NUREG-1547

NISTIR 5842

Methodology for Developing and

Implementing Alternative

Temperature-Time Curves for

Testing the Fire Resistance of

Barriers for Nuclear Power

Plant Applications

L. Y. Cooper, K. D. Steelder, NIST

Prepared for:

U.S. Nuclear Regulatory Commission

Office of Nuclear Reactor Regulation

Washington, DC 205554001

U. S. Department of Commerce

Technology Administration

National Institute of Standards and Technology

Gaithersburg, MD 20899
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NUREG-1547

NISTIR 5842

Methodology for Developing and

Implementimg Alternative

Temperature-Time Curves for

Testing the Fire Resistance of

Barriers for Nuclear Power

Plant Applications

ManuscrptIComleted:Iayi199

Manuscript Completed: May U%9

Date Published: August 1996

L Y. Cooper, K. D. Steckler, NIST

Building and Fire Research Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899

Division of Systems Safety and Analysis

Office of Nuclear Reactor Regulation

US. Nuclear Regulatory Commission

Washington, DC 20555-0001

National Institute of Standards and Technology

Technology Administration

U.S. Department of Commerce

ABSTRACT

Advances in fire science over the past 40 years have offered the potential for developing

technically sound alternative temperature-time curves for use in evaluating fire barriers for areas

where fire exposures can be expected to be significantly different than the ASTM E-1 19 standard

temperature-time exposure. This report summarizes the development of the ASTM E-119,

standard temperature-time curve, and the efforts by the federal government and the petrochemical

industry to develop alternative fire endurance curves for specific applications. The report also

provides a framework for the development of alternative curves for application at nuclear power

plants. The staff has concluded that in view of the effort necessary for the development of nuclear

power plant specific temperature-time curves, such curves are not a viable approach for resolving

the issues concerning Thermo-Lag fire barriers. However, the approach may be useful to licensees

in the development of performance-based fire protection methods in the future.

iii NUREG-1547

CONTENTS

Ene

ABSTRACT .. . . . . . . . . . . . .. . . . . . . . . . . . i"

EXECUTIVE SUMMARY ......................................... ix

PART 1. HISTORY AND USE OF THE ASTM STANDARD

TEMPERATURF-TIME CURVE .................................... I

LIST OF FIGURES - PART I ................................... 2

SUMM4ARY - PART I ......................................... 3

1.1 OBJECTIVE ............................ *..... ........... 4

1.2 OVERVIEW OF THE ASTM E 119 TEST METHOD ................. 4

1.3 HISTORY OF ASTM E 119 STANDARD TEMPERATURE-TIME

CURVE .................... ............................. 5

1.4 RELEVANCE OF THE STANDARD TEMPERATURE-TIME CURVE .... 6

1.5 APPLICATION OF THE STANDARD TEMPERATURE-TIME CURVE ... 7

1.6 CONCERNS ABOUT STANDARD TEMPERATURE-TIME CURVE/FIRE

LOAD METHODOLO*GY .................................. 10

1.6.1 Equal-Area Hypothesis ................................ 10

1.6.2 Factors Other than Fire Load ............................ 10

1.6.3 The Standard Curve Itself (outside context of equal-area

hypothesis) ......................................... 11

1.7 CONCLUSIONS ..... 0... 00... .. .. ..... 0 .. ........ .. . 13

1.8 REFERENCES .......................................... 14

APPENDIX - PART 1. CHRONOLOGY OF FIRE TESTING PRIOR TO THE

ADOPTION OF THE STANDARD TEMPERATURE-TIME

CURVE FOR U.S. FIRE TESTS .......................... 25

REFERENCES - APPENDIX - PART 1 .......................... 30

V NUREG-1547
Contents

PART 2. HISTORY AND USE OF ALTERNATIVE TEMPERATURETI1 E CURVES ........ .... ......................................... 31

LIST OF FIGURES - PART 2 .................................. 32

SUMMARY - PART 2 ........................................ 33

2.1 OBJECTIVE .................................. . . . . 34

2.2 INTRODUCTION ....................... ................ 34

2.3 ALTERNATIVE TEMPERATURE-TIME CURVES REPRESENTING

ROOM FIRES... 35'"

2.3.1 Qualitative Curves or Curves Based on Qualitative Arguments ..... 35

2.3.1.1 Corson's Curves . r.e.......o........ ............. 35

2.3.1.2 Seigel's Qualitative Curves ..... ....... ......... 36

2.3.2 Curves Determined by Mathematical Models .. . o..o. .. o... ... 37

2.3.3 Curve Determined by Full-Scale Burnout Experiments - the NBS

Recreation Room Curve ..... 39

2.4 TEMPERATURE-TIME CURVES REPRESENTING OPEN-AIR

PETROCIEMNICAL CURVES ..... ................. ...... ... 40

2.5 CONCLUSIONS ..... .. . ........... .................. . 44

2.6 REFERENCES . . ......... o.............. .. . .. .. . ... 45

PART 3. FEASIBILITY OF DEVELOPING AND IMPLEMENTING

NPP-SPECIFIC DESCRIPTIONS OF FIRE ENVIRONMENTS FOR USE IN

EVALUATING THE FIRE RESISTANCE OF FIRE BARRIERS .... ....... 61

LIST OF FIGURES - PART 3 .... ... .. 62

LIST OFITABLES-PART3 .................................. 62

SUMMARY - PART -A 3 3................................... 63

3.1 INTRODUCTION: THE OBJECTIVE AND AN OVERVIEW. ......... 63

3.2 TWO TYPES OF FIRE BARRIERS: STRUCTURAL BARRIERS

AND WRAP ASSEMBLIES ................................. 65

NUREG-1547 Vi
Contents

3.3 ASTM E 119 AND THE ASTM E 119-Type APPROACH TO THE

EVALUATION OF STRUCTURAL-TYPE BARRIER PERFORMANCE.. 66

3.4 STRUCTURAL BARRIERS, WRAP ASSEMBLIES AND ASTM E 119 ... 66

3.4.1 A Characteristic that Distinguishes Between Structural

Barriers and Wrap Assemblies ........................... 66

3.4.2 Describing Fire Exposure and Evaluating Fire Barrier

Performance ........................................ 67

3.5 EXPOSURE TO THE INDIRECT AND DIRECT EFFECTS

OF THE FIRE: TWO TYPES OF THREAT .................... 68

3.5.1 The One- or Two-Layer Description of the Fire Environment

and Zone-Type Compartment Fire Modeling .................. 68

3.5.2 The Indirect and Direct Effects of the Fire and Applicability

of ASTM E 119-7ype Test Methods ........................ 69

3.5.3 Using the Direct and Indirect Fire Exposure to Define the

Fire-Barrier Threat ............................ .... 70

3.6 A CONCEPT FOR EVALUATING THE FIRE PERFORMANCE OF FIRE

BARRIERS IN NUCLEAR POWER PLANTS - AN ALTERNATIVE TO

TRADITIONAL APPLICATION OF ASTM E 119 ................. 71

3.6.1 An Alternative to ASTM E 119 .......................... 71

3.6.2 A Concept For Evaluating the Fire Performance of Fire Barriers

In Nuclear Power Plants ................................ 72

3.7 A COMPARTMENT FIRE MODEL FOR USE IN EVALUATING

THE DIRECT AND INDIRECT THREATS TO FIRE BARRIERS

IN NPPs .............................................. 73

3.7.1 Mathematical Fire Models and Associated Computer Codes ....... 73

3.7.2 General-Purpose and Special-Purpose Compartment Fire Models ... 74

3.7.3 Features of a Compartment Fire Model Suitable For Evaluating

Direct and Indirect Threats to NIP Fire Barriers ............... 75

3.7.4 Suitability of Using COMPBRN HI for the Evaluation of

Fire Barrier Performance ................................ 78

vii NUREG-1547
Contents

3.7.5 Suitability of Using the Fuel Load/Ventilation Method

Fire Model for the Evaluation of Fire Barrier Performance ........ 79

3.8 THE THREATENING COMBUSTIBLES FOUND IN NPPs;

IMPLEMENTING THE ASSUMPTION OF FULLY-DEVELOPED

BURNING ............................................. 80

3.8.1 The Threatening Combustibles ........................... 80

3.8.2 Simulating Fully-Developed Burning of Cable Trays ............ 81

3.8.3 Simulating Burning of Combustible/Flammable Liquid Pool Fires ... 85

3.9 A NPP-SPECIFIC FIRE MODEL AND SCENARIO-SPECIFIC THREATS

TO FIRE BAR IER ... 0.... 0... 0*000...... 0........... ..... ... 87

3.9.1 NPP-Speciflc ASTM E 119-Type Temperature-Time Curves and

Indirect Threat Exposures ........................... . 87

3.9.2 The Bounding-Temperature Principle and NPP-Specific Fire Curves . 88

3.9.3 Developing Multiple, NPP-Specific, Temperature-Time Curves ..... 89

3.9.4 NPP-Specific Direct Fire Exposures ........... 90

3.10 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS ......... 90

3.10.1 General Sunmary and Conclusions ....................... 90

3.10.2 Recommended Tasks ................................. 92

3.11 NOMENCLATURE ..................................... 93

3.12 REFERENCES ........................................ 94

APPENDIX A POLICY ISSUE COMMISSION INFORMATION PAPER:

NUCLEAR POWER PLANT-SPECIFIC TIME-TEMPERATURE

CURVES FOR TESTING AND QUALIFYING FIRE BARRIERS . 103

NUREG-1547 o°fi
EXECUTIVE SUMMARY

BACKGROUND

As part of its regulatory responsibilities, the U.S. Nuclear Regulatory Commission (NRC)

requires certain provisions for fire protection in U.S. nuclear power plants (NPPs). One aspect

of NPP fire safety has to do with the performance of fire barriers designed explicitly to protect

components, equipment, etc. on the protected side of the barrier, from potential threatening fire

environments on the fire-exposed side. These fire barriers, which are in the form of walls,

floor/ceilings, partitions, cladding, or wrapping, are designed to prevent the passage of fire, or

excessive heat from a fire, for the time period necessary to shutdown critical equipment safely,

or to control the fire prior to its threatening structural elements or areas critical to general plant

safety. This property of a barrier to withstand fire or give protection against it is known as fire

resistance and the elapsed time during which a barrier exposed to fire continues to exhibit fire

resistance is known as fire endurance [ES-I].i Despite the distinction, the two terms often are

used interchangeably.

In the United States, ASTM E 119, "Standard Test Methods for Fire Tests of Building

ConstructionMaterials" [ES-2], or one of its counterparts for doors [ES-3], windows [ES-4], or

through-penetration fire stops [ES-5], is the generally accepted standard method for evaluating

and rating the fire resistance of building fire barriers. The method involves furnace-fire exposure

of a portion of a full-scale fire-barrier specimen. The furnace-fire environment follows a particular,

monotonically-increasing,temperature-time history, which is defined in the test-method document.

This temperature-time history (or curve) of the standardASTM E 119 fire is consideredto represent

a "severe" fire exposure. The test method specifies explicit acceptance criteria2 that involve the

measured response of the barrier test specimen at the time into the standard fire exposure that

corresponds to the desired barrier rating. For example, a barrier design is said to have a 3-hourfireresistance rating if the tested specimen meets all specified acceptance criteria during at least three

hours of a standard fire exposure. The fire-resistance rating, in turn, qualifies the barrier design for

certain uses. Here the term qualifies is intended to mean that the barrier design meets or exceeds

the fire-resistance requirements of a building code or other regulation.

In 1976, NRC adopted the ASTM E 119 methodology to regulate fire barriers protecting redundanttrain safe-shutdown systems in NPPs [ES-6]. NRC set the required ASTM E 119 fire-endurance

rating at 3-hours, consistent with requirements for U.S. industrial properties classified by insurance

INumbers in brackets with designation [ES-#] refer to literature references listed at the end of this

Executive Summary.

2 These criteria include non-failure of structural components in the test assembly, passage of neither flame

nor very high temperature gas to the unexposed side of the barrier, and limits on the temperature rise on

the unexposed-surface of the test assembly above its initial temperature.

ix NUREG-1547
Executive Summary

carriers as "Improved Risk" or "Highly Protected Risk" properties [ES-6]. The safe-shutdown

regulations supplemented existing, ASTM E 119-based, general regulations for fire barriers in areas

critical to general plant safety [ES-7].

Advances in fire science over the past 40 years have offered the potential for developing technically.

sound alternative temperature-time curves for use in evaluating fire barriers for areas where fire

exposures can be expectedto be significantly different than the standard temperature-time exposure.

For example, during the 1970s and 1980s, several countries, including the United States, developed

and implemented technically-sound alternative curves for testing fire barriers that might be

subjected to open-air hydrocarbon pool fires (see Part 2, section 2.4). The NRC staff has initiated

the current effort to investigate the feasibility of developing alternative temperature-time curves for

the qualification of fire barriers used to protect cabling and equipment necessary to achieve safe

shutdown on the basis of realistic fire hazards found in NPPs.

OBJECTIVE

The aim of the current study is to propose a methodology for developing and implementing NPPspecific descriptions of fire environments and associated ASTM-type furnace test methods. Here

the terminology ASTME 119-type test is used to refer to a test method that basically follows the

ASTM E 119 test procedures, but where the ASTM E 119 standard temperature-time curve is

replaced with a relevant, NPP-specific, alternative curve.

APPROACH

The approach taken in the current study consists of three steps or tasks: 1) review the history of the

ASTM E 119 temperature-timecurve to assess its current applicability and limitations in simulating

real fires; 2) review the history of efforts to develop alternative curves and the methodologies used;

and 3) use the findings from (1) and (2), knowledge of NPP construction, fuel types and loads, and

state-of-the-art fire science to propose a methodology for developing and implementing NPPspecific descriptions of fire environments and associated ASTM-type temperature-time curves and

test methods.

NUREG-1547 x
Executive Summary

RESULTS AND CONCLUSIONS

Part 1. History of Standard Temperature-Time Curve

1. The historical evidence indicates that the ASTM E 119, standard, temperature-time curve was

prescribed in 1917 with very little knowledge of the levels and the temporal development of

temperatures in actual room fires. The standard curve was basically an idealization of temperaturetime curves measured in furnaces at various laboratories, was deemed to represent a severe fire, and

was intended only to provide a basis for comparing the fire endurance of building assemblies using

a simple test. Full-scale room burnout tests conducted at the National Bureau of Standards in the

1920s establishedthat the actual temperature histories of room fires differed significantly from the

standard curve.

2. Using fire-endurance ratings, which were based on the standard curve, and the equal-area fireseverity hypothesii, Simon Ingberg at the National Bureau of Standards correlated the fire load in

a room with the fire endurance necessary to withstand complete burnout of the room. The result was

a relatively simple, albeit somewhat technically-weak, system of analysis for determining the fireendurance required for building elements. Technical shortcomings of the analysis include: no

technical basis for the equal-area hypothesis; real room fire intensities are not a sole function of fire

load; and temperatures of real fires can rise faster than the standard curve. In spite of these

shortcomings, the fuel-load/equal-area/standard-curve method was widely accepted and, indeed,

remains a landmark development in the history of fire protection engineering. It simply was better

than anything else available. Subsequent work, reported in Part 2 of this document, showed that

temperatures of real fires often exceeded the standard curve. Today, the National Fire Protection

Association acknowledges that although the fire-load method is technically obsolete, it is still useful

in situations that do not involve high heat-release rate combustilies and in which fire conditions do

not produce temperatures significantly higher or lower than the standard temperature-time curve

[ES-s].

3. Fire-endurancetesting in the United States has not changed substantially since the publication of

the standard temperature-time curve in 1918. Continued use of the fire-load/standard-curve

methodology has been justified on the bases that the analysis: is judged to be conservative (more

severe) with respect to the maximum fire exposure in many occupancies; has a proven record with

respect to safety; is tied to a standard test; and is relatively simple to use. Arguments against

changing the standard curve include: that a large amount of experience has been gained with the

existing standard temperature-time curve; re-radiation from the exposed surface makes the exact

temporal details of the curve unimportant; and no other curve will eliminate all the objections.

4. Critics argue that real fires often rise faster and/or exceed the standard curve. Although the

repercussions of the inaccuracies of the standard fire-load analysis were not so great in the era of

relatively massive fire barriers (1920s and 30s), the consequences could be much greater now that

'The equal-area fire-severhy hypothesis is that the area beneath a temperature-time exposure curve is a

measure of the intensity or severity of a fire, and all fires with equal-area exposures are equally severe.

xi NUREG-1547
Executive Summary

lighter-weight barriers must withstand fires fueled by modem synthetic materials. Moreover,

improved analytical techniques are available that can avoid at least some of the objectionable

assumptions inherent in the standard fire-load analysis.

Part 2. History of Alternative Curves

1. The known shortcomings of the standard fire-load analysis/standard temperature-time curve and

the availability of new analytical tools provided the impetus in the 1950s and 60s to seek alternatives

to the standard analysis/curve.

2. Hydrocarbon Processing Industry (HPI) temperature-time curves (e.g., [ES-9]) appear to be the

only alternative curves widely used for fire-resistancetesting. The HPI (or, for short, hydrocarbon)

curves simulate the direct-impingement exposure from a open-air hydrocarbon pool fire and are a

more severe exposure than the standard curve.

3. The NBS recreation-room study in the 1970s produced an alternative curve for testing the fire

resistance of residential floor constructions in a furnace. NBS recommended that the new curve be

used in a new ASTM fire-resistancetest for rating residential floor constructionswhich require fireresistance ratings of less than one hour. This recommendation, however, was rejected on the basis

that the sixteen full-scale room-burnout experiments were too limited in scope. Although achieving

consensus in the highly diversified ASTM fire-protectioncommunity may be more difficult than in

the smaller, more focused, NPP industry, this NBS experience at least cautions that changes might

not be accepted by the NPP industry without substantial experimental underpinnings.

4. No record has been found to indicate that fire-resistance testing has been conducted to qu.ajify a

building element or assembly using any curve other than the standard temperature-time curve or a

hydrocarbon curve. In the course of the present review, the only record of fire-resistance testing of

building elements or systems in a furnace using an alternative curve was that found in the NBS

recreation-room study, which was a research, rather than a qualification program.

5. The literature suggests that the development of any new temperature-time curves for compartment

applications likely will rely heavily on mathematical modeling (the tool that became available in the

1950s) because of the large number of configurations that need to be considered. Indeed,

temperature-time curves generated by mathematical room-fire models are sanctioned by Swedish

regulations for design calculations as alternatives to the more traditional classification of building

components; namely, by furnace-testingbuilding components against the ISO 834 [ES- 10] standard

temperature-time curve, which is very similar to the ASTM E 119 standard curve.

6. To date, a major weakness of room-fire models lies in their inability to simulate accurately

burning rates of real fuels under real-fire conditions. For example, much of the modeling work

aimed at creating alternative temperature-time exposures has been based on the burning

characteristics of cellulosic fuels. Also, such models only simulate the simplest type of room

ventilation - an opening to the outside environment. It seems that the development of credible

temperature-timecurves for nuclearpower plant (NPP) applicationswill require better understandirg

NUREG-1547xi xii
Executive Summary

of the burning behavior of NPP-specific fuel packages - for example, cable bundles - and more

advanced ventilation considerations.

Part3. Feasibility ofDeveloping and ImplementingNPP-specific Descriptions

of Fire Environments for Use in Evaluating the Fire Resistance of Fire

Barriers.

1. NPP fire barriers include structuralbarriers, useful in isolating a compartment of fire origin from

adjacent spaces, and wrap assemblies, used to isolate and protect plant equipment, cables, etc.,

within a compartment of fire origin, from the effects of exposure to the fire environment.

2. The nature of fire-barrierexposure to an NPP-compartment fire environment can be categorized

as a) direct exposure to the most extreme zones of the fire environment, e.g., direct, sustained

exposure of the barrier to the flame, and b) indirect exposure, where the fire barrier is mainly

exposed to the average properties of the overall fire environment. Both kinds of threats need to be

addressed. ASTM E 119-type test methods, employing new, alternative, temperature-time curves,

are appropriate for the indirect threats. Other tests may need to be devised to simulate direct

exposure threats. Here, the hydrocarbon exposure curve may play a role. Compartment-fire model

simulations should be useful in defining quantitatively both the indirect and direct test exposures.

3. A methodology for evaluating the fire resistance of NPP fire barriers is presented that removes

weaknesses of and/or introduces flexibifity to the traditional ASTM E 119 approach. This relies on

a combined experimental and analytic approach that involves the Bounding-Temperature Principle.

(i.e., if the temperature-timecurve of one fire environment bounds that of another, then, relative to

the threat to structural integrity of a NPP fire barrier, the bounding-curve environment is the more

severe.) Fire-resistanceexperiments would involve ASTM E 119-type tests employing alternative

temperature-time furnace fires deduced from reliable fire-model simulations. Analysis would

involve compartment fire modeling methods, where computer simulations would be carried out with

a new, advanced, special-purpose, zone-type fire model.

Specifically:

a. The new compartment-firemodel must include features particularlyrelevant to simulating

fire environments that threaten NPP fire barriers, from the point of view of boh direct and

indirect fire exposure. These included: the simulation of fully-developed burning of

extensive dense arrays of cable trays (i.e., all exposed surfaces of a combustible cable are

supplying fuel (losing mass) due to either heating by the fire environment or surface

combustion), both under fuel-controlled and ventilation-controlled conditions; the

simulation of combustible/flammable liquid pool fires; and advanced means of modeling

ventilation and radiation-heat-transfer-related phenomena. A new special-purpose model

with these features could be developed as a customized advanced version of an existing, twolayer, multi-room, zone-type fire model, e.g., CFAST [ES-1 1].

xii. xiii NUREG-1547
Executive Summary

b. The new model would be used to simulate a wide variety of potential fire scenarios in

rooms of fire origin of selected NPPs. The simulations would lead to new insights on the

characteristics of real, fire-barrier-threatening, NPP fire environments. Based on

applications of the Bounding-TemperaturePrinciple, the simulated fire scenarios would lead

to a series of NPP-specific test fire curves covering a wide range of NPP-type fire severities.

c. An experimental study on available ASTM E 11 9-type test furnaces would be carried out

to establish that these new test fire curves (instead of the standard ASTM E 119 fire curve)

were achievable for use in ASTM E 11 9-type barrier rating tests. Then ASTM E 11 9-ype

tests, using the new NPP-specific test fire curves, would be established as the method of

evaluating the fire performance of NPP fire barriers.

4. A significant effort will be required to carry out this plan. Due to knowledge gaps in critical areas

such as burning rate and ventilation effects in NPP-specific environments, the modeling work will

require a substantial experimental component. Indeed, the experimental aspects, including full-scale

burnout of fuel packages and furnace fire-resistancetests, are similar in scope to the NBS recreationroom study (see Part 2, section 2.3.3), which was a multi-year effort.

RECOMMENDATIONS

Consistent with the above, it is recommended that the following tasks be carried out with the goal

of establishing a reliable methodology for evaluating NPP fire barrier performance:

1. Develop a new, special-purpose, NPP-specific fully-developed fire model capable of

simulating fire environments that threaten NPP fire barriers. It is recommended that this be

developed as an advanced version of an existing multi-room compartment fire model, e.g.,

CFAST [ES- 11]. The new model should include the advanced modeling features identified

in the section 3.7.3, "Features of a Compartment Fire Model Suitable For Evaluating

Direct and Indirect Threats to NPP Fire Barriers." These include: the simulation of

fully-developedburning of extensive dense arrays of cable trays, both under fuel-controlled

and ventilation- controlled conditions; the simulation of combustible/flammable liquid pool

fires; the simulation of the fire environment in multi-room facilities (at least two adjacent

spaces); and advanced means of modeling ventilation and radiation-heat-transfer-related

phenomena

2. Carry out full-scale experimental verification of the advanced modeling methods of item 1,

especially those aspects of the new model associated with the simulation of burning cable

trays and combustible/flammablefuel fires in enclosed spaces. Also, carry out experiments

to better evaluate and characterize the fire hazard in NP~s introduced by electrical

panels/cabinets.

3. Use new model simulations to determine the direct-exposure threat to fire barriers, and use

these to establish experimental methods to evaluate barrier fire performance relative to the

direct exposure threat.

NUREG-1547 xiv
Executive Summary

4. Use the new model to carry out an extensive simulation study of selected NPP fire areas.

Results of this would be used to establish the characteristics of real, fire-barrier-threatening,

NP? fire environments and to identify a series of NPP-specifictest fire curves to replace the

ASTM E 119 standard fire curve.

5. Carry out an experimental study on available ASTM E 11 9-type test furnaces to establish that

the new test fire curves of item 4 are attainable and reproducible.

6. Use the results of items 4 and 5 to establish an ASTM E 119-type method of evaluating the

performance of structural fire barriers relative to the indirect exposure; establish

corresponding methods for wrap-assembly fire barriers.

ACKNOWLEDGEMENTS

Funding for this project was provided by the U.S. Nuclear Regulatory Commission, Office of

Nuclear Reactor Regulation. The authors wish to thank K. Steven West, Patrick M. Madden, and

Edward A. Connell, all of the NRC, for their valuable advice throughout the course of this work.

REFERENCES

[ES-i] ASTM E 176-93a, Standard Terminology of Fire Standards, Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and Materials,

Philadelphia, PA, 1995.

[ES-2] ASTM E 119-88, Standard Test Methods for Fire Tests of Building Construction

Materials, Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American

Society for Testing and Materials, Philadelphia, PA, 1994.

[ES-3] ASTM E 152-81a, Standard Methods of Fire Tests of Door Assemblies, Annual Book of

ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and

Materials, Philadelphia, PA, 1994.

[ES-4] ASTM E 163-84, Standard Methods of Fire Tests of Window Assemblies, Annual Book

of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and

Materials, Philadelphia, PA, 1994.

[ES-5] ASTM E 814-94b, Standard Teat Method for Fire Tests of Through-PenetrationFire Stops,

Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for

Testing and Materials, Philadelphia, PA, 1995.

[ES-6] Appendix A to Nuclear Regulatory Commission Branch Technical Position APCSB 9.5-1,

"Guidelines for Fire Protection for Nuclear Power Plants Docketed Prior to July 1, t977",

August 1976.

XV NUREG-1547
Executive Summary

[ES-7] Nuclear Regulatory Commission, Title 10, Part 50, Appendix A, Criterion 3, Fire

Protection Design Criteria, Federal Register, Vol. 36, p. 3256, February 20, 1971.

