel dispersion, be it in liquid or vapor form, cannot

be separated from the fire phenomenology. None of the past investigations

on LNG spills on water have taken into account this fact. Each of the

phases, i.e., pool spreading, vapor dispersion, and fire, is treated

separately. Thus in addition to the large disparities that exist between

various predictions, there is the uncertainty of the effect of coupling

between fuel dispersion and fire phenomenology. Therefore, a major aim of

the current study is to investigate, at least in a global manner, the

interaction between spill, fuel dispersion, and fire phenomena and to

provide a more realistic quantitative estimate of the threats produced by

the LNG spill environment.

The most prevelant hazard to a nuclear power plant from an LNG spill is due

to the fire threat. As pointed out earlier, ignition can occur during any

stage of fuel dispersal, and the resulting conflagrations will, in general,

cover a large area. Thus significant thermal loads may result at the

nuclear power plant, even when the fire is some distance away, simply

through thermal radiation.

When a flammable vapor cloud actually reaches the plant site, then in a

resulting fire the radiation loads are augumented by direct convective

heating of structures, systems, and components. In case of vapor cloud

explosion or rapid deflagration, the resulting overpressure loads may be

quite high. Since the energies involved are very large, blast durations and
8

hence drag loads may also be significant. Should a portion of the vapor

cloud drift on site and infiltrate some structure volume detonations may

result. Even if ignition does not occur, the LNG vapor in itself may

constitute a hazard to plant personnel through oxygen depletion. Similarly,

the combustion gases constitute a hazard because of their possible toxicity

and oxygen displacement. Finally, for a sufficiently close spill, the LNG

liquid may represent a hazard, either through ingestion into water intake

systems or through its effect on structures, e.g., metal imbrittlement,

because of its extremely low temperature.
9

3. LNG SPILL AND POOL DYNAMICS

The spill and spreading of LNG on a water surface involve many complex

hydrodynamic and thermodynamic phenomena. LNG, being lighter than water

(density about 45% that of water), will spread away from the spill source

driven by buoyant forces and form an ever-increasing pool. The spreading is

influenced by both the characteristics of the spill source and the LNG-water

interactions at the pool boundaries. In addition, the cryogenic LNG

experiences intense heating on contact with the much warmer water and

surrounding air. This results in vaporization of the LNG and is the source

of the LNG cloud formation. The mass loss from the pool due to vaporization

is sufficiently large that it must be taken into account in any realistic

representation of the pool dynamics. A schematic of the LNG pool spreading

as it might occur when a shipping container is breached is shown in Fig.

1. A complete mathematical formulation and solution of the LNG pool

spreading, even under axisymmetric assumption, is very difficult.

Neglecting spill source details, viscous forces, and gradients in the

vertical direction results in a simplified differential formulation of the

gravity spreading for the LNG pool (see Appendix B). To make such a model

tractable, the vaporization is represented as a constant liquid mass loss

rate per unit area. The boundary conditions at the spill source depend on

the actual LNG release mechanisms, and those at the pool's leading edge are

dependent on the LNG-water interaction. The system of equations for the

described model is hyperbolic and can be solved numerically using the method

of characteristics. However, even for such a solution, the many

uncertainties involving the boundary conditions require additional

assumptions. Further, the complexity of the solution makes it difficult to

integrate such a formulation into a global approach for the prediction of

the entire LNG spill phenomenology. A detailed examination of this

formulation under steady-state conditions (see Appendix B) i.e., when the

inflow equals the vaporization mass loss and the pool radius is constant,

reveals some interesting characteristics of this model. Most important of

these is the result that, under steady-state conditions, physically

meaningful solutions are only obtainable when the pool depth decreases with

the pool radius. The steady-state solution can be applied to the late

stages of pool development when the leading edge of the pool is at

considerable distance from the spill source and the behavior is essentially

quasi-static.

To treat the early transient behavior of the pool spreading at a level

commensurate with a global-model approach, it is further assumed that the

pool depth at any instant of time is uniform over the entire pool expanse.

This eliminates the spatial dependence of the pool dynamics. Some
N

N

N

N

N

N

N

N

N

N

N

N

N

N

N

N

N

N

*In

v

r 4~L .*;*;*;~;~*:*:.:.:.:.x.:..J%~..~......... c~r

X.: ý-A

CL) |

I I

i ý r

Figure 1. Model of LNG Pool Spread.
11

experimental evidence exists that this simple model provides reasonable

estimates of the pool size as a function of time [1-4]. The equations of

motion governing the gravity spreading of a circular pool are then simply

reduced to

r -c/ C (1)

and

2 V n 2" 2

Vp - rrh + 2rrhr -i(t) - wr v. (2)

Here the first equation relates the pool spreading rate r to the spatially

uniform pool depth h; the second equation expresses the mass conservation in

the LNG pool, with r the pool radius, Vin the time-dependent volumetric

inflow to the pool, and v the constant volumetric evaporation rate per unit

area. The constant C represents the gravitational effect modified by the

relative density (or a simple multiple thereof), i.e.,

g(Pw - (3)

where Pw is the density of water, PL is the density of the LNG and g is the

acceleration of gravity. The pool will begin to break up when its thickness

reaches a critically small value. This critical value is roughly defined by

experimental data. A major drawback of the spatially uniform pool-thickness

model is that such a breakup would occur instantly over the entire pool

expanse. Similarly, another physically unreal situation arises in a pool of

uniforn depth when the entire pool disappears suddenly due to

vaporization. To avoid these anomalies, the LNG pool description at late

times, when the pool thicknesses are small, is switched from the uniformdepth model to the quasi-steady model described above (see also Appendix

B). Although such a change in modeling is arbitrary and produces some

abrupt changes in pool behavior, it does allow for an orderly pool shrinkage

and receding of the pool leading edge until the pool completley disappears.

The details of the spill source may be expected to influence the dynamics of

the LNG pool spreading. For a fixed total (or maximum) release quantity,

the major variable is the rate of release. Since LNG is shipped essentially

at atmospheric pressure, the release is totally due to gravity flow and is

governed by geometric variables such as tank size and shape, as well as the

cross-sectional area of the rupture and its location relative to the water

surface and tank bottom. Because the LNG contents are not pressurized, it
12

is reasonable to assume that the size of the rupture is a constant during

outflow. Typical results for the outflow from a fully vented tank of

uniform cross-sectional area are shown in Fig. 2, where the rupture is

assumed to be above the water line (for more details, see Appendix A). The

geometric and other primary variables used in defining the dimensionless

parameters of the figure are the following: t-time, V-tank volume, As-area

of opening, H-height of uniform section tank, L-height of opening, 1-height

of fluid above opening, g-gravitaional constant, oLdensity of liquid

(LNG). The outflow velocity U and mass flow rate m are then, respectively,

U - 2-•'g (3a)

and

im = PLA 0f-- (3b)

Finally dimensionless parameters are defined as follows:

= L/H, (4a)

tA 02H (4b)

V

and

C W . (4c)

pLA 2 2-g-H

The initial dimensionless mass flow rate is designated as Go and the

dimensionless release duration as Ef. For gravity outflow from a tank of

uniform cross-section, the height of liquid above the opening I decreases

proportional to t 2 . Hence, the mass outflow decreases linearly with time as

shown in Fig. 2. Also indicated are the dependences of mass flow release

duration on the location of the opening relative to the tank bottom.

When the cross section of the tank is not uniform, some departure from the

above-described behavior may be expected. However, as shown (in Appendix A,

Plates 12 and 13) for the extreme of a spherical tank, a linear decrease of

mass outflow rate with time is a reasonable approximation for much of the

release duration under most conditions of gravity flow. If, the tank

rupture occurs some distance below the water line, the outlfow is further

modified in that the primary release ceases when the outside water height
I .

U -

I I T

r'f G

0 0

0 Tf 2 0 I

Figure 2. Gravity Outflow from Tank of Uniform Cross Section.
14

above the opening balances the LNG height above the opening inside the tank

(see Plates 14 and 15 of Appendix A). At this point, a secondary release

may occur due to the infiltration into the tank of water, which has a higher

density than the LNG. This water will sink to the bottom of the tank,

displacing the LNG above it and causing an additional release. This process

will continue until the water level inside the tank reaches the opening.

For a tank of uniform cross section, again the secondary release mass flow

varies linearly with time. In most circumstances the secondary release is

expected to be a slow process complicated by the rapid evaporation of LNG in

intimate contact with the water. Hence, the overall spill will be dominated

by the primary release due to direct gravity flow from the tank.

The effect of spill rate on LNG pool size for a linearly decaying

(triangular) release is examined in Fig. 3. The influence of the spill rate

is seen to be significant during the early pool development, particulary if

instantaneous release (tf = 0) is considered. These differences are

markedly reduced by the time the pool reaches its final size, i.e.,

disappears due to complete vaporization. The variations in the times at

which maximum pool growth occurs is even less pronounced. The influence of

both rate of release and iorm of release Is illustrated in Fig. 4, by the

example of a decaying triangular and a constant mass outflow. Note that the

release rate, or duration of spill, has a more pronounced influence on the

pool size than the release form. Again, differences are large during the

early stages of pool development.

The vaporization rates and vaporization history of the pool are Important

features for the study of the generation, growth, and spread of the LNG

vapor cloud. As indicated earlier, the model adopted here assumes a

constant vaporization rate. Figure 5 shows the vaporization history for

various spill rates, assuming a vaporization rate of 3 x 1i- 4 m/s. The

effect of spill rate (duration) on the total amount of vapor generated is

seen to be significant. For a spill duration of 500 and 1000 s, the pool is

completely evaporated before the end of the spill. A secondary pool is then

formed, and vaporization continues. This is indicated by the abrupt change

in slope in the curves for these spills. The effect of a higher

vaporization rate, 6 x 10-4 m/s, is illustrated by the dashed line.

Based on the foregoing results and additional data and analysis given in

Appendixes A and B, a two-stage global pool spreading and spill model is

adopted. A linearly decaying spill rate appears to be most plausible. The

transient stage of the pool spreading is determined under the assumption of

spatially uniform average pool thickness, constant vaporization rate, and

constant Froude number F - (r/gI). The model then adjusts to a quasi-
15

1000

100 E

I0

3

I 1O iO0o 1000

TIME , S

Figure 3. Influence of Release Duration on Pool Spread;

Total Volume of Spill - 25,000 m3 .
500 iCOMPLETE I I I

EVAPORATION'

0 /

-- 00 /00 /

-0000 t f 500S

30 - I I I I

10 100 500

TIME, s

Figure 4. Influence of Release Form on Pool Spread;

Total Volume of Spill = 25,000 m3 .

I .
3

X/

'C /

INSTANTANEOUS /

2I

- /

L&J

WC4

o /

> t = I0000

0~

u.4

L- 50- /

0 ,00 4 0

0

//•/ 250

0 200 40O0 6O00

TIME, s

Figure 5. Pool Vaporization Histories; Variation with Spill Duration.

Vaporization Rate: 0.003 m/s, 0.006 m/s.
18

steady solution when the transient pool thickness suddenly tends to zero due

to complete vaporization. This is illustrated in Fig. 6 for a total spill

of 25,000 m 3 over a time duration of 600 s. Also shown is the secondary

pool growth that would result if the dynamic uniform-thickness pool modeling

was assumed valid for the duration of the spill. Although the adopted

global model has some shortcomings, such as the abrupt change in size once

the maximum dynamic radius is reached, it is believed to give a reasonable

representation of a very complex phenomenon and to provide a good basis for

the more complex integrated modeling of the LNG vapor cloud spreading and

fire phenomena.
LU

4i~

C,,

LUJ

-j

LUJ

83.3

0

RELEASE RATE

0 600

-TIME,s

DYNAMIC

'0

400

'< 200

-J

0- 0

0 200 400 600

TIME, s

Figure 6. Pool Dynamics.
20

4. LNG VAPOR-CLOUD DYNAMICS

The problem of a vapor cloud spreading as a result of a spill of cryogenic

liquid has been studied by many investigators, and a number of these efforts

specifically address the dispersion of LNG vapors. The models developed

range from sophisticated three-dimensional computer codes [8] to analytical

global representations of the phenomena [I] and also include direct

experimental correlations [9,10]. The uncertainties and discrepancies of

these past efforts are vividly illustrated in Ref. 11, where for the same

spill scenario the predictions for the distance to the lower flamability

limit (LFL), obtained by the various models, range from approximately 2.5 to

80 km.

Currently, no published nonproprietary model exists that will reliably

predict the size and downwind drift of the LNG cloud resulting from a

massive LNG spill on water. At the same time, little guidance can be

obtained from the available small-scale experimental data. Limited

published results and evaluations of a sophisticated three-dimensional code

SIGMET [8] indicate that the cloud geometry is relatively constant, being a

somewhat elongated cylinder which drifts downwind. At any instant of time,

the primary cloud region apt, ears to be quite uniform in its thermodynamic

state. This information suggests that a global model of LNG vapor cloud

spreading which assumes a cloud of uniform thermodynamic state and simple

shape, and which moves under the influence of wind, may provide a reasonable

approximation to the true phenomena. It is further apparent that spreading

of the cloud mass is primarily due to gravity and that air entrainment plays

an important role [1,8]. From simple thermal-equilibrium considerations, it

can be shown (see Appendix B) that, at the upper flamability limit (UFL, 15%

by volume), the weight ratio of air to LNG vapor is 12, while at the lower

flamability limit (LFL, 5% by volume) this ratio increases to 35. Hence the

properties of the ambient air are important in defining the behavior of the

vapor cloud. In particular, the humidity of the air, because of the high

heat content in the water vapor, may be expected to have significant

influence. Another factor that affects the cloud behavior is the heat

transfer between the vapor-cloud and the water surface. It can also be

shown (Appendix B) that the vapor-cloud spread rate by gravity does not vary

much with the dilution of the vapor cloud by the entrained air. This

insensitivity to the air dilution suggests a similarity in the cloud-growth

histories.

4.1 Vapor Cloud Model

It is assumed that the cloud dynamics can be adequately represented by a
21

global model. The LNG vapor and entrained air are assumed to be intimately

mixed and to be in thermal equilibrium so that the entire cloud expanse is

in a uniform state spatially at any instant of time. The model shown in

Fig. 7 assumes a cylindrical cloud shape of radius R and height H. The

cloud grows laterally due to the action of gravity while retaining it

cylindrical shape and drifts downwind under the action of the wind. LNG

vapor enters the cloud as a result of time-dependent vaporization from the

spreading LNG pool. Air is assumed to be entrained at the upper surface of

the cloud region due to the local relative velocity between the cloud vapors

and the adjacent ambient air and due to atmospheric turbulence. The effects

of air humidity and heat transfer from the water surface may be included.

The cloud size is initially of zero radius and height at the start of the

LNG release process, and it grows simultaneously with the spill process and

the associated LNG pool growth. A gradient wind model is used to describe

the atmospheric wind field. To obtain physically meaningful results at all

times, it is assumed that if the LNG pool grows radially more rapidly than

the corresponding vapor cloud can grow, under the action of gravity, then

the cloud radius is governed by the pool spread rate and pool size, which

control the region over which LNG vapor is injected into the atmosphere.

The air, LNG vapor, and water vapor (if present) are all treated as perfect

gases having a constant average value of specific heat. An energy and mass

balance determines the temperature of the vapor cloud and thus its density

(see Section 3 of Appendix B). For large relative humidities in the ambient

air, the latent heats of the water vapor can have a significant effect on

the heat balances.

The cloud dynamics is governed in part by the gravity-spreading velocity

relationship, which is based upon the density intrusion concept of Yih [12]

(see also Appendix B):

dR g P H (5)

where R is the cloud radius, p is the cloud density, pa is the air density,

t is time, H is cloud height, and g is the gravitational constant. The

factor B may be called a "Froude Multiplier," and sensitivity studies with

the cloud dispersion model suggest a nominal value of B = 2 (see Appendix

B). The LNG vapor production is controlled by the pool dynamics described

in the preceding chapter and is given in differential form as
CLOUD

AIR ,,AI R

I

R I

R H

(WATER)

POOL -

RELEASE POINT

R

Figure 7. Basic Cloud-Dispersion Model.

I I
23

dt (6)

where r is the time-dependent pool radius, v is the vaporization rate per

unit area of LNG liquid, PL is the density of LNG at the saturated liquid

state, and M is the mass of LNG vapor in the cloud.

The air entrainment details are similar to the approach used by Germeles and

Drake [1], i.e., air is entrained due to the local relative motion between

cloud mass and air. Again, in differential form the following equation

applies:

dM'W awR2 P e (7)

dt ae

where pa is the density of ambient air, Ue is the entrainment velocity, M'

is the mass of air in the cloud, and a is the entrainment coefficient. The

last quantity can be interpreted as being made up of two parts: entrainment

due to shear layer effects (between air and vapor cloud) and entrainment due

to atmospheric turbulence. While the specific value of a is an input to the

model, parametric studies and extrapolation of experimental data (see

Appendix B) indicate values of about 0.1. Implicit in Eq. 7 is the neglect

of air entrainment at the vertical (side) surfaces of the cloud-mass

cylinder. In general, however, this area represents less than 5% of the

total area and is therefore neglected.

The entrainment velocity Ue in Eq. 7 is obtained by properly integrating and

averaging the local relative velocities over the entire top surface of the

cloud (see Appendix B). These relative velocities are the vector sum of the

local expansion velocity and the wind velocity. Depending on the wind

velocity, the ratio of the entrainment to the spreading velocity Ue/R is

found to vary from a value of 2/3, when the wind velocity W - 0, to a value

approaching W when the latter is large. To facilitate numerical

calculations an analytic fit of this relationship is formulated (see

Appendix B):

Ue/R - W/R ÷ exp [1.62(W/R)] . (8)

The simultaneous integration of Eqs. 5-7, subject to the overall energy and

mass balance, together with the condition of thermal equilibrium i.e.,

uniform cloud mixture temperature, yields the cloud volume, V, as a function

of time. The cloud radius, R, is directly determined from Eq. 5, and the

cloud height H is computed from the geometry, i.e.,
24

H = V/rR2 . (9)

Figure 7 illustrates the basic cloud-dispersion model; it includes the LNG

pool, which is the source area for the (time-dependent) LNG vapor injection

into the cloud. The pool of radius r spreads dynamically at a rate t, in

accordance with the pool-dynamics model described in the preceding

chapter. Both cloud and pool are of zero height (or depth) and zero radius

at the time of the start of the spill process (t - 0). Thereafter the cloud

and pool grow simultaneously until the pool starts to shrink again because

the vapor formation outstrips the LNG inflow at the source.

The gradient-height model for the wind field used in the computations is

shown in Fig. 8, together with other model details that concern the drift

effects of the cloud-dynamics model. The wind magnitude, W, varies with

height, H, in accordance with the relation

W - W*(H/H*n), 0 H H1*, (10)

where W* is the nominal wind magnitude at the gradient height H*, and n is

the profile exponent. Above the gradient height the wind Is constant at the

nominal value W*. The recommended values for a water surface are [13]: H* =

228.6 m (750 ft) and n = 0.16.

The drift motion of the LNG cloud is in part due to the horizontal momentum

convected into the cloud mass by the entrainment of the moving atmospheric

air. The LNG vapor that enters the cloud has no net horizontal momentum;

hence, its addition to the cloud mass decelerates the moving cloud.

Furthermore, the cloud, which represents a large intrusion into the

atmospheric wind field, will distort the streamline of the surrounding air

(see Fig. 8) and induce an aerodynamic drag force on the cloud system. This

force is here defined in the classical manner, namely

F 1 U(2RH)C (11) d 2 ParH d'

Here Fd is the total drag force, Cd is the drag coefficient, Ur is the

relative velocity between the cloud and the wind, i.e., Ur = W - U, and U is

the center of mass cloud velocity.

In the above equation the wind magnitude at the top of the cloud Is used;

this is a conservative assumption, since it results in the largest drag

force and hence the largest cloud drift. The factor (paUr 2 /2) is the

dynamic pressure and 2RH is the projected area in the direction of wind
I

HEIGHT STREAMLINE

/- CLOUD

1-n

WIND PROFILE POOL

RELEASE POINT

0

Figure 8. Drift Effects on Cloud-Dispersion Model.
26

motion. A momentum balance yields the change in cloud velocity, which is

given in differential form as

Fddt + f(W - U) dM' - UdM

dU = , (12)

M' + R

where Fddt Is the impulse due to the aerodynamic drag force f(W - U)dM' is

the momentum increment due to the entrainment of an incremental amount of

air (dM'), and UdM represents the momentum that must be imparted to the

injected Incremental amount of LNG vapor (dM). The denominator of Eq. 12

represents the current total mass of the cloud system. A momentum factor,

f, is used to allow for some momentum loss due to internal resistive

mechanisms such as mixing. While this is an input parameter, a value close

to unity is usually used. On the basis of the cylindrical shape of the

cloud, a nominal value of 0.3 is assigned to the drag coefficient. At

present, no ground-surface shear resistance is included in the cloud

model. Finally, the displacement x (in the wind direction) is computed from

its differential form definition

dxd-x= U. (13)

Initially, both x = 0 (the spill release location) and U = 0.

Extensive computational studies were performed to determine the sensitivity

of the LNG vapor-cloud dynamics model to the expected range of parameter

values and input variables (Section 3 of Appendix B and Section 4 of

Appendix C). Using what are considered to be the most typical values of the

entrainment coefficient (0.10), wind velocity (4.48 m/s = 10 mph), spill

duration (600 s), and the standard spill quantity at 25,000 m3 we obtain the

pool and upwind and downwind cloud trajectories shown in Fig. 9. The result

is for dry air and no heat transfer from the water surface. As indicated

earlier, the sudden change in pool size at 400 s occurs because the pool

description at this point is changed from a spatially uniform-depth dynamic

model to a variable-depth quasi-static model. Although this is not

completely satisfactory, it does result in the maximum possible pool size

and also permits an orderly shrinking of the pool. The most noteworthy

result of this computation is that the vapor cloud reaches its LFL in a

relatively short time (500 s) and at the moderate downwind distance of 2.8

km. As will be seen later in the result summary these are indeed

representative values. Figure 10 is graphic representation of the LFL

contours for the same standard spill parameters but different wind speeds.

For all but the highest wind speed (8.96 m/s - 20 mph), there is overlap

between the pool and cloud at the LFL. In the latter case, some trailing

vapor configuration reaching back to the pool should be expected.
27

3

2

E

.Ac

I

0

-I

0 200 400 600 800

TIMES

FIgure 9. Cloud and Pool Boundary Trajectories.
+2

WIND

"DIRECTION

SILa,

z

0

p-

(n SPILL

:•LO CAT 10N 0.

z

-2

-2 0 2

DISTANCE FROM SPILL LOCATION, x DIRECTION, km

4

Figure 10. Plan View of LFL Contours.
29

Since no experimental data exist on large-scale LNG spills on water, the

global model developed cannot be validated here. Some limited comparisons

with other models were attempted however, (see Section 3 of Appendix B).

Figure 11 compares the downwind distance to the LFL for various spill

quantities, between the current model, SIGMET-code results [8], and the

Germeles and Drake model [1]. Two values of the entrainment coefficient a

[0.10 and 0.05] are used for the current model. An instantaneous spill,

zero humidity, and a wind speed of 2.24 m/s are assumed. Although the

results of all models are reasonable close at moderate spill quantities, the

Germeles-Drake model seems to diverge sharply from the other results at the

large spill quantities of interest in the current effort. Although this

favorable comparison of the present model, with the elaborate computer

calculations using the SIGMET code, can by no means be construed as a

validation of the vapor-cloud model, it provides some confidence and

verification that the obtained predictions are not unreasonable.

The major difference between the current formulation and other global models

is the incorporation of global drift of the entire cloud mass. Models that

neglect this aspect clearly violate momentum consideration. Numerous

parameter studies indicate that, under most conditions, the vapor cloud

remains negatively buoyant until the LFL is reached. Dispersion by

atmospheric turbulence and the subsequent formation of vapor plumes are

expected to play a minor role in the spreading of the LNG vapor in its

flammable range. The current formulation therefore does not incorporate any

of the Gaussian plume models so frequently advocated for the modeling of

cryogenic vapor dispersion [14]. As shown in the sensitivity studies

(Section 3 of Appendix B and Section 4 of Appendix C) fuel dispersion to

distances larger than those indicated by gravity spreading can be readily

accounted for by the global drift of the cloud mass. For predominantly

negatively buoyant clouds, the latter dispersion mechanism is much more

credible than plume dispersion.

