131-04 

GEMS 1288209 

Consequence Assessment Methods for 

Incidents Involving Releases from 

Liquefied Natural Gas Carriers 

May 13, 2004

This work was performed by ABSG Consulting Inc. for the Federal Energy Regulatory 

Commission under contract number FERC04C40196. 
NOTICE 

This report was prepared by ABSG Consulting Inc. (ABS Consulting) solely for the benefit of 

the Federal Energy Regulatory Commission (FERC). Neither ABS Consulting, FERC, nor any 

person acting in their behalf makes any warranty (express or implied), or assumes any liability to any 

third party, with respect to the use of any information or methods disclosed in this report. Any thirdparty recipient of this report, by acceptance or use of this report, releases ABS Consulting and FERC 

from liability for any direct, indirect, consequential, or special loss or damage, whether arising in

contract, tort (including negligence), or otherwise. 

ABS Consulting and its employees, subcontractors, consultants, and other assigns cannot, 

individually or collectively, predict what will happen in the future. We have made a reasonable 

effort to review the available consequence assessment methods for some types of incidents involving

liquefied natural gas spills and to recommend modeling approaches for use by FERC. Appropriate 

application of the presented methods will require selection of scenarios to model and selection of 

some site-specific and scenario-specific modeling parameters. Such application of the models should

be done by an experienced consequence analyst. However, as described in this report, the 

recommended methods have limitations, including uncertainty in the results that they produce. Even

when applied appropriately, the recommended methods will provide only estimates of the 

consequences of postulated scenarios involving LNG spills on water. ABS Consulting accepts no 

liability for any adverse impacts stemming from use of information in this report. 

ii 

 
EXECUTIVE SUMMARY 

Liquefied natural gas (LNG) has been transported by sea since 1959 in specially designed 

LNG carriers. These vessels have a remarkable safety record and provide an essential link in the 

movement of LNG from production locations to consumer locations. However, stakeholders 

recognize that there are possibilities for some serious incidents involving LNG carriers, 

particularly in light of increased awareness and concern about potential terrorist actions. 

The Federal Energy Regulatory Commission (FERC) sponsored this study with the goal of 

identifying appropriate consequence analysis methods for estimating flammable vapor and 

thermal radiation hazard distances for potential LNG vessel cargo releases during transit and 

while at berth. 

This work considers the flammable vapor and thermal radiation hazards created by

unconfined LNG spills on water resulting from an LNG cargo release. This includes review of 

literature on experimental LNG spills and on consequence assessment methodologies that are 

applicable to modeling of incidents involving LNG spills on water. 

The key modeling issues addressed in this report are: 

• Rate of release of LNG from a ship 

• Spread of an unconfined pool on water 

• Vapor generation for unconfined spills on water 

• Thermal radiation from pool fires on water 

• Distance for flammable vapor dispersion following spills on water 

• Rapid phase transitions (RPTs) 

• Effects of thermal radiation on people and structures 

Where possible, methods are recommended for assessing carrier spills on water; however, in 

the case of RPTs no available consequence methods could be identified. 

The descriptions of the recommended analysis methods are left to the body of this report. In 

general these methods can provide only rough estimates of the magnitude of effects for incidents 

involving large LNG releases on water. This is typically the case with consequence assessments, 

and it is important to keep in mind that the recommended methods cannot provide precise 

estimates of effects because of variability in actual incident circumstances as well as uncertainty

inherent in the methods used. 

In the particular case of the methods of interest here (i.e., methods for large releases from 

LNG carriers), some important issues include: 

• No release models are available that take into account the true structure of an LNG 

carrier, in particular the multiple barriers that the combination of cargo tanks and the 

double hulls in current LNG carriers provide 

• No pool spread models are available that account for wave action or currents 

iii 

 
• Relatively few experimental data are available for validation of models involving LNG 

spills on water, and there are no data available for spills as large as the spills considered

in this study

In making the modeling recommendations in this report, our goal was to select methods that 

provide the most accurate estimates possible; however, we recognize the limitations of the

models we suggest. As a result of these limitations, the project team made some selections that 

we believe lead to conservative estimates (i.e., tend to overestimate the consequences of LNG 

releases). Clearly there is an opportunity to develop pool spread models that consider more

realistic analysis of the spill behavior on the water surface. In the long term, additional research 

will need to be performed to develop more refined models, and additional large-scale spill tests 

would be useful for providing better data for validation of models. However, definition of the 

specific research tasks and objectives needed to accomplish that research was not in the scope of 

this project.

No release models were identified that account for the multi-hull structure of an LNG carrier 

and the physics of a release of cryogenic LNG. In the absence of models specifically for LNG 

carriers, the orifice model is recommended. 

Several models for spreading of pools on water were found in the literature, most of which 

are based on the assumption that spreading is dominated by the balance of the gravitational 

spreading force and inertial resistance to spreading. One methodology, developed by Webber

and his colleagues (as described in TNO 1997), also accounts for resistance to spreading as a 

result of frictional forces. Our analysis of these methodologies indicates that frictional effects 

are important for scenarios involving large releases in short periods of time. Therefore, 

Webber’s methodology is recommended. 

For a given hole size (or release rate), the upper limit on pool spread depends on the duration 

of release. For long duration releases, the pool will spread to the point where evaporation rate 

(or burning rate in case of a fire) matches the release rate into the pool. This is based on a simple

mass and heat balance for the pool, and effectively defines an upper limit for long-term releases. 

For short duration releases, the dynamics of the release are more important in identifying the

upper limit. The pool spread is slowed by inertial and frictional effects but is only stopped once 

the pool is thin enough for surface tension effects to become important. During the spread, 

evaporation reduces the volume of the pool, and therefore also slows the spread. As a result of 

these dynamics, it is difficult to establish an upper limit for short duration releases without 

modeling the full spread process. 

Two methods are identified for estimating heat flux to a boiling LNG pool on water: (1) 

select a value estimated from spill tests or (2) calculate a value based on heat transfer theory. 

Based on review of the literature, it is recommended to estimate heat flux using heat transfer 

theory. Key reasons for this choice are that (1) the experimental data show significant spread, 

(2) the experimental data for heat flux were typically calculated from other data (rather than

measured directly), and (3) a viable theoretical model from heat transfer theory is available to

help fill the gap in experimental data. 

iv 

 
Available pool fire models for hazard assessment purposes are predominantly based on the 

solid flame model. The fire is represented as gray-body emitter, and the geometry is 

approximated by an upright cylinder that is tilted as a result of wind. The point source model is 

also sometimes used. In this model, energy generated by combustion is assumed to emanate 

from a single point at the center of the pool fire. The calculated energy is multiplied by a 

fraction that accounts for the fact that only part of the energy will be emitted as thermal 

radiation. In some circumstances the point source model can be appropriate, but its chief 

disadvantage is that it does not produce valid estimates for receptors close to the fire. Therefore,

the solid flame model is recommended. 

The extent to which people are injured by exposure to thermal radiation depends on both the 

incident heat flux and the exposure time. A variety of data are available for estimating effects on 

people, including data from experiments with humans and animals and review of historical data. 

Section 2.7.1 summarizes some of the injury criteria developed based on these data. Like effects

of thermal radiation on people, effects on structures also depend on incident heat flux and the 

exposure time. With structures, effects also depend strongly on the materials of construction 

(e.g., wood, steel, concrete). Section 2.8 presents some criteria from various sources for 

structural damage. 

The body of the report also presents some example consequence calculations, which are 

intended to illustrate use of the various analysis methods recommended. Input parameter values

(e.g., atmospheric conditions, surface roughness) were chosen to allow comparison with other 

methods (Fay [2003], Lehr [2004], and Quest [2001]) or to match values specified by federal 

regulations for LNG facility siting (49 Code of Federal Regulations [CFR] 193). When using the

presented analysis methods for evaluating a specific facility, input parameter values based on 

actual site conditions should be used. 

For the example dispersion calculations presented in this report, the DEGADIS software was 

used. This model accounts for dense-gas effects and was originally developed for simulation of 

cryogenic flammable gas dispersion, particularly for LNG. It has been validated against a wide 

range of available laboratory and field test data. 

It is also important to note that this study addresses the potential consequences of large scale 

LNG cargo releases without regard to the sequence of events leading to such an incident or their 

probabilities of occurrence. As such, this report does not and was not intended to provide a 

measure of risk to the public. A thorough risk assessment would consider both the probabilities 

and consequences of hazardous events. And finally it should not be assumed that the levels of 

hazards presented in this study are the assured outcome of an LNG vessel release, given the 

conservatisms in the models and the level of damage required to yield such large-scale releases.

Consistent with the scope of work defined by FERC, this report is highly technical and is 

focused on the currently available methods for modeling potential releases from LNG vessels. In 

the scientific community, there is room for debate and different opinions about the most 

appropriate way to model possible outcomes of incidents. The state of the science will continue

to evolve, leading to ongoing iterations of better models/approaches in coming years. However, 

practical and defendable guidance is needed now to help understand the currently available 

modeling methods, including their limitations, so that reasonable decision-making processes are

v 

 
used. This work ultimately recommends practical models/approaches (with appropriate caveats)

for use in guiding facility siting decisions. 

While studying the results of this report, readers should keep the following key points in 

mind: 

Models have limitations. The recommended models/approaches represent a reasonable set of 

tools to aid decision making. However, it should be recognized that modelshave an inherent 

level of uncertainty, as described in this report. In some cases, the scientific community’s 

understanding of the fundamental physical phenomena is better than its ability to simulate such

phenomena. Also, models cannot make decisions for us. They inform decision makers, who 

must integrate information on many different factors in their decision-making process. 

Consequence assessment is only one piece of the risk picture. Understanding risk requires an 

understanding of (1) what can go wrong, (2) what the consequences might be, and (3) how likely 

the losses are to occur. This report focuses only on consequence modeling of potential release

scenarios, not how likely such scenarios are to occur. Decision making related to scenario risks 

should be considered in the context of both potential consequence and expected likelihood (or 

frequency). With regard to potential attack scenarios, the expected likelihood is a function of 

both threat (the likelihood that someone would try to carry out a specific type of attack) and

vulnerability (the likelihood that a specific type of attack would be successful and produce the 

expected consequences of concern). As noted in the report, LNG vessel and associated facility

operations are highly regulated and closely monitored/controlled by authorities, so many layers 

of protection exist against losses. The dependability of these layers of protection was not 

addressed in this project, but are important considerations in understanding the total risk picture. 

Risk perception and risk acceptance are complex issues. How individuals and groups of 

individuals perceive or accept risks depends on many factors, which are often subjective with no

clear right and wrong answers. Even when very precise/certain risk information is available, 

different people often react to the information in different ways. This project made no attempt to 

place value judgments on what risks people should or should not accept. 

vi 

 
TABLE OF CONTENTS 

EXECUTIVE SUMMARY ................................................................................................................iii 

LIST OF TABLES AND FIGURES..................................................................................................ix

ACRONYMS........................................................................................................................................ x

SCIENTIFIC SYMBOLS AND UNITS............................................................................................xi

1 INTRODUCTION ......................................................................................................................... 1

1.1 BACKGROUND....................................................................................................................... 1

1.2 SCOPE OF WORK ................................................................................................................... 1

1.3 PROPERTIES OF LNG ............................................................................................................ 2

1.4 LNG HAZARDS....................................................................................................................... 2

1.4.1Fire Hazards...................................................................................................................... 3

1.4.2Explosions.........................................................................................................................5

1.4.3Rapid Phase Transition (RPT) .......................................................................................... 5

1.4.4Cryogenic Effects..............................................................................................................5

2 CONSEQUENCE ASSESSMENT METHODS.......................................................................... 7

2.1 RATE OF RELEASE OF LNG FROM A SHIP....................................................................... 7

2.1.1Review of Modeling Approaches...................................................................................... 7

2.1.2Modeling Recommendations ............................................................................................ 8

2.2 SPREAD OF AN UNCONFINED POOL ON WATER........................................................... 9

2.2.1Review of Modeling Approaches...................................................................................... 9

2.2.2Modeling Recommendations .......................................................................................... 13

2.3 VAPOR GENERATION FOR UNCONFINED SPILLS ON WATER ................................. 14

2.3.1Review of Modeling Approaches.................................................................................... 14

2.3.2Modeling Recommendations .......................................................................................... 16

2.4 THERMAL RADIATION FROM POOL FIRES ON WATER............................................. 19

2.4.1Review of Modeling Approaches.................................................................................... 20

2.4.2Modeling Recommendations .......................................................................................... 22

2.5 FLAMMABLE VAPOR DISPERSION FOLLOWING SPILLS ON WATER..................... 25

2.5.1Review of Modeling Approaches.................................................................................... 25

2.5.2Modeling Recommendations .......................................................................................... 25

2.6 RAPID PHASE TRANSITIONS ............................................................................................ 26

2.6.1Literature Review............................................................................................................ 26

2.6.2Modeling Recommendations .......................................................................................... 27

2.7 EFFECTS OF THERMAL RADIATION ON PEOPLE ........................................................ 29

2.7.1Review of Data and Effects Models................................................................................ 29

2.7.2Selecting Levels of Concern ........................................................................................... 32

2.8 EFFECTS OF THERMAL RADIATION ON STRUCTURES.............................................. 32

2.8.1Review of Data and Effects Models................................................................................ 32

2.8.2Selecting Levels of Concern ........................................................................................... 33 

vii 

 
TABLE OF CONTENTS (cont’d) 

3 CONSEQUENCE ASSESSMENT EXAMPLES...................................................................... 36

3.1 RELEASE RATES.................................................................................................................. 36

3.2 POOL FIRES...........................................................................................................................36

3.3 FLAMMABLE VAPOR DISPERSION ................................................................................. 38

4 LNG RELEASE PREVENTION AND MITIGATION MEASURES....................................40

4.1 INTRODUCTION...................................................................................................................40

4.2 EQUIPMENT OVERVIEW.................................................................................................... 40

4.3 INCIDENT PREVENTION MEASURES.............................................................................. 40

4.3.1Regulatory and Industry Approaches for LNG Carrier Design ...................................... 40

4.3.1.1 Definitions..........................................................................................................41

4.3.1.2Design specifications for LNG tankers...............................................................42 

4.3.2Design Features Pertinent to Release Prevention............................................................ 43

4.3.3Operational Measures Pertinent to Accidental Release Prevention................................ 47

4.3.4Security Measures that Help Prevent Release Incidents Due to Deliberate Attacks.......47

4.4 MITIGATION MEASURES TO HELP RESPOND TO RELEASES ................................... 49

5 CONCLUSIONS..........................................................................................................................51

5.1 REVIEW OF WORK SCOPE................................................................................................. 51

5.2 OBSERVATIONS................................................................................................................... 54

6 REFERENCES ............................................................................................................................55

Appendix A Film Boiling Heat Transfer Calculations.................................................................A-1 

Appendix B Example Release Rate Calculations..........................................................................B-1 

Appendix C Example Pool Fire Calculations................................................................................C-1 

Appendix D Example Source Term Calculations for Flammable Vapor Dispersion................D-1 

viii 

 
LIST OF TABLES 

Table Page

Table 1.1 Fire-related Properties of LNG and Other Light Hydrocarbon Fuels......................... 4

Table 2.1 Summary of Test Data for LNG Spilled on Water ...................................................... 15

Table 2.2 Thermal Radiation Burn Injury Criteria from FEMA (1990).................................... 30

Table 2.3 Permissible Thermal Radiation Exposure for Flares from API 521 (1997) ..............30

Table 2.4 Thermal Dose Data for Exposure to Fireballs from Prugh (1994)............................. 31

Table 2.5 Probit Models for Injury by Thermal Radiation from TNO (1992)........................... 31

Table 2.6 Effects on People for 1,600 BTU/hr/ft2 (5 kW/m2) Thermal Radiation......................32

Table 2.7 Structure Damage Criteria for Thermal Radiation Exposure from World

Bank (1988)...................................................................................................................... 33

Table 2.8 Damage Resulting from Thermal Radiation for Various Materials from 

TNO (1992) ...................................................................................................................... 33

Table 2.9 Various Thermal Radiation Limits for Structures from Lees (1996)......................... 34

Table 3.1 Summary of Results for Example Pool Fire Calculations........................................... 38

Table 3.2 Summary of Results for Example Dispersion Calculations ........................................ 39

LIST OF FIGURES 

Figure Page

Figure 3.1 Estimated LNG Release Rates Based on Orifice Model .............................................. 37 

ix 

 
ACRONYMS 

AIChE American Institute of Chemical Engineers 

AIS automatic identification system

ANOA Advanced Notice of Arrival 

API American Petroleum Institute 

CFD computational fluid dynamics 

CFR Code of Federal Regulations 

COTP Captain of the Port 

DoS declaration of security

EPA Environmental Protection Agency

FERC Federal Energy Regulatory Commission 

GRI Gas Research Institute 

IGC Code International Code for Ships Carrying Liquefied Gases in Bulk 

IMO International Maritime Organization 

ISM Code International Management Code for the Safe Operation of Ships and for Pollution 

Prevention 

ISPS Code International Ship and Port Facility Security Code 

LFL lower flammability limit

LNG liquefied natural gas 

LPG liquefied petroleum gas 

MARVS maximum allowable relief valve setting 

MIE minimum ignition energy 

MSO Marine Safety Office 

NFPA National Fire Protection Association 

PSC Port State Control 

RPT rapid phase transition 

SIGTTO Society of International Gas Tanker and Terminal Operators 

SOLAS Safety of Life at Sea 

SSA ship security alarm

STCW Standards of Training, Certification, and Watchkeeping 

TNT trinitrotoluene 

UFL upper flammability limit 

USCG United States Coast Guard 

x 

 
SCIENTIFIC SYMBOLS 

Cd discharge coefficient 

CF turbulent or laminar resistance force 

FG gravitational force 

FI inertial resisting force 

g gravitational acceleration 

gr reduced gravitational acceleration 

h mean pool height 

f h pool height at the leading edge 

H liquid height above hull breach 

Q flow rate 

r pool radius 

R radius of hull breach 

Φ coefficient for gravity term

ρ l density of LNG 

ρ w water density

UNITS 

BTU British thermal unit kJ kilojoule 

BTU/ft2 British thermal unit per square foot kJ/m2 kilojoule per square meter 

BTU/hr/ft2 British thermal unit per hour per square

foot 

kW kilowatt

°C Celsius kW/m2 kilowatt per square meter 

cm centimeter lb/ft2

/s pound per square foot per 

second 

°F Fahrenheit lb/s pounds per second

ft foot m meter

ft/s feet per second m/s meter per second 

ft3 cubic feet m3 cubic meter 

K Kelvin m3

/s cubic meter per second 

kg kilogram min minute

kg/m2

/s kilogram per square meter per second mph miles per hour 

kg/s kilograms per second psi pounds per square inch 

sec second

xi 

 
1 

 

1 INTRODUCTION 

1.1 BACKGROUND 

Liquefied natural gas (LNG) has been transported by sea since 1959 in specially designed LNG 

carriers. These vessels have a remarkable safety record and provide an essential link in the 

movement of LNG from production locations to consumer locations. 

