UNITED STATES OF AMERICA 

FEDERAL ENERGY REGULATORY COMMISSION 

 Docket No. AD04-6-000 

NOTICE OF AVAILABILITY OF STAFF’S RESPONSES 

TO COMMENTS ON THE CONSEQUENCE ASSESSMENT METHODS 

 FOR INCIDENTS INVOLVING RELEASES FROM LIQUEFIED 

 NATURAL GAS CARRIERS 

(June 18, 2004) 

 The Federal Energy Regulatory Commission’s staff researched and reviewed 

methodologies for modeling liquefied natural gas spills on water. The final report 

entitled Consequence Assessment Methods for Incidents Involving Releases from 

Liquefied Natural Gas Carriers was made available to the public on May 14, 2004, with 

comments due by May 28. The staff’s responses to the comments are now available in 

PDF format from the Commission’s website (http://www.ferc.gov/industries/gas/indusact.asp). 

 Magalie R. Salas 

 Secretary 

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STAFF’S RESPONSES TO COMMENTS ON CONSEQUENCE ASSESSMENT 

METHODS FOR INCIDENTS INVOLVING RELEASES FROM LIQUEFIED NATURAL 

GAS CARRIERS

Comments on the ABSG Consulting Inc. report were filed by 49 parties, including 22 

individuals, nine industry groups, three local governments, three environmental 

organizations, and 12 from the scientific community. The issues raised by the 

commenters included: the need for additional large scale testing and research; the 

development of new models; the need to consider liquefied natural gas (LNG) vessel 

design characteristics; and the need to proceed with a rulemaking process. While some 

commenters contend that the report should not be used for any realistic evaluation of 

existing or proposed LNG projects until further developed, other commenters assert that 

the report supports their claim that LNG projects should be denied in their area. The 

comments and staff’s responses follow: 

Comment 1 

A number of commenters requested that the current regulations for LNG facility siting 

should be updated given the hazard impact areas stated in the report. Alternatively, the 

distances described in the report should be used as the minimum distance between new 

facilities and population, and that exclusion zones should be established around LNG 

tankers. In addition, the numbers suggested by the study indicate that locating these 

facilities near population is unacceptable and that they should be located offshore. 

Response: 

The intent of this study was to recommend modeling methods to be used by FERC staff 

in the site specific National Environmental Policy Act (NEPA) review of proposed LNG 

import facilities. The results shown do not provide a generic site assessment for all LNG 

import facilities. Credible worst-case scenarios, based on the most recent information 

available, will be used in site-specific analyses of each LNG import facility proposed 

before the Commission and will be issued in FERC NEPA documents. As stated in the 

final environmental impact statement (FEIS) for the Freeport LNG Project (Docket No. 

CP03-75-000), it should not be assumed that the hazard distances identified are the 

assured outcome of an LNG vessel accident or attack, given the conservatisms in the 

models and the level of damage required to yield such large scale releases. Further, these 

estimated “worst case” scenarios should not be misconstrued as defining an exclusionary 

zone. Rather the “worst case” scenarios provide guidance in developing the operating 

restrictions for LNG vessel movements within each shipping channel, as well as in 

establishing potential impact areas for emergency response and evacuation planning. 

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Comment 2 

Information used in the determination of LNG facility site acceptability, including onshore exclusion zones, and the Federal safety zones for LNG vessels, should be available 

to the public before any siting decisions are made. The public has a right to know the 

risks involved. 

Response: 

The results of every site-specific analysis performed with the recommended methods will 

be included in the Commission’s NEPA documentation. Further opportunity for review 

and comment will be provided during the scoping period in each case and during the 

comment period for each of the Commission’s environmental impact statements. 

Comment 3

The study is too limited in the review of tanker failure and should examine more than just 

the loss of one cargo hold. Multiple hull breaches and the loss of multiple tanks should 

be considered. The use of 1- and 5-meter diameter holes should be justified and proven 

to be credible worst-case damage scenarios from an intentional attack. Real-life events, 

such as the USS Cole and the Limburg tanker attack, as well as the Skikda explosion, 

should be incorporated into the models. Clearly these events indicate that a terrorist 

could successfully attack an LNG ship and that LNG vapor clouds do explode. 

Response: 

The study was a review of consequence modeling methodology, rather than an evaluation 

of damage scenarios or a determination of the robustness of LNG carriers. The cargo 

tank hole sizes presented in the study were selected to represent values used in other 

studies and it was not an attempt to quantify the probability or extent of credible damage 

scenarios. In the final EIS for the Freeport LNG Project (Docket No. CP03-75-000, 

issued in May 2004), we identified a security analysis that analyzed a range of potential 

attack scenarios and estimated consequences. Subsequently, a detailed evaluation of the 

consequences of a terrorist attack on a modern membrane LNG tanker was prepared by 

Lloyds Register North America for the Weaver’s Cove LNG Project (Docket No. CP04-

36-000). These provide a basis for estimating the potential magnitude of a hazard from a 

successful terrorist attack, and for developing LNG vessel and waterfront security plans. 

There are other studies currently underway, by other government agencies and the natural 

gas industry, which are evaluating credible ship damage and effects. As those findings 

become available, we will continue to refine the site-specific analyses FERC staff is 

performing for each proposed project. 

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Comment 4 

Modeling of the cryogenic effects of the spill on the tanker and a consideration of the 

effects of LNG vapors in tanker compartments should also be included. 

Response: 

The purpose of the study was to recommend methodology for modeling the consequences 

of an LNG spill on water. There are other studies underway that are evaluating credible 

ship damage scenarios. 

Comment 5 

The study is solely focused on the consequences of hypothetical cargo tank hole sizes 

which are not credible given the multiple layers of ship safety and security measures. 

Response: 

In addition to reviewing consequence methodology, section 4 of the study reviews LNG 

release prevention and mitigation measures, including design features of LNG vessels 

and security measures to deter potential attack scenarios. 

Comment 6 

The report does not assess the potential risks associated with LNG shipping, which must 

include the probability of potential incidents. 

Response: 

The focus of the study is the methodology that FERC staff will use to calculate the sitespecific hazards from potential marine spills. The study also reviews the various layers 

of protection built into the LNG vessel and associated terminal operations that serve to 

avoid or minimize the likelihood of an incident, but was not intended to provide a 

measure of risk to the public. The operational safety of LNG ships is under the 

jurisdiction of the U.S. Coast Guard, an agency within the Department of Homeland 

Security (DHS). In a letter dated April 15, 2004 to Representative Edward Markey, the 

DHS stated that it believes that LNG tanker operations can continue to be safely 

conducted in U.S. ports. 

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Comment 7 

The FERC should release information concerning the Skikda event and should consider 

the possibility of such an occurrence in siting new LNG facilities. 

Response: 

The intent of this study was to establish guidance on modeling methods to be used by 

FERC staff in analyzing marine spills of LNG. Information concerning the incident at 

the Skikda onshore liquefaction facility was summarized in the final EIS for the Freeport 

LNG Project. Further information, as it becomes available, will be included in 

subsequent NEPA documents for other projects undergoing Commission review. 

Comment 8 

Given the complexity of this issue, several parties stated that an extension of the comment 

period is needed to allow more thorough review of the study. 

Response: 

The intent of this study was to review existing consequence modeling methods, which in 

most cases have already been subjected to standard academic peer review. In addition, 

the results of every site-specific analysis performed with the recommended methods will 

be included in the Commission’s NEPA documentation. Further opportunity for review 

and comment will be provided during the comment period for each of the Commission’ 

environmental impact statements. The format of the study breaks the methodology into 

individual modeling components so that factors can be adjusted as the scientific 

community provides new experimental data or improved modeling approaches. 

Comment 9 

The report is limited in the number of models examined and in the depth of analysis 

(models and literature). Other research work should be used. There may be other 

studies that were not available at the time of the study and this information should be 

included in the review. 

Response: 

This report is a starting point in developing consequence assessment methodology that 

can be applied to projects before the Commission using supportable data and models, 

while recognizing that the results are conservative. We also recognize that improved 

modeling and additional research will allow for more refined consequence methodology. 

In effect, the development of our consequence models will be ongoing. With every 

advancement in these models, we will incorporate those improvements in our analyses. 

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Comment 10 

Several commenters suggested that FERC, as well as the natural gas industry and other 

Federal agencies, fund or develop more experimental data to better quantify the 

“complex and disparate physical properties” used in modeling these spills. 

Response: 

Page iv of the Executive Summary recognizes the need for additional research and largescale spill tests to provide better data and refine the models. 

Comment 11 

Commenters recognized that several modeling recommendations were made in an 

attempt to provide a measure of conservatism. Some commenters felt that the study was 

overly conservative, while others did not feel that every conservative assumption had 

been made. 