4.2 Result Summary of LNG Vapor-Cloud Spreading

The extensive sensitivity studies carried out [see Appendixes B and C),

together with data from the literature [1,3,8,12], provide a basis for

standardizing certain of the problem parameters. Although, the selection of

the appropriate values involves some engineering judgement, the impact on

the overall behavior of the variables that are fixed in the current

formulation is in general found to be minor. These values are listed in

Table 1; also shown are the variables that must be provided as input to the

model and the output that may be expected from a typical computation. Other

variables that are fixed for the current application are the total spill
30

INSTANTANEOUS SPILLS W = 2.24 m/s

6

5

2

-J

&-J

0

,I

C:)

I--

C,,

Cl

z

0

0

4

3

2

I

0

SPILL VOLUME, m3

Figure 11. Comparison of Maximum Downwind LFL Distances.
31

TABLE 1

LIST OF VARIABLES FOR CLOUD DYNAMICS MODEL

INCLUDING CODE DETAILS

PARAMETERS:

Froude multiplier for pool

Froude multiplier for cloud

Pool breakup depth

Molecular weight of methane

Molecular weight of air

Molecular weight of water

Density of water

Density of liquid LNG

Temperature of saturated LNG

Specific heat for air

Specific heat for LNG vapor

Specific heat of water vapor

Specific heat of liquid water

Atmospheric pressure

Wind gradient height

Exponent for wind profile

Momentum factor

Drag coefficient

2

2

.0018 m

16.0

28.967

18.0

1.0 g/cm3

.4487 g/cm3

111.7 K

.24 cal/g K

.48 cal/g K

.48 cal/g K

1.0 cal/g K

.101325 MPa

228.6 m

.16

.90

.30

INPUT VARIABLES:

Peak release rate (VOD)

Release duration (TO)

Temperature of air (TEMPA)

Temperature of surface (TEMPS)

Relative humidity (RHUM)

Molecular weight factor (FMOLWT)

Entrainment factor (ALFA)

Wind speed at gradient height (WIND)

Vaporization rate (VAPR)

r s

MS

OC

z

m/s

rn/s

OUTPUT VARIABLES:

Time

Cloud radius

Cloud height

Pool radius

Pool depth

Cloud spread rate

Concentation (by volume)

Cloud velocity

Cloud drift

(Plus many other variables such as:

Cloud Temperature, Energy, Density, etc)

s

m

m

m

m

M/s

m/s

M.
32

volume (25,000 m3 ) and the form of the LNG release, i.e., linearly decaying

outflow. Thus the initial maximum spill rate is uniquely determined by the

total spill volume and the spill duration.

Among the input variables the influence of ambient temperature, at least for

dry air, is found to be minor. Although the effects of some of the other

variables may be significant, e.g., vaporization rate, there appears to be

sufficient experimental evidence to choose values that are realistic. Thus

these parameters were not included in the sensitivity studies. To provide a

basis for comparison and to evaluate the influence of various parameters,

standard reference conditions were chosen; these are given in Table 2. Also

listed in Table 2 are the variables evaluated in the sensitivity studies.

The molecular weight factor is used to model, in a approximate way, the

influence of LNG composition.

The results of the parameter studies are primarily time histories of such

variables as volumetric concentration, cloud radius and height, spreading

rate, drift velocity, and drift distance. They are summarized here in a

series of tables and are presented in detailed graphs in Appendixes B and

C. SI units are used in the tables; thus, all lengths are in meters, time

in seconds, and velocities in m/s. The tables present the physical sizes of

the vapor cloud at the UFL and LFL as well as the drift distance of the

cloud center of mass. To obtain the total downwind range (distance) of the

cloud, the cloud radius must be added to the drift distance. Variables not

specified in the tables take on the values specified above as reference

conditions.

Table 3 summarizes the effect of entrainment coefficient and wind speed. As

may be expected, the entrainment coefficient, a, influences primarily the

cloud size, i.e., its height and radius, particularly at the LFL. A similar

effect is noted (in Appendix B) in the variation of cloud LNG concentration

with time. Although a lower value of a appears to give a more conservative

answer (larger cloud dimensions), the most realistic experimental estimate

of this parameter is around a value of 0.10 (see Appendix B and Ref. I).

When, as is proper, drift is included in the model,the wind speed is seen to

have essentially no effect on the cloud dimensions, which are solely

goverened by the gravity spreading and air entrainment. The major influence

of wind speed is on the drift velocity and cloud travel distance. The drift

velocity reaches a value of about two-thirds of the wind speed at the time

the LFL is reached. Similarly, the drift distance is proportional to wind

speed. This is illustrated in Fig. 12, which shows both the drift and the

total downwind distance as a function of wind speed.
33

TABLE 2

REFERENCE CONDITIONS AND VARIABLES FOR LNG SPREADING

REFERENCE CONDITIONS:

Total spill volume

Spill duration

Entrainment coefficient

Vaporization rate

Molecular weight factor

Wind velocity

Ambient temperature

Surface temperature

Relative humidity

VARIABLE PARAMETERS:

Entrainment factor

Spill duration

Wind speed

Relative humidity

External heating

Molecular weight factor

25,000 m3

600 s

0.10

0.0003048 m/s

1.00 (Pure methane)

4.48 m/s

21.1 °C

(Assume no external heating)

0.0
34

TABLE 3

INFLUENCE OF WIND AND ENTRAINMENT COEFFICIENT

ON CRITICAL CLOUD DIMENSIONS

25,000 m3 SPILL

ALFA WIND m/s RADIUS, m HEIGHT, m DISTANCE, m

UFL LFL UFL LFL UFL LFL

** WITH DRIFT **

0.10

0.10

0.10

0.10

0.05

0.05

0.05

0

2.24

4.48

8.96

0

2.24

4.48

664

666

657

622

1646

1644

1645

1643

2099

2103

2095

16.0

16.0

16.0

15.9

16.0

16.0

16.0

37.9

37.7

37.9

38.7

24.1

24.1

24.1

0

207

412

801

0

383

772

0

609

1225

2501

0

747

1497

1339

1338

1331

** WITHOUT DRIFT **

0.10

0.10

0.10

2.24

4.48

5.15

570 1545

397 1353

352 1251

15.6

14.7

14.3

41.6

50.8

52.0

CONCENTRATION: UFL = 15%

LFL = 5%
4

j -.. • j

IDRIFT

C1

-n

=0

0 2 4 6 8 1O

WIND SPEED, m/s

Figure 12. Influence of Wind Speed on Downwind Transport.
36

The dynamic behavior of a typical cloud variable is illustrated in Fig. 13,

which gives the variation of volumetric LNG concentration as a function of

time for various wind speeds. Initially, the concentration is 100%. It

then decreases rapidly, since the cloud spreading, at early times, is

controlled by the pool spreading. As the pool spread rate decreases,

gravity spreading of the cloud dominates and the concentration increases

again until a maximum of about 50% is reached. Thereafter the dilution

(i.e., air entrainment) process dominates, and the concentration decreases

until the flamability limits are reached. Ultimately, the concentration

reaches a safe diluted state. The LFL is reached in about 500 s for the

reference spill duration. This state is indicated by the dot in Fig 13. It

is also clear that the concentration history is not significantly influenced

by the wind speed. A similar (three-phase) time history is indicated for

the other dynamic variables such as the cloud spreading velocity. This and

other detailed results are presented in Appendix C.

The effect of spill duration on the cloud dimensions and drift is summarized

in Table 4. The results are for the reference wind speed of 4.48 m/s and

dry air. All the dimensions appear to attain their peak values in the

vicinity of the reference duration, i.e., 600 s. The same holds for all the

dynamic variables. This also implies that the time to reach the LFL is the

longest for this case. For more rapid spills, including the instantaneous

spill, the decrease from the peak values is quite small. However, for long

spill durations, the decrease from the peak values is quite significant, so

that for spill durations of the order of 10 hours the cloud dimensions are

reduced by more than a factor of four. This is illustrated in Fig. 14 for

the cloud radius and the total downwind travel.

The influence of relative humidity on the cloud dynamics is investigated

together with the effects produced by variations in the average molecular

weight of the LNG, i.e., by the presence of some fraction of hydrocarbons

other than methane. As indicated earlier, this fraction is controlled in

the model simply by a molecular weight factor; all other properties of the

LNG are assumed to be the same as those of pure methane. The results are

summarized in Table 5. A careful examination of this data indicates that

the effects of humidity in changing the cloud dimensions at the LFL and the

drift are opposed by the increase of molecular weight.

To understand this phenomenon, we must examine the influence of water vapor

alone on the cloud behavior. This influence is complex, because the water

in contact with the cryogenic LNG can exist in all three phases (solid,

liquid, and vapor). The primary effect of the water is to alter the

thermodynamic state of the mixture, and although the quantity of water may
o

0

'4

0

4.I

U

C.,

0

U

0

o 

o

*0

(.0

0

I.'

'-I

0

uJ p.

I-. 0

0 .00

o p.

U

41

0

*6.4

0

*4-a

CJ

0

0

I

I-I

.p4

(.0

0

0 0 0 0 0 to e~J

1N33~I3d 'NOI1V~I1N33NO3
38

TABLE 4

INFLUENCE OF SPILL DURATION ON

CRITICAL CLOUD DIMENSIONS

25,000 m3 SPILL

DURATIONS, s RADIUS, m HEIGHT, m DISTANCE, m

.06

.6

6

60

120

360

600

1200

1800

3600

5400

6000

7200

18000

36000

60000

UFL

836

850

891

901

906

751

678

540

482

388

344

339

315

234

185

156

LFL

1524

1562

1591

1637

1649

1667

1648

1435

1260

980

826

798

718

514

390

316

UFL

20.3

20.6

21.0

21.3

21.8

18.2

16.1

13.1

11.8

9.5

8.5

8.1

7.7

5.7

4.6

3.9

LFL

35.1

36.0

36.7

37.7

38.0

38.4

37.5

33.1

29.1

22.7

19.1

18.7

16.7

12.0

9.2

7.5

UFL

439

443

460

473

505

452

421

361

336

289

267

264

252

205

174

153

LFL

999

1023

1069

1087

1092

1181

1216

1156

1059

887

775

785

689

552

455

392

CONCENTRATION: UFL = 15%

LFL = 5%
4

E

-3

-TOTAL DRIFT

U, 0

ecn

0 1

0.1 I 10 o00 1000

SPILL DURATION to (min)

Figure 14. Influence of Spill Duration on Cloud Radius and Drift;

Wind Speed - 4.48 m/s.
49

TABLE 5

INFLUENCE OF HUMIDITY AND MOLECULAR WEIGHT

ON CRITICAL CLOUD DIMENSIONS

25,000 m3 SPILL

RELATIVE

HUMIDITY, %

MOL. WEIGHT

FACTOR RADIUS, m HEIGHT, m

UFL LFL

DISTANCE, m

UFL UFL LFL

0

25

50

60

70

75

80

85

90

100

0

50

70

75

80

90

100

0

50

75

100

1.0

1.0

1.0

1.0

1.0

1.0

1.0

1.0

1.0

1.0

1.1

1.1

1.1

1.1

1.1

1.1

1.1

1.2

1.2

1.2

1.2

678

803

1076

1089

1076

984

910

863

827

773

508

723

900

959

1018

1038

1043

408

530

650

820

1648

1653

1657

1652

1627

1420

1144

926

897

855

1575

1598

1607

1606

1603

1588

1147

1490

1533

1556

1558

16.1

19.3

26.6

27.6

29.0

35.1

41.4

45.8

50.1

57.9

11.8

17.5

22.3

23.9

25.5

26.5

27.4

9.3

12.6

15.9

20.5

37.5

37.5

38.4

38.5

41.0

52.1

80.7

122.8

130.6

144.6

36.0

36.5

36.9

37.0

37.3

37.9

72.6

33.4

34.8

35.6

36.2

421

546

907

1024

1222

1476

1638

1741

1839

2026

282

486

691

767

855

943

1052

209

314

430

620

1216

1247

1576

1829

2690

4112

5772

6752

7195

8159

i111

1262

1418

1569

1707

2193

4150

953

1107

1281

1633

CONCENTRATION: UFL = 15%

LFL = 5%
41

be small, the effect may be significant because the latent heat of

vaporization is very large. For pure methane and high relative humidities,

it is even possible to obtain mixture densities that are slightly smaller

than that for air. Thus a stall condition in the cloud spreading may be

achieved. This stall process may be only temporary, since the wind effect

(if present) will allow the dilution process to continue this dilution,

together with the associated thermodynamic effects, may result in sufficient

increase in mixture density to permit the spreading by gravity to

continue. This process is best understood by examining the radius-time

histories for pure methane LNG at various relative humidities, as shown in

Fig. 15. For relative humidities up to 50%, the vapor-cloud spreading is

little affected. However, for relative humidities in excess of 70%, the

above-described stall condition occurs and the cloud radius remains

essentially unchanged (spreading velocity is zero). For a relative humidity

of 75%, this is seen to be a temporary condition and thereafter the cloud

radius increases again rapidly. For 100% humidity, the stall situation

persists and the radius remains essentially constant until the LFL condition

(dot) is reached at about 7200 s. Significant cloud height growth will

occur during stall due to continued entrainment of air at a constant cloudradius condition. If subsequent spreading occurs, the increased cloud

height may drive the cloud to large velocities and some cloud height

reduction may actually occur. Finally, the cloud drift distance at the LFL

increases, simply (see Table 5) because a longer period of time is needed to

achieve this condition.

The cloud radius history for a heavier LNG composition (1.2) is shown in

Fig. 16. A much more conventional behavior is observed, and although some

differences do occur over the full humidity range, they are significantly

reduced as compared to pure methane. Clearly, the mixture density is always

greater than the ambient air density. Figure 17 shows the total downwind

distance (drift plus radius) as a function of relative humidity for three

gas compositions, i.e., pure methane and LNG which is 10% and 20% heavier

than methane. The influence of the relative humidity on this critical

transport distance is dramatic for pure methane, yielding a downwind

distance about three times as large as for dry air. For the intermediate

gas composition, stall conditions are just developing when the humidity

reaches 90-100%. This causes the downwind distance to be doubled when

compared with that for dry air.

The effects of heat transfer from the water surface in combination with

other conditions, specifically relative humidity, LNG composition, and spill

duration are summarized in Table 6. In general, the influence of heating is

insignificant. The only observable effect is on the onset of stall; i.e.,
3000

2500

2000

Cl, 0

o75

1500 J7

o25 0/

1000 - LFL

100%

500

0 I I

0 500 1000 1500 2000 2500 3000

TIME, s

Figure 15. Effect of Relative Humidity on Vapor Cloud Radius for Pure Methane; Wind

Speed - 4.48 m/s.
3000 I

2500

2000

-50

1500

00

. 1000 -75 _

5000

0

.0 200 400 600 800 1000

TIME, S

Figure 16. Effect of Relative Humidity on Vapor Cloud Radius for LNG (1.2); Wind

Speed - 4.48 m/s.
8

E

-J

LL~

--J

0

6

4

2

0

0 20 40 60 80 100

RELATIVE HUMIDITY,%

Figure 17. Influence of Relative Humidity and LNG Composition on Downwind Distance

to LFL; Wind Speed - 4.48 m/s.
45

TABLE 6

INFLUENCE OF EXTERNAL HEATING ON

CRITICAL CLOUD DIMENSIONS

25,000 CU M SPILL

RELATIVE

HUMIDITY, %

MOL. WEIGHT

FACTOR RADIUS, m

UFL LFL

HEIGHT, m

UFL LFL

DISTANCE, m

UFL LFL

** WITH EXTERNAL HEATING **

0

50

75

100

0

50

75

100

1.0

1.0

1.0

1.0

1.2

1.2

1.2

1.2

698

1083

923

752

412

541

660

834

** WITHOUT EXTERNAL HEATING **

1651

1656

1225

841

1509

1534

1556

1557

1648

1657

1420

855

1490

1533

1556

1558

16.6

26.9

40.1

61.0

9.4

12.9

16.2

20.9

16.1

26.6

35.1

57.9

9.3

12.6

15.9

20.5

38.1

38.4

69.8

149.6

33.6

35.1

35.6

36.4

37.5

38.4

52.1

144.6

33.4

34.8

35.6

36.2

442

936

1601

2099

216

342

441

637

421

907

1476

2026

209

314

430

620

1287

1625

5362

8498

949

1129

1292

1692

1216

1576

4112

8159

953

1107

1281

1633

0

50

75

100

0

50

75

100

1.0

1.0

1.0

1.0

1.2

1.2

1.2

1.2

678

1076

984

773

408

530

650

820

SPILL DURATION, s RADIUS, m

UFL LFL

HEIGHT, m

UFL LFL

DISTANCE, m

UFL LFL

** WITH EXTERNAL HEATING **

6

60

600

6000

** WITHOUT EXTERNAL HEATING **

6

60

600

6000

899

915

698

352

891

901

678

339

1590

1636

1651

808

1591

1637

1648

798

21.3

21.6

16.6

8.5

21.0

21.3

16.1

8.1

36.8

37.7

38.1

18.9

36.7

37.7

37.5

18.7

470

487

442

288

460

473

421

264

1092

1120

1237

822

1069

1087

1216

785

CONCENTRATION: UFL = 15%

LFL - 5%
46

stall occurs at a slightly lower relative humidity, and hence the cloud

drift is somewhat increased.
47

5. LNG FIRES

The fire that may develop following a massive spill of LNG on water will

depend on a number of factors. Foremost of these is the type, location, and

timing of the ignition source. Others are the gas composition, cloud

homogeneity, cloud turbulence, and possible confinement. Considering only

the fire of unconfined, combustible LNG vapor clouds, a recent review of the

subject [15] states that the ability to predict the thermal radiation for

such a cloud has not been demonstrated. However, other investigators (see

section 4 of Appendix B) have shown that, aside from a direct LNG pool fire,

the specific fire scenarios that may occur depend primarily on the degree of

premixing of fuel vapor and air. Thus, for the fuel-rich mixtures,

turbulent diffusion flames may be expected; a premixed cloud, i.e., one with

concentrations below the UFL, has a tendency to burn as a fireball or a

rapidly moving flame sheet. Using these facts, a simplified LNG global fire

model is synthesized that is compatible with the global treatment of LNG

pool and vapor-cloud spreading. This model essentially addresses the ,regime

of relatively low flame speeds when the primary hazard is the thermal flux.

5.1 Fire-Model Description

The integrated fire model is based on the current state of the akt and

treats three separate fire regimes: pool fires; diffusion flames

propagating through a fuel-rich vapor cloud, i.e., a cloud having an average

concentration above the UFL; and a reaction zone propagating through a

premixed cloud, i.e., within the flammability limits.

5.1.1 Pool Fires

The literature [16] identifies three techniques for modeling the heat

radiation from LNG fires: (1) rigorous solutions of the gas-radiation

problem, (2) point-source models, and (3) emitting surface models. The last

approach is found to be most appropriate for the current application (see

Appendix B), providing a reasonable balance between model simplicity and

accuracy.

The emitting surface model considers the flame to be a solid body of simple

geometry, e.g., a cylinder or a cone, which emits radiant energy from its

surface as a gray body with an effective emissivity Cf given by:

1f "-aD (14)

where a is the absorption coefficient for the flame and D is the flame
48

diameter. For large diameters, the emissive power Ef is effectively that of

a black body. The radiant energy flux leaving the flame is then given by;

ClEf and the energy flux reaching the target is given as

- TFef Ef (15)

where T is the atmospheric transmissivity and F is the configuration factor.

The emitting surface model is illustrated in Fig. 18, and the configuration

factor is defined as

FI F Cs t dA• (16)

- WrS2

Here the subscripts s and t refer to the source and target, respectively,

and 0 is the angle between the ray connecting the incremental source surface

element with the target and the appropriate surface normal. The length of

the ray is S, and A. is the surface area of the source. The configuration

factor thus defines the flame geometry (height, diameter, and tilt) and the

distance and orientation of the target surface. Note that the latter is

really a surface element of the total target geometry.

The emitting surface model captures the essential features of the pool fire

problem and yet maintains adequate simplicity. Expressions for each of the

terms in Eq. 15 are identified either from the literature or from simple

physical considerations where the literature is incomplete. Thus, based on

the best data available, a value of Ef- 210 kW/m 2 is selected for the

optically thick emissive power, and the flame-absorption coefficient a

0.16 m-1 is used (see Section 4.2 of Appendix B for details).

The configuration factor F, needed for computing the heat flux from a pool

fire to an arbitrarily oriented surface element of a target, can be

rigorously defined by integrating Eq. 16. This integration can in most

cases be carried out only numerically, and some computer programs have been

developed for simple shapes (see Appendix B). Two options have been

provided for the current modeling of pool fires. The incremental target

surfaces are assumed to lie in either the vertical or horizontal planes and

at specified distances and elevations from the pool center (but not engulfed

by the flame). In the first option, a numerical integration of Eq. 16 is

performed over a fire surface represented by a number of stacked disks of

equal diameter which are staggered along the flame-tilt axis to account for

wind drag. This is illustrated in Fig. 19a. This approach is very accurate

and should be used for targets situated near the fire column. The other

option uses simple analytical expressions for a flat rectangular radiating

plane which may also be tilted from the vertical. The geometry is shown in
3-

Figure 18. Emitting-Surface Fire Model.
50

TILT AXIS

A4 ELEVATION

(a) NUMERICAL

TILT AXIS

j ELEVATION

(b) ANALYTICAL

Figure 19. Configuration-Factor Models.
51

Fig. 19b and is generally accurate for large relative distances to the

target. The latter must be downwind when wind tilt is used.

For circular pool fires, which is the assumption in the current model, the

flame diameter is merely the LNG pool diameter. To define the flame height,

it is necessary to rely on experimental correlations (see Appendix B). The

most appropriate for the current application is given by Thomas [17], who

identified the dimensionless parameters for correlating experimental flame

length data. Thus a dimensionless vaporization rate u is defined as

; m (17)

where m is the mass vaporization rate per unit area, pa is the ambient air

density, g is the gravitational acceleration, and D is the flame diameter.

The flame length to diameter ratio L/D is then given as

L/D - 2610.40 0.1 < U,

L/D - 42P0.61, 0.007 < P < 0.1, (18)

L/D - 561067, U < 0.007.

To account for the effect of wind on plume length, the following expression

is used:

L/D -70p0.86 (W2)-0. (19)

gD--

where W2 /gD is the Froude number and W is the wind speed.

The use of the above corrolations requires the knowledge of the vaporization

rate. As indicated earlier, the best experimental data currently available

[18] suggests as the liquid regression rate for a burning pool a value four

times as great as for nonburning pools, i.e., 0.0012 m/s. This regression

rate defines the vaporization rate for areas of the pool engulfed by the

fire. It corresponds to a heat flux of about 210 kW/m 2 , which is the same

as the emissive power used for the pool fire flame.

To define the flame-tilt angle 0 , experimental correlations must again be

used [19]; they are

tn0 0.07 _2 0.0 8 P -0.6

t -. 3.3 (2--) (W-) (-A) (20)

a gD

where va is kinematic viscosity of ambient air and pg is the density of the
52

cold vapor. Flame trailing, i.e., increased fire-base width, is neglected

in the current model (see Appendix B). The final parameter needed in the

pool-fire model, namely atmospheric transmissivity T, is a function of

specific atmospheric conditions. Based on some studies (see Appendix B),

the effect is not thought to be very strong. Thus, a representative

transmissivity-versus-distance relation is used for an atmosphere of 50%

relative humidity and 21* C, namely,

T = 1.0 - 0.16 log1 0 S, (21)

where S is the distance from the emitting surface to the target given in

meters.

5.1.2 Diffusion Flames

The second fire regime considered in the current application is a diffusioncontrolled flame for the fuel-rich vapor clouds. The wall-fire model of Raj

[20] was adapted for this purpose. The flame is assumed to propagate at the

turbulent flame velocity. The relationship between flame height Hf and

flame width Wf can be obtained from experimental data [21] . The flame

width Wf can then be calculated, at least for pure fuel-vapor clouds and a

known vapor cloud height H, from the expression [21]

U2

W f = 20 [Uf ( )wEwap) 1/3 (2

H-- fgH" a (1 )(22)

Ah -1

with w -[l + c -I- 

&C T

aa

Here Uf is the turbulent flame velocity, g is the gravitational

acceleration, p is the density of the vaporPa is the ambient air density, E

is the stoichiometric ratio, Ahc is the heat of combustion, Ca is the

specific heat of the ambient air, and Ta is the ambient temperature.