The combination of recent interest in expanding or building new facilities to receive LNG 

carriers (i.e., as a result of increases in demand for LNG in the United States), along with increased 

awareness and concern about potential terrorist actions, has caused stakeholders to raise questions 

about the potential consequences of incidents involving LNG carrier operations. For example, the 

size and extent of a possible fire or the distance a vapor cloud may extend are important factors in 

gauging the acceptability of a new facility. 

Several recently published studies on releases or spills from LNG carriers employ varying 

methodologies and assumptions. As a result, these studies report some widely varying estimates of 

potential flammable vapor and thermal radiation hazard distances. 

In this context, the Federal Energy Regulatory Commission (FERC) sponsored this study with 

the goal of identifying appropriate analysis methods for estimating flammable vapor and thermal 

radiation hazard distances for potential LNG vessel cargo releases during transit and while at berth. 

1.2 SCOPE OF WORK 

This work considers the flammable vapor and thermal radiation hazards created by unconfined 

LNG spills on water resulting from an LNG cargo release. This includes a search for and review of 

literature on experimental LNG spills and on consequence assessment methodologies that are 

applicable to modeling of incidents involving LNG spills on water. Key consequences of concern 

include the distance from a spill that flammable vapors might travel and, in the case of a pool fire, the 

distance from a spill that might receive harmful thermal radiation. 

The specific scope of work, as defined by FERC, is as follows: 

• An evaluation of published models to calculate the rate of cargo release and the spread of 

unconfined LNG spills on water; 

• A determination of the effect of cargo tank hole diameter on cargo release and pool spread 

(including establishment of an upper limit on pool spread); 

• An evaluation of these models to calculate the rate of vapor generation for unconfined LNG 

spills on water; 

• An evaluation of the models to calculate thermal radiation distances from fires involving 

unconfined LNG spills on water; 

• An evaluation of the models to calculate the distance traveled by flammable vapors from 

unconfined LNG spills on water; 

• An evaluation of the theoretical basis for current models used to calculate each of the above, 

and a review of the applicability of existing computer modeling to field test data; 

• A review of the effects of thermal radiation levels on population and structures; 
2 

 

• A recommendation, including the selection rationale, of specific methods and a model FERC 

staff may use to calculate the flammable vapor and thermal radiation hazards associated with 

marine transportation of LNG; and 

• A comprehensive review of design, operational, safety, and security features and possible 

mitigating measures that an LNG facility and/or vessel operator employs or could provide to 

minimize impacts from: accidents; attacks; leaks; spills; fire; rapid phase transitions; or the 

spread of negatively buoyant, flammable vapor clouds. 

A key goal of this work is to search for and review published methodologies for consequence 

assessment that are applicable to LNG spills on water. As part of this review, FERC requested the 

project specifically address the methods applied in four reports: Fay (2003), Lehr (2004), Quest 

(2001), and Vallejo (2003). Section 2 of this report discusses the modeling approaches in Fay 

(2003), Lehr (2004), and Quest (2001). However, the Vallejo (2003) report does not provide 

sufficient technical information and is not discussed. 

1.3 PROPERTIES OF LNG 

LNG is transported and stored at temperatures as low as -259 oF (-162 oC). LNG is typically 85 

to 96% methane on a volume basis, with the balance being mostly other light hydrocarbons such as 

ethane, propane, butane, and up to 1% nitrogen. Methane is flammable in air at 5 to 15% (by 

volume). At a pressure of 1 atmosphere and temperature of -259 oF (-162 oC), its normal boiling 

point, LNG can evaporate, forming vapor, which has a specific gravity of 1.7. However, LNG vapor 

at ambient temperatures is lighter than air, and its specific gravity relative to air is 0.55 (once thermal 

equilibrium is reached). Hence, LNG vapor from a liquid release will tend to stay near the surface of 

the ground or water until it mixes with air and warms to a temperature of approximately -162 °F 

(-108 oC), at which point it will become less dense than air and tend to rise and disperse more 

rapidly. 

1.4 LNG HAZARDS 

LNG’s principal hazards result from its (1) cryogenic temperature -259 oF (-162 oC), (2) 

flammability, and (3) vapor dispersion characteristics. As a liquid, LNG will neither burn nor 

explode. Methane, the primary component of LNG, is colorless, odorless, and tasteless, and is 

classified as a simple asphyxiant for human exposure. LNG vaporizes rapidly when exposed to 

ambient heat sources such as water, producing 620 to 630 standard cubic feet of natural gas for each 

cubic foot of liquid. 

When spilled onto water, LNG will initially produce a negatively buoyant vapor cloud (i.e., the 

cold vapors are more dense than air and stay close to the water or ground). As this cloud mixes with 

air, it will warm up and disperse into the atmosphere. If not ignited, the flammable vapor cloud 

would drift downwind until the effects of dispersion dilute the vapors below a flammable 

concentration. At a 5 percent concentration of gas in air, LNG vapors are at their lower flammability 

limit (LFL). Below this vapor/air ratio, the cloud is too dilute for ignition. At a 15 percent 

concentration of gas in air, LNG vapors are at their upper flammability limit (UFL). Above this 

vapor/air ratio, the cloud is too rich in LNG for ignition. 

The downwind distance that flammable vapors might reach is a function of the volume of LNG 

spilled, the rate of the spill, and the prevailing weather conditions. Also, in order to disperse to 
3 

 

significant downwind distances, a vapor cloud must avoid ignition. Evaluation of ignition 

probability is beyond the scope of this study; however, it is noted that the large releases from an LNG 

carrier would likely require a significant energy source to initiate (i.e., to puncture the outer hull, 

inner hull, and cargo tank). An event of sufficient magnitude to rupture an LNG cargo tank may also 

provide ignition sources. If a flammable cloud is ignited by the initiating event or by other ignition 

sources (e.g., on the ship, on other nearby vessels, or on shore), the flame will burn back to the vapor 

source, and the flammable cloud would not travel a significant distance over land. 

If a flammable vapor-air mixture from an LNG spill is ignited, it may result in a flash fire, which 

is a short duration fire burning the vapors already mixed with air in flammable concentrations. The 

flame front will burn back through the vapor cloud to the spill site, provided the vapor concentration 

along this path is high enough to continue burning. The rate at which this flame front travels relative 

to the unburned gas is called the laminar burning velocity. An unconfined methane-air mixture will 

burn slowly, tending to ignite combustible materials within the vapor cloud, whereas fast flame 

speeds tend to produce flash burns rather than self-sustaining ignition. 

Although LNG vapors can explode (i.e., create large overpressures) if ignited within a confined 

space, such as a building or structure, there is no evidence suggesting that LNG is explosive when 

ignited in unconfined open areas. Experiments to determine whether unconfined methane-air 

mixtures will explode have been conducted and, to date, have been negative. 

The principal LNG hazards of interest for this study are those posed by flammable vapor 

dispersion and thermal radiation. Secondary hazards, such as cryogenic burns and asphyxiation, are 

typically localized to LNG transport and storage areas and are outside the scope of this study. 

LNG is less hazardous than liquefied petroleum gas (LPG) and liquefied ethylene, which have (1) 

higher specific gravities, (2) a greater tendency to form explosive vapor clouds, (3) lower minimum 

ignition energies (MIEs), and (4) higher fundamental burning velocities. LNG is not toxic, and it 

rapidly evaporates; therefore, long-term environmental impacts from a release are negligible if there 

is no ignition of natural gas vapors. 

1.4.1 Fire Hazards 

LNG vaporizes quickly as it absorbs heat from the environment, and the resulting vapor is 

flammable when mixed in air at concentrations from 5 to 15% (volume basis). Its fire-related 

properties are comparable to other light hydrocarbon fuels (see Table 1.1). The only significant 

difference is that its molecular weight is considerably less than air, so once it warms above 

approximately -162 oF (-108 oC) it will become less dense than air and tend to rise and disperse more 

rapidly. However, LNG vapor at its normal boiling point -259 oF (-162 oC) is 1.5 times more dense 

than air at 77 oF (25 oC). Typically, LNG released into the atmosphere will remain negatively 

buoyant until after it disperses below its LFL. 

Three types of fires — pool fires, jet fires, and flash fires — are postulated for the purposes of 

this study. 

Pool Fire – When a flammable liquid is released from a storage tank or pipeline, a liquid pool 

may form. As the pool forms, some of the liquid will evaporate and, if flammable vapor finds an 

ignition source, the flame can travel back to the spill, resulting in a pool fire, which involves burning 

of vapor above the liquid pool as it evaporates from the pool and mixes with air. 
4 

 

Jet Fire – If compressed or liquefied gases are released from storage tanks or pipelines, the 

materials discharging through the hole will form a gas jet that entrains and mixes with the ambient 

air. If the material encounters an ignition source while it is in the flammable range, a jet fire may 

occur. For LNG stored at low pressure as a liquid, as it is in an LNG carrier, this type of fire is 

unlikely. Jet fires could occur during unloading or transfer operations when pressures are increased 

by pumping. Such fires could cause severe damage but will generally affect only the local area. This 

report focuses on large spills on water, and analysis of jet fires is outside the scope. 

Flash Fire – When a volatile, flammable material is released to the atmosphere, a vapor cloud 

forms and disperses (mixes with air). If the resultant vapor cloud is ignited before the cloud is 

diluted below its LFL, a flash fire may occur. The combustion normally occurs within only portions 

of the vapor cloud (where mixed with air in flammable concentrations), rather than the entire cloud. 

A flash fire may burn back to the release point, resulting in a pool or jet fire but is unlikely to 

generate damaging overpressures (explode) when unconfined. 

Table 1.1 Fire-related Properties of LNG and Other Light Hydrocarbon Fuels 

Material Ethylene1,2 Gasoline1,3 LNG/methane1,2 Propane (LPG)1,2

Flash Point (oF) -186 -40 to -49 <-259 <-155 

Flammability Limits (% LFL 2.7 1.3 5 2.1 

in air) UFL 36 7.1 15 9.5 

Autoignition Temperature (oF) 910 820 1000 840 

Minimum Ignition Energy 

(BTU) 6.6 × 10-8 Not 

reported 2.5 × 10-7 2.4 × 10-7

Fundamental Burning Velocity4

(ft/s) 2.6 1.3 1.3 1.5 

Vapor Specific Gravity 0.97 < 2 0.55 1.5 

Data sources: 

1. NFPA (1995) 

2. Lees (1980) 

3. NFPA (2001) 

4. NFPA (2002) 

Terms: 

Flash Point – The minimum temperature at which a liquid gives off vapor in sufficient concentration to form an 

ignitable mixture with air near the surface of a liquid, as specified by test. 

Flammability Limits – The lowest (LFL) and highest (UFL) concentrations of a combustible substance in a 

gaseous oxidizer that will support burning. 

Autoignition Temperature – Temperature at which a flammable mixture will spontaneously ignite. 

Minimum Ignition Energy – The minimum amount of energy released at a point in a combustible mixture that 

causes flame propagation away from the point, under specified test conditions. 

Fundamental Burning Velocity – The burning velocity of a laminar flame under test conditions. When ignited in 

a flammable vapor cloud, substances with lower fundamental burning velocities will tend to generate lower 

overpressures. 

Vapor Specific Gravity – Ratio of the molecular weight of the material/compound to the molecular weight of air 

(based on an assumed composition of 79 vol% N2 and 21 vol% O2). 
5 

 

1.4.2 Explosions 

As discussed in the previous section, a flash fire can occur if LNG is released into the atmosphere 

and ignited. If ignited in open (unconfined) areas, pure methane is not known to generate damaging 

overpressures (explode). However, if some confinement of the vapor cloud is present, methane can 

produce damaging overpressures. Confinement can be provided by spaces within the ship or nearby 

structures, such as a building on shore or another ship. Areas congested with equipment and 

structures can also facilitate damaging overpressures if a vapor cloud is ignited within such an area. 

For example if a vapor cloud infiltrates a chemical process plant in an area with various vessels, 

structures, and piping and the cloud ignites, the portion of the cloud within that congested area may 

generate damaging overpressures. 

A larger volume fraction of heavier hydrocarbons in the LNG (1) reduces the minimum ignition 

energy required for detonation and (2) increases the specific gravity of the hydrocarbon mixture (and 

hence reduces the tendency to rapidly disperse). Both of these effects increase the likelihood of 

generation of damaging overpressures. 

1.4.3 Rapid Phase Transition (RPT) 

RPT is the term used to describe a phenomenon recognized in some LNG release experiments 

involving the nearly instantaneous transition from the liquid to vapor phase and an associated rapid 

pressure increase. This topic is further discussed in Section 2.6. 

When LNG forms a pool on water, the heat from the water rapidly vaporizes the LNG; however, 

this boiling is not the phenomenon referred to as RPT. In an RPT, a portion of the spilled LNG 

changes from liquid to gas virtually instantaneously. 

Although the physical mechanism is not well understood, RPT is attributed to the superheating of 

the LNG due to the lack of nucleation sites (sites that help with the formation of gas bubbles and 

promote boiling). An RPT may result in two types of effects: (1) overpressure resulting from the 

rapid phase change and (2) dispersion of the “puff” of LNG expelled into the atmosphere. Rapid 

phase changes have caused numerous steam explosions in foundries and other industrial operations, 

but have not resulted in any known major mishaps involving transport of LNG. 

1.4.4 Cryogenic Effects 

LNG carriers are designed to prevent LNG from contacting the inner and outer hulls, but 

incidents can be postulated that would place LNG in contact with the hulls. It is possible that a 

release of liquid LNG to the inner hull would cause low temperatures for areas of the hull structure 

that are not designed for cryogenic temperatures. (The international ship design rules require areas 

where cargo tank leakage might be expected to be designed for contact with cryogenic LNG.) 

Assessing the likelihood or result of cold temperature on hull structural members was not within the 

scope of this study. 

A study by Lloyds (2001) includes brief descriptions of 10 LNG spills involving LNG carriers 

that occurred between 1965 and 1989. Seven of these ten spills led to brittle fracture of the deck or 

tank covers, but none of them caused serious structural damage. Given where the damage occurred 

(i.e., to the deck or tank cover), it is likely that all of these releases occurred from piping systems 
6 

 

used during LNG transfers. Also, vessels built since 1976 have to comply with current ship design 

rules and are designed with steel rated for low temperature in areas where LNG leakage might be 

expected to contact decking or internal structures. (See Section 4.3.1.2 for a discussion of the design 

requirements for LNG tankers.) 
7 

 

2 CONSEQUENCE ASSESSMENT METHODS 

This section presents a review of available methods for modeling some specific aspects of events 

involving large releases of LNG on water. Also presented are recommended approaches for 

modeling, which are based on: 

• The reviews of available methods, 

• The types of events of interest (e.g., large LNG releases on water leading to pool fires and/or 

flash fires), and 

• The intended uses for the consequence assessment results (e.g., facility siting decision 

making, emergency planning). 

Section 3 presents some example calculations that follow the methods recommended in this section. 

2.1 RATE OF RELEASE OF LNG FROM A SHIP 

If an LNG cargo tank is punctured below the liquid level, the LNG will flow through the hole. 

This section presents an evaluation of methods to calculate how fast the liquid will spill. 

2.1.1 Review of Modeling Approaches 

Fay (2003) – Fay presents two release rate models, one for holes above the seawater level and 

one for holes below the seawater level. The model for holes above the seawater level is equivalent to 

the well-known relationships for flow through an orifice, driven by gravity. The orifice model is 

presented in many references on consequence assessment, such as AIChE (2000) and TNO (1997), as 

well as most basic textbooks on fluid mechanics. This model represents flow from a hole in the side 

of a single hull cargo tank that allows the LNG to flow directly from the tank into the water. A 

discharge coefficient is often used to represent the frictional loss accompanying flow through the 

orifice; Fay does not include a discharge coefficient, which is equivalent to assuming a value of 1.0 

(no frictional loss). This model does not attempt to account for the multi-hull construction of LNG 

carriers, and therefore may overestimate the rate at which LNG would escape through a breach. 

Fay’s model for holes below the seawater level is an idealistic approach in which the difference 

in density of the fluids causes (1) the LNG to flow out the hole to the water surface and (2) water to 

flow into the ship. However, this method ignores the fact that the boiling point of LNG is well below 

the water temperature. In reality, the LNG would flash to vapor as it exited the hole and came into 

contact with the water, rapidly rising to the surface. Water that did enter the ship would likely freeze 

and might affect the rate of LNG release. This model also does not attempt to account for the multihull construction of LNG carriers. 

Lehr (2004) – Lehr neither presents nor uses release rate models. This paper presents a single 

release example that is an instantaneous release of 500 m3

 of LNG. 

Quest (2001, 2003) – This study postulates 1-m and 5-m diameter holes but does not describe the 

method used to calculate the flow rate from the holes. 
Other Methods – Our literature review did not identify any additional methods for estimating

release rates for spills from LNG carriers. No methods were identified that attempt to account for the

multiple barriers to release that are present in LNG carriers. 

2.1.2 Modeling Recommendations 

In LNG carriers, the LNG is contained inside an insulated cargo tank that is contained within the

inner hull of the ship. The scenarios of interest for this study involve incidents that puncture both the

inner and outer hulls, the insulation layer, and the LNG cargo tank. No existing release model 

accounts for these multiple barriers to release or for a ship’s possible response to a large release of 

LNG (e.g., into the inner compartments of a ship). 

In addition, the existing models assume that air can enter the space at the top of the LNG 

compartment or LNG can be vaporized to replace the exiting liquid at a volumetric rate equal to the 

discharge rate from a hole in the side wall of the tank. That is, they ignore the possibility of low 

pressure in the vessel reducing the rate of outflow. For small diameter holes, or for scenarios where 

a hole extends above the liquid level, this may be a valid assumption. For large diameter holes, the

liquid release rate may be higher than the rate at which vapor can be generated or at which air can 

enter through the vacuum relief valve. If this occurs, a partial vacuum may be created in the tank,

reducing the outflow rate and possibly further damaging the compartment (if it is not designed for 

full vacuum).

In the absence of models that account for the true structure of an LNG carrier and the physics of a

release of cryogenic LNG, the simple orifice model is recommended: 

Q Cd l R 2gH 2 = πρ

where Q = flow rate 

 = discharge coefficient (taken as 1.0) Cd

ρ l = density of LNG 

R = radius of hull breach 

H = liquid height above hull breach 

When using this model it is vital to keep in mind that the model does not reflect actual carrier 

construction very well, and the results should be interpreted as a rough guide to the rate of release for 

a given hole size. It is also important to note that cargo spilling between hulls is not accounted for. 

The orifice model represents the double-hulled ship as a single-wall vessel, and therefore will tend to 

overestimate the LNG outflow for many scenarios. 

The orifice model is presented in many references on consequence assessment, such as AIChE

(2000) and TNO (1997), as well as most basic textbooks on fluid mechanics. 

There are no appropriate models available for releases through a hole in the tank wall that is 

below the seawater level. As described above, Fay (2003) does present a model for underwater 

punctures, but his model does not account for the (1) actual physics of a release of cryogenic liquid

directly into seawater or (2) actual geometry of an LNG carrier. In the absence of an appropriate

model, the orifice model may be applied to provide a rough guide to the rate of release. 