[ES-8] Fire Protection Handbook, Seventeenth Edition, Arthur E. Cote, Editor-in Chief, National

Fire Protection Association, Quincy, MA, pp. 6-76, 6-77, July 1991.

[ES-9] ASTM E 1529-93, Standard Test Methods for Detennining Effects of Large Hydrocarbon

Pool Fires on Structural Members and Assemblies, Annual Book of ASTM Standards, Sec.

4, Construction, Vol. 4.07, American Society for Testing and Materials, Philadelphia, PA,

1993.

[ES-10] International Standard 834, Fire-resistancetests - Elements in Building Construction, Ref.

No.ISO 834-1975 (E), November, 1975.

[ES-11] Peacock, R.D. et al, CFAST, the Consolidated Model of Fire Growth and Smoke

Transport, NIST TN 1299, National Institute of Standards and Technology, Gaithersburg,

1993.

NUREG-1547 xvi
PART 1. HISTORY AND USE OF THE ASTM STANDARD

TEMPERATURE-TIME CURVE
History of Standard Temperature-Time Curve

LIST OF FIGURES - PART 1

Figure 1-1.

Figure 1-2.

Figure 1-3.

Figure 1-4.

Figure 1-5.

Figure 1-6.

Comparison between proposed temperature-time curve for column

tests at Underwriters' Laboratories [1-18] and current temperaturetime curve [1-1] (main figure reproduced from [1-18]) ....... 18

Standard temperature-time curve used in furnace tests, cooling

curves, and curve representing the average temperature in a typical

office occupancy test (reproduced from [1-20]) ............ 19

Fire endurance versus fire load (combustible content) ....... 20

Average Gas Temperatures in Compartments, where 60 (%) denotes

60 kg/rn2 fire load and 50 percent open wall (reproduced from

[1-30]) ....................................... 21

Schematic of standard temperature-time curve and hypothetical

curves for room fires .............................. 22

Theoretical results of Kanury and Holve [1-33] demonstrating the

effect of the form of the temperature-time curve on unexposed-surface temperature .................................... 23

NUREG-1547 2
History of Standard Temperature-Time Curve

SUMMARY - PART 1 -

The standard temperature-timecurve defines the furnace-fire exposure for rating the fire resistance

of fire barriers by ASTM E 119, Standard Test Methods for Fire Tests of Building Construction

Materials [1-1] . The historical evidence indicates that the standard curve was prescribed in 1917

with very little knowledge of the levels and temporal development of temperatures in actual room

fires. Indeed, it appears that the curve was deemed to represent a severe fire and was intended

only to provide a basis for comparing the fire resistance (endurance) of building assemblies.

Full-scale room burnout tests in the 1920s established that the actual temperature histories of room

fires were different from the standard curve. The standard curve, however, was "rescued" by the

concept offire severity, which was considered a measure of the intensity of the fire exposure, and

was defined as the area beneath the exposure temperature-time curve. The hypothesis is that all

fires with equal-area exposures are equally severe. Severity, therefore, provided the connection

between a real fire exposure and the standard temperature-time exposure used to evaluate the fire

endurance of building assemblies. In addition, the intensity or severity of a real room fire was

assumed to be a function of the fire (fuel) load alone. The result was a relatively simple system

of analysis for determining the fire endurance required to withstand a complete burnout of a room

containing a given fire load.

Technical shortcomings of this system of analysis include: no scientific basis exists for the equalarea hypothesis; real room fire intensities are not a sole function of fire load; and temperatures

of real fires can rise faster than the standard curve. In spite of these shortcomings, the fuelload/equal-area/standard-curve method was widely accepted and, indeed, remains a landmark

development in the history of fire protection engineering. Later work, which is reported in Part

2 of this document, demonstrated that temperatures of real fires often exceeded the standard curve.

Today, the National Fire Protection Association acknowledges that although the fire-load method

is technically obsolete, it is still useful in situations that do not involve high heat-release rate

combustibles and in which fire conditions do not produce temperatures significantly higher or

lower than the standard temperature-time curve [1-2].

Continued use of the fire-oad/standard-curvemethodology has been justified on the bases that the

analysis: is judged to be conservative (more severe) with respect to the maximum fire exposure

in many occupancies; has a proven record with respect to safety; is connected to a standard test;

and is relatively simple to use. Arguments against changing the standard curve include: a large

amount of experience has been gained with the existing standard temperature-time curve; reradiation from the exposed surface makes the exact temporal details of the curve unimportant; and

no other curve will eliminate all the objections.

Critics argue that real fires often rise faster and/or exceed the standard temperature-time curve.

Although the consequences of the inaccuracies of the standard fire-load/standard-curve analysis

apparently were not so great in the era of its inception, when fire barriers were relatively massive 0

4 Numbers in brackets with designation [14/] refer to literature references listed at the end of Part I of this

document.

3 NUREG-1547
History of Standard Temperature-Time Curve

and failed at high temperatures, the consequences could be much greater now that lighter-weight

barriers must withstand fires fueled by modem materials. Moreover, for the past three decades,

improved analytical techniques have been available that can avoid at least some of the

objectionable assumptions inherent in the standard fire-load analysis.

The fact remains that fire-resistancetesting in the United States has not changed substantially since

the adoption of the standard temperature-time curve in 1918.

1.1 OBJECTIVE

The purpose of Part 1 of this document is to review the literature concerning the origin,

justification, and use of the ASTM E 119 standard temperature-time curve [1-1] for determining

fire resistance according to American Society for Testing and Materials (ASTM) proceduress.

The scope is limited to the fire-temperature specifications. The merits of a heat flux-time

exposure specifications relative to a temperature-time exposure [1-9,1-10,1-111 are not addressed

in this review.

1.2 OVERVIEW OF THE ASTM E 119 TEST METHOD

In the U.S., ASTM E 119 is the generally accepted standard method for evaluating and rating the

fire resistance of structural-typebuilding fire barriers. The method involves faunace-fire exposure

of a portion of a full-scale fire barrier specimen. The furnace-fire environment follows a monotonically-increasing, temperature-time history, which is specified in the test method document The

latter defines the temperature-time history of the standard ASTM E 119 fire. The test method

specifies explicit acceptance criteria that involve the measured response of the barrier test specimen

at the time into the standard fire exposure, referred to as thefire resistance of the barrier design, that

corresponds to the desired barrier rating. For example, a barrier design is said to have a three-hour

fire-resistanceradng if the tested specimen meets specified acceptance criteria during at least three

hours of a standard fire exposure. The fire-resistancerating, in turn, qualifies the barrier design for

certain uses. Here the term qualifies is intended to mean that the barrier design meets or exceeds

the fire-resistance requirements of a building code or other regulation.

5 The standard curve is employed in ASTM E 119 [1-1], ASTM E 152 [1-3], ASTM E 163 [1-4], and

ASTM E 814 [1-5]. The standard curve is also used in National Fire Protection Association (NFPA)

standards [1-6 to 1-8], which are nearly identical to first three of these ASTM standards.

NUREG-1547 4
History of Standard Temperature-Time Curve

1.3 HISTORY OF ASTM E 119 STANDARD TEMPERATURE-TIME

CURVE

Several accounts of the history of the ASTM E 119 fire resistance (endurance) test have been

published; for example [1-1,1-9,1-10,1-12,1-13,144,1-15,1-16]. Of these, the works by Shoub [I13] and Babrauskas and Williamson [1-15,1-16] are the most complete. Unless noted otherwise, the

following section presents a summary of the Babrauskas and Williamson account of the development

of the standard temperature-time curve.

Prior to 1903, there were no widely accepted standards specifying the conditions for testing building

materials and construction systems for fire endurance. A review of the history of fire testing before

1903 (see APPENDIX - PART 1) shows that each laboratory specified its own test conditions,

including the temperatures to which the test materials and assemblies were exposed. In most cases,

the furnace temperature, averaged over the test period, had to equal or exceed some specified level.

In 1903, the British Fire Prevention Committee (BFPC) proposed the first widely sanctioned

standard test procedure. The standard defined three fire-endurance classifications: Full, Partial, and

Temporary protection from the burnout of a room's contents. A minimum exposure temperature and

minimum test duration were assigned to each class. For example, the "Full" rating, associated with

the most severe of the three classifications, required that the material or assembly not fail when

exposed to a minimum temperature of 982 °C for 4 hours. The "Temporary" rating, for the least

severe classification, required that the assembly not fail when exposed to an 816 °C environment

for 45 minutes.

In 1899, the New York Building Code defined the first standard for fire-endurance testing in the

United States. This was not intended to be a national standard. Nevertheless, with the Baltimore

conflagration of 1904 as an impetus, a national standard [1-17], very similar to the New York

standard, was instituted by the American Society for Testing and Materials (ASTM), committee C-5

(later E-5) in 1907. Like the New York standard, the ASTM standard required that floor assemblies

be exposed to a fire in a furnace "hut" for 4 hours at an average temperature of no less than 926 *C.

Ira Woolson, a pioneer fire researcher who built the first permanent U.S. facilities for testing fire

resistance at Columbia University, was chairman of this ASTM committee.

In 1909, a separate test for walls was included and furnace control was specified further. The

furnace was to be heated to 926 °C within the first 30 minutes and then held at that level for the

remaining 1.5 hours of the test.

During this period, the BFPC standard of 1903 was gaining international recognition.

Acknowledging this, a National Fire Protection Association (NFPA) committee on standards,

chaired by Ira Woolson (no longer at Columbia .University) recommended in 1914 that further

development of the U.S. standard be stopped and that NFPA adopt the BFPC standard with some

modifications. NFPA did not accept this recommendation.

Instead, eleven organizations met to discuss U.S. fire test methods. In 1916, the first of these

meetings produced a recommendation for a standard temperature-time curve for the testing of

5 NUREG-1547
History of Standard Temperature-Time Curve

columns at Underwriters Laboratory in Chicago [1-18]. The motivation for the curve was the

realization that furnace transients needed to be quantified to achieve reproducible results. The

proposed curve was basically an idealization of temperature-time curves measured in furnaces at

various laboratories.

Babrauskas and Williamson report that the curve published in 1916 has not been changed since then.

This apparently is not entirely correct. Figure 1-1, which compares the 1916 curve with the current

curve presented in ASTM E 119-88, shows some difference during the initial steep transient period.

Babrauskas and Williamson report that in the course of establishing the standard curve, the

committee increased the heating rate during the first 10 minutes in order to address concerns that

fires might rise faster in some occupancies. It appears that these adjustments were made subsequent

to publication of the 1916 document, but prior to the issuance of ASTM C 19 (later to become E 119)

on February 24, 19176. The standard curve has not been changed since then.

Since the chosen curve was closest to the New York/Columbiacurves obtained after 1902, when the

average temperature criterion was dropped from 1093 °C to 926 °C7, the standard curve was known

as the "Columbia Curve" in honor of Ira Woolson.

1.4 RELEVANCE OF THE STANDARD TEMPERATURE-TIME CURVE

Ryan [1-14], citing a Southwest Research Institute (SWRI) internal report [1-19], says that, Professor

Woolson at Columbia University established a temperature-time curve at the turn of the century for

severe fires based on data he obtained from visiting fire scenes, talking with the fire service, and

comparing fire debris with known melting points. Since the SWRI report was not available for

review', the origin of the statement could not be traced. Shoub makes a similar statement that the

1917 E 119 curve

"... apparently was based on temperatures found in the various stages of growth of actual

fires in buildings using referencessuch as the observed time offiusion of materials of known

melting points. " [1-13] (emphasis added)

No reference is given. In his 1928 report on room-fire experiments, Ingberg makes the related

statement

6 Semantics has confused somewhat the date of ASTM C19. Babrauskas and Williamson [1-16J report that

ASTM C19 (later renumbered E119) was issued on February 24, 1917 but later refer to the "1918

standard". Ryan [1-14] and Shorter [1-9] report that ASTM E 119 was adopted in 1918. The current

ASTM E 119-88 standard [1-1] states that C19 was first published in 1918 as C19-1917T, which, in current

ASTM notation, would indicate that the standard was approved in 1917 and first published in 1918.

7 Indeed, the average temperature over the first 4 hours of the standard curve is 945 °C, which is close to

the New York requirement of an average of 926 °C over the same period.

a The report could not be located at SWRI.

NUREG-1547 6
History of Standard Temperature-Time Curve

"Indications ofthe intensity of buildingfires have been obtainedfrom fused metals andfrom

general fire effects on materials on which information is extant as to their reaction to

temperature or fire exposure such as in test fires. The fire ruins or reports of fwes give,

however, little information on the duration of the temperatures in any given portion of the

building. The absence of data to enable constructions or devices giving a certain

performance in the standard test to be applied as protection against fire conditions in

buildings with as much precision as results of strength tests are applied for load carrying

purposes, led me to consider the possibility of conducting burning-out tests in suitably

designed structures to obtain the needed information." [1-20] (emphasis added)

Although the last statement neither confirms nor denies a direct link between room-fire temperatures

inferred from fire debris and the standard curve, it does point out that little information on the

temporal development of a real fire can be gleaned from such data.

Regarding a systematic experimental basis, Ryan states that "Thepresent curve at its inception was

not based onfitll scale test data." [1-14]

On this subject, Babrauskas and Williamson report that in 1903 Woolson was conducting tests at

926 IC because, in Woolson's words, "This particular temperature was chosen because it is given

by the New York Building Code as approximatelythe heat ofa burning building." They add that "To

complete the circle, one only needs to know that the New York Building Code used 926 'C as the

temperature of a burning building because Constable' ran his fire tests at that temperature." [1-16]

The systematic measurement of temperatures during room burnout experiments was not initiated in

the U.S. until 1923 11-16], 6 years after the standard temperature-time curve was prescribed.

Although burnout experiments had been conducted in Europe prior to 1917, Babrauskas and

Williamson report [1-16] that ASTM committee C-5 was unaware of them. They conclude that '"..

the standard curve was prescribed in 1917 without the knowledge of what actual temperatures in

buildings might be." [1-16]

Nevertheless, the standard curve was (and still is) considered to represent a fairly severe fire

[1-12,1-21]. At the time of its adoption, the curve was intended only to provide a basis for

comparing the fire endurance of building assemblies using a simple fire test [ 1-14]. Techniques for

more sophisticated analyses were not available.

1.5 APPLICATION OF THE STANDARD TEMPERATURE-TIME CURVE

In 1923, burnout tests were commenced by Simon Ingberg at the U.S. National Bureau of Standards

(NBS). Ingberg was interested in the relationship between the extant standard temperature-time

curve and the duration of temperatures in a room subjected to the burnout of its contents. Office

occupancies were simulated using papers and wooden and steel furniture, which were fairly uniformly dispersed throughout the lower portion of the room. The independent variable of the tests wasfire

9 Superintendent of Buildings in New York, circa 1896.

7 NUREG-1547
History of Standard Temperature-Time Curve

load, which is defined as the average fuel mass per unit floor area of the space. The findings were

published in 1928 [1-20].

Figure 1-2 shows spatially-averagemperature results for one of the occupancy tests. Also shown

is the standard curve that is followi1 during a farnace test. Attached to the standard curve are postshutdown cooling curves obtained from furnace tests at NBS. The substantial differences between

the experimental and standard-plus-cooling curves led Ingberg to formulate his equal-area fireseverity hypothesis;

"An approximate comparative measure of severity is obtained by assuming that the area

under the latter curve'0 [L e., the occupancy test], expressed in degree-hours, gives severity

equivalent to an equal area under the standard exposure curve and the cooling curve

applicable for the given period." [1-20]

"Severity" became the connection between actual fire intensities and fire test exposures based on the

standard temperature-time curve. Severity "validated" the standard curve.

The fire-severity concept allowed Ingberg to translate the full-scale test results into equivalent fire

endurances (determined by furnace tests) required by the walls, etc., to withstand complete burnout

of the spaces. Analysis of the experiments in this fashion led to Ingberg's 1928 correlation between

fire load and fire endurance [1-20], which is displayed as the solid curve in Fig. 1-3." It was also

concluded from these experiments that the standard curve represented the maximum severity of a

fire resulting from the burnout of a brick wood-joistedbuilding and its contents [1-21 ]. At least one

room-fire experiment, however, produced a faster initial temperature 'ise than the standard curve [1-.

20].

The simplicity of the fire-load method, its link to a standard fire test, and Ingberg's publication of

detailed fire-load data for a wide range of occupancies [1-22,1-23] led to near-universal acceptance

by building code authorities [1-9,1-14].

Since the fire load method of analysis represents a landmark development in the history of fire

protection engineering, it is importantto summarize the conditions/assumptionson which it is based,

as well as the reasoning behind the assumptions.

* Equal-area assumption: Although Ingberg was aware of technical problems with this

hypothesis (see below), he justified its use by the lack of an alternative simple method.

• Limited experimental basis: Ingberg's experiments involved only wood and paper fuels

spread out fairly uniformly over the lower region of the room. He was aware that other

10 The area was computed with respect to a baseline of either 150 'C or 300 °C depending upon the

combustibles present.

11 In this figure, fire load is expressed as combustible content (potential heat), which is the sum of the

products of the mass of each fuel and its heat of combustion, divided by the floor area.

NUREG-1547 8
History of Standard Temperature-Time Curve

occupancies can have different fuels with different calorific content and possibly different

burning rates [1-20]. Nevertheless, he concluded that a lot of materials have calorific

contents in the range of wood and paper [1-22]. For example, he assigned wood, paper,

cotton, wool, silk, straw, grain, sugar, and similar organic compounds their actual weights.

Materials with higher calorific contents were assigned larger wood-equivalent weights. For

example, he assigned animal and vegetable oils, fats, waxes, petroleum products, asphalt,

bitumen, paraffin, pitch, alcohol, and napththelene twice their actual weights.

- Assumption that fire intensity is solely a function of fire load: The intent of the fire load

concept was to provide a simple basis for comparisons that was better than anything available at the time. No attempt was made to account for other factors, such as ventilation and

the nature and disbursement of the combustibles, all known to influence the behavior of the

fire [1-10,1-14].

Despite these technical shortcomings, the validity of the method was not questioned until the 1960s

[1-9] when improved analytical techniques were developed. Also, the advent of lighter-weight

building assemblies raised concern. Because the heavy, non-melting", fire-resistant materials

available in Ingberg's time usually had to reach high temperatures to fail, the consequences of the

inaccuracies of the method were less important in that era [1-24]. Also, there is the issue of the

effect that rapid temperature rise and attendant thermal shock might have on the integrity of lighter

barriers. (See section 2.4 and footnote in section 3.9.2).

Nevertheless, the fire load method is still in use today. The current fire load/endurance relationship

is shown in Fig. 1-3 as a dashed curve. After first acknowledging ihat the method is "technically

obsolete", the 1991 NFPA Fire Protection Handbook [1-2] states that the method still is useful in

many situations because

"In many cases, this original fire severity/fire load relationship was more severe than is

indicated by more accurate analysis." [1-2]

and

"Although the technique has its limitations, the fire severity/fire load relationship still

provides an approximate but conservative estimate of the probable maximum fire severity

in residential, institutional, and some commercial occupancies. Fire loads should not be

used as an approximate indicator of fire severity with combustibles having a high heatrelease rate and when fire conditions can produce temperatures signi.ficantly higher or

lower than the standard temperature-time curve." [1-2] (emphasis added)

1 Here and in section 1.6.1, Babrauskas uses the term "melting" to indicate a phase change in the barrier

material.

9 NUREG-1547
History of Standard Temperature-Time Curve

1.6 CONCERNS ABOUT STANDARD TEMPERATURE-TIME CURVE/

FIRE LOAD METHODOLOGY

1.6.1. Equal-Area Hypothesis

Ingberg was aware of the approximate nature of the hypothesis when he wrote

"...that equal area under temperature-timefire exposure curves standfor equivalent severity

of exposure is an approximation only, since in the heat conductivity equation applicablefor

the case the exposing temperature enters directly as a factor in the expression for

temperature obtaining at any point within an exposed body, while the time, which is the

other factor in the time-temperature area, enters as an exponent. " [1-20]

Babrauskas raises four physical objections to the equal-area hypothesis:

"1) The outstanding example is when materials can undergo a phase change at some

temperature T, Consider two fires, one which heats up some portion ofa building assembly

beyond its melting point and one which does not.." It is clearly unreasonable to say that

those two fires might somehow be equated

2) If some building assembly is combustible, its rate of mass loss, and thereby degradation,

can usually be expressed by an equation of the form

hnmdt = A exp(-EIRI)

This relationship is patently not linearly dependent on the gas temperature.

3) The main mechanism of heat transfer to the wall, at temperatures above 500 IC is

radiatioL The radiant flux is proportional to 7V not 7".

4) Finally, some building materials derive their protection primarily from a latent heat of

hydration. Gypsum wallboard is the most common example of this kind of protection. For

a material ofthat kin4 degradation is proportional to the heat input, which is not a linear

function of temperature." [1-24]

The hypothesis simply cannot be defended on scientific grounds.

1.6.2. Factors Other Than Fire Load

It has been known for a long time that fire load is not the only important factor that determines the

intensity of a fire in a room. Commenting on the accounts of Ingberg's experiments (1-20],

1

3 Although this is certainly a theoretical possibility, no information was found during the present review

that indicates the extent to which it might be a real problem. To be a practical problem, it seems the

critical temperature, Tc, would have to be fairly high, so that it might be reached in one fire but not

another.

NUREG-1547 110
History of Standard Temperature-Time Curve

Robertson and Gross found clear evidence that ventilation and the nature and disbursement of the

room contents "... had an important influence in modifying fire behavior..." [1-10]

Lie reports in a 1968 document that factors affecting room-fire temperatures include

"1. the average amount of combustible material present per unit floor area (so-called fire

load)

2. the form in which the combustible material is present

3. the size of the ventilation openings

4. the dimensions of the interior space

5. the insulating capacity of the walls." [l7-25]

Odeen identifies a similar list of factors in his statement

"When designing a construction from a fire resistance point of view it is necessary to

determine the temperature levels to which the construction might be exposed The extremely simplified method adopted today, assuming the temperature influence upon a

construction member to be dependent, in principle, on the quantity of combustion material

available (the fire load), is then not satisfactory. Thus no consideration is taken of such

factors as rate of combustion (which is influenced by e.g. the air supply and the degree of

fineness and particle geometry of the combustible material) the thermal properties of

construction enclosing the fire cell, as well as those enclosed within the fire cell and by the

dimensions of the fire cell."' 4

Many reports have documented the influence of ventilation on room-fire behavior [1-10,1-27,1-

28,1-29,1-30]. For example, Kawagoe [1-29] found that a small window-area-to-fuel ratio

produced long duration fires of moderate intensity, but large window-area-to-fuel ratios produced

shorter duration fires that often exceeded the standard temperature curve. Similar results reported

by Butcher et al. [1-30] are presented in Fig. 1-4.

1.6.3. The Standard Curve Itself (outside context of equal-area hypothesis)

Babrauskas warns

"It is sometimes asserted that even though under many conditions the standard curve

exposure will not be similar to the expected realistic exposure, it is still justified to use the

curve. The argument usually runs " we know the test results will not be the same as

endurance time in fire, but so long as the test exposure is fully standardized, all materials

will be testedfairly and adequate ranking established" It should be adequately clear that

such a viewpoint is untenable. Compare, for instance, an assembly using materials which

14 Odeen, Kai, Theoretical Study of Fire Characteristics in Enclosed Spaces, thesis, Royal Institute of

Technology, Stockholm, Sweden, unpublished.

I1I NUREG-1547
History of Standard Temperature-Time Curve

are good insulators and have low T,15 with one using poorly insulating, high T, materials.

When tested under appropriately low temperatures the first assembly will prove superior,

but at higher temperatures the second will be better. In general, there is no way of

assuring that even relative rank will be preserved; in consequence testing under conditions

greatly differingfrom those of the expectedfire is not a suitable design philosophy." [1-24]

Minor and Berry [1-3 1] advocated keeping the standard curve, but avoiding use of the equal-area

conversion. They suggest using a given barrier only in situations where the anticipated exposures

will never exceed the temperature or test duration of the standard curve to which the barrier was

actually exposed. Referring to Fig. 1-5, a 1-hour barrier would be acceptable for exposure A, but

would be unacceptable for exposures B and C. A logical extension of this approach would be to

establish alternative "standard" curves for qualifying barriers for short "hot" fires and long "mild"

fires. Indeed, Corson [1-21] in 1953 and Siegel [1-32] in 1967 proposed using a series of different

curves to represent different fire loads. The history and use of alternative curves are discussed in

Part 2 of the present document.

Regarding the standard curve, it is worth reiterating that the NFPA Fire Protection Handbook

cautions that

"Fire loads should not be used as an approximate indicator of fire severity with

combustibles having a high heat-release rate and when fire conditions can produce

temperatures significantly higher or lower than the standard temperature-time curve."

[1-2]

Kanury and Holve [1-33] performed a theoretical analysis of fire-endurance testing of wood and

gypsum board panels. This was a thermal model that accounted for charring and desorption

processes under the assumption that the degraded panel remained intact; that is, the overall

thickness of the panel remained constant. For analysis purposes, they adapted two of the ASTM E

119 critical points as follows:

ASTM E 119 Critical Point I Adapted to

Average unexposed-surface Unexposed surface reaches 1.5

temperature reaches 139 *C times its initial absolute

above its initial temperature. temperature.

Flame or hot gases penetrate Significant pyrolysis occurs at

barrier. unexposed surface.

Using the adapted criteria, they report that, for a wood barrier, the exact shape of the

temperature-time curve has little effect on the theoretical fire-endurance time because surface re15 T, denotes a critical temperature, such as the temperature at which a phase change occurs.

NUREG-1547 12
History of Standard Temperature-Time Curve

radiation compensates for fairly large changes in exposure conditions16 . For example, Fig. 1-6

shows results of calculations using both the standard exposure and a "peaked" (firegrowth/decay) exposure. The difference in fire-endurance time is about 12 percent. From these

and other calculations involving realistic peaked exposures, they conclude that

"... there is no reason to discard the standard T/(t) curve as a specified exposure source

for fire performance evaluation of materials, even though it superficially fails to be a

realistic duplicate of any one particularfull-scale enclosure fire exposure history." [1-33]

It should be noted that the conclusion is drawn from limited results from a theoretical model that

excludes mechanical effects.