Response: 

Although the goal of the study was to select methods that provide the most accurate 

estimates possible, the lack of large-scale historical incidents and the need to extrapolate 

small-scale field test data lead to the use of conservative assumptions. As the scientific 

community provides new experimental data or improved modeling approaches, the 

methodology, which is broken into individual modeling components, can be adjusted to 

provide a more refined analysis. 

Comment 12 

Over-reliance on security protection to justify poor siting location decisions, such as 

siting in populated areas, can be highly deceptive. The U.S. Coast Guard cannot stop 

terrorist attacks on all ships. 

Response: 

The focus of the study is the methodology that FERC will use to consistently calculate 

the site-specific hazards from potential spills. Credible worst-case scenarios, based on 

the most recent information available, will be used in site-specific analyses of each 

facility proposed before the Commission and will be issued in FERC NEPA documents. 

The operational safety of LNG ships is under the jurisdiction of the U.S. Coast Guard. 

While the risks associated with the transportation of any hazardous cargo can never be 

entirely eliminated, they can be managed. 

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Comment 13 

The heat flux levels of concern listed in the report are not appropriate. The 1,600-, 

3,000-, and 10,000 Btu/hr-ft2

 levels listed in NFPA 59A do not provide protection for the 

public, and are established by a committee which does not include fire service members. 

A lower level heat flux should be used as an acceptable level. How does the 10,000 

Btu/ft2-hr level protect property when this level can ignite structures? This level is also 

too high for fire-fighting and rescue attempts. 

Response: 

The intent of this study was to establish guidance on modeling methods to be used by 

FERC staff in the NEPA review of proposed LNG import facilities. In selecting incident 

flux levels to evaluate damage to structures, including residences, and injuries to the 

exposed public from LNG spills on water, the flux levels specified by the U.S. 

Department of Transportation’s (DOT) regulations under 49 CFR Part 193, which 

incorporates NFPA 59A, have been used in the recently issued Freeport LNG Project 

final EIS. These specific incident flux levels have more than a 25-year regulatory 

history, including several rulemaking processes with multiple opportunities for public 

input. Since the safety of both onshore LNG facilities and marine transportation are 

assessed in a single environmental impact statement, the use of the NFPA 59A levels 

allows a consistent treatment of hazards from both components of an LNG import 

facility. 

Comment 14 

FERC should apply this study to existing facilities such as Everett LNG facility in Boston 

and should periodically review response and evacuation plans for existing facilities. 

Response: 

As the Lead Federal Agency responsible for the preparation of NEPA documentation 

required for the siting and construction of onshore LNG facilities, the FERC conducts 

environmental, safety, and security reviews of proposed LNG plants and related pipeline 

facilities. This analysis includes LNG tanker operations, which are not under the 

jurisdiction of the FERC, as part of the overall safety review shared with the DOT and 

the U.S. Coast Guard in accordance with the 2004 Interagency Agreement. The intent of 

this study was to establish guidance on modeling methods to be used by FERC staff in 

the NEPA review of proposed LNG import facilities. Once a proposed facility has been 

constructed and placed in operation the marine hazards are managed by the U.S. Coast 

Guard through port-specific LNG operating plans. The U.S. Coast Guard exercises 

regulatory authority over LNG vessels which affect the safety and security of port areas 

and navigable waterways and is responsible for matters related to navigation safety, 

vessel engineering and safety standards. 

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Comment 15 

The Boston Fire Department questioned the continued use of acceptable risk levels that 

were determined for LNG transit in Boston Harbor since these were based solely on the 

probability of ship accidents. The threat of a deliberate action such as sabotage 

precludes such risk estimates. They also stated that structures used as emergency 

shelters may burn, increasing the number of potential fatalities. 

Response: 

These comments concern issues which are outside the scope of this study. 

Comment 16 

The Boston Fire Department also commented that agency jurisdiction over the Boston 

facility is unclear and asked for clarification of the regulatory roles between the FERC, 

the U.S. Coast Guard, the Department of Energy, and the State of Massachusetts. 

Response: 

The FERC, the U.S. Coast Guard (within the DHS), and the Office of Pipeline Safety 

under the Research and Special Programs Administration (RSPA) within the DOT are 

responsible for exercising regulatory authority over certain aspects of liquefied natural 

gas facilities, and related land and marine safety and security issues. The FERC is 

responsible for authorizing the siting and construction of onshore LNG facilities under 

Section 3 of the Natural Gas Act (15 U.S.C. § 717 et seq.). FERC also authorizes the 

construction and operation of interstate natural gas pipelines that may be associated with 

the LNG facilities under section 7 of the Natural Gas Act. RSPA has authority to 

promulgate and enforce safety regulations and standards for the transportation and 

storage of LNG in or affecting interstate or foreign commerce under the pipeline safety 

laws (49 U.S.C. Chapter 601). RSPA’s authority extends to the siting, design, 

installation, construction, initial inspection, initial testing, operation, maintenance of 

LNG facilities. The U.S. Coast Guard exercises regulatory authority over LNG facilities 

which affect the safety and security of port areas and navigable waterways under E.O. 

10173, the Magnuson Act (50 U.S.C. § 191), the Ports and Waterways Safety Act of 

1972, as amended (33 U.S.C. § 1221, et seq.) and the Maritime Transportation Security 

Act of 2002 (46 U.S.C. Section 701). The U.S. Coast Guard is responsible for matters 

related to navigation safety, vessel engineering and safety standards, and all matters 

pertaining to the safety of facilities or equipment located in or adjacent to navigable 

waters up to the last valve immediately before the receiving tanks. The U.S. Coast Guard 

also has authority for LNG facility security plan review, approval and compliance 

verification as provided in Title 33 CFR Part 105, and siting as it pertains to the 

management of vessel traffic in and around the LNG facility. 

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Comment 17 

The Boston Fire Department questioned why this study was not funded until 2 years after 

9/11. They also requested information as to whether periodic reviews of existing 

facilities were conducted using the latest technical information. 

Response: 

This study was intended to review modeling methodology so that FERC staff could 

prepare its site-specific review of each proposed terminal. 

Comment 18 

The Boston Fire Department faulted FERC for having little interaction with the public, 

for not committing to applying the report’s conclusions to existing facilities, and for not 

pursuing a “vigorous” program of system improvement. They recommend that FERC 

prepare a report that analyzes the risks and hazards of a worst case release. The report 

should include consideration of damage to surrounding infrastructure and should detail 

the on-site capabilities required at a LNG terminal as well as a review of existing LNG 

facilities to determine if they meet the latest safety standards. They maintain that this 

report should cover both land and water based vehicles that transport LNG. Also, the 

operator of each facility should provide funds for the local community to conduct their 

own independent review of the report. 

Response: 

These comments concern issues which are outside the scope of this study. 

Comment 19 

The report states that there are no models to account for many of the phenomena and 

scenarios to be examined. However, higher-level modeling tools, such as Computational 

Fluid Dynamics (CFD) modeling could be used to perform these analyses, but is not 

addressed in the study. 

Response: 

The report states that fire modeling can be preformed using CFD methods, but that they 

are not normally used for typical pool fire hazard assessment 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. 

Similarly, the federal regulations for siting LNG facilities in 49 CFR Part 193 allow the 

use of both the DEGADIS, a dense gas dispersion model, and FEM3A, a CFD model. 

While recognizing that FEM3A may more realistically depict terrain and building effects, 

DEGADIS provides acceptable and conservative results that are widely used rather than 

the more complex FEM3A. For additional discussion, see the responses to Analytical & 

Computational Engineering, Inc. in Comment 27. 

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Comment 20 

The thermal radiation effects on people in Section 2.7 showing the 99% fatality 

conditions need to show both distance and time. This information indicates that 

emergency service responses, as well as evacuation plans, would be ineffective given the 

large impact areas associated with a potential release. 

Response:

The thermal dosage data in Table 2.4 represents the cumulative heat received over time 

and corresponding type of injury. For example, exposure to 12 kW/m2

 for 100 seconds 

results in the 99% fatality level. Table 2.4 is not used to define the intensity of a specific 

fire or the distances to specific thermal radiation levels. 

Comment 21 

The study did not include a comparative analysis between LNG and other hazardous 

cargos also carried in large tankers through shipping channels and deepwater ports. 

Response: 

The purpose of the study was to provide the methods for calculating site-specific hazards 

from an LNG spill on water. The modeling of other hazardous cargoes and a comparison 

with LNG was beyond the scope of the study. 

Comment 22 

According to a U.S. Coast Guard study, vapor clouds from mixtures of LNG containing 

85% methane are susceptible to explosion. However, the report states that LNG vapor 

will not explode. 