8 

 
9 

 

A qualitative discussion of what might happen in a release below water level is useful in 

understanding the difficulty in accurately modeling such a situation. To a certain level below the 

seawater surface, the LNG pressure will exceed the water pressure, so the LNG will tend to flow out 

the hole and vaporize rapidly as it contacts the water. For holes further below seawater level (or once 

the LNG level drops), a point will be reached where the LNG pressure and water pressure match, and 

seawater may begin entering the hole. Any seawater that entered the vessel would tend to freeze 

rapidly, and the heat it transfers to LNG within the vessel will vaporize LNG, possibly increasing 

pressure in the vessel and forcing LNG and/or water out the breach. Other possible scenarios might 

involve both LNG and seawater spilling into the void space between the tank and outer hull. 

For many scenarios, the orifice flow model is expected to overestimate the release rate because it 

does not account for factors such as (1) a reduction in tank pressure as liquid exits (i.e., when outflow 

exceeds rate at which volume can be replaced by air or LNG vaporization) or (2) the additional 

barriers to release provided by the design of the cargo system and the structure of the ship. On the 

other hand, the orifice model does not address any additional ship damage that might occur as a 

result of a postulated release. 

Two other aspects of LNG spills were also considered in this review: 

Effects of LNG Composition – Cumber (2002) assessed the discrepancy between estimates of 

outflow for LNG using the correct composition and using an approximation of the material as pure 

methane. This report finds that composition can “change the shape and location of the boundary of 

the two-phase envelope,” which can significantly affect outflow. He presents an example with LNG 

that is 90% methane held at 44 psi (3 bar) and –244 oF (120 K). For this example, the pure methane 

approximation underpredicts initial outflow by nearly a factor of two. Within roughly 300 seconds, 

the estimated flow rates converge to approximately the same value. In the case of an LNG carrier, 

LNG is stored very near the normal boiling point, so two-phase flow effects are expected to be 

minimal. Considering the overall uncertainty introduced by applying the orifice model to postulated 

LNG release scenarios, approximating the outflow using pure methane properties will introduce 

negligible additional uncertainty. 

Transient Nature of Spills – For the idealized case of a gravity-driven release of liquid from a 

hole in a vessel, the flow rate will decrease as the height of the liquid above the hole decreases, and 

the flow will stop once the liquid level reaches the bottom of the hole. If appropriate for the scenario 

postulated, this transient flow rate should be applied when modeling LNG carrier releases. For other 

scenarios, it may be more appropriate to conservatively assume a constant outflow rate (i.e., at the 

initial rate predicted by the orifice model). 

2.2 SPREAD OF AN UNCONFINED POOL ON WATER 

LNG is less dense than water and when spilled on water will form a floating pool that will spread 

and evaporate. This section presents an evaluation of the methods for estimating the rate and extent 

of LNG pool spread on water. 

2.2.1 Review of Modeling Approaches 

Fay (2003) – Fay presents a model based on a balance of fluid inertia and gravity force. This 

model assumes the water surface is smooth. 
Lehr (2004) – Lehr applies a model based on a balance of fluid inertia and gravity force; the 

model includes empirical constants, which he takes from Briscoe (1980). For large spills, Lehr 

points out that spread may become controlled by a balance of viscous and gravity forces. However,

he concludes that the time at which this would occur would likely be longer than the expected burn

time for an LNG pool fire. This model assumes the water surface is smooth. 

Quest (2001, 2003) – Quest also applied a model for spread based on a balance of fluid inertia 

and gravity force. In addition, Quest attempted to account for the effects of waves on spread. The 

waves were approximated by a cycloid function, and computations were performed to apply 

boundary conditions on the basic gravitational spreading model. The effect of these boundary 

conditions is that the pool stops spreading once the height of the LNG above the water level drops

below 60% of the wave height. An adjustment was also made for the increased water surface area

caused by the waves, which results in an increase in vaporization flux of 27%. 

Other Methods – Several pool spread models are currently in use. Early spread models

published by Hoult (1972), Fay (1973), Shaw (1978), and others are based on the steady state 

Bernoulli equation and approximate axisymmetric spread on water by: 

g h

dt

dr r = 2

where r = pool radius 

w

w l

r g g

ρ

ρ − ρ =

h = mean pool height 

ρ w = water density

ρ l = density of liquid in pool 

With this approach, spread is driven strictly by gravity, and the rate is given as a function of pool 

height only.

Raj and Kalelkar (1974) derived a different spreading relationship, based directly on equating 

gravitational force ( ) and inertial resisting force ( ), given by: FG FI

G r F πrh ρg

2 = and 2

2

2 ( )

dt

d r F C r h I = − π ρ

The constant C is introduced to account for the fact that the inertia of the entire pool is a fraction 

of that given by use of the acceleration of the leading edge ( 2

2

dt

d r ). Equating these forces and solving 

for acceleration gives: 

Crρ

g h

dt

d r r = − 2

2

This approach also defines spread as driven strictly by gravity.

10 

 
Both of these approaches effectively ignore resistance to spreading as a result of friction. Some

authors acknowledge this fact and mention that viscous effects may be important for “large” LNG 

spills on water. 

One method, developed by Webber (as described in TNO 1997), does account for friction effects. 

This approach is based on self-similar solutions of the shallow water equations and lubrication

theory. Webber’s formulation results in the following description of spreading: 

F

r C

r

g h

dt

d r − Φ = 4

2

2

 

inertia gravity resistance 

where: 

Φ = coefficient, which is a function of h

h f

= pool height at the leading edge f h

= turbulent or laminar resistance force CF

In this formulation, resistance by turbulent or laminar friction effects is included, and Webber 

provides methods for estimating the various values needed (e.g., Φ , ). CF

To better understand the characteristics of these methods and to assess the importance of friction

effects for spills of interest in this report, two of these methods were implemented using an integral 

solution approach. The methods implemented are those of Shaw (1978) and Webber (TNO 1997). 

The method of Webber was chosen because it is the only one that includes friction effects. The 

method of Shaw was chosen as representative of the methods based solely on gravitational spreading.

As shown by Otterman (1975), while the formulations of these gravitational spread methods vary,

they give very similar results. Otterman’s analysis included methods of Fay, Hoult, and Raj/ 

Kalelkar; note that the methods of Fay and Hoult included are equivalent to the method of Shaw as

presented above. 

Implementation of the Shaw and Webber methods was accomplished with the assistance of the 

Mathcad computer software (©1986-2000 MathSoft, Inc.). For the case of an LNG spill on water, 

the solutions are simplified by the fact that pool temperature remains constant (at the boiling point)

and the evaporation heat flux is constant (based on either film boiling heat flux or burning rate). The 

numerical solutions applied here also allow accounting for various types of spill sources, including 

those that are transient. 

Example pool spread scenarios and results from both the Shaw and Webber methods are shown

below. These scenarios involve spills of 4.4 × 105

 ft3 (12,500 m3) through 3.3-ft (1-m) and 16-ft 

(5-m) diameter holes. These hole diameters were selected for the sensitivity analysis since they are

representative of potential ship damage as found in other vulnerability assessments. Results are 

provided for two cases (1) assuming that a fire occurs immediately (at the beginning of the release) 

and (2) assuming that no fire occurs. The spill duration is the time required to empty the volume of

liquid that is above the hole in the cargo tank. 

11 

 
12 

 

Pool Spread Scenario for Comparison of Shaw and Webber Models – Fire Case

Hole diameters: 3.3 ft (1 m) and 16 ft (5 m) 

Initial liquid height above hole: 43 ft (13 m) 

Total spill quantity: 4.4 × 105

 ft3 (12,500 m3) 

Burning rate: 0.058 lb/s/ft2

 (0.282 kg/s/m2) 

Pool Spread Results for Comparison of Shaw and Webber Models – Fire Case

Hole diameter 3.3 ft (1 m) 16 ft (5 m) 

Initial spill rate 11,700 lb/s 

(5,300 kg/s) 

290,000 lb/s 

(130,000 kg/s) 

Total spill duration 33 min 1.3 min 

Spread model Shaw Webber Shaw Webber 

Maximum pool radius 250 ft 

(76 m) 

240 ft 

(74 m) 

790 ft 

(240 m) 

440 ft 

(130 m) 

Total duration of fire 33 min 33 min 3.2 min 6.9 min 

Pool Spread Scenario for Comparison of Shaw and Webber Models – Nonfire Case 

Hole diameters: 3.3 ft (1 m) and 16 ft (5 m) 

Total spill quantity: 4.4 × 105

 ft3 (12,500 m3) 

Air temperature: 71 °F (22 °C) 

Water temperature: 70 °F (21 °C) 

Pool Spread Results for Comparison of Shaw and Webber Models – Nonfire Case 

Hole diameter 3.3 ft (1 m) 3.3 ft (5 m) 

Initial spill rate 11,700 lb/s 

(5,300 kg/s) 

290,000 lb/s 

(130,000 kg/s) 

Total spill duration 33 min 1.3 min 

Heat transfer to LNG pool 11,700 BTU/hr/ft2 (37 kW/m2) 

Spread model Shaw Webber Shaw Webber 

Maximum pool radius 450 ft 

(140 m) 

430 ft 

(130 m) 

1,100 ft 

(340 m) 

550 ft 

(170 m) 

Total evaporation duration 34 min 34 min 6.1 min 18 min 

Conclusions from the analysis of the Shaw and Webber methods are that: 

• For some of the LNG spills of interest in this report, Webber’s methods indicate that friction 

effects on the LNG pool spreading over water are important, in particular, for large, shortterm releases. In a scenario involving a release of 4.4 × 105

 ft3 (12,500 m3) through a 16-ft 

(5-m) diameter hole, Webber’s method predicts a maximum pool radius that is approximately 
half that predicted by the method of Shaw. On the other hand, for longer duration spills, 

friction effects appear much less important. For a scenario involving a release of 4.4 × 105

 ft3

(12,500 m3

) through a 3.3-ft (1-m) diameter hole, the methods provide very similar results. 

• When Webber’s method is modified to ignore friction effects, it produces results very similar

to those of Shaw. As TNO (1997) points out for this special case, Webber’s method reduces 

to the same form as the basic gravitational methods (like Shaw), differing only by the 

coefficient on the right side. Shaw effectively has a coefficient of 2 = 1.41, and Webber 

uses a coefficient of 1.64 (empirically derived). 

• Webber’s method for determining the turbulent resistance term, , involves taking the

larger of estimates for laminar or turbulent flow. The implementation here included both 

factors, but for all of the scenarios examined, the turbulent factor dominated.

CF

• A numerical solution to the spreading problem is desirable because it allows handling 

transient spill sources, and therefore avoids artificially characterizing a spill as either 

“instantaneous” or “continuous.” 

13 

 

2.2.2 Modeling Recommendations 

Based on review of the literature and example implementation of the Shaw and Webber spread 

methodologies, the method of Webber (as described in TNO 1997) is recommended. For many

scenarios (e.g., long duration releases), this method will provide results that are very similar to those 

from the various methods based solely on gravity-driven spread. However, for some scenarios (e.g.,

large, short-duration releases), Webber’s method predicts significantly slower pool spread (as friction 

effects become more important). Because Webber’s method (1) has a more sound theoretical basis 

and (2) accounts for friction effects, these results are believed to be more realistic than those from the

simpler gravitational models that ignore friction effects. 

It is also recommended that pool spread modeling be implemented in a way that avoids the need

to characterize spills as either “instantaneous” or “continuous.” The integral approach taken in the 

example implementation for this report accomplishes this. If desired, it can accommodate an

“instantaneous” release (i.e., a specified pool volume in the water at time zero) and/or a fixed or

transient release rate into the pool. 

It is also useful to note that for some scenarios, detailed modeling of pool spreading is not

necessary. In particular, with a long-term (i.e., “continuous”) release, the pool will spread until the 

evaporation rate (or burn rate, in the case of a pool fire) matches the rate of addition from the spill.

In a case such as this, the pool area can be estimated using the spill rate and the evaporation rate per

unit area. Of course, as with the Webber model, this is predicated on the assumption of smooth, 

quiescent water. 

Because the recommended spread model does not account for wave action or currents, it can only 

provide an approximation of pool behavior. All LNG tankers transit waterways in tidal contact with 

the ocean, and therefore are usually subject to tidal currents and wave action. However, when

unloading cargo at a marine terminal (in a protected slip or turning basin) tidal currents and wave

action may at times be minimal. Wave action will tend to increase the rate of heat transfer to the 

spreading pool of LNG, which will increase the rate of LNG vaporization and limit the size of the 

pool area. Water currents (tidal, river, etc.) and wave movement will tend to stretch the pool in the 

direction of current flow and increase the size of the spill. Likewise, winds will tend to drag the LNG 
14 

 

liquid pool in the direction of the wind flow, increasing the size of the spill area. However, 

increasing wind will also increase the wave action on the spill. No existing model for an LNG spill 

accounts for the simultaneous impacts of wind, current, and wave action on the spread of the spill. 

The wave model of Quest (2001, 2003) appears to take a step in the right direction, but is not 

recommended at this time. This method may lead to low estimates of pool area and evaporation 

because it models waves as stationary objects, effectively assuming that the effects of wave 

movement, wind, and currents are negligible. In the examples presented by Quest for wave heights 

from 0.575 to 1.24 m, the pool (circle) area was reduced by more than 90%. The model does 

recognize the increased water surface area within the circle, which Quest states results in an increase 

in vaporization flux of 27%. However, because the pool radius is reduced so dramatically, the model 

results in lower estimates for both pool radius and total evaporation rate (i.e., mass per unit time). 

These results may be correct, but until the model is better developed and supported (experimentally 

and/or theoretically), a more conservative approach is recommended. 

This work did not include a comprehensive review of computer software for performing 

spreading calculations. However, several available software programs implement gravitational 

spreading methods and would be expected to provide reasonable results for scenarios where friction 

effects are minimal. Examples of such software include SOURCE5, which implements the methods 

of Raj and Kalelkar (1974) and LPOOL (part of HGSYSTEM), which implements the methods of 

Cavanaugh (1994). However, without applying a more complete model (like Webber’s), it may be 

difficult to determine if friction effects are important. Webber’s work included development of a 

computer model called GASP for the Health and Safety Executive in the U.K., but this software is 

neither publicly or commercially available. 

As a general rule, it may often be difficult to apply existing computer models if they provide 

limited ability to specify input parameters. For example, if a software package is designed to 

calculate heat transfer itself and does not calculate an appropriate value for LNG on water (as 

discussed in Section 2.3), then the software could not be used for modeling the scenarios considered 

in this study. 

2.3 VAPOR GENERATION FOR UNCONFINED SPILLS ON WATER 

As an LNG pool spreads on a water surface, the heat transferred from the water and other sources 

will cause the liquid to evaporate. This section presents an evaluation of the methods for estimating 

the rate of evaporation for spills on water. 

2.3.1 Review of Modeling Approaches 

Fay (2003) – This paper does not address vapor generation in a detailed manner, but for the 

example calculations presented, a burn regression rate for pool fires is taken as 8 × 10-4 m/s 

(evaporation rate of 0.34 kg/m2

/s). This value is based on contributions to evaporation caused by 

heat from the fire, giving a regression rate of 1.9 × 10-4 m/s (evaporation rate of 0.080 kg/m2

/s), and 

heat from the water, giving a regression rate of 5 × 10-4 to 7 × 10-4 m/s (evaporation rate of 0.21 to 

0.30 kg/m2

/s). 

Lehr (2004) – This report primarily addresses pool fires and provides information about 

vaporization only in the context of fire. 
15 

 

Quest (2001, 2003) – Evaporation is assumed to be controlled by film boiling. Calculations 

provide a heat flux to the LNG of 92 kW/m2

 and evaporation rate of 0.18 kg/m2/s. As mentioned in 

Section 2.2, Quest attempted to account for the effects of waves on vapor generation by 

approximating the waves as a cycloid function. Boundary conditions were applied to the spread 

model, which resulted in smaller pool diameters. An adjustment was also made for the increased 

water surface area caused by the waves, which results in an increase in vaporization flux of 27%. 

Other Methods – One method available for estimating vapor generation for an LNG pool on 

water is to calculate the heat transfer into the pool. Because the pool is at its normal boiling point, 

heat added goes directly to vaporize a portion of the LNG. Those who have taken this approach 

assume film boiling of the LNG and use correlations from heat transfer theory, such as that given by 

Klimenko (1981), to calculate the heat transfer. During film boiling, a layer (or film) of pure methane 

vapor forms between the liquid LNG pool and the water surface beneath it. Heat transfer between 

the water and the LNG occurs by conduction and radiation through the vapor layer. 

Models reviewed also assume that the water remains at a constant temperature that does not 

decrease with time. Consequently, heat transfer from the water to the LNG pool can be assumed to 

remain constant. No models were identified that attempt to quantify the convection heat transfer in 

the water. 

A second method for estimating vaporization from an LNG pool is simply to select a value from 

the available experimental data. Some of the available data for spills of LNG on water are provided 

in Table 2.1. 

Table 2.1 Summary of Test Data for LNG Spilled on Water 

Reference Description 

Vaporization 

Rate 

(kg/m2/s) 

Associated 

Heat Flux 

(kW/m2) 

Burgess (1970) U.S. Bureau of Mines tests 

Tests in an aquarium (2 ft by 1 ft by 1 ft) 

Ice film formed after approximately 20 

seconds 

0.181 92 

Burgess (1972) U.S. Bureau of Mines tests 

Tests on a pond (70 m diameter and 8 m deep) 

No ice formed 

0.155 79 

Boyle (1972) Shell Research Tests 

Laboratory scale 0.024 to 0.195 12 to 99 

McQueen 

(1972) 

Esso Tests 

Spill volumes from 0.8 to 10.8 m3 0.195 99 

Koopman 

(1979) 

5 m3 LNG spill tests at China Lake 

Pool sizes ranged from 6.8 to 7.8 m radius 0.12 to 0.15 61 to 78 
16 

 

Table 2.1 Summary of Test Data for LNG Spilled on Water (cont’d) 

Reference Description 

Vaporization 

Rate 

(kg/m2/s) 

Associated 

Heat Flux 

(kW/m2) 

LNG Research Center at MIT 

Laboratory scale 

Pure methane 

Range is over a period of 100 seconds 

Ice began to form 

0.05 to 0.23 25 to 120

ValenciaChavez (1979) 

Tertiary LNG (82.9% C1, 10.1% C2, and 7% 

C3) 

Range is over a period of 90 seconds 

Ice began to form 

0.02 to 0.28 10 to 140

2.3.2 Modeling Recommendations 

The literature reviewed consistently agrees that the rate of vaporization for a spill of a cryogenic 

liquid on water will be controlled by the rate of heat transfer from the water into the LNG. Because 

the pool is already at its normal boiling point, any heat added will go directly to vaporize the liquid. 

Factors that affect the heat flux from the water include: 

• Convective heat transfer cells formed in the water below the spill 

• Water currents below the spill 

• Waves that could increase surface area 

• Ice formation between the spill and the water 

• The degree of spill confinement 

• The heat transfer phenomenon occurring at the surface (e.g., the boiling mechanism) 

None of the literature reviewed attempt to account for all of these factors. 

Regarding convection and water currents, the consensus approach is to assume that convection 

and water currents are sufficient to keep the water well mixed and maintain the water temperature at 

its initial level. As long as the quantity of water available is large compared to the spill volume, this 

is a reasonable assumption. 

Regarding waves, the wave model of Quest (2001, 2003) is not recommended at this time (see 

discussion in Section 2.2.2), and no other models were identified. 