Other proponents for the existing standard temperature-time curve argue that it should not be

changed because a great deal of experience has been gained using this curve, and there is no

other curve that will eliminate all the objections [1-31]. The use of multiple standard curves is

viewed as a great complication to evaluation of the fire-endurance of constructions that would

make comparisons of individual constructions difficult [1-9] and would not be well received by

testing laboratories without a rigorous demonstration of its value [1-10].

1.7 CONCLUSIONS

1. The historical evidence indicates that the standard temperature-time curve was prescribed in

1917 with very little knowledge of the levels and the temporal development of temperatures in

actual room fires.

2. It appears that the standard curve was deemed to represent a severe fire and was intended

only to provide a basis for comparing the fire endurance of building assemblies using a simple

test.

3. Although full-scale room burnout tests in the 1920s established that the actual temperature

histories of room fires were different from the standard curve, the standard curve was "rescued"

by the technically-flawed equal-area hypothesis; namely, that all fires with equal-area

temperature-time exposures are equally severe or intense.

4. Using fire-endurance ratings, which were based on the standard curve, and the equal-area

hypothesis, Ingberg correlated the fire load in a room with the fire endurance necessary to

withstand complete burnout of the room. The result was a relatively simple, albeit somewhat

technically-weak, system of analysis for determining the fire-endurance required for building

elements.

16That is, the integrated net absorbed heat flux (incident flux minus the reflected and re-radiated fluxes),

not integrated exposure temperature, is of primary importance.

13 NUREG-1547
History of Standard Temperature-Time Curve

5. Despite technical shortcomings (such as: having no scientific basis for the equal-area

hypothesis; real room fire intensities are not a sole function of fire load; and real fires can rise

faster than the standard curve) the fuel-load/equal-area/standard-curve method was widely

accepted and, indeed, remains a landmark development in the history of fire protection

engineering. Today, the National Fire Protection Association acknowledges that although the

fire-load method is technically obsolete, it is still useful in situations that do not involve high

heat-release rate combustibles and in which fire conditions do not produce temperatures

significantly higher or lower than the standard temperature-time curve.

6. Fire-endurance testing in the United States has not changed substantially since the publication

of the standard temperature-time curve in 1918. Continued use of the fire-load/standard-curve

methodology has been justified on the bases that the analysis: is judged to be conservative (more

severe) with respect to the maximum fire exposure in many occupancies; has a proven record

with respect to safety; is connected to a standard test; and is relatively simple to use.

Arguments against changing the standard curve include: that a large amount of experience has

been gained with the existing standard temperature-time curve; re-radiation from the exposed

surface makes the exact temporal details of the curve unimportant; and no other curve will

eliminate all the objections.

7. Critics argue that real fires often rise faster and/or exceed the standard curve. Although the

repercussions of the inaccuracies of the standard fire-load analysis were not so great in the era of

relatively massive fire barriers, the consequences could be much greater now that lighter-weight

barriers must withstand fires fueled by modem materials. Moreover, improved analytical

techniques are available that can avoid at least some of the objectionable assumptions inherent in

the standard fire-load analysis.

1.8 REFERENCES

[1-1] ASTM E 119-88, Standard Test Methods for Fire Tests of Building Construction

Materials, Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American

Society for Testing and Materials, Philadelphia, PA, 1994.

[1-2] Fire Protection Handbook, Seventeenth Edition, Arthur E. Cote, Editor-in Chief,

National Fire Protection Association, Quincy MA, pp. 6-76, 6-77, July 1991.

[1-3] ASTM E 152-81a, Standard Methods of Fire Tests of Door Assemblies, Annual Book of

ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and

Materials, Philadelphia, PA, 1994.

[1-4] ASTM E 163-84, Standard Methods of Fire Tests of Window Assemblies, Annual Book

of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and

Materials, Philadelphia, PA, 1994.

NUREG-1547 14
History of Standard Temperature-Time Curve

[1-5] ASTM E 814-94b, Standard Teat Method for Fire Tests of Through-Penetration Fire

Stops,Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American

Society for Testing and Materials, Philadelphia, PA, 1995.

[1-6] NFPA 251, Standard Methods of Fire Tests of Building Construction and Materials,

National Fire Codes, Vol. 6, 1995 Edition, National Fire Protection Association,

Quincy, MA.

[1-7] NFPA 252, Standard Methods of Fire Tests of Door Assemblies, National Fire Codes,

Vol. 6, 1995 Edition, National Fire Protection Association, Quincy, MA.

[1-8] NFPA 257, Standard for Fire Tests of Window Assemblies, National Fire Codes, Vol.

6, 1995 Edition, National Fire Protection Association, Quincy, MA.

[1-9] Shorter, G.W., The Fire Protection Engineer and Modem Building Design, Fire

Technology, National Fire Protection Association, pp. 206-213, August 1968.

[1-10] Robertson, A.F., and Gross, D., Fire Load, Fire Severity, and Fire Endurance, ASTM

STP 464, American Society for Testing and Materials, pp. 3-29, 1970.

[1-11] Harmathy, T.Z., and Lie, T.T., Fire Teat Standard in the Light of Fire Research, Fire

Test Performance, ASTM STP 464, American Society for Testing and Materials, pp. 85-

97, 1970.

[1-12] Bird, E.L., and Docking, S.J., Fire in Buildings, Adam & Charles Black, London,

1949.

[1-13] Shoub, H., Early History of Fire Endurance Testing in the United States, ASTM STP

301, American Society for Testing and Materials, pp. 1-9, 1961.

[1-14] Ryan' J.F., Perspective on Methods of Assessing Fire Hazards in Buildings, ASTM STP

502, American Society for Testing and Materials, pp. 11-23, 1972.

[1-15] Babrauskas, V., and Williamson, R.B., The Historical Basis of Fire Resistance Testing

- Part I, Fire Technology, No. 3, pp. 184-194, August 1978.

[1-16] Babrauskas, V., and Williamson, R.B., The Historical Basis of Fire Resistance Testing

- Part II, Fire Technology, No. 4, pp. 304-316, November 1978.

[1-17] Standard Test for Fire-Proof Floor Construction, ASTM Proceedings, Vol. 7, pp. 179-

180, 1907.

[1-18] Fire Tests of Building Columns, NFPA Quarterly, Vol. 9, No. 3, pp. 253-260, January

1916.

15 NUREG-1547
History of Standard Temperature-Time Curve

[1-19] Bickerdorf, F.W., and Yuill, C.H., An Investigation of the Hazards of Combustion

Products in Building Fires, Southwest Research Institute Project 3-1237-3, October 14,

1963.

[1-20] Ingberg, S.H., Tests of the Severity of Building Fires, NFPA Quarterly, Vol 22, 1, pp.

43-61, July 1928.

[1-21] Corson, R.C., The Significance of Fire Loading, NFPA Quarterly, Vol. 47, No. 1, pp.

65-72, July 1953.

[1-22] Ingberg, S.H., Fire-Resistance Classifications of Building Constructions, Building

Materials and Structures Report BMS92, Superintendent of Documents, Washington,

DC, October 7, 1942.

(1-23] Ingberg, S.H., Dunham, J.W., Thompson, J.P., Combustible Contents in Buildings,

Building Materials and Structures Report 149, Superintendent of Documents,

Washington, DC, July 25, 1957.

[1-24] Babrauskas, V., Fire Endurance in Buildings, Report No. UCB FRG 76-10, Ph.D.

dissertation, University of California, Berkeley, November 1976.

[1-25] Lie, T.T, Fire Resistance in Buildings, National Research Council of Canada Technical

Translation 1334, Ottawa, 1968.

[1-26] Not used.

[1-27] Heselden, A.J.M., Parameters Determining the Severity of Fire, Behavior of Structural

Steel in Fire Symposium No. 2, Proceedings of a Symposium held at the Fire Research

Station, Boreham Wood, Herts, on January 24, 1967, Her Majesty's Stationery Office,

pp. 19-27, 1968.

[1-28] Butcher, E.G., and Law, M., Comparison Between Furnace Tests and Experimental

Fires, Behavior of Structural Steel in Fire Symposium No. 2, Proceedings of a

Symposium held at the Fire Research Station, Boreham Wood, Herts, on January 24,

1967, Her Majesty's Stationery Office, pp. 46-55, 1968.

[1-29] Kawagoe, K., Fire Behavior in Rooms, Report No. 27, Building Research Institute,

Japan, September 1958.

[1-30] Butcher, E.G, Chitty, T.B., and Ashton, L.A., The Temperature Attained by Steel in

Building Fires, Fire Research Technical Paper No. 15, H.M. Stationery Office, London,

1966.

NUREG-1547 16
History of Standard Temperature-Time Curve

[1-31] Minor, E.E., and Berry, D.L., Nuclear Power Plant Fire Protection - Fire Barriers

(Subsystems Study Task 3), NURBG/CR-0468 SAND78-1990 RP, Sandia Laboratories,

Albuquerque, NM September 1979.

[1-32] Seigel, L.G., The Severity of Fires in Steel Framed Buildings, Behavior of Structural

Steel in Fire Symposium No. 2, Proceedings of a Symposium held at the Fire Research

Station, Boreham Wood, Herts, on January 24, 1967, Her Majesty's Stationery Office,

pp. 57-63, 1968.

[1-33] Kanury, A.M., and Holve, D.J., A Theoretical Analysis of the ASTM E-119 Standard

Fire Test of Building Construction and Materials, Final Report, Contract 4-36105,

Stanford Research Institute, Menlo Park, CA, August 1975.

17 NUREG-1547
History of Standard Temperature-Time Curve

PNOPOSED TIME-TEAMPE.fWTURE CURVE"

• "' :' -' .• " '" : 'k .. " • • • " " ", " "'. " 'I' . I

I I I!

r

R

Figure 1-1. Comparison between proposed temperature-time curve for column tests

at Underwriters' Laboratories [1-18] and current temperature-time

curve [1-1] (main figure reproduced from [1-18]).

NUREG-1547 18
History of Standard Temperature-Time Curve

TCMPCPATURC IN DOCCA. rAHXN'T"

-A

R1R

-A-1

- -I-IIA

a

FC

t_

7J

| i A |X

- .1 - -1.. 4-1 -~-.*~i--~'~ I - - I - - t -' - t - r - I - '? -r

= i I j ! i i I i. m I W R V w • a _ . . V -1

- -- 4- -1- -4

-~ iiIi~ ~6

te .4-I--.-'. I

L

I

I +! ! ! 11Its

I -f I i I f ! !* I

TrMwPc.RATUR im EccRecs cDCTIcRA~c

Figure 1-2. Standard temperature-time curve used In furnace tests, cooling curves, and curve representing the average temperature in a typical office

occupancy test (reproduced from [1-20]).

19 NUREG-1547
History of Standard Temperature-Time Curve

co

.~J.

C,,

2qN.

eqS

- N.

~0

Ii~

6 * 1 * I *

0

0

IV

0O

0o

0

0

00

04

0q

0

0

0

0

[sinoq] eous.npa aji~d

N 0

Figure 1-3. Fire endurance versus fr-e load (combustible content).

NUREG-1547 20
History of Standard Temperature-Time Curve

E

1500

I000

B00

600

400

200

0 10 20 30 40 50 60

Time (min)

Figure 1-4. Average Gas Temperatures in Compartments, where 60 (1⁄2) denotes 60

kg/mn2 fire load and 50 percent open wall (reproduced from [1-30]).

21 21 NUREG-1547
History of Standard Temperature-Time Curve

4)

02x4BIxodmaiL

Figure 1-5. Schematic of standard temperature-time curve and hypothetical curves

for room res.

NUREG-1547 22
History of Standard Temperature-Time Curve

0

CO,

to z

C%1

0O

0

o 0 0 0 0 o 0 0 0 0 00] Wd Cn4 einjejedwel

0

Figure 1-6. Theoretical results of Kanury and Holve [1-331 demonstrating the effect

of the form of the temperature-time curve on unexposed-surface

temperature.

23 NUREG-1547

APPENDIX - PART I

CHRONOLOGY OF FIRE TESTING PRIOR TO THE

ADOPTION OF THE STANDARD TEMPERATURE-TIME CURVE

FOR U.S. FIRE TESTS.

APPENDIX - PART 1. CHRONOLOGY OF FIRE TESTING PRIOR TO

THE ADOPTION OF THE STANDARD TEMPERATURE-TIME CURVE

FOR U.S. FIRE TESTS.

THIS SUMMARY IS LIMITED TO HISTORICAL REFERENCES TO THE EXPOSURE TEMPERATURES USED IN

FIRE-ENDURANCETESTING AND THE NBS WORK RELATING FIRE LOAD TO FIRE ENDURANCE. UNLESS

OTHERWISE NOTED, THE FOLLOWING WAS EXTRACTED FROM THE ACCOUNTS OF BABRAUSKAS AND

WILLIAMSON [AI-1,A1-2]' 7

1884 Column tests were conducted in Munich, Germany by Professor J. Bauschinger. The

column was heated in a horizontal orientaticn in a wood-fired furnace. Column temperature rather

than furnace temperature was measured. The column was heated to 300 °C, doused with water,

reheated to 400 *C or 500 *C, doused with water, and reheated to 600 °C and doused again.

1892-1895 Additional column testing was conducted in Hamburg by F.A. Meyer. The columns

were tested under load in a upright position in a gas-fired furnace. The columns were heated

symmetrically to 1200 °C to 1400 °C for up to 7 hours. A standard temperature curve was not

followed.

1890 Floor fire tests were conducted in Denver to compare three competing floor systems for use

in the Denver Equitable Building. Each system was built over a pit, loaded down, and subjected to

a coal fire contained in the pit. The fire exposure was maintained at an average temperature of

.9815 0C for 24 hours.

1891 Wall tests were conducted by Professor Bohme at Charlottenburg, Germany. The test

partition was erected between a bum room and observation room. A fixed mass of petroleumsoaked logs was used to fuel the fire. The test lasted 1 hour. The average gas temperature was

1000 *C

1891 Floor fire test was conducted in St Louis, MO, by architects for the Wainwright Building.

Following the initial heat-up period, the gas-fired furnace was maintained at approximately 815 °C

for about 6.5 hours. This is one of the first known gas-fired tests.

1893 Floor testing was initiated in Germany. These were burnout rather than standard tests.

Realistic furniture served as the fuel for the tests which were conducted in a Berlin building about

to be demolished. Temperatures ranged from 1000 °C to 1300 °C.

1894 A German floor fire test was conducted using wood, coal, and coke as fuel. The exposure

temperature was less than 700 *C.

17Letter-nurneral combinations in brackets with designation [Al-#] refer to literature references listed at

the end of this Appendix.

27 NUREG-1547
Appendix - Part 1

1895-1900 Wall tests continued at Charlottenburg, Germany. Only maximum exposure

temperatures were reported. They varied from 1000 °C to 1100 *C.

1896 Fourteen!' floor structures were tested by the New York Building Department using a wood

fire, 5-hr exposure maintaining an average temperature of 2000 °F (1093 °C) during the last 4

hours [Al-3]. In practice the maximum temperature ranged from 1975 OF to 2575 OF.

1896 Column testing was begun in the U.S. using a gas-fired furnace located at the Continental

Iron Works in Brooklyn. Temperatures (presumably column temperatures) reached up to 840 °C

and test periods ranged from 25 minutes to more than 2 hours.

1897 British Fire Protection Committee (Edwin 0. Sachs) originally built eight"9 brick chambers

for testing the fire endurance of building materials and systems of construction [A1-3]. This was

the first attempt to compile fire resistance data on materials and systems used in building construction [A1-4]. The exposure simulated a slow smoldering period followed by an increase to about

1093 °C [Al-I].

1899 British Fire Protection Committee began testing walls. Same temperature control as floor

test.

1899 New York Building Code included a the first fire-test standard in the U.S. for testing floors.

1901 New York Department of Buildings conducted a series of partition tests. Furnace

temperature was raised to 926 °C in 30 minutes and held at that level for the remainder of the test

(30 minutes).

1902 Floor testing resumed in New York and the average temperature was lowered from 1093 to

926 °C.

1902 Combined floor, partition, and column tests were conducted for the New York Building

Department. Tests were conducted for 4 hours at an average temperature of 930 °C.

1902 First permanent facilities in U.S. for testing the fire resistance of building components were

built by Professor Ira Woolson, Columbia University (2 large-scale furnaces; floor and wall).

Ryan reports [A1-4] that, at the turn of the century Professor Ira Woolson, Columbia University,

established a temperature-time curve for severe fires. He based this curve on data from building

fire investigationsincluding interviews of firefighters and comparison of materials gathered at fire

scenes with melting-point data. Other curves were developed at. this time.

1903 U.L. tested doors and windows in a gas-fired furnace. Furnace temperature was raised to

926 °C during the first 30 minutes and then held at that level for an additional 90 minutes.

1

8 Babrauskas and Williamson report [AI-1] that 16 tests were conducted in this series.

19 Babrauskas and Williamson report [Al-i] that 3 huts were built originally.

NUREG-1547 28
Appendix - Part 1

1904 Unloaded columns were tested in Chicago in a wood-fired furnace operating at 800 °C to

1000 *C. Tests lasted 3 hours.

1916 A description of plans for cooperative fire tests of columns was published in 1916. This

included the temperature-time curve that would be employed.

1917 The first standardized column tests were conducted at UL Chicago. Simon Ingberg, from

the National Bureau of Standards (NBS), was the director of the program. The National Board of

Fire Underwriters (NFBU) and the Factory Mutual companies participated in the program. More

than 100 columns were tested under load in a gas-fired furnace. Furnace temperature control was

standardized.

1917 ASTM C19 was issued2

" at the February 24, 1917 meeting of a conference made up of 11

organizations.

1918 ASTM C19 was adopted [A1-3,A1-4].2 Single curve was adopted in recognition of the

need for performance evaluation and economy in testing [A1-6].

1922 Burnout tests were conducted at NBS in specially constructed buildings [A1-6]. The

objective was to relate the adopted fire test exposure to the actual fire exposures in occupied buildings.

1926 NBS burnout tests representing office occupancies using wood and steel furniture, filing

cabinets were described in a brief report [A1-6]. The study found that the decrease following the

peak gas temperature in the room was much slower than the cool-down in the typical furnace test.

The relationship between the two exposure conditions would require further study. The tests were

conducted to obtain information on the intensity and duration of actual fires so that proper

exposure conditions could be used in fire endurance tests of construction materials.

1927 First report of Ingberg's correlation for office fire loads of 10 to 160 lb/ft2 [AI-6].

1928 The most complete report of NBS work was issued [A1-6]. It included a table correlating

fire load with fire endurance measured in a standard test. Ventilation was recognized to be

important but was not quantified. The furnace cooling process was recognized as part of the

thermal exposure. Ingberg recognized technical problems but assumed that matching the area of

2o Semantics has confused somewhat the date of ASTM C19. Babrauskas and Williamson [A1-2] report

that ASTM C19 (later renumbered E119) was issued on February 24, 1917 but later refer to the "1918

standard-. Ryan [A1-4] and Shorter [A1-3] report that ASTM E 119 was adopted in 1918. The current

ASTM E 119-88 standard [AI-5] states that C19 was first published in 1918 as C19-1917T, which, in

current ASTM notation, would indicate that the standard was approved in 1917 and fiwst published in

1918.

21 See previous footnote.

29 NUREG-1547
Appendix - Part 1

the average burnout curve with the area below the combined furnace heating and cooling curve

would yield equal severity exposures in the furnace test and actual fire.

192847 NBS expanded fire-load surveys to residences, schools, medical buildings, mercantile,

and manufacturing buildings [A1-6].

REFERENCES - APPENDIX - PART 1

[Al-] Babrauskas, V., and Williamson, R.B., The Historical Basis of Fire Resistance Testing -

Part I, Fire Technology, No. 3, pp. 184-194, August 1978.

[AI-2] Babrauskas, V., and Williamson, R.B., The Historical Basis of Fire Resistance Testing -

Part II, Fire Technology, No. 4, pp. 304-316, November 1978.

[AI-3] Shorter, G.W., The Fire Protection Engineer and Modem Building Design, Fire

Technology, National Fire Protection Association, pp. 206-213, August 1968.

[A1-41 Ryan, J.F., Perspective on Methods of Assessing Fire Hazards in Buildings, ASTM STP

502, American Society for Testing and Materials, pp. 11-23, 1972.

[AI-5] ASTM E 119-88, Standard Test Methods for Fire Tests of Building Construction Materials,

Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American Society for

Testing and Materials, Philadelphia, PA, 1994.

[AI-6] Robertson, A.F., and Gross, D., Fire Load, Fire Severity, and Fire Endurance, ASTM STP

464, American Society for Testing and Materials, pp. 3-29, 1970.

NUREG-1547 330
PART 2. HISTORY

TEMPERATURE-TIME

AND USE OF ALTERNATIVE

CURVES
LIST OF FIGURES - PART 2

Figure 2-1.

Figure 2-2.

Figure 2-3.

Figure 2-4.

Figure 2-5.

Figure 2-6.

Figure 2-7.

Figure 2-8.

Figure 2-9.

Figure 2-10.

Standard temperature-time curves used in various countries for testing of

building elements (reproduced from [2-5] and "incorrect" label added). See

footnote 22, section 2.2 .................. . . . .... .... . .. 49

Possible classification of building contents for fire severity and duration. The

straight lines indicate the length of fire based upon amounts of combustibles

involved. The curved lines indicate the severity expected for the various

occupancies, see table.. There is no direct relationship between the straight

and curved lines, but, for example, 10 pounds of combustibles per square foot

will produce a 90-minute fire in a "C" occupancy and a fire severity following

the time-temperature curve "C" might be expected (reproduced

from [2-7]) ................................................. 50

Ingberg's curves used by Corson (reproduced from [2-7]) .........

Temperatures in compartment fires (reproduced from [2-9D) .......

Suggested severity curves for fire tests (reproduced from [2-9]) ......

51

52

53

Relationship between the fire compartment temperature T, during the heating

phase (flame phase) and the time t according to Swedish Building Regulations

SBN 67 for some opening factors AhV/A, for a fire load mainly of the wood

type. T. = fire compartment temperature at t=0. The dashed curve is the

standard fire curve ... (reproduced from [2-17]) ................ 54

Calculated gas temperature-time curves for complete fire processes for

different fire loads q and opening factors Abl/At in fire compartment Type A

(standard fire cell), the fuel being of wooden type ... (reproduced from [2-17]) ........................................ 55

Predicted fires in offices (reproduced from [2-20D) .............. 56

The range of variation in average gas temperatures for residential room fires

(reproduced from [2-22D) ....................... .... .. 57

Comparison of fire exposure curve derived from room fire test 9 (BSMTO9)

and ASTM E 119 (reproduced from [2-22]) ................... 58

Figure 2-11. Mobil temperature-time curve ............................ 59

Figure 2-12. Hydrocarbon curves compared with ASTM E 119 curve .......... 60

NUREG-1547 32
History of Alternative Curves

SUMMARY - PART 2

Alternative temperature-time curves for evaluating the fire-endurance of building elements and

systems have been proposed since the 1950s, but it appears that only the alternative curves

developed for the hydrocarbon processing industry are widely used for fire-endurance testing.

Indeed, no record has been found to indicate that fire-endurance testing has been carried out IQ

Qualify a building element or assembly using any curve other than the ASTM E 119 standard

temperature-time curve, one of its foreign counterparts, or a hydrocarbon temperature-time

curve.

During the 1970s, an alternative curve was derived from a series of full-scale burnout experiments

involving residential recreation-room occupancies. That study, which was conducted at the U.S.

National Bureau of Standards - now the National Institute of Standards and Technology, included

tests of floor-ceiling assemblies in a furnace capable of following either the standard or alternative

curve. Although the results showed some dramatic differences in fire-endurance performance for

the two exposures, the study was judged too limited in scope to serve as a basis for changing fireendurance test standards. No follow-up work occurred. In the course of the present review, the

only record of fire-endurance testing of building elements or systems with an alternative curve

(other than a hydrocarbon curve) was that reported in the recreation-room study.

Since the 1960s mathematical modeling of room fires has been proposed as a means for

determining more realistic exposure conditions for evaluating fire endurance of building elements

and systems. In general, the fire-science literature suggests that, to.be practical, the development

of any new temperature-time curves for compartment applications will have to rely heavily on

mathematical modeling. This approach has been developed in Sweden to the extent that

temperature-time curves generated by mathematical room-fire models are sanctioned by Swedish

regulations for certain design calculations.

To date, a major weakness of room-fire models, however, lies in their inability to simulate

accurately burning rates under real-fire conditions. For example, much of the modeling work

aimed at creating alternative temperature-time exposures has been based on the burning

characteristics of cellulosic fuels. Also, such models only simulate the simplest type of room

ventilation- an opening to the outside environment. Development of credible temperature-time

curves for nuclear power plant (NPP) applications will require better understanding of the burning

behavior of NPP-specific fuel packages - for example, cable bundles - and more advanced

ventilation considerations.

33 NUREG-1547
History of Alternative Curves

2.1 OBJECTIVE

The purpose of Part 2 of this document is to review the literature concerning the history of the

development of alternatives to the ASTM E 119 temperature-time curve [2-1]" for assessing the

fire-endurance of construction elements and assemblies. In particular, this review attempts to

convey the methodology that was used to define non-standard alternative curves and the extent to

which these alternative curves are or were used.

2.2 INTRODUCTION

Part 1 of this document dealt with the technical basis for the ASTM E 119 standard temperaturetime curve. It is important to note that standard curves similar to the ASTM E 119 curve have

been and continue to be used in other countries to test building elements. These national standard

curves are shown in Fig. 2-1 [2 -5 ]fl. Curve 3 is the ASTM E 119 curve, used in North America,

and curve 2 is the ISO 834 curve [2-4r2, which is used in most of EuropeU. Note that the

differences among these curves are generally small. Indeed, Lie [2-5] argues that all the curves

likely are based on the same data as the 1916 curve [2-6] for column testing at Underwriters

Laboratories. Consequently, the concerns raised in Part I about the technical basis of the ASTM

E 119 curve are expected to apply to all of the curves displayed in Fig. 2-1.

Beginning in the 1950s, those concerns led to re-evaluations of what actually constituted a

reasonably accurate representationof real fire exposures for assessing fire-endurane requirements

of building elements. Impetus for the work included the availability of new analytical tools for

21 Numbers in brackets designated by [2AI] refer to literature references listed at the end of Part 2 of this

document.