Response: 

The report notes that a larger volume fraction of heavier hydrocarbons in LNG reduces 

the minimum ignition energy required for detonation and increases the likelihood of 

generating damaging overpressures. 

Comment 23 

A wider range of comparison of dispersion modeling should have been performed and 

should reference more recent work on the validation of these models. Recent work was 

not considered. 

Response: 

The report surveyed a representative sample of flammable vapor dispersion models 

focusing on those that are widely used in available software. For example, the federal 

regulations for siting LNG facilities in 49 CFR Part 193 specify the use of DEGADIS as 

one of the two approved models. 

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Comment 24 

The models should recognize that rapid phase transitions (RPTs) are explosions and 

should be modeled with a great deal of conservatism. 

Response: 

Section 2.6 of the report was devoted to a review of the current knowledge of RPTs, 

including the explosive yield from RPTs that was observed in the field. 

Recommendations for modeling the effects of RPTs on vapor dispersion were performed 

using several scenarios. 

Comment 25 

Given the errors and shortcomings of the ABSG Consulting study identified by 

commenters, FERC should issue a corrected version of the study before using it as 

guidance in siting decisions. 

Response: 

As discussed in greater detail in the response which follows, ABSG Consulting has 

revised various components of its consequence assessment methodologies. These 

changes include: the orifice discharge coefficient for calculating spill rates has been 

changed from 1.0 to 0.65; the approximate pool shape of an uncontained LNG spill on 

water is now represented as a semicircle instead of a circle; the estimated effects of 

friction between the LNG pool and the water surface on pool spread have been reduced; 

the relationship between decreasing spill rate and pool size has been refined; the rate of 

heat influx from water has been set at 85 kW/m2

; and the solid flame model has been 

modified to represent a two-zone pool fire. As new research, data and improved 

modeling techniques appear, the methodology will be revised as appropriate. 

Comment 26 

We received several comments on factors in the model regarding: A) selection of the 

discharge coefficient; B) pool shape; C) film boiling effects on viscous spreading; D) 

effect of decreasing spill rate on pool spread; E) film boiling heat flux; F) pool fire flame 

length; G) pool fire smoke production; H) pool fire burning rate and surface emissive 

power. 

Response: 

Individual discussion of items A-H are provided below. A summary of the modifications 

to the recommended consequence modeling methods, as well as revised example 

calculations, are also included.

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A. Selection of Discharge Coefficient 

In the orifice flow model used to estimate outflow from a hole in an LNG tank, a 

discharge coefficient can be used to account for the fact that friction retards the flow. 

The report employed a discharge coefficient of 1.0; however, several comments 

recommend use of a lower value. 

The literature review did not identify any methodology or empirical results to support a 

lower value, particularly with regard to holes in ships, so a value of 1.0 was used. This is 

also consistent with guidance of Lees (1996), which states “For cases where the discharge 

coefficient is unknown or uncertain, use a value of 1.0 to maximize the computed flows 

to achieve a conservative result.” This same guidance is repeated in Crowl (1990) and 

AIChE (2000). 

In spite of this, an argument can be made for using a value lower than 1.0, which 

represents the ideal (frictionless) case. 

The comments of the International LNG Alliance and the International Gas Union 

suggest a value of 0.65, and the comments of the Center for LNG suggest that “a 

reasonable, conservative estimate of discharge coefficient ranges from 0.6 to 0.8.” 

However, references were not provided for these values. 

As additional guidance, Lees (1996) states that: 

(1) For sharp-edged orifices and for Reynolds numbers greater than 30,000, the 

discharge coefficient approaches the value 0.61. For these conditions the exit 

velocity is independent of the hole size. 

(2) For a well-rounded nozzle the discharge coefficient approaches unity. 

(3) For short sections of pipe attached to a vessel with a length-diameter ratio not less 

than 3, the discharge coefficient is approximately 0.81. 

AIChE (2000) also presents an estimate of the discharge coefficient based on the 2-K 

method. In this method the coefficient is calculated and accounts for contributions to 

friction from the entrance to the hole and exit from the hole. The calculated discharge 

coefficient value is 0.63. 

In the case of a rough, irregular hole that would be expected in a spill from an LNG 

carrier, the friction would be expected to be larger than in the case of a sharp-edged 

orifice. On this basis, and the guidance of Lees (1996) and AICHE (2000), the suggested 

value of 0.65 can be defended as a reasonable estimate, and this value is recommended. 

B. Pool Shape 

Dr. Fay’s comments recommend modeling the LNG pool on water as a “semicircle with 

base diameter located at the vessel side and center of radial symmetry at the vessel 

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orifice.” He presents this approach in a recent paper (Fay 2003). Drs. Havens and Spicer 

also comment that “ship tank spills would more probably than not occur to one side of the 

vessel, resulting in the spill resembling an expanding half-circle (bounded by the side of 

the ship) more than a circular spill.” 

In general, the pool shape employed in modeling should be selected based on the 

specifics of the scenario being evaluated. In the case of the types of spills from LNG 

carriers that are being addressed, a semicircle is a more realistic representation. 

Dr. Fay’s methods of estimating spread (Fay 2003) is applicable to this situation. Also, 

as he points out in his comments, the methodology employed for the examples in the 

report can be modified easily to model a semicircle. In fact, this is simply a matter of 

introducing a factor of 1/2 in the appropriate equations to account for the fact that the 

area of the pool is half the area of a circle with the same radius. This approach is 

recommended as it still allows use of the integral solution, which has the advantage of 

accommodating a transient spill source and avoiding the need to characterize spills as 

either instantaneous or continuous. 

C. Film Boiling Effects On Viscous Spreading 

The comments of Drs. Havens and Spicer and comments of Dr. Fay question one specific 

aspect of the application of the Webber pool spread methodology (which is outlined in 

TNO 1997). In particular, the commenters indicated that the friction between the water 

and spreading LNG pool was overestimated, causing a tendency to retard spreading of the 

pool. While less specific, the comments of the Center for Liquefied Natural Gas also 

express the concern that the Webber model may overestimate viscous effects. 

The estimate of friction is based on equations 3.70 and 3.71 in TNO (1997), along with 

the recommended turbulent friction coefficient of 0.0015. Commenters pointed out that 

this approach does not account for the fact that film boiling is taking place. In film 

boiling, a thin layer (film) of LNG vapor forms between the water surface and liquid 

LNG pool. This vapor has a viscosity that is two orders of magnitude less than the 

viscosity of the liquid LNG and three orders of magnitude less than the viscosity of the 

water. Webber does take the friction coefficient to be a constant value, which is probably 

justified for many circumstances. However, he also points out that the value “may 

depend on the nature of the surface under the pool and whether or not film boiling is 

taking place” (Webber 1987, p. 5). 

The thin layer of vapor would be expected to significantly reduce the friction between the 

two liquid surfaces, and further review of Webber’s work indicates that the methodology 

employed overestimated the friction. 

For purposes of analysis, many authors (Fay 1969, Hoult 1972, Webber 1986, and others) 

discuss pool spread in terms of various flow regimes, such as the ‘gravity-inertia’ (or 

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‘inviscid’) regime and ‘gravity-viscous’ regime. In the gravity-inertia regime (i.e., early 

in a spill), gravitational spreading forces are balanced by fluid inertia. In the gravityviscous regime, gravitational spreading forces are balanced by viscous forces. These are 

useful concepts, but it is important to keep in mind that they are simplifications. 

In the gravity-inertia regime, frictional forces are still present but are generally small 

compared to inertia forces. Likewise, later in a spill, inertial forces are still present but 

are generally small compared to viscous forces. The point of transition between these 

regimes is typically taken as the point where inertial and viscous resistance forces are 

equal. So, at that instant, ignoring one of these two forces means ignoring half of the 

resistance force. 

Because LNG vaporizes rapidly when spilled on water, most spills will evaporate 

completely before friction forces become important. However, very large and rapid spills 

may exist long enough for frictional forces to become important. 

The literature review did not identify a recommended methodology for estimating friction 

in the case of film boiling. However, a straightforward approach is to assume complete 

slip between the LNG pool and water surface and estimate the friction from the shear 

stress in the vapor film. The assumption of complete slip means that water surface 

remains stationary and that the velocity profile in the leading edge of the pool is uniform. 

This approximation is justified by the fact that the viscosity of the vapor is much lower 

than that of either of the liquids. 