Regarding ice formation, the literature consistently indicates that ice can form when cryogenic 

material is spilled on water in laboratory-scale experiments. However, in larger-scale experiments 

(such as those on ponds), ice formation is minimal. For large LNG spills from carriers, it is 

recommended to assume no ice formation. 

Confinement should be considered on a site-specific basis. For most scenarios involving large 

spills from LNG carriers, it is assumed that the spill will be unconfined. However, for narrow 
17 

 

shipping channels, the spill will be confined by the shoreline, and the predicted unconfined pool 

radius may exceed the width of the channel. 

The literature indicates that LNG will boil on water in the film boiling regime for the majority of 

the evaporation time, until almost all of the methane is evaporated. The film is a thin layer of pure 

methane vapor between the liquid LNG pool and water surface below. 

Chemical composition can also significantly impact LNG vaporization, as concluded by Conrado 

(2000). However, this work, which simulated vaporization of (1) pure methane and (2) a 90% 

methane and 10% ethane mixture, indicated that assuming pure methane underestimates total 

evaporation time by 10 to 15%. In addition, the vaporization rates for pure methane and the 

methane/ethane mixture match closely until much of the methane is evaporated. Given the overall 

uncertainty in the analysis of large LNG spills on water, attempting to include effects of composition 

is not warranted, at least for currently typical LNG compositions. 

Considering all of the factors discussed above, it is recommended that vaporization be estimated 

directly from the film boiling heat flux from the water to the LNG. The heat flux is the rate at which 

heat is transferred from the water to the LNG. The water temperature is assumed constant and the 

pool remains at the normal boiling point, so a constant heat flux value can be used. The evaporation 

rate is then calculated directly from the heat flux and heat of vaporization for LNG. 

The heat flux could be taken directly from experimental data; however, the available data for 

LNG show significant spread. As noted by Waite (1983) regarding experimental work, “estimates of 

the heat flux from water to liquefied gas pools have ranged over a factor of 4, from 25 kW/m2

 to 100 

kW/m2

.” Our review here showed some even higher estimates. Notable reasons for wide variation in 

experimental results include the (1) differences in the experimental setups, (2) differences in 

measurement techniques, and (3) fact that evaporation rates were typically inferred from other data 

(rather than measured directly). 

Because of (1) the uncertainty in the experimental data and (2) the availability of a viable 

theoretical model from heat transfer theory to help fill the gap in experimental data, it is 

recommended that heat flux be estimated using film boiling heat transfer theory. This choice is also 

supported by Waite (1983), who performed a simulation of some of the experiments for which data 

are available, using a “typical film boiling heat flux of 25 kW/m2

” and using a heat flux of 100 

kW/m2

, “often used in LNG spill simulations,” for comparison. The conclusion of that comparison 

was that the simulation using a film boiling heat flux better matched the experimental data than the 

simulation with 100 kW/m2

. 

A correlation developed by Klimenko (1981) can be used to estimate the film boiling heat flux. 

This method is also presented in TNO (1997) and Conrado (2000), and Appendix A provides a 

sample calculation. In addition to the thermophysical properties of the substance, this method 

depends upon the water temperature. For a water temperature of 68 oF (20 oC), the method estimates 

a heat flux of 11,600 BTU/ft2

/hr (37 kW/m2), which will generate an evaporation rate of 0.015 lb/ft2/s 

(0.072 kg/m2

/s). 

Of course, the heat transfer rate must also be used along with an appropriate method for 

estimating pool spread, as discussed in Section 2.2. In the recommended spread model, heat flux to 

the pool directly drives evaporation and, since the evaporation removes material from the pool, it also 
18 

 

affects the rate and extent of pool spread, which in turn affects the downwind distance for flammable 

vapor travel. 

Given the relative uncertainty in the choice of heat flux, it is useful to examine the sensitivity of 

final consequence analysis results (such as distance to the LFL predicted by an atmospheric 

dispersion model) to the heat flux. To examine the sensitivity, calculations were performed with 

three different heat flux values to see the effects on pool radius, evaporation time, and dispersion 

distance. The heat flux values examined are 7,900 BTU/ft2

/hr (25 kW/m2), 11,600 BTU/ft2/hr (37 

kW/m2

), and 31,700 BTU/ft2/hr (100 kW/m2). The scenario and results are as follows: 

Scenario for Evaporation Heat Flux Sensitivity Analysis

Hole diameter: 3.3 ft (1 m) 

Initial liquid height above hole: 43 ft (13 m) 

Total spill quantity: 4.4 × 105

 ft3 (12,500 m3) 

Air temperature: 71 °F (22 °C) 

Relative humidity: 50% 

Wind speed: 6.7 mph (3.0 m/s) and 4.5 mph (2.0 m/s) 

Surface roughness: 0.03 ft (0.01 m) 

Pasquill-Turner stability class: D and F 

Averaging time: 0 sec (i.e., peak concentrations are used) 

Results for Evaporation Heat Flux Sensitivity Analysis

Hole diameter 3.3 ft (1 m) 

Initial spill rate 11,700 lb/s (5,300 kg/s) 

Total spill duration 33 min 

Spread model Webber 

Film boiling heat flux to pool 7,900 BTU/hr/ft2

(25 kW/m2) 

11,700 BTU/hr/ft2

(37 kW/m2) 

31,700 BTU/hr/ft2

(100 kW/m2) 

Evaporation rate 0.010 lb/ft2/s 

(0.049 kg/m2

/s) 

0.015 lb/ft2

/s 

(0.072 kg/m2

/s) 

0.040 lb/ft2

/s 

(0.20 kg/m2

/s) 

Maximum pool radius 500 ft 

(150 m) 

430 ft 

(130 m) 

280 ft 

(87 m) 

Total evaporation duration 34 min 34 min 34 min 

Wind speed and stability class 6.7 mph (3.0 m/s) and D stability 

Downwind distance to LFL 6,200 ft 

(1,900 m) 

6,600 ft 

(2,000 m) 

4,300 ft 

(1,300 m) 

Time at which LFL reaches 

maximum distance 

38 min 33 min 33 min 

Time at which entire cloud drops 

below LFL 

42 min 37 min 37 min 
19 

 

Results for Evaporation Heat Flux Sensitivity Analysis (cont’d) 

Wind speed and stability class 4.5 mph (2.0 m/s) and F stability 

Downwind distance to LFL 9,200 ft 

(2,800 m) 

11,000 ft 

(3,300 m) 

5,600 ft 

(1,700 m) 

Time at which LFL reaches 

maximum distance 

38 min 44 min 26 min 

Time at which entire cloud 

drops below LFL 

43 min 48 min 29 min 

The downwind distance to the LFL is the maximum distance at which some of the released LNG 

vapor is within the flammable range. Note that the vapor cloud continues to travel downwind and 

disperse even after the spill stops and the pool completely evaporates. 

This example shows that the chosen heat flux affects the distance to the LFL and the time at 

which the maximum distance is reached.. For the heat flux values examined, downwind distances to 

the LFL vary by approximately 54% in the case with D stability and approximately 94% in the case 

with F stability. This example also illustrates the importance of selecting a best-estimate value for 

the heat flux (as opposed to a “conservatively high” estimate). In this case, the heat flux value of 

11,700 BTU/hr/ft2

 (37 kW/m2), as calculated using heat transfer theory, results in the longest 

downwind distances. 

This example also illustrates the importance of carefully characterizing the release source. 

Dispersion results are sensitive to mass evaporation rate, source area, and release duration. In 

addition, the example shows the importance of considering the transient nature of the large spills 

addressed in this report. In this case, the total evaporation duration is approximately the same in all 

three cases, but with the higher heat flux values, mass evaporation increases more rapidly, and the 

majority of the total quantity is released in a shorter period of time. This effect is particularly evident 

in this example because the spill rate decreases with time (i.e., a large fraction of the total spill 

amount is added to the pool early in the release). 

While this example provides valuable insights about the recommended methodology, it is 

important to note that it does not represent a comprehensive sensitivity analysis; sensitivity may be 

different for other types of scenarios and with other input parameter values. 

2.4 THERMAL RADIATION FROM POOL FIRES ON WATER 

A pool fire can result when LNG is spilled on water, forming a pool, and vapor evolving from the 

pool is ignited. The fire above the pool continues to be fueled by evaporation from the pool. The 

flames that form the pool fire emit thermal radiation, which can injure people and damage structures. 

On water, LNG will spread and burn rapidly, so the fire will typically burn out shortly following the 

end of the spill into the pool, at which time the thermal radiation hazard will cease. This section 

presents an evaluation of the methods for estimating the distances at which thermal radiation levels 

of concern might be reached if an LNG pool fire occurs. 
20 

 

2.4.1 Review of Modeling Approaches 

Fay (2003) – For the example calculations presented in this paper, the point source model is 

employed. In this model, the energy generated by combustion is assumed to emanate from a single 

point at the center of the pool fire. A key parameter for this model is the fraction of heat that is 

emitted as thermal radiation. As Fay correctly points out, this point source model only applies at 

large distances from the pool fire. A burn regression rate for pool fires is taken as 8 × 10-4 m/s 

(evaporation rate of 0.34 kg/m2

/s). This value is based on contributions to evaporation caused by 

heat from the fire, giving a regression rate of 1.9 × 10-4 m/s (evaporation rate of 0.080 kg/m2

/s), and 

heat from the water, giving a regression rate of 5 × 10-4 to 7 × 10-4 m/s (evaporation rate of 0.21 to 

0.30 kg/m2

/s). The fraction of heat that is emitted as thermal radiation is taken as 0.15. 

Lehr (2004) – Lehr employs the solid flame model. The fire is represented as an upright 

cylinder, and a form of the Thomas equation (Brown, 1974) is used to estimate the flame (cylinder) 

height. Geometric view factors are used to compute the heat flux at a given receptor location, and 

attenuation of the thermal radiation by the atmosphere is estimated using the work of Glastone 

(1977). The paper states that experiments involving LNG pool fires on water show burn regression 

rates that vary from 4 × 10-4 to 1 × 10-3 m/s (evaporation rate of 0.17 to 0.42 kg/m2

/s) and references 

Raj (1979). It further states that for larger continuous releases, the burn rate has been estimated as 2.5 

× 10-4 m/s (evaporation rate of 0.11 kg/m2/s); for this value, the report references an unpublished 

internal memo. 

Quest (2001) – Quest also employs the solid flame model; however, in this case the fire is 

represented as an upright (but tilted) elliptical cylinder. The elliptical shape and tilt attempt to 

account for the effects of wind on the fire, which can increase the distance that damaging thermal 

radiation can reach in the downwind direction. The Thomas (1965) correlation is used to calculate 

flame length. Flame tilt is estimated using an empirical correlation from Welker (1970), and flame 

drag (increasing the downwind dimension of the base of the flame) is calculated based on the work of 

Moorhouse (1982). Quest also divides the flame into two zones: a clear zone and a smoky zone, in 

which a portion of the flame is obscured by smoke. The length of the clear zone is calculated based 

on work by Prichard (1992). The surface flux is calculated using an equation that depends on the 

maximum surface flux, an extinction coefficient, and the pool diameter; however, no information is 

given on the values used for the maximum surface flux or extinction coefficient. In the smoky zone, 

the surface flux for the smoky areas is taken as 20 kW/m2

, and the clear areas are taken as having the 

same surface flux as the clean burning zone. The fraction of smoke in the smoky zone is taken as 

10%. Geometric view factors are used to compute the heat flux at a given receptor location, and 

attenuation of the thermal radiation by the atmosphere is estimated using the work of Wayne (1991). 

Other Methods – Available pool fire models for hazard assessment purposes are predominantly 

based on the solid flame model. The fire is represented as gray-body emitter, and the geometry is 

approximated by an upright cylinder that is tilted as a result of wind. Geometric view factors are 

available to calculate the thermal radiation that a given receptor would be exposed to, based on the 

geometry of the cylinder. For hazard assessment purposes, hazard zones are often depicted as 

circles, based on the maximum downwind distance to the level of concern (i.e., in the direction that 

the flame is tilted) and centered on the base of the pool. However, it can be noted that in an actual 

fire, flame tilt will result in a hazard zone with a roughly elliptical pattern, with upwind distances 

being shorter than downwind distances. 
21 

 

Detailed descriptions of solid flame models are available in many references. Perhaps the best 

overall reference on the topic is Section 3/Chapter 11 of NFPA (1995); others include TNO (1997) 

and Rew (1996). 

The point source model is also sometimes used in hazard analysis. As mentioned above, in this 

model the energy generated by combustion is assumed to emanate from a single point at the center of 

the pool fire. The calculated energy is multiplied by a fraction that accounts for the fact that only 

part of the energy will be emitted as thermal radiation. 

Fire modeling can also be performed using computational fluid dynamics (CFD) methods (e.g., 

finite element analysis). These methods are often based on numerical solutions to the Navier-Stokes 

equations but also often require use of some empirical data. CFD models are not normally used for 

typical pool fire hazard assessments because they require significantly more effort to apply but 

provide little or no benefit over the solid flame model when the goal is prediction of heat flux at 

significant distances from the fire. CFD models do have a distinct advantage in cases where it is 

necessary to model effects on objects engulfed in fire and in modeling fires with irregular geometry. 

Detailed examination of CFD methods is beyond the scope of this study. 

While LNG has been studied more than many substances, there are still relatively few data 

available for use in determining the behavior of LNG pool fires. This is particularly true for very 

large pool fires on water. No correlations are available to predict burning rate or surface emission 

power for various pool sizes. 

The data identified in this review are as follows: 

From TNO (1997) 

• Thomas (1963) correlation for flame length 

• For an experimental LNG pool fire of 16.8-m diameter, a mass burning flux of 

0.255 kg/s/m2

 is given (TNO Table 6.10) 

• Mass burning fluxes for several flammable materials vs. pool diameter on water (TNO Table 

6.11) 

o LNG on water – up to 13.7-m diameter: 0.94 kg/s/m2

o LNG on water – 20-m diameter: 0.106 kg/s/m2

o LNG on water – 35-m diameter: 0.14 kg/s/m2

From NFPA (1995) 

• For unconfined, continuous spills, diameter will increase until burning rate matches spill rate 

• Summary of Radiation Data on Hydrocarbon Pool Fires (NFPA Table 3-11.2) 

o Raj (1979) – 10 to 15 m pools: burning rate 4 to 10 x 10-4 m/s, 1400 K, average 

emissive power of 220 kW/m2

o Mizner (1983) – 20 m (diameter), LNG and LPG, burning rate 2.4 x 10-4 m/s 

• Measured Emissive Powers and Radiation Temperatures for Various Liquid Hydrocarbon 

Pool Fires (NFPA Table 3-11.4) 

o LNG on water: 8.5 to 15.0 m, emissive power 210 to 280 kW/m2

, radiation 

temperature 1500 K (estimated using narrow-angle radiometer data and spectral data) 
22 

 

From Mizner (1983) 

• Regarding China Lake tests, states that “From the LNG tests, surface emissive powers were 

determined to be about 210 kW/m2

 for both the pool and cloud fires” 

• Data reported for four tests at Maplin Sands in 1989: three continuous and one instantaneous. 

All tests involved a cloud fire, and only one involved a pool fire. Report indicates that the 

one pool fire probably did not last long enough to become fully developed. Continuous 

releases were 3.2, 5.8, and 4.7 m3

/s. Instantaneous release was 12 m3. Wind speeds ranged 

from 4 to 6 m/s 

• Surface emissive powers reported in “range from 137 kW/m2 to 225 kW/m2 with an overall 

mean of 174 kW/m2

” 

• For Trial 39 (the only pool fire), emissive powers ranged from 178 kW/m2 to 248 kW/m2, 

with a mean of 203 kW/m2

. Estimated pool diameter was approximately 30 m, flame length 

ranged from 50 to 78 m, and tilt angle from 27 to 35 degrees 

From Rew (1996) 

• Work included a review of recent developments in hydrocarbon pool fire modeling and 

development of a computer program called POOLFIRE6 

• For LNG on water, they chose the following parameters, based on a review of available test 

data: 

o Maximum burning rate: 0.282 kg/s/m2

o Emissive power: 265 kW/m2

2.4.2 Modeling Recommendations 

While the point source model is computationally easier to apply, its chief disadvantage is that it 

does not produce valid estimates for receptors close to the fire. At long distances (and assuming 

other parameters such as burning rate, emissive power, and fraction of heat radiated are matched), 

both the point source model and the solid flame model will provide equivalent results. Therefore, the 

solid flame model is recommended, with the pool fire represented as a tilted cylinder. 

Based on the literature reviewed, the values chosen by Rew (1996) for burning rate and emissive 

power are recommended, namely a burning rate of 0.058 lb/s/ft2

 (0.282 kg/s/m2) and emissive power 

of 84,000 BTU/hr/ft2

 (265 kW/m2). The authors performed a reasonable review of the experimental 

data, and our review for this work concurs with their choices. Comparisons only to data for the 

larger scale spill tests tend to indicate that these are relatively conservative choices. Remaining 

details of the recommended approach (e.g., correlations for flame height and tilt) are given in the 

example calculations in Appendix C. 

It is also important to note that the methods recommended here have generally been validated 

against a broad range of the available experimental data. However, there are no experimental data 

for pool fires as large as some of the sizes postulated for LNG carrier incidents. These are the best 

currently available methods, but it is important to keep in mind that the methods are being used well 

outside the range where they have been validated. 