22 Curve 7 is incorrect. The Japanese curve, which is defined in JIS A 1304 [2-2], has always been within

a few degrees of the E 119 curve and is not defined beyond 4 hours. The erroneous curve 7 may be a

result of an improper evaluation/extensionof the following equation which was developed at the Japanese

Building Research Institute to approximate the AS temperature-time curve (up to 4 hours) [T. Tanaka,

Building Research Institute, Japan, Personal Communication, 1995]:

T(t) = 1080 - 340 exp(-0.8t) - 130 exp(-5.0t) - 610 exp(-19t), where T [°C] and t [hours].

The JIS curve is used for testing all constructions except 3-story wood-house constructions. The latter are

tested using the ISO 834 curve [2-4].

2 Although the ISO 834 curve and the ASTM E 119 curve are similar, the type of thermocouples used are

different. Therefore, the difference in the actual thermal exposure to a test assembly between the two

methods is greater than the curves indicate.

' Countries participating in the European economic normalization will be required to use the ISO 834

curve.

NUREG-1547 34
History of Alternative Curves

studying compartment fires, much larger window areas in modem buildings, and a great increase

of non-cellulosic fuels in the buildings.

In general these efforts sought to identify realistic heating/cooling temperature-time curves for use

in controlling fire-test furnaces or serve as input to mathematical models designed to simulate the

response of fire barrier materials. It is these latter, more realistic curves, rather than the standard

curves in Fig. 2-1, that are the subject of Part 2.

2.3 ALTERNATIVE TEMPERATURE-TIME CURVES REPRESENTING

ROOM FIRES

2.3.1 Qualitative Curves or Curves Based on Qualitative Arguments

2.3.1.1 Corson's Curves

Using the work of Ingberg at NBS as a basis, Corson [2-7] in 1953 proposed a classification

system to relate fire duration to fire load in a given occupancy. The system consists of five

categories and is defined by the following table in conjunction with Fig. 2-2.

Fire Severity Expected by Occupancy [2-7_

Expcted

Fire Severity Category Typical Occupancy TemperatureTime Curve

(see Fig. 2-2)

Slight Well-arranged office, metal furniture, non-combustible Curve A

building.

Welding areas containing slight combustibles.

Non-combustible power house.

Non-combustible buildings, slight amount of combustible

occuancy.

Moderate Cotton and waste paper storage (baled) and well- arranged, non- Curve B

combustible building.

Paper making processes, non-combustible building.

Non-combustible institutional buildings with combustible

occupancy.

Moderately Severe Well-arranged combustible storage , e.g. wooden patterns, non- Curve C

combustible buildings.

Machine shop having non-combustible floors.

Severe Manufacturing areas, combustible products, non-combustible Curve D

building.

I Congested combustible storage areas, non-combustible building. 1 _1

35 NUREG-1547
History of Alternative Curves

ASTM Standard Fire Flammable liquids. Curve E

Exposure - Very Woodworking areas.

Severe Office, combustible furniture and buildings.

Paper working, printing, etc.

Furniture manufacturing and finishing.

Machine shop having combustible floors.

The straight lines in Fig. 2-2 relate fire load to fire-endurance time for each of the categories,

while the curved lines denote attendant "expected" temperature-time curves. Note that straight

line E is Ingberg's 1928 relationship between fire load and the fire endurance necessary to

withstand burnout of that load in an office occupancy, and curved line E is the standard ASTM

E 119 temperature-time curve. The justifications for the remaining curves (A through D),

however, remain unclear. The shapes chosen for these new temperature-time curves appear to

have been motivated by Ingberg's experimental burnout curves which are displayed in Fig. 2-3.

Regarding the latter curves, Corson states "It is interesting to note that in certain of these tests

a rapid rise of temperature occurred at the start, while in others the rapid rise was delayed. Also,

in most cases, temperatures at the start or very soon thereafter reached the intensity represented

by the standard ASTM "time-temperature curve." [2-7] This observation appears to be the sole

basis for the new curves.

Despite the weak scientific basis for the temperature-time curves A through D, they continue to

appear in the NFPA Fire Protection Handbook [2-8]. In the course of the current review, no

record was found of the actual use of these curves.

2.3.1.2 Seigel's Qualitative Curves

In 1967 Seigel [2-9] concluded that the standard temperature-time curve does not provide a

satisfactory representation of fire conditions in modern steel-framed buildings. The fire loads in

these buildings usually consist mainly of furnishings that can be easily ignited once a fire occurs.

He states that 'A well-ventilated fire in such a building will tend to reach its maximum

temperature quickly and will usually burn out in a short time because there will be no contributian

from the non-combustible structure." [2-9] Results of limited tests [2-9] conducted in room-size

compartments (Fig. 24) support this statement. He also concluded that a high-intensity shortduration fire may constitute a more severe exposure for a fire barrier than a low-intensity longduration exposure. "If the temperatures are high enough, explosive spalling of concrete and

destructive shrinkage of gypsum may result after a limited exposure. Other materials may also

be seriously deteriorated, so that they cannot perform their intended function for the duration of

the fire. " [2-9].

In view of these serious deficiencies, Seigel proposed that considerationbe given to replacing the

standard temperature-time curve with multiple curves such as those shown in Fig. 2-5. He noted

that further study would be necessary to quantify the curves to reflect maximum severity and

duration for a given fire load or occupancy.

NUREG-1547 36
History of Alternative Curves

2.3.2 Curves Determined by Mathematical Models

The next signifiant achievement toward a quantified understanding of room-fire intensities was

made by Kawagoe and co-workers [2-10 to 2-13] in Japan. By applying mathematical modeling

techniques to room-fire experimental results, they identified ventilation and wall properties as two

more important factors (in addition to Ingberg's fire load) that greatly influence the severity of

a fire in a room. They used their room-fire model to calculate the temperature-time curve

produced by ventilation-controlledfires up to the point of peak temperature (when fuel was gone)

and then assumed a temperature decay rate of 10 °C/min. They did not deal with fuel burning

rate (mass loss rate), but rather accounted for fuel effects solely by fire load.

With the mathematical model as a tool, it became possible, in principle, to calculate deterministic

temperature-time curves for specific room sizes, geometries, fuels, and ventilation conditions.

Here deterministic means that all the necessary variables are known and specified. Kawagoe

produced such curves, but then used Ingberg's equal-area hypothesis [2-14] to convert to an

equivalent standard temperature-time exposure for fire-endurance testing purposes.

In Sweden 6deen [2-15] developed a room-fire model similar to Kawagoe's except that burning

rate was included as an input variable. For a given amount of fuel, Odeen demonstrated that the

temporal development of the calculated gas temperature was influenced strongly by the burning

rate. Model temperature-time results compared well with experiments conducted in a

compartment with wood or kerosine fuels.

6deen's model was the basis for changes in Swedish Building Regulations, SBN 67 [2-16] in 1967

(effective January 1, 1968) that allowed choosing alternative temperature-time curves for some

situations [2-15]. If the fuel is primarily wood or other cellulosic materials and a ventilation

facto9• can be assigned to the compartment of interest, then Fig. 2-6, which was generated from

the model using the assumption that the bounding surfaces are brick or concrete block, can be

used to identify an alternative temperature-time heating curve for fire-protection design. The

regulations include procedures for determining fire duration and for appending a cooling curve

to the heating curve.

Subsequent work by Pettersson et al. [2-17], which is based on the work of Magnusson and

Thelandersson [2-18], produced a refined set of curves which superseded those in Fig. 2-6 in the

Swedish standard. Like the Kawagoe model, the Swedish model assumed a fully-developed,

ventilation-limited, wood-fueled fire in a compartment with bounding surfaces with specified

thermal properties, and then calculated the temperature-time curve for the contained gas. Figure

2-7 presents results for a room with boundary surfaces having "average" thermal properties.

Other curves were determined for a wide range of specific boundary surfaces. Each temperaturetime curve, therefore, represents results for a specified ventilation factor, boundary surface, and

fire load.

' For example, the ventilaton factor for a single vertical opening in a compartment is defined as

Ah'/A,, where A is the area of the opening, h is its height, and A, is the total internal surface area of the

fire compartment.

37 NUREG-1547
History of Alternative Curves

The Swedish regulations also allow for making a case for the effect other fuel properties/

geometries may have on the thermal exposure produced by a compartment fire. According to

Pettersson et al. [2-17], an alternative exposure defined as follows is permitted by the regulations:

"... the Swedish Building Code permits a design procedure which is finctionally better

substantiated and more rational [than the classification system based on the standard

temperature-time curve]. This is based on the gas temperature-time curve relating to the

complete fire process. This is determined in the individual case from heat or mass balance

equations or in some other way, consideration being given to the combustion characteristics

of the fire load, the ventilation characteristics of the fire compartment, and the thermal

properties of the structures enclosing the fire compartment and contained in this." [2-17]

Regardless of how it is determined, the resulting thermal exposure is used as input to a basic heattransfer model to determine the temperature-time response of key structural elements (usually steel

beams, columns, or rods ) which are protected by a fire barrier. If this structure is load bearing,

then consistent with the appropriate temperature-time curve either the minimum load bearing

capacity or the time of failure for the given load is determined by calculation or comparison with

available test results. The load bearing determination is usually omitted if the structure only

performs a separation function.

In the U.S., Babrauskas and Williamson [2-19] developed a more sophisticated mathematical

room-fire model that includes fuel-controlled burning and the ability to switch between fuel and

ventilation control. Although their model could be applied in a purely deterministic mode to

produce a specific temperature-time curve for a specific set of input data, they sought an

alternative approach that would produce a more general result. One such alternative is a series

of solutions based on a parameterized fuel load coupled with a limited range of worst-case

ventilation conditions. For a given compartment size, fuel load, and boundary thermal properties,

the ventilation condition is adjusted, within realistic limits, to produce the highest possible gas

temperature at each step in the calculation. Babrauskas and Williamson refer to this process as

ventilation pessimization. The result is a reduction of the dimensionality of the problem

(ventilation has been removed), and a more general, more conservative*, but less accurate

solution.

Figure 2-8 presents results of calculations applied to an office occupancy. Ventilation was

pessimized in the range 0.00 - 0.17 mt'n [2-20], corresponding to no-window and one-wall-missing

respectively. The fuel-percentilevalues shown in Fig. 2-8 represent fuel loads of 20, 35, 50, and

100 kg (wood equiv.)/Im.

26 In the course of studying Babrauskas' reports [2-19,2-20] for the present review, evidence emerged that

suggests that the ventilation pessimization can produce a room gas temperature that is near but less than

the absolute worst-possibletemperature. For example, see Fig. 8 in reference [2-19]. It appears that the

calculated result may be "path dependent;" that is, affected by the history of the calculation [Babrauskas,

V, personal communication, 1995.]. Further investigation of this issue is beyond the scope of the present

review.

NUREG-1547 338
History of Alternative Curves

Babrauskas [2-20] suggests that this approach could be used to develop a limited set of

temperature-time curves of the type Corson [2-7] was seeking. No evidence has been found,

however, that this suggestion was acted upon.

2.3.3 Curve Determined by Full-Scale Burnout Experiments - The NBS Recreation Room

Curve

In the mid to late 1970s, Fang and Breese [2-22] at the U.S. National Bureau of Standards (NBS)

(now NIST) conducted a study to establish rational test procedures for testing the fire endurance

of residential floor-ceiling assemblies. This work was sponsored by the U.S. Department of

Housing and Urban Development (HUD) and was motivated by the concern that the ASTM E 119

temperature-time exposure, which had been adopted 60 years earlier, was not representative of

fires in modern residences having furnishings made from synthetic materials and larger window

areas.

The study focused on fire behavior in typical recreation (family) rooms in single-family

residences. The aim was to characterize the severity of fires originating in these rooms and

develop a set of exposure conditions for testing the fire endurance of floor-ceiling systems.

A total of 16 burnout experiments were conducted in two instrumented test rooms, 3.3 m x 3.3

m x 2.6 m (high), and 3.3 m x 4.9 m x 2.4 m (high), furnished with modern household furniture,

and lined with typical wall and ceiling finish materials. The 16 experiments represented a range

of fire loads, ventilation conditions, and finish materials.

The chosen fire loads, materials, and room geometries were based on a survey of 70 homes

located in the Washington, DC area. Total fire loads (movable plus structural) of 21, 28, and 42

kg/rmr were selected since the survey demonstrated that they represented low, average, and high

loads, respectively. Four ventilation conditions were examined: door open with and without

forced ventilation supplied to the room in a fashion simulating a forced air heating/cooling system,

and door closed with and without forced ventilation.

Figure 2-9 displays the range of average gas temperatures measured in the upper region of the test

room for a variety of fuel loadings and configurations. All of these tests were conducted with the

door open. The peak temperatures were about the same, but the times to reach the peak varied

considerably. Moreover, the experimental temperature-time curves were notably different from

the ASTM E 119 curve also shown in the figure. The experimental fires became hotter faster than

the ASTM E 119 exposure and began to decay within 15 to 30 minutes after ignition.

Fang and Breese considered Test BSMT09 as representing "standard" test conditions; namely, a

3.3 m x 3.3 m recreation room with plywood paneling walls, gypsum board ceiling, open door,

and 23 kg/i 2 movable loading of household furnishings. They fit a polynomial to 7 points on

temperature-time curve from Test BSMT09 to produce the "derived curve" shown in Fig. 2-10

and judged the result to be an approximate average over the range of test conditions employed in

their study. They concluded that

39 NUREG-1547
History of Alternative Curves

"The rate of development and the intensity of real fires involving the burning of typical

furniture and interior linings in a room during the first 20 minutes may be significantly

greater than those defined by the ASTM E 119 time-temperature curve. A more realistic time

temperature curve for residential occupancies is [the "derived curve" in Fig. 2-10.]" [2-22]

Subsequent work at NBS [2-23] included another series of burnout tests under the standard

conditions defined above, but with a set of selected residential floor/ceiling assemblies. These

same systems were later tested by Fang [2-24] in a gas-fired furnace under both an ASTM E 119

exposure and the newly-derivedtemperature-timecurve. This allowed comparisons between the

two types of furnace exposure and between the furnace and room burnout tests.

Among Fang's conclusions were

"The wood joist floors exposed to the newly developed fire conditions in the gas-firedfuirnace

had a shorter time to failure compared with the earlier residential room fire tests on the same

floor constructions. This was due primarily to the increased burning rates of the combustible

materials in the test structure with the excess air inside the test furnace. " [2-24]

"Individual test assemblies resisted flame penetration in the furnace fire tests for a time

approximately 40 percent shorter when tested under the newly developed time-temperature

curve as compared with the ASTM fire exposure." [2-24]

They recommended that the newly developed temperature-time curve should be used in a new

ASTM fire endurance test for rating residential floor constructions which require endurance times

of less than one hour. This recommendation, however, was rejected on the basis that the roomburnout experiments were too limited in scope 27.

2.4 TEMPERATURE-TIME CURVES REPRESENTING OPEN-AIR

PETROCHEMICAL CURVES

In the 1970s, the hydrocarbon processing industry (HPI) began working on a fire test that

realistically represented direct exposure to a petroleum fire (combusti'ble/flammable liquids and

gases). This was motivated by the need to protect fire-exposed structural steel for up to I hour

and cable trays for up to 30 minutes. The tray protection was needed to allow for orderly

shutdown of equipment [2-26].

Warren and Corona [2-26] found that the ASTM E 119 test was an unsuitable representation of

direct petroleum fire exposure because its temperature rise was too slow and the heat flux to the

test assembly is mainly radiative. They report that petroleum fires rise rapidly to about 1100-1200

°C (about 2000-2200 OF), have a high convective flux component, and often totally involve

27Fang, J.B., personal communication, 1994.

NUREG-1547 40
History of Alternative Curves

"5. Bring together 3 and 4 above to derive a design fire for offshore structures and to define

its characteristics, in particular with regard to heat flux and temperature .... [2-29]

This analysis identified two salient fire scenarios:

"a) A spill on the deck of the structure. This would be a fairly small fire, perhaps involving

a minimum of 0.05 m3 of fuel. Such fires occur fairly frequently, with a probability of

occurrence of 10" to 10 2 per installation year. It would consist of fuels such as aviation fuel

crude oil, methanol or natural gas liquids which are stored on board offshore structures in

varying quantities. Temperatures around 800 °C to 1100 'C were indicated from the data

examined and maximum total heat fluxes of around 150 kWI/ 2. These values would be

reached in, typically, two minutes.

b.) A spill from a tanker on the sea. This could involve 120,000 m3 of almost any fuel,

carried by a tanker and therefore include LNG (liquid natural gas]. The probability of such

an occurrence involving LiVG was not estimated, but would be significantly lower than 101

per installation year. However the density of sea traffic in the North Sea is such that it was

considered to be appropriate to include an LWG spill. If LNG is considered a possible fuel

then temperatures around 1000-1200 OC were indicated by the data examined and maximum

total heat fluxes of around 220 kW/M2. Again these values would be achieved within two

minutes although maintained for only 10 minutes." [2-28]

FRS chose to propose the following design fire based on a combination of both of the above

scenarios:

SFRS DESIGN FIRE

Temperature Total Heat Flux Convective

I_ I Heat Flux

1200 "C within 2 min. and maintnined 220 kW/m2 achieved within 2 min. 40 kW/m2

for another 10 1m. and maintained for 10 min.

Drop to 1100 OC at 12 min. and Drop to 150 kW/ 2 at12min. and 40 kW/&2

maintain for the duration of the fire maintain for the duration of the fire

This FRS design curve is plotted in Fig. 2-12.

Meanwhile, in the United States ASTM Subcommittee E05.11 was working on a similar task to

standardize hydrocarbon-fire simulations which would lead eventually to ASTM E 1529 "Standanr

Test Methods for Determining Effects of Large Hydrocarbon Pool Fires on Structural Members

and Assemblies" [2-30]. It was recognized that a hydrocarbon fire gets hotter faster (158 kW/m2

at 5 minutes) than the ASTM E 119 standard exposure (35 kW/&i at 5 minutes and 118 kW/m&

at 60 minutes).

NUREG-1547 42
History of Alternative Curves

The approach was to rely on engineering judgement to select a single set of test conditions to

represent a reasonable worst case for hydrocarbon-production installations for facility-design

purposes. Consequently, with respect to radiant heat transfer from the fire to an object within or

adjacent to the fire, worst case conditions would prevail when the view factor between the flame

and object was 1.0, the flame emissivity was 1.0, and time continuity (fraction of time the object

is located in a flame that generally fluctuates greatly in a given space) was 100 percent.

Bader [2-31] measured heat fluxes in large pool fires and modeled the results. The maximum

time-integrated total heat flux measured with a slug calorimeter was 150 kW/m2 . He found that

for modeling purposes a blackbody temperature of 1010 °C (corresponding to a 154 kW/m2

blackbody radiant flux) satisfactorily accounted for the combined effect of radiation and

convection from a real pool fire. Other studies [2-32 to 2-39] covering a range of fuels and fire

sizes report either radiant or total cold wall's heat fluxes to engulfed objects, or objects on the

perimeter of the fire, in the range 158-161 kW/m2 . Typically, the test object represented a

dangerous and/or high-value item such as a weapon or weapon container [2-32,2-33], hazardous

material container [2-34], or aircraft fuselage section [2-35].

On the bases of these experimental data and analyses, a total heat flux of 158 kW/m2 was chosen

as being representative of a reasonable worst case exposure. Although convective heat flux is not

specifically called out in the standard, calculations indicate it should be about 10 percent of the

total flux to a vertical column. This is consistent with the 9 to I ratio observed by Mansfield [2-

36] for large pool fires.

In addition to heat flux, the standard specifies the temperature of the furnace gases;' that produce

much of tie exposure heat flux. The temperature range of 1010-1180 °C was selected for two

reasons a.) actual pool fires range from 927-1260 °C in the luminous region, and the specified

range is in middle of this wider range, and b.) specifying a range allows the test operator

flexibility for attaining the required heat flux.

I ASTM E 1529 FIRE EXPOSURE

Temperature Total Cold-Wail Heat Flux Convective

I IHeat Flux

At least 815 °C after the first 3 min. Not specified during first 3 min. Not

Specified

Between 1010-1180 °C at5 min. and 158 ± 8kW/ 2 at5 rin. and Not

maintain for the duration of the fire. maintain for the duration of the fire specified

The ASTM E 1529 temperature-time curve is plotted in Fig. 2-12.

28 The heat transfer rate to an object depends on the object's temperature. Cold wall heat fluxes usually

refer to fluxes measured with water-cooled gauges located at the surface of the object.

" More accurately, the average temperature of furnace thermocouples, located in the furnace gases about

150 mm from the exposed surface of the specimen, is specified.

43 N3NUREG-1547
History of Alternative Curves

In 1989 Underwriters Laboratories Inc. issued a similar standard, UL 1709 [2-40], for testing

protection materials for structural steel. The test exposure is listed in the table below and the

temperature-time curve is plotted in Fig. 2-12. Note the temperature band permitted by UL 1709

is almost identical to that allowed by ASTM E 1529, but the attendant cold-wall heat fluxes differ

by about 30 percent. Apparently the higher heat flux specified in UL 1709 reflects the actual

furnace-specific characteristics of the UL furnace (wall/flame emissivity, etc.) when operating at

the specified temperature.

UL 1709 FIRE EXPOSURE

Temperature Total Cold-Wal Heat Flux Convective

I I Heat Flux

1093 ± III "C within 5 mitnutes and 204 ±- 16 kW/m: within 5 minutes Not

maintined for the duration of the test. and maintained for the duration of Specified

the test.

2.5 CONCLUSIONS

1. Hydrocarbon temperature-time curves appear to be the only alternative curves widely used

for fire-endurance testing.

2. No record has been found to indicate that fire-endurance testing has been conducted toAUali•

a building element or assembly using any curve other than the standard temperature-time

curve or a hydrocarbon curve. In the course of the present review, the only record of fireendurance testing of building elements or systems with an alternative curve was that found

in the NBS recreation-room study.

3. The literature suggests that the development of any new temperature-time curves for

compartment applications likely will rely heavily on mathematical modeling because of the

need to consider a wide variety of configurations.

4. Alternative temperature-time curves generated by mathematical room-fire models are

sanctioned by Swedish regulations for design calculations.

5. To date, a major weakness of roon-fire models lies in their inability to simulate accurately

burning rates under real-fire conditions. For example, much of the modeling work aimed at

creating alternative temperature-time exposures has been based on the burning characteistics

of cellulosic fuels. Also, such models only simulate the simplest type of room ventilation - an opening to the outside environment. Development of credible temperature-time curves for

nuclear power plant (NPP) applications will require better understanding of the burning

behavior of NPP-specific fuel packages - for example, cable bundles - and more advanced

ventilation considerations.

NUREG-1547 44
History of Alternative Curves

2.6 REFERENCES

[2-1] ASTM E 119-88, Standard Test Methods for Fire Tests of Building Construction

Materials, Annual Book of ASTM Standards, Sec. 4, Construction, Vol. 4.07, American

Society for Testing and Materials, Philadelphia, PA, 1994.

[2-2] Japanese Industrial Standard Method, Fire Resistance Test for Structural Parts of

Buildings, JIS A 1304-1975, (English translation), Japanese Standards Association,

Tokyo, Japan, 1975.

[2-3] Not used.

[2-4] International Standard 834, Fire-resistance tests - Elements in Building Construction,

Ref. No.ISO 834-1975 (E), November, 1975.

[2-5] Lie, T.T., Fire and Buildings, Applied Science Publishers, Ltd. London, 1972.

[2-6] Fire Tests of Building Columns, NFPA Quarterly, Vol. 9, No. 3, pp. 253-260, January

1916.

[2-7] Corson, R.C., The Significance of Fire Loading, NFPA Quarterly, Vol. 47, No. 1,

pp. 65-72, July 1953.

[2-8] Fire Protection Handbook, Seventeenth Edition, Arthur E. Cote, Editor-in Chief,

National Fire Protection Association, Quincy MA, p. 6-78, July 1991.

[2-9] Seigel, L.G., The Severity of Fires in Steel Framed Buildings, Behavior of Structural

Steel in Fire Symposium No. 2, Proceedings of a Symposium held at the Fire Research

Station, Boreham Wood, Herts, on January 24, 1967, Her Majesty's Stationery Office,

pp. 57-63, 1968.

[2-10] Kawagoe, K., Fire Behavior in Rooms, Report No. 27, Building Research Institute,

Japan, September 1958.

[2-11] Kawagoe, K. and Sekine, T, Estimation of Fire Temperature-Time Curve in Rooms,

Occasional Report No. 11, Building Research Institute, Japan, June 1963.

[2-12] Kawagoe, K. and Sekine, T., Estimation of Fire Temperature-Time Curve in Rooms,

Second Report, Occasional Report No. 17, Building Research Institute, Japan, March

1964.

[2-131 Kawagoe, K., Charts for Estimating the Equivalent Fire Duration on the Standard

Temperature-Time Curve, Research Paper No. 45, Building Research Institute, January

1971.

45 NUREG-1547
History of Alternative Curves

[2-14] Ingberg, S.H., Tests of the Severity of Building Fires, NFPA Quarterly, Vol 22, 1, pp.

43-61, July 1928.

[2-15] Odeen, K., Standard Fire Endurance Tests - Discussion, Criticism, and Alternative, Fire

Test Performance, ASTM STP 464, American Society for Testing and Materials, pp. 30-

56, 1970.

[2-16] Statens Planverk, Svensk Byggnorm 67, (The National Board of Urban Planning,

Regulations, Recommendations, and Directions Concerning the Swedish Building

Ordinance), Stockholm, 1967.

[2-17] Pettersson, 0., Magnusson, S., and Thor, J., Fire Engineering Design of Steel

Structures, Publication 50, Swedish Institute of Steel Construction, Stockholm, Sweden,

1976.

[2-18] Magnusson, S.E. and Thelandersson, S., Temperature-TimeCurves of Complete Process

of Fire Development - Theoretical Study of Wood Fires in Enclosed Spaces, Acta

Polytechnica Scandinavica, Civil Engineering and Construction Series, No. 65,

Stockholm, 1970.

[2-19] Babrauskas, V., and Williamson, R.B., Post-flashover Compartment Fires -

Applications of a Theoretical Model, Fire and Materials, Vol. 3, No. 1, pp. 1-7, 1979.

[2-20] Babrauskas, V., Fire Endurance in Buildings, Report N4o. UCB FRG 76-10, Ph.D.

dissertation, University of California, Berkeley, November 1976.

[2-21] Not used.

[2-22] Fang, J.B., and Breese, J.N., Fire Development in Residential Basement Rooms, NBSIR

80-2120, National Bureau of Standards, U.S. Department of Commerce, Washington,

DC, October 1980.