In the present fire modeling, Eq. 22 is used to estimate the maximum

attainable flame width (or optical thickness) for diffusion flames in the

vapor cloud and all concentrations above the UFL. The ratio Hf/Wf is

treated as an empirical parameter with a nominal value of 0.5. Similarly,

the nominal value of w is 0.1 [21]. The diffusion-flame velocity Uf is
53

defined relative to the unburned gas and is also treated as an empirical

parameter with values ranging from 1 to 20 m/s. The value of Wf calculated

from the equation is compared to a value characteristic of the cloud volume

consumed and cloud extent, and the minimum of these is used. The flame

height is then computed from the assumed value of Hf /Wf, but is not allowed

to be lower than the cloud height. In this way the flame geometry takes

account of the finite dimensions of the cloud when most of the vapor is

burned-off. It is tacitly assumed that the target is exposed to the maximum

emitting surface area presented by the diffusion flame; therefore emissivity

is based upon Wf.

5.1.3 Premixed Vapor Flames

Based on observations from a test series [22] with flammable mixtures of

hydrocarbon fuels, the simple flame propagation model illustrated in Fig. 20

is developed. As shown, the flame thickness is xo and the face area is

Ao. The flame is moving at a specified velocity Uf, and a linear

temperature profile across the flame is assumed. The model is based on the

premise that, for steady-state flame propagation (no acceleration), the

energy release must equal the total energy loss. For a flame velocity Uf

and combustion efficiency n, the total energy release is

4total (23)

A PoUfXfhc 0 xo

0

where po is the ambient density of the fuel-air mixture ahead of the flame,

Xf is the ratio of fuel mass to total mass of gas in the cloud, and Ahc is

the heat of combustion.

The total energy loss is taken as the radiant loss from the six faces of the

rectangular flame slab. This is believed to be the dominant loss; although

convective losses are also present, they are considered less Important.

While the heat loss can be obtained for any temperature profile (see

Appendix B), the model is further simplified by assuming a uniform

temperature profile T across the flame. The heat loss then becomes

Qtotal 2T 4 [1 - e 0 +-. (l - e -ay) + (1-e 0)] ,(24) A-0 YO 0 X0 -

yoz0

where xo, yo, and zo are the dimensions of the flame slab (see Fig. 20),

is the Stefan-Boltzmanu constant, and a is the attenuation coefficient,

which has a nominal value of 0.07 m- 1 [22].

The turbulent flame velocity Uf relative to the unburned gas is of the order

of 5.8 m/s [22] and is arrived at (see Appendix B) by the relationship
54

!-as - X0 1--i

T1

YO x0

Figure 20. Model for Flame Propagating Through a Premixed Cloud.
55

uf Pa U (25)

fPf L

Hear UL is the laminar flame speed (about 0.37 m/s for a 10 vol % mixture of

methane in air); * is the empirical turbulence enhancement factor, which has

a value of about 2.1, and pa/Pf represents the thermal expansion of the

gases behind the flame front (value about 7.63). In the actual

computational model, the flame-propagation velocity Uf is treated as a basic

empirical parameter, which is varied. Similarly, the combustion efficiency

n is treated as an empirical parameter with values between 0.4 and 1.0. In

all computations, the stoichiometric fuel concentration for methane, i.e.,

10 vol %, is used, which gives a mass ratio of Xf - 0.058. The flame

thickness will vary with the flame velocity and flame temperature, and the

maximum allowable thicknesses is defined by the iterative solution of eqs.

23 and 24 for xo. Typical thickness at Uf - 1 m/s and Tf - 2148 K are 20 m;

the width can be as much as 140 m for colder flame zones.

5.2 Integrated LNG Dispersion and Fire Model

It was pointed out earlier that existing models for the estimation of LNG

accident consequences are only suited for gross predictions of the possible

fire hazard. These models also take no account of the coupling between the

dynamic spreading and dispersion of the LNG and the fire phenomenology. A

major aim of the current effort has been to overcome some of these

shortcomings, and in particular to couple the LNG dynamics and fire

phenomena. In the simple global model developed here the coupling is not

complete, e.g., while the effect of fire on LNG vaporization is taken into

account the effect on cloud buoyancy is not included. Similarly the vapor

cloud spreading model takes no account of the asymmetry produced by the

propagating flame boundary.

The present combined LNG dynamics and fire modeling assumes that the vapor

cloud can, at any stage of its spreading, encounter an ignition source.

Accordingly, ignition is assumed to occur along either the upwind or

downwind cloud-boundary trajectory. Ignition at the instant of spill is

also accounted for an leads to a direct pool fire.

The flame is modeled as propagating through all or a portion of the vapor

cloud as either a diffusion or premixed flame at selected values of

consumption speed (flame speed relative to the unburnt vapor). The flame is

convected by the local cloud spreading and drift and therefore may or may

not ignite the pool. The dynamics of the cloud spreading is modified

globally to adjust for the effect of cloud consumption. Also, dynamic
56

influences of the fire upon the cloud are accounted for globally through the

estimation of the flame consumption or relative flame speeds. Fire-scenario

characteristics include:

o Pool fire only; immediate ignition condition.

o Diffusion to premixed flame transition in the

cloud at the UFL.

o Flame extinguishment in the cloud when (1) the LFL

is reached, or (2) burn-off of the cloud occurs.

o Delayed pool ignition when the propagating flame

reaches the pool boundary.

The existence of a pool fire modifies both the pool and cloud growths due to

an increased (and initially transient) vaporization rate induced by the heat

load into the pool.

The capabilities of the combined LNG dynamics and fire model are best

illustrated by the distance-time representations of four scenarios given in

Figs. 21-24. All are for wind speeds of 4.48 m/s (10 mph) and ignition

times of 100 s after the ;tart of the spill. Figure 21 shows the flame

trajectory and transition for an assumed consumption speed of 1 m/s and

downwind ignition. The pool does not ignite, and only a small portion of

the total LNG burns off. At a consumption speed of 5 m/s (Fig. 22), the

cloud burns off and the pool is ignited; all LNG spilled is consumed.

Upwind ignitions under similar conditions produce a pool fire with the cloud

flame extinguished at the LFL for I m/s and cloud burn-off at 5 m/s; these

are shown in Figs. 23 and 24, respectively. The selected flame speeds in

these results are assumed to apply to both the diffusion flame and the

burnoff of premixed vapor clouds. It should also be pointed out that real

LNG fires may progress somewhat differently and that the results to some

extent reflect the idealizations of the model. In particular the assumption

of a spatially uniform cloud composition leads to a sudden fire

extinguishment at the LFL. In a real cloud some variation in vapor

composition is expected, with the heaviest concentrations probably around

the cloud center. Hence a more gradual fire extinguishment should be

expected.

For any fire scenario the integrated model will also yield the incident

radiant flux histories for targets located at arbitrary distances in the

downwind or upwind direction. Two flux histories are computed at each

location, namely, for vertical and horizontal targets, and the elevation of

the target relative to the base of the fire may be specified. Typical

radiant flux histories for a pool fire (immediate ignition) and targets at
57

2

IGNITION

DI FFUSION

-IL

0 200 400 600

FLAMEs

Figure 21. Flame Path for Cloud Fire; Downwind Ignition;

Ignition Time - 100 s, Flame Speed - 1 m/s,

Wind Speed - 4.48 m/s.
58

IGNITION

DIFFUSION

E FLAME

,..J .W .

S0 z

POOL FIRE

CLOUD

c:)

-I

0 200 400 6

TIME, s

Figure 22. Flame Path for Cloud Fire; Downwind Ignition,

Ignition Time = 100 s, Flame Speed = 5 m/s,

Wind Speed - 4.48 m/s.

00
59

2

E

W

C-, M

I

0

-I

0 200 400 600

TIME, s

Figure 23. Flame Path for Cloud Fire; Upwind Ignition,

Ignition Time - 100 s, Flame Speed - 1 m/s,

Wind Speed - 4.48 m/s.
60

2

CLOUD

0

POOL FIRE

0

IGNITIONDIFFUSION

FLAME

-I

0 200 400 600

TIME,s

Figure 24. Flame Path for Cloud Fire; Upwind Ignition,

Ignition Time - 100 s, Flame Speed = 5 m/s

Wind Speed = 4.48 m/s.
61

zero elevation are shown in Fig. 25. As expected, the fluxes diminish with

distance from the fire, and vertical targets (solid lines) receive more

incident radiation then horizontal targets (dashed lines). The results

shown in Fig. 26 are for vertically oriented targets, downwind ignition at

100 s, flame speed of 5 m/s, and wind speed of 4.48 m/s. In this case,

upwind targets (negative sign) receive in general less radiation than

downwind targets at comparable distances from the spill site. Detailed flux

histories for a wide variation in the fire, spreading, and ignition

parameters are presented in Appendices B and C.

The combined LNG dispersion and fire model computes only the heat fluxes

arising from radiation at some distance from the fire. Thus targets

engulfed by the fire are not treated by the model. Both radiation and

convective heat fluxes are important in this case, and the target response

cannot be decoupled from the heat input. The heat transfer must be based on

the temperatures of the flame and target as well as the heat-transfer

coefficient at the target surface. An upper boundary on the heat transfer

for this case is readily estimated by assuming that the target-surface

temperature quickly (instantaneously) becomes equal to the flame

temperature.

5.3 Result Summary for LNG Fires

The foregoing description of the fire models indicates that a great

multitude of fire scenarios may occur. It is therefore difficult to

assemble a comprehensive result summary, particularly for the expected

radiation-flux histories. Even when the problem parameters that only

influence the vapor cloud spreading are fixed at their nominal values, it is

found (see Appendix B) that at least five major variables influence the

incident heat flux at any particular location:

o Flame speed

o Wind speed

o Spill duration

o Ignition time

o Ignition location (upwind or downwind)

Although it is not possible to state categorically the order of importance

of these variables, it does appear (Appendix B), that the total heat flux is

least sensitive to the spill duration, at least for relatively short

durations (under 1000 s). Obviously, in addition to the fire variables, the

location (relative to the spill site) and orientation of the target will

determine the incident heat flux. From the many computations carried out
40

30

U- 20

I0

0

0 200 400 600 800

TIME, S

Figure 25. Thermal Radiation Histories for Pool Fire;

Spill Duration - 600 s, Wind Speed - 4.48 m/s,

Target Evaluation - 0.
50

40

N 5

-J

LL

I0

30

20

15

I0

5

0

0 100 200 300 400 500 600 700 800

TIMEI S

Figure 26. Thermal Radiation Histories at Vertical Targets;

Downwind Ignition at 100 s, Spill Duration - 600 s,

Wind Speed - 4.48 m/s, Flame Speed - 5 m/s.
64

(Appendixes B and C), it has also been found that a typical radiation-flux

history cannot be defined; i.e., even qualitatively the radiation-time

variations are not similar and depend strongly on the fire variables and the

target location.

To arrive at certain generalizations, consider first the case of fixed spill

duration (600 s) and zero wind speed, which in turn implies no cloud

drift. Figures 27-29 show approximately the flame regions for this case as

functions of the ignition delay time. The assumed flame speeds are 1, 5,

and 10 m/s, respectively, and apply to all fire regimes, i.e., pool fire,

diffusion flame, and premixed vapor cloud fire regions. The principal

boundaries of the flame region are (1) the ignition boundary along which

ignition is assumed to occur at some ignition delay time, (2) the cloud

burn-off boundary that is reached if the entire vapor cloud is consumed, and

(3) the LFL burnout boundary where the cloud fire is extinguished as its

concentration drops below the LFL. Targets within about 2 km of the spill

center can be expected to be within range of the flame, and this distance

can, of course, increase significantly in the downwind direction if wind is

present.

The time durations for the pool, vapor cloud, and total fire at a flame

speed of 5 m/s are presented in Fig. 30. The pool and vapor cloud fire

generally overlap, and except for spikes due to the cloud fire, the total

fire durations are similar at all flame speeds (see Appendix C). The total

fire durations are quite moderate, and the maximum duration is associated

with immediate ignition, i.e., a direct pool fire. If ignition is delayed

too long, the vapor cloud disperses below the LFL and no fire occurs. For

similar no-wind conditions but a shorter spill duration (60 s), the total

fire durations decrease substantially to about 200 s (see Appendix B). This

is due to the fact that for shorter spill durations it takes less time to

reach the LFL and the cloud radii are smaller than for the longer spill

times.

Another measure of fire severity is the total thermal load incident at a

particular location. Its value is obtained by integrating the incident

radiant flux over the appropriate fire-duration interval. Figures 31-33

give the total thermal load again for a no-wind condition and the three

reference flame speeds of 1, 5, and 10 m/s, respectively. Ignition delay

times range from zero (immediate ignition-pool fire) through 400 s in steps

of 100 s. Negative distances in these figures indicate the upwind direction

or in this case the direction 1800 out of phase with the ignition source.

Because ignition can occur only at one edge of the cloud, the total thermalenergy distribution, except for immediate ignition, is not symmetric about
2

LaJ

z

la0 1; I0 zoo 300 400 500

o IGNITION DELAY TIME, s

LFL BURNOUT BOUNDARY

CLOUD BURNOFF BOUNDARY

FLAME SPEEDS 1 rm/s -2

Figure 27. Extent of Flame Region as a Function of Ignition Delay Time for Flame

Speeds of 1 m/s; Wind Speed - 0.
2

I

IGNITION BOUNDARY

e

LAd

-j

UC-,

a

0

I -

300 400 500

IGNITION DELAY TIME, s

-LFL BURNOUT

BOUNDARY

FLAME SPEEDS = 5m/s

0~'

0*'

CLOUD BURNOFF

BOUNDARY

-2

Figure 28. Extent of Flame Region as a Function of Ignition Delay Time for Flame

Speeds of 5 m/s; Wind Speed = 0.
1 4

2

IGNITION BOUNDARY

re.

LUI

I-M

-J-

-J 0 .* * . . . . ..........

U. 0 ......... .. ..-

:•400 500

LLFL

<: BURNOUT

UBOUNDARY

CLOUD BURNOFF

BOUNDARY FLAME

SPEEDS= 10 m/s

-2

Figure 29. Extent of Flame Region as a Function of Ignition Delay Time for Flame

Speeds of 10 m/s; Wind Speed - 0.
68

600

500

Ii,

Ca.

C>

P

400

300

200

100

0 100 200 300 400 500

IGNITION DELAY TIME, s

Figure 30. Time Durations of Pool, Cloud, and Total Fires as a Function of Ignition

Delay Time for Flame Speed of 5 m/s.
I I

IGNITION DELAY TIME= 0 0 IN~S

1 1 200S

,IGNITION SIDE

NO WIND

FLAME 5;PEFEDS : Iiiml

=4

lOOs

-3 -2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER,kkm

Figure 31. Total Thermal Load Received by Vertical Targets at Ground Level as a

Function of Distance and Ignition Delay Time for Flame Speeds of 1 m/s.
IGNITION DELAY TIME=O

200s

0 IOOs 200s

a a - 300s

400s

IGNITION SIDE

NO WIND

FLAME SPEEDS: 5m/s

4.

4400

- --

-3 -2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER, km

Figure 32. Total Thermal Load Received by Vertical Targets at Ground Level as a

Function of Distance and Ignition Delay Time for Flame Speeds of 5 m/s.
IGNITION DELAY TIME = 0

200S IOOsI 300s I II

0 1 s 200s 300s

400s

IGNITION SIDE

NO WIND

FLAME SPEEDS:I10m/s

-3 -2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER,km

Figure 33. Total Thermal Load Received by Vertical Targets at Ground Level as a

Function of Distance and Ignition Delay Time for Flame Speeds of 10 m/s.
72

the spill origin. In all cases, the total incident thermal load drops very

rapidly as the location of interest moves to about 500 m outside the actual

flame region. The contribution of the pool fire only to these total loads,

expressed as a ratio of pool fire to total load, is given in Fig. 34 for

ignition delay times of 100 and 200 s and for flame speeds of 5 and 10

m/s. In many cases at longer distances the pool fire provides the dominant

portion of the heat load. Figure 35 compares the total incident radiation

loads received by vertically and horizontally oriented targets. A flame

speed of 5 m/s is assumed, and comparisions are shown for immediate ignition

and an ignition delay of 200 s. As indicated earlier, the vertical targets

in general receive larger heat loads, particularly at some distance away

from the flame.

Still another measure of fire severity is the peak thermal flux received at

a specific location. Figure 36 is a plot of these peak thermal fluxes as a

function of distance for several ignition delay times and an assumed flame

speed of 5 m/s. The peak loads are seen to vary in a manner similar to the

total heat loads. Although the scale in the figure extends to a maximum

value of 100 kW/m2, directly adjacent to the flame the peak radiation flux

is much higher. Assuming a flame temperature of 1100°C, the peak flux is in

excess of 200 kW/m 2 . This provides some measure of how very rapidly the

peak radiation flux decays with distance from the flame surface.

The peak-to-average radiation flux ratio is shown in Fig. 37 as a function

of distance from the spill location. The average flux is determined by

dividing the total thermal load received at a location by the appropriate

total duration of the fire. The results are for several ignition delay

times and for flame speeds of 5 and 10 m/s. The most noteworthy feature of

these plots is that the ratio appears to be nearly constant at locations

well outside the flame region, with values of the ratio ranging from less

than 2 to about 6. Close to the flame region the ratio becomes quite large,

in excess of 10, indicating that short-duration peak fluxes occur at these

locations.

The results presented above may be used directly to estimate the gross fire

environment resulting from the ignition of LNG under no-wind conditions.

Although similar results can be readily produced for other LNG spill and

wind conditions, it is also possible to use the no-wind results in

conjunction with cloud drift and other dispersion parameters to infer the

general nature of the fire environment that my exist in these altered

situations. Table 7 compares the total thermal load estimates at fixed

distances from the spill source, between a no-wind condition and a wind

speed of 4.48 m/s. Both the flame speed and ignition time are varied. For
73

LU

01

LU

CL

100

50

0 -

-3

100 -

50

0-

-3

-2 -I 0 i 2 3

-2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER, km

Figure 34. Pool Fire Contribution to Total Thermal Load Received by Vertical

Targets at Ground Level with Ignition Delay Time as Parameter.
IGNITION DELAY TIME =0 0

I 1200s

200s I I

I - I VERTICAL TARGET

1 1 1 --- HORIZONTAL TARGET I III

I I 611 ,IGNITION SIDE

=I NO WIND

IFLAME \5m/s SPEEDS 

I~4

/ j

/i / \

-3 -2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER,km

Figure 35. Comparison of Total Thermal Loads Received by Vertical and Horizontal

Targets at Ground Plane.
IGNITION DELAY TIME =1OOs 0

200s I

IOOs

300s

400s

IGNITION SIDE

NO WIND

FLAME SPEEDS =5m/s

-3 -2 0 I 2 3

DISTANCE FROM SPILL CENTER,km

Figure 36. Peak Thermal Load Flux at Vertical Targets at Ground Level as a Function

of Distance and Ignition Delay Time for Flame Speeds of 5 m/s.
76

FLAME SPEEDS=IOm/s

CC

UI.-

Cx

LAJ

a-.

I

-3 -2 -I 0 2 3

FLAME SPEEDS = 5 m/s I1

200 s

I I I I I

-3 -2 -I 0 I 2 3

DISTANCE FROM SPILL CENTER, km

Figure 37. Peak-to-Average Radiant Flux Ratio as a Function of Distance

and Ignition Delay Time.
77

TABLE 7

TOTAL THERMAL LOADS RECEIVED BY VERTICAL

SURFACE TARGETS-600s SPILL DURATION

Target to

Spill Center

Distance (m)

Total Thermal

Ignition

Downwind

200 300

Load (kJ/m2 )

Time (s)

0 100 100

Upwind

200 300

Wind Speed - 0 m/s

Flame Speeds = 1 m/s

2500

2000

1500

-1500

-2000

311

498

903

903

498

268

471

997

189

126

249

512

1506

90

62

Flame Speeds = 5 m/s

191

466

2292

41

29

640

1113

x

x

366

Same

as

downwind

2500

2000

1500

-1500

-2000

311

498

903

903

498

552

923

1783

1487

801

579

910

X

x

6446

Same

as

downwind

Wind Speed = 4.48 m/s

Flame Speeds - 1 m/s

2500

2000

1500

-1500

-2000

326

531

985

985

531

Flame Speeds = 5 m/s

913

5677

x

103

74

1019

1885

4218

1541

878

1812

x

x

50

37

952

1716

4045

819

529

4060

x

x

21

16

335

529

936

1143

618

224

347

597

946

494

547

928

1969

1080

611

54

83

138

293

147

365

576

1017

964

540

2500

2000

1500

-1500

-2000

326

531

985

985

531

2204

7060

x

243

181

386

661

1849

985

544

x Indicates vapor cloud reaches target before ignition; model not

in this case.

- Indicates distance in upwind direction; when windspeed is zero,

away from the ignition source is indicated.

applicable

the direction

Note: Ignition is assumed at cloud edge.
78

a direct pool fire (immediate ignition), the differences between the no-wind

and finite-wind-speed conditions are minimal. However, when ignition is

delayed, the situation changes radically. At slow flame speeds and long

ignition delays, many of the downwind locations become engulfed by the flame

region in the presence of wind-induced cloud drift. As expected, due to the

cloud drift, higher thermal loads are encountered at large distances from

the spill source.
79

6. LNG EXPLOSIONS

The subject of explosion of unconfined LNG/air vapor clouds is currently

controversial and poorly understood. On the one hand, there is some

indication [231 that enormous amounts of energy are required to explode an

unconfined methane/air mixture, and on the other hand, there is experimental

evidence that LNG vapor/air clouds may sustain a detonation wave [6]. It

has also been shown [24] that to generate substantial overpressures it is

not necessary to detonate a fuel-vapor cloud; fast and/or accelerating

deflagration is sufficient. In fact, at least for spherical vapor/air

mixture clouds, the overpressures and impulses resulting at large distances

from the source do not depend significantly on the speed of the combustion

reaction, as long as the flame speed exceeds some minimum value [251. It is

therefore reasonable to assume that a large unconfined LNG vapor-air cloud

may explode due to the influence of a number of factors including the

following:

o A significant fraction of hydrocarbons heavier than methane

(e.g., propane, ethane) in an LNG cloud.

o The possibility of stcong ignition sources.

o The effects of local confinement and flame-acceleration

mechanisms (e.g., the presence of various obstacles in the path of

the cloud).

o The potential size of the cloud itself.

Since the velocity of the combustion in an unconfined vapor cloud explosion

is ill-defined, it is conservative to assume the worst case of an ideal

detonation. This will result in the highest overpressures within the cloud

and in its close proximity, while as stated earlier, the pressure fields at

large distances are very much alike. It is then necessary to examine the

effects of the explosion geometry itself. As indicated earlier, the

expected vapor clouds are "pancake like" with a radius many times larger

than the cloud height, a ratio of radius to height of 50 being not

uncommon. Therefore, as a detonation wave propagates through such a cloud,

significant pressure unloading may be expected from the bottom to top

surface of the cloud. The most important aspect of the explosion problem is

to determine the coupling of the detonation energy to the surrounding air

medium and thus the estimation of air-blast wave as a function of time and

location. Although this problem can be solved by numerical methods (i.e.,

using an appropriate computer code), this approach would be very

uneconomical because of the unusual geometry of the explosion field. Thus a
80

two-stage approach is taken here. First, the actual detonation within the

vapor cloud is solved by semianalytical methods. The results of this

solution are then used as initial conditions for a two-dimensional

(axisymmetric) hydrodynamic computer calculation to determine the blast wave

propagating into the surrounding air. Blast-wave pressures based on a onedimensional hemispherical explosion source have been used by other

investigators [26] to estimate the pressure field from vapor-cloud

explosions. Because of the flat pancake geometry of the LNG cloud, this

approach is not felt to be valid in the current application.

6.1 Vapor-Cloud Detonation Parameters

The pressure field inside an exploding vapor cloud and the air blast

generated in the near field environment depend both on the thermodynamics of

the explosion and on the geometry of the detonating mass. As pointed out

earlier, an ideal detonation is assumed in the current application. The

parameters governing such an explosion include the ambient conditions,

cloud-mixture properties, and the Chapman-Jouget (CJ) detonation state. The

general equation that relates the thermodynamic state behind the detonation

wave traveling in an ideal gas to the energetics of the gas is given as

- 1)e [2 y- 1 +1) /2 (25) p e 0 + - 1) e1 (2e 0 + 1

where

p = P d/P.

and

P

eo = (y + ntCm ec

The descriptions and the values of the variables defining the CJ state and

the ambient cloud conditions, assuming a stoichiometric mixture (= 10 vol %

of LNG) are tabulated below:
81

Description Variable Value

Detonation pressure Pd 1.5513 MPa

(225 psia)

Detonation wave speed

Detonation particle speed

Detonatin sound speed

Detonation temperature

D

Ud

Cd

1695 m/s

741 m/s

954 m/s

Td 2511 K

Detonation density

Ratio of specific heats of detonation

products

Combustion conversion efficiency

2.052 kg/m3

X 1.29

0.71

Low heat of combustion

Ambient pressure

Mixture temperature

49.98 MJ/kg

P..