For rapid releases onto an unconfined water surface, pool spread can be estimated using a pool 

spread model, as described in Section 2.2. However, in the case of a fire, heat transfer to the pool 

from both the water and the fire itself must be included. The additional heat transfer (from the fire) 

results in greater evaporation than in the nonfire case; this greater evaporation rate can affect pool 
23 

 

spread. The chosen burning rate of 0.058 lb/s/ft2 (0.282 kg/s/m2) already accounts for both of these 

contributions and corresponds to a heat transfer rate of approximately 46,000 BTU/ft2

/hr (140 

kW/m2

). The model accounts for the resulting evaporation in estimating the pool spread. Given the 

relative uncertainty in the choices of burning rate and surface emitted flux, it is useful to examine the 

sensitivity of final consequence analysis results (such as distance to a heat flux level of concern) to 

the values of these parameters. To examine the sensitivity, calculations were performed using the 

methods recommended in this report for an example scenario. The burning rates examined are 0.041 

lb/s/ft2

 (0.20 kg/s/m2), 0.057 lb/s/ft2 (0.28 kg/s/m2), and 0.074 lb/s/ft2 (0.36 kg/s/m2). The surface 

emitted fluxes examined are 63,000 BTU/ft2

/hr (200 kW/m2), 84,000 BTU/ft2/hr (265 kW/m2), and 

95,000 BTU/ft2

/hr (300 kW/m2). The scenario and results are as follows: 

Scenario for Burning Rate and Surface Emitted Flux Sensitivity Analysis

Hole diameter: 3.3 ft (1 m) 

Initial liquid height above hole: 43 ft (13 m) 

Total spill quantity: 4.4 × 105

 ft3 (12,500 m3) 

Air temperature: 80 °F (27 °C) 

Relative humidity: 70% 

Wind speed: 20 mph (8.9 m/s) 

Results for Burning Rate Sensitivity Analysis

Hole diameter 3.3 ft (1 m) 

Initial spill rate 11,700 lb/s (5,300 kg/s) 

Total spill duration 33 min 

Spread model Webber 

Flame surface emitted flux 84,000 BTU/ft2/hr (265 kW/m2) 

Burning rate 0.041 lb/s/ft2

(0.20 kg/s/m2) 

0.057 lb/s/ft2

(0.28 kg/s/m2) 

0.074 lb/s/ft2

(0.36 kg/s/m2) 

Maximum pool radius 280 ft 

(90 m) 

240 ft 

(74 m) 

220 ft 

(66 m) 

Total fire duration 33 min 33 min 33 min 

Flame length (height) 790 ft 

(240 m) 

910 ft 

(280 m) 

1,000 ft 

(310 m) 

Flame tilt at maximum radius 34 deg 35 deg 37 deg 

Downwind distance to 

12,000 BTU/hr/ft2 (38 kW/m2) 

1,300 ft 

(380 m) 

1,200 ft 

(370 m) 

1,200 ft 

(370 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

1,500 ft 

(450 m) 

1,500 ft 

(450 m) 

1,500 ft 

(440 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

2,000 ft 

(600 m) 

2,000 ft 

(600 m) 

2,000 ft 

(600 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

2,800 ft 

(860 m) 

2,800 ft 

(860 m) 

2,800 ft 

(850 m) 
24 

 

Results for Surface Emitted Flux Sensitivity Analysis

Hole diameter 3.3 ft (1 m) 

Initial spill rate 11,700 lb/s (5,300 kg/s) 

Total spill duration 33 min 

Spread model Webber 

Burning rate 0.057 lb/s/ft2 (0.28 kg/s/m2) 

Maximum pool radius 240 ft (74 m) 

Total fire duration 33 min 

Flame length (height) 910 ft (280 m) 

Flame tilt at maximum radius 36 deg 

Flame surface emitted flux 63,000 BTU/ft2/hr 

(200 kW/m2

) 

84,000 BTU/ft2

/hr 

(265 kW/m2

) 

95,000 BTU/ft2

/hr 

(300 kW/m2

) 

Downwind distance to 

12,000 BTU/hr/ft2 (38 kW/m2) 

1,100 ft 

(330 m) 

1,200 ft 

(370 m) 

1,300 ft 

(400 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

1,300 ft 

(400 m) 

1,500 ft 

(450 m) 

1,600 ft 

(470 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

1,800 ft 

(540 m) 

2,000 ft 

(600 m) 

2,100 ft 

(630 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

2,500 ft 

(760 m) 

2,800 ft 

(860 m) 

3,000 ft 

(900 m) 

The results for burning rate sensitivity show that the distance estimates are not highly sensitive to 

burning rate. There are competing factors at work here: 

• As the burning rate rises, the maximum pool radius drops (since the LNG evaporates more 

quickly, reducing the opportunity for spread). 

• Flame length is a function of both burning rate and pool radius. Increasing either of these 

results in a longer flame length. In the example above, the effect of burning rate dominates 

this competition, and flame length drops with burning rate. 

In the example above, increasing burning rate also increases flame length but decreases pool 

radius. These are competing effects on estimated flux at downwind distances, and in this example, 

the effect of increased pool diameter slightly dominates this competition. As a result, increasing 

burning rate causes a decrease in the estimated distances to thermal flux levels of concern. For other 

scenarios, distances may increase with increased burning rate, but as in this example, the distances 

are expected to vary by relatively small percentages over this range of burning rates. 

The results for surface emitted flux sensitivity show that the distance estimates are affected by 

the value chosen. However, increasing the value by 50%, from 63,000 BTU/ft2

/hr (200 kW/m2) to 

95,000 BTU/ft2

/hr (300 kW/m2), increases the distance estimates by only 18% to 21%. 

So, while the burning rate and surface emitted flux are important parameters in the pool fire 

analysis, it is useful to note that the distance results are not highly sensitive to the values chosen. 
25 

 

In these examples, it can also be noted that the total fire duration matches the total spill duration. 

This occurs because the pool spreads to the point where the burn rate matches the spill rate into the 

pool, so following the initial spreading of the pool, the LNG burns at the same rate that it is spilled. 

While this example provides valuable insights about the recommended methodology, it is 

important to note that it does not represent a comprehensive sensitivity analysis; sensitivity may be 

different for other types of scenarios and with other input parameter values. 

This work did not include a comprehensive review of computer software for performing pool fire 

consequence assessment. However, the majority of software currently available for modeling pool 

fires is based on the same basic approach recommended here (solid flame model, tilted cylinder). 

This includes software packages like LNGFIRE and POOLFIRE6. However, as was stated in the 

case of software for pool spread estimation in Section 2.2.2, existing software for pool fire modeling 

may be difficult to apply to large LNG spills on water if they provide limited ability to specify input 

parameters. 

2.5 FLAMMABLE VAPOR DISPERSION FOLLOWING SPILLS ON WATER 

If LNG is spilled on water but not ignited immediately or soon after the spill, the LNG will 

evaporate. The vapor will mix with air and move downwind. The resulting vapor cloud could cause 

a flash fire if ignited before it is diluted below the LFL. This section presents an evaluation of the 

methods for estimating distances that flammable vapors might reach. 

2.5.1 Review of Modeling Approaches 

Fay (2003) – This paper does not address vapor dispersion. 

Lehr (2004) – This paper does not address vapor dispersion. 

Quest (2001) – To estimate the distance to the LFL, Quest used a dispersion modeling software 

package that it sells commercially, which it states is based on the SLAB model. The DEGADIS 

dispersion model was also used. 

Other Methods – For flash fires, the level of concern is typically defined as either the LFL or 1⁄2 

the LFL for the substance. Therefore, modeling of flash fires is primarily a matter of applying a 

dispersion model. This work did not involve a detailed review of dispersion models, but it is clear 

that for large releases of LNG, dense-gas effects will be important and must be accounted for in the 

model used. Several such models are available in the form of freely available software, including 

DEGADIS, SLAB, and HGSYSTEM. In addition, several commercial software packages include 

models for dense-gas effects. 

For further information, reviews and validation studies for various dispersion models have been 

conducted by a variety of organizations. Examples include EPA (1991) and API (1992). 

2.5.2 Modeling Recommendations 

When considering large releases of LNG, dense-gas effects are important and must be accounted 

for in any model selected. Among the dispersion models appropriate for large LNG spills, 
26 

 

DEGADIS was selected for use (see example problems in Section 3). DEGADIS accounts for densegas effects and was originally developed for simulation of cryogenic flammable gas dispersion, 

particularly for LNG. It has been validated against a wide range of laboratory and field test data. 

Chapter 6 of GRI (1990) provides information on the comparison of DEGADIS predictions with 

field test data from the Burro and Maplin Sands tests. In addition, API (1992) indicates that 

DEGADIS tends to overpredict by less than a factor of two for simulations with short averaging 

times. This is based on comparisons with field test data, including data from the Burro, Coyote, 

Desert Tortoise, Goldfish, Maplin Sands, and Thorney Island tests. 

Furthermore, the federal siting requirements for onshore LNG facilities (49 CFR 193) specify the 

use of DEGADIS for the determination of dispersion distances. As recognized by 49 CFR 193, 

DEGADIS does not account for additional dispersion caused by “the complex flow patterns induced 

by tank and dike structure(s).” For cases where dense gas flow obstructions are a concern, the 

regulation allows dispersion distances to be calculated with the FEM3A model (see GRI 1997). 

However, for most hazard assessment purposes, including facility siting analyses, modeling of 

complex terrain is not necessary. 

For flammable vapor dispersion distance calculations, the level of concern is sometimes taken as 

the LFL, and there is some regulatory basis for this choice in EPA’s risk management program rule 

(40 CFR 68). However, the level of concern is also often defined as 1⁄2 the LFL to account for 

localized pockets of higher gas concentrations that may occur in an actual release. Use of 1⁄2 the LFL 

for LNG is also supported by 49 CFR 193, which specifies use of an average gas concentration in air 

of 2.5 percent for onshore exclusion zones. For the example calculations presented in Section 3, 

distances are provided for both the LFL and 1⁄2 the LFL. 

As is the case any time one models dispersion of flammables, it is recommended that a short 

averaging time (not more than a few seconds) be used because a flammable cloud need only be 

within the flammable range for a very short time to be ignited. It is also important to appropriately 

represent the topography downwind of the release point. For many dispersion models, topography is 

characterized by a surface roughness value, which the model uses in accounting for the effects of 

terrain on dispersion of the cloud. In general, a rougher surface will tend to cause more mixing with 

ambient air, and therefore more rapid dispersion of a vapor cloud. 

2.6 RAPID PHASE TRANSITIONS 

An RPT may occur during an LNG spill when a portion of the pool changes from liquid to gas, 

virtually instantaneously. This vaporization will rapidly expel LNG vapor into the atmosphere and 

can generate an overpressure. The vaporization may also increase the downwind distance where the 

LFL is reached or it may affect the size of a pool fire. RPTs have occurred during LNG spill tests on 

water. An RPT is a physical explosion of a liquid that does not involve fire or other chemical 

reactions. 

2.6.1 Literature Review 

Fay (2003), Lehr (2004), and Quest (2001) – These documents do not address RPTs.
27 

 

Other Literature – The literature reveals that RPTs are a poorly understood phenomenon. They 

have occurred in small and large simulated spill releases of LNG on water. The water is very hot with 

respect to the LNG, which has a normal boiling point of approximately -260 °F. Typical sea water 

temperatures range from 30 °F to 80 °F. The large temperature difference between the sea water and 

the LNG results in a very rapid vaporization of the LNG. 

Sometimes, the RPT occurs after the spill, as the LNG vaporizes. Some have speculated that the 

differential boiling (i.e., change in composition as methane boils off first) may play a role in the RPT 

initiation. In some incidents, more than one RPT has occurred during a single spill event. 

Although the physical mechanism for the rapid vaporization is not well understood, if it proceeds 

fast enough, an overpressure wave may be formed. In some test cases, overpressures generated were 

strong enough to damage test equipment in the immediate vicinity of the LNG release point. The 

sizes of the overpressure events have been generally small, with a maximum TNT equivalence 

estimated at several pounds of TNT. Such a small overpressure event is not expected to cause 

significant damage to an LNG vessel. However, the RPT may increase the rate of LNG pool 

spreading and the LNG vaporization rate. 

Atallah (1997) surveyed the literature on RPTs for the Gas Research Institute. Atallah's study 

reviewed available experimental and theoretical studies of RPTs for liquid metal-water and LNGwater systems. Atallah’s work provides a good summary of what is known about RPTs. 

The survey of the LNG-water releases yields the following conclusions: 

1. Rapid phase transitions have been observed in many field tests of spills of LNG on a water 

surface 

2. Increasing the spill rate increases the effective explosive yield of an RPT event (Atallah, pg. 

3-17, Figure 3.7) 

3. The largest explosive yield of an RPT event observed in a field test was equivalent to an 

explosion of 5.5 kg of TNT (Atallah, pg. 3-21, Table 3.11) 

4. RPT events did not occur in every spill test 

5. No existing theory for RPT events can account for all of the observed test data 

6. There is no existing theoretical method to predict the severity of an RPT event involving the 

release of the LNG from an LNG carrier onto water 

2.6.2 Modeling Recommendations 

Modeling of RPTs could be useful for two types of effects: (1) overpressure resulting from the 

rapid phase change and (2) dispersion of the “puff” of LNG expelled into the atmosphere by the 

RPT. 

Although the possibility of an RPT igniting a vapor cloud was mentioned by Vallejo (2003), the 

experimental results reviewed do not indicate that ignition is likely. Unfortunately, no theoretical or 

experimental basis could be identified for modeling either the overpressure or source term effects. 

Regarding overpressure, the literature suggests that damage might be limited to the immediate 

vicinity of the RPT event; that is, that overpressures decrease rapidly with distance. On the other 

hand, none of the spill tests reported involve spills of the magnitude of interest in this report. 

Regarding the “puff” of LNG entering the atmosphere in an RPT, an example dispersion analysis 

was performed. The purpose of this analysis is to examine the importance of an RPT in estimating 
28 

 

downwind flammable concentrations. The analysis looks at three scenarios: (1) a base case releasing 

1.8 × 105

 ft3 (5,000 m3) through a 9.8-ft (3-m) hole without an RPT occurring, (2) the base case 

modified by assuming that an RPT occurs at 2 minutes and the entire mass remaining in the pool is 

released to the atmosphere within 30 seconds and that the source area used as input into a dispersion 

model is twice the pool area at the time of the RPT, and (3) the base case modified by assuming that 

an RPT occurs at 2 minutes and the entire mass remaining in the pool is released to the atmosphere 

within 30 seconds with the source area remaining equal to the pool area at the time of the RPT. The 

source area is the area over which the vapor cloud originates; in the case of an evaporating pool 

(without an RPT), this is simply the pool area. 

These scenarios are designed to at least roughly follow some of the anecdotal information from 

the literature, in particular, the cases where a spill occurs and the RPT follows within a couple of 

minutes. This example is pursued simply to explore the possible effects of RPT on dispersion of a 

flammable cloud; there is no precedent in the literature for the chosen modeling approach. 

The scenario parameters and results are as follows: 

Scenario for Analysis of RPT Influence on Vapor Dispersion

Hole diameter: 9.8 ft (3 m) 

Total spill quantity: 1.8 × 105

 ft3 (5,000 m3) 

Air temperature: 70 °F (21 °C) 

Water temperature: 70 °F (21 °C) 

Relative humidity: 50% 

Wind speed: 6.7 mph (3.0 m/s) 

Surface roughness: 0.03 ft (0.01 m) 

Pasquill-Turner stability class: D 

Averaging time: 0 sec (i.e., peak concentrations are used) 

Results for Analysis of RPT Influence on Vapor Dispersion

Scenario 

Base Case 

(No RPT) 

RPT at 2 min with 

Pool Radius 

Doubled 

RPT at 2 min with Pool 

Radius Held Constant 

Maximum pool radius 

(from spread calculations) 

410 ft 

(125 m) 

310 ft (93 m) 

spread at 2 min 

310 ft (93 m) 

spread at 2 min 

Maximum source radius for 

dispersion modeling 

410 ft 

(125 m) 

610 ft (186 m) 

twice pool radius at 

2 min 

310 ft (93 m) 

pool radius at 2 min 

Total evaporation duration 13 min 2.5 min 2.5 min 

Downwind distance to LFL 6,600 ft 

(2,000 m) 

7,500 ft 

(2,300 m) 

7,500 ft 

(2,300 m) 
29 

 

In this example, the “puff” of LNG released into the atmosphere by the RPT does have some 

impact on downwind distances since it generates vapor more rapidly and results in longer distances 

to the LFL. However, the assumptions here are probably more severe than would be expected in an 

actual RPT since the majority of the original volume goes into the atmosphere in a 30-second period. 

This is not a comprehensive analysis; results could be different for different scenarios and/or input 

parameters. However, it gives some indication that for scenarios like this short-term release, an RPT 

may not result in estimates of dispersion distances that are much longer than what is expected with 

similar scenarios that do not involve RPTs. 

2.7 EFFECTS OF THERMAL RADIATION ON PEOPLE 

This section provides an overview of data and methods for estimating the effects on people and 

structures that result from thermal radiation exposure. 

2.7.1 Review of Data and Effects Models 

The extent to which people are injured by exposure to thermal radiation depends on both the 

incident heat flux and the exposure time. The data used in estimating effects on people include those 

from: 

• Experiments with humans (at lower levels of thermal radiation) 

• Experiments with animals 

• Review of historical data from actual thermal radiation burn incidents 

Based on these data, various researchers have presented estimates of effects primarily in the form 

of (1) tabulated radiation levels and exposure times for various effects and (2) probability unit 

(probit) models that allow estimating the probability of various effects, typically based on a thermal 

load (i.e., the time integral of radiation intensity to an empirically derived power). 

Detailed summaries of available data and models are provided in references such as Lees (1996) 

and TNO (1992). This report does not provide as detailed a summary as these references; rather, it 

presents examples of data that are appropriate for hydrocarbon fires and are frequently used when 

selecting levels of concern for hazard assessment purposes, such as facility siting analyses. 

Tabulated effects for various radiation levels and exposure times include the following: 

• Table 2.2 – Burn injury criteria from the Federal Emergency Management Agency (FEMA, 

1990) 

• Table 2.3 – Exposure limits recommended by API (1997) for exposure of workers to thermal 

radiation from flares 

• Table 2.4 – Burn injury estimates for fireballs given by Prugh (1994). Note that these data 

are based on thermal dose, which is the time integral of radiation intensity. These data are 

also based on short-duration exposures (i.e., on the order of a minute or less) 
30 

 

Table 2.2 Thermal Radiation Burn Injury Criteria from FEMA (1990) 

Thermal Radiation Intensity 

BTU/hr/ft2 kW/m2

Time for Severe 

Pain (sec) 

Time for Seconddegree Burns (sec) 

300 1 115 663 

600 2 45 187 

1000 3 27 92 

1300 4 18 57 

1600 5 13 40 

1900 6 11 30 

2500 8 7 20 

3200 10 5 14 

3800 12 4 11 

Table 2.3 Permissible Thermal Radiation Exposure for Flares from API 521 (1997) 

Thermal Radiation Intensity 

BTU/hr/ft2 kW/m2 Type of Damage 

500 1.6 Permissible level at any location where personnel are continuously 

exposed 

1,500 4.7 Permissible level in areas where emergency actions lasting several 

minutes may be required by personnel without shielding but with 

appropriate clothing 

2,000 6.3 Permissible level in areas where emergency actions lasting up to 1 

minute may be required by personnel without shielding but with 

appropriate clothing 

3,000 9.5 Permissible level in areas where exposure to personnel is limited to 

a few seconds, sufficient for escape only 

The values shown in Table 2.3 are measures of heat per unit area as a function of time. Table 2.4 

shows cumulative heat received over time. For example, exposure to a flux level of 5 kW/m2

 (approx. 

1,500 BTU/hr/ft2) for 30 seconds results in a thermal dose of 150 kJ/m2 and causes second-degree burns.
Table 2.4 Thermal Dose Data for Exposure to Fireballs from Prugh (1994) 

Thermal Dose 

kJ/m2 BTU/ft2 Type of Injury 

40 3.5 Threshold of pain

100 8.8 Sunburn (first-degree burn)

150 13 Blisters (second-degree burn) 

250 22 1% fatal (third-degree burn)

500 44 50% fatal (third-degree burn) 

1200 106 99% fatal (third-degree burn) 

Available probit models include the relations shown in Table 2.5 (TNO, 1992). In these 

empirically derived relations, 3

4

tI is referred to as the thermal load. The probit value (Pr) is a 

Gaussian-distributed, random variable with a mean value of 5.0 and a standard deviation of 1.0. The 

probit value is related to probability by the following: 

du

2

1 Probability

Pr 5

2

u2

∫

−

−∞

−

= e

π

Table 2.5 Probit Models for Injury by Thermal Radiation from TNO (1992) 

Probit Equation Effects 

Pr 39.83 3.0186 ln( ) 3

4

= − + tI First-degree burns 

Pr 43.14 3.0186ln( ) 3

4

= − + tI Second-degree burns 

Pr 36.38 2.56ln( ) 3

4

= − + tI Lethality

In addition to radiation intensity and time, it is important to keep in mind that other factors can 

strongly affect the extent of injury to people that is caused by thermal radiation: 

• Protection afforded by shelter 

• Protection afforded by clothing 

• Contribution of solar radiation to total exposure (250-330 Btu/hr-ft2) 

• Susceptibility of individual exposed 

• Response of individual (e.g., ability to take shelter) 

31 

 
32 

 

2.7.2 Selecting Levels of Concern 

As in all hazard assessment activities, thermal radiation levels of concern must be chosen with 

the nature of (1) the potentially exposed population and (2) the potential fire events. For example, 

permissible levels could be defined differently for (1) workers in a process plant who are wearing 

protective clothing, (2) areas where people are generally not present but could have access, or (3) 

sensitive populations, such as the elderly. 