As Fay points out in his comments, the thickness of the vapor film can be approximated 

as 

Q

v Tw Tp ( − ) = λ

δ (1) 

where 

δ = Thickness of the vapor film 

λv = Thermal conductivity of the vapor 

Tw = Temperature of the water 

Tp = Temperature of the LNG pool 

Q = Film boiling heat flux from the water to the LNG pool 

The shear stress in the film is then 

δ

μ τuv = (2) 

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where 

τ = Shear stress 

μ v = Viscosity of the vapor 

u = Speed of the LNG pool front 

This shear stress represents the friction force on the pool, per unit area of pool surface. 

The corresponding acceleration on the pool provides a resistance term analogous to 

equation 3.70 in TNO (1997). 

h

C

l

FF

ρ

τ = (3) 

where 

CFF = Frictional resistance term with film boiling 

ρ l = Density of the liquid LNG 

h = Pool depth 

Before this method can be applied, an appropriate heat flux must be selected for use in 

estimating the vapor film thickness. Choosing a flux of 85 kW/m2 (see Item E), assumes 

that localized breakdown of the film accounts for a portion of the heat transfer, so basing 

the film thickness estimate on this heat flux value is inappropriate. However, where 

breakdown of the film does not occur, the heat flux from heat transfer theory probably 

provides a reasonable estimate of the flux and can allow a reasonable estimate of the film 

thickness (in areas where breakdown does not occur). Where film breakdown does 

occur, friction will be higher, so basing friction on a fixed film thickness (calculated 

using heat transfer theory) is a reasonable and probably somewhat conservative approach. 

The thickness is calculated using equation 1 above, along with the method of Klimenko 

(1981) for film boiling heat transfer. The thickness is a very weak function of water 

temperature, and over the range of water temperatures from 10 °F (-12 °C) to 90 °F (32 

°C), the calculated film thickness is 0.063 mm; it is therefore taken as a constant 

(independent of water temperature) for purposes of LNG consequence assessment. Note 

that this value can be used for both the fire case and non-fire case and is used only for 

purposes of estimating friction between the LNG pool and water surface. 

The Webber spread method was modified to use this friction estimate, and calculations 

were made for the two example scenarios presented in Section 3.3 of the report. For 

comparison purposes, calculations were also made for these scenarios with the friction 

term set to zero. The resulting maximum pool radii and evaporation times are shown in 

Table 1. Note that the calculations of these results also incorporate the modifications 

discussed in Items A, B, D, and E. 

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Table 1 Comparison of Pool Behavior for Various Friction Models 

Hole diameter 1 m 5 m 

Friction model Ignoring Film 

Boiling1

Film 

Boiling2

Inviscid 

Flow3

Ignoring Film 

Boiling1

Film 

Boiling2

Inviscid 

Flow3

Maximum pool 

radius 

420 ft 

(130 m) 

420 ft 

(130 m) 

420 ft 

(130 m) 

600 ft 

(180 m) 

1,100 ft 

(350 m) 

1,200 ft 

(370 m) 

Total evaporation 

duration 51 min 51 min 51 min 13 min 5.3 min 5.2 min 

1

 Original method in the report, which ignores the effect of the vapor film on friction 

2

 Friction based on shear stress in the vapor film 

3

 Setting the friction term to zero 

Comparing the results for the 1-m diameter hole show that the results are identical (to 

two significant figures) regardless of which of these models is chosen for friction. The 

results for the 5-m hole show that basing friction on shear stress in the vapor film results 

in a pool radius 83% larger than when effects of film boiling are ignored, and they show 

that accounting for friction based on shear stress in the vapor film does have a small 

effect over assuming inviscid flow. 

These results support the conclusion that friction is unimportant for many spills, 

including the 1-m hole example shown here. They also support the conclusion that for 

large and rapid spills, friction effects do begin to have an effect on pool spread. Friction 

effects would also become more important if modeling an even larger or more rapid spill. 

While the effects of friction will be small in many cases, accounting for these effects can 

improve the analysis for large, rapid spills, and the straightforward method presented here 

(with friction based on shear stress in the vapor film) is recommended. 

D. Handling of Effect of Decreasing Spill Rate on Pool Spread 

Drs. Havens and Spicer commented that “the assumption that the spreading will stop 

when the evaporation rate of the pool becomes equal to the release rate is incorrect.” 

This comment refers to a simplification made in the spread calculations for spills from 1-

m holes. In this scenario, as the pool spreads, the spill rate into the pool is decreasing (as 

the LNG liquid height in the storage tank decreases). As the comment indicates, the 

report methodology made the simplification that spreading stops once the pool grows to 

the point where the evaporation rate is equal to the current spill rate. In reality, the pool 

would continue to spread for a short time and then regress to the point where the 

evaporation and spill rates matched. Due to computational difficulties caused by the 

dynamic nature of the spread model combined with the decreasing spill rate, the 

recommended methods take the simplification of assuming that the pool stops spreading. 

Without the simplification, the pool spreads to the point that all of the LNG evaporates 

before the spill has stopped, which is also clearly incorrect. 

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While this simplification has little impact on the overall consequence assessment results, 

it may result in slight underprediction of peak pool size. Therefore, the recommended 

approach has been modified to handle the computational difficulties in a different and 

probably more realistic manner. 

Specifically, the methodology is modified to allow the pool to continue to spread beyond 

the point where the evaporation and spill rates match. Beyond this point, the evaporation 

rate will exceed the spill rate, so the pool will begin to thin rapidly. The computations 

monitor for the condition where the pool would be emptied while the spill is still in 

progress, and the pool size is decreased immediately to the size that can be supported by 

the spill rate occurring at that moment. From this point on, the pool size will decrease as 

the spill rate decreases. 

This revised method is also a simplification but is more realistic than the previous 

approach. It approximates the situation where the pool thins at the edge (after the spill 

rate decreases below the evaporation rate) and the radius of the pool decreases rapidly to 

the value sustainable by the spill rate at that point. 

E. Film Boiling Heat Flux 

The comments of Dr. Koopman and the comments of Dr. Raj point out that the heat flux 

estimate based on heat transfer theory (~37 kW/m2

) is significantly lower than estimates 

from LNG spill tests, such as some of those listed in Table 2.1 of the report. This is true, 

but the value is also higher than some of the test results. 

As pointed out in the report, the spread in experimental values causes uncertainty in the 

appropriate value, and heat transfer theory is a reasonable approach to help fill in the gap 

in experimental data. Other researchers have taken this approach, and Waite (1983) 

concluded that simulation using a film boiling heat flux better matches experimental data 

than simulation with a 100 kW/m2

 flux. 

On the other hand, Dr. Raj’s comments point out three factors that may result in higher 

heat flux than predicted with film boiling heat transfer theory: 

(i) the water layer immediately beneath the LNG layer cools off rapidly resulting 

in the temperature difference between the water and LNG to be reduced. This will 

result in the occurrence of nucleate boiling phenomenon with very large increase 

in heat flux. Convective cells in water will not form quickly enough to prevent the 

nucleate boiling to occur. (ii) the film in the film boiling is very unstable even in 

laboratory conditions. The break up of film is more or less certain in ocean 

conditions with currents and waves, (iii) The violence of boiling creates smaller 

chunks of LNG which may intimately “mix” with water thus increasing the 

effective boiling area compared to the nominal area of spreading. 

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On the whole, this is a convincing argument for why the heat flux is expected to be 

higher than predicted by film boiling alone. Unfortunately, this still does not point to an 

appropriate selection. 

The experimental data do indicate that the value is probably below 100 kW/m2

 (assuming 

ice does not form). Dr. Koopman’s comments state that “at China Lake an evaporation 

rate of about 0.17 kg/m2/s was observed for the tests, which corresponds to a heat flux of 

about 85 kW/m2.” This value appears reasonable given the inferred upper and lower 

bounds (37 to 100 kW/m2

) and is recommended in place of the previously recommended 

heat transfer theory. 

F. Pool Fire Flame Length 

Several comments suggest that the use of Thomas’ correlation for flame height 

significantly overestimates this height. 

For very large fires, Thomas’ correlation may overestimate flame length; however, there 

are no data for such fires to confirm or deny this hypothesis. The argument that reduced 

air entrainment (that accompanies increased pool diameter) will tend to reduce the flame 

height is a reasonable argument; however, the review of current data and modeling 

methods did not identify a way to account for this in large LNG pool fires on water. 

In addition, both Mudan (in SFPE 1995) and Rew (1986) have reviewed the available 

correlations for flame length, and both concluded that Thomas’ correlation is the best 

available choice. The use of Thomas’ correlation remains the recommended method. 

G. Pool Fire Smoke Production 

Several comments point out that for large LNG pool fires, significant smoke production 

is expected, and this will tend to obscure the flame and reduce the thermal radiation 

emitted from the fire. Even with a ‘clean burning’ fuel like LNG, this is expected and 

can be seen in still and motion pictures of pool fire tests. 