(14.7

0.10135 MPa

psia)

292 K

1.153 Mixture kg/m3 density

T.

Pa.

Mixture mass concentration CM 0.06

6.2 Two-Dimensional Explosion Models

The vapor-cloud geometries associated with a large spill of LNG on water

have been shown to be rather thin "pancake-like" configurations. Since such

extreme geometries have a marked impact upon the airblast environment, it is

essential to factor in the source (cloud) shape in computing the resulting

pressure fields. To accomplish this, a nominal or generic cloud

configuration is selected, and the two-step solution procedure mentioned

earlier is adopted. Some aspects of the solution are scalable in terms of

the cloud height; hence the solution possesses some universality, and the
82

results obtained can be extended to most LNG clouds resulting from a spill

on water.

A typical cloud configuration is illustrated (to scale) in Fig. 38. The

cloud radius is 1.1 km and its height is 25 m; hence its radius-to-height

ratio is 44. The volumetric concentration of the fuel at this stage of

cloud development is 10%, which corresponds approximately to stoichlometric

conditions. The equivalent hemispherical source based upon equal volume is

also illustrated in this figure. The equivalent spherical source radius is

350 m. Thus the actual cloud extends out to 3.14 equivalent spherical

radii. The extreme geometric dissimilarity between the two sources is quite

apparent.

As indicated earlier, it is conservative to assume that an initiation source

of sufficient strength can occur and that the cloud composition and size are

such that a stable detonation can develop. The assumed explosion scenario

is illustrated in Fig. 39 and defines four distinct phases. The first is

the initiation phase (a); the initiation can occur anywhere within the cloud

region and terminates when a self-sustaining flame or detonation-wave

situation occurs. The second phase (b) is a transition phase associated

with flame acceleration and the establishment of a stable detonation-wave

system. Pressure waves will exist during this phase, and they will

spacially precede the chemical-release process. The third phase (c) is the

establishment of a stable detonation-wave system in which the detonation

front has caught up to any preceding pressure waves in the undetonated

mixture and where the compression waves in the air have coalesced into a

single air-shockfront. This detonation-wave system will be quasi-steady in

nature, and the local explosion environment will be defined by the

detonation state and the pressure relief caused by the impedance mismatch

with the air above the cloud. The fourth phase (d) is the airblast phase,

which starts when local burn-out occurs as the detonation wave reaches the

end of the cloud. Thereafter, the chemically released energy is

redistributed hydrodynamically, and an outward propagating blast wave will

be established. The characteristics (i.e., peak overpressure, impulse, and

duration) of this blast environment at locations near the cloud are of prime

interest.

Most probably ignition will occur locally, i.e., from a point source. This

will lead to a three-dimensional wave system, and hence detonation and airblast environments that vary with respect to the horizontal azimuth. Nonetheless the scenario described in Fig. 39 will still apply, and the blast

environment in the direction of detonation-wave propagation will be most

severe. A complete three-dimensional solution of the problem is beyond the
EQUIVALENT SPHERICAL SOURCE

TYPICAL CLOUD CONFIGURATION

2-D COMPUTATIONAL REGION

0)

0 Ikm1 GROUND PLANE

Rs 2Rs 3Rs 4Rs

Figure 38. Typical Cloud Configuration at 10% Fuel Concentration.
84

, _-INITIATION SOURCE

(a)

COMPRESSION WAVE

(b) ACCELERATING FLAME

QUASI-STEADY

DETONATION

FIELD

-SHOCK WAVE ,,- PLANT

(d)

Figure 39. Assumed Explosion Scenario.
85

scope of the present effort. However, a two-dimensional solution can embody

most of the salient features of the phenomenon and represent a conservative

(worst-case) estimate of the expected explosion environment.

The two-step approach used to develop the two-dimensional air blast solution

consists of first establishing the quasi-steady detonation environment

defined in Phase c (Fig. 39c) and then to use this flow field as an initial

condition for a two-dimensional axisymmetric hydrocode calculation which

starts at the time of local burnout, i.e., when the detonation wave reaches

the cloud edge. The computational region used for the nominal cloud

configuration is shown in Fig. 38.

6.2.1 Quasi-steady Detonation Field

The large aspect ratio of the LNG vapor cloud, i.e., ratio of radius to

height, will tend to make the quasi-steady detonation field planar.

Unloading from the top of the cloud because of upward expansion of the

reaction products will predominate, making pressure decay due to cylindrical

divergence and Taylor-wave expansion near the detonation source less

important (see Section 3 of Appendix C). Since, in addition, the detonation

wave will propagate through the cloud at constant speed, the estimation of

the detonation field can be reduced locally to the solution of a twodimensional planar and quasi-steady problem. The number of independent

variables is thus reduced from three to two; that is, time is eliminated

explicitly. The independent variables are then the vertical and horizontal

distances z and y; and the only characteristic length associated with the

problem is the cloud height H. The latter can be used to define

dimensionless similarity variables

= z/H

and (26)

- y/H.

Since the detonation wave and the related pressure field propagate

horizontally at a constant speed D, the horizontal distance y represents a

measure of time, i.e.,

y - Dt. (27)
86

Thus n can also be considered to be the independent time variable. The

above scaling renders the results more general, making them applicable to a

cloud of any height.

The resulting pressure-wave system is illustrated in Fig. 40 using the

dimensionless variables. The system is rendered stationary by using a

coordinate system attached to the detonation wave. Thus the combustible

mixture approaching the detonation-wave boundary with a constant speed D is

converted into reaction products at high pressure and sonic particle

velocity. The upward expansion of these products create a shock wave in the

air above the cloud and a fan of rarefaction waves in the gases that reflect

from the rigid ground surface. A curved interface separates the air from

the reaction products. The pressures in the two gases across the interface

are equal; however, the tangential velocities on either side may vary due to

slip. The air shock wave strength, while variable along its length, is

continuous. The solution of this planar, but complex wave system is

obtained using the conventional method of characteristic techniques [27].

It provides the pressures and velocities in both the air and combustion

gases of the detonation field (see Section 3 of Appendix C).

A specific result of this calculation is the spatial pressure distribution

along the ground plane behind the detonation front. This is presented in

Fig. 41 in terms of the normalized absolute pressure, which is obtained by

dividing the pressure by the (absolute) detonation pressures. Since the

horizontal distance variable is also a time variable, the curve can be

Interpreted as a pressure history. The integration of this result with

respect to time yields the impulse per unit area that a surface element

experiences as the wave system moves past a given location. This impulse,

adjusted to the ambient pressure is

0. 8 APdH

I - , (28) cd

where APd is the detonation overpressure, cd is the sound velocity of the

detonation state, and H is the cloud height.

6.2.2 Airblast Environment Computations

An existing two-dimensional axisymmetric Eulerian hydrocode called BUNIE

[28] is used to compute the airblast environment that exists outside the

vapor cloud after the detonation wave reaches the edge of the cloud. The

calculation is carried out in terms of height units and thus scales with

respect to cloud height. A nominal radius of curvature is used,

corresponding to a location of the cloud edge at 50 height units from the
00

-4

DETONATION

WAVE

5 4 3 2 1 0

n7

Figure 40. Wave Diagram for Quasi-Steady Detonation Field.
I I IOII II

0 -

0 2 4

Figure 41. Dimensionless Pressure History/Profile Behind Detonation Front at Ground

Level.

3
89

origin. Thus blast attenuation due to spatial divergence is approximately

accounted for in the solution. Since the computer code is only capable of

treating a single continuous perfect-gas region, a compromise value of the

ratio of specific heats I - 1.35 is used for both the air and combustion

gases. This approach has been found satisfactory in earlier applications

[28).

The computational region extends up 9 height units, 10 height units

upstream, and 16 height units downstreamn. A variable spacing grid is used

to provide good resolution in the region of airblast wave formation and at

the same time to avoid long computer runs. The total calculation is carried

out for 500 time steps corresponding to 33 ms of real time per meter of

cloud height. This is the time when the shock wave along the ground passes

out of the computational grid. Initial conditions, i.e., pressure,

velocities and thermodynamic variables in each computational cell are given

by the quasi-steady explosion field obtained from the analytical approach

described earlier. Since the finite difference technique used relies on

artificial viscosity methods to accommodate shock waves, some smearing of

these waves occurs. The computer code is based on the FLIC computational

methodology, which used a "donor cell" technique to avoid certain

calculational instabilities [28]. This in turn produces a certain amount of

unavoidable numerical dispersion. However, the computed results are thought

to capture the salient features of the air blast field in the close

proximity of the detonating vapor cloud.

6.3 Airblast Result Summary

Once burnout of the vapor cloud occurs and energy input to the system

ceases, the expanding airblast is characterized by a rapid decrease of peak

overpressure; e.g., after propagating about one height unit along the ground

range, the peak shock has already decayed to 55Z of its initial value. The

pressure field adjacent to the cloud is illustrated by the pressure-time

histories given in Figs. 42 and 43 for locations along the ground and at two

cloud height units. The horizontal locations correspond to positions 0, 1,

2, ... , to 15 height units downstream from the edge of the cloud. The peak

overpressure decays rapidly along the ground immediately in front of the

cloud. Note that the overpressure scale differs in the two figures and that

the pressure history for the horizontal location four height units

downstream from the cloud edge is repeated in both figures for comparison

purposes.

The basic characteristics of the airblast environment are best represented

in terms of peak overpressure and impulse variations as a function of
1.5I

1.0

0.5 71 0

= 0 5 10 15 20 25 30 35

USCALED TIME, ms/M

LU

cc 1.5 0 a. IIII1 cc 710 •;=0

0

10 -

0 5

3 4

0 10 15 20 25 30 35

SCALED TIME, ms/m

Figure 42. Pressure Histories for LNG Cloud Explosion.
0 .2 ,7: 7:=4 1I I 

0.1

= 0 5 10 15 20 25 30 35

cn SCALED TIME, ms/m

UJ, 0.2 7.1•:4 { :0 1

6 "

0.1 0 7 8 1

0 I0WI 1

0 10 15 20 25 30 35

SCALED TIME, ms/m

Figure 43. Pressure Histories for LNG Cloud Explosion (Continued).
92

relative range measured in cloud-height units. These are shown in Figs. 44

and 45 for nominal elevations of zero (ground level), one, and two cloud

heights. At the cloud edge, all the curves become very flat for relative

ranges less than 1. As a reference value, the detonation overpressure, APd

- 1.45 MPa, is indicated in Fig. 44. The interaction of the detonation wave

with the air at the edge of the cloud results in an instantaneous drop in

overpressure to 1.16 MPa because of the impedance mismatch of the two gases;

this is shown in the results for zero elevation. Close to the detonating

cloud, the peak overpressure is highest along the ground level. At relative

ranges greater than two, the influence of elevation becomes quite small,

with some indication that the overpressure at higher elevations is slightly

larger. This effect can be observed in the overpressure histories given in

Fig. 43. In this range the overpressure decays approximately as the

reciprocal of the horizontal distance. At the limit of the calculations or

a range of 15 height units, the overpressure decays to about 0.05 MPa (7

psig). Extrapolating the results to the lowest level of interest or about

0.007 MPa (1 psig), a range of about 100 height units is indicated. For a

nominal cloud height of 25 m, this implies a distance of 2.5 km from the

cloud edge.

The side-on impulse, i.e., the impulse in the incident blast wave, as a

function of relative range is shown in Fig. 45. The values given are in

units of MPa ms per meter of cloud height. The impulse obtained from the

quasi-steady detonation solution Id M 1.22 HPa ms/m is indicated in the

figure. The impulse decay with range is slow until a relative range value

of approximately 8 is reached. Thereafter the decay rate increases

sharply. Since the calculations were terminated at a time of 33 ms/m, the

impulse values at relative ranges greater than 10 are estimated values and

it is not feasible to extrapolate the results significantly beyond the range

of the computations. The influence of elevation on impulse is small, some

differences existing at close-in locations.

The third important blast-wave characteristic, the duration of the

overpressure, has to be estimated from the pressure-time histories. Because

of the very flat overpressure decay at late times (see Fig. 43), this

quantity is difficult to evaluate. In fact, during the entire computer

calculations the overpressure never decays to zero. However, setting an

arbitrary low overpressure of about 0.007 MPa (1 psig) as the limit, one

finds that the blast-wave durations are 10-20 ms per meter of cloud height.

Applying the above scaled results to the specific example of the vapor cloud

shown in Fig. 40, we can make a direct comparison with the often-used

equivalent one-dimensional blast estimates. As indicated earlier, this
93

5 I I I 111111 I I III I I I I I II II I I I I I I I I I I I I I I

Il~

C,,

we

I

0.1

A Pd

"'- t = 0

t=21 iii

0.01

I Itilill I I I I I IIiTi 0.005 i

0.I I 10

RELATIVE RANGE, H UNITS

100

Figure 44. Peak Overpressure Variation for LNG Cloud Explosion.
I0_o I 11111 •• lw• I I IIII_

Cto

22

0.1 1 10 100

RELATIVE RANGE, H UNITS

Figure 45. Impulse Variation for LNG Cloud Explosion.

J( 0.
95

cloud has a height of 25 m and a radius of 1.1 km, while the radius of the

equivalent hemispheral cloud is 350m. The peak overpressure comparison is

shown in Fig. 46. Here the range is measured from the cloud center

(origin), where initiation is assumed to occur. The current two-dimensional

results are labeled "LNG Cloud" and show a peak overpressure within the

cloud equal to the detonation value of 1.45 MPa. Outside the cloud the

pressure decays rapidly, reaching a value of 0.05 HPa (7 psig) at 1.5 km and

extrapolating to the 0.007 HPa (1 psig) level at around 3.6 km.

Using a TNT equivalency concept [29], the curve labeled "TNT Equivalence" in

Fig. 46 is obtained. Since this curve is based on matching overpressures in

the 0.007 MPa (1 psi) range, it can clearly not be applied at significantly

higher pressures and definitely takes no account of the geometric

distortions embodied in the two-dimensional model. The overpressures are

greatly overestimated at close-in ranges, then underestimated at

intermediate distances, i.e., ranges of the order of the cloud radius. The

curve labeled "Det." corresponds to hemispherical explosions of propane

[30], and although the energetics of this system are slightly different than

those of LNG, it is clear that this data also grossly underestimates the

overpressure at intermediate ranges. The other curves in the figure

represent a volume detonation and calculated overpressure curves for various

deflagration speeds all assuming a spherical source [30]. It is evident

that none of these results are adequate to predict the pressure field

resulting from the explosion of a "pancake-like" vapor cloud.

The impulse results for the two-dimensional calculations and for the TNT

equivalency method are presented in Fig. 47. (No impulse information is

available for the detonating hemispherical gas source.) The impulse

estimates obtained from the point-source TNT equivalency methodology are

wholly inadequate over the entire range of interest. Thus the current twodimensional results are the only meaningful impulse estimate available.

Similarly, the positive phase durations of the overpressure waves, which are

in the range of 0.25-0.50 s for the example cloud, are the only meaningful

duration estimates.
96

tO u I as'' I * * l IuII.,

TNT

EQUIVALENCE

SDET LNG CLOUD

a- IS0.1

CL

0.0

0.01 \lPS \

R 170 mIs

LR :210 m/s

0.1 1 10

RANGE0, km

Figure 46. Comparisons of Peak Overpressure Estimates for LNG Cloud Explosion.
97

100 * , , , ,E , , I , I,-

LNG

(f" "CLOUD

E

"10

U,

-J .%

ZTNT

EQUIVALENCE -

3 .3 !1101 I 3, I J I l

OI I IO

RANG E, km

Figure - 47. Comparisons of Impulse Extimates for LNG Cloud Explosion.
98

7. LNG SPILL THREATS

The methods for determining the environment resulting from a massive spill

of LNG on water were outlined in the preceding chapters. A result summary

was also presented and more detailed data can be found in Appendixes A, B,

and C. Based on this information, an attempt is made to estimate the nature

and magnitude of the threat from such a spill. Although the term "threat"

is not very precise, it is defined here as the extent, intensity, and total

magnitude of a particular effect arising from the LNG spill. Thus of

interest are the distances and durations of a specific effect as well as

fluxes and their time integral. The intent here is to define the credible

maxima of these quantities, assuming that the total spill is limited to

25,000 m3 . The threats considered are the LNG liquid, i.e., the pool, the

vapor cloud, fire, and explosion. While the actual susceptibility of

nuclear power plants will be discussed later, the threat definition provided

here obviously takes cognizance of the ultimate application of the results.

7.1 LNG Pool

The major concerns with respect to the LNG liquid is that it may reach a

vital plant installation or that is may serve as the fuel source for a pool

fire. Thus the size or extent of the LNG pool is of primary importance.

This distance depends both on the spill duration and evaporation rate (see

Chapter 3 and Appendix A). Because of the latter dependence, the maximum

pool radii occur when ignition of the LNG does not take place. Using the

most credible evaporation rate (without fire) of 0.000305 m/s, maximum pool

radii of the order of 400 m may be expected. The persistence of the LNG

pool and quantity of liquid involved are abviously position-dependent, with

the longest durations and largest quantities at the spill source. This

location is of interest primarily in case of immediate ignition, i.e., for a

direct pool fire. At locations some distance from the source, the

persistence of the LNG pool can be estimated from the duration of the pool

spread. The latter quantity appears to peak at a value of 600 s (see

Appendix A) and gives an indication of the upper limit for pool persistence

for most locations, except in the immediate vicinity of the source where the

duration is equal to the spill duration. Since the average velocity of pool

spreading is of the order of 1 m/s, the actual persistence of the LNG pool

at most locations will be considerably shorter than the 600 s indicated

above. The spreading distance of a spill on water in a particular direction

may be enhanced due to current and wave action. However, these are slow

effects, which require considerable time [31]. Because of the cryogenic

nature of LNG and hence rapid evaporation, the influence of current and wave

action is expected to be minor.
99

7.2 Vapor Cloud

The LNG vapor cloud poses a direct threat to nuclear power plants primarily

through its asphyxiating effect on personnel at high enough

concentrations. Since LNG usually containsabout 85-90% of methane and 15-

10% higher hydrocarbons, it can be assumed to act as a simple asphyxiant,

i.e., only by the inert dilution of air and thus reduction in oxygen

level. The threshold of .personnel impairment for such an asphyxiant has

been estimated at a volumetric concentration of 28%, which corresponds to an

oxygen concentration of 15% [32]. Examining the computational results

(Appendixes B and C), we find that for most cases this concentration level

of the vapor cloud is reached rather quickly, i.e., usually in times less

than 250 s after spill initiation. Variables such as wind speed and spill

duration have little effect on concentration. The major effects come from

the relative humidity of the air and the initial gas composition. An

extreme case arises for high relative humidities (- 100%) and a pure methane

gas. It takes nearly 450 s to reach an LNG concentration of 28%. The

corresponding cloud radius at this time is about 750 m. The cloud-drift

distance is strongly influenced by wind speed and can be as much as 2200 m

for a wind speed of 8.96 m/s (20 mph) (see Appendix C). Thus the furthest

possible downwind distance at which LNG vapor concentrations are

sufficiently high to result in some personnel impairment due to reduced

oxygen levels is about 3000 m from the spill source.

For more usual LNG compositions, i.e., compositions including higher

hydrocarbons such that the molecular weight exceeds that of methane by 10%,

it is estimated that the critical composition is reached in about 250 s.

The cloud radius at this time is estimated to be 800 m, and the cloud drift,

even at a high wind speed of 8.96 m/s, does not exceed 500 m. Thus a most

plausible distance to the threshold of personnel impairment due to oxygen

depletion is about 1300 m.

Since the cloud composition varies only with time, the longest possible

duration associated with a condition of severe oxygen depletion is equal to

the time when the critical threshold composition is reached and is limited

to a region in the close proximity of the spill source. At loctions away

from the source, this duration is equal to the time interval starting with

the cloud arrival and ending when the threshold composition is reached

throughout the cloud. For the extreme case, i.e., pure methane, 100%

relative humidity, and 8.96 m/s wind speed, the total durations of oxygen

depletion are estimated as follows:
100

Distance from Spill

Source m Duration s

0 450

1000 300

2000 175

3000 0

Using the more realistic LNG composition, the condition of oxygen depletion

appears to last at most 250 s, and at a distance of 1000 m from the spill

source, this duration is reduced to about 30 s.

7.3 Fire Potential

The LNG vapor cloud also poses an indirect threat to any installation in

that it constitutes the fuel source for severe fires and explosions. Of

primary interest, therefore, are the distances and times to reach the

flammability limits, i.e., the LFL and UFL. The first of these indicates

the maximum extent of possible fire and explosion hazards; the UFL

represents the boundary between premixed and diffusion flames and delineates

the upper boundary for explosion regions.

In describing the combined LNG dispersion and fire model, we showed how

strongly the extent of the flame region depends on the time of ignition (see

Figs. 27-29). In defining the threat, this aspect is ignored, however, in

order to arrive at a conservative estimate. Figure 17 and Table 5 show the

distances to the LFL for various LNG compositions and relative hunidities of

the air. These values are for a single wind speed of 4.48 m/s (10 mph).

Although wind speed has no effect on the cloud-radius, it does influence the

cloud drift. As seen in Table 3, an increase in the wind speed from 4.48 to

8.96 m/s adds an additional 1300 m to the LFL cloud drift for the zerohumidity case.

The times to reach the flammability limits are very sensitive to the LNG

composition and relative humidity, but are essentially independent of wind

speed. For pure methane and 100% relative humidity it takes 2700 s to reach

the LFL and only 800 s to attain the UFL. Thus the cloud remains

combustible for 0.75 h, and the risk of an intense fire, in a premixed

cloud, and of explosion persists for more than half an hour. Using the more

realistic composition (molecular weight 10% higher than methane), the time

to LFL is estimated at 1450 s and to UFL at 470 s. Thus the time for

possible cloud flammability is reduced to less than 0.5 h and the risk of

explosion to about 0.28 h. For lower relative humidities of the air, the
101

times during which the vapor cloud represents a fire and explosion risk are

much reduced. Using the time to reach LFL and estimating the average drift

velocity at 6 m/s when the wind speed is 8.96 m/s, it is calculated that,

for the worst-case (pure methane, 100% relative humidity), the LFL extends

to about 17 km (or in excess of 10 miles). The corresponding distance for

the UFL is approximately 4.4 km using an average drift speed of 4.5 m/s.

When a more realistic LNG composition is assumed, i.e., a molecular weight

10% larger than that of methane, the maximum distances are: for the LFL,

9.8 km and for the UFL, 3.2 km. All these distances are to the farthest

edge of the cloud.

Again, reducing the relative humidities will reduce the distances. However,

the assumption of very high relative humidity, i.e., close to 100%, is not

unrealistic, considering that the main interest is in a layer of air above a

large water surface. Vapor clouds under consideration will probably not

keep their coherence f~r some of the long times required to reach the LFL.

This is particularly true for the pure methane case where the time is 0.75

h. Dispersion by air turbulence, which this model neglects, may play some

role for these cases. Therefore the most credible estimate of the threats,

using the current model, should be based on the heavier LNG composition.

7.4 Fire-Threat Magnitudes

The fire scenarios that may result from an LNG spill on water depend on a

large number of variables. It therefore becomes difficult to define a

credible fire threat. In all of the examples of flux histories (see Chapter

5 and Appendixes B and C) the durations of the fires are typically less than

600 s, the peak heat fluxes vary from 20 to 50 kW/m 2 , and the total thermal

loads range from less than 2 KY/m 2 to about 8 lj/m2r. These results depend

strongly on the time and the location of ignition and are representative for

distances of 500-2500 m from the fire source (burning cloud or pool). The

fire threat thus appears to be moderate at these locations. However, as

shown earlier (Chapter 5), the fire severity increases rapidly as the fire

source is approached. Both the fire-peak intensity and the total thermal

energy become large. Taking into consideration the long durations of cloud

flammability and cloud-drift distances that occur under certain conditions

of high humidity and LNG composition (see preceding section), severe fire

threats may arise in certain situations.