For purposes of onshore facility siting analysis, 49 CFR 193 and NFPA 59A specify a level of 

concern of 1,600 BTU/hr/ft2

 (5 kW/m2). Based on the information presented in Section 2.7.1, 1,600 

BTU/hr/ft2

 (5 kW/m2) could be expected to cause the effects summarized in Table 2.6. 

Table 2.6 Effects on People for 1,600 BTU/hr/ft2

 (5 kW/m2) Thermal Radiation 

Effect 

Exposure Time 

(seconds) Data Source 

Severe pain 13 Table 2.2 

First-degree burns 20 Table 2.4 

(5 kW/m2

 for 20 seconds corresponds to a 

thermal dose of 100 kJ/m2

) 

30 Table 2.4 

(5 kW/m2

 for 30 seconds corresponds to a 

thermal dose of 150 kJ/m2) 

Second-degree burns 

40 Table 2.2 

Third-degree burns (1% 

fatality) 

50 Table 2.4 

(5 kW/m2

 for 50 seconds corresponds to a 

thermal dose of 250 kJ/m2

) 

72% probability of first-degree 

burns 

40 Table 2.5 

This level of 1,600 BTU/hr/ft2

 (5 kW/m2) is applicable to short duration events, such as fireballs. 

It is also applicable to the early stages of longer duration events, such as pool fires, provided that the 

potentially exposed population will have both opportunity and capability to quickly take cover. 

2.8 EFFECTS OF THERMAL RADIATION ON STRUCTURES 

2.8.1 Review of Data and Effects Models 

Like effects of thermal radiation on people, effects on structures also depend on incident heat 

flux and the exposure time. With structures, effects also depend strongly on the materials of 

construction (e.g., wood, steel, concrete). 

Tabulated effects for various radiation levels and exposure times include the following: 

• Table 2.7 – Damage criteria from the World Bank (1988) based on observations of large fires 

• Table 2.8 – Critical intensity levels presented by TNO (1992) 

• Table 2.9 – Various thermal radiation limits presented by Lees (1996) 
33 

 

Unless otherwise noted, these tabulated intensity levels and effects are applicable for long-term 

exposure, which is very generally considered to be on the order of 30 minutes or more. 

Table 2.7 Structure Damage Criteria for Thermal Radiation Exposure from World Bank 

(1988) 

Thermal Radiation Intensity 

BTU/hr/ft2 kW/m2 Type of Damage 

11,890 37.5 Sufficient to cause damage to process equipment (e.g., steel 

structures and equipment) 

7,930 25.0 Minimum energy required to ignite wood at indefinitely long 

exposures without a flame 

3,960 12.5 Minimum energy required to ignite wood 

with a flame or cause melting of plastic tubing 

Table 2.8 Damage Resulting from Thermal Radiation for Various Materials from TNO (1992) 

Critical Radiation Intensity1 (kW/m2) 

Type of Damage Damage Level 12 Damage Level 23

Steel 100 25 

Wood 15 2 

Synthetic materials 15 2 

Glass 4 – 1 Radiation intensity that can cause damage from long-term exposure. 2 Damage Level 1: Surfaces of exposed materials catch fire and structural elements collapse or rupture. 3 Damage Level 2: Surfaces of exposed materials experience serious discoloration as well as paint peeling, and 

structural elements undergo substantial deformation. 

A more accurate assessment of structural damage requires performing heat balance calculations 

to determine the resulting surface temperature for comparison with acceptable temperatures for the 

particular structure. TNO (1992) presents a discussion and example of such calculations. In addition, 

Chapters 9, 10, and 11 of NFPA (1995) present analytical methods for determining fire resistance of 

steel, concrete, and timber members. 

2.8.2 Selecting Levels of Concern 

For large incidents involving LNG spills on water, selection of levels of concern is complicated 

by the fact that fire incidents may be brief, perhaps on the order of 5 minutes for a pool fire involving 

a rapid spill of the contents of a single membrane tank from a carrier. However, the general guidance 

on damage to structures (e.g., intensity levels presented in Section 2.8.1) is based on long-term 

exposure (perhaps 30 minutes or more). 
34 

 

Table 2.9 Various Thermal Radiation Limits for Structures from Lees (1996) 

Thermal Radiation Intensity Limit 

(BTU/hr/ft2) (kW/m2) Limit Description 

Design Guidance from Kletz (1980) 

12,000 38 Intensity on storage tanks 

4,000 12.5 Intensity on wood or plastics 

1,600 5 Intensity on people performing emergency operations 

Design and Assessment Guidance from British Standard 5908 (BS 5908, 1990) 

12,000 37.5 Intensity at which damage is caused to process equipment 

7,900 25 Intensity at which nonpiloted ignition of wood occurs 

4,000 12.5 Intensity at which piloted ignition of wood occurs 

Design and Assessment Guidance from Mecklenburgh (1985) 

4,400 14 Intensity which normal buildings should be designed to 

withstand 

3,200 – 3,800 10-12 Intensity at which vegetation ignites 

Assessment Guidance from Dinenno (1982) 

9,500 30 Spontaneous ignition of wood 

4,800 15 Piloted ignition of wood 

6,300 20 Ignition of No. 2 fuel oil in 40 seconds 

3,200 10 Ignition of No. 2 fuel oil in 120 seconds 

5,700 - 6,300 18-20 Cable insulation degrades 

3,800 12 Plastic melts 

12,000 37.5 Equipment damage 

2,900 9 Equipment damage – conservative value used in flare 

system design 

The data presented in Section 2.8.1 show some variations in intensity levels and associated 

effects. However, the data are relatively consistent, and for long-term exposure, some reasonable 

choices for levels of concern and associated types of damage are as follows: 

• 12,000 BTU/hr/ft2 (38 kW/m2) – Damage to process equipment and storage tanks 

• 7,900 BTU/hr/ft2 (25 kW/m2) – Ignition of wood without direct flame exposure 

• 3,800 BTU/hr/ft2 (12 kW/m2) – Piloted ignition of wood, melting of plastic, ignition of 

vegetation 

Of course, the appropriate levels of concern for a specific site depend on the materials present. 

For example, in an area with a wooden structure, a value of 7,900 BTU/hr/ft2

 (25 kW/m2) might be 

appropriate if there are no combustibles in the area that are more vulnerable. However, if vegetation 

or other more vulnerable combustibles are present, a level of 3,800 BTU/hr/ft2

 (12 kW/m2) may be 

more appropriate, particularly if ignition of these materials could initiate a larger fire. 
35 

 

These levels could be used as general guidelines for fire events of any duration, but they will tend 

to be very conservative for short-duration fires. For example, some steel structures could be exposed 

to much higher intensities for short durations without their temperatures rising to damaging levels. 

However, as noted previously, heat balance calculations are generally required to evaluate the 

intensities/durations that a specific structure can withstand. To minimize the need for heat balance 

calculations, the above levels of concern could be employed, and detailed analysis could be used on a 

case-by-case basis when structures have the potential for exposure to fires of shorter duration. 
36 

 

3 CONSEQUENCE ASSESSMENT EXAMPLES 

This section presents some example consequence analysis results using the methods 

recommended in Section 2. The scenarios presented here are chosen to illustrate the various analysis 

methods and are not intended to model any specific facility. For these examples, site-specific 

parameters (e.g., atmospheric conditions, surface roughness) were chosen either to match values used 

in other studies (for comparison purposes) or values required by 49 CFR 193 for LNG facility siting. 

For actual analysis of a site, these parameters must be selected based on site conditions and the goals 

of the analysis. 

Detailed computations and computer output data for the samples in this section are provided in 

Appendices B (release rates), C (pool fires), and D (vapor dispersion). The computations presented 

in those appendices were performed with the assistance of the Mathcad computer software (©1986-

2000 MathSoft, Inc.). 

3.1 RELEASE RATES 

As mentioned in Section 2, the orifice model is recommended, although only for providing a 

rough guide to the rate of release for a given hole size. As an example, Figure 3.1 presents estimated 

release rates for a few hole sizes. 

For these calculations, the height of the LNG above the hole was taken as 43 ft (13 m). This is an 

approximate value for the height of the LNG above the water level in a membrane-type carrier. The 

discharge coefficient was assumed to be 1.0. The calculations of these values are presented in 

Appendix B. 

3.2 POOL FIRES 

This section presents the results of example pool fire calculations for an LNG spill on water. The 

scenarios examined are fires following spills from 3.3-ft (1-m) and 16-ft (5-m) holes in an LNG 

carrier just above the waterline. In both cases, the holes are assumed to be uniform in diameter and 

to penetrate through the outer and inner hulls and the cargo tank. The 3.3-ft (1-m) diameter hole 

represents a relatively long duration event, and the 16-ft (5-m) diameter hole represents a more rapid, 

short duration release. These example calculations are intended only as demonstrations of the 

modeling methods. The results should not be taken as a consequence assessment for any specific 

facility. Evaluation of a specific facility requires input parameter values based on site-specific 

conditions, and analysis of different or additional scenarios may be appropriate. 
Figure 3.1 Estimated LNG Release Rates Based on Orifice Model 

For these examples, it is assumed that the amount of LNG above the hole is 4.4 × 105

 ft3 (12,500

m3

), and the orifice model is used to estimate outflow, with flow rate dropping as the liquid level 

above the hole drops. It is assumed that the spill is ignited immediately upon release. The scenarios 

and input parameter are summarized as follows: 

Scenarios

Hole diameters: 3.3 ft (1 m) and 16 ft (5 m)

Initial liquid height above hole: 43 ft (13 m)

37 

 
38 

 

Total spill quantity: 4.4 × 105 ft3 (12,500 m3) 

Air temperature: 80 °F (27 °C) 

Relative humidity: 70% 

Wind speed: 20 mph (8.9 m/s) 

Flame surface emitted flux: 84,000 BTU/hr/ft2

 (265 kW/m2) 

Burning rate: 0.058 lb/s/ft2

 (0.282 kg/s/m2) 

Table 3.1 summarizes the results of the pool fire calculations for these scenarios. The detailed 

calculations are provided in Appendix C. Note that the results for the 3.3-ft (1-m) hole show that the 

total fire duration matches the total spill duration. This occurs because the pool spreads until the 

burning rate matches the spill rate into the pool. For the 16-ft (5-m) hole, the spill is much more 

rapid, lasting only 1.3 minutes, and the pool does not have enough time to spread to the point where 

the burn rate matches the spill rate. However, the large mass in the pool does result in a much larger 

pool, and therefore the spilled material burns in a much shorter time. 

Table 3.1 Summary of Results for Example Pool Fire Calculations 

Hole diameter 3.3 ft (1 m) 16 ft (5 m) 

Initial spill rate 11,700 lb/s (5,300 kg/s) 290,000 lb/s (130,000 kg/s) 

Total spill duration 33 min 1.3 min 

Maximum pool radius 240 ft (74 m) 440 ft (130 m) 

Total fire duration 33 min 6.9 min 

Flame length (height) 910 ft (280 m) 1,400 ft (430 m) 

Flame tilt at maximum radius 35 deg 31 deg 

Downwind distance to 

12,000 BTU/hr/ft2 (38 kW/m2) 

1,200 ft (370 m) 2,000 ft (600 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

1,500 ft (450 m) 2,400 ft (720 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

2,000 ft (600 m) 3,200 ft (980 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

2,800 ft (860 m) 4,600 ft (1,400 m) 

3.3 FLAMMABLE VAPOR DISPERSION 

This section presents the results of example dispersion calculations for an LNG spill on water. 

The scenarios examined are spills that are not immediately ignited and disperse downwind. Spills 

are from 3.3-ft (1-m) and 16-ft (5-m) holes in an LNG carrier just above the waterline. For these 

examples, it is assumed that the amount of LNG above the hole is 4.4 × 105

 ft3 (12,500 m3) and the 

orifice model is used to estimate outflow, with flow rate dropping as the liquid level above the hole 

drops. 
39 

 

These are the same scenarios presented in Section 3.2 for pool fires, except in this case it is 

assumed that ignition does not occur immediately and ambient conditions are different. Also, as 

stated above for pool fires, these example calculations are intended only as demonstrations of the 

modeling methods. The results should not be taken as a consequence assessment for any specific 

facility. Evaluation of a specific facility requires input parameter values based on site-specific 

conditions, and analysis of different or additional scenarios may be appropriate. 

The scenarios and input parameters are summarized as follows: 

Scenarios

Hole diameters: 3.3 ft (1 m) and 16 ft (5 m) 

Total spill quantity: 4.4 × 105

 ft3 (12,500 m3) 

Air temperature: 71 °F (22 °C) 

Water temperature: 70 °F (21 °C) 

Relative humidity: 50% 

Wind speed: 4.5 mph (2.0 m/s) 

Surface roughness: 0.03 ft (0.01 m) 

Pasquill-Turner stability class: F 

Averaging time: 0 sec (i.e., peak concentrations are used) 

Table 3.2 summarizes the results of the source term and dispersion calculations for these 

scenarios. The detailed source term calculations are provided in Appendix D, and dispersion 

calculations were performed using DEGADIS. 

Table 3.2 Summary of Results for Example Dispersion Calculations 

Hole diameter 3.3 ft (1 m) 16 ft (5 m) 

Initial spill rate 11,700 lb/s (5,300 kg/s) 290,000 lb/s (130,000 kg/s) 

Total spill duration 33 min 1.3 min 

Heat transfer to LNG pool 11,700 BTU/hr/ft2 (37 kW/m2) 

Maximum pool radius 430 ft (130 m) 550 ft (170 m) 

Total evaporation duration 34 min 18 min 

Downwind distance to LFL 11,000 ft (3,300 m) 13,000 ft (3,900 m) 

Time at which LFL reaches maximum distance 44 min 35 min 

Time at which entire cloud 

 drops below LFL 

48 min 38 min 

Downwind distance to 1⁄2 LFL 16,000 ft (4,800 m) 18,000 ft (5,500 m) 

Time at which 1⁄2 LFL reaches 

maximum distance 

53 min 45 min 

Time at which entire cloud 

drops below 1⁄2 LFL 

58 min 48 min 
40 

 

4 LNG RELEASE PREVENTION AND MITIGATION MEASURES 

4.1 INTRODUCTION 

This section outlines current measures that the international community, individual governments, 

shipping companies, and associated facilities employ to prevent and mitigate potential LNG releases 

from LNG carriers. Specifically, the section covers the following areas: 

• Equipment overview 

o Brief descriptions of the basic types of LNG carriers (Section 4.2) 

• Prevention measures 

o Regulatory and industry approaches for LNG carrier design (Section 4.3.1) 

o Design and operational features pertinent to release prevention (Sections 4.3.2 and 

4.3.3) 

o Security measures pertinent to release prevention (Section 4.3.4) 

• Mitigation measures 

o Measures that can help mitigate releases once they occur, whether they are accidental 

releases or releases caused by security incidents (Section 4.4) 

4.2 EQUIPMENT OVERVIEW 

LNG is transported at sea in three basic types of carriers – membrane ships, independent 

prismatic tank ships, and independent spherical tank ships. Each of these types of carriers must meet 

international and U.S. Coast Guard (USCG) requirements before they receive the various approvals 

that are required to allow them to operate in U.S. waters. Because there are different cargo tank 

designs that fall into the three classes of LNG carriers referenced above, this report does not describe 

the specifics of LNG carrier tank design. Since this report was not developed to reflect a specific 

cargo tank design, that information is not within the scope of this study. 

However, all current LNG cargo tanks either provide two barriers designed for cargo 

containment at low temperature (i.e., membrane type designs) or a more robust single barrier with 

provisions to handle potential leakage (including international requirements specifically for hull 

structures in locations where they could be exposed to potential leaks and USCG rules that require 

crack-arresting steel in specific locations). Also, the LNG cargo tanks for all three types of LNG 

tanks are separated from the external environment by a double hull. This means external damage to 

an LNG carrier hull would not result in an LNG release unless the damage was extensive enough to 

puncture the inner hull and then the cargo tank. 

4.3 INCIDENT PREVENTION MEASURES 

4.3.1 Regulatory and Industry Approaches for LNG Carrier Design 

LNG carriers and most other oceangoing vessels are governed by rules and regulations 

established by many different entities. Those entities are generally categorized as the International 

Maritime Organization, the flag state, the port state, and classification societies. Each of the entities 
41 

 

has specific areas of responsibility, and in some cases the responsibilities overlap. The result of these 

various rules and regulations is that most aspects of the design, construction, and operation of LNG 

tankers are addressed by the various entities. LNG carriers are unique in that LNG is a cryogenic 

liquid; accordingly, specialized materials, construction methods, and operating procedures are needed 

to safely handle LNG. The general rules and regulations that govern ships at sea do not address the 

particular concerns of LNG; therefore, specific rules and regulations have been developed by the 

various entities in order to ensure the safety of LNG tankers and their ports of call. 

4.3.1.1 Definitions 

The International Maritime Organization (IMO): A branch of the United Nations that operates 

under the auspices of the United Nations Convention on the Law of the Sea. The IMO Convention 

entered into force in 1958, and the new Organization met for the first time the following year. The 

purposes of the Organization, as summarized by Article 1(a) of the Convention, are "to provide 

machinery for cooperation among Governments in the field of governmental regulation and practices 

relating to technical matters of all kinds affecting shipping engaged in international trade; to 

encourage and facilitate the general adoption of the highest practicable standards in matters 

concerning maritime safety, efficiency of navigation and prevention and control of marine pollution 

from ships." The Organization is also empowered to deal with administrative and legal matters 

related to these purposes. IMO serves to write international requirements for safety and pollution 

prevention for ships in international service. These rules are then administered in various ways by 

Flag and Port States. 

Flag States: The country of registry for the vessel, such as the United States, Panama, Bahamas, 

etc. The flag state is the country of the national flag the vessel flies. In IMO Conventions, the flag 

state is sometimes referred to as the "Administration." The flag state establishes regulations for the 

construction and operation of vessels registered under its flag. Many of the flag state requirements 

are based on the vessel complying with regulations of the IMO. 

Port State: The authority that has jurisdiction over the port area and waters under national 

control. In the United States, this is typically the USCG, supplemented by state and local 

authorities. As defined by IMO, Port State Control (PSC) is the inspection of foreign ships in 

national ports to verify that the condition of the ship and its equipment comply with the requirements 

of international regulations and that the ship is manned and operated in compliance with these rules. 

The USCG defines Port State Control as “the process by which a nation exercises its domestic and or 

international authority over foreign vessels when those vessels are in waters subject to its 

jurisdiction.” 