[2-23] Fang, J.B., Fire Performance of Selected Residential Floor Constructions Under Burnout

Conditions, NBSIR 80-2134, National Bureau of Standards, U.S. Department of

Commerce, Washington, DC, December 1980.

[2-24] Fang, J.B., Fire Endurance Tests of Selected Residential Floor Construction, NBSIR 82-

2488, National Bureau of Standards, U.S. Department of Commerce, Washington, DC,

April 1982.

[2-25] Not used.

[2-26] Warren, J.H., and Corona, A.A., This Method Tests Protective Coatings, Hydrocarbon

Processing, Vol. 54, No. 1, pp. 121-122, 126-128, 130, 132, 137, January 1975.

NUREG-1547 46
History of Alternative Curves

[2-27] Schiff, M., Hydrocarbon Fire Standard for OffShore Installations, Fire Surveyor, Vol.

13, No. 6, pp. 4-7, December 1984.

[2-28] Schiff, M., Hydrocarbon Fire Standard for OffShore Installations, Building Research

Establishment Paper 164/165, K3, IP 15/84, Building Research Establishment,

Department of the Environment, UK, September 1984.

[2-29] Shipp, M.P., A Hydrocarbon Standard - An Assessment of Existing Information, Fire

Research Station Report to Department of Energy OT/R/8294, January 1983.

[2-30] ASTM E 1529-93, Standard Test Methods for Determining Effects of Large

Hydrocarbon Pool Fires on Structural Members and Assemblies, Annual Book of ASTM

Standards, Sec. 4, Construction, Vol. 4.07, American Society for Testing and Materials,

Philadelphia, PA, 1993.

[2-31] Bader, B.E., SC-R-64-1366AHeat Transfer in Liquid Hydrocarbon Fuel Fires, Chemical

Engineering Progress Symposium Series, Sandia Corp., Albuquerque, NM, Vol 61, No.

56, 1965.

[2-32] Russell, L.H., and Canfield, J.A., Experimental Measurement of Heat Transfer to a

Cylinder Immersed in a Large Aviation-Fuel Fire, Journal of Heat Transfer, August

1973.

[2-33] Cornell, G., Fire Technology Programs at the Naval Air Warfare Center Weapons

Division, China Lake, Proceedings of International Conference on CQmposites

Engineering (ICCE/1), pp. 103-104, August 1994.

[2-34] Bainbridge, B.L., and Keltner, N.R., Heat Transfer to Large Objects in Large Pool Fires,

Journal of Hazardous Materials, Vol. 20, pp. 21-40, 1988.

[2-35] Brown, L.J., Jr., Cabin Hazards from a Large External Fuel Fire Adjacent to an Aircraft

Fuselage, Report No. FAA-RD-79-65,August 1979 Final Report, Prepared for U.S. Dept.

of Transportation, Federal Aviation Administration, Systems Research & Development

Service, Washington, DC.

[2-36] Mansfield, J.A., Tank Car Fire Program, Report No. DOT-NASA-3-1, May 1974 Interim

Report, Prepared for Dept. of Transportation, Federal Railroad Administration,

Washington, DC.

[2-37] Henshaw, J., et al, Fire Protective Materials Application Program, Vol. 1, Final Report,

AVCO Systems Div. for NASA-Ames, March 1972.

[2-38] Gregory, J.J, Keltner, N.R.,and Mata, R., Thermal Measurements in Large Pool Fires, Heat

and Mass Transfer in Fire, ASME-HTD Vol. 73,1987 and Journal of Heat Transfer, Vol.

111, pp. 446-454, May 1989.

47 NUREG-1547
History of Alternative Curves

[2-39] Schneider, M.E., and Kent, L.A., Measurements of Gas Velocities and Temperatures in a

Large Open Pool Fire, Heat and Mass Transfer in Fire, ASME-HTD Vol. 73,1987 and Fire

Technology, Vol. 25, No. 1, February 1989.

[2-40] UL 1709, Standard for Rapid Rise Fire Tests of Protection Materials for Structural Steel,

First Edition, Underwriters Laboratories Inc., February 22, 1989.

NUREG-1547 48
History of Alternative Curves

1300

1200

1100

0.'•1000

900

L800

w 7 0 0

600

*500=

400

Figure 2-1.

AUSTRALIA

GREAT BRITAIN

NEW ZEALAND

[BELGIUM

DENMARK

FINLAND

FRANCE

NETHERLANDSNORWAY

SWEDEN

WEST GERMANY

2 3 4

Ua

w

w

cc

DURATION, h

Standard temperature-time curves used in various countries for testing of

building elements (reproduced from [2-5] and "incorrect" label added). See

footnote 22, section 2.2.

49 NUREG-1547
History of Alternative Curves

M~y W~U 'Ll *bS /flO=%=

S CV 94 - -

\ FAH -a04

'1 N

i

I - &'~ ' - .4- .4-4. -~ 4~4. * I -- ý

\ A\ -__ J'

N *1

I - I 6- 4 h I I - i - - i i I i 41 It Ik i A

N '*'N'N \

\I \ K .--

Eli ~

S23 a .U

Figure 2-2. Possible classification of building contents for fire severity and duration. The

straight lines indicate the length of fire based upon amounts of combustibles

involved. The curved lines indicate the severity expected for the various

occupancies, see table.. There is no direct relationship between the straight and

curved lines, but, for example, 10 pounds of combustibles per square foot will

produce a 90-minute fire in a "C" occupancy and a fire severity following the

time-temperature curve "C" might be expected (reproduced from [2-71).

NUREG-1547 50
History of Alternative Curves

z! 10

C 0-

1 a.

.24J

*-a

.~I

-DIU

.3!

*C Z

CIE

0 %0 0 to

1!4u-r)40J Sawba(] ui aijflDJdwa.

Figure 2-3. Ingberg's curves used by Corson (reproduced from [2-7]).

51 NUREG-1547
History of Alternative Curves

0 0 0> 0 0 0

C4 0

o 0 C 00 0

I -u 

-

-o

'C 0

110

'S

1/)

1<I

'S

I

I

'S

'S

* I

I

'S

p. p.

~o *= .= -

- %% %%~

~ ~O D

'S

~ C~4 -

N%

-C

cl4

LO 0 0 0 0 0

U-C 0 C> €-• a

P V c C

0

Figure 24. Temperatures in compartment fires (reproduced from [2-9]).

NUREG-1547 52
History of Alternative Curves

TEMPI

TIME-.-

Figure 2-5. Suggested severity curves for fire tests (reproduced from [2-91).

53 3 NLTREG-1547

History of Alternative Curves

1z <~ ck

0

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Regulations SBN 67 for some opening factors Ah"/A, for a fire load mainly of

the wood type. T. = fire compartment temperature at tO--. The dashed curve is

the standard fire curve ... (reproduced from 12-17]).

NUREG-1547 54
History of Alternative Curves

Figure 2-7. Calculated gas temperature-timecurves for complete fire processes for different

fire loads q and opening factors Ah"/A, in fire compartment Type A (standard

fire cell), the fuel being of wooden type ... (reproduced from [2-17]).

55 NUREG-1547
History of Alternative Curves

FUEL LOAD PEROSNTILE

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FUEL LOAD DATA FROM ;tULVER

VZHTILATION t PESSIMIZED 40

0

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Figure 2-8. Predicted fires in offices (reproduced from [2-20]).

NUREG-1547 56
pomnpajdaJ) tuoij -([Z-ZJ 

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History of Alternative Curves

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Figure 2-10. Comparison of fire exposure curve derived from room fire test 9 (BSMT09) and

ASTM E 119 (reproduced from [2-22]).

NUREG-1547 558
History of Alternative Curves

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59 NUREG-1547
History of Alternative Curves

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NUREG-1547 660
PART 3. FEASIBILITY OF DEVELOPING AND

IMPLEMENTING NPP-SPECIFIC DESCRIPTIONS OF

FIRE ENVIRONMENTS FOR USE IN EVALUATING THE

FIRE RESISTANCE OF FIRE BARRIERS
LIST OF FIGURES - PART 3

Figure 3-1.

Figure 3-2.

Figure 3-3.

Figure 34.

NPP compartment fire involving fully-developed burning of an extensive dense

array of cable trays . ................................... 99

NPP compartment fire involving burning of a combustible/flammable liquid. .... V............................................. 100

NPP compartment rwe involving burning of a combustible/flammable liquid

and direct fire exposure of unprotected and wrap-assembly-protected cable

trays and of the compartment ceiling structure ................. 101

Multi-room NPP compartment fire involving fully-developed burning of an

extensive dense array of cable trays and a combination of different types of

ventilation configurations ................................ 102

LIST OF TABLES - PART 3

Table 3-1. A Preliminary Combustible Load Summary of the Spaces at the Watts Bar

NPP Having the Largest Density of Combustibles ............... 98

NUREG-1547 662
SUMMARY - PART 3

Part 3 of this document assesses the feasibility of developing and implementing Nuclear Power

Plant (NPP)-specific descriptions of fire environments and alternative ASTM E 119-type

temperature-time fire curves for evaluating the fire resistance of fire barriers, including structural

building components. Advances in fire science over the past 40 years have offered the potential

for developing technically-sound alternative curves for use in areas where fire exposures can be

expected to be significantly different than the standard temperature-time exposure. For example,

during the 1970s and 1980s, several countries, including the United States, developed and

implemented technically-sound alternative curves for testing fire barriers that might be subjected

to open-air hydrocarbon pool fires.

The NRC staff has initiated the current effort to investigate the feasibility of developing alternative

temperature-time curves for the qualification of fire barriers used to protect cabling and

equipment necessary to achieve safe shutdown on the basis of realistic fire hazards found in NPPs.

The present work 1) begins with a discussion of the problem of defining the fire exposure with

which to evaluate fire barrier performance, 2) identifies NPP-specific arrays of combustibles and

compartment-configuration characteristics that critically influence fire severity, and 3) proposes

a methodology for evaluating NPP-specific fire-barrier performance. This methodology would

incorporate: an advanced, practical, mathematical compartment model to simulate NPP-specific

fire envirorxnents; ASTM E 119-type full-scale furnace tests; and additional test methods, to be

developed and applied as required, that would simulate barrier response to exposures of sustained

direct flame impingement. Key elements of the model are outlined.

3.1 INTRODUCTION: THE OBJECTIVE AND AN OVERVIEW

One aspect of fire safety in NPPs has to do with the performance of fire barriers designed

explicitly to protect components, equipment, etc., on the protected side of the barrier, from

potential threatening fire environments on the fire-exposed side. It is the objective of this work

to propose a methodology of evaluating NPP fire barrier performance that takes into account NPPspecific fire environments.

Advances in fire science over the past 40 years have offered the potential for developing

technically-sound alternative curves for use in areas where fire exposures can be expected to be

significantly different than the standard temperature-time exposure (see Part 2). For example,

during the 1970s and 1980s, several countries, including the United States, developed and

implemented technically-sound alternative curves, for testing fire barriers that might be subjected

to open-air hydrocarbon pool fires [3-1,3-2]29 (see Part 2, section 2.4).

9 Numbers in brackets with designation [3-f] refer to literature references listed at the end of Part 3 of this

document.

63 NUREG-1547
NPP-Specific Descriptions of Fire Environments

Part 3 of this document will assess the feasibility of developing and implementing NPP-specific

descriptions of fire environments for use in evaluating the fire resistance of fire barriers. These

include structural barriers, useful in isolating a compartment of fire origin from adjacent spaces,

and wrap assemblies [3-3], used to isolate and protect plant equipment, cables, etc., within a

compartment of fire origin, from the effects of exposure to the fire environment.

A discussion on the nature of fire barrier exposure to the compartment fire environment will be

presented. This will distinguish between direct exposure to the most extreme zones of the fire

environment, e.g., direct, sustained exposure of the barrier to the flame, and indirect exposure,

where the fire barrier is mainly exposed to the average properties of the overall fire environment

The problem of evaluating the integrity of fire barriers to both kinds of threats will be addressed.

The problem of defining the fire exposure with which to evaluate fire barrier performance is

discussed. It is shown that the use of ASTM E 119-type test methods, where the standard ASTM

E 119 fire [34] is replaced by alternative temperature-time furnace fires and where the alternathi

fires would be deduced from reliable fire model simulations, has promise.

A methodology for evaluating the fire performance of NPP fire barriers will be presented. As

will be seen, this relies on a combined experimental and analytic approach that involves the

Bounding-Temperature Principle (i.e., if the temperature-time curve of one fire environment

bounds that of another, then, relative to the threat to structural integrity of a NPP fire barrier, the

bounding-curve environment is the more severe.) Experiments would involve ASTM E 119-type

tests, to address indirect exposure threats, and other tests that would be devised to deal with direct

exposure threats. Analysis would involve compartment fire modeling methods. Computer

simulations would be carried out with a new, advanced, special-purpose, zone-type fire model

developed to include features particularly relevant to simulating fire environments that threaten

NPP fire barriers, from the point of view of both direct and indirect fire exposure.

Based on a review of the literature of NPP-specific combustibles and previously-developed NPPspecific fire models, special features required of the new fire model will be identified. These

include: the simulation ofd ly-developed burning of extensive dense arrays of cable trays (i.e.,

all exposed surfaces of a combustible cable are supplying fuel (losing mass) due to either heating

by the fire environment or surface combustion), both under fuel-controlled and ventilationcontrolled conditions; the simulation of combustible/flammable liquid pool fires; the simulation

of the fire environment in multi-room facilities (at least two adjacent spaces); and advanced means

of modeling ventilation and radiation-heat-transfer-related phenomena. It is proposed that a new

special-purpose model with these features be developed as a customized advanced version of an

existing, two-layer, multi-room, zone-type fire model.

The new model would be used to simulate a wide variety of potential fire scenarios in rooms of

fire origin of selected NPPs. The simulations would lead to new insights on the characteristics

of real, fire-barrier-threatening, NPP fire environments, including both indirect and direct

exposures. Based on applications of the Bounding-Temperature Principle, the simulated fire

scenarios would lead to a series of NPP-specific test fire curves covering a wide range of NPPtype fire severities. An experimental study on available ASTM E 119-type test furnaces would

NUREG-1547 64
NPP-Specific Descriptions of Fire Environments

be carried out to establish that these new test fire curves (instead of the standard ASTM E 119

fire curve) can be used in ASTM E 119-type barrier rating tests. Then ASTM E 119-type tests

using the new NPP-specific test fire curves would be established as the method of evaluating the

fire performance of NPP fire barriers against indirect threats.

Based on all the above, six specific tasks are proposed that would lead to a reliable methodology

for evaluating NPP fire barrier performance.

3.2 TWO TYPES OF FIRE BARRIERS: STRUCTURAL BARRIERS AND

WRAP ASSEMBLIES

In NPPs, two types of fire barriers are of interest, structural fire barriers and the fire barrier wrap

assemblies.

The structural-type fire barrier is common to all types of building facilities, including NPPs.

Reference here is to certain wall/ceiling/floor elements, designed to define and separate one

compartment or room of a facility from another. Such a barrier may or may not be load-bearixg.

It would be designated as a fire barrier in the sense that, for a specified time interval, a successful

design would protect the contents of a compartment, on one side of the barrier, from the effects

of a fire in the compartment on the other side of the barrier.

To the extent that the collapse of isolated load-bearing columns and beams would typically lead

to the failure of wall/ceiling/floor fire barriers, such columns and beams are also considered to

be structural fire barriers. As such, successful designs of these latter structural elements must also

be able to sustain exposure to an appropriate compartment fire environment, for a design-specified

time interval, without threat of collapse.

The second type of fire barrier, the wrap assembly fire barrier, is particularly important in NPPs.

This includes barriers that are more-or-less fully contained within a NPP compartment of potential

fire origin. Such barriers enclose and separate protected equipment from the rest of the compartment. They must perform their protecting function when installed under realistic field conditions

and while exposed to the compartment fire environment for a design-specified time interval.

Thus, in the event that a fire develops within the compartment, a successful fire-barrier wrapassembly design would be one which will maintain its integrity to the extent that the environment

within the barrier enclosure will not be so severe as to lead to fire damage of the protected

equipment. In NPPs, a typical wrap-assembly fire barrier is the barrier enclosure that isolates one

redundant critical set of cables in a NPP compartment of interest from the other set, the latter set

being installed without barrier protection, somewhere within the compartment [3-5].

65 NUREG-1547
NPP-Specific Descriptions of Fire Environments

3.3 ASTM E 119 AND THE ASTM E 119-TYPE APPROACH TO THE

EVALUATION OF STRUCTURAL-TYPE BARRIER PERFORMANCE

The ASTM E 119 test method involves three major elements: a facility to carry out furnace testing

of full-scale test specimens, a standard temperature-time curve that specifies the furnace fire

environment, and specified acceptance criteria for establishing structural integrity of the exposed

specimens. (For some types of structural elements, acceptance criteria include the ability of a test

specimen to sustain a fire and hose stream test. This aspect of the ASTM E 119 test will not be

addressed further.) In general, a clear correspondence between the standard fire and real

threatening fires is unavailable (see Part 1). However, it should be possible to purposefully

develop alternative, achievable, furnace-fire environments, for use in ASTM E 119-type test

methods, that correspond by design to real fire scenarios of interest. The idea of developing and

using such alternative temperature-time curves to replace the standard temperature-time curve

in ASTM E 119 has already been used in the case of simulating exposures to large hydrocarbon

pool fires [3-1].

Here the terminologyASTME 119-type test method is used to refer to a test method that basically

follows the ASTM E 119 test procedures, but where the ASTM E 119 standard temperature-time

curve is replaced with a relevant alternative curve. Of interest here are NPP-specific temperaturetime curves.

3.4 STRUCTURAL BARRIERS, WRAP ASSEMBLIES AND ASTM E 119

3.4.1 A Characteristicthat Distinguishes Between Structural Barriers and Wrap Assemblies

For the purpose of the present discussion, a characteristic that distinguishes structural-type from

wrap-assembly-type fire barriers is whether or not it is possible for ASTM E 119 to be applied

without ambiguity. In the case of structural-type barriers, it is possible. In the case of wrapassemblies, the nature of the barriers' physical design would preclude an unambiguous application

of ASTM E 119.

The above discussion leads to the following questions:

Structura-type_ fire barriers. Relative to evaluating the fire performance of NPP

structural-type fire barriers by ASTM E 119, are there situations where the test method

should be modified and/or replaced, and if so, what is the alternative?

Wrap-assembly-type fire barriers. Relative to evaluating the fire performance of NPP

wrap-assembly-typefire barriers, are the basic features of ASTM E 119 applicable? If so,

then how should the details of the test method (e.g., requirements for type of furnace,

specimen extent and placement in the firnace, conditions for acceptance, etc.) be changed

to accommodate the characteristics of a particular wrap-assembly-type of barrier of

interest? If not, then what is the alternative?

NUREG-1547 66
NPP-Specific Descriptions of Fire Environments

In the case of wrapassemblies the above questions are addressed in NRC Generic Letter 86-10,

Supplement 1 [3-8]. It is also noteworthy that the ASTM E5.11 Committee on Construction

Assemblies is presently working to respond partially to the above questions on ASTM E 119

applicability. In this regard, the E5.11.08 Sub-Committee on Cable Tray Protection has

developed and is now evaluating the new draft standard: Standard Test Methods for Fire Tests of

Fire-Resistive-Barrier Systems for Electrical System Components [3-91. Similarly, the

Underwriters Laboratory is in the process of developing a new standard: Standard for Thermal

Barrier Systems for Electrical System Components [3-10].

3.4.2 Describing Fire Exposure and Evaluating Fire Barrier Performance

In principle, evaluations of the performance of a particular fire barrier design should be made

relative to: 1) fire environments that could conceivably develop on the fire-exposed side of the

barrier; 2) time intervals required for protection (i.e., the entire duration of the fires, assuming

that they are not suppressed, or some other specified time intervals); and 3) specified bounds that

characterize an acceptable environment in the space on the protected side of the barrier, an

environment in which satisfactory functioning of the protected equipments is assured. In practice,

such performance evaluations must also include the use of good engineering judgement, leading

to appropriate levels of margins of safety, i.e., providing enough protection to account for

uncertainty in analytic methodology, degradation of a constructed barrier over time, etc.

Assume that a detailed description of the compartment of fire origin and of additional

compartments that would potentially influence fire development in it are known. (The description

of the compartments would include: dimensions of the spaces; quanfity, arrangement, and material

properties of combustibles; HVAC system characteristics and its status; characteristics of barrier

penetrations, including the opening status of doors and other variable-area vents; and, if relevant,

exterior weather conditions, especially wind conditions.) Assume that the details of an ignition

source of a particular fire scenario of interest are also known. Then, in principle, one would hope

to be able to: estimate characteristics of the developing fire environment; determine the thermal

and structural response of the fire-exposed barrier and the environment in the protected space;

and, finally, evaluate barrier integrity. However, in practice, it is beyond the scope of current

science and technology to describe from basic principles several of the fundamental and critical

processes that drive the overall phenomenology of interest. For example, the solution from first

principles to the critical generic problem of fire spread and growth through realistic arrays of

practical combustibles is not feasible (e.g., fire spread from a specified open flame across a

stuffed chair of specified design, or along cables of a specific design, placed, according to a

specified cable density and arrangement, in stacked steel trays of specified design and orientation)

Fortunately, the most important problems of fire safety can typically be expressed in terms that

successfully avoid the above type of technically intractable formulations, and that can yield

solutions of satisfactory reliability. Indeed, a major role and challenge of fire safety science and

technology is to formulate tractable problems whose solutions result in fire safe designs and/or

design practices that lead to improved reliability and/or economy. For example, while fire safety

in a facility can be achieved generally if ignition or early fire growth is completely eliminated,

in practical terms such elimination is usually not possible. Therefore, instead of dealing with the

67 NUREG-1547
NPP-Specific Descriptions of Fire Environments

issues of eliminating ignition, one often deals with problems of fire safety in facilities by first

positing a realistic design-fire threat, and then formulating and solving a now-tractable problem

of fire safety in terms of determining how to provide the desired level of safety "given initiation

of the design fire."

Examples of currently accepted practices with potential for improved reliability and/or economy

are those that depend on the fire barrier ratings methods based on equivalency rules and ASTM

E 119. Thus, it is possible that currently accepted practices are unnecessarily conservative, and

in some instances, perhaps not conservative at all. When such is the case, application of advanced

fire safety science and technology can be used to guide the development of improved practices.

As will be discussed below, careful problem formulation and current technology together can

provide important opportunities for future advances and refinements in may areas of fire safety,

including the problem area of interest here, viz., the evalualion of NPP fire barrier performance.

3.5 EXPOSURE TO THE INDIRECT AND DIRECT EFFECTS OF THE

FIRE: TWO TYPES OF THREAT

Introduced above was the idea that fire barrier performance is evaluated relative to exposure to

a particular specified fire environment or a class of fire environments, i.e., the fire or class of

fires that could develop in the barrier-bounded compartments of fire origin. In developing this

idea further, it will be useful to distinguish between two aspects of such fire exposures, indirect

and direct. This is introduced with the following discussion of fire environments and their

simulation.

3.5.1 The One- or Two-Layer Description of the Fire Environment and Zone-Type

Compartment Fire Modeling

As a fire grows and spreads through the combustibles in a compartment, it has been observed that

the environment of the bulk of the volume of the compartment can be reasonably described by two

relatively-uniform stratified layers, a relatively-hot and smokey upper layer and a relatively-cool

and uncontaminated lower layer, or, as is the case in many important fire scenarios, by a single

relatively-uniform layer that fills the entire compartment. This observation is the basis of twolayer zone-type mathematical compartment fire models.

When a two-layer or single-layer description of the fire environment is appropriate, the thickness

of the layers as well as the properties of the layers or layer are typically time-dependent. (In the

case of a two-layer environment, when one of the layers effectively shrinks to zero thickness, the

bulk environment of the entire compartment is approximated by the spatially-averaged properties

of the single remaining layer.) The time dependence is the result of:

the action of fire combustion zones which continuously generate rising, literal-entraining

plumes of hot combustion products, where the plumes typically rise to the ceiling of the

compartment and are deposited and mixed into the upper layer;

NUREG-1547 68
NPP-Specific Descriptions of Fire Environments

* convective and radiative heat transfer exchanges between bounding surfaces of the layers

(including the walls, ceiling, and floor of the compartment), the layer gases, and the

combustion zone; and

cross-vent (e.g., cross-door, -window, -HVAC-diffuser or HVAC-exhaust vents)

convective heat and mass transfer exchanges between the layers of the compartment of fire

origin and adjacent spaces.

Many of these phenomena typically take place within and across relatively small-volume

(compared to the volume of the compartment) zones, e.g., the combustion zone, the plume, zones

defined by plume-driven near-ceiling boundary flows or near-wall boundary flows, etc., where

these zones are submerged within the layers (3-11,3-12]. Fire scenarios involving a burning array

of cable trays and a pool fire are depicted in Figs. 3-1 and 3-2, respectively.

In terms of predicting the developing fire environment, the two-layer, zone-type, compartmentfire-modeling methodology has proven to be very useful in the analysis of a variety of different

kinds of problems of fire safety, and many such fire models have been developed over the last few

decades. Such models are constructed by modeling mathematically the above transfer phenomem

in ways that are consistent with conservation of energy, momentum, species, etc. and the basic

two-layer description. In this regard, perhaps the most significant differences between the

various available models are associated with the number of and sophistication with which

individual physical phenomena are taken into account. As a result of such differences, the

inevitable limitations of zone models will correspondingly differ one from the other. Examples

of zone-type compartment fire models that were developed specifically to simulate fire environmerits in NPPs are COMPBRN HI [3-13] (a two-layer model) and the Fuel Load/Ventilation

Model [3-14] (a single-layer model).

3.5.2 The Indirect and Direct Effects of the Fire and Applicability of ASTM E 119-7ype

Test Methods

From the above discussion, it is clear that for any particular compartment fire scenario, the

fire-environment exposure of the surface of a barrier in question will vary as a function of

position and time. However, even though a full-scale compartment fire experiment can, in

principle, yield data on time- and position-dependent exposure for a particular fire scenario

and barrier design, and even though it may be possible to simulate such exposure by

mathematical modeling methods, it is rarely practical that such detail would be useful in

specifying a fire exposure with which to evaluate the performance of a multi-use barrier

design. Also, even in a particular application, the nature of possible real fire scenarios

typically have significant stochastic aspects (e.g., the actual amount of combustibles, the

location of the fire relative to the barrier surface, etc.)