It has been shown (see Chapter 5) that flame temperatures for hydrocarbon

fuels burning as diffusion flames (fuel rich) are about 1373 K, and

stoichiometric premixed vapor clouds may burn with flame temperatures as

high as 2148K. Assuming Ideal conditions, i.e., black body radiation, the
102

corresponding radiation heat fluxes in the immediate vicinity of the flames

(no attenuation) are 201 and 1207 kW/m 2 respectively. The first of these

values agrees well with the measured flux of 210 kW/m 2 obtained in pool-fire

experiments with LNG [18]. Should a particular target be actually engulfed

by the fire, then in addition to radiation heating, convective heating will

take place. The latter will depend on the heat-transfer coefficient at the

surface of the target. For turbulent flames the maximum value of this

coefficient is estimated at 852 W/m2 K (150 Btu/h ft 2 *F) [33]. Using this

value the maximum convective heat flux to a target engulfed by the fire is

estimated at 920 kW/m 2 for the diffusion flame and 1,580 kW/m 2 for the

premixed vapor burning (2148 K). All the cited heat fluxes represent the

maximum possible flux, assuming an initial target-surface temperature of

20°C. As the target-surface temperature increases, the heat flux to the

target from both radiation and convection will diminish. However, the

quoted heat-flux values provide upper bounds for estimating the threat to

targets in contact with the combustion gases.

The second aspect that must be considered is the duration of the fire

threat. Given a fixed spill scenario, this duration is primarily determined

by the location and time of ignition, as well as the speed of the flame

traversing the fuel source. Although most LNG vapor clouds and pools remain

flammable only for times of about 10 min, it was shown that for compositions

approaching pure methane and for 100% relative air humidity, a cloud may be

flammable for as much as 0.75 h. Such a cloud reaches a near constant

radius of about 800 m in 500 s (see Appendix C, Figure 4.14). Thus if

ignition takes place after this time at one edge of the cloud and the flame

speed has a low value say 1 m/s, the duration of the fire, which must burn

across the diameter of the cloud, is 1600 s. However, note that the cloud

under consideration reaches the UFL at 800 s and that the flame velocity

would most likely increase at this point. Assuming a flame speed of 2 m/s

in the premixed region reduces the burning time to 950 s while a flame speed

of 5 m/s reduces the time to 560s.

For a more realistic LNG composition (molecular weight 10% higher than

methane) the nearly constant cloud radius of 1100 m is reached in 450 s.

Thus a fire starting at this time and having a flame speed of 1 m/s could

burn for about 1000 s until the LFL is reached (1450 s). Since most of the

burning takes place in a premixed cloud (UFL occurs at 470 s) higher flame

speeds may be expected. At 2 m/s, the fire duration remains at 1000 s,

because the LFL is- reached before the cloud is burned off. However, at a

flame speed of 5 m/s burn-out occurs after a fire duration of 450 s. Note

that the above fire threats are essentially independent of the LNG pool. In

all cases, to obtain the longest fire durations, ignition is assumed to be
103

delayed to a time when LNG evaporation is essentially complete.

The range, or distance from the spill site, at which the above fire threats

may occur is primarily a function of the cloud drift due to wind action. As

shown earlier, the maximum distance for a pure methane cloud at 100%

relative humidity and a wind speed of 8.96 m/s is 17 km. For the heavier

LNG composition, the corresponding range is 9.8 km. These distances are

very large; however, there exists a mitigating factor when the wind velocity

is high; namely, the cloud residence time at a given location is short.

This in turn reduces the fire threat.

The fire durations estimated above are the total times required to burn off

or extinguish the flame regardless of the location and motion of the vapor

cloud and flame front. In estimating the threat, however, one is interested

in the intense fire duration at a fixed location. The vapor clouds under

consideration have a finite size with a diameter of about, 2000 m.

Therefore, even for a stationary vapor cloud, a fixed target will experience

the high-intensity flame only for a fraction of the total fire duration,

because the flame will be either approaching or receding from the target

and, as shown earlier, the heat fluxes diminish rapidly as the distance from

the target increases. In addition, for a fire approaching the target, the

vapor cloud itself will provide some shielding from the heat flux.

With cloud drift present, the situation becomes more complex, since now the

residence time of the cloud itself at a given location is a function of the

drift velocity. Thus for a 2000 m diameter cloud and a wind speed of 8.96

m/s, which results in an average drift velocity of 6 m/s, the residence time

for the cloud at a fixed location is 333 s. For a wind speed of 4.48m/s

(drift speed, 3 m/s) this residence time is 667 s. Again as the burning

cloud drifts by the target, the actual location of the flame front and thus

the source of the heat flux may be aproaching, receding from, or passing the

target, depending on the flame speed and ignition location. Therefore, in

general, the heat flux received at the target will be smaller than the

maximum. The full residence time of the cloud at the target becomes equal

to the fire duration at the target only in case the flame speed (relative to

the unburned cloud) is equal to the drift speed, and ignition occurs at the

downwind edge of the cloud as it first contacts the target. Since drift

velocities and flame speeds are of the same order of magnitude, this is

probably the most realistic but worst-case fire scenario possible. The

farthest location from the spill source, where this full fire duration can

be realized, is at least one cloud diameter less then the above-cited

distances to the LFL. At greater distances, cloud dilution below the LFL

would shorten the fire duration. For the pure methane cloud, 100% relative
104

humidity, and a wind speed of 8.96 m/s, the farthest location would be 15 km

with a fire duration of 333 s. For the more realistic (heavier) LNG

composition, the location is about 8 km. When the wind speed is 4.48 m/s

and the flame and drift speed are reduced to 3.0 m/s, the farthest target

locations are 6.9 km for methane and about 2 km for the heavier LNG, with

durations increasing to 667 s. Since for all these cases the cloud

cQmposition is in a premixed state, i.e., below the UFL, the target could

experience the highest possible heat flux corresponding to a flame

temperature of 2148 K. These estimates are based on assuming a constant

drift velocity that is reasonable for the times and wind speeds under

consideration (see Appendix C). However, for smaller wind speeds this would

not apply. These cases are, however, of lesser interest, because the

distances would approach the spill source. The above estimates are thought

to provide upper bounds for the expected fire threat due to LNG spills on

water.

7.5 Explosion Threat

The magnitudes of the overpressure, impulse, and duration from the explosion

of a "pancake" shaped LNG cloud are shown in Chapter 6. To ascertain the

threat from the explosion, one must determine the distances relative to the

spill source at which such explosions may take place. Since the explosion

propagates through the cloud at a detonation speed of 1680 m/s, the entire

process will be completed even for a large cloud in about 1 s. Relative to

the cloud spreading and drift, this is essentially an instantaneous

process. The clouds present on explosion threat during the entire time they

reside in a premixed state (between UFL and LFL) and out to distances at

which the LFL is reached.

While for many spill scenarios these distances and times are moderate, it

has been shown (Section 7.3) that circumstances exist (high relative

humidity, nearly pure methane) when the threat persists a long time and with

appropriately high wind velocities can extend out to long ranges. Thus, for

a wind speed of 8.96 m/s (20 mph), a target may experience the detonation

overpressure and impulse at a distance of nearly 17 km from the source,

assuming pure methane. For a more realistic heavier LNG composition, the

range is 9.8 km.

The distance from the cloud edge to the safe overpressure level of 0.007 MPa

(1 psi) [29] was earlier estimated as equal to 100 cloud-height units (see

Chapter 6). For the cases considered above, these distances would be 14.5

km for the pure methane cloud, which has a height of 145 km and 7.3 km for

the heavier cloud, which has a height of 73 m at the LFL (see Table 5).
1.05

However, these distances are based on results obtained for a cloud with an

aspect ratio (radius to height) of 44 for which planar detonation is a good

approximation. For the clouds considered here, the aspect ratio is 5.9 for

the methane cloud and 15.8 for the heavier composition. Thus, the above

distances are probably overestimates of the true distances. A lover limit

for this distance may be obtained by considering the hemispherical explosion

of the cloud. The cloud contains approximately a fuel mass of 10O7 kg: with a

TNT mass equivalency of 2.4 [29]. The 1 psi level from such an explosion is

reached at a scaled distance of 18 rn/kg 113 [34). This yields a distance of

5.2 km from the explosion center and should be compared with 15.3 km for the

methane cloud and 8.4 k for the cloud of heavier composit ion, which are

obtained with a pancake cloud model. The latter distances are nov also

referred to the center of the cloud. As pointed out earlier, the estimates

based on the hemispherical cloud assumption represent a lover limit of -the

distance to the I psi overpressure level. The actual distance lies probably

somewhere between the two extremes. Considering the large drif t of the

cloud, one finds that the range from the spill source to a completely safe

overpressure level is quite large. For the extreme assumption of 100%

humidity and pure methane using a credible wind speed of 8.96 m/s (20 mph),

this range lies somewhere between 21.4 and 31.5 km. Using the heavier LNG

composition and assuming the same atmospheric conditions, the distance to

the safe overpressure level is estimated to lie between 13.9 and 16.0 km.

As pointed out before, it is somewhat doubtful that for the pure methane

composition the cloud will maintain its coherence over the long drift

distance predicted by the current model. Atmospheric turbulence coupled

with wind action may be expected to cause earlier dispersion below the LFL.

In case an LNG cloud explosion does take place, some degree of damage may be

expected, for targets such as a nuclear power plant, at all locations closer

to the spill source than the safe 0.007 MPa (1 psi) overpressure level. The

degree of damage will. depend on the overpressure level, the blast duration,

and the hardness of the target. Damage sufficiently severe to affect the

safe functioning of a nuclear power plant should probably be expected for

overpressures in excess of 0.207 HMPa (30 psig). This is the level at which

severe damage normally occurs to reinforced concrete windowless structures

if the duration of the blast duration is of the order of 0.5 to 2.0.s

[35]. From Fig. 43 It is seen that the blast duration for an LNG cloud

explosion at this pressure level approaches about 20 ins/rn of cloud height.

Hence, a cloud height of 75 m will yield a duration of about 1.5 s. While

such a cloud height develops only under conditions of high relative humidity

(see Table 5), heights of about 50 m and hence durations of 1 s are not

unusual. Therefore the 0.207 HMPa overpressure level appears to be a
106

reasonable threshold of severe damage to nuclear power plant structures and

functioning.

In case of a "pancake-like" cloud, this threshold occurs at about four

cloud-height units from the edge of the exploding cloud (see Figs. 42 and

43). For most cloud configurations, this represents a distance of less than

200 m from the cloud edge, and only in case of high humidity and pure

methane may the distance approach 500 m. Therefore, this pressure threshold

will depend primarily on the cloud drift. For the most severe case (100%

humidity, methane, 8.96 m/s wind speed), the distance from the spill source

is 17.6 km, and for the heavier LNG composition, the distance becomes 10.1

km. When a TNT equivalency approach or a detonating hemispherical cloud is

used to estimate the pressures from an LNG cloud explosion, the threshold

level of interest (30 psi) lies at a distance of 600-650 m from the center

of the cloud (see Fig. 46). This, however, is a location that actually lies

within the vapor cloud at the time LFL is reached (see Table 5) and thus in

reality would experience much higher overpressures. Since the 30 psi

overpressure level occurs such a short distance from the cloud edge (4

height units), the estimates based on the pancake cloud should be reasonably

accurate and will be used in the current application.

7.6 Threat Summary and Application

The most accurate way of establishing the LNG spill threat in a particular

application is to actually perform specific computations for the expected

separation distances (spill to target) and the prevailing atmospheric

conditions. This, however, is. a time-consuming procedure and requires

specific site information. To arrive at a first estimate of the expected

threat, the results presented in the preceding sections can be used

directly. Similarly, in the current application the threat as discussed

above will be used to estimate the vulnerability of nuclear power plants.

For convenience these results are summarized in Table 8. Both "worst-case"

and "typical" threats are presented. The latter approximately corresponds

to moderate winds and relative humidities (4.48 m/s and 50%, respectively)

as well as a realistic LNG composition (molecular weight 10% larger than

methane). The "worst-case" results are for 100% humidity and 8.96 m/s wind

velocity and again a realistic LNG composition. The data obtained for pure

methane are also listed for the worst case and appear in parentheses.

Because of the strong fire-target interaction, it is not possible to

estimate the total thermal energies in case of targets engulfed by the

fire. It is seen, however, that the initial peak heat fluxes are very

intense in this case. Thus the surface temperature of a target engulfed by

the fire will increase very rapidly. To estimate the heat transfer into the
107

TABLE 8

SUMMARY OF LNG THREATS

1. Spill and Pool

Spill duration resulting in maximum pool and cloud size

Maximum pool duration, no fire

Maximum pool radius

600 s

600 s

400 m

2. Vapor Cloud

(a) Asphyxiant (28% LNG by volume)

Max. Range, km Total Duration, s

1.3 250

(3.0) (450)

(b) Fire Potential

Duration at 1 km

from Spill Source, s

30

(300)

LFL(5%)

Max. Range, km Time, s

3.5 600

9.8 1450

(17.0) (2700)

Typical Case

Worst Case

UFL(15%)

Max.Range, km

0.8

3.2

(4.4)

Time, s

170

470

(800)

3. Fire

(a) Typical Targets Not Engulfed by Fire

Distance from spill source, km

Ignition Delay Times, s

Fire Duration, s

Peak Heat flux, kw/m 2

Total Thermal Loading MJ/mr2

(b) Targets Engulfed by Fire

(i) Fire Conditions

Flame Temperature, K

Max. Radiation Flux, IU/m 2

Max. Convective Flux, kW/m 2

0.5 to 2.5

0 to 400

600

20 to 50

2 to 8

Diffusion Flame Premixed Flame

1373 2148

201 1207

920 1580
108

TABLE 8 (cont'd)

(b) Targets Engulfed by Fire (cont'd)

(ii) Worst Case Fire

Max. Range, km 9.8 (17.0)

Total Duration, s 1000 (1600)

(iii) Fire Durations at Targets

Relative Humidity, %

Wind Speed, m/s

Distance, km

Duration, s

50

4.48

2.0

300

100

4.48

2.0(6.4)

670

100

8.96

8.0 (15.0)

330

4. Explosion

Detonation: 1.45 MPa

Distance, km

Impulse, MPa-ms

Duration, s (Estimate)

Severe Damage: 0.207 MPa (30 psi)

Distance, km

Impulse, MPa-ms

Duration, s (Estimate)

Safe Overpressure: 0.0007 MPa (1 psi)

Distance, km

Result Assumptions

LNG Molecular Weight, % of Methane

Wind Speed, m/s

Relative Humidity, %

Typical

3.5

37.0

0.100

3.7

25.0

0.5

6.5

Typical

110

4.48

50

Worst Case

9.8(17.0)

89(176)

0.250(0.500)

10.1(17.5)

45(89)

1.5(2.9)

15(26)

Worst Case

110(100)

8.96(8.96)

100(100)

Note: Numbers in parenthesis always refer to pure methane case.
109

target it is, for most applications, conservative and acceptable to assume

that the target surface temperature is equal to the flame temperataure vhen

the 'target is engulfed by the flame. This approach is used and is justified

by more elaborate calculations in the current study.
110

8. Survey of Safety Systems for Nuclear Power Plants

To ascertain the vulnerability of a nuclear power plant to the threats posed

by a large LNG spill on water, one must first identify the vital components

of the plant that may be affected by the spill environment. Of primary

interest are safety related plant systems particularly those that are

required for the safe shutdown and heat removal from the plant so as to

avoid any significant incidents that may create a safety hazard for the

general public. The current study is not concerned with a particular power

plant, i.e., is not site specific; rather, the LNG hazards are to be

evaluated in a generic manner. Currently operational nuclear power plants

[36] located along water transportation routes or along shorelines are

listed in Table 9. As can be seen, there is significant variability in

design, size, and age of the plants. Both P•#R and BWR systems are

represented. The general susceptibility of these plants to the threats

posed by an LNG spill may therefore be expected to vary. However, keeping

in mind the main concerns, i.e., safe shutdown and heat removal, we can

arrive at some generalizations.

To accomplish these operations in a nuclear power plant, the plant must

first of all be controllable. This requires not only that certain

instrumentation and control systems remain operational, but also that an

adequate supply of electric power be maintained and that key operating

personnel of the plant be able to function. For plant heat removal, at

least one of the major cooling circuits must be available. Also maintained

must be an ultimate heat sink, component cooling, and some of the airconditioning. This in turn implies the continued functioning of certain

equipment, such as pumps and valves, and obviously the availability of

electric power. It should also be noted that while the durations of the

acute threats from an LNG spill may be relatively short (see Chapter 7), all

the above-mentioned functions must be maintained for much longer periods

following the acute stage. This is most obvious with respect to electric

power supply. The period of acute threat may well be handled by the

emergency batteries, but in the long run (approximately 2 hours), another

power source must become available to the plant to continue the decay-heat

removal. Thus severe damage and loss of all offsite power and the emergency

onsite diesel generators would indeed constitute a severe incident.

As indicated earlier, the major threats from an LNG spill on water are

thermal loads produced by fire and overpressures in case of explosion.

Other hazards are oxygen depletion by the vapor cloud, and toxic effects of

either combustion gases or vapors released on site, e.g., chlorine, due to

the effects of the primary threats. Finally, in some cases a direct threat
TABLE 9

NUCLEAR POWER PLANTS ALONG WATERWAYS

NAME

CALVERT CLIFFS

CRYSTAL RIVER

DIABLO CANYON

INIDAN POINT

MILLSTONE

PILGRIM

SAINT LUCI

SALEM

SAN ONOFRE

SHOREHAM

SURRY

TURKEY POINT

OUTPUT

UNIT TYPE MANUF. MWt

1&2 PWR CE 2700

3 PWR B&W 2452

1 PWR W 3338

2 PWR W 3411

1 PWR B&W 615

2 PWR W 2758

3 PWR W 3025

1 BWR GE 2011

2 PWR CE 2700

1 BWR GE 1998

1&2 PWR CE 2570

LOCATION

Chesapeak Bay, MD

Gulf of Mexico, 70 ml N of Tampa,FL

Pacific Ocean, 10 mi from San Luis

Obispo, CA

East Bank of Hudson

River, 24 mi N of New York

Long Island Sound-North Shore

40 mi SE from Hartford, CT

Cape Cod Bay, Western Shore,

Plymouth, MA

Hutchison Island near Fort Pierce,FL

Delaware River, 8 mi SW of Salem, NJ

Pacific Ocean Camp Pendleton Marine

Station, S of San Clemente, CA

North Shore of Long Island, NY

just N of BNL

James River (Gravel Neck Point), VA

Biscayne Bay, 25 ml S of Miami, FL

REMARKS

LNG terminal at Cove Point

approximately 5 km away

Steel Enclosure Bldg.

around containment.

'-I

I-.

1

1

PWR

PWR

W

W

3350

1347

Steel Containment Concrete

Shield Bldg.

30 ft seawall, Containment

Sphere Enclosure Bldg.

4000 acre, 168 mi long grid

of closed cycle cooling

ponds and canals.

I BWR GE 2436

l&2 PWR

3&4 PWR

W

W

2441

2200
112

may arise from the LNG liquid pool, either through possible ingestion into

the cooling-water system or through the effects of the extremely cold liquid

on plant structures located at the water line. As was shown in Chapter 7,

both the type of threats and their severity depend on many factors, the most

important of which are: distance between the spill location and the plant

site, wind conditions, relative humidity, LNG composition, and ignition time

and location.

A survey of the vital NPP systems and components [37, 38] readily indicates

that most of them are housed inside substantial, usually Class I,

structures. Thus, the effects of the LNG spill environment on these systems

are in general only indirect, unless the protection provided by the

structure is breached. Also, the state of the plant at the time the spill

environment or its effects reach the site may be of utmost importance; e.g.,

has the reactor been shut down, have all ventilation system dampers been

closed, has all unnecessary equipment been shut down, has the operating

personnel been warned, etc.? In specific cases, an accident scenario must

therefore be assumed in order to evaluate the hazard. This in turn, would

depend on such plant features as detection and warning systems. The

detailed consideration of all these factors is clearly beyond the scope of

the present effort. In case the LNG in either vapor or liquid form does not

reach the site, the threats are limited to thermal loads due to the

radiation from offsite fires and to overpressure effects from a possible

offsite explosion. The major hazard in this case may be the possible

structural or functional failures of vital systems and components. Ignition

of combustibles may also occur, and there is some possibility that heat

loads could become excessive for the available cooling capacities. Among

safety-related systems, the offsite power supply appears to be the most

vulnerable.

If the LNG vapor and/or liquid reach the plant site, additional threats must

be considered. Included are oxygen depletion, toxicity, and the earlier

mentioned effects of LNG liquid. The fire threat is now enhanced by

convective heating, and explosions are more likely and severe because of

possible confinement. In particular, volume detonations may occur when

combustible mixtures ignite inside a nearly enclosed space. Again the

offsite power supply appears to be most vulnerable, but in addition, any

space containing vital equipment which is not completely closed off or which

cannot withstand the severity of the loading may be vulnerable. This could

include auxiliary diesel power, emergency batteries, intake structures,

diesel fuel storage, or even the control room and cable spreading room, via

their air intake and ventilating systems. Tables 10 and 11 summarize the

possible threats, effects, and affected safety systems for offsite and
113

TABLE 10

SYSTEMS SUSCEPTIBLE TO OFF-SITE LNG THREATS

Threat Possible Effects Susceptible Items Most Likely Affected

Safety Systems

Fire Ignition of On-site Fuels (exposed) Off-site power

(Radiation Combustibles supply

Only) Wooden Structures

Construction Materials

Electrical Cables

Excessive Heating of Steel Structures Off-site power

Structures and Compon- Concrete Structures supply

ents - Structural Storage Tanks

Failures Electrical Equipment

(insulators, transformers etc.)

Excessive heat loads - Ventilation & air con- Off-site power

Insufficient heat re- ditioning systems supply

moval capability Electrical Equipment

Personnel Injuries Exposed Personnel

Burns

Explosion- Structural & Function- Electrical Equipment Off-site power

Fast al Failures Ventilation Systems supply

Deflagra- Storage Tanks

tion Exposed Pipine Systems

(Over- All Structures &

pressure- buildings (except

Impulse) reinforced concrete)

Personnel Injuries Exposed Personnel
114

TABLE 11

SYSTEMS SUSCEPTIBLE TO ON-SITE LNG THREATS

Most Likely Affected Threat Possible Effects Susceptible Items Sikely Sfsted Safety Systems

LNG Pool Structural & Functional Intake Structure Cooling water cirFailures Cooling Water Circulating culating pumps

pumps

LNG Vapor Oxygen Depletion Personnel-Asphyxiation Control Room

Cloud personnel

Fire Ignition of Combusti- Fuels Off-site Power

(Radiation & bles (exposed) Wooden Structures supply

Convection) Construction Materials

Electrical Cables

Fire Propagation into Ventilation Systems Emergency power batPlant Interior Air Intakes and Exhausts tery room, Control

room, cable spreading room

Structural & Functional Exposed Steel Structures Off-site power

Failures Concrete Structure supply

Storage Tanks

Piping Systems

Electrical Equipment

Excessive Heat Loads - Ventilation &Air Condi- Off-site power

Insufficient Heat tioning Systems, Elec- supply, Control

Removal trical Equipment, room

Cooling Systems

Toxicity-Combustion Exposed or sheltered Control Room PerGases Personnel sonnel

Personnel Injuries - Exposed or sheltered Control room PerBurns Personnel sonnel

ExplosionFast Deflagration (overpressure -

Impulse)

Structural & Functional

Failures

All exposed equipment Storage Tanks

Ventilation Systems

All Structures

Piping Systems

In order of increasing severity of

blast:

Off-site power

Main Steam or cooling water lines

(exposed)

Diesel fuel storageblock house

Intake structure

Auxiliary building

Nuclear fuel storage

Containment building

- I I

ExplosionFast Deflaggration(overpressure -

Impulse)

(cont'd)

Overpressure Propagation into Plant

Ventilation Systems

Air Intake and Exhaust

Systems

Control Room

Emergency Power Battery Room

Emergency Diesel

Systems

Control Room

Personnel Personnel Injuries Exposed and Sheltered

Personnel
115

onslte LNG situations. It is possible to imagine a whole gamut of effects,

and their likelihood and severity, as pointed out earlier, depends very much

on the state of NPP, i.e., both its operational status and its degree of

Isolation from the outside environment. Although many items are identified

in the hazards compilations (Tables 10 and 11), only the most likely and

severe hazards, i.e., selected systems, components, and personnel, are

analyzed for their susceptibilities.
116

9. Nuclear Power Plant Susceptibility

To estimate the effects of the LNG spill environment on a nuclear power

plant, it is first necessary to define the susceptibility of such a plant to

the types of threats posed by the spill. As shown in the preceding chapter,

only a limited number of plant systems need be considered to assume safe

shutdown and heat removal. The approach taken in determining

susceptibilities is based on using available information from the literature

concerning the generic response or resistance of a system or component to a

specific threat, e.g., fire, blast. Thus the process amounts to

establishing simple damage criteria, and no detailed analyses are

undertaken. It is well known [391 that, regardless of the system considered

and the type of threat, the susceptibility may be a function of both the

threat intensity and of the time integral of the intensity. The latter

provides a measure of the total dosage, impulse, energy, etc., delivered to

the system. For each threat and system considered, there usually exist

threshold values for both the intensity and its time integral below which

the threat will have no effect on the system. Similarly, when a fixed isodamage level for the system is considered, there again exist minimum values

for both intensity and its time integral, which must be exceeded to produce

this damage. Therefore in establishing susceptibilities of a system in

general, one should consider both factors, i.e., intensity and total

amount. Only in the extreme case of a threat of very long duration is it

usually possible to define damage criteria or susceptibility limits based

only on the intensity. Conversely, for threats of very short durations

relative to the response time of the system, criteria may be established

using as a basis only the total amount (time integral) of the threat.