The difficulty lies in quantifying the effect of smoke on emission of thermal radiation, 

not accounted for in the recommended methods, principally because of the limited 

amount of available experimental data for large LNG pool fires. 

Upon further review, the two-zone flame model, as recommended by Rew (1996), can be 

defended for use with large LNG pool fires. This model divides the cylindrical fire into 

two zones: (1) a lower ‘clear’ zone, where smoke does not obscure the flame and (2) an 

upper zone where smoke does obscure the flame and reduce the amount of thermal 

radiation emitted. 

A correlation used by Pritchard and Binding and recommended by Rew (1996) is 

recommended for estimating the length of the lower (clear) zone. While the data for 

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LNG alone are probably insufficient to support such a correlation, the smoke production 

phenomenon is a result of processes not confined to LNG, and this correlation’s basis on 

a broader set of data can be defended for use with LNG. The correlation (Rew 1996, 

equation 2.27) is a function of pool diameter, non-dimensional burning rate, nondimensional wind speed, and carbon to hydrogen ratio for the fuel. 

Rew also reviewed data for estimating the effect of smoke (i.e., in the upper zone) on 

emission of thermal radiation and recommends values of the unobscured ratio for pool 

diameters less than 10 m (0.77), 10 to 20 m (0.69), and more than 20 m (0.55). These 

values are adopted. 

H. Pool Fire Burning Rate and Surface Emissive Power 

Several comments suggest that the selected burning rate (0.282 kg/m2

/s) and selected 

surface emissive power of the flame (265 kW/m2

) are overly conservative (i.e., too high). 

As pointed out in the report, these may be somewhat conservative, based on review of the 

experimental data for the larger pool fires. However, these are reasonable choices 

(adopted from the work of Rew 1996), and as Dr. Koopman pointed out in his comments, 

generally consistent with the China Lake experiments. 

In addition, given the uncertainty in the two-zone flame model, it is recommended that 

these burning rate and emissive power values be retained to help ensure that the model is 

balanced and can still be easily defended as leaning toward conservatism. 

As a final note, the sensitivity analyses in the report, as well as in this response, indicate 

that final results are relatively insensitive to the choices of these values (particularly the 

burning rate). 

G. Summary of Modifications to Recommended Consequence Modeling Methods 

In accordance with the discussions in the previous sections of this response, the following 

modifications to the report’s consequence analysis methods are recommended: 

• Apply a discharge coefficient of 0.65 to the calculations of outflow from the ship; 

• Approximate the pool shape as a semicircle, rather than a circle; 

• Estimate friction between the LNG pool and water surface based on shear stress in 

the vapor film; 

• Use improved method of handling effect of decreasing spill rate on pool spread; 

• Adopt a value of 85 kW/m2

 for heat flux from the water to the LNG pool; and 

• Incorporate two-zone pool fire model. 

With these refinements, the methods suggested here will produce more realistic results 

than those described in the original report but are still defendable and will still tend to 

provide conservative estimates. Furthermore, there appears to be little room for further 

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improvement in the short term (i.e., based on currently published methods). Modeling 

techniques must be defendable based on science but not to the extent that they are grossly 

conservative; the study’s recommendations strike an appropriate balance between these 

conflicting goals. 

Evaporation Heat Flux Sensitivity Analysis 

Given the recommended modifications to modeling methods, it is important to revisit the 

sensitivity analysis for evaporation heat flux presented in Section 2.3.3 of the original 

report. This analysis examines 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 15,800 BTU/ft2

/hr (50 kW/m2), 26,900 BTU/ft2/hr (85 

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) 

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Results for Evaporation Heat Flux Sensitivity Analysis

Hole diameter 3.3 ft (1 m) 

Initial spill rate 7,600 lb/s (3,400 kg/s) 

Total spill duration 51 min 

Film boiling heat flux to pool 15,800 

BTU/hr/ft2

(50 kW/m2

) 

26,900 

BTU/hr/ft2

(85 kW/m2

) 

31,700 

BTU/hr/ft2

(100 kW/m2

) 

Evaporation mass flux 0.020 lb/ft2/s 

(0.098 kg/m2

/s) 

0.034 lb/ft2

/s 

(0.17 kg/m2

/s) 

0.040 lb/ft2

/s 

(0.20 kg/m2

/s) 

Maximum pool radius 

(semicircular pool) 

520 ft 

(160 m) 

420 ft 

(130 m) 

390 ft 

(120 m) 

Total evaporation duration 51 min 51 min 51 min 

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

Downwind distance to LFL 6,300 ft 

(1,900 m) 

6,500 ft 

(2,000 m) 

6,500 ft 

(2,000 m) 

Time at which LFL reaches maximum 

distance 

18 min 16 min 15 min 

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

Downwind distance to LFL 10,000 ft 

(3,100 m) 

11,000 ft 

(3,300 m) 

11,000 ft 

(3,300 m) 

Time at which LFL reaches 

maximum distance 

29 min 29 min 29 min 

For this scenario, the downwind distances are insensitive to variations in the heat flux 

estimate. This is largely a result of the fact that in all of these cases the evaporation is 

limited by the spill rate. That is, for a significant portion of the spill (following the initial 

spread), the evaporation rate is the same for all three heat fluxes. The different heat flux 

values result in different pool areas, each giving the same evaporation rate. 

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. 

Burning Rate and Surface Emitted Flux Sensitivity Analysis 

This section presents a sensitivity analysis like that presented in Section 2.4.2 of the 

original report. It examines the sensitivity of final consequence analysis results (such as 

distance to a heat flux level of concern) to the values of the burning rate and surface 

emitted flux. To examine the sensitivity, calculations were performed using the revised 

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methods recommended in this response for an example scenario. As in the original 

report, 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 7,600 lb/s (3,400 kg/s) 

Total spill duration 51 min 

Flame surface emitted flux 84,000 BTU/ft2/hr (265 kW/m2) 

Burning rate 0.041 lb/s/ft2

(0.20 kg/s/m2

) 

0.058 lb/s/ft2

(0.28 kg/s/m2

) 

0.074 lb/s/ft2

(0.36 kg/s/m2

) 

Maximum pool radius 390 ft 

(120 m) 

340 ft 

(100 m) 

310 ft 

(93 m) 

Total fire duration 51 min 51 min 51 min 

Maximum flame length (height) 790 ft 

(240 m) 

920 ft 

(280 m) 

1,000 ft 

(310 m) 

Clear flame length at maximum 430 ft 

(130 m) 

590 ft 

(180 m) 

730 ft 

(220 m) 

Flame tilt at maximum radius 35 deg 36 deg 37 deg 

Downwind distance to 

12,000 BTU/hr/ft2 (38 kW/m2) 

920 ft 

(280 m) 

910 ft 

(280 m) 

900 ft 

(280 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

1,100 ft 

(340 m) 

1,100 ft 

(340 m) 

1,100 ft 

(340 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

1,500 ft 

(450 m) 

1,500 ft 

(460 m) 

1,500 ft 

(470 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

2,100 ft 

(630 m) 

2,100 ft 

(650 m) 

2,200 ft 

(670 m) 

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Results for Surface Emitted Flux Sensitivity Analysis

Hole diameter 3.3 ft (1 m) 

Initial spill rate 7,600 lb/s (3,400 kg/s) 

Total spill duration 51 min 

Burning rate 0.058 lb/s/ft2 (0.28 kg/s/m2) 

Maximum pool radius 340 ft (100 m) 

Results for Surface Emitted Flux Sensitivity Analysis (cont’d)

Total fire duration 51 min 

Maximum flame length (height) 920 ft (280 m) 

Clear flame length at maximum 590 ft (180 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) 

780 ft 

(240 m) 

910 ft 

(280 m) 

970 ft 

(300 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

980 ft 

(300 m) 

1,100 ft 

(340 m) 

1,200 ft 

(360 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

1,400 ft 

(410 m) 

1,500 ft 

(460 m) 

1,600 ft 

(490 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

1,900 ft 

(580 m) 

2,100 ft 

(650 m) 

2,200 ft 

(690 m) 

As discussed in the original report, variation in burning rate has competing effects on 

pool size and flame length. As a result, final consequence results (distances) are 

insensitive changes in burning rate. 

Also as shown in the original report, results are affected more by the choice of surface 

emitted flux. However, the results are still somewhat insensitive: 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 19% to 25%. 

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. 

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Consequence Assessment Examples for Pool Fires 

This section presents the example scenarios from Section 3.2 of the report with 

calculations modified as described herein. 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. 

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 sitespecific conditions, and analysis of different or additional scenarios may be appropriate. 