In many practical fire scenarios, all of a barrier surface, or most of a barrier surface for most

of the time will be removed from direct exposure to the relatively-small, but particularly

threatening zones of a compartment fire environment (i.e., the combustion zone, near-surface

boundary zones in the vicinity of plume/surface impingement points above the fire plumes,

69 NUREG-1547
NPP-Specific Descriptions of Fire Environments

etc.). Under such circumstances, for evaluating barrier fire performance it is often reasonable

to let the spatially-average d temperature history of the (upper layer) fire environment define

the fire exposure. Exposures to such spatially-averaged environments lead to what are

designated here as the indirect effects of the fire.

When the specified temperature-time description of a furnace fire simulates, in some sense, the

spatially-averagedtemperature of the threatening fire environment, it would appear from the above

that use of an ASTM E 119-4ype test method to evaluate barrier fire performance is well justified.

Thus, when the ASTM E 119 standard fire is a reasonable and conservative surrogate for the

temperature-time history of the average compartment fire environment of real, structurethreatening fire scenarios, ASTM E 119 evaluations of structural barriers would appear to yield

valid determinations of barrier performance relative to exposures from the indirect effects of the

fire. (It is noteworthy that the methodology developed in reference [3-14] is consistent with this

reasoning.) However, it is clear that such evaluations discount the importance of possible severe

fire exposures from direct and sustained flame/barrier impingement. Such kinds of intense

exposures lead to what are designated here as the direct effects of the fire.

The latter, unusual, sustained, and problematic type of fire threat to barrier integrity, defined by

the direct fire exposure, may be important in NPP applications.

3.5.3 Using the Direct and Indirect Fire Exposure to Define the Fire-Barrier Threat

As mentioned, where the ASTM E 119 test method is used to evaluate fire-barrier performance,

sustained direct fire exposures in the real fire scenario are not takeniinto account. It would appear

that the justification for this is that combustibles in building facilities are typically wel-distributeI

throughout the spaces, and that during the course of expected fire scenarios, combustion zone(s)

would tend to shift from place to place within the compartment. Under such common

circumstances, sole use of an ASTM E 1194ype test method to evaluate fire barrier performance

would be adequate.

However, there are special building compartment designs where sustained localized fuel

concentrations exist, and where the ASTM E 119-type approachto evaluating barrier fire performance may not be adequate. For example, if the source of most of the fuel load in a building

compartment was from a diked, ruptured, diesel-generator fuel tank, then the threatening fire

scenario would be defined as that which would be generated by. one or a series of geometricallywell-defined pool fires. For such scenarios, an evaluation, e.g., of ceiling fire-barrier performance that depended solely on exposure to an ASTM E 119-type furnace fire environment, that

simulated well the average temperature-timeof the real fire environment, may not be valid. Also

recommended would be a second test, where, for some extended specified time interval, say for

an interval that corresponds to the duration of the fire (that would have to be determined), the

ceiling fire barrier was exposed to a test fire environment that simulated sustained directflame/plume impingement from the pool fire which is itself contained in the compartment fire

environment.

NUREG-1547 70
NPP-Specific Descriptions of Fire Environments

ASTM E 1529 [3-1] is an example of a test method that is designed to simulate direct-flame/plunr

exposure, but in outside fire scenarios, rather than in compartment fire scenarios. This test

method is used to simulate the fire performance of structural members and assemblies when

exposed to large hydrocarbon fires. As reported in Part 2, section 2.4, the test exposure was

developed from theoretical considerations and test data from a wide range of real-fire scenarios

involving open-air spill and open-air fuel fires. ASTM E 1529 is similar to ASTM E 119, except

that instead of using a temperature-time curve, the furnace test fire is specified in terms of heat

flux to the exposed surface of the test specimen. The specified flux level must be attained within

the first 5 min of testing, and it must be maintained for the duration of the evaluation. Reference

(3-2] explains that a furnace with optically opaque walls can lead to a specified ASTM E 1529 fire

exposure if the furnace-gas temperature rises to 1020 "C within the first five minutes of the

initiation of the test, and if the gases are transparent and have in-furnace flow speeds of 10 m/s.

It would appear that this is a more severe exposure than that resulting from the standard ASTM

E 119 fire, which is specified as having a furnace gas temperature of 538 'C at 5 min, and which

does not rise to 1020 'C until about 2.25 hours into the test.

Examples of scenarios in NPP spaces where the direct fire exposure would be significant are

depicted in Figs. 3-1 to 3-3. Figure 3-1 involves fidly-developed burning of an extended dense

array of cable trays. The term fully-developed burning is intended to mean that all exposed

surfaces of a combustible object are supplying fuel (losing mass) due to either heating by a fire

environment or surface combustion. In this scenario, only the ceiling structure is exposed to the

direct effects of the fire. The protected and unprotected pairs of cable trays at right side of the

compartment are only exposed to relatively indirect effects of the fire, i.e., the upper layer

environment. Thus, although they are shown as being submerged in the near-ceiling boundary

flow, they are located far enough from fire/ceiling impingement that the somewhat more intense

environment there, over and above the average upper layer environment, may be regarded as not

significant. Figures 3-2 and 3-3 involve a large ruptured-fuel-tank-type pool fire, which can

produce sustained and particularly intense exposures to the flame zone by virtue of the fact that

the location of the pool, and therefore the flames, is assumed to be fixed (e.g., to a diked area of

the pool and above it) throughout the entire fire scenario. In the scenario depicted in Fig. 3-2,

only the ceiling above the pool is directly exposed to the flames. In Fig. 3-3, a pair of

unprotected cable trays and a pair of wrap-assembly-protectedcable trays are also directly exposed

to the flames.

3.6 A CONCEPT FOR EVALUATING THE FIRE PERFORMANCE OF

FIRE BARRIERS IN NUCLEAR POWER PLANTS - AN ALTERNATIVE

TO TRADITIONAL APPLICATION OF ASTM E 119

3.6.1 An Alternative to ASTM E 119

The critical features of the alternative methodology are that it be rational (i.e., related to real

NPP-specific fire environments and having a firm technical basis), appropriately advanced (uses

the most advanced, appropriate, modeling methods), and practical (i.e., is based on tools of

71 NUREG-1547
NPP-Specific Descriptions of Fire Environments

analysis that are readily available to and useable by fire safety practitioners, and realizable

experimental methods that use readily available facilities).

3.6.2 A Concept For Evaluating the Fire Performance of Fire Barriers In Nuclear Power

Plants

All the previous discussion suggests the following three-point concept for evaluating the fire

performance of fire barriers in NPPs:

1. DetermineNPP fire environments Determine the characteristics of the real fire environments (direct and indirect exposures) in the compartmentsbounded by a (structural) barrier

of interest, or within which a (wrap assembly) barrier of interest is installed. In making

such determinations a mathematical modeling approach will be emphasized here, and

modeling concepts for such an approach will be developed below. However, in special

cases, full-scale compartment fire experiments may be necessary.

2. Given the fire environment. define/identify methods for evaluating NPP fire barrier

performance. Define and/or identify methods for determining barrier fire performance

when exposed to the real-fire environments determined from the above modeling or experimental methods.

Indirect effects of the fire: Define an ASTM E 1 19-type test method for evaluating fire

performance of the fire barrier when it is exposed to the indirect effects of the fire. In the

test method, the temperature-time history of the furnace fire environment would simulate

or, if that is not possible, bound the temperature-time histories of the real fire environments, determined according to the above methods. Here, the ASTM E 119-type test is

emphasized. However, in special cases, and especially for structural barriers, it may be

feasible and appropriate to determine the thermal and structural response of the barrier

from mathematical modeling analyses. In such cases, the present task would be to identify

the appropriate method of analyzing the fire performance (e.g., available finite-element

computer models for simulating the thermal and structural response of the barrier [3-15]).

Direct effects of the fire: If the nature of the barrier and compartment designs is such as

to make sustained direct fire exposure likely, then the characteristics of this exposure,

including its duration, would have been obtained from above methods. The task would

be to develop a fire-barrier exposure fire test method and test duration that simulates or

bounds these direct fire exposures. This would be used to evaluate fire performance of

the fire barrier when it is exposed to the direct effects of the fire. -Here, an

experimental/test approach is emphasized. However, as in the case for indirect fire

exposure, in special cases it may be feasible and appropriate to determire the thermal and

structural response of the barrier to direct fire exposures from mathematical modeling

analyses. (Note that the methodology presented in reference [3-3] represents an approach

that models mathematically the direct fire threat. However, the modeling used there is

limited in the sense that it does not include the effects of full room fire involvement.) In

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such cases, the present task would, again, be to identify the appropriate method of

carrying out the evaluation of barrier response to the specified exposure.

3. Evaluate NPP fire barrier performance Evaluate the fire performance of fire barriers

by applying the test method(s) and/or method(s) of analysis of items 1 and 2 above.

The remainder of Part 3 will be concerned with the development of modeling ideas consistent with

the above concept.

3.7 A COMPARTMENT FIRE MODEL FOR USE IN EVALUATING THE

DIRECT AND INDIRECT THREATS TO FIRE BARRIERS IN NPPs

It is a major objective of this work to propose the development of a special-purpose compartment

fire model suitable for use in evaluating the performance of fire barriers in NPP facilities exposed

directly and/or indirectly to threatening pre- and post-flashover fire environments. Important

special features of a such a model are identified in this section.

3.7.1 Mathematical Fire Models and Associated Computer Codes

Over the years, many general- and special-purpose mathematical models and associated computer

codes for predicting the dynamics of compartment fires have been developed. For the most part,

these can be divided into two categories, field models and zone models.

Field models incorporate global partial differential equations, which describe the relevant flow

and heat transfer processes. They formulate and solve initial/boundary value problems for the

unknown variables of compartment fire scenarios of interest.

Zone models describe the phenomena of the fire scenario in terms of coupled submodel algorithms

or equations sets. Each algorithm describes a fire-generated process taking place in a particular

physical zone of the compartment of interest. The individual algorithm equation sets typically

involve the solution of algebraic equations (e.g., to estimate entrainment into and flow rate of a

fire plume), integration of algebraic equations (e.g., to estimate the instantaneous total rate of

convective heat transfer between a plume-driven ceiling jet and a ceiling surface), or the solution

of ordinary or partial differential equations (e.g., to simulate heat conduction through a wall,

ceiling, or floor). The mathematical coupling of the algorithm equation sets corresponds to the

interdependence of the individual physical processes that occur throughout the different zones of

the overall compartment fire environment.

The results of field model simulations typically provide significantly more detail of the firegenerated environment than do zone model simulations. Also, since the governing equations used

in field model can describe from first principles the actual physical phenomena being simulated,

field-type modeling has the potential for yielding the most accurate possible simulations. (An

example of an important class of phenomenon that, in practical terms, currently defies accurate

modeling is the combustion process itself.)

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In general, field models are significantly more computational intensive than zone-type models and,

for any particular simulation, they generally require significantly more effort to develop and

implement input data. As a result of this, field models are typically not well-suited for use in

parametric studies, e.g., in the trial-and-error-types of analyses often used to solve design

problems. Such studies are efficiently addressed by zone-type models.

In view of the above, it is concluded that a multi-room, two-layer, zone-type, compartment-firemodel analysis, with sufficiently detailed submodel algorithms, would provide the best possible

means of predicting real fire environments for use in determining the fire performance of NPP fire

barriers.

3.7.2 General-Purpose and Special-Purpose Compartment Fire Models

A goal of the technology of zone-type compartment fire modeling is to develop a general-purpose

model that is so rich in detail as to have near-universal utility and high reliability. A generalpurpose model would be capable of simulating the early growth of fire conditions (i.e., preflashover, when not all exposed combustible surfaces are pyrolyzing, and/or burning) and the

onset of smoke spread within and beyond compartments of fire origin, phenomena that are

typically characterized by time scales of the order of seconds and minutes. Such a model would

also be expected to simulate the relatively long-term quasi-steady aspects of fire environments,

when available combustibles in compartments of fire origin may be fully-involved in the

combustion process (i.e., post-flashover) and where important variations in the changing fire

environment are typically characterized by intervals of tens of minutes to a few hours. Besides

dealing with large differences in time scales, a general-purpose model would, ideally, also be

capable of simulating details of fire growth and smoke spread, taking into account a wide variety

of: types and arrangements of combustible assemblies, room-grouping arrangements and

interconnections, physical properties of exposed surfaces, forced ventilation designs, etc.

Development toward high-quality general-purpose models has always been, and continues to be

a major focus of fire research. In spite of this, when dealing with a particularly important

problem area there is often justification to develop special-purpose models that focus on classes

of fire phenomena and issues of special interest. When carrying out such development, appropriate proven concepts used in the general-purposemodels are typically adopted, improved on and/or

otherwise customized to satisfy the particular modeling requirements.

An important example of a class of special-purpose or customized compartment fire model is the

fully-developed post-flashover fire model. Consistent with the concept introduced above ('m

section 3.6.2, "A Concept For Evaluating the FIre Performance of Fire Barriers in Nuclear

Power Plants"), fully-developed fire models have been developed specifically to predict the longterm fire environments that threaten structural fire safety, where the simulations are used in the

evaluation of structural fire performance. Such models are of particular interest here, since they

deal with problems directly related to the class of problem at hand, i.e., determining the fire

performance of fire barriers in NPPs. An example of a relatively-simple fully-developed fire

model, developed explicitly to predict post-flashover fire environments in NPPs, is the Fuel

Load/Ventilation Method fire model of reference [3-14].

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Perhaps the most ambitious fire model developed explicitly to simulate NPP fire scenarios, is

COMPBRN MI [3-13]. This follows existing general-purpose types of models in that it includes

the capability to simulate fire growth and early smoke spread phenomena, as well as the relatively

late-time environment, provided pre-flashover conditions continue to prevail.

The suitability of using or advancing COMBRN IH or the Fuel LoadNentilation Method fire

model for the evaluation of fire barrier performance will be discussed in sections 3.7.4 and 3.7.5,

respectively.

3.7.3 Features of a Compartment Fire Model Suitable For Evaluating Direct and Indirect

Threats to NPP Fire Barriers

As discussed earlier, the fire barrier threat would be specified from knowledge of both direct and

indirect fire exposures. Furthermore, a realistic determination of the direct fire exposure threat

is dependent on, and would be determined from outputs of compartment fire model simulations

of the overall fire environment, where these outputs would typically provide an explicit

description of the indirect fire exposure. Thus, the same zone-type model that simulates the layer

temperature histories of the Figs. 3-1 to 3-3 scenarios, would also provide the basis for estimating

the intensity of the fire environment local to directly-exposed fire barriers. The main focus of the

proposed model will be on achieving realistic simulations of the fire environments, from which

both direct and indirect exposures would be determined.

Major features of a compartment fire model that can be expected to provide the necessary

simulations include:

1. Simulating fully-developed burning of the most sidgificant combustibles typically

found In NPPs. Because of the fact that NPPs typically contain a large variety of fuel

arrangements and materials, general analytic methods that are available (or that can

reasonably be expected to be developed over the next several years) to describe the details

of fire growth and spread are not (or would not be) reliable. Similarly, because of the

large variety of relevant fuel arrangements, etc., detailed empirical descriptions of fire

growth and spread that one might hope to develop from experiments on burning arrays of

real combustibles (e.g., like the experiments described in references [3-16 to 3-21]), would

generally require an unrealistically large amount of resources. Even then, as would be the

case with analytic descriptions, such empirical descriptions would be suspect, e.g.,

because of questions on the significance of interactions between the burning combustibles

and the developing compartment fire environment, etc.

While the prediction of fire growth and spread in NPP fire scenarios would generally

require solutions to intractable problems, it seems that empirically-based analytic models

for estimating fully-developed burning of the kinds of combustibles that threaten fire

barriers in NPPs are achievable and that it would be reasonable to use such models to

simulate the complicated combustion processes occurring in these NPP fire scenarios of

interest. Also, as will be seen below (in section 3.8.2, "Simulating Fully-Developed

75 NUREG-1547
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Burning of Cable Trays"), for the important class of scenario involving burning of

extensive dense arrays of cable trays, it is possible that limited, achievable experimental

studies would reveal that the time to reach fully-developed burning is generally small

compared to the total fire duration of interest. This would support an analysis based on

the modeling assumption that fully-developed burning is attained immediately upon

ignition. Furthermore, it is reasonable to expect that an assumption of instantaneouslyattained fully-developedburning is conservative in the sense that artificially rapid spread

to the point of complete and sustained involvement of all combustible surfaces would lead

to fire environments that are most threatening to the fire barriers whose integrity is under

evaluation. For this reason, it is recommended that a fully-developed fire assumption be

adopted for the proposed model. Note that this recommendation is consistent with the

opinion expressed in reference [3-14] that "total involvement of the combustibles in a room

(i.e., post-flashover) represents the most severe condition for containing a fire with

barriers."

In summary, the modeling assumption proposed here is that significant threats to the

integrity of fire barriers in NPPs are a result of direct and/or indirect exposures to fire

environments generated by sustained fully-developed burning, when all exposed

combustible surfaces in the compartment of fire origin are fully involved. Detailed ideas

on implementing the fully-developed-fire assumption for specific classes of relevant

combustibles will be presented section 3.8 below (under the heading, "THE

THREATENING COMBUSTIBLES FOUND IN NPPs; IMPLEMENTING THE

ASSUMPTION OF FULLY-DEVELOPED BURNING").

2. Simulations of the fire environment in multi-mrom fadlities This involves a capability

for simulating the development of the threatening fire environment on both sides of

structural barriers, including the environment in a room of fire origin, on one side of the

barrier, and in an adjacent space, on the other side of the barrier, into which high

temperature gases can be transported via significant ventilation penetrations (i.e., a

capability for simulating the development of the fire-generated environment in different

interconnected rooms of a multi-room facility).

3. Simulation of forced and natural ventilation. The nature of the fully-developed fire

environment is strongly dependent on the rate of inflow of ventilation air, i.e., oxygen,

to the comrtment of fire origin. Because of the wide variety of different spaces found

in a typical NPP facility, ventilation to spaces of fire origin during the course of a fullydeveloped fire scenario involves one or a combination of any or all of the following types

of ventilation, some of which are indicated in the sketch of Fig. 3-4:

Vertical vents to the outside, including wind: Natural, i.e., buoyancy-driven,

venting through vertical vents to the outside environment, e.g., through doors or

windows, including the effect of wind [3-22 to 3-25].

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Vertical vents to adjacent spaces: Natural venting through vertical vents to

adjacent spaces of the facility [3-22 to 3-25]. Here, an accounting of multi-room

considerations are critical. See Fig. 3-4.

Horizontal vents to the outside, including wind: Natural venting through

horizontal vents to the outside environment, e.g., holes in floors or ceilings [3-

26,3-27]. See Fig. 3-4.

Horizonal vents to adjacent spaces: Natural venting through horizontal vents to

adjacent spaces of the facility [3-26,3-27]. Here, as in item 2, an accounting of

multi-room considerations is critical. See Fig. 3-4.

Forced ventilation: Forced ventilation via HVAC systems [3-22,3-23,3-25,3-28],

including a capability of simulating effect of the fire environment on

specified/designed flow rates.

In predicting the fully-developed fire environment, an accounting for each of the above

types of ventilation requires special considerations (refer to the latter-mentioned references

[3-21 to 3-28]).

In summary, the proposed ventilation-simulation feature involves a capability for

simulating a variety of different room-to-room and/or room-to-outside ventilation flow

phenomena, including those associated with forced ventilation, natural-ventilation flow

exchanges through penetrations in vertical and horizontal partitions, and fire-driven

variations to design flow rates delivered by forced-ventilation systems.

4. Capability to distinguish at some non-trivial level of refinement, and to simulate the

changing spacial variation in temperature of partition/barriersurfaees. This includes

a capability of simulating: radiative heat transfer from bounding surfaces of a room to

arbitrarily located targets, and the variability across the different partition/barrier

structures of their in-depth thermal responses. As will be seen below (in section 3.8.3,

"Simulating Burning of Combustible/Flammable Liquid Pool Fires"), this capability

would be used in the simulation of fire scenarios involving combustible/flammableliquids.

This modeling would be achieved by implementing and extending the ideas presented in

References [3-29] and [3-301. Each wall/floor/ceiling partition of a room would be

divided into "several" rectangular segments, each with a pair of rectangular surface

elements, one on each side of the partition/barrier. In each room, energy conservation

considerations would account for radiative heat transfer exchanges between gas layers,

combustion zones and partition/barrier-surfaceelements. Heat transfer through the depth

of the segmented partition/batrrier structure elements would be simulated by an appropriale

wall submodel. Beside providing a reliable closure on the model equations of energy

conservation, as required the latter would also provide the basis for reliable simulations

of partition/barrier structural response.

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The fully-developed fire modeling features of item 1, the more challenging ventilation-simulation

capabilities of item 3, and the relatively sophisticated heat transfer analysis of item 4 each require

additional submodel/algorithm development and implementation and/or experimental validation.

In contrast to this, the multi-room modeling capability of item 2 and some of the basic ventilationsimulation capabilities of item 3 would tend to already be included in the features of any existing

successful multi-room fire model. It is therefore recommended that the proposed model be

developed, as a matter of economy, as an advanced version of such an existing model.

3.7.4 Suitability of Using COMPBRN MI for the Evaluation of Fire Barrier Performance

As mentioned above, the fully-developed post-flashover fire is particularly threatening to the

satisfactory fire performance of fire barriers. Therefore, when there is a possibility that a fire in

a particular NPP enclosure can develop to a flashover condition, it is the potential flashover-fire

environment that must be predicted and used to evaluate fire barrier performance. Unfortunately,

COMPBRN M does not include a capability for making such predictions ("The assumptions made

in COMPBRN MI are geared towards the modeling of relatively small fires in large enclosures or

fire scenarios involving large fuel loads early during their pre-flashover burning period." [3-13]).

For this reason alone, COMPBRN HI in its current state of development is not suitable for use

in the fire barrier evaluations of post-flashover fires.

Regarding experimental validation of COMPBRN M, reference [3-13] provides comparisons

between data from pre-flashover enclosure pool-fire experiments and corresponding COMPBRN

I1 simulations. Model predictions of steady-state hot-gas-layer temperatures are described as

"reasonably accurate" in tests involving steady combustion of a methane burner. In tests

involving a heptane pool fire, the model is described as being "able to bracket actual experimental

values" when, in the simulations, the combustion efficiency is taken to be 1.0 and 0.7. Tbe input

parameters chosen for the computations analyzed in reference [3-13] are tabulated in that work.

It is noted there that "these parameters are empirical and depend on ... detailed characteristics of

the particular fire scenario ... " Further, it is explicitly stated that "these parameters are not wellknown for the arbitrary scenario" and it is concluded that "significant input uncertainties exist

even in the case of well-controlled experiments." Even when used in specific applications for

which it was designed, the above observations, reported in what can be considered as a

COMPBRN MI source document, do not instill confidence in the accuracy and reliability of

COMPBRN MI simulations.

There are additional aspects of the COMPBRN II model that would significantly limit its

applicability in fire barrier evaluation. Based on the model description in reference [3-13], some

of these are:

COMPBRN III is a one-room fire model (the compartment of fire origin) where the only

vent available is a vertical vent to a quiescent outdoor ambient environment. Thus, for

example, none of the ventilation flows depicted in the sketch of Fig. 3-4 can be simulated

by COMPBRN III.

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In COMPBRN IM, there is no mechanism for contamination of the lower layer of the

compartment of fire origin. ("The lower region is assumed to be thermally inert and its

temperature remains at ambient room temperature all the time.") In real fire scenarios,

in the one-room/one-vent configuration, and even more so in more complicated

configurations, mixing and beat transfer mechanisms will lead to both contamination (e.g.,

reduction of oxygen concentration) and heating of the lower gas layer. All this could have

important implications on the simulated room fire environment and on the simulated threat

to barriers.

In COMPBRN III, all radiation exchange between the gas layers of the compartment of

fire origin and its inside bounding surfaces is assumed to occur between the upper layer

and the assumed spatially-uniform-temperatureceiling surface. ("... only the ceiling area

is used to compute the radiative heat loss of the hot gas layer to the room boundaries.")

Especially at higher upper layer temperatures, and even in the pre-flashover state,

radiation exchanges with all exposed surfaces of the compartment of fire origin will have

a large effect on the compartment fire environment. Such radiation exchanges will not

usually be adequately simulated by the ceiling-radiation limitation.

There are phenomena modeled in COMPBRN Ill (e.g., wall jet mass flow rate and

doorway mixing rate) that depend on roughly defined values of empirical parameters,

where the use of such parameters often have a weak theoretical basis and/or are not wellknown. (Refer to above comments on the experimental validation of COMPBRN Ill.)

Furthermore, it would appear that variations of the values of these factors within

recommended ranges could lead to significant variations in results of model simulations.

Advanced modeling ideas can be used to improve the simulation of these phenomena.

In view of the above shortcomings and limitations of COMPBRN MI, and in view of the

recommendation that the model for evaluation of fire barrier performance be developed as an

advanced version of an existing multi-room fire model, it is concluded that COMPBRN HI is not

suitable for use in its current state of development, and it is not suitable as a candidate for

advancement.

3.7.5 Suitability of Using the Fuel Load/Ventilation Method Fire Model for the Evaluation

of Fire Barrier Performance

A major and reasonable assumption of the Fuel Load!Ventilation Method fire model [3-14] is that

fully-developed, post-flashover, fire conditions always prevail in the compartment of fire origin.

However, as with COMPBRN Ill, a shortcoming of the Fuel Load/Ventilation Method model, is

that it is a one-room fire model that only allows for a single vertical vent to a quiescent outdoor

ambient environment. Also, the basic assumption of the model's fuel mass-loss rate (i.e., a

specified constant value for fuel-controlledburning, proportional to ventilation rate for ventilationcontrolled burning) is such that it can not be expected to provide reliable results when simulating

the burning of combustible/flammableliquids fires; and this in spite of the fact that, as discussed

in the next section, combustible/flammableliquid fires lead to one of the two major threats to fire

barrier integrity. Finally, the model is based on assumptions that are claimed to generally lead

79 NUREG-1547
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to conservative results, in the sense that the simulated environments will presumably always

represent a more severe threat to fire barriers than actual environments. Even if true, the level

of conservatism of the simulations is not quantified. For this later reason alone, use of the Fuel

Load/Ventilation Method fire model in the present application is problematic.