In many cases it is possible to define the susceptibility of a system, or at

least its first-order estimate, simply in terms of the level of the threat,

i.e., its intensity and duration. This usually requires the existence of a

large body of experimental data concerning the effects of the hazard under

consideration. This approach is usually applicable where the response of

the system has little or no effect on the intensity of the threat

environment, i.e., the threat environment and the response are not

interactive. On the other hand, when the threat and system response are

interactive and the times required to produce a response are relatively

long, it is difficult to arrive at susceptibilities in terms of the threat

levels alone. In these cases it is first necessary to define damage levels

or criteria in terms of the response of the system. These can then be used

in conjunction with analytical or computational procedures to determine the

threat magnitudes required to produce the specified damage level. This

approach can be used in all cases. However, as pointed out above, it is
117

often simpler, when the appropriate data base exists, to go directly to

criteria defined in terms of the threat level.

Concerning this latter more direct approach, note that much of the data

required for the current application comes from the weapons-effects

literature [35,40]. Therefore, the intensities and durations of the blast

and thermal threats are not always commensurate with the possible hazard

environments created by an LNG spill, and care must be exercised when using

these data for the current application. Finally, since the interest here

lies in safe operation and shutdown of nuclear power plants, the

susceptibility limits must be chosen conservatively, so as to preclude any

doubt that the intended functions can be carried out.

9.1 Personnel Exposure Limits

All the threats produced by a large LNG spill on water may to some extent

affect or incapacitate the personnel at a nuclear power plant. Many of the

effects may be strongly mitigated or completely eliminated for personnel

housed inside a structure such as the typical nuclear power plant auxiliary

building, which normally contains the control room. The thresholds provided

here are therefore primarily for personnel who are found in unprotected

areas of the power plant.

The response of the human system, primarily the skin, to thermal loads is a

function of both the heat-flux intensity and the duration. Interest here,

however, lies in intensity levels that will produce immediate (in less than

a few minutes) effects. Although there is some disparity in the literature

data, it does appear that intensity levels of 3-5 kW/m 2 provide the

threshold for first-degree burns in less than 1 min exposure [39,41,42].

Thus a level of 4 kW/m 2 will be taken to indicate the limit for personnel

exposure below which there will be no impairment of function.

It was pointed out earlier that LNG can act as an asphyxiant through oxygen

depletion in the atmosphere. Again, the response of the human body to this

threat is fairly immediate. Thus, a criterion based solely on the intensity

level, in this case, concentration, is acceptable. It is fairly well

established that, for a simple asphyxiant, a concentration by volume of 42%

or an oxygen depletion in the atmosphere to 12% constitutes the threshold

for mental impairment 132]. However, for the current application, where

operator performance and not injury is the issue, this oxygen depletion

level is considered too low [431. Therefore a level of 15% of oxygen in the

atmosphere will be taken as the limit for the impairment of operator
118

function. This corresponds to a level of 28 vol % of LNG vapor in the

atmosphere.

No direct toxic effects are produced by LNG. However, some secondary or

indirect toxic effects may arise from the burning of LNG vapors. The most

important of these is the toxicity of combustion gases, which in turn means

the concentrations of carbon dioxide and carbon monoxide. The latter

produces significant physiological effects, even for concentrations as low

as 0.1% [32]. However, long exposures (hours) are required at these low

concentrations to produce these effects. Thus, 0.1% of 00 is taken as the

limit for the current application. While there are some noticeable

physiological effects for carbon dioxide concentrations of 2 vol %, e.g.,

increase in the depth of respiration, no impairment of function is expected

at this level [321. Hence the 2% level of C02 is accepted as the limit for

the current purpose.

Airblast can produce multiple effects and injuries on exposed personnel.

The level of injury strongly depends on both the overpressure and

duration. It was seen earlier that the expected blast durations from LNG

cloud explosions are quite long (on the order of 0.5-1.0 s); i.e., the

duration is comparable with those from nuclear weapons explosions in the

high-kiloton or low-megaton range. Since the response time of the body is

much shorter than these durations, simple overpressure criteria are

acceptable for the determination of susceptibility. In addition to the

direct blast effects on personnel, injuries are possible due to secondary

effects produced by impacting missiles and debris 135,39). Also, tertiary

injuries arise when the human subject is accelerated and displaced by the

blast drag forces and impacts a stationary object or is dragged along the

ground surface. However, the aim here is not to determine severity, but to

establish thresholds under which the power plant personnel can keep on

functioning effectively. An overpressure of 0.0345 MPa (5 psi) is therefore

taken as the limit of personnel exposure, since it represents the threshold

both for eardrum damage and for injuries due to small-missile (10 g) impact.

9.2 Airblast Effects on Structures and Components

The structures at typical nuclear power plants may range from heavy

reinforced-concrete containment buildings to light-wall cinder block

structures. Obviously they vary greatly in their resistance to air blast

effects and in their natural frequencies. Concerning the latter, even for

the most massive structures the response times are of the order of 1 s.

These are times comparable in length to the durations of the expected blast

waves. It is also a well-documented fact that massive reinforced-concrete
119

buildings are primarily overpressure sensitive [35,40]. Therefore, simple

overpressure criteria will again be used to establish the susceptibility of

these structures. Some of the other equipment items of interest, e.g.,

electric power systems, switchyards, high-voltage power lines, and pipe

lines, are primarily drag-sensitive. However, the blast durations are

sufficiently long, in comparision to their response times, to permit again

the use of a simple overpressure criterion. Using information from the

nuclear weapons effects literature [35,391, one arrives at the

susceptibility criteria presented in Table 12. These values are considered

threshold values at which some damage to the system in question is to be

expected. However, its functionality should not be significantly impaired.

In at least one case, namely for reinforced-concrete structures, the

overpressure criterion of 0.2069 HPa (30 psi) was thoroughly examined, using

structural loading and response analysis techniques that are widely accepted

in the design of protective construction [44]. It was found that for a

building with dimensions and construction representative of an auxiliary

building (20 m high, 20 m deep and 40 m long; 0.45 m thick walls, 0.3 m

thick roof, and 3% reinforcing), the displacements produced by a 30 psi peak

overpressure blast wave having a duration of 1 s, and approximating that of

an LNG explosion, are almost completely in the elastic range. Similar for

concrete panels 4.6 m (15 ft) on the side, 0.3 m thick, and with 1%

reinforcement, the displacements and stresses were found to be acceptable.

The computations provide some confidence that the established

susceptibilities are realistic. Further confirmation of the criteria comes

from the literature [45], where typical containment structures subjected to

blast loadings are analyzed in detail using finite element codes.

9.3 Thermal Susceptibility of Structures

It was brought out earlier that the susceptibility cannot always be defined

in terms of threat-environment variables. When the response of the systems

is closely interactive with the environment, it is necessary to define

criteria in terms of the response variables and then to provide estimates

for each case of the system response, given a prescribed environment. The

response of structures and components to thermal loads is a good example of

this. The heat transfer to the structure is so dependent on the structure

itself and on the relative geometry of the fire and structure that it is

difficult to define susceptibility criteria in terms of the fire variables

alone.

The most widely accepted design criterion for steel structures and

structural elements, such as beams and columns, requires that the average
120

TABLE 12

AIR BLAST SUSCEPTIBILITY CRITERIA -

STRUCTURES AND COMPONENTS

Type of Structure

Reinforced Concrete Containment Buildings

Reinforced Concrete Structures:

Auxiliary Building, Intake Structures

Fuel Storage Structures, etc.

Heavy Industrial Type Building (steel frame):

Turbine Building

Light Industrial Type or Brick Building:

Shops, Offices, Stores

Wood Frame or Cinder Block Buildings:

Offices, Stores, Guard Posts

Large Pipes and Pipe Bridges

Electrical Power Systems:

Switchyard, substation, high voltage lines

Telephone lines

Ventilation Stacks (metal)

Overpressure

MPa

.3448

.2069

.0828

.0345

.0172

.0483

.0345

.0241

.0207

Level

(psi)

50.0

30.0

12.0

5.0

2.5

7.0

5.0

3.5

3.0
121

temperature of the elements not exceed 538°C (1000*F) [33,46]. This

criterion corresponds to a reduction of the ultimate strength of structural

steel to 46% of its value at ambient temperature. It can also be shown [46]

that, for a typical steel column under steady-state conditions, i.e., longduration fires, and under the assumption that the structure is not engulfed

by the flame, the above criterion implies a radiation thermal load of 29.5

kW/m 2 . Other investigators [42] have suggested that, for uncooled steel

structures, the radiation thermal load from long-duration fires be limited

to 24 kW/m 2 . This corresponds to an average steel temperature of 493*C.

Finally, if most (90%) of the ultimate strength of a steel structure must be

retained, a limit of 400*C for the steel has been suggested [33). Again

this corresponds to a steady-state thermal load for a steel column of 14.9

kW/m 2 . For the general case of unsteady thermal loads, it is best, however,

to rely directly on the temperature criteria.

Most of the important buildings on a typical nuclear power plant site are of

reinforced concrete construction.' Criteria and design guides for reinforced

concrete under thermal loading have been developed by many investigators and

are well established [33,47,48,49]. Most of these criteria are based on the

concept of a critical temperature for the reinforcing or prestressing steel

of reinforced concrete structures. When the critical temperatures are

exceeded, a structure or structural element is considered unsafe. Although

there is some variation among the values used by various investigators or in

various countries, these differences are minor. For the current

application, a conservative set of temperature criteria has been selected

from the literature. Their values are given below and are applicable to

both concrete walls and structural elements such as beams and columns.

Temperature Criteria for Reinforced Concrete Structures

Maximum allowable temperature on unexposed

face of concrete wall 1390C (250*F)

Maximum temperature of prestressing steel 427-C (800-F)

Maximum temperature of reinforcing steel 500°C (932-F)

A value of 593°C (1100°F) is often used in the literature for the maximum

temperature of reinforcing steel [33]. This, however, appears to be too

high a value, since it corresponds to a reduction of ultimate strength to

20% of its ambient value. The first criterion is equivalent to the heating

of all the concrete in the wall to the above-indicated temperature and

implies [491 possible ignition of combustibles on the unheated side of the

wall.
122

The above criteria are developed using standard fire resistance tests [48]

in which the average heat flux is 27.5 kW/M 2 . The criteria also carry with

the assumption that no spalling of the concrete occurs. However, the

potential for spalling due to the buildup of steam pressure always exists

when the concrete temperature exceeds 100°C. The evaporation of free water

in the concrete, the migration of the water vapor, and its possible

recondensation are a very complex phenomenon. The potential for spall

increases as the critical temperature (100°C) moves to greater depth. This

is due to both the greater amount of free water at greater depth and the

increase in resistance to steam outflow from the heated surface. The larger

amount of water at greater depth is a result of recondensation of steam

liberated during the heating of shallower layers. The critical depth for

spalling appears to be about 5 cm from the heated surface (33,48).

Explosive spalling may occur early in the fire or at later times, 0.25 to

0.50 hour and may result in exposing the main steel to high temperatures.

This can lead to the serious weakening of a structural member, particularly

for prestressed beams and heavily reinforced columns. In the current

application, the above temperature criteria, together with the critical

depth for spalling, will be used as a measure of susceptibility for concrete

structures.

9.4 Ignition of Combustibles

The ignition of combustible and flammable materials at a nuclear power plant

site is of interest because such ignition may lead to secondary fires

directly at the site, and this in turn may result in the damage of a vital

plant component. Also, ignition of some materials, e.g., electrical

insulation, can directly lead to a failure affecting the plant's

operation. Although there may be a variety of combustible materials at a

plant site, only a small number of representative materials will be treated

here.

In general, wood can be considered as representative of construction

materials and may actually constitute part of some less important on-site

buildings. Its ignition response under thermal radiation has been

thoroughly documented [50], and correlations for the required heat fluxes

and times have been derived. Assuming an average thermal inertia for wood,

one can obtain simple equations for these correlations [32], for both

piloted and spontaneous ignition. The former requires the presence of an

ignition pilot, such as a small flame or spark. The expressions governing

the ignition of wood are given below. The nomenclature used in these

expressions is as follows:
123

k: Thermal conductivity

p: Density

C: Thermal capacity

Ir: Incident radiant flux

Ip: Minimum radiant flux

for pilot ignition

is: Minimum radiant flux for

spontaneous ignition

t p: Time to pilot ignition

ts: Time to spontaneous

ignition

teff: Effective duration of

incident radiation

Average thermal inertia of wood:

kpC - 8.76 x 105 J 2 /(m 5 _ s

Pilot ignition:

I = 13.4

p

(W/m K)

(kg/m3)

(J/kg -K)

(kW/m 2 )

(kW/m 2 )

(kW/m 2 )

(s)

(s)

(s)

- K5 )

(29)

=p 76.4 3/2

r p

Criterion:

I I ; t t

r p eff p

Spontaneous ignition:

I = 25.4

5/4

t (f83.7 )

r s

(30)
124

Criterion:

I I ; t t

Ir >s teff s

In either case, two criteria must be met: (1) the incident radiation must

be equal to or larger than the threshold value and (2) the duration must

exceed the computed minimum value.

For more readily combustible materials, the limits for pilot and spontaneous

ignition are often indistinguishable. Some of these examples taken from the

literature [32] are provided below. These may be considered as

representative of the most easily ignitable materials on site.

Cotton:

I - 12.5 s

t - 7.17 x 103 1-1.64 (31)

s r

c-cellulose (paper):

I = 16.72

S

t - 4.41 x 106 1-3.64 (32)

s r

Again the criteria for ignition to occur are

I I ,t )t

r s " teff s

Another group of materials, the ignition of which is of interest, are

plastic polymers. These are considered representative of materials used in

electrical insulation. Because of their importance in industrial

applications, the thermal and ignition response of polymers has been studied

extensively [51, 52]. A correlation relationship based on experimental data

has been derived for the time required for ignition to occur.

160(k C) 0 .75 (T -T)

ti a (33)

r

This equation is dimensional and cannot be converted to arbitrary units.

The appropriate units are:
125

ti: ignition time: s

k: thermal conductivity: cal/g s *C

density: g/cm3

C: heat capacity: cal/g "C

Ti: ignition temperature: 6C

Ta: ambient temperature: OC

Ir: radiant flux cal/cm2 s

To ascertain if ignition will occur, the values of the ignition temperature

must also be known; these can be obtained from the literature [32]. Some

representative values are listed below.

Material kpC Ti'OC

Silicone rubber 2.363 x 10-4 720

PVC 1.68 x 10-4 220

21 ~ ~rteff 1/2

T Z T + r (34)

n(kpC)I

1 2

The criteria for ignition to occur are

T > T i ; t eff > tI .

The actual procedure is first to use Eq. 34 to estimate the surface

temperature and to check if it exceeds Ti. Then Eq. 33 is used to compute

the time required for ignition at the given radiation level. This time is

finally compared to the effective duration of the thermal radiation to see

if ignition is possible.

9.5 Fracture of Brittle Materials

Brittle materials such as ceramics play a large role as insulators in

electrical substations and power lines. Such materials may fail by fracture

when subjected to large temperature changes. The actual stresses generated

in an electric insulator subjected to sudden thermal load depend in a

complex manner on the geometry and material characteristics of the
126

insulator. Precise estimates of the response are therefore difficult to

obtain. However, using some limiting assumptions, one can get first-order

estimates of the response [54-56]. These are based on the concept of

thermal-stress-resistance factors. Assuming that either the surface heattransfer coefficient is very large or that the heat load is so overwhelming

that the surface temperature of the body, to a good approximation, changes

instantaneously to a new higher value, the thermal resistance factor is

given as

R =s( - (35)

Ea

where S can be the tensile strength (St) or the shear strength (Ss); also, v

is the Poisson's ratio, E the modulus of elasticity, and a the coefficient

of linear thermal expansion. The case of both tensile and shear failure

must be considered. The latter may be the more likely mode of failure in

case of sudden heating at the surface. The temperature difference that

leads to failure is the expressed as

ATf = A R, (36)

where A is a factor that depends on the geometry of the body and type of

failure mode, i.e., tensile or shear. For a sphere, the factor A takes on a

value of 3/2 for failure at the center by tension due to uniform surface

heating, or a value of 2 if failure is by shear [54]. This is in reasonable

agreement with the values quoted by other investigators [56]. For

instantaneous temperature increase at the surface, the maximum tensile

stress at the center can be shown to be [57]

o - 0.386 R(Ts - Ta ) (37)

where T. is the suddenly applied surface temperature. This maximum stress

occurs at a time t given by

b2C

b 2Cp

t = 0.0574 k p (38)

where b is the radius of the sphere, Cp the heat capacity, p the density,

and k the thermal conductivity. The maximum temperature difference to which

a sphere can be subjected by convective heating over a wide range of Biot

Numbers (0 - bh/k) is approximately given as [56]

2.5S t(1 - v)

AT max [1 + 2/8] • (39) c
127

Here h is the heat-transfer coefficient. When 0 is large, Eq. 39 reduces to

Eq. 37.

If the surface heat-transfer coefficient h is constant and the heating is

moderate, the resistance parameter becomes 154,56]

kS (1 - v)

R' E (40)

where the nomenclature is the same as used above.

failure then is [54]

ATfi R'A' I f hb

The temperature at

(41)

where b is a characteristic dimension of the body under consideration. For

a sphere, a value of A' - 5.0 may be used [56]. Typical of the materials

used for insulators in power lines and distribution stations is porcelain.

Although some variability in the material constants of insulators is to be

expected, the following representative values taken from the literature [54-

56] are used in making response estimates.

Density p - 3996 kg/m3

Heat capacity

Heat conductivity (at 400°C]

Linear thermal-expansion coefficient

Modulus of elasticity

Poisson's ratio

Tensile strength

Cp- 904.3 J/kg • K

k = 1.76 W/m - K

a - 6 x 10- 6 K-1

E - 6.2 x 104 MPa

v- 0.3

St . 55 HPa

Shear strength Ss -M 3St

Thermal resistance

(Constant surface temperaturetension)

Thermal resistance

(Constant heat-transfer

coefficient--tension)

R - 103 K

R' - 182 W/m
128

It is not possible to obtain accurate results of the thermal response of

insulators, without performing detailed stress analyses. This holds

particularly when the heating is not uniform and when complex geometric

shapes are considered. This kind of analysis is, however, beyond the scope

of the current effort. Therefore, the above simple expressions will be used

to make vulnerability evaluations. The results of such computations should

provide order-of-magnitude estimates of the hazard.
129

10. Vulnerability of Nuclear Power Plants

To establish the vulnerability of nuclear power plants to the hazards

arising from large LNG spills on water it is necessary to apply the LNG

threats to the identified important plant systems and compare the response

with the earlier defined susceptibilities. As mentioned before, it is

difficult to perform such an analysis without considering actual spill

scenarios including assumptions concerning the state of the plant.

Concerning the latter, it is here, in general, assumed that plant personnel

are cognizant of the existing hazard and that proper actions have been taken

to mitigate as much as possible the consequences of the spill. Hence, among

other things, it is assumed that by the time any effects arrive at the plant

site, the reactor and all unnecessary equipment has been shut down and that

all plant openings such as vent ilat ion-duct dampers, doors, and hatches have

been secured. Where, in specific cases, these assumptions are relaxed, it

is so stated.

on examining the hazards associated with an LNG spill (see Chapter 7), one

sees that some of them pose much greater threats than others. The analysis

here will concentrate on the former. It is f irst noted that the threat

posed by the LNG liquid (pool) Is probably insignificant. Considering that

maximum pool diameters are of the order of 400 m, it not likely that LNG

liquid will reach tbe plant, and even if it does it appears that, at such

close-in distances, the effect on the plant arising from other threats such

as fire and explosion will be much more severe than those associated with

the LN~G pool. A similar conclusion can be reached concerning the effects of

oxygen depletion where the maximum credible distance to the asphyxiation

limit is estimated as 1300 m and durations of threat are very short (about

30 s). With a closed-up plant these effects are then truly minimal.

Therefore these two threats, i.e., liquid LNG and oxygen depletion, are not

considered in establishing plant vulnerability.

it has also been shown [321 that, even for very large fires, one is not

likely to obtain, at ground level, concentrations of combustion products

that are toxic. It Is possible to envision fire propagating into the plant

interior under some circumstances and thus creating a toxicity hazard.

However, with the assumption of a closed-up plant this is not likely to be

the case. Should - f ire or smoke nevertheless ingress the plant, then

combustion-product toxicity could indeed represent a severe hazard.

However, nuclear power plants are in general, at least in the control-room

area, equipped with smoke detectors [381. Thus plant personnel can take

evasive actions, eog., using emergency breathing apparatus, in order to

mitigate the combustion-product toxicity threat. To make a reasonable
130

estimate of the toxicity threat, one must estimate infilitration and airexchange rates into the plant areas occupied by personnel. Due to the

uncertainties concerning the plant configuration and state, it is not

possible to make any meaningful estimates of these quantities on a generic

basis. Because of this and also because some protective action can be taken

against the toxicity hazard, no further consideration is given to this

threat. It should, however, be emphasized that this threat is not thought

to be negligible and may have to be considered in specific cases.

In estimating the plant vulnerability to other threats, one can organize the

results in two ways: (1) by type of threat and (2) by plant system.

Although it is convenient to use the latter approach with respect to plant

personnel, a presentation by type of threat is used for all other plant

systems. Note also that all maximum vulnerability distances presented in

the following are obtained using a realistic LNG composition (molecular

weight 10% in excess of the value for pure methane), high relative humidity

(approaching 100%), and a wind speed of 8.96 m/s (20 mph). Should the LNG

vapor composition approach that of pure methane, the maximum vulnerability

distances may have to be increased substantially, i.e., on the order of 7-8

km.

10.1 Personnel Vulnerability

Besides the above-discussed threats of oxygen depletion and toxicity, plant

personnel are susceptible to thermal loads from fire and overpressure and

impulse from explosion-generated airblast. Concerning thermal loads, a

limit of 4 kW/m 2 is used here (see Chapter 9) as the threshold value for

impairment of personnel functioning. From the results presented earlier

(see Chapters 5 and 7), it is readily established that this rather low flux

value may extend to a distance of up to 3.0 km from the edge of the fire.

At the same time, a realistic estimate for the maximum cloud drift to reach

the LFL is 9.8 km. Thus, exposed (in the open) plant personnel may be

vulnerable to thermal loads from LNG fires on water up to distances of 13 km

from the spill location. On the other hand, personnel located inside a

substantial structure, such as the auxiliary building, are essentially

completely protected from the threat of thermal loads. The only exception

to this is the case of a power plant which has not been closed up before the

LNG vapor cloud drifts on site and ignites. Then it is possible for a

combustible mixture to enter the plant ventilation system and ultimately

pose a threat to personnel. The maximum credible distance from the spill

source for such a scenario is about 9 km. However, its occurrence is

considered highly unlikely.
131

The threshold for personnel injury from airbiast is 0.0345 MPa (5 psi),

which occurs about 1.1 km from the edge of an exploding vapor cloud. This

estimate is based on the assumption of a cloud height of approximately 50 m

and is reduced for lover clouds and somewhat Increased for greater cloud

heights. Considering the possible cloud drift, it is estimated that exposed

personnel remain vulnerable to airblast up to about 11 km from the spill

source.