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 parameters are listed below, and Table 2 

summarizes the results of the pool fire calculations for these scenarios. 

Scenarios

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) 

Air temperature: 80 °F (27 °C) 

Relative humidity: 70% 

Wind speed: 20 mph (8.9 m/s) 

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Table 2 Summary of Results for Example Pool Fire Calculations 

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

Initial spill rate 7,600 lb/s (3,400 kg/s) 190,000 lb/s (86,000 kg/s) 

Total spill duration 51 min 2.0 min 

Maximum pool radius (semicircular pool) 340 ft (100 m) 1,000 ft (310 m) 

Total fire duration 51 min 4.2 min 

Maximum flame length (height) 920 ft (280 m) 2,100 ft (630 m) 

Clear flame length at maximum 590 ft (180 m) 890 ft (270 m) 

Flame tilt at maximum radius 36 deg 27 deg 

Downwind distance to 

12,000 BTU/hr/ft2 (38 kW/m2) 

910 ft (280 m) 2,000 ft (620 m) 

Downwind distance to 

7,900 BTU/hr/ft2 (25 kW/m2) 

1,100 ft (340 m) 2,500 ft (760 m) 

Downwind distance to 

3,800 BTU/hr/ft2 (12 kW/m2) 

1,500 ft (460 m) 3,500 ft (1,100 m) 

Downwind distance to 

1,600 BTU/hr/ft2 (5 kW/m2) 

2,100 ft (650 m) 5,000 ft (1,500 m) 

Consequence Assessment Examples for Flammable Vapor Dispersion 

This section presents the example scenarios from Section 3.3 of the report with 

calculations modified as described herein. The scenarios examined are spills that are not 

immediately ignited and subsequently 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. 

These are the same scenarios presented in the previous section 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 listed below, and Table 3 summarizes the results 

of the source term and dispersion calculations for these scenarios. 

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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) 

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 Summary of Results for Example Dispersion Calculations 

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

Initial spill rate 7,600 lb/s (3,400 kg/s) 190,000 lb/s (86,000 kg/s) 

Total spill duration 51 min 2.0 min 

Maximum pool radius (semicircular pool) 420 ft (130 m) 1,100 ft (350 m) 

Total evaporation duration 51 min 5.3 min 

Downwind distance to LFL 11,000 ft (3,400 m) 13,000 ft (4,100 m) 

Time at which LFL reaches maximum distance 29 min 29 min 

Time at which entire cloud 

 drops below LFL 

54 min 30 min 

Downwind distance to 1⁄2 LFL 15,000 ft (4,600 m) 19,000 ft (5,900 m) 

Time at which 1⁄2 LFL reaches 

maximum distance 

35 min 37 min 

Time at which entire cloud 

drops below 1⁄2 LFL 

56 min 38 min 

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Comment 27 

We also received numerous specific comments on the technical aspects of the ABSG 

Consulting study. 

Response: 

This section provides responses to some specific technical comments on the report. This 

does not include responses to all comments; in particular, it does not include responses to 

issues already addressed above. Responses are limited primarily to modeling issues that 

are within the scope of the original report. 

Responses to Comments of Dr. Jerry Havens and Dr. Tom Spicer 

Comment 

Our review of the flame radiation modeling procedure used in the report suggests 

that an error has been made in the computer coding of the equations that assign 

values of the transmissivity factor. We have notified ABS of this concern and have 

asked for clarification. 

Response 

Drs. Havens and Spicer are correct in that the calculations for the example 

problems included an error in the equations for calculating the transmissivity of 

the atmosphere. The result of the error caused the calculation to underestimate the 

amount of thermal radiation that would be absorbed by the atmosphere (i.e., 

between the fire and the downwind distance to the level of concern). This error 

resulted in reporting hazard zone distances that were longer than they should have 

been. It is important to note that this was only a calculation error in the example 

problems and not an error in the actual recommendation of modeling methods. 

This error has been corrected. 

Responses to Comments of Dr. Phani K. Raj 

Comment 

The models used are not balanced or accurate to the same degree of confidence. 

The description of the source of LNG (i.e., the rate of spill, quantity of spill, 

interaction between LNG and any water that may flood the inter-hull space, etc) is 

too simplistic, devoid of realism, whereas the spread of liquid on the water is 

described by a model that can be solved only numerically. Similar is the situation 

with respect to the consideration of the evaporation rates, pool fire size and the 

thermal radiation calculations. Real phenomena have been neglected in 

preference to ease of calculations. 

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Response 

The models recommended are based on currently available techniques, which 

some people might consider simplistic. However, these techniques can help 

provide decision makers with useful information about the possible impact of 

large incidents involving spills of LNG on water. The source characterization is 

more appropriate for some types of scenarios that can be postulated than for 

others. For example, in the case of a scenario involving a significantly larger hole 

in the outer hull than in the inner hull, with the outer hole perhaps reaching the 

water level, the recommended approach is a reasonable approximation. 

Comment 

In paragraph 2 of section 1.4, page 2, the following statement is made; “As the 

cloud mixes with air, it will warm up and disperse into the atmosphere.” 

Dispersion in the atmosphere has nothing to do with warming up. It is aided by the 

entrainment of air due to atmospheric turbulence. Also, the implied statement that 

the cloud warms up and therefore becomes buoyant is a thermodynamically 

incorrect statement. When a cold vapor (LNG vapor) mixes with warm air, the 

vapors heat up; but the air that is mixed with the vapor cools down. The net result 

is that the mixed gas (vapor-air mixture) is ALWAYS heavier than air if the 

original vapor is heavier than air. The only way a heavy gas, such as LNG vapor, 

can become buoyant after mixing with air is if an additional source of heat exists, 

such as the ocean or the sun. In the experiments in hot desert conditions no lifting 

of the LNG vapors was observed. 

Response 

For the types of LNG spills considered, dense-gas effects will remain important 

until after the cloud disperses below the lower flammability limit. However, at 

some point the cloud will be neutrally buoyant; stated simply, the cloud will no 

longer have sufficient density to affect its own dispersion. 

Comment 

The description of the flame spread in a vapor air mixture cloud indicated in 2nd

paragraph of page 3 is also incorrect. Flame spread rate in the open atmosphere 

is NOT a function of the laminar-burning rate; it is entirely dependent on the wind 

speed and the intensity of turbulence generated either by wind or obstructions in 

the path of the flame. 

Response 

The statement in the report is that “The rate at which this flame front travels 

relative to the unburned gas is called the laminar burning velocity.” This 

statement is correct and consistent with commonly used terminology; examples of 

the definition of burning velocity are available in many references, such as NFPA 

(2002) which defines burning velocity as “the rate of flame propagation relative to 

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the velocity of the unburned gas that is ahead of it.” Similarly TNO (1997) 

defines burning velocity as “the velocity of a propagating flame, measured relative 

to the unburnt gases immediately ahead of the flame front.” 

Comment 

No discussion has been provided to contrast the commonly used term “explosion” 

and detonation. The discussion in section 1.4.2, paragraph 1 is too simplistic and 

confusing. The statements imply that if there is some obstruction or small 

confinement, methane air mixture will “explode.” For a detonation wave to be 

sustained in a methane air mixture requires the simultaneous occurrence of 

several conditions, namely, long transition length for deflagration to transit to 

detonation, methane concentration between 9 % and 12 % in air, and no venting 

of the pressure in front of the deflagration fire. None of these conditions are likely 

to be fulfilled in an open or semi open atmospheric environment. That is why one 

does not have any records of methane-air detonations in the open. However, 

explosion is a phenomenon in which the local pressure increases due to the hot 

gases (produced by the fire) confined in a weak enclosure whose walls give way 

quickly relieving the pressure. Building collapses due to methane (natural gas) 

releases into basements and the subsequent ignition is an example of the 

commonly referenced “explosion.” What should be of interest for the LNG hazard 

is the occurrence or not of a detonation type of combustion of a vapor cloud which 

then propagates a blast wave in the air affecting structures and people at large 

distances from the fire. This type of burning has not happened in methane-air 

mixtures even when ignited by an explosive charge. Therefore, the allusion (by 

the authors of the FERC report) to a potential methane air explosion implying that 

“ ... the portion of the cloud within that congested area may generate damaging 

overpressures” is inappropriate. 

Response 

Use of the term “explosion” can be problematic. In some literature, the term 

explosion is used to mean “detonation,” while in other literature the term is used 

more generally to refer to an event that leads to generation of a pressure wave. 