Classification Society: A classification society is an industry organization, other than a flag 

state, that issues certificates of class and/or International Convention Certificates. The certificates of 

class are based on rules published by the classification society that govern the design and 

construction of ships and offshore installations. A classification society has specific procedures 

regarding the level of design review and survey that are required to allow a vessel to be “classified.” 

Classification indicates that the vessel met applicable class rules, international requirements, and 

specific national requirements. Also, some flag states delegate certain additional review and 

inspection responsibilities to classification societies. 

Membrane Tanks: Non-self-supporting tanks that consist of a thin layer (membrane) supported 

through insulation by the adjacent hull structures. The membrane is designed in such a way that 
42 

 

thermal and other expansion or contraction is compensated for without undue stressing of the 

membrane. 

Independent Tanks: Self-supporting tanks; they do not form part of the ship’s hull and are not 

essential to the hull strength. There are three categories of independent tanks: 

Type A — Designed primarily using recognized standards for classical ship structural analysis 

procedures 

Type B — Designed using model tests, refined analytical tools, and analysis methods to 

determine stress levels, fatigue life, and crack propagation characteristics 

Type C — Tanks meeting pressure vessel criteria and designed for specific vapor pressure 

criteria (not typically applicable to LNG carriers) 

4.3.1.2 Design specifications for LNG tankers 

The IMO, through its signatory nations, provides the basis for the framework of international 

regulations for shipping. The IMO publishes documents that outline the international requirements 

for vessel design and construction, operation, safety security, and crewing. There are several IMO 

publications and or conventions regarding design, construction, and operation of LNG tankers that 

are allowed to operate in U.S. waters. These conventions are: 

• Convention on the International Regulations for Preventing Collisions at Sea, 1972/1981 

• Safety of Life at Sea (SOLAS), 1974/1981 

• Code for the Construction and Equipment of Ships Carrying Liquefied Gases in Bulk (Gas 

Carrier Code), 1983 

• International Code for Ships Carrying Liquefied Gases in Bulk (IGC Code), 1993 

• 1994/1996 Amendments to the IGC (replaced the Gas Carrier Code) 

• International Convention on Standards of Training, Certification and Watchkeeping (STCW) 

for Seafarers, 1978 

• International Management Code for the Safe Operation of Ships and for Pollution Prevention 

(ISM Code) – adopted by IMO Resolution A.741 (18) in 1994 

The most important of these international requirements for new LNG carrier design is the IGC 

Code. This provides the basic requirements that flag states and classification societies then 

supplement with more detailed requirements. 

Flag states are responsible for ensuring that vessels intending to fly their flags are designed and 

constructed in accordance with the applicable international standards and any unique national 

requirements. Flag states generally discharge their responsibilities by (1) combining design review 

and inspection by their personnel and (2) requiring the vessels to be constructed in accordance with 

the design and survey rules of a classification society of which they approve. Flag states then have 

an ongoing responsibility for safe vessel operation and maintenance. This requires the vessel to be 

periodically inspected and certified by the flag state and likely requires periodic survey by the 

classification society also. Such inspections also confirm that the ship is adequately staffed (based 

on international and national requirements) and that the vessel crew has received specific training 

and certification (where required). 
43 

 

Port states impose laws and regulations specific to their individual ports and areas of jurisdiction. 

In the case of vessels wanting to enter the United States, the USCG exercises port state control. As 

part of port state control measures, a port state can impound or delay a vessel for a number of 

reasons, including, but not limited to, safety issues, pollution issues, operational 

restrictions/concerns, vessel condition, crew member issues such as immigration status, and security 

issues. 

Industry organizations (like the vessel charterers and other participants in the LNG value chain) 

may also perform periodic audits of vessel safety and compliance with applicable requirements. 

These inspections serve not only to help ensure safety but also contribute to reliable operation in the 

LNG trade. A contributor to the safety of the industry is the Society of International Gas Tanker and 

Terminal Operators (SIGTTO), which publishes recommended practices in areas important to the 

safety of LNG vessels and terminals. 

The remainder of this section discusses some of the design features of LNG carriers that help 

prevent significant releases if incidents should occur. 

4.3.2 Design Features Pertinent to Release Prevention 

Double Hull Construction: All LNG carriers (both currently in service and planned) provide 

double hull construction, providing a significant measure of protection against LNG release due to 

external damage, such as could occur in groundings, collisions, or allisions (i.e., striking a fixed 

object). The IGC Code requires cargo tanks to be protected from damage due to collision or 

grounding by being located at least a specified minimum distance inboard from the ship’s shell 

plating. Also, gas carriers are required by the IGC Code to survive the normal effects of flooding 

that could be caused by external damage to the hull. Typically a grounding, collision, or allision 

occurs due to a (1) failure of the propulsion and/or steering system, (2) failure of navigational 

equipment, or (3) human error. The various authorities extensively regulate propulsion and steering 

systems. Maintenance and inspection of such equipment are the first step in preventing failures that 

would lead to the grounding of an LNG carrier. To ensure that such actions occur, flag states require 

periodic inspections of the vessels in their registry. Also, port states may perform their own 

inspections or require evidence of routine inspections to allow the vessel entry into ports under their 

jurisdiction. Entry into port would not typically be allowed if there were deficiencies in this 

equipment. Also, LNG carriers are generally guided into ports and restricted waterways by local 

pilots with extensive experience in the port area. This reduces the likelihood of navigational failures 

or human errors by normal tanker crews. Docking and departing of LNG tankers also generally 

involve the use of harbor tugs to assist in maneuvering the LNG carrier. Also, in the United States, 

LNG carrier arrivals and departures are subject to operational plans developed and enforced by the 

cognizant USCG Marine Safety Office (MSO). These plans generally mandate safety zones and 

restrict traffic in the area of any LNG tanker. 

Segregation of Cargo Areas and Piping: The IGC Code requires that hold spaces (i.e., the 

enclosed space in the ship structure where the cargo containment system is situated) be separated 

from machinery and boiler spaces, accommodations, service spaces, control stations, chain lockers, 

drinking and domestic water tanks, and stores. There are also specific requirements for segregation 

of cargo piping systems from other piping systems in order to prevent the transfer of cargo or cargo 

vapor into or through other piping systems. Appropriate locations for accommodations, service and 

machinery spaces, and control stations are also defined in the IGC Code. 
44 

 

Accessibility for Inspection Access: In order to facilitate inspection of various structures and 

equipment in gas carriers, the IGC Code requires arrangements that allow visual inspection of at least 

one side of the inner hull structure and insulation located in hold spaces. Adequate access to cargo 

tanks to allow internal inspection is also required. The ability to access these areas allows periodic 

surveys and inspections to be performed in order to identify any damage or deterioration that has 

occurred. This in turn helps any identified problems to be corrected before they become safety issues 

for the vessel. 

Leak Detection for Gas Carrier Hold Spaces: The IGC Code requires hold spaces and 

insulation areas where cargo leakage could occur to be equipped with gas detection and low 

temperature alarms. 

Cargo Containment Design Requirements: The IGC Code mandates development of a cargo 

containment design that addresses a design vapor pressure for the containment system that considers 

the cargo vapor pressure and the maximum allowable relief valve setting. Also, the selection of 

materials for cargo tank construction must be based on a defined minimum design temperature, and it 

must be shown that the cargo tank temperature cannot be lowered below the design temperature. 

Design loads for the cargo containment must be based on specific combinations of potential loads 

(e.g., internal pressure, external pressure, dynamic loads due to ship motion, thermal loads, sloshing 

loads, ship deflection loads, weight distribution). 

Structural Analysis Requirements: The IGC Code mandates that structural analyses be 

performed for each type of cargo containment design to address the specific loads of concern for that 

type of cargo containment design. The IGC Code also defines the allowable stresses for independent 

cargo tanks. 

Secondary Barrier Provisions: LNG carriers built to the IGC Code require a secondary cargo 

barrier. The design intent for the secondary barrier is to contain an envisaged leakage of liquid cargo 

(generally for up to 15 days) to prevent lowering of the temperature of the ship structure to an unsafe 

level. Membrane and Type A independent tanks require a full secondary barrier. Independent Type 

B tanks (i.e., self-supporting tanks that are designed using model tests, refined analytical tools, and 

analysis methods) only require a partial secondary barrier. In all cases, the vessel inner hull that is 

adjacent to cargo tanks must be protected against contact from liquid cargo (e.g., either a full second 

barrier or appropriate spray shield and routing of potential leakage). 

Low Temperature Provisions: In addition to the requirements to handle envisaged leakage, the 

IGC Code requires a combination of proper material selection, adequate insulation, and use of 

heating systems (which must include standby heating in the event of a failure of the normal heating 

system) to ensure that hull structures are not exposed to temperatures below their design temperature. 

The design must consider that the secondary barrier is at the cargo temperature with the lowest 

design ambient air and seawater temperatures. 

Construction and Testing Requirements for Cargo Containers: The IGC Code includes 

construction and testing requirements that address welding, workmanship, quality assurance, 

hydropneumatic or hydrostatic pressure testing, leak testing, and nondestructive testing of welds. 

Pressure Vessel and Piping System Requirements: The IGC Code contains provisions for 

general design requirements, piping thickness, design pressure rating, allowable stresses and stress 

analyses, type testing, fabrication and joining, welding, postweld heat treatment, nondestructive 
45 

 

testing, hydrostatic testing, and leak testing of piping systems and pressure vessels. These 

requirements are designed to ensure that the piping is appropriately designed and fabricated. Other 

IGC requirements specify the types of isolation valves required on connections to cargo tanks. 

Emergency Shutdown Valves and Shutdown Systems: The IGC Code requires remotely 

controlled emergency shutdown valves on the ship for stopping liquid and vapor transfers between 

the ship and shore. Cargo pumps and compressors must be arranged to shut down automatically if 

the emergency shutdown system closes the emergency shutdown valves. In addition, there must be a 

remotely operated shutdown valve at each cargo hose connection used in transfers. The emergency 

shutdown control system must be capable of being actuated by a single control in either of two 

locations on the ship. It must also be actuated by fusible links that will respond in the event of a fire 

in locations, including tank domes and loading stations. 

Cargo Pressure/Temperature Control: The IGC Code requires pressure control for the cargo 

system in order to keep the cargo tank pressure under the tank design pressure or the maximum 

allowable relief valve setting (MARVS). The pressure control system can (1) employ mechanical 

refrigeration or (2) use the boil-off vapor as fuel for shipboard use or for a waste heat system. Per the 

IGC Code, a flag administration can allow other means to control cargo tank pressure, such as 

venting LNG vapor to the atmosphere. Of course, operators would only consider this as a secondary 

means of pressure/temperature control. However, the USCG requires LNG cargo systems to be used 

in U.S. waters to be capable of maintaining the cargo for at least 21 days without venting. This is 

intended to eliminate the need to vent flammable vapors while in U.S. ports. 

Pressure Relief Systems: The IGC Code requires pressure relief systems for cargo tanks that 

are independent of the pressure control systems described previously. For LNG cargo tanks of the 

size typically used on LNG carriers, the IGC Code requires at least two pressure relief valves of 

appropriate capacity. The valves have to be provided with vent piping that minimizes the potential 

for accumulation of LNG vapor on deck or ingress of vapor into areas where it might create a 

dangerous condition. The IGC Code mandates that the pressure relief valves have a combined 

relieving capacity of the greater of (1) the maximum capacity of the cargo tank inerting system or (2) 

the vapor generation rate due to fire exposure of the tank. 

Vacuum Protection Systems: The IGC Code requires that any LNG cargo tank not designed for 

the maximum pressure differential that could be developed by (1) pumping out a tank without vapor 

return, (2) operating a refrigeration system, or (3) sending boil-off vapor to machinery spaces, be 

provided with a vacuum protection system. Acceptable vacuum protection systems under the IGC 

Code include use of two independent pressure switches that alarm and actuate to stop any suction of 

liquid or vapor from the cargo tank and shut down refrigeration equipment. Another vacuum 

protection option is to provide vacuum relief valves that are sized for the maximum discharge rate 

from the cargo tank. Since operation of a vacuum relief valve would introduce air into the tank, in 

practice, ship pressure control systems are designed to prevent negative pressure situations in the 

tank well before vacuum relief valves operate. 

Vessel Fire Protection Systems: Per the IGC Code, LNG carriers require a firewater main 

system and hydrants/hose stations with the ability to supply at least two jets of water to any part of 

the deck in the cargo area and parts of the cargo containment and tank covers above the deck. The 

firewater supply system also has to supply water for a water spray system that provides water for 

cooling, fire prevention, and crew protection in specific areas. LNG carriers also require a fixed dry 

chemical powder-type extinguishing system to allow the crew to fight fires on the deck in the cargo 
46 

 

area and cargo handling areas. In addition, the cargo compressor room, which is located on deck, is 

fitted with a CO2 smothering system. 

Cargo Tank Instrumentation: In order to allow an LNG carrier crew to monitor cargo tank 

conditions, the IGC Code requires a variety of instrumentation. These include: 

• Gas detection equipment capable of alerting the crew to leakage of LNG from the 

primary barrier (i.e., instruments to detect LNG vapor between the primary and 

secondary barriers) 

• At least one liquid level gauging device for each tank 

• A high liquid level alarm (independent of other level devices) that gives audible and 

visual warning and shuts off flow into the cargo tank when it is actuated. This serves to 

help prevent the tank from becoming liquid-full 

• Cargo tank pressure gauges, marked with maximum and minimum pressures, providing 

both high and low pressure (if the tank requires vacuum protection) 

• Two temperature indicating devices (one at the bottom and one at the top of the cargo 

tank) marked to show the minimum cargo tank design temperature 

• Temperature indicating devices and alarms to monitor for leaks of LNG into the 

insulation between primary and secondary LNG cargo tank barriers or onto the hull 

structure adjacent to the cargo tank 

This instrumentation, along with the pressure relief and vacuum pressure protection systems 

described earlier, provides a major role in preventing LNG cargo tank upsets or human errors from 

damaging the tank or resulting in a large scale release. 

Vessel Gas Detection Systems: In addition to gas detection discussed previously, the IGC Code 

requires a permanently installed gas detection system with audible and visual alarms for areas, 

including cargo compressor rooms, motor rooms, cargo control rooms, enclosed spaces in the cargo 

area, specific ventilation hoods and gas ducts in the propulsion machinery space for the fuel gas 

system, and air locks. These detectors are required to activate alarms when 30% of the LFL is 

reached. Ships also are required to have at least two sets of portable gas detectors. 

Gas Detection and Safety Shutdown Systems for Terminal Operations: LNG unloading 

systems and docks are equipped with LNG vapor detection, fire detection, and associated safety 

shutdown systems that shut down pumping operations and close valves to isolate the transfer lines. 

The shutdown systems can be actuated by the ship’s crew or LNG terminal personnel. In most cases, 

these systems also respond automatically to any detection of LNG in the atmosphere by shutting 

down pumping operations and closing valves to isolate the LNG transfer lines. The systems can also 

be actuated by the ship’s crew or LNG terminal personnel. This serves to limit the amount of LNG 

that would be released in a leak that occurred during the unloading process. 

Emergency Release Couplings: Current designs for LNG terminals have emergency release 

couplings that are fitted between the ship’s cargo manifold and the receiving station. These couplings 

are designed to release if vessel movement exceeds predetermined limits. If the couplings release, the 
47 

 

resulting LNG loss to the atmosphere is designed to be very small (e.g., approximately 1 gal). As 

LNG is a cryogenic liquid that has the capacity to cause severe embrittlement of steel structures, the 

areas of an LNG carrier deck likely to be contacted by a leak of LNG are protected with drip trays 

constructed of materials suitable for withstanding LNG exposure. 

4.3.3 Operational Measures Pertinent to Accidental Release Prevention 

In addition to design features, the international and national maritime authorities have recognized 

the need for operational measures that prevent or at least reduce the likelihood of LNG incidents. 

These measures include: 

• Training 

• Procedures 

• Inspections

Training – Minimum training requirements for crews of LNG carriers are specified in the IMO 

STCW Convention referenced previously. This level of training is generally exceeded by the 

training specified by the flag state and the shipping company. LNG carrier crews are considered 

some of the most professional and competent in the international shipping business, recognizing the 

important role they play in safe operations. 

Procedures – International ships over 500 gross tons subject to SOLAS (including LNG carriers) 

are required by IMO to meet the ISM Code. This safety management approach requires development 

of safety and environmental policies, along with the assignment of responsibilities, development of 

procedures, performance of training, and periodic audits and reviews to evaluate compliance with 

those requirements. The ISM Code explicitly addresses the need to prepare for responding to 

emergency situations like fire and LNG releases. 

Inspections – Inspections for vessels occur from a variety of sources. Flag states and 

classification societies both have requirements for periodic inspections and surveys. Port state 

control activities often include safety inspections of vessels desiring to enter a port. These thirdparty inspections are in addition to inspections and audits required under vessel and shipping 

company ISM programs or by charterers concerned with ensuring safe, reliable voyages on which the 

success of their investment depends. 

4.3.4 Security Measures that Help Prevent Release Incidents Due to Deliberate Attacks 

Some of the design features and operational measures described in the previous sections play a 

role in preventing LNG releases even when terrorist acts are considered (e.g., double hull structures). 

However, this section describes the international and U.S. requirements for security measures 

applicable to LNG carriers. 

After the events of 9-11, the IMO added Chapter XI-2 to the SOLAS Convention to provide the 

International Ship and Port Facility Security (ISPS) Code. Applicable to LNG vessels, the ISPS 

Code requires actions by flag states, port states, shipping companies, and port facilities to reduce the 

likelihood of terrorism acts against maritime assets. Effective July 1, 2004, the ISPS Code requires: 

• Security levels 

• Ship security plans 
48 

 

• Ship security alarm systems 

• Automatic identification systems 

• Port security plans 

• Declarations of security 

• Facility security plans 

These security measures are discussed below. 

Security levels – Flag states and port states are required to provide security-related information to 

their vessels and port facilities in the form of a designation of security levels. The three security 

levels represent the assessment of the likelihood of a terrorist event being attempted against maritime 

assets. Each level is intended to trigger specific security measures that are defined in ship, port, and 

port facility security plans. 

Ship security plans – Flag states are required to review and approve ship security plans 

developed by the operator of each SOLAS vessel in international service that is larger than 500 gross 

tons (which includes all LNG carriers). These plans must address security measures for: 

• Access to the ship by ship personnel, passengers, visitors, etc. 

• Restricted areas on the ship 

• Handling of cargo 

• Delivery of ship’s stores 

• Handling unaccompanied baggage 

• Monitoring the security of the ship 

These measures are intended to protect the ship and to prevent the ship from being used as a 

weapon against other maritime targets (e.g., other vessels, bridges, terminals). The ship security plan 

implementation must be evaluated by an onboard verification by the flag state or a security 

organization recognized by the flag state before an International Ship Security Certificate (ISSC) can 

be issued for that vessel. 

Ship security alarm systems – Each vessel subject to the ISPS Code must carry on board 

equipment to provide a ship security alarm (SSA) that the ship’s crew can activate if terrorists board 

the vessel. The alarm, which is does not sound aboard the vessel, is intended to alert responsible 

authorities (e.g., the USCG or other designated authority) that a security incident is underway. 