Again, personnel stationed inside a substantial building are essentially

immune to air blast effects, unless the building itself is severely damaged

by blast. Although pressure waves can propagate Into the plant interior

occupied by personnel via a breached ventilation system, it is 'difficult to

estimate the exterior blast-wave overpressure that will lead to this

scenario. Details of the ventilation system design, i.e., its geometry and

blast resistance, must be known to arrive at a realistic estimate. However,

it can be surmised that the occurrence of such an event is not likely. For

ordinary ventilation systems, the intake or exhaust stacks may fail at an

overpressure as low as 0.0207 MPa (3 psi). Similarly, the flow dampers,

which close up the plant, are not expected to act as blast protection

devices. Thus, they may be expected to fail at relatively low

overpressures. Finally, the ducting itself Is probably not designed to

resist significant overpressure and will fail when a pressure wave

propagates into it. This latter occurrence actually can provide protection,

because upon duct failure most of the blast energy will spread out into the

space surrounding the duct. Thus, unless the ventilation ducting is very

short or can resist substantial overpressures, little blast energy will

arrive at the location occupied by personnel. Secondly, even if substantial

energies and overpressures propagate into' the plant via a ventilation

system, only personnel directly adjacent to the duct termination will be

severely affected. In the remainder of the room the pressure buildup would

be much more gradual and hence much less damaging. Thus it is believed that

the free-field overpressures which would lead to injuries of protected

personnel (inside the plant) would be of the same level as the pressures

that would severely damage the structure itself. For reinforced concrete

buildings, these pressures are, about 0.2 !4Pa (30 psi), which occurs only

near the edge of an exploding vapor cloud (200 m from the cloud edge for a

50 m cloud height). Thus such events can only occur when the cloud has

nearly drifted to the plant site, which for the worst case is about 10 km.

10.2 Ignition Vulnerability

In Chapter 7 (Table 8) it is shown that a duration of a typical thermal

exposure for targets outside the vapor cloud is 600 s. For targets inside
132

the burning vapor cloud, the durations are shorter, about 300 s, but the

flux intensities are so extreme (in excess of 200 kW/m 2 ) that all

combustibles engulfed by such a fire are expected to ignite.

For off-site fire the empirical formulas presented in the preceding chapter

are used to establish if ignition of combustibles occurs. It is assumed

that fire durations are about 600 s and that the peak heat flux (whatever

its value) persists for half of this time, i.e., 300 s. Examining typical

fire scenarios (see Chapter 5) and the variations of heat flux with distance

from the fire, these are believed to be conservative assumptions. Ignition

is postulated when the peak radiation flux at the target exceeds the

threshold value of the ignition flux for a given material and the time to

ignition is less than 300 s. For polymer ignition, the surface temperature

is computed from Eq. 34 using the peak radiation flux and an effective

duration of 300 s. The obtained value is compared with the ignition

temperature specified in the literature. At the same time, Eq. 33 is used

to check if the time to ignition for a given flux is less than the assumed

300 s duration. For easily combustible materials, the times for ignition to

occur are very short, and hence the radiation heat flux is the governing

criterion. Thus for cotton and a-cellulose, ignition may be expected at 12

and 16 kW/m 2 , respectively, regardless of the fire duration. Peak

radiation-heat-flux values of this magnitude may occur for the case of no

wind at distances up to 2 km from the spill center, or about 1 km from the

cloud edge. When the vapor cloud is allowed to drift, due to wind, the

distance from the cloud edge is more difficult to establish, but can again

be taken as about 1 km. Considering that the cloud drift may be as much as

9.8 km under strong wind conditions (8.96 m/s), the ignition hazard extends

up to 11 km from the spill source. The same holds for the pilot ignition of

wood, where it is estimated that 15 kW/m 2 of heat flux is required for

ignition to occur in less than 300 s. Spontaneous ignition of wood requires

a heat flux of about 27 kW/m 2 . A conservative distance for this flux level

from the cloud edge is again about 1 km, and the total distance from the

spill, allowing for maximum cloud drift, is 11 km.

The ignition temperatures for polymers that are representative of electrical

insulation materials are 220°C for polyvinalchloride (PVC) and 720*C for

silicon rubber [32]. With the specified conditions, i.e., 300 s peak flux

duration, it takes 17 kW/m 2 to ignite PVC and approximately 35 kW/m 2 to

ignite silicon rubber. Although these heat fluxes are for pilot ignition,

they are accepted here as being threshold values for all ignition. Since

these values are higher than for some of the other combustibles, this

assumption is not unreasonable; i.e., ignition sources will probably be

present. Again, conservative distances from the burning vapor cloud to the
133

target are about 1 km for ignition of the PVC and somewhat less for silicon

rubber. Hence, the longest distance from the spill source for ignition of

either insulating material is again taken as 11 km.

The ignition of combustibles such as wood and fibers does not in itself

constitute a hazard to any of the important power plant systems. The threat

from such fires is indirect, in that the fire may spread and eventually

endanger a plant system. However, since most important systems are housed

in substantial buildings, the likelihood of such an occurrence is small. A

direct hazard may arise from the ignition of polymer insulation of electric

cables, which are exposed to thermal loads. Should the ignition lead to

cable failure, important components in an exposed switchyard or substation

may eventually fail also. Most important among the systems that could be

affected is off-site power. Although its failure alone does not constitute

a direct safety hazard, it is the first major link in a chain of events that

can lead to such a hazard. Therefore such a failure constitutes a

significant increase in the vulnerability of the plant to other effects.

10.3 Thermal-Load Vulnerability

In the preceding section it is shown that the expected durations of LNG

fires resulting from spills on water are about 600 s. For these relatively

short time spans, it is not expected that steady-state heat-transfer

conditions will be established in any of the important plant components and

structures that are exposed to the fire thermal load. Thus, to estimate

their response to the thermal loading, one must rely on transient analysis.

Most of the vital plant systems required for shutdown and heat removal are

housed in rather substantial buildings and are therefore shielded from

direct thermal loads. It then becomes Important to ascertain if the

structures housing this equipment may be sufficiently affected by the

thermal loading to preclude its proper operation.

10.3.1 Response of Ceramic Electrical Insulators

A notable exception to this is ceramic electric Insulators, which may be

used in transmission lines and substations on the plant site. These are

directly exposed to the thermal loads, and their failure may lead to a loss

of off-site power. As indicated before, the thermal-stress analysis of such

complex geomteric bodies represents a difficult problem. Here the simple

expressions given in the preceding chapter will be used to obtain firstorder estimates of their response. Since ceramic insulators are poor heat

conductors, it may be expected that under severe thermal loading their outer
134

surface will heat up rapidly, while the interior will remain unaffected.

This in turn will result in thermal stresses, which could lead to failure by

tension toward the center of the body [57]. In the limit it is conservative

to assume that the surface temperature changes instantaneously. Using Eq.

36 or 37 together with the material parameters for porcelain (given in the

preceding chapter), it is estimated that failure may be expected for

temperature differences on the order of 200-300*C. The time required for

the surface temperature to increase depends on the thermal load. For a flux

intensity of 200 kW/m 2 , the time will vary from 4 to 10 s (see Eq. 34)); for

an intensity of 100 kW/m 2 , 17-38 s is required. It is also estimated, using

Eq. 38, that for a body with a characteristic dimension of 50 mm, e.g.,

radius of a cylinder or sphere, about 250 s is required to reach the maximum

or failure stress. This duration is an order of magnitude or more longer

than the time required to reach the appropriate surface temperature. Thus

the assumption of an instantaneous temperature rise is reasonable. Failure

should be expected at these flux levels if they persist for longer than the

indicated 250 s. This indeed is the case under the assumptions made

earlier. Flux levels of this magnitude, i.e., larger than 100 kW/m 2 , occur

only directly adjacent to the fire or in the region engulfed by fire.

Allowing for maximum vapor-cloud drift, this could occur at distances of up

to 10 km from the spill source.

At lower heat-flux values, the assumption of an instantaneous surfacetemperature rise is not valid. It is then more difficult to estimate the

effect of thermal stress on ceramic insulators, in particular, since the

heating is due to thermal radiation. To overcome this difficulty, it is

assumed that the radiation flux can be replaced by an equivalent convective

heat transfer with a constant heat-transfer coefficient h [53]. The latter

is simply estimated by dividing the incident radiant flux by the initial

temperature difference between the radiation source (flame) and the target,

i.e., the ambient temperature. Assuming the flame temperature at 1100C

(diffusion flame), the ambient temperature at 20*C, and a characteristic

dimension (radius) of 50 mm, the maximum allowable temperature difference

(surface to center) can be computed for various heat fluxes q using Eq.

39. The results are as follows:

,kWm2 h, w/im2°K Tmax' °C

50 46.3 649

30 27.8 909

20 18.5 1237

10 9.3 2207
135

The last two temperature-difference values are obviously Impossible. It is

also not very plausible that temperature differences as large as those

presented above for the first two cases can be attained with the fire

durations (600 s) considered here. This is corroborated by graphical

results for a flat-plate configuration given in Ref. 54. Thus it can be

concluded that failures for ceramic electrical insulators should be expected

only when the fire is directly adjacent to, or on, the actual plant site.

As indicated above, considering vapor-cloud drift, this results in a maximum

distance of 10 km from the spill location.

10.3.2 Effects on Steel Structures and Components

Many of the secondary structures at a typical nuclear power plant site may

be of steel-frame construction. Similarly, at some plant designs, steel

pipe bridges carrying feedwater or steam lines may be encountered. Although

failures of these structures may not represent a safety hazard in

themselves, they do contribute to the general plant vulnerability.

Because of the variety and complexity of geometric shapes of structural

steel components, it is difficult to arrive at a simple and universal means

of estimating their response to thermal loads. Another complication arises

from the fact that structural steel has a relatively high thermal

conductivity. This precludes the usage of some of the simplifying

assumptions used in the previous section; e.g., it is not reasonable to

assume, even for high heat fluxes, that the surface temperature changes

instantaneously to the value of the surroundings. The only means of making

accurate predictions of the thermal response of steel structures are

detailed numerical computations. Due to the geometrical and other

complexities of steel components, such an approach is considered beyond the

scope of the current effort. Numerical data in the literature 153,58] can

be used to obtain a first-order estimate of the temperature rise of steel

members subjected to thermal loads. Specifically, results for radiation

from a medium at a temperature of Tf to a simple slab geometry originally at

a temperature Ta can be used. This procedure again replaces the more

complex radiation problem, where the heat-transfer coefficient is estimated

by dividing the incoming radiant heat flux by the initial temperature

difference between the radiation source (flame) and the structure. In

calculating the temperature rise in the steel, the following average

material properties are used [59]:

Density p - 7724 kg/m3

Heat capacity Cp = 618 J/kg-K
136

Thermal conductivity k - 41.5 W/m*K

Thermal diffusivity a - 8.694 x 10-6 m2 /s

Typical results obtained using the numerical and graphical data from the

literature [58] are shown in Fig. 48. The temperature rise in steel plates

of various thicknesses is given as a function of heat flux, assuming a

constant flux and a duration of 600 s. The temperature values are for the

center of a plate of given thickness heated on both faces. They also apply

for a plate of half the thickness heated on one face and insulated, or

having minimal heat loss, at the other face.

In these results it was noted that, due to the high thermal conductivity,

the temperature varied little (less than 10%) across the thickness of the

steel plate, in particular at high-heat-flux values. An approximate

solution based upon the assumption of a spatially uniform (across the plate

thickness) temperature therefore suggests itself. For the equivalent

convective problem, the equation governing the response of the steel slab is

then simply

s P E .f + (42)

where if and ia represent, respectively, the heating of the plate and any

heat losses that may occur. They can be expressed as

qf - hf(Tf - T)

4a ha(Ta h -T). (43)

In Eqs. 42 and 43, T represents the temperature, t is time, s is the plate

thickness, p is the steel density, and Cp is the specific heat. Subscripts

f and a refer to the heat source (flame) and the ambient (or initial) conditions, respectively. The heat-transfer coefficients hf and ha are assumed

to be constant. The former is obtained by the procedure outlined earlier;

the latter is assumed to have a value typical for the cooling of a steel

plate in the atmosphere [58], i.e., ha = 11.35 W/m2 -K - 2 Btu/h-ft 2 o -F.

Under these assumptions Eq. 42 can be integrated analytically:

b - (a - b T)e-bt/spC a[a a p] (44)
137

1000

800

C>)

0

- 600

LU

cLU

`- 400

I--

200

0 40 80 120 160 200

HEAT FLUX, kW/rM2

Figure 48. Temperature Rise in Steel Plate vs. Heat Flux;

Flux Duration 600 s.
138

where

ff a a -h f T f + h a Ta

and

b h hf + ha

The results from Eq. 44, for a slab of 40 mm thickness (or 20 mm thickness

when heated on one side), are shown as the solid dots in Fig. 48. The

agreement with the exact solution is quite good, particularly for high heatflux values. Encouraged by this result, and also because there is

considerable doubt as to the equivalency of convective heat transfer as a

substitute for radiation, the actual problem is treated again by assuming a

uniform spatial temperature across the steel slab. When a structure is not

engulfed by fire, the radient flux is incident only on one face of the slab;

however, heat losses by reradiation and convection occur on both faces. The

equation governing the thermal response of such a steel plate can then be

written as

___ 4 SpC dT q -2[o(T 4 - T) + ha(T - T (45)

p d r -a a a

The notation is the same as used earlier, qr is the specified incoming

radiant flux. The factor 2 accounts for losses on both faces of the slab.

An analytical solution to Eq. 45 is not readily evident. Therefore a

numerical solution is obtained with the assumption of constant heat flux ir

and heat-transfer coefficient ha. Using the simplest possible integration

scheme, we can express Eq. 45 in finite-difference form as

T T+ [ /2 + aT4 + haT - (o + hT )] (46) n+l n r T + ha a o n aT~ n

p

Here At is the time step and the subscripts n and n + 1 refer to the

preceding and current time step, respectively. Figure 49 shows the

temperature rise as a function of time for two values of heat flux and two

plate thicknesses, and Fig. 50 presents the temperature increase as a

function of heat flux for flux durations of 600 and 300 s. Comparing the

results of Fig. 50 with those of Fig. 48 for the equivalent convective

problem, i.e., the 40 mi curve (20 mm if heated on one face), we see that

for heat fluxes of up to 100 kW/m 2, the agreement is good. However, at

higher heat fluxes the actual radiation solution results in higher

temperatures than the equivalent convective solution. The numerical

radiation solution is therefore used to estimate the thermal response of

steel structures. In Figs. 49 and 50, the maximum allowable temperature for
1000

800

U' 600

U400

Ui..

200

0

I-h

100 200 300 400 500 600

TIME, s

Figure 49. Temperature Rise

Constant Radiant

in Steel Plate not Engulfed by Fire;

Heat Flux.
140

1000

600s DURATION

800 - 300s DURATION

0 MAX. ALLOWABLE

TEMPERATURE

600

, 400

200 -----

0 40 80 120 160 200

RADIANT HEAT FLUX, kW/m 2

Figure 50. Temperature in Steel Plate not Engulfed by Fire as Function of Heat

Flux.
141

structural steel, i.e., 538°C (see Chapter 9), is indicated where

appropriate. For heating durations of 300 s or less, it appears (see Fig.

50) that heat fluxes up to 180 kW/m 2 can be tolerated even for very thin

steel sections (20 mm). Thicker steel sections (40 mm) can withstand

heating at similar levels for 600 s. For these longer durations, a 20 mm

plate can tolerate up to 100 kW/m 2 . Reviewing the results on fire threats

(Chapter 7), we find that these high-level fluxes occur only in the direct

vicinity of the fire. Hence, taking vapor cloud drift into account, they

may occur at distances of up to 10 km from the spill source. Since for most

real scenarios the heating duration at peak levels is much shorter than the

times indicated here (see Chapter 5), typical steel structures should be

relatively immune to the thermal loads from an LNG fire on water unless they

become actually engulfed by the fire.

To examine the response of steel components engulfed by fire, we again

assume that the spatial variation in temperature, across the steel section,

can be ignored. Heat input to the steel is now both by radiation and

convection. For the latter it is necessary to define a heat-transfer

coefficient hf. There is considerable variation in the literature as to

the appropriate value of hf. A conservative value of hf = 170 W2. K = 30

Btu/hoft 2 *"F is used for the current appliction [46]. The equation

governing the heat transfer to a steel slab engulfed by fire can be written

as

d T 4 4 spC dTdt - a(Tf - T ) + hf(Tf - T) , (47)

where Tf is again the temperature of the flame. All shielding by the

combustion gases, which could diminish the radiant flux, is neglected in

this expression. As written, Eq. 47 represents heating on one face of the

slab with the other face insulated or experiencing minimal heat loss. For

reasons of symmetry, the results obtained from the integration of this

equation are equally applicable to a slab of twice the thickness, totally

engulfed by fire, and heated on both faces. The simplest finite difference

form is

T T + [T + hfT - OT 4 + hfT)] At (48) n+1 n Lf f F n f n p

Figure 51 shows numerical results obtained for the temperature rise, as a

function of time, in various slab thicknesses. As given, the thicknesses

are for steel plates completely engulfed by fire, i.e., heated on both

faces. The results of each curve are equally applicable to slabs of onehalf the indicated thickness which are heated only on one side. The assumed
142

v

1200

1000

800

SAP

600

400

200

0 100 200 300 400 500 600

TIME ,s

Figure 51. Temperature Rise in Steel Plate Engulfed by Fire.
143

fire temperature of 1373 K is that of a turbulent diffusion flame. Again

the maximum allowable temperature for structural steel is shown in the

figure. It is clearly evident that thin steel sections can tolerate the

fire environment only for short durations. Since the maximum duration of a

vapor-cloud fire at a fixed location is about 300 s (see Chapter 7), it

appears that a steel section of 40 mm, if heated on one side, or 80 mm if

totally engulfed by the fire, is required in order to avoid serious

weakening of the structure. Most nuclear power plant structures are

expected to be contiguous without any free-standing structural components.

Thus the 40 mm slab section and heating from one side appear to be indicated

as the safe limit. Since the structure must be engulfed for some time to

experience a substantial temperature rise, it is estimated that taking cloud

drift into account, the furthest distance from the spill location at which

such a scenario is possible is about 9 km.

It is realized that considerable variability prevails in the size and

geometry of structural components, none of them being truly flat plates.

However, the results given here permit a first-order estimate of the

response and vulnerability of steel structures at nuclear power plants

subjected to thermal load from LNG fires. Using the reasonable assumption

of uniform temperature across a steel section, we can readily carry out more

accurate calculations for actual structural steel shapes. The procedure is

also easily adapted to the case of time-varying heat flux, where the latter

may be obtained from an actual spill, ignition, and fire scenario. The most

important result of the current analysis is that steel structures at nuclear

power plants are probably immune to thermal loads from LNG fires, unless the

plant is actually engulfed by the fire. In the event the later is true, a

closer examination is necessary, in particular for cases in which important

structural components have sections with thermal inertias less than that of

40 mm thick flat slab.

10.3.3 Thermal Response of Concrete Walls

Most important safety-related equipment in modern nuclear power plants is

housed inside substantial reinforced concrete buildings. The response to

the expected thermal loads of concrete structures is thus of utmost

importance in analyzing the vulnerability of such plants. Typical material

constants of concrete that are important for the thermal response are [33]:

Density p - 2082 kg/m3

Specific heat Cp - 837.4 J/kg-K
144

Thermal conductivity k - 2.77 W/m-K

The most significant of these parameters is the rather low value of the

thermal conductivity. It suggests that, unless heating persists for very

long times (hours), severe thermal response of a concrete structure will be

limited to a rather shallow layer adjacent to the heated surface. The wall

thicknesses in nuclear power plant structures vary substantially from

building to building, but are not expected to be less then 0.3 m.

Preliminary estimates for such a wall, based on data for the literature

[53], indeed indicated that for thermal-load durations typical of LNG fires,

say less than one-half hour, significant heating is limited to the vicinity

of the exposed surface.

The above findings suggest certain conservative simplifications in

calculating the response of concrete walls. For a structure engulfed by the

fire, it is therefore assumed that the exposed surface temperature changes

instantaneously to the flame temperature. Calculations of the thermal

response are then readily carried out, for any wall configuration, using

existing heat-transfer computer codes [60]. For the estimates used here to

assess the vulnerability of concrete structures, the response of a simple

slab is computed.

When the structure is not engulfed by the fire, heating is due to radiation,

and the incident heat flux can be obtained from the integrated LNG spreading

and fire model. Reradiation to the surroundings, which are at ambient

temperature must, however, be taken into account, in order to avoid a gross

overestimate of the heating effect. An existing procedure [61] for the

computation of the thermal response in a semi-infinite slab is adapted for

the prediction of the temperature rise in a concrete wall subjected to a

radiant heat flux. Figure 52 shows typical results of the temperature, as a

function of time, at the heated surface and at various depths for a constant

heat flux of 50 kW/m 2 and in Fig. 53 for a heat flux of 201.4 kW/m 2 . The

latter flux value corresponds to radiation from a diffusion flame (1373 K)

without any loss or attenuation. Already, at a depth of 100 mm, the

temperature rise even for the higher flux is quite moderate. The

calculations also show, that at a depth of 300 mm the temperature remains

essentially unchanged after 1800 s of heating. While some temperature rise

at these greater depths may be expected at later times after the radiant

flux has ceased, the increases are minimal. This is born out by the results

given in Fig. 54, which represent the response to a variable heat flux of

finite duration with the extreme peak value of 100 kW/m 2 .

Figure 55 shows the temperature rise, as a function of incident radiant heat
600

500

C.)

C

ILii

0~

Lii

I400

300

200

,1-.

100o

1800

TI MEs

Figure 52. Temperature Histories in Concrete Slab at Various Depths;

Radiant Flux - 50 kW/m 2 .
1200

-' 800

LU ta~n,-20 mm

~600

~400

I~l50 mm

200 100 mm

00 M

0 200 400 600 800 1000 1200 1400 1600 1800

TIME,s

Figure 53. Temperature Histories in Concrete Slab at Various Depths;

Radiant Flux - 201.4 kW/m2 (TR - 1373 K).

.1 0, .
CIQ

E

1000- - I200 --

800 HEAT FLUX

0j I, -- 150 ° I

" 600 - I SURFACE

I00.

"- 400 2

200 -50

00 0 -

0 500 1000 1500 2000 2500 3000

TIME,s

Figure 54. Temperature Histories in Concrete Slab at Various Depths.
1000-

3000 s

o- 800 u.T • 1800 s

LU "" •00 s

<--.= = 600 180 s (TIME)

LU. CM 400-

LU

200

0 - 111 1 1 1 1

0 20 40 60 80 100 120 140 160 180 200

RADIANT HEAT FLUX, kW/m 2

Figure 55. Surface Temperature of Concrete Slab as Function of Incident Radiant

Heat Flux.

41 .'
149

flux, at the surface of the slab and for various durations. Similar results

for a depth of 50 and 100 mm from the heated surface are given in Figs. 56

and 57, respectively. Since expected fire durations are of the order of 600

s and peak fluxes persist for considerably shorter times, the heating of the

layer of concrete adjacent to the exposed face appears to be very

moderate. The maximum allowable temperature increase for reinforcing steel

is 500*C. This temperature is only exceeded at a very shallow depth of 20

mm for an extreme heat flux (200 kW/m 2 ) and an exposure duration in excess

of 360 s. In most massive concrete structures, the first layer of

reinforcing is expected to lie at a depth of about 50 mm, for which an

exposure duration of 1660 a is needed to exceed the critical temperature.

Thus the heating appears to have little detrimental effect on the

reinforcing and thus structural integrity. Similarly, prestressing steel,

which can tolerate a temperature rise up to 400°C (see Chapter 9), should be

unaffected, since it is normally located at depths greater than 100 mm.

Some surface spalling may be expected, however, since the water-vaporization

temperature (1000C) is exceeded for larger heat fluxes at the critical depth

of 50 mm. The effect of such spalling is difficult to estimate. However,

direct exposure to the radiant heating of the reinforcing could lead to

substantial weakening of the structure. A mitigating circumstance is the

fact that, even for the severest flux (200 kW/m 2 ), a duration in excess of

300 s is required to raise the temperature to 100 0C at the 50 mm depth. At

a lesser heat flux (50 kW/m2), this time is in excess of 600 s. Thus,

considering the actual durations of the fires, it is not expected that, for

the off-site fire, the combination of concrete spalling and reinforcement

exposure to heat flux will lead to any significant degradation of the

concrete structures.