The key point of the paragraph 1.4.2 is that when ignited in open areas, large 

overpressures are not expected; however, when ignited in confined areas or areas 

congested with equipment, higher overpressures can occur. This is consistent with 

guidance on use of the so-called multi-energy techniques widely used for 

explosion modeling (see TNO 1997). While NOT intended to imply that 

congestion will cause a detonation, the statement that portions of the cloud within 

a congested area may generate damaging overpressures is appropriate and 

supported by information in Table 5.5 of TNO (1997). 

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Comment 

The statement “The combustion normally occurs within only portions of the vapor 

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

cloud.” in the paragraph on Flash Fire, page 4, is incorrect. Methane air mixtures 

can be IGNITED only when the methane concentration in the mixture is between 5 

% and 15%. However, when a cloud is ignited it burns even the vapor that is at 

concentrations higher than 15% by ingesting air into the burning plume. Hence, 

once the cloud is ignited (at any concentration between 5% and 15%) it will burn 

all vapors above the LFL concentration and not just the ones in the flammable 

range as implied by the authors. 

Response 

The statement that “combustion normally occurs within only portions of the vapor 

cloud” is supported by evidence from LNG spill experiments. However, as Dr. 

Raj points out, ingestion of air into the burning plume can result in combustion in 

a larger portion of the cloud than is within the flammability limits at the time of 

ignition. 

Comment 

The spreading model attributed to Weber is “solved” numerically without 

describing the procedure properly. For example, the procedure requires the 

specification of the value of radius and the velocity of spread at the initial time 

(second order differential equation in the radius). The authors seem to use an 

initial radius of Volume (1/3) as the starting radius. It is unclear what spill 

volume they use initially. Similarly, the “spread front” thickness is estimated by 

assuming that the layer is viscous/surface tension force controlled. It is well 

known that when the spread starts it is controlled by inertial forces and not 

viscous or surface tension. There are a number of questions on the use of the 

Weber model that have not been explained by the authors.

Response 

A full description of the Webber model may be found in TNO (1997) and the 

original work published by Webber (1986, 1987). The implementation presented 

in the report does estimate the radius at the very beginning of the spill as the 

volume to the 1/3 power. The starting volume is the amount spilled during the 

first time-step. After the first time-step, the radius is calculated using the Webber 

model equations. Estimating the starting radius using volume to the 1/3 power is 

an approximation, but because it only affects the radius very early in the spill, it is 

judged a reasonable approach. It is notable that others have also used volume to 

the 1/3 power as a starting point, including the authors of SOURCE5 who use this 

approximation for the starting radius for ‘instantaneous’ spills. Regarding control 

of spread by viscous or inertial forces, the Webber model allows the user to avoid 

artificially choosing which of these dominates by including both throughout the 

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spread. It is true that the inertial forces dominate initially, but that does not mean 

that friction forces are not present. 

Comment 

After going through significant spread calculations using numerical solution 

techniques, the authors make the following statement (last but one paragraph of 

page 13): “It is also useful to note -----. In a case such as this, the pool area can 

be estimated using the spill rate and the evaporate rate per unit area...” This 

statement has no meaning. When the spill rate is changing with time (as the 

square root of the head of liquid depth in the tank) what spill rate is to be used to 

calculate the maximum area? Certainly not the initial maximum spill rate! 

Secondly, the authors do not define quantitatively the conditions under which the 

above “approximation” can be used to calculate the maximum pool dimensions. 

Response 

As the example problem presented for a spill from a 1-m hole shows, there can be 

a point in time where the spill and evaporation rates match, after which the pool 

radius decreases with the decreasing spill rate. However, without performing 

detailed spreading calculations, it is not possible to identify when or if this will 

occur. 

Comment 

The authors provide a table of comparison of LFL distances obtained for different 

assumptions for the value of the heat flux from water to LNG pool. The results 

presented in the table on page 18 are some what counter intuitive. As the heat flux 

increases (making the vapor generation rate higher) the down wind flammable 

distance for an unignited cloud decreases. This is completely contrary to what one 

expects; as the source rate increases, all things about the weather conditions 

being the same, the down wind LFL distance increases. 

Response 

While the spill rate and duration are the same for all three cases, the transient 

evaporation rate looks different for each case because of the differences in heat 

flux values and corresponding differences in pool spread. Figure 1 shows the 

evaporation rate versus time for all three heat flux values. 

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Figure 1 Evaporation Rate Versus Time for Heat Flux Sensitivity

These differences in transient evaporation account for the counterintuitive trend in 

heat flux and downwind distance to the LFL. 

However, this issue is somewhat moot because the modeling improvements 

recommended in this document (see Items C and D in particular), result in 

evaporation rate curves that are more realistic and different from these. 

Comment 

The authors base all of the hazard area calculations on the basis of the maximum 

pool diameter. It is noted that the fire lasts only a fraction of a second at this 

maximum diameter after which it extinguishes itself. All thermal radiation hazard 

criteria have some exposure time requirements (skin burn about 30 seconds). An 

exposure time of 30 s or 45 s forms a substantial fraction of the pool spread time. 

Hence, basing the hazard area on the maximum diameter is incorrect. 

Response 

For purposes of providing information to assist decision makers with facility siting 

issues, a thermal flux level of concern is appropriate. The recommended value of 

1,600 BTU/hr/ft2

 (5 kW/m2) is also used in other regulatory applications, such as 

with onshore facility siting (49 CFR 193 and NFPA 59A) and EPA risk 

management programs (40 CFR 68). 

0

500

1000

1500

2000

2500

3000

3500

4000

4500

5000

0 500 1000 1500 2000

Time (s)

Evaporation Rate (kg/s)

25 kW/m2 Flux 37 kW/m2 Flux 100 kW/m2 Flux

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It is true that the pool fire is predicted to last for only a very short time at the 

maximum diameter, and as discussed in the report, it is also important to consider 

exposure time; Table 2.6 in the report lists some expected effects of a 1,600 

BTU/hr/ft2

 (5 kW/m2) thermal flux level for various exposure times. For example, 

second-degree burns are expected in approximately 30 seconds. So, basing the 

level of concern on thermal flux at the maximum diameter is conservative. 

A useful question then is: just how much conservatism is introduced by use of the 

maximum diameter? The answer of course depends on the scenario being 

considered. If the thermal flux is close to the level of concern for a significant 

time, then basis on maximum diameter does not introduce much additional 

conservatism, and if the thermal flux is not close to the level of concern except for 

a very short time, basis on maximum diameter is more conservative. 

Looking at the example problems presented in Item I can provide some insights to 

the degree of conservatism. To do this, a thermal dose criterion of 13 BTU/ft2

(150 kJ/m2

) for second degree burns, as suggested by Prough (1994) and given in 

Table 2.4 of the report, was selected. This thermal dose corresponds to an 

exposure to 1,600 BTU/hr/ft2

 (5 kW/m2) for 30 seconds, so with short exposure, it 

can be argued that this dose criterion is roughly equivalent to the 1,600 BTU/hr/ft2

(5 kW/m2

) flux criterion. Using an exposure time of 30 seconds and the actual 

(transient) pool diameters for the example problems, the distance to the thermal 

dose criterion was calculated. In the case of the 1-m hole, the 30 seconds was 

taken as 15 seconds before and 15 seconds after the peak thermal flux, and in the 

case of the 5-m hole, the 30 seconds was taken as the final 30 seconds of the fire 

event (since the maximum occurs at the end of the event). The resulting distances 

to these levels of concern are shown in Table 4. 

Table 4 Comparison of Thermal Flux and Thermal Dose Levels of Concern 

Hole diameter1 3.3 ft (1 m) 16 ft (5 m) 

Downwind distance2 to thermal flux of

1,600 BTU/hr/ft2 (5 kW/m2) 

2,139 ft (652 

m) 

5,008 ft (1,527 

m) 

Downwind distance2 to thermal dose 

of 

13 BTU/ft2 (150 kJ/m2) 

2,096 ft (639 

m) 

4,936 ft (1,505 

m) 

1 Scenario details are given in Item I. 2

 Calculated downwind distance values are shown for comparison purposes; values would not 

normally be reported with this number of significant digits

These results show that for the types of scenarios considered here, basing the level 

of concern on thermal flux introduces very little additional conservatism, even 

though the fire is predicted to be at its maximum diameter for only a very short 

time. 

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Comment 

It is unclear whether the authors used the criterion of heat energy dosage (and the 

probit equations) that are discussed in the report in calculating the hazard zone 

dimensions instead of the damage heat flux criterion. If the dosage models are not 

used, then they should not be indicated in the report. 

Response 

Section 2.7 of the report provides an overview of the data and methods available 

for estimating the effects on people and structures that result from thermal 

radiation exposure. Dosage or probit methods were not used in the examples in 

the report, but these are important data to include in any overview of the topic. 

Responses to Comments of Analytical & Computational Engineering, Inc. 