3.8 THE THREATENING COMBUSTIBLES FOUND IN NPPs; IMPLEMENTING THE ASSUMPTION OF FULLY-DEVELOPED BURNING

3.8.1 The Threatening Combustibles

There are a great variety of ftnctions and types and sizes of spaces that make up an operating

NPP. In these, it is possible to group all potentially significant combustibles into five categories:

cable trays, combustible/flammableliquids, electrical panels/cabinets, waste, and miscellaneous.

Based on a preliminary combustible-load summary of the Watts Bar NPP'0 , spaces having the five

largest density of combustibles were identified and tabulated in Table 1. The primary and

secondary combustibles in these spaces are also identified in the table.

Not surprisingly, Table 3-1 indicates that the most significant combustible of NPPs are the cable

trays. This continues to hold true for most other categories of relatively high-combustible -

density spaces of Watts Bar, not appearing in the table. Also, combustible/flammable liquids

are seen to be the major contributor of combustibles in the Diesel Generator/Lube Oil Storage

areas, which have the second largest average density of combustibles. Finally, although the

miscellaneous category of combustibles appears as the primary contributor in the case of the

Shift Engineer office, this particular space has a relatively small area, and other spaces where

"miscellaneous" is the primary category of combustibles is only found infrequently throughout

the rest of the plant (and then, only with considerably reduced average combustible density).

It is evident that waste and miscellaneous combustibles, and even relatively small amounts of

combustible/flammable liquids, can play a significant role in the ignition and in early fire

growth and spread of threatening NPP fires.

Results of full-scale experiments on growth and spread of fires initiated in NPP control cabinets

are presented in references [3-19] and [3-20]. These indicate that "a cabinet fire can propagate

within a single cabinet; however for the limited conditions tested it does not appear that the fire

poses a threat outside the burning cabinet except [for] the resulting smoke." [3-19]

Nevertheless, *many potential fire-spread paths were not investigated" in the experiments. In

this regard, results of the tests suggested that "partial or incomplete barriers and unsealed cable

•Fire Protection Staff of the Watts Bar Nuclear Power Plant, Combustible Load Summary for Watts Bar

Nuclear Power Plant (Draft), private communication.

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penetrations can be expected to allow further spread of fire, given a fully involved cabinet fire."

[3-20]

It seems reasonable to expect and it is tentatively assumed that, as in the case of control

cabinets, the general electrical-panels/cabinets category of combustible can play an important

potential role as a generic site of fire initiation and spread.

Based on the above, it is reasonable to assume that fire in electrical panels/cabinets and in

relatively small amounts of combustible/flammable liquids, waste, and miscellaneous

combustibles can play a significant role in the ignition and in early fire growth and spread of

threatening NPP fires. In this sense, for example, electrical cabinets such as switchgear and

motor control centers, must be recognized as significant fire hazards. However, based on

Table 1, and compared to cable trays and combustible/flammable liquids, they are expected to

be relatively small contributors to the fully-developed sustained fires that threaten fire barrier

integrity.

It is assumed that the Watts Bar NPP is representative of all NPPs considered in this work.

In view of all the above, it is concluded that there are two categories of combustibles that

represent a significantly greater fire threat to the integrity of fire barriers in NPPs than all

others; namely, cable trays (e.g., see Figs. 3-1 and 3-4) and combustible/flammable liquids

(e.g., see Figs. 3-2 and 3-3). For this reason, it is necessary to consider only these two in the

modeling and analysis of fully-developed fire-generated environments of interest here. The next

two sections will propose the basis of methods on how to model fully-developed burning of

cable trays and combustible/flammable liquids, respectively.

3.8.2 Simulating Fully-Developed Burning of Cable Trays

Fully-developed burning of extensive dense ara of cable tays: a uniform-mass-loss-rate

model. The cable-tray category of combustibles, so pervasive in NPPs, is deployed primarily

in configurations involving extensive dense arrays of loaded cable trays. Compared to more

typical distributions of combustibles which are based mainly near floor elevation, arrays of

loaded cable trays tend to be well-distributed within and mainly, but not entirely, confined to

the upper portion of the volume of the space, i.e., from just above head-height to near-ceiling

elevations.

The cable-tray category is exceedingly complex and varied. For any particular compartment

of potential fire origin, a detailed description would involve: the properties of a variety of (both

combustible and non-combustible) material and the configuration of these that make up relevant

cable sub-assemblies; the somewhat haphazard grouping of these different cables types onto one

or more types of cable tray fixture; and the different possible custom configurations of the cable

tray fixtures as they are mounted throughout the compartment. As stated earlier, the problem

of predicting fire growth and spread through such arrays of combustibles is not generally

tractable. Moreover, the location of many cable trays in overhead regions that, in the event of

a fire, are highly susceptible to hot vitiated gases, further complicates the problem since fire

81 81NUREG-1547
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spread and growth in vitiated atmospheres is not well understood. Clearly, some judicious

simplifications are required if the pyrolysis/burning of cable-tray combustibles is to be

simulated.

Using theoretical considerations together with data from grouped-cable-tray fire experiments

[3-18], it is the purpose here to present a simpI peculati-e model, a uniform-mass-loss-rate

model, for the fully-develop•d burning (all exposed combustible surfaces are losing mass) of

such arrays. Additional experiments would be required to validate, improve, or provide the

basis for an alternative to this model. Guidelines for such experiments are proposed below (in

sub-section, "Experiments to validate, improve, or provide the basis for an alternative to

the uniform-mass-loss-rate model").

In the case of cable tray combustibles, the proposed fire scenario threat is one where,

immediately after ignition, all cable trays (i.e., the exposed surfaces of all individual cable

bundles within the array) in the compartment of fire origin are fully involved in the combustion

process. Since the generic configuration involves extensive and dense groupings of cable trays,

most of the exposed combustible surfaces within the array exchange radiation with surfaces of

near-identical properties that completely surround them. The situation is analogous to the

burning of dense cribs in ventilated compartments, where, at least for fuel-controlled burning

(i.e., when the rate of ventilation air to the compartment is greater that the rate required to

support stoichiometric burning of the combustibles being pyrolyzed), mass-loss rate (i.e., fuelproduction rate) is mainly independent of the compartment fire environment and phenomena

outside the crib. Thus, for fuel-controlled burning of cribs, mass-loss rate is approximately

independent of: the temperature of the smoke layers, the temperatures of bounding surfaces of.

the compartment, and the radiation exchanges between these latter types of surfaces and the

outer bounding surfaces of the cnib. Other evidence, presented below, suggests that the crib

analogy may be appropriate for ventilation-controlled burning/pyrolysis of cable trays as well.

Fuel-controlled burning. Under conditions of fuel-controlled burning, and consistent with

Table 2-1.3, Fig. 2-1.3, and Eq. (4a) of reference [3-32], the total mass loss rate of a dense

cnrb, fibm, is proportional to the total exposed surface area of all *sticks" of the crib, Arx, and

a characteristic regression rate of the surfaces. The latter regression rate is proportional to a

stick-material-dependent constant, Csm, and D. 6 , where D is a characteristic dimension of the

stick section.

cm• -= Cc=AEx; CcRIB = CsMD'°'6; Csm = Cm(crib material) (1)

Developing further the analogy between the burning of a crib and an extensive dense array of

cable trays, consider the total exposed surface area of the latter:

Associated with a single loaded cable tray of the array, is the outward exposed surface area of

the combustible jacketing material of the cable bundle contained in the tray. It is reasonable to

assume that whatever the mix of type and size of cabling, all cable trays of the array contain

few-to-several full layers of cabling. Thus, for typically shallow trays (i.e., HT < < WT, where

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NPP-Specific Descriptions of Fire Environments

HT and WT are the height and width of the tray, respectively) the exposed area of the cable

bundle of one tray is approximately 2WTLr, where LT is the length of the cable tray. (The

factor of two is used when the cable tray is mostly open on all sides, e.g., from below and

above, in the case of horizontal trays, where both lower and upper surfaces of the cable bundle

are freely exposed to the environment surrounding the tray. When and if the cable tray is closed

on the top or bottom, the factor of two is not used, etc.) Also involved is a relatively small

additional exposed area corresponding to the exposed sides of the bundle, i.e., 2H-Lr, where

H. is the height of the bundle. Here, assume that the average height of the bundle in the cable

tray is half the height of the tray, i.e., H. = HT/2. Assume, further, that the contained cable

bundles and all trays of the entire array are approximately the same. Then, analogous to Eq.

(1), the total mass loss rate for a burning array of cable trays is

rhAmy = C.AyAEx; Ax = (2WT + HT)LrToA (2)

where C~m is a function of the cable array design, and Lro is the total length of all cable

trays in the array.

C .Ay of Eq. (2) is analogous to Cc= of Eq. (1). In general, it is expected to depend

on the combustible cable insulation and jacket material properties and on the characteristic

grouping of cable trays within the cable tray array (i.e., on the dimension of the tray section and

the relative spacing and configuration of trays within the array). However, it seems that the

characteristics of cable tray grouping configurations in NPPs are fairly uniform (approximately

as in the experiments of [3-16 to 3-18], with tray height and width, 0.1 in and 0.5 in,

respectively; vertical tray-to-tray separation, 0.3 m; and side-to-side tray separation, 0.2 i).

Therefore, it is reasonable to assume a weak dependence of Ccm on grouping geometry.

It is possible to extract preliminary estimates of Cny from results of two compartment fire

experiments involving the burning of relatively large arrays of cable trays. These experiments,

the ones designated as Test 2 and Test 3 in reference [3-18], used cable with polyethylene (PE)

insulation and a polyvinyl chloride (PVC) jacket. It will be assumed that a 50-50 mix of PE and

PVC was used, with a resulting average heat of combustion of 30x10 6 J/(kg burned). For the

cable trays used, WT = 0.46 m and HT = 0.08 m. To enhance fire spread in the tests, an

extra tight arrangement of cables, including weaving of cables between trays, was utilized.

Both tests involved two adjacent (0.15 in side-to-side separation gap) stacks (0.27 m top-tobottom separation gap) of six, 2.44 m-long, horizontal, loaded cable trays (as depicted in Figs.

3-1 and 3-4) with two(Test 2)-to-three(Test 3) adjacent vertical cable trays. Test 3 involved

additional interweaving of tray-to-tray cables. It also included an additional horizontal cable

tray directly overhead, but 2.44 in above the main array. (This was connected to the main array

by the cables contained in the vertical cable trays.) In both tests, the fire was initiated from a

combustible/flammable liquid pool fire below one of the two lowest cable trays. The fire spread

through most of the array by the time the first sprinklers were actuated at 289 s (Test 3) and

368 s (Test 2) into the tests. As indicated in Fig. A-1 of [3-18], in the absence of the sprinklers,

continued growth to what would have been a fully-developed fire would have been expected.

83 NUREG-1547
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From that figure, it is estimated that for both tests the fire would have been fully-developed by

approximately 900-1200 s. Assuming a 70% efficiency for burning of pyrolysis products, the

Fig. A-1 plots lead to the following estimate

CC = 0.05 kg/(m2s) for PE(insulation)/PVC(jacket) cable (3)

where a determination of possible significant variations in the above value of CA y for other

cable constructions would require additional study.

According to references [3-16] and [3-17], another full-scale test involving two adjacent stacks

of seven, heavily-loaded, horizontal cable trays was carried out. In this, "all the cables in the

test were completely consumed." [3-17] During the course of the present analysis it was not

possible to obtain and include any additional results that may be included in the original report

[3-33] of this test. To the extent that such results exist and can be made available, further study

of this test is merited.

Ventilation-controlled burning. For burning wood cribs, the mass-loss rate changes when the

air supply to the compartment of fire origin is reduced to the point that ventilation-controlled

burning prevails [3-32]. However, according to reference [3-34], the "burning [i.e., mass-loss

rate] of non-charring fuels [in crib-like arrays] is virtually unaffected by ventilation rate." This

statement is based on a set of 15 burn experiments on single non-charring-plastic 31 cribs with

specified ventilation rates. In the tests, ventilation was varied over a range from 0.5 to 1.3 of

the ventilation-controlled limit (the exact ventilation rate required for stoichiometric burning).

The mass-loss rates of the cribs for all test runs were identical to within ten percent of the mean

value.

The latter crib-bum behavior is to be compared to the well-known commonly-used behavior

observed in burning wood cribs, whereby, for ventilation rates below the ventilation-controlled

limit, mass-loss rate is approximately proportional to ventilation rate [3-35].

Because of the crib-like features of the cable tray arrays, it seems reasonable and it is tentatively

proposed to adopt the latter result, and to extend to under-ventilated burn scenarios the Eq. (2)

results for over-ventilated cable-tray-array burns. Note that such a combustion model should

be adopted with confidence only after it is supported and validated by results of an experimental

program involving full-scale under-ventilated burns of cable tray arrays. Also, there has to be

some low-ventilation-rate limit where the Eq. (2) model becomes invalid, and where additional

considerations have to be taken into account. Thus, at some sufficiently low ventilation rate,

perhaps when the oxygen concentration near the elevation of the base of the burning array of

cable trays is reduced to some specified, non-zero value, associated with flame extinction, it is

expected that all combustion within the array will cease, or become insignificant [3-36].

31 Sticks formed from a resin of approximately 60% unsaturated polyester and 40% cross-linking agent,

mainly styrene monomer, and 1% hardener (60% methyl ethyl ketone and 40% diethyl phthalate) [3-341.

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Experiments to validate, improve, or provide the basis for an alternative to the uniformmass-loss-rate model. The above-proposed uniform-mass-loss-rate model is speculative and

it is based on and supported by only a minimum experimental data base. It is important that

additional experiments be carried out to validate, improve, and, if necessary, provide an

alternative to this model. The new experiments should include some free-burn-type scenarios

(i.e., burning of a relatively large-density grouping of loaded cable trays in a space that is so

large and/or so well ventilated that the effect of the environment in the enclosure can not be

suspected of modifying significantly the burn characteristics of the grouped cable trays), but

mainly they should involve enclosed spaces, similar in room scale (possibly somewhat smaller)

to that used in reference [3-18] (i.e., 12 m x 12 m x 6 m high). They should involve both

natural- and forced-ventlation scenarios, both in the fuel controlled and ventilation-controlled

regimes, and scenarios with fuel densities comparable to those indicated in Table 1. Because

of the latter fuel-density requirement, such experiments would involve substantially greater

amounts of combustibles than those used in the studies of references [3-16 to 3-18].

The experimental fires should be allowed to develop to a burn-out, or at least a burn-down

stage. The enclosure structure would have to be robust enough to survive the significant fire

environment that can be expected to develop within the test space.

An experimental program that follows the above guidelines would involve a significant effort,

and it should based on a well-developed plan with clearly stated objectives.

Until the uniform-mass-loss-rate model is validated or until an improved cable-tray burn model

is established, when carrying out model simulations during ventilation-limited combustion, it

would be prudent to also carry out parallel simulations that use a traditional crib-type model

(i.e., mass-loss rate proportional to ventilation rate during ventilation-limited combustion).

Such a model can be developed with full generality (i.e., not confined to scenarios where

ventilation is from a single natural vertical vent to an outside quiescent environment) following

the ideas presented (on p. 310) in reference [3-35]. However, it is important to point out that

for other than cellulosic crib-like combustibles, and in the absence of experimental verification,

confidence in the traditional approach is not warranted [3-37].

Note that using the uniform-mass-loss rate model or an improved alternative model, instead of

the traditional model, would likely lead to a significant difference in the resulting simulated fire

scenarios. It would lead to predictions of much more rapid pyrolysis of the cable bundles,

without corresponding increased combustion in the compartment of fire origin. Significant rates

of pyrolysis products would be predicted to spread to and burn in adjacent spaces.

3.8.3 Simulating Burning of Combustible/Flammable Liquid Pool Fires

A basis for the model. In the fire scenarios of interest here, flame spread across the surface

of a combustible/flammable liquid pool fire and any important transient characteristics of the

pool itself develop so rapidly that the fully-developed fire assumption is always relevant. Then,

the net mass-loss rate of the fuel at the surface of a pool, rietOL, is determined by QPOOL, the net

rate of heat transfer absorbed at the pool surface,

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IhPOOL = QPOOL/L; QPOOL = QF + QE- QL (4)

where L is the heat of gasification of the combustible/flammable liquid, QF and QE are the heat

transfer from the flame and from external sources (the gas layer and bounding compartment

surfaces), respectively, and QL represents the heat transfer losses through the pool surface [3-

35]. Thus, ihpO0L does not depend explicitly on the rate of ventilation to the compartment, or

on whether the compartment fire scenario involves fuel-controlled or ventilation-controlle d

burning.

A single-room, single-vertical-vent fire model for ventilation-controlled burning of a steadystate pool fire is presented in reference [3-37]. The model accounts for radiation heat transfer

exchanges between the pool surface, the compartment gases, the inside bounding surface of the

compartment, and the vent opening. The pool mass-loss rate is modeled by Eq. (4).

Experimental data have been acquired and presented in reference [3-38] for pool fires (both

combustible/flammable liquids and thermoplastics, with areas of 0.186 m2 and 0.372 m2) in a

compartment (2 m wide x 1 m x 1 m) with a single vent. The experimental configuration

corresponded to that modeled in reference [3-37]. In the experiments it was found that,

compared to unconfimed burning of the pools (i.e., burning in the open, outside the confines of

a compartment), for ventilation-controlled burning, the high temperature environment of the

compartment had a very strong influence on increasing the radiation to, and the mass-loss rate

of the pool. For example, mass-loss rate for an ethanol pool was increased up to 6 times. Good

correlation was obtained between measured radiation flux data, tpowL data, and the Eq. (4)

model. However, the data indicated that an attenuating layer of fuel vapor immediately above

the pool surface could be significant in absorbing and attenuating the radiation incident to the

pool surface. The effect was observed to be significant in only one of the tests, involving the

larger of the PE pools, where the data indicated that only 2/3 of the radiation penetrated the

conjectured attenuating layer to actually penetrate the pool surface itself [3-38].

In view of the above, it is proposed that an advanced reference [3-37]-type model, which also

incorporates Eq. (4), be used to model combustible/flammable liquid pool fires in the present

NPP application. The model should be capable of predicting with reasonable accuracy the net

radiation absorbed by the surface of the pool. It should be able to simulate both confined

(diked) pools and relatively-unconfined large-area spills. This would require relatively detailed

modeling of the radiation exchanges between compartment boundary surfaces, the gas layers,

the flame, and the pool surface, a level of detail that goes beyond that existing in any currentlyavailable zone-type compartment fire model and beyond the level of detail used in the reference-

[3-37] model. An appropriate level of detail in the modeling of the radiation exchanges would

be achieved with the N-surface-element approach of reference [3-29], or an N-surface

advancement to the approach of reference [3-30]. The model should be capable of dealing with

both ventilation- and non-ventilation-controlled burning, where ventilation could result from

flow exchanges through multiple vents to the outside or to multiple adjacent spaces, including

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any of the various vent configurations discussed earlier. These model features are consistent

with and included in the list of model features outlined above in section 3.7.3.

Note that the comparison between data and theory reported in reference [3-38] is between plots

of Eq. (1) and plots of radiation vs ihiOL, from experimental values. Comparisons between

actual rhpooL predictions from the compartment fire model of reference [3-37] and

corresponding rhpooL measurements from reference [3-38] are not provided. In this regard,

it is important that the validity of the model, developed according to the above guidelines, be

verified with relatively-large-scale reference [3-38]-type compartment pool-fire data.

3.9 A NPP-SPECIFIC FIRE MODEL AND SCENARIO-SPECIFIC

TMREATS TO FIRE BARRIERS

It is proposed that the above modeling ideas be developed and implemented. The result would

be a new, special-purpose, NPP-specific, fully-developed fire model capable of simulating fire

environments that threaten NPP fire barriers. The outputs of such a model would include an

estimate of the indirect threat exposure and the variables required to estimate possible significant

direct threat exposures.

3.9.1 NPP-Specific ASTM E 119-7ype Temperature-Time Curves and Indirect Threat

Exposures

In simulating a fire scenario defined by a particular NPP facility, room of fire origin, and

ventilation configuration, the new model would predict, along with a variety of other output

variables, the temperature-time curve that characterizes the elevated-temperature history of the

upper layer. In principle, this curve would represent a scenario-specific alternative to the

temperature-time curve of the standard ASTM E 119 fire. If it were feasible to replicate such

a temperature-time-curve environment in an ASTM-type test furnace, the curve would be the

basis for an ASTM E 119-type test for verifying the integrity of fire barrier designs against

indirect fire exposures, used in the scenario-specific NPP room of fire origin.

Even if the new fire model could be relied on to produce accurate simulations, and if ASTM E

119-type test furnaces could be controlled to generate perfect replications of the simulated

temperature-time curves, a strict application of the above approach can never be practical. For

example, when considering a particular facility and room of fire origin, including specified

combustibles, it is appropriate to consider possible variations in the room ventilation, e.g., one,

two, or no doors may be opened to a relatively large adjacent space or to the outside, ceiling

vents may be open or closed, forced ventilation systems can be operating or not, etc. Such

variations can lead to significantly different fire environments and to different levels of threat

to the fire barrier whose performance is being evaluated. Thus, in principle, a barrier design

under evaluation would have to be tested multiple times to show that it performs adequately

when exposed to each of the simulated threats. Also, a particular barrier design would typically

be used in multiple locations throughout a NPP (each location requiring its own set of

simulations), indeed, it may be used throughout a variety of different NPPs. The number of

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required temperature-time curve simulations and corresponding ASTM E 119-type tests of the

barrier would typically be very large and completely impractical.

3.9.2 The Bounding-Temperature Principle and NPP-Specific Fire Curves

The multiple-test difficulty can be resolved by invoking the following "Bounding-Temperature"

Principle:

"Bounding-Temperaftre" Princidle: Consider threats to the integrity of a fire barrier

design when it is exposed to two different fire environments, each characterized by its

own temperature-time curve. If one of the two curves bounds the other at all times up

to a time, t, then up to that time the fire environment associated with the bounding curve

poses the greater of the two threats to the fire barrier.

The bounding-temperature principle can be applied as follows:

1. Draw a bounding temperature-time curve over all, or a distinguished group of the aboveconjectured simulated temperature-time curves (e.g., the distinguished group of curves

can be those curves that characterize all simulated fires associated with those rooms of

the original NPP where the barrier design of interest will be deployed).

2. Let the bounding curve define a design-basis temperature-time fire environment (i.e.,

compared to the simulated real fire environments) that is reproducible in an available

ASTM E 119-type test furnace. This is designated as the new, NPP-speciftc, testfire

curve.

3. Evaluate the integrity of the barrier while it is subjected to an ASTM E 119-type test

carried out with the new test fire curve. From the test, determine that the integrity of

the barrier is maintained for the fire resistance rating time interval, tR < t. Then the fire

barrier in question is given a fire rating of tR relative to the new test fire curve. It

follows from the bounding-temperature principle that the fire barrier in question is also

considered to have (at least) a tR fire rating relative to all individual simulated fires of

the distinguished group of fire threats.

Although not explicitly stated in the above terms, the bounding-temperature principle was used

as an integral part of the Fuel Load/Ventilation Method fire model [3-14], and it seems

reasonable to accept generally its validity when evaluating the performance of NPP fire barriers.

Indeed, the above concept of a NPP-specific fire resistance rating system and a NPP-specific

test fire is dependent on the general validity of the bounding-temperature principle.

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While it is reasonable to expect that the bounding-temperature principle is valid and useful for

evaluating the performance of most fire barriers, it is should be mentioned that the principle is

not universally valid.3

3.9.3 Developing Multiple, NPP-Specific, Temperature-Time Curves

In the above discussion, it was seen how a single group of simulated fire curves could be used

to define a single, bounding, NPP-specific test fire curve. In carrying out model simulations

for the different fire scenarios expected in a variety of different NPPs, it would, in an analogous

way, be possible to identify meaningful multiple groups of such simulated fire curves, and

corresponding multiple NPP-specific test fire curves. The most appropriate one of these test fire

curves (the one that bounds all relevant simulated temperature-time curves) could then be

recommended for selection and use in an NPP-specific ASTM E 1 19-type furnace test to

evaluate the fire performance of an arbitrary existing or proposed NPP fire barrier design.

Development of the above concept is dependent on the availability of the proposed specialpurpose NPP-specific fire model. With this in hand, implementation of the above ideas involves

two complementary efforts.

First would be a significant modeling effort involving simulations of a wide variety of

conjectured fire scenarios in at least a few NPPs. The simulations would provide invaluable

information on the general characteristics of real, fire-barrier-threatening, NPP fire

environments. By studying, categorizing, etc. the broad range of these simulations, it is

envisaged that meaningful NPP-specific distinguished groupings of them would be identified.

These would then be used to define the sought-after NPP-specific test fire curves.

After the modeling effort, there would be an experimental study to determine the capabilities

of existing ASTM E 119 test furnaces relative to generating reproducible, temperature-time,

test-fire environments other than that of the standard ASTM E 119 fire. In this regard, for a

given furnace facility the most important characteristics of such achievable furnace fire

environments would be those associated with the maximum (bounding), achievable, temperaturetime curve: the peak temperature; the minimum time-to-peak-temperature; and the actual

temperature-time curve corresponding to the fire initiated and sustained with the maximum

3 That the bounding temperature principle is not universally valid is proven by the following counterexample: Consider a nonreacting unconstrained structural barrier exposed to a fire environment along its

length, on one side, and insulated on the other. Consider two temperature-time fire curves, both of which

start at ambient temperature and reach the same maximum steady temperature. Let one curve grow

linearly with time with an arbitrarily small slope. Let the second one not grow at all until relatively late

in time, when there is a near step-change in temperature to a value somewhat below the first curve. In the

case of the first curve, the temperature will always be effectively uniform through the depth of the

structure, and it will always have an arbitrarily small stress field. For the second curve, a non-uniform

temperature distribution and a non-zero stress field, possibly leading to failure, will develop. Therefore,

in this particular case the bounding-temperature principle is not valid, since the structural response in the

second scenario is more severe than that of the first even though the fire curve of the first bounds that of

the second.

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possible supply rate of fuel. gas. It is anticipated that such a study of existing test furnaces

would lead to the identification of a limited number of easily reproduced (from facility-tofacility) test fire curves having a wide range of "severity." (In the sense of the boundingtemperature criterion, a set of curves having a wide range of "severity" would be associated

with a set of curves having a wide range of "bounding capability." A reasonable measure of

a curve's bounding capability would be its peak temperature and time-to-peak-temperature. )

These test fire curves would be selected for their relevance to the previously-determined, NPPspecific, simulated-fire groupings.