The thermal response of concrete structures engulfed by the fire is based on

the assumption of a sudden surface temperature jump to the value of the

flame temperature. Two cases are considered: a diffusion flame fire at

1373 K and the fire of a premixed LNG vapor cloud at 2148 K. Both are

illustrated by the example of a 0.3 m thick wall. Results for the former

are shown in Figs. 58 and 59. The first of these presents numerical results

[60] for the temperature rise with time at various depths from the heated

surface, and Fig. 59 is the temperature profile through the wall after a

fire duration of 1800 s. First, note that, even after this long time, the

temperature at the far face is still unaffected by the thermal exposure. It

takes in excess of 1350 s to reach the critical temperature (500*C - 773 K)

for reinforcement steel at a depth of 50 mm, and the critical temperature

(400 *C - 673 K) for prestressing steel is never reached at the 100 mm depth

within the first half-hour of fire exposure. At a shallow depth of 20 mm
V

V4

C) C) 40 

000

_C)

-jL-J

00

('.4

Lii

C)'-

60

C)~

CD~~a C> C DC

CD C) C> CD CD C

CD 00 o ';5

Do MnivUM-.
500

C-)

0

ILii

0~

w

U'

I200

RADIANT HEAT FLUX, kW/m 2

Figure 57. Temperature in Concrete Slab as Function of Incident Radiant Heat Flux;

Depth - 100 mm from Heated Surface.
152

1200 20 mm FROM HOT FACE

1000

50 mm FROM HOT FACE L- 800

600

C1.

•- 4 0 0 ....................

200 100 mm FROM HOT FACE

O'O 300 600 900 1200 1500 1800

TIME,s

Figure 58. Temperature Rise with Time in Concrete Wall 0.3 m Thick; Constant

Surface Temperatures - 1373 K, Initial Temperature - 293 K.
153

1400

LU

1-

300

DISTANCE FROM HEATED FACEmm

I Figure 59. Temperature Variation Across 0.3 m Thick Concrete Wall at Time t=1800 s;

Constant Surface Temperature - 1373 K, Initial Temperature - 293 K.
154

the 500*C limit is reached within 230 s. However, as pointed out earlier,

the concrete cover over the reinforcement is expected to be substantially

thicker than the 20 mm. Again spalling of concrete is to be expected, but

considering the approximately 300 s duration of the fire (see Chapter 7), no

significant structural degradation is anticipated. In general, this case is

very similar, both in thermal load and structural response, to the case of

high radiant heat flux.

Finally, the burning of a premixed vapor cloud on the plant, site is

analyzed. The temperature rise with time at various depths from the heated

surface is presented in Fig. 60, and the temperature profile through the

wall after one-half hour is given in Fig. 61. Again, the far face of the

wall is seen to be unaffected. Similarly, the critical temperature for

prestressing steel (400*C) is never reached at a depth of 100 mm or

larger. The 500*C limit for reinforcing steel now obtains in 600 s at a

depth of 50 mm. Thus, if fires of this duration should occur, then

significant structural effects might be expected. However, this is not a

likely event for a premixed cloud. Although the time between UFL and LFL

may exceed the 600 s under some circumstances very slow flame motion and

cloud drift, relative to the stationary site, would be required to achieve

such durations. In addition, the slow cloud drift would limit the

occurrence of such an event to the close proximity of the spill location.

Thus this is not considered a viable scenario. Again, the spalling

conditions at a depth of 50 mm are reached in a time somewhat in excess of

200 s. At shallower depth this occurs much earlier. Hence, spalling is a

likely occurrence, and the remote possibility exists that a combination of

spalling and heating may ultimately lead to some structural degradation.

The effect of fire engulfment followed by a rapid decay of surface

temperature is illustrated in Fig. 62. It is seen that at some distance

from the heated surface (100 mm), the maximum temperature occurs long after

the surface temperature has decayed. After a long time, the entire slab

reaches an essentially uniform temperature.

Based on the foregoing, it is ascertained that concrete structures can

sustain the effects of LNG fires very well. Although some spalling of

concrete is to be expected, no structural degradation is in general

anticipated. Even the spalling can only be significant when the heat fluxes

are large, which implies on-site or near-site fires. Considering cloud

drift, such effects may occur out to distances of 9-10 km from the spill

source.
155

2000

20 mm FROM HOT FACE

1600

~ 10050 mm FROM HOT FACE 1200AC

800

400

......mm 10 FROM HOT FACE

0 I I I 1

0 300 600 900 1200 1500 1800

TIME,s

Figure 60. Temperature Rise with Time in Concrete Wall 0.3 m Thick; Constant

Surface Temperature - 2148 K, Initial Temperature - 293 K.
156

LLU

LuJ

QL

LUJ

2400

2000

1600

1200

800

101°

Oo

50 100 150 200 250 300

DISTANCE FROM HEATED FACE,mm

Figure 61. Temperature Variation across Concrete Wall 0.3 m Thick at Time t=1800 s;

Constant Surface Temperature = 2148 K, Initial Temperature = 293K.
157

2400

LiJ

~ 1200-

S800

LiI"-

400

0-0

Figure 62. Temperature

600 1200 1800 2400 3000 3600

TIME,s

Rise with Time in Concrete Wall 0.3 m Thick; Initial

Temperature - 293 K.
158

10.4 Air Blast Vulnerability

A detailed structural response analysis of the important nuclear power plant

buildings subjected to air blast from an LNG vapor-cloud explosion is beyond

the scope of the current effort. Therefore, the structural resistance

criteria summarized in Table 12 will be used to assess the power plant

vulnerability. As pointed out earlier (Chapter 9), the expected blast

durations are sufficiently long to permit the employment of simple

overpressure criteria. The distance to a particular overpressure level from

the edge of the exploding vapor cloud is a function of cloud height (see

Chapter 6), which realistically may be expected to vary from 25 to 75 m.

The following estimates of distance are based on the average value of cloud

height, i.e., 50 m.

Starting with the most severe blast scenario, i.e., an actual cloud

detonation on the plant site that produces an overpressure of 1.45 MPa, it

can readily be seen from Table 12 that all the plant structures would be

extremely

vulnerable. In fact, destruction or severe damage may be expected for all

buildings. The consequences of such an event are difficult to ascertain,

but significant radiation release is quite probable.

A more moderate scenario is one involving volume explosions on the plant

site. Again this requires the vapor cloud to drift to the site.

Infiltration by the premixed LNG vapor into nearly completely confined

spaces and subsequent ignition may lead to this event. The overpressure

generated by a volume explosion is approximately one-half the detonation

value [29], or 0.73 MPa in the present case. Again none of the plant

structures are designed to withstand such an internal overpressure. Thus,

severe damage would result. All structures that are not closed off from the

atmosphere, are susceptible to such an event. Since LNG vapor, even in the

premixed state, is heavier than air, infiltration by gravity into any

opening engulfed by the cloud is to be expected. The containment is

generally closed to the atmosphere and is therefore not a likely candidate

for such an event. The most vulnerable structures appear to be concrete

block houses used for the protection of diesel fuel and condensate water

tanks against the effects of tornados. The intake structure could also be

affected in this manner, as well as areas of the auxiliary building, e.g.,

battery rooms if not closed off or positively ventilated. The only

mitigating circumstance is the fact that some plant openings may not be

reached by the vapor mixture because of the limited cloud height. Both of

the described events, i.e., on-site detonations and volume explosions,

require LNG vapor cloud drift to the site location. Under strong wind
159

conditions (8.96 m/s), it is estimated that the farthest distance from the

location of spill at which these events can occur is about 9.5 km.

The effects from a volume explosion could serve as the initiation source for

a general vapor-cloud explosion, thus involving additional structures.

Another secondary effect of consequence could be the debris problem.

Although debris is generated whenever blast pressure results in significant

structural damage, the problem could be particularly severe for volume

explosions. The reason for this is that in such an event, rather

substantial concrete structures may

be blown apart, resulting in very massive high-velocity missiles. The

impact of such missiles may lead to secondary damage of important safetyrelated components and structures. Particularly vulnerable appear to be

exposed pipe lines and electrical equipment, such as distribution stations

and off-site power installations.

The maximum overpressure level to which reinforced concrete containment

buildings can be exposed without sustaining significant damage is estimated

at 0.345 MPa (see Table 12). For substantial reinforced concrete

structures, such as the auxiliary building, the corresponding overpressure

level is 0.207 MPao The distances from the cloud edge to these overpressure

levels are, respectively, 2.5 and 4.0 cloud-height units or 125 and 200 m.

Since the maximum cloud drift is 9.8 km (see Table 8), the maximum distance

from the spill source can, for both thresholds, be taken as about 10 km.

Any substantial damage to the concrete structures, and particularly to the

containment building, would indeed be serious.

Moderate damage to heavy industrial-type structures, of which the turbine

building may be representative, is estimated to occur at overpressures of

about 0.083 MPa (see Table 12). Such overpressure levels can occur at

distances of 450 m (9 cloud-height units) from the edge of the exploding

vapor cloud. Accounting for cloud drift, the maximum distance from the

spill source is 10.3 km. Although damage to such buildings by itself may

not constitute a safety hazard, it is again possible to envision secondary

effects, caused by debris or other actions, which may result in such a

hazard. In this category would be any disruption of the cooling system.

Since main steam lines often run exposed through the turbine building, such

a possibility cannot be excluded.

A similar possibility exists at even lower overpressures such as 0.048 MPa

(see Table 12), when main coolant lines (feedwater and stem) run exposed

between buildings as is the case in some older existing plants. This

overpressure persists at a distance from the cloud edge of 850 m (17 cloud-
160

height units) or including cloud drift about 10.7 km from the spill source.

A loss-of-coolant accident is always considered a grave matter, but may be

particularly dangerous when coupled with failures of other vital

equipment. For air blast overpressure, this may indeed be the case, because

electrical power systems may sustain damage of overpresures as low as 0.035

MPa (see Table 12). The vulnerable items are substations, switchyards, and

high-voltage lines, the damage of which can result in the loss of off-site

power. Since the off-site power supply is more susceptible to blast than

large piping runs or the turbine building, failure of power supply should be

expected when the latter items sustain damage. The distance from the cloud

edge at which electrical systems may be damaged is estimated at 1100 mm (22

cloud-height units). Including cloud drift the maximum distance from the

spill source for such an event is 10.9 km.

Damage to other buildings, systems, and components can result at still lower

blast overpressures (see Table 12). The loss or damage of none of these

items, however, appears to pose a direct or indirect threat to the safety of

the power plant. The dynamic pressures, and hence, drag forces at these

pressure levels, i.e., 0.035 NPa or below, are comparable with those

resulting from the design tornado [62]. Hence, at these overpressures, the

expected damage to power plant structures should be equivalent to that

resulting from a toranado.

As with plant personnel, some potential for damage exists due to the

propagation of pressure waves into the interior of the plant via a breached

ventilation system, access hatch, or door. The overpressure levels for

breaching to occur and the resulting damage are so design-dependent that it

is not possible to make any meaningful generic vulnerability estimates. The

purpose of it being mentioned here is simply to indicate that this problem

must be considered and to point out that approaches and methodologies exist

[63,64] to make reasonable predictions of the blast-wave propagation into

the plant interior in specific cases.

10.5 Summary of Vulnerabilities

Only a limited number of hazardous effects arising from a large LNG spill on

water were considered In the generic vulnerability analysis of the preceding

sections. Many other vulnerabilities can probably be identified for

individual power plants. This, however, requires the consideration of

specific plant design and siting details and will not be pursued here.

One additional generic question that may be asked is how well a typical
161

plant can cope with the overall thermal-loading incident on the plant or a

specific building. That this is a legitimate question can be readily seen

by comparing the cooling or heat-rejection capability provided in a typical

containment cooling (air-conditioning) system with the heat received by the

building from a severe fire. For typical modern PWR designs, the maximum

instantaneous containment cooling is about 7 x 104 kW [38]. On the other

hand, the instantaneous heat load on the building, when engulfed by a

massive LNC fire, can readily be 200 kW/m 2 , or for typical containment

building dimensions (diameter 30 m, height 50 m), a total load of about 10 6

kW. Thus the ratio of the instantaneous heat load to the cooling capacity

is nearly 15.1. It was shown earlier that the durations of intense LNG

fires are rather short, say 600 s. Assuming an average concrete wall

thickness of 0.75 m for the containment building, we find that, even

neglecting all heat losses, the building possesses sufficient heat capacity

to limit the average temperature rise in the wall to less than 85*C. This

is an acceptable value, and hence it is thought that no significant hazard

arises from the overall heating of concrete buildings, which house most of

the critical systems of a modern nuclear power plant.

To facilitate an overall evaluation of nuclear power plant vulnerability,

Table 13 summarizes the important generic hazards due to an LNG spill on

water. The hazards are listed together with their threshold vulnerability

distances, i.e., the maximum distances from the spill source at which a

plant may be vulnerable to a particular effect. Items marked with an

asterisk are considered to constitute a direct safety hazard. Essentially

all these items pertain to blast effects on the power plant. Thus, if LNG

vapor cloud explosions are possible and occur in the vicinity of a nuclear

power plant, then the safety of the plant may be endangered.

On the other hand, the effects produced by fire and thermal loading on the

nuclear power plant seem, in general, to be tolerable. Although specific

safety-related' components might be affected, the plant's backup and

engineered safety system should permit a safe shutdown and subsequent heat

removal. The major reason for this limited effect of thermal loads is the

short duration of the expected fires. Heat loads similar in magnitude to

those produced by the LNG fire, but of much longer duration, could not be

tolerated as well.

The data in Table 13 establishes also another very important point; namely,

the vulnerability threshold distances for all the considered hazards are

approximately the same. Using a single value of about 10.0 km for all of

them would result in very little error, at least for the most severe

effects. The reason for this is twofold. First, the thermal and blast
162

TABLE 13

SUMMARY OF NUCLEAR POWER PLANT VULNERABILITY

Plant System, Threat and Effect (Comment) Standoff Distance, km

Personnel

Exposed - Fire - Skin Burns (4 kw/m 2 ) 13.0

Exposed - Air Blast (.0345 MPa) 11.0

Protected - Air Blast (.2069 MPa) 10.0*

Ignition

All combustibles including electric insulation 11.0

Thermal Loads

Ceramic insulators (100 kw/m2) 10.0

Steel Structures - Not engulfed (100 kw/m 2 ) 10.0

Steel Structures - Engulfed (40 mm thickness safe) 9.0

Concrete Structures - Not engulfed (200 kw/mIm2 ; 360 s) 10.0

Concrete Structures - Engulfed (300 s safe, no spalling) 9.0

Air Blast

Detonation on Site (1.45 MPa - Severe Damage all Structures) 9.5*

Volume Detonation (Severe Damage-Concrete Bunkers) 9.5*

Containment Building (.3448 MPa) 10.0*

Concrete Structures (.2069 MPa) 10.1*

Heavy Industrial Structures (.0828 MPa) 10.3*

Exposed Piping Systems (.0483 MPa) 10.7*

Electrical Equipment (.0345 MPa) 11.0*

Direct Effect on Safety System
163

effects are only severe enough to affect the plant in the close vicinity of

the fire/explosion source, I.e., at distances of about 12.0 km from the edge

of the LNG vapor cloud. Second, the overall vulnerability threshold

distance (maximum distance from spill source) is dominated by the LNG vapor

cloud drift. The cloud drift distance represents, in general, 90% or more

of the total distance.

On the basis of Table 13 and the foregoing analysis, it is also observed

that, among the safety-related plant systems, exposed electrical equipment

and specifically off-site power appears to be most vulnerable to the LNG

spill hazards. Ignition of cable insulation, failure of ceramic insulators,

and damage by relatively low overpressure blast waves are considered here.

Other damage mechanisms seem also plausible, e.g., overheating and

subsequent failure of oil-filled transformers. From a safety point of view,

off-site power is primarily important for long-term decay heat removal. As

mentioned earlier, the failure of off-site power alone does not constitute a

direct safety hazard. However, in combination with other failures, e.g.,

emergency power diesel failures or loss of all diesel fuels,. the loss of

off-site power, for any extended period of time, may result in a grave

situation.
164

11. Discussion, Conclusions, and Recommendations

The analysis outlined in Chapters 3-6 of this report permits an overall

description of the phenomena associated with a large spill of LNG on

water. The models are based on simple engineering analysis and rely on a

global approach to arrive at a reasonable definition of a complex process.

Although the limitations of the models are recognized, it is believed that

the most salient features of the problem have been included. Sensitivity

studies were used to establish the importance of various parameters.

One major achievement of this effort was the development of a combined LNG

spill, spreading, and fire model. This approach permits the simultaneous

and coupled computation of all the major aspects of an LNG fire. A tool is

thus available to carry out detailed analyses of the fire threat for any

postulated scenario and specific power plant arrangements and sites.

Another very important result of the analysis effort is the development of

an approach for the prediction of air blast in the vicinity of an exploding

"pancake-like" vapor cloud (Chapter 6). The method has its shortcomings in

that it is not a unified procedure and depends on both simplified analytical

calculations and the use of two-dimensional hydrocodes. However, the

obtained results are scalable with cloud height and therefore are applicable

to a broad class of problems. Although the limitations of the procedure are

recognized, it is believed that the predictions are a much better

approximation of the air blast field in the vicinity of the vapor cloud than

those, from the normally employed TNT equivalency concept, which is based on

a hemispherical vapor cloud-shape assumption.

In carrying out the effects and vulnerability analysis (Chapter 10), we use

simple susceptibility criteria for the various plant systems and

components. Also, the analysis itself is much less detailed than the

descritpion of the LNG spill environment. This approach was taken because

of limitations on the program effort and also because detailed response

analyses are thought to be more appropriate for site specific studies. Most

of the blast-response vulnerabilities are established by comparing

overpressure criteria for various plant systems with the blast

environment. Simple calculations are carried out to define the thermal

response.

Based on this analysis, it is found that the thermal loads resulting from a

large LNG fire on water can in general be tolerated by the important safetyrelated plant systems. This is primarily due to the short durations of the

expected fires. On the other hand, it is shown that nuclear power plant
165

systems are very vulnerable to blast effects from an exploding LNG vapor

cloud. It is also concluded that all serious effects are limited to the

immediate vicinity of the fire/explosion source, i.e', the LNG vapor

cloud. Because of this, the standoff distance, i.e., the distance from the

LNG spill location, which is required to ensure complete safety of the

important nuclear power plant systems, is primarily dependent on the windinduced vapor-cloud drift. For most of the severe hazardous effects and

under the assumption of a strong wind (8.96 m/s - 20 mph), the required

standoff distance in the downwind direction is found to be somewhat in

excess of 10 km.

The results of the vulnerability analysis are all based on the predictions

of the models that describe the LNG spill environment. To gain some

confidence as to the correctness of the results, verification of these

models is required. This would best be accomplished by comparison with

experimental data. Although some small-scale tests of LNG spills on water

have been conducted [6,9,101, the data is either inadequate or not

documented well enough to provide the needed verification. The validity of

scaling this information to the large LNG spill is also questionable. The

other approach to verification, of the methods used here, is by comparison

with much more detailed analytical and computational procedures. One such

comparison with an elaborate computer code [8] has been carried out and

found acceptable. However, much more is required to obtain true

verification of the employed methodologies. Because of their dominant

importance in defining' plant vulnerability, two aspects in particular

deserve precise scrutiny. These are: (i) the wind-drift predictions for

the LNG vapor clouds and (ii) the assumption of LNG detonability

(explosion).

Nuclear power plants are only affected by the hazards from an LNG spill when

the vapor cloud Is directly on the plant site or immediately adjacent to

it. This suggests that safety measures may be taken to mitigate at least

some of the hazardous effects. Since nuclear power plants are not likely to

be located in the immediate area of the LNG spill source, it should be

possible to institute a warning system that will make the operating

personnel aware of the impending danger. This in turn will permit some

mitigating actions.

Actions that come to mind are plant isolation and elimination of possible

ignition sources. The first of these would involve the closing of all

dampers, hatches, doors, etc., to prevent the ingress of combustible

mixtures into the plant interior. Shutdown of the plant and the

deenergizing of all unnecessary power sources, such as electrical equipment
166

and diesel engines will remove many of the possible ignition sources, thus

reducing the hazards further. Isolation of the plant from the exterior

should be quite effective against fire hazards. However, with current plant

designs it is not believed that isolation will prevent the propagation of

air blast pressure waves into the plant interior, e.g., via ventilation

shafts.Effective isolation against air blast can only be provided by blasthardening of the structures and components, as in military protective

design. Although this approach does not appear practical for the entire

plant complex, it may be applicable to selected plant systems.

As pointed out earlier, the most vulnerable safety-related system at a

typical plant site appears to be the off-site power supply. It therefore is

a prime canditate for blast hardening, e.g., by providing a completely

underground power supply to the plant. Multiple redundancy of safety

systems is another approach to reducing nuclear power plant vulnerability.

Finally, the best protection is provided by sufficient standoff distance

from the LNG spill source. Based on the current analysis result, this

distance is in excess of 10 km and may not always be achievable for specific

plant sites.

It should also be remembered, that the conclusions of the current study are

based on a purely deterministic analysis. Thus worst-case assumptions are

often made, and no consideration is given to the likelihood of the

occurrence of an event. In applications to actual power plants, probability

considerations must be included, in particular when a preliminary evaluation

of the effects indicates significant hazard levels. In many cases, the

stochastic analysis may indicate that the risks are still acceptable.

The methodology and results of the current study represent a significant

improvement in the treatment of the hazard problem arising from large LNG

spills on water. The method can be readily adapted to the analysis of

hazards in specific cases. It is particularly valuable as a rapid and

inexpensive screening procedure when many possible spill scenarios and

environmental conditions must be considered. Coupled with probability

analysis, it can become a valuable tool in assessing the risks to nuclear

power plants arising from large LNG spills on water transportation routes.

Verification of the results in specific cases' can be obtained by comparing

some aspects of the predictions with more detailed and precise deterministic

procedures; e.g., cloud development and drift can be compared with the

results of precise hydrocode calculations.

The main improvement in the accuracy of the prediction for specific cases

can be achieved by introducing more elaborate and detailed structural and
167

thermal-response analysis. This can be readily achieved with existing

procedures and computer codes. Improvements in the descrption of the LNG

spill environment can also be introduced. Further sensitivity and parameter

studies may provide a better definition of the various problem constants

used in the analysis. It is also believed that the methods are readily

adaptable to treating the influence of solid boundaries, such as shore

lines, on the development of the LNG pool, vapor cloud, and fire scenario.

Finally the methods developed here for the treatment of LNG spills on water

can -be extended to land spills and may also be used to analyze spills of

other hazardous materials. Specifically any material whose spill results in

the formation of negatively buoyant vapor clouds may be treated by the

approach used here for LNG.
168

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174

APPENDICES

The appendices of this report contain the detailed phenomenological studies

which lead to the development of global models describing all aspects of a

large LNG spill on water. Since these models have been amply outlined in

Chapters 3, 4, 5 and 6 and because of the large volume of the detailed

material, the appendices are here included in microfiche form.

The studies presented in the appendices were conducted for Argonne National

Laboratory (ANL) by the IIT Research Institute (IITRI) Chicago, Illinois.

Each of the appendices addresses different aspects of the LNG spill

phenomenology and corresponds to an IITRI Technical report submitted to ANL

as follows:

Appendix A

T. V. Eichler, A. H. Wiederman, R. Pape, -Study of Liquid Natural

Gas (LNG) Spill Phenomenology on Water," IITRI Final Report J6456,

December 1978, Revised May 1979.

Appendix B

T. V. Eichler, A. H. Wiedermann, R. Pape, "Study of Liquid Natural

Gas (LNG) Spill, Dispersion, and Combustion Phenomenology," IITRI

Final Report J6481, May 1980.

Appendix C

T. V. Eichler, A. H. Wiedermann, R. Pape, "Study of Explosions and

Fast Deflagrations of LNG Vapor," IITRI Final Report J6517,

September 1980.
175

Distribution for NUREG/CR-2490 (ANL-CT-81-17)

Internal:

E. S. Beckjord

C. E. Till

R. S. Zeno

G. S. Rosenberg

M. G. Srinivasan (2)

H. C. Lin

J. B. van Erp

C. A. Kot (35)

ANL Patent Dept.

ANL Contract File

ANL Libraries (2)

TIS Files (3)

External:

NRC, for distribution per RI and XA (220)

DOE-TIC (2)

Manager, Chicago Operations Office, DOE

President, Argonne Universities Association, Argonne, Ill.

Components Technology Division Review Committee:

A. Bishop, U. Pittsburgh, Pittsburgh, Pa. 15261

F. W. Buckman, Consumers Power Co., 1945 Parnall Rd., Jackson, Mich. 49201

R. A. Greenkorn, Purdue U., West Lafayette, Ind. 47907

W. M. Jacobi, Westinghouse Electric Corp., P. 0. Box 355, Pittsburgh,

Pa. 15230

M. A. Schultz, 124 Lakeshore Drive, North Palm Beach, Fla. 33408

E. E. Ungar, Bolt, Beranek and Newman, Inc., 50 Moulton St., Cambridge,

Mass. 02138

J. Weisman, U. Cincinnati, Cincinnati, 0. 45221

R. M. Bernero, Div. Risk Analysis, USNRC, Washington (50)