Comment 

Submitted comments include several recommendations for use of computational 

fluid dynamic (CFD) modeling. 

Response 

While the recommended methods can provide useful information to decision 

makers, there is certainly room for future improvement of modeling of events 

involving LNG spills on water, and CFD approaches could well prove to be useful 

tools in such efforts. However, the work described here focused on identifying 

currently available methods for assessing consequences of such spills. With 

regard to the problems of spread of LNG on water and modeling of pool fires, the 

literature review did not identify any CFD models that could be readily applied 

and defended. 

With regard to pool fires, the statement in the report that CFD models provide 

little or no benefit over the solid flame model is based largely on the conclusions 

of Cowley (1992). This work was sponsored by 29 (mostly industry) 

organizations and published by the Health and Safety Executive. It involved an 

evaluation of semi-empirical models (like the recommended solid flame model), 

field models (e.g., CFD models), and integral models. This report states that: 

Well validated, soundly based semi-empirical models are currently the best 

models which are available for the prediction of heat fluxes to objects 

outside flames. They can only be used however within the range of 

conditions given by the experimental data used in their derivation. Thus, 

for example, they are only applicable for open fires which are unaffected by 

the presence of structures nearby or within the flame. 

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In addition, regarding field models the report states that: 

Current models are essentially research tools. In general they have weak 

descriptions of the gaseous combustion, the soot production and oxidation 

and the radiative heat transfer sub-models. These are not criticisms of the 

models as research tools under development. As such they test combustion 

knowledge and provide pointers for experimentalists to tackle. They 

demand a fundamental knowledge of combustion and heat transfer hitherto 

somewhat ignored in large-fire science. 

Our criticisms are directed rather towards the use of these models as 

commercial hazard analysis tools. For most of the problems where they 

have been applied they have not been validated properly. They are not

capable of producing accurate quantitative predictions. 

While it is likely field models have improved significantly in the 12 years since 

publication of this report, as previously stated, the literature review did not 

identify any CFD models that could be readily applied and defended. 

Cowley’s caution to only use the semi-empirical models with the range of 

conditions given by the experimental data is not taken lightly, and in 

recommending such an approach for very large LNG pool fires, it is recognized 

that the empirical correlations used are based on smaller pool fires than the sizes 

considered here. However, the application of the methods recommended by the 

study will tend to be conservative (i.e., overestimate rather than underestimate the 

hazard zone). 

For those not familiar with CFD, it is worthwhile to mention that such modeling 

approaches do not provide a panacea for some of the tough problems in modeling 

LNG spills. For example, in modeling pool spread and evaporation, the 

recommended approach uses a value for heat flux from the water to the pool that is 

an estimate based largely on experimental data. In applying CFD methods, an 

analyst will have essentially two options (1) select a heat flux value as was done in 

the study or (2) attempt to model the complex boiling process in a detailed 

manner. If the first option is selected, the results will suffer from uncertainty 

because of uncertainty in the selected value (as the recommended approach does). 

If the second option is selected, the analyst has significant challenges in modelling 

the complex process correctly, which will include identification of the correct set 

of physics equations (i.e., from first principals) that must be solved and applying 

the correct boundary conditions. With a turbulent, multi-phase problem such as 

this, the analyst may also be required to use some empirical data to solve the 

problem, even numerically. 

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CFD models do have the potential to provide more detailed information, such as a 

more realistic pool shape (accounting for factors such as wind and currents). In 

addition, one particular area where CFD methods have potential to improve 

modeling results (or at least decrease uncertainty) is in applying more fundamental 

physics to help overcome the need to use empirical correlations (e.g., Thomas’ 

correlation for flame length) outside the range for which they were designed. 

Comment 

Please provide backup calculation to the assertion that LNG will stay negatively 

buoyant until after it disperses below the LFL. ACE performed simple hand 

calculation and obtained a much different result. 

Response 

This assertion was based primarily on review of the results of the dispersion 

modeling performed for the examples included in the report. 

Responses to Comments of Project Technical Liaison Associates, Inc. 

Comment 

The spread on water uses essentially smooth surface and only mentions the effect 

of a non-smooth surface. But the effect of even small waves was shown to be 

dramatic. There is no guidance as to how FERC should take into account waves. 

As even small waves will effect the peripheral portions of the pool, this factor must 

be considered. 

Response 

It is not clear where the commenter believes that “the effect of even small waves 

was shown to be dramatic.” The only model of waves identified was that 

presented by Quest (2001, 2003), and the smallest waves considered in that study 

were over 1.89 ft (0.575 m) tall. Typical harbor areas will often have waves 

significantly lower than this, and in many harbors calm waters are not uncommon. 

Comment 

The model for thermal radiation flux from a sustained pool fire may be 

satisfactory but the application to a spreading pool of LNG is not a convincing 

picture of reality, considering: 

• The pool is not ignited until it reaches maximum diameter 

• A pool of a few millimeters thickness at the edge will not create a fully 

developed cylindrical flame above the pool or burn long enough to 

establish the model’s flame geometry 

• Adequate air will not be available to create the same flame size and 

intensity as modeled. 

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Response 

In the pool fire examples presented, it is assumed that the fire is ignited 

immediately upon release. The pool spread algorithm takes this into account in 

estimating the pool radius and pool thickness as a function of time. Regarding 

adequacy of air, the recommended modification to use a two-zone flame model 

helps address this issue. 

Comment 

The common portrayal of the vapor dispersion hazard area is a circle around the 

release point whose radius is the maximum dispersion distance at the most 

unfavorable conditions. The actual area at risk for an incident is the area of the 

plume (based on width at distance) of the LFL concentration. No consideration of 

this limitation is given in the report. 

Response 

Carefully defining the area that could be within a flammable cloud would be an 

important part of a detailed quantitative risk assessment. However, when 

providing facility siting decision makers with information about distances to 

which effects of interest might reach, representing the hazard zone as a circle will 

usually be adequate. 

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References 

AIChE (2000) 

Center for Chemical Process Safety of the American Institute of Chemical Engineers, 

Guidelines for Chemical Process Quantitative Risk Analysis, Second Edition, ISBN 0-

8169-0720-X, New York, 2000. 

Cowley (1992) 

L.T. Cowley and A.D. Johnson, Oil and Gas Fires: Characteristics and Impact, OTI 

92596, Health and Safety Executive, 1992. 

Fay (1969) 

James. A. Fay, “The Spread of Oil Slicks on a Calm Sea”, Oil on the Sea, ed. D. P. Hoult, 

Plenum, New York, 1969. 

Hoult (1972) 

David P. Hoult, “Oil Spreading on the Sea,” Annual Review of Fluid Mechanics, 1972. 

Lees (1996) 

F.P. Lees, Loss Prevention in the Process Industries, Second Edition, ISBN 0-7506-

1547-8, Butterworth-Heinemann, Oxford, 1996. 

NFPA (2002) 

National Fire Protection Association, NFPA 68 Guide for Venting of Deflagrations, 

National Fire Protection Association, 2002. 

Prugh (1994) 

R.W. Prugh, “Quantitative Evaluation of Fireball Hazards,” Process Safety Progress, The 

American Institute of Chemical Engineers, Vol. 13, No. 2, pp. 83-91, April 1994. 

Quest (2001, 2003) 

Quest Consultants, Inc., Modeling LNG Spills in Boston Harbor, Letter from John 

Cornwell of Quest Consultants to Don Juckett of the U.S. Department of Energy, October 

2, 2001; Letter from John Cornwell of Quest Consultants to Don Juckett of the U.S. 

Department of Energy, October 3, 2001; and Letter from John Cornwell of Quest 

Consultants to Clifford Tomaszewski of the Office of Natural Gas & Petroleum Import & 

Export Activities, November 17, 2003. 

TNO (1997) 

C.J.H. van den Bosch and R.A.P.M Weterings, eds., Methods for the Calculation of 

Physical Effects (TNO Yellow Book), TNO, The Hague, The Netherlands, Third edition 

1997. 

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Waite (1983) 

P.J. Waite, R.J. Whitehouse, and E.B. Winn, “The Spread and Vaporization of Cryogen 

Liquids on Water,” Journal of Hazardous Materials, Vol. 8, pp. 165-184, 1983. 

Webber (1986) 

D.M. Webber and P.W.M. Brighton, Similarity solutions for the spreading of liquid 

pools, SRD R371, United Kingdom Atomic Energy Authority, Safety and Reliability 

Directorate, July 1986. 

Webber (1987) 

D.M. Webber and P.W.M. Brighton, An integral model for spreading, vaporising pools. 

SRD R390. United Kingdom Atomic Energy Authority, Safety and Reliability 

Directorate. August 1987. 

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