When evaluating the fire performance of NPP fire barriers, multiple NPP-specific temperaturetime fire curves determined according to the above ideas would replace the standard ASTM

E 119 fire curve.

3.9.4 NPP-Specific Direct Fire Exposures

As is the case with indirect fire exposure threats, when evaluating the performance of a particular barrier component relative to sustained direct fire exposure, inordinate numbers of

possible NPP facility fire scenarios would lead to inordinate numbers of different possible direct

exposure threats. While these could be calculated with the proposed advanced fire model, it is

again clear that implementing the multitudes of corresponding direct exposure simulation tests

that could be devised, in principal, would not be practical. Rather, a presumably conservative

"Bounding Direct-Fire-Exposure" Principle, corresponding to the Bounding-Temperature

Principle, would be adopted, and applied. For example, in cases where a barrier design can be

subjected to actual sustained flame exposure, perhaps ASTM E 1529 test exposures, applied for

a calculated time interval, say the predicted maximum possible duration of the considered fire

scenarios, could be shown to be a valid basis for such a principle.

3.10 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS

3.10.1 General Summary and Conclusions

Part 3 of this document assessed the feasibility of developing and implementing Nuclear Power

Plant (NPP)-specific descriptions of fire environments for use in evaluating the fire resistance

of fire barriers. These include structural barriers, useful in isolating a compartment of fire

origin from adjacent spaces, and wrap assemblies, used to isolate and protect Plant equipment,

cables, etc., within a compartment of fire origin, from the effects of exposure to the fire

environment.

A discussion on the nature of fire barrier exposure to the compartment fire environment

distinguished between direct exposure to the most extreme zones of the fire environment, e.g.,

direct, sustained exposure of the barrier to the flame, and indirect exposure, where the a fire

barrier is mainly exposed to the average properties of the overall fire environment. The

problem of evaluating the integrity of fire barriers to both kinds of threats was addressed.

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The problem of defining the fire exposure with which to evaluate fire barrier performance was

discussed. This included ASTM E 119-type test methods (i.e., exposure of full-scale barrier

specimens to specified furnace environments that simulate indirect-type fire exposures), the

need to develop other tests that would be devised to simulate direct exposure threats, and the

applicability of compartment fire model simulations.

A methodology for evaluating the fire performance of NPP fire barriers was presented that

removes weaknesses of and/or introduces flexibility to the traditional ASTM E 119 approach.

This would rely on a combined experimental and analytic approach that involved the BoundingTemperature Principle. (I.e., if the temperature-time curve of one fire environment bounds that

of another, then, relative to the threat to structural integrity of a NPP fire barrier, the boundingcurve environment is the more severe.) Experiments would involve ASTM E 1 19-type tests,

where the standard ASTM E 119 fire is replaced by alternative temperature-time furnace fires,

and where the alternative fires would be deduced from reliable fire model simulations. Analysis

would involve compartment fire modeling methods, where computer simulations would be

carried out with a new, advanced, special-purpose, zone-type fire model. This model would be

developed to include features particularly relevant to simulating fire environments that threaten

NPP fire barriers, from the point of view of both direct and indirect fire exposure.

Based on a review of the literature of NPP-specific combustibles and previously-developed NPPspecific fire models, special features required of the new fire model were proposed. These

included: the simulation of fully-developed burning of extensive dense arrays of cable trays,

both under fuel-controlled and ventilation- controlled conditions; the simulation of

combustible/flammable liquid pool fires; and advanced means of modeling ventilation and

radiation-heat-transfer-related phenomena. It was proposed that a new special-purpose model

with these features be developed as a customized advanced version of an existing, two-layer,

multi-room, zone-type fire model.

The new model would be used to simulate a wide variety of potential fire scenarios in rooms of

fire origin of selected NPPs. The simulations would lead to new insights on the characteristics

of real, fire-barrier-threatening, NPP fire environments. Based on applications of the BoundingTemperature Principle, the simulated fire scenarios would lead to a series of NPP-specific test

fire curves covering a wide range of NPP-type fire severities. An experimental study on

available ASTM E 119-type test furnaces would be carried out to establish that these new test

fire curves (instead of the standard ASTM E 119 fire curve) can be used in ASTM E 119-type

barrier rating tests. Then ASTM E 119-type tests using the new NPP-specific test fire curves

would be established as the method of evaluating the fire performance of NPP fire barriers.

A significant effort will be required to carry out this plan. Due to knowledge gaps in critical

areas such as burning rate and ventilation effects in NPP-specific environments, the modeling

work will require a substantial experimental component. Indeed, the experimental aspects,

including full-scale burnout of fuel packages and furnace fire-resistance tests, are similar in

scope to the NBS recreation-room study (see Part 2, section 2.3.3), which was a multi-year

effort.

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3.10.2 Recommended Tasks

Consistent with the above, it is recommended that the following tasks be carried out with the

goal of establishing a reliable methodology for evaluating NPP fire barrier performance:

1. Develop a new, special-purpose, NPP-specific fully-developed fire model capable of

simulating fire environments that threaten NPP fire barriers. It is recommended that this

be developed as an advanced version of an existing multi-room compartment fire model,

e.g., CFAST [3-25]. The new model should include the advanced modeling features

identified in the section 3.7.3 "Features of a Compartment Fire Model Suitable For

Evaluating Direct and Indirect Threats to NPP Fire Barriers." These include: the

simulation of fully-developed burning of extensive dense arrays of cable trays, both

under fuel-controlled and ventilation-controlled conditions; the simulation of

combustible/flammable liquid pool fires; the simulation of the fire environment in multiroom facilities (at least two adjacent spaces); and advanced means of modeling

ventilation and radiation-heat-transfer-related phenomena

2. Carry out full-scale experimental verification of the advanced modeling methods of

item 1, especially those aspects of the new model associated with the simulation of

burning cable trays and combustible/flammable fuel fires in enclosed spaces. Also, carry

out experiments to better evaluate and characterize the fire hazard in NPPs introduced

by electrical panels/cabinets.

3. Use new model simulations to determine the direct-exposure threat to fire barriers, and

use these to establish experimental methods to evaluate barrier fire performance relative

to the direct exposure threat.

4. Use the new model to carry out an extensive simulation study of selected NPPs. Results

of this would be used to establish the characteristics of real, fire-barrier-threatening,

NPP fire environments and to identify a series of NPP-specific test fire curves to replace

the ASTM E 119 standard fire curve.

5. Carry out an experimental study on available ASTM E 119-type test furnaces to establish

that the new test fire curves of item 4 are attainable and reproducible.

6. Use the results of items 4 and 5 to establish an ASTM E 119-type method of evaluating

the performance of structural fire barriers relative to the indirect exposure; establish

corresponding methods for wrap-assembly fire barriers.

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3.11 NOMENCLATURE

AJEx total exposed surface area in a crib or in an array of loaded cable

trays.

CAU.Ay constant of cable array design.

CCRM Eq. (1).

CSM crib-material-dependent constant, Eq. (1).

D characteristic dimension of a crib's "stick" section.

L heat of gasification of the fuel.

Lr length of cable tray.

LoTAL total length of all cable trays in the array.

lhk"Ay mass loss rate of array of cable trays.

rh0 , mass loss rate of array of crib.

1hP0oL net mass loss rate of fuel at the pool surface.

QE rate of heat transfer from external sources (the gas layer and

bounding compartment surfaces).

QF rate of heat transfer from the flame.

QL rate of heat transfer lost through the pool surface.

QPOOL net rate of heat transfer absorbed at the pool surface.

WT width of cable tray.

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3.12 REFERENCES

[3-1] Standard Test Methods for Determining Effects of Large Hydrocarbon Pool Fires on

Structural Members and Assemblies, ASTM E 1529 - 93, American Society for Testing

and Materials, Philadelphia.

[3-2] Zwiers, R.I. and Morgan, B.J., Performance of Concrete Members Subjected to Large

Hydrocarbon Pool Fires, PCI Journal, pp. 120-135, 1989.

[3-3] Fire Induced Vulnerability Evaluation Methodology (FIVE) Plant Screening Guide,

Electric Power Research Institute, Palo Alto CA, 1992.

[3-4] Standard Methods of Fire Test of Building Construction and Materials, ASTM E 119,

American Society for Testing and Materials, Philadelphia.

[3-5] Thermo-Lag 330-1 Combustibility Evaluation Methodology Plant Screening Guide,

prepared for the Nuclear Management and Resources Council by Stone and Webster

Engineering Corporation, 1993.

[3-6] ISO, Fire Resistance Tests - Elements of Building Construction, International

Organization FOR Standardization, ISO-834 Part 1, Switzerland, 1992., Geneva.

[3-7] Babrauskas, V. and Williamson, R.B., The Historical Basis of Fire Resistance Testing

-Part I (pp. 184-194) and Part II (pp. 304-316), Fire Technology, 14, 1978.

[3-8] Generic Letter (GL) 86-10, Supplement 1: Fire Endurance Test Acceptance Criteria for

Fire Barriers Used to Separate Redundant Safe Shutdown Trains Within the same Fire

Area, Enclosure 1 of March 25, 1994 letter from NRC to All Holders of Operating

Licenses or Construction Permits for Nuclear Power Reactors, Nuclear Regulatory

Commission, Washington.

[3-9] Standard Test Methods for Fire Tests of Fire-Resistive-Barrier Systems for Electrical

System Components: Draft 19, Proceedings of the Meeting of the E5.11.08 SubCommittee on Cable Tray Protection, Phoenix, December 14, 1995, American Society

for Testing and Materials, Philadelphia.

[3-10] Proposed Requirements and Proposed Effective Date for the First Edition of the Standard

for Tests of Thermal Barrier Systems for Electrical System Components, UL 1724,

Appendix A of March 24, 1995 UL Memorandum from C.H. McCall to Electric and

Fire Councils of UL and to Subscribers to UL's Classification Service, Underwriters

Laboratory, Northbrook.

[3-11] Cooper, L.Y., Compartment Fire-Generated Environment and Smoke Filling, Section

2/Chapter 7, SFPE Handbook of Fire Protection Engineering, SFPE, Boston, 1988.

NUREG-1547 94
NPP-Specific Descriptions of Fire Environments

[3-12] Cooper, L.Y., Smoke Movement in Rooms of Fire Involvement and Adjacent Spaces,

Fire Safety Journal, Vol. 7, pp. 33-46, 1984.

[3-13] Ho, V., Siu, N., and Apostolakis, G., COMPBRN III - A Fire Hazard Model for Risk

Analysis, Fire Safety Journal, Vol. 13, pp. 137-154, 1988.

[3-14] Berry, D.L., Analysis of Fire Barriers Within Nuclear Power Plants, Nuclear

Technology, 53, pp. 204-216, 1981.

[3-15] Sullivan, P.J.E, Terro, M.J., and Morris, W.A., Critical Review of Fire-Dedicated

Thermal and Structural Computer Programs, Journal of Applied Fire Science, Vol. 3(2),

pp. 113-135, 1993-94.

[3-16] Klamerus, L.J., Cable Tray Fire Test, Proceedings of the Winter Meeting of the IEEE

Power Engineering Society, New York NY, Feb. 4-9 1979.

[3-17] Chavez, J.M. and Klamerus, L.J., Cable Tray Fire Experimentation, Proceedings of the

CSNI Specialist Meeting on Interaction of Fire and Explosion with Ventilation Systems

in Nuclear Facilities, LA-9911-0, Vol. 1, Los Alamos National Laboratory, Los Alamos

NM 87545, pp. 375-389, 1983.

[3-18] Hill, J.P., Fire Tests in Ventilated Rooms - Extinguishmen of Fire in Grouped Cable

Trays, Factory Mutual Research Corp. interim report (NP-2660, Research Project 1165-

1) prepared for EPRI (EPRI Report NP-2660), Dec. 1983.

[3-19] Chavez, J.M., An Experimental Investigation of Internally Ignited Fires in Nuclear

Power Plant Control Cabinets: Part I: Cabinet Effects Tests, Report SAND86-0336 for

U.S. NRC, Sandia National Lab., April 1987.

[3-20] Chavez, J.M. and Nowlen, S.P., An Experimental Investigation of Internally Ignited

Fires in Nuclear Power Plant Control Cabinets: Part H: Room Effects Tests, Report

SAND86-0336 Vol. 2 for U.S. NRC, Sandia National Lab., Nov. 1988.

[3-21] Nowlen, S.P., A Summary of Nuclear Power Plant Fire Safety Research at Sandia

National Laboratories, 1975 - 1987, Report SAND89-1359 for U.S. NRC, Sandia

National Lab., Dec. 1989.

[3-22] Cooper, L.Y. and Forney G.P., The Consolidated Compartment Fire Model (CCFM)

Computer Code Application CCFM.VENTS - Part I: Physical Basis, NISTIR 90-4342,

National Institute of Standards and Technology, Gaithersburg MD, 1990.

[3-23] Cooper, L.Y. and Forney G.P., Eds., The Consolidated Compartment Fire Model

(CCFM) Computer Code Application CCFM.VENTS - Part Ill: Catalog of Algorithms

and Subroutines,, NISTIR 90-4344, National Institute of Standards and Technology, Gaithers-burg MD, 1990.

95 NUREG-1547
NPP-Specific Descriptions of Fire Environments

[3-24] Cooper, L.Y., A Note on Calculating Flows Through Vertical Vents in Zone Fire

Models Under Conditions of Arbitrary Cross-Vent Pressure Difference, Combustion

Science and Technology, 63, 1-3, pp. 43-50, 1989.

[3-25] Peacock, R.D. et al, CFAST, the Consolidated Model of Fire Growth and Smoke

Transport, NIST TN 1299, National Institute of Standards and Technology,

Gaithersburg, 1993.

[3-26] Cooper, L.Y., Combined Buoyancy- and Pressure-Driven Flow Through a Horizontal,

Circular Vent, Journal of Heat Transfer,Vol. 117, pp. 659-667, August 1995.

[3-271 Cooper, L.Y., VENTCF2: An Algorithm and Associated Computer Subroutine for

Calculating Flow Through a Horizontal Ceiling/Floor Vent in a Zone-Type

Compartment Fire Model, NISTIR 5470, National Institute of Standards and

Technology, Gaithersburg MD, August 1994.

[3-28] Klote, J.H. and Cooper, L.Y., Model of a Simple Fan-Resistance Ventilation System

and Its Application to Fire Modeling, NISTIR 89-4141, National Institute of Standards

and Technology, Gaithersburg, September 1989.

[3-29] Yamada, T. and Cooper, L.Y., Surface Heat Transfer Algorithm, Summaries of Center

for Fire Research In-House Projects and Grants - 1990, NISTIR 4440, Cherry, S.M.,

Ed., National Institute of Standards and Technology, Gaithersburg, 1990.

[3-30] Forney, G.P., Computing Radiative Heat Transfer Occurring in a Zone Fire Model,

NISTIR 4709, National Institute of Standards and Technology, Gaithersburg, 1991.

[3-31] Not used.

[3-32] Babrauskas, V., Burning Rates, Section 2/Chapter 1, SFPE Handbook of Fire Protection

Engineering, Society for Fire Protection Engineers, Boston, 1988.

[3-33] Klamerus, L.J., A Preliminary Report on Fire Protection Research Program (July 6,

1977 Test)," SAND77-1424, October 1977.

[3-34] Harmathy, T.Z., Experimental Study on the Effect of Ventilation on the Burning of Piles

of Solid Fuels, Combustion and Flame, Vol. 31, pp. 259-264, 1978.

[3-35] Drysdale, D., An Introduction to Fire Dynamics, Wiley, 1985.

[3-36] Tewarson, A. and Pion, R.F., Flammability of Plastics. I. Burning Intensity,

Combustion and Flame, Vol. 26, pp. 85-103, 1976.

NUREG-1547 96
NPP-Specific Descriptions of Fire Environments

[3-37] Bullen, M.L, A Combined Overall and Surface Energy Balance for Fully-Developed

Ventilation-Controlled Liquid Fuel Fires in Compartments, Fire Research Journal, Vol.

1, pp. 171-185, 1977/78.

[3-38] Bullen, M.L. and Thomas, P.H., Compartment Fires with non-Cellulosic Fuels, lIM

Symposium (International) on Combustion, Combustion Institute, Pittsburgh, pp. 1139-

1148, 1978.

97 NUREG-1547
NPP-Specific Descriptions of Fire Environments

Table 3-1. A Preliminary Combustible Load Summary of the Spaces at the Watts Bar

NPP Having the Largest Density of Combustibles.

Building/Room Area Avg. Density Primary

Name of (Secondary)

Combustibles Contributor

Control/Spreading

Room

Diesel Generator

Lube Oil Storage

Auxiliary/Mechanical

Equipment

Control/Shift Eng.

Office

Auxiliary/480-V

Board Room

823 m2 601x101 kJ/m2

38 m2 484xl07 j/rn 2

65 m2 346x10 7 kJ/m 2

29 m2 286x10 7 kj/m2

200 m2 277x107 kJ/m2

cable trays

(panels/cabinets, < 10%)

flammable liquid

(miscellaneous, < 1%)

cable trays

(miscellaneous, < 1%)

miscellaneous

cable trays

(panels/cabinets, < 10%)

NUREG-1547 98
NPP-Specific Descriptions of Fire Environments

Inddirect fire exposures of:

unprotected cable trays direct fire exposu

of ceiling structu harre

ire

wail structural

wrAp-assembfy-p roteted barrier

cable trays

flammable liquid pood fire

FIgure 3-1. NPP compartment fire involving fully-developed burning of an extensive

dense array of cable trays.

99 99 NUREG-1547
NPP-Specific Descriptions of Fire Environments

fully-developed burning

in the crib-like cable tray fuel array

Indirect fire exposures of:.

unprotected cable trays.

wail structural Xbarrier

wvapassembly-protected

cable trays

Figure 3-2. NPP compartment fire involving burning of a combustible/flammable liquid.

NUREG-1547 100
.NPP-Specific Descriptions of Fire Environments

direct fire exosues of:

unprotected cable trays

Indimet fire exMosures of:

wrap - assembly - protected cable trays /

unprotected cable trays

wall structural

Vy -rotetedbarrier

flammable liquid pool fire /

Figure 3-3. NPP compartment fire involving burning of a combustible/flammable liquid

and direct fire exposure of unprotected and wrap-assembly-protected cable

trays and of the compartment ceiling structure.

101 NUREG-1547
NPP-Specific Descriptions of Fire Environments

FIgure 3-4. Multi-room NPP compartment fire involving fully-developed burning of an

extensive dense array of cable trays and a combination of different types of

ventilation configurations.

NUREG-1547 102
APPENDIX A

POLICY ISSUE COMMISSION INFORMATION PAPER:

NUCLEAR POWER PLANT-SPECIFIC TIME-TEMPERATURE CURVES

FOR TESTINGAND QUALIFYING FIRE BARRIERS

POLICY ISSUE

(Information)

July 19, 1996 SECY-96-162

FOR: The Commissioners

FROM: James M. Taylor

Executive Director for Operations

SUBJECT: NUCLEAR POWER PLANT-SPECIFIC TIME-TEMPERATURE CURVES FOR

TESTING AND QUALIFYING FIRE BARRIERS

PURPOSE:

To inform the Commission that the U.S. Nuclear Regulatory Commission (NRC)

staff (1) has completed a feasibility study regarding the development of

nuclear power plant specific time-temperature curves for testing and

qualifying fire barriers and (2) will provide the study to the Nuclear Energy

Institute (NEI) for industry consideration.

BACKGROUND:

On May 12, 1994, the NRC staff submitted SECY-94-127, "Options for Resolving

the Thermo-Lag Fire Barrier Issues." In that paper, the staff suggested that

a possible way to help resolve the Thermo-Lag fire barrier issues was to

qualify Thermo-Lag fire barriers on the basis of representative nuclear power

plant fire hazards rather than the time-temperature curves specified in

industry fire endurance test standards. The staff proposed that it

Investigate the feasibility of developing alternative nuclear power

plant-specific time-temperature curves for qualifying nuclear power plant fire

barriers. The staff stated that if it found that such curves were feasible,

NOTE: TO BE MADE PUBLICLY AVAILABLE IN 5

WORKING DAYS FROM THE DATE OF THIS PAPER

Contact: Edward Connell, HRR

301-415-2838

105 NUREG-1547
The Commissioners - 2-

the curves should be developed by industry. The Commission, in a Staff

Requirements Memorandum (SRI4) of June 27, 1994, directed the staff to evaluate

the feasibility of qualifying fire barriers on the basis of representative plant fire hazards. The Commission also stated that the responsibility for

developing alternative curves should rest with industry.

DISCUSSION:

The NRC staff obtained technical assistance from the National Institute of

Standards and Technology (NIST), Building and Fire Research Laboratory. NIST

documented the study in a report entitled ' Methodology for Developing and

Implementing Alternative Temperature-Time Curves for Testing the Fire

Resistance of Barriers for Nuclear Power Plant Applications," June 1996. In

its report, NIST documented the history of fire endurance testing, summarized

other attempts to develop alternative fire exposures, and proposed a

methodology for developing technically sound alternative nuclear power plant- specific time-temperature fire test curves. These alternative fire exposures would be based on the representative nuclear power plant fire hazards (plus margin) and state-of-the-art fire science methods would be used.

On the basis of its work with NIST and its review of NIST's report, the staff

concluded that it is technically feasible to develop nuclear power plant specific time-temperature curves that can be used to test and qualify fire

barriers. The approach recommended by NIST would involve identifying and

characterizing representative plant fire hazards, modifying an existing

computer fire model, and developing and validating fire endurance testing methods. *The fire modeling would require a full-scale experimental fire test

program to provide the data and information needed to revise, verify, and

validate the model. After the alternative time-temperature curves and fire

endurance testing methods are developed, any nuclear power plant fire barrier

that a licensee proposes to qualify using an alternative nuclear power

plant-specific time-temperature curve would need to be tested against the

alternative curve. This effort would require significant financial and human

resources. On the bases of discussions with NIST and its experience with fire

test programs and fire modeling, the staff estimates that if industry undertook this effort, it could take 3 to 5 years to complete.

In accordance with SECY-94-127 and the SRN of June 27, 1994, the staff will

provide the NIST report to NEI for industry consideration. In view of the

time and effort that would be required to develop nuclear power plant-specific time-temperature curves and to test the Thermo-Lag fire barriers against the

curves, the staff has concluded that such curves are not a viable approach for

resolving the Thermo-Lag issues within licensees' schedular commitments.

However, an approach using alternative time-temperature curves may be useful

to licensees that plan to use performance-based methods in the future for

other purposes.

In a letter to the Chairman of March 15, 1996, the Advisory Committee on

Reactor Safeguards (ACRS) stated that it had been told that alternitive

time-temperature curves had been produced by the insurance industry. As part

NIJRB -1547 106
The Commissioners -3 -

of its work, NIST reviewed the literature for other efforts to develop alternative curves. The alternate curves mentioned in the ACRS letter were

not included in the available literature. After it received the ACRS letter,

the staff discussed the availability of alternative curves with its contacts

in the nuclear insurance industry. These contacts were not familiar with such

work. The staff has informed the ACRS staff of this and has requested that

the ACRS provide specific information on the availability of alternative

time-temperature curves from the insurance industry. If the staff finds

information on alternative time-temperature curves that is publicly available

and germane to the work discussed here, it will provide this information to

NEI.

es . Tlor

cutive Director

for Operations

DISTRIBUTION:

Commissioners

OGC

OCA

ACRS

ASLBP

EDO

SECY

107 NUREG-1547

NRC FORM 335 UI. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

P(Assigned by NRC. Add Vol.. Supp, Rev.

=1.C3110 BIBUOGRAPHIC DATA SHEET 3201. 5202

(See . .. .. Cf on . N U R E G -1547

2. TITLE AND SUBTITLE

Methodology for Developing and Implementing Alternative Temperature-Time Curve for Testing -

the Fire Resistance of Barriers for Nuclear Power Plant Applications MONTH YED

August 1996

4. FIN OR GRANT NUMBER

S. AUTOR(M) 6. TYPE OF REPORT

LY. Cooper, K.D. Steckler, NIST

7. PERIOO COVERED &ftsqx•,Dtefs)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (ArNRC, pvovie hkoce woRegwio, U.S. Abdow rReWula, Commission, and maft edds=s feoranato,

p'mvie name d mailing addus&

National Institute of Standards and Technology

Technology Administration

U.S. Department of Commerce

Gaithersburg, MD 20899

9. SPONSORING ORGANIZATION - NAME AND ADDRESS trNR,qn type Zerme as abon', I condrec pvd ARC &Wivhy Ofre or R~egion U. & Mie ftW•r CVrh1,

and maAfg sd*-us.

Division of Systems Safety and Analysis

Office of Nuclear Reactor Regulation

U.S. Nuclear Regulatory Commission

Washington. DC 20555-0001

10. SUPPLEMENTARY NOTES

11. ABSTRACT a00 wads ark=)

Advances in fire science over the past 40 years have offered the potenti for developing

technically sound alternative temperature-time curves for use in evaluating fire barriers for areas

where fire exposures can be expected to be significantly different than the ASTM E-1 19 standard

temperature-time exposure. This report summarizes the development of the ASTM E-119,

standard temperature-timecurve, and the efforts by the federal government and the petrochemical

industry to develop alternative fire endurance curves for specific applications. The report also

provides a framework for the development of alternative curves for application at nuclear power

plants. The staff has concluded that in view of the effort necessary for the development of nuclear

power plant specific temlperature-time curves, such curves are not a viable approach for resolving

the issues concerning Thermo-Lag fire barriers. However, the approach may be useful to licensees

in the development of performance-based fire protection methods in the future.

12. KEY WORDSIDESCRIPTORS g(s wods arphrases Utaf wXaS3W eaamher In /=&V be JPrtj 1& AVALABILITY STATEMENT

ASTM El19, cables, fire barriers, fire endurance, fire models, fire resistance, histories, nuclear power unlimited

plants, temperature, zone models 14. SE=-f ro.ASSlFlAn"

unclassified

M93~ Repo;V

unclassified

15. NUMBER OF PAGES

16. PRICE

NR FRM33 IIIII lilim II icoaal IIdue In EteFdrlon IsImo

NRC FORM 335 CUM TWs form was elecuonscafly produced by Elb Federal Forrm, Ino.