Automatic identification systems (AISs) – AIS equipment has been in the process of being 

introduced over the last few years. It consists of a transponder on each vessel that provides a signal 

identifying vessel name, location, cargo, heading, etc. Originally intended to help reduce vessel 

collisions and make port state control activities easier, it is recognized as another step that contributes 

to port security. As such, application of AISs was accelerated for some vessels as part of the changes 

to SOLAS made when the ISPS Code was introduced. For U.S. ports where vessel traffic systems or 

other AIS base stations are implemented, the AIS transponders allow the USCG to have real-time 

radar systems that provide a higher level of information than just the radar images. 

Port security plans – Each port served by international vessels subject to the ISPS Code is 

required to have a port facility security officer and to develop a port facility security plan. The plan 

must specify the security measures expected at local port facilities and the measures expected by 

international vessels that interface (e.g., load, unload, refuel, transfer passengers) with those 

facilities. In the United States, the applicable USCG Captain of the Port (COTP) serves as the port 
49 

 

facility security officer and, in conjunction with a port security committee, is responsible for 

developing a port security plan. Among other things, the port security plan identifies the individual 

facilities within the port that will require their own facility security plan. Also, port security plans 

define what operational controls (e.g., security zones, escorts, sea marshals) apply to various kinds of 

vessels to enter the port. The USCG also uses the requirement for an Advanced Notice of Arrival 

(ANOA) from vessels entering U.S. ports from foreign departure locations as a tool in port state 

control. The ANOA information allows the USCG to select vessels based on their route, cargo, 

security, and safety history, and other pertinent information in order to plan security inspections, 

escorting, and boarding of sea marshals or other USCG personnel as appropriate. 

Declarations of security – Nationwide, the USCG has mandated that all visits by vessels like 

LNG carriers require a declaration of security (DoS) that must be prepared by the ship and the 

facility it is to call upon. A DoS defines the security responsibilities for both the facility and the ship 

for the period of time the vessel is at that facility. It has to be coordinated between the ship security 

officer and the facility security officer and signed prior to any interface activities (e.g., cargo 

transfers or personnel movement) between the vessel and the facility. 

Facility security plans – Under the USCG maritime security regulations (33 CFR, Subchapter 

H), LNG facilities that receive LNG carriers will have to develop a security plan. Like the ship plans 

that have to meet the ISPS Code, the USCG regulations define areas the facility security plans have 

to address, including: 

• Security administration and organization of the facility 

• Personnel training 

• Drills and exercises 

• Records and documentation 

• Response to change in security level 

• Procedures for interfacing with vessels 

• DoS 

• Communications 

• Security systems and equipment maintenance 

• Security measures for access control, restricted areas, handling cargo, delivery of vessel 

stores and bunkers, and monitoring 

• Security incident procedures 

• Audits and security plan amendments 

Like ship security plans, facility security plans have to be approved by the responsible 

government, in this case, the port state. For U.S. facilities, that requires submission of the facility 

security plans to the USCG for review and approval. 

4.4 MITIGATION MEASURES TO HELP RESPOND TO RELEASES 

For LNG releases, the industry has traditionally focused on release prevention; however, there 

are some measures that are pertinent to help mitigate releases once they occur. This section discusses 

those measures, which apply to both accidental releases and releases caused by security incidents 

(e.g., terrorist acts). 

The primary focus in responding to LNG releases is to protect public and employee safety and 

secondarily to minimize the amount of damage to vessel and facility property. Some of the features 

previously discussed as design elements for preventing significant releases play a role once a release 
50 

 

occurs. These include gas detection, safety shutdown systems, and fire protection systems. Beyond 

those elements there are other mitigation features such as: 

• Safety and security zones 

• Ship and facility emergency response plans 

• Coordination with USCG and local emergency responders 

• Evacuation plans and procedures 

Safety and security zones – Safety zones around vessels play a role in reducing the likelihood of 

collisions or the need for an LNG vessel to try to avoid other port traffic. Security zones help vessel 

and USCG resources recognize if a vessel near an LNG carrier is not willing to abide by established 

rules and allow the vessel to keep watch for unusual activity. They also allow the USCG to intervene 

with a vessel if resources are available to do so. However, either of these types of zones help in that 

if there is an event, they limit the proximity of recreational and commercial traffic to the LNG vessel 

and berthing area. 

Ship and facility emergency response plans – Both LNG carriers and LNG terminals are 

required to develop emergency response plans, complete with personnel training and exercises. 

These plans serve to help ensure that prompt and effective action results if a release scenario occurs. 

Emergency response actions include release isolation (if possible), minimization of ignition sources 

(e.g., by shutting down equipment or restricting activities), and evacuation of ship crew or members 

of the public if deemed appropriate. 

Coordination with USCG and local emergency responders – Because the ship crew may not be 

able to deal with all potential contingencies or possible release events, emergency response actions 

are typically coordinated with USCG and other local emergency response personnel located near 

LNG terminals where transfers will occur. Coordination, which might include joint drills, exercises, 

or other types of training, helps improve understanding of the types of events that could occur, the 

expected consequences, and how the resources of the various entities can best be combined to 

address those events. 

Evacuation plans and procedures – If necessary, an emergency response plan will call for 

evacuation of locations where the emergency personnel are concerned that threats to personnel may 

develop. Particularly for an unignited release, emergency responders may involve law enforcement 

or other public resources to help notify personnel in the downwind direction to take action and move 

away from the area of the hazard. 
51 

 

5 CONCLUSIONS 

5.1 REVIEW OF WORK SCOPE 

• An evaluation of published models to calculate the rate of cargo release and the spread 

of unconfined LNG spills on water 

No models were identified that account for the multi-hull structure of an LNG carrier and the 

physics of a release of cryogenic LNG. In the absence of models specifically for LNG 

carriers, the orifice model is recommended. However, when using this model it is vital to 

keep in mind that the model does not match the reality of carrier construction very well, and 

the results should be interpreted as a rough guide to the rate of release for a given hole size. 

Because it does not account for the multi-hull construction of carriers, the orifice model will 

tend to overestimate the LNG outflow for many scenarios. See Section 2.1 for details of the 

release rate modeling review and recommendations. 

Several models for spreading were found in the literature, most of which are based on the 

assumption that spreading is dominated by the balance of the gravitational spreading force 

and inertial resistance to spreading. One methodology, developed by Webber and his 

colleagues (as described in TNO 1997), also accounts for resistance to spreading as a result 

of frictional forces. Our analysis of these methodologies indicates that frictional effects are 

important for scenarios involving large releases in short periods of time. Therefore, 

Webber’s methodology is recommended. 

It is important to note that this methodology for estimating pool spread does not account for 

some aspects of spills on open water that could be important such as wind, water currents, 

and waves. However, no existing model for an LNG spill appropriately accounts for these 

effects. It should be recognized that the recommended model is based on the assumption of 

smooth quiescent water. 

See Section 2.2 for details of the pool spread modeling review and recommendations. 

• A determination of the effect of cargo tank hole diameter on cargo release and pool 

spread (including establishment of an upper limit on pool spread) 

The example release rates presented in Section 3.1 provide some information about the effect 

of cargo tank hole diameter. However, the results of the sample calculations must be 

considered no more than a rough guide to potential release rates. 

For a given hole size (or release rate), the upper limit on pool spread depends on the duration 

of release. For long duration releases, the pool will spread to the point where evaporation 

rate (or burning rate in case of a fire) matches the release rate into the pool. This is based on 

a simple mass and heat balance for the pool, and effectively defines an upper limit for longterm releases. For short duration releases, the dynamics of the release are more important in 

identifying the upper limit. The pool spread is slowed by inertial and frictional effects but is 

only stopped once the pool is thin enough for surface tension effects to become important. 

During the spread, evaporation reduces the volume of the pool, and therefore also slows the 

spread. As a result of these dynamics, it is difficult to establish an upper limit for short 

duration releases without modeling the full spread process. 
52 

 

This raises the question of how to decide if a given release should be treated as long duration 

or short duration for modeling purposes. It is recommended to avoid the necessity of 

selecting either of these by modeling the pool spread in a way that handles a transient release 

into the pool. The integral solution to the spreading problem implemented for the example 

problems in this report accomplishes this and allows modeling scenarios using (1) an initial 

pool volume and/or (2) a constant or transient flow rate into the pool. 

• An evaluation of these models to calculate the rate of vapor generation for unconfined 

LNG spills on water 

LNG spilled from a carrier onto water will form a pool at its boiling point (at atmospheric 

pressure) of approximately -259 oF (-162 oC). At this temperature the LNG will rapidly 

absorb heat from the water, and the heat absorbed will go directly to vaporize the LNG (at a 

rate determined by the heat of vaporization of the LNG). Therefore, modeling of 

vaporization of the LNG essentially reduces to the modeling of heat flux into the pool from 

the water. Of course the dynamics of pool spread (as discussed above and in Section 2.2) are 

also important because total evaporation rate (mass per unit time) depends on the pool area. 

Two methods are identified for estimating heat flux to a boiling LNG pool on water: (1) 

select a value estimated from spill tests or (2) calculate a value based on heat transfer theory. 

Based on review of the literature, it is recommended to estimate heat flux using heat transfer 

theory. Key reasons for this choice are that (1) the experimental data show significant 

spread, (2) the experimental data for heat flux were typically calculated from other data 

(rather than measured directly), and (3) a viable theoretical model from heat transfer theory is 

available to help fill the gap in experimental data. 

As noted above for pool spread, it is important to note that this methodology for estimating 

vapor generation does not account for some aspects of spills on open water that could be 

important such as wind, water currents, and waves. 

See Section 2.3 for details of the vapor generation modeling review and recommendations. 

• An evaluation of the models to calculate thermal radiation distances from fires 

involving unconfined LNG spills on water 

Available pool fire models for hazard assessment purposes are predominantly based on the 

solid flame model. The fire is represented as gray-body emitter, and the geometry is 

approximated by an upright cylinder that is tilted as a result of wind. The point source model 

is also sometimes used. In this model energy generated by combustion is assumed to 

emanate from a single point at the center of the pool fire. The calculated energy is multiplied 

by a fraction that accounts for the fact that only part of the energy will be emitted as thermal 

radiation. 

In some circumstances the point source model can be appropriate, but its chief disadvantage 

is that it does not produce valid estimates for any people or structures located close to the 

fire. Therefore, the solid flame model is recommended. 

Two key parameters required to apply the solid flame model are the burning rate and surface 

emitted flux. Based on a review of experimental data the recommended values for these 

parameters are a burning rate of 0.058 lb/s/ft2

 (0.282 kg/s/m2) and emissive power of 84,000 

BTU/hr/ft2

 (265 kW/m2). There is significant spread in the experimental data for these 
53 

 

parameters; however, comparisons only to data for the larger scale spill tests tend to indicate 

that these are relatively conservative choices. 

See Section 2.4 for details of the pool fire modeling review and recommendations 

• An evaluation of the models to calculate the distance traveled by flammable vapors 

from unconfined LNG spills on water 

This work did not involve a detailed review of dispersion models. For large releases of LNG, 

the cold LNG vapors are more dense than the surrounding air and tend to remain at surface. 

These dense-gas effects are important and must be accounted for in the model used. Several 

such models are available in the form of freely available, as well as commercial, software. 

For the example dispersion calculations presented in this report, the DEGADIS software was 

used. This model accounts for dense-gas effects and was originally developed for simulation 

of cryogenic flammable gas dispersion, particularly for LNG. It has been validated against a 

wide range of available laboratory and field test data. 

See Section 2.5 for details of the dispersion modeling review and recommendations.. 

• An evaluation of the theoretical basis for current models used to calculate each of the 

above, and a review of the applicability of existing computer modeling to field test data 

The theoretical basis is an important factor that was considered as part of the overall review 

of each modeling area and is addressed throughout Section 2. 

The time available for this work did not allow a detailed review of existing computer 

programs that implement the recommended methods; however, use of some existing software 

is discussed in Section 2. In general, validation of computer models (as well as their 

theoretical bases) has been very limited with regard to large LNG spills on water. There are 

simply very few data to validate models against, and there are no test data for spills of the 

large size considered by this report. 

• A review of the effects of thermal radiation levels on population and structures 

The extent to which people are injured by exposure to thermal radiation depends on both the 

incident heat flux and the exposure time. A variety of data are available for estimating 

effects on people, including data from experiments with humans and animals and review of 

historical data. Section 2.7.1 summarizes some of the injury criteria developed based on these 

data. 

Like effects of thermal radiation on people, effects on structures also depend on incident heat 

flux and the exposure time. With structures, effects also depend strongly on the materials of 

construction (e.g., wood, steel, concrete). Section 2.8 presents some criteria from various 

sources for structural damage. Most of the criteria presented are for long-term exposure, 

which is very generally considered to be on the order of 30 minutes or more. More accurate 

assessment of structural damage requires performing heat balance calculations to determine 

the resulting surface temperature for comparison with acceptable temperatures for the 

particular structure. 
54 

 

• A recommendation, including the selection rationale, of specific methods and a model 

which FERC staff may use to calculate the flammable vapor and thermal radiation 

hazards associated with marine transportation of LNG 

The recommendations are made throughout Section 2, and example assessments are 

presented in Section 3. The detailed calculations for the examples are provided in 

Appendices A through D. 

• A comprehensive review of design, operational, safety, and security features and 

possible mitigating measures that an LNG facility and/or vessel operator employs or 

could provide to minimize impacts from: accidents; attacks; leaks; spills; fire; rapid 

phase transitions; or the spread of negatively buoyant, flammable vapor clouds 

Section 4 of this report is dedicated to this item. 

5.2 OBSERVATIONS 

The methods recommended in this analysis can only provide rough estimates of the magnitude of 

effects for incidents involving large LNG releases on water. This is typically the case with 

consequence assessments, and it is important to keep in mind that the recommended methods cannot 

definitively describe the effects because of variability in actual incident circumstances, as well as 

uncertainty inherent in the methods used. 

The example consequence calculations presented in Section 3 are intended only to illustrate use 

of the various analysis methods recommended in Section 2. The scenarios were chosen simply to 

exercise the various methods and are not necessarily the correct scenarios to model for specific 

facility siting or any other hazard assessment purpose. Furthermore, some input parameters (e.g., 

atmospheric conditions, surface roughness) were chosen for the examples only and may be 

inappropriate for a particular site and/or particular hazard assessment. When using the presented 

analysis methods for an actual analysis, it is important to choose scenarios and input parameters 

based on the analysis requirements and site conditions. 
55 

 

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57 

 

Klimenko (1981) 

V.V. Klimenko, “Film Boiling on a Horizontal Plate – New Correlation,” International Journal of 

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R.P. Koopman, B.R. Bowman, and D.L. Ermak, Data and Calculations of Dispersion on 5 m3 LNG 

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F.P. Lees, Loss Prevention in the Process Industries, Second Edition, ISBN 0-7506-1547-8, 

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58 

 

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59 

 

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Appendix A 

Film Boiling Heat Transfer Calculations 
A-2 

 

Tb = 200.988R

Vapor density

(at Tb, using ideal gas law)

ρ v

pa

Ru

M

Tb ⋅

:= ρ v 1.751 kg

m3

= ρ v 0.109

lb

ft3 =

Vapor viscosity (at Tb

) μv 4.362 10− 6 := ⋅ Pa sec

Kinematic viscosity

of vapor(at Tb

)

νv

μv

ρ v

:= νv 2.491 10− 6 × m2

sec =

Density of Liquid (at Tb

) ρ l 422.5 kg

m3 := ⋅

Heat of vaporization hfg 509331.9 joule

kg := ⋅

Surface Tension (at Tb

) σ 0.0133

newton

m :=

Thermal conductivity of vapor λv 0.01269

watt

m K⋅ :=

Heat capacity of vapor Cp 2075.56 joule

kg⋅K :=

Vapor thermal diffusivity αv

λv

ρ v⋅Cp

:= αv 3.491 10− 6 × m2

sec =

LNG Boiling Heat Flux for Spill on Water

Ambient Conditions

Ambient temperature Ta := ( ) 100 + 459.67 ⋅R Ta = 559.67R

Ambient pressure pa := 101325⋅Pa pa = 14.7psi

Temperature of water Tw := ( ) 20 + 273.15 ⋅K Tw = 293.15K

Constants and Conversion Factors

Unit conversions kJ ≡ 1000⋅joule lbmole ≡ 1 kgmole kg

lb ≡ ⋅lbmole kW ≡ 1000⋅watt

Universal gas constant Ru 1545

ft⋅lbf

lbmole⋅R ≡ ⋅

Material Properties

Molecular weight M 16.043 kg

kgmole := ⋅

Normal boiling point Tb := 111.66⋅K
Qfilm 36.6

kW

m2

Heat flux Qfilm := hfilm⋅( ) Tw − Tb =

hfilm 201.8

watt

m2

⋅K

hfilm λv = Nu

lc

Heat transfer coefficient := ⋅

Nu 0.19 ( ) Ar⋅Prv Nu = 179.4

1

3 ⋅ ⋅F1 Ar 108 ⋅( ) ≤ 0.0086⋅ Ar Prv

1

3 ⋅ Ar 108

:= + ⋅( ) >

F2 1 L⋅( ) vR ≤ 2

LvR

2

⎛

⎜

⎝

⎞

⎠

1

2

F1 1 L⋅( ) vR ≤ 1.4 := + ⋅( ) LvR > 2

LvR

1.4

⎛

⎜

⎝

⎞

⎠

1

3

:= + ⋅( ) LvR > 1.4

Nusselt number

LvR LvR = 1.352

hfg

Cp⋅( ) Tw − Tb

Dimensionless heat of vaporization :=

Prv Prv = 0.713

νv

αv

Prandtl number :=

Ar 5.452 108 Ar Ga = × ( ) ρ l − ρ v

ρ v

Archimedes number := ⋅

Ga 2.269 106 Ga = × g lc

3 ⋅

νv

2

Galileo number :=

l lc 2π c = 0.011m σ

g⋅( ) ρ l − ρ v

⎡

⎢

⎣

⎤

⎥

⎦

.5

Length scale for vapor bubble formation :=

Use the model/correlation of Klimenko, as given by TNO (1997).

Film Boiling Heat Flux

A-3 

 
 

Appendix B 

Example Release Rate Calculations 
B-2 

 
B-3 

 
 

Appendix C 

Example Pool Fire Calculations
C-2 

 
C-3 

 
C-4 

 
C-5 

 
C-6 

 
C-7 

 
C-8 

 
C-9 

 
C-10 

 
C-11 

 
C-12 

 
C-13 

 
C-14 

 
C-15 

 
C-16 

 
C-17 

 
C-18 

 
C-19 

 
C-20 

 
C-21 

 
C-22 

 
C-23 

 
C-24 

 
C-25 

 
C-26 

 
C-27 

 
C-28 

 
 

Appendix D 

Example Source Term Calculations for 

Flammable Vapor Dispersion 
D-2 

 
D-3 

 
D-4 

 
D-5 

 
D-6 

 
D-7 

 
D-8 

 
D-9 

 
D-10 

 
D-11 

 
D-12 

 
D-13 

 
D-14 

 
D-15 

 
D-16 

 
D-17 

 
D-18 

 
D-19 

 
D-20 

 
D-21 

 
D-22 

 
D-23 

 
D-24