Report of Experimental Results for the International 

Fire Model Benchmarking and Validation Exercise #3 

Anthony Hamins 

Alexander Maranghides 

Rik Johnsson 

Michelle Donnelly 

Jiann Yang 

George Mulholland 

 Robert L. Anleitner 

Building and Fire Research Laboratory

November 2003 

revised June 2005 

 NIST Special Publication 1013-1 

U.S. Department of Commerce

Donald L. Evans, Secretary 

Technology Administration

Michelle O’Neill, Acting Under Secretary for Technology

National Institute of Standards and Technology

Hratch Smerjian, Acting Director 

ii

NIST Special Publication 1013-1 

Report of Experimental Results for the International 

Fire Model Benchmarking and Validation Exercise #3 

Anthony Hamins 

Alexander Maranghides 

Rik Johnsson 

Michelle Donnelly 

Jiann Yang 

George Mulholland 

 Robert L. Anleitner 

 Building and Fire Research Laboratory
iii

Executive Summary 

National Institute of Standards and Technology Special Publication 1013-1 

Natl. Inst. Stand. Technol. Spec. Publ. 1013-1, 175 pages (June 2005) 

CODEN: NSPUE2 

U.S. GOVERNMENT PRINTING OFFICE 

WASHINGTON: 2003

_________________________________________ 

For sale by the Superintendent of Documents, U.S. Government Printing Office 

Internet: bookstore.gpo.gov — Phone: (202) 512-1800 — Fax: (202) 512-2250 

Mail: Stop SSOP, Washington, DC 20402-0001

Certain commercial entities, equipment, or materials may be identified in this 

document in order to describe an experimental procedure or concept adequately. Such 

identification is not intended to imply recommendation or endorsement by the 

National Institute of Standards and Technology, nor is it intended to imply that the 

entities, materials, or equipment are necessarily the best available for the purpose. 
iv

NUREG/CR-6905 

NIST SP 1013-1

Report of Experimental Results for the International 

Fire Model Benchmarking and Validation Exercise #3 

Manuscript Completed: November 2003 

Date Published: May 2006 

Prepared by 

A. Hamins, A. Maranghides, 

R. Johnsson, M. Donnelly, J. Yang, 

G. Mulholland, R. L. Anleitner 

National Institute of Standards and Technology 

Building and Fire Research Laboratory 

100 Bureau Drive 

Gaithersburg, MD 20899 

J. Dreisbach, NRC Project Manager 

Prepared for: 

Division of Risk Assessment & Special Projects 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001 

NRC Job Code Y6688 
v

Abstract 

The U.S. Nuclear Regulatory Commission (NRC) and the National Institute of Standards and Technology 

(NIST) are participating in an International Collaborative Fire Model Project (ICFMP) to assess 

and validate fire computer codes for nuclear power plant applications. This assessment is being 

conducted through a series of benchmarking and validation exercises. The goal of the present 

study is to provide data from a large-scale fire test series of a simulated nuclear power plant 

compartment conducted in summer 2003. 

The experiments consisted of hydrocarbon spray fires varying from 350 kW to 2 MW in heat 

release rate, burning in a single compartment 7 m (23 ft) wide, 22 m (72 ft) long, and 4 m (13 ft) 

tall. The experimental results were composed of 15 tests with more than 370 channels of data 

per test. 

This report includes a description of the measurement methods and experimental results for tests 

conducted as part of the International Fire Model Benchmarking and Validation Exercise #3. 

The experimental data are available electronically to accompany this report. The data can be 

used as a comparison with fire models as part of the international benchmark exercise. 
vi

FOREWORD 

As part of its Fire Risk Research Program, the U.S. Nuclear Regulatory Commission (NRC) 

collaborates with the National Institute of Standards and Technology (NIST) in a broader 

International Collaborative Fire Model Project (ICFMP) to assess and validate computer codes 

for fire modeling of nuclear power plant (NPP) applications. The aim of this collaborative project 

is to share information and resources, and jointly develop reports documenting the assessment 

and validation of fire modeling computer codes. Toward that end, the NRC, in collaboration with 

NIST, conducted realistic full-scale fire tests simulating a typical NPP compartment. This report 

provides data from these full-scale fire tests. 

The NRC uses fire modeling and fire dynamics calculations in a number of regulatory applications. 

Licensees use these models and calculations to demonstrate compliance with, and to request 

exemptions to, the regulatory requirements for fire protection. The NRC uses them in the 

Significance Determination Process (SDP), a part of the NRC’s inspection program. In addition, 

the NRC uses them in the risk-informed performance-based (RI/PB) voluntary fire protection 

licensing requirements established under 10 CFR 50.48(c). The RI/PB method is based on the 

National Fire Protection Association (NFPA) Standard 805, “Performance-Based Standard for 

Fire Protection for Light-Water Reactor Generating Plants.” 

This report provides data and technical documentation that are referenced and included in draft NUREG1824, “Verification and Validation of Selected Fire Models for Nuclear Power Plant Applications,” 

which was published for public comment in January 2006. This report documents some of the 

experimental data used in draft NUREG-1824, including experimental uncertainty estimates. 

The results presented in this report were obtained from experiments using hydrocarbon spray fires 

varying from about 350 kW to 2 MW in heat release rate (HRR), burning in a single compartment 

that was 7 m (23 ft) wide, 22 m (72 ft) long, and 4 m (13 ft) tall. The experimental results include 

more than 370 channels of data for each of the 15 tests conducted as part of this test series. This research 

provides valuable and extensive test data needed to assess the predictive capabilities of fire modeling 

tools, which may be used to evaluate fire hazards in NPPs and make regulatory decisions. 

 _________________________________

 Carl J. Paperiello, Director 

 Office of Nuclear Regulatory Research 

 U.S. Nuclear Regulatory Commission
vii 

TABLE OF CONTENTS 

ABSTRACT.....................................................................................................................................v 

FOREWORD ................................................................................................................................. vi 

TABLE OF CONTENTS.............................................................................................................. vii 

LIST OF FIGURES ....................................................................................................................... ix 

LIST OF TABLES....................................................................................................................... xiv 

ACRONYMS............................................................................................................................... xvi 

1 Introduction...................................................................................................................................1 

2 Experimental Configuration and Conditions ................................................................................3 

2.1 Test Compartment..................................................................................................................3 

2.2 Compartment Contents ..........................................................................................................3 

2.3 Compartment Openings and Ventilation ...............................................................................7 

2.4 Fuel Delivery System and Fire ..............................................................................................8 

2.5 Background Documentation ..................................................................................................8 

3 Heat Release Rate .......................................................................................................................13 

3.1 Heat Release Rate Based on the Fuel Flow .........................................................................13 

3.2 Heat Release Rate Based on Oxygen Consumption Calorimetry........................................17 

3.3 Energy Balance in the Compartment ...................................................................................25 

3.4 Summary..............................................................................................................................26 

4 Heat Flux Measurements ............................................................................................................29 

4.1 Heat Flux Measurement Devices.........................................................................................29 

4.2 Uncertainty Analysis............................................................................................................30 

4.3 Results and Discussion ........................................................................................................33 

5 Smoke Concentration..................................................................................................................39 

5.1 Principle of Operation..........................................................................................................39 

5.2 Design ..................................................................................................................................39 

5.3 Uncertainty...........................................................................................................................40 

5.4 Temperature Uncertainty .....................................................................................................41 

5.5 Results..................................................................................................................................41 

6 Target Temperatures ...................................................................................................................47 

6.1 Description of Measurements ..............................................................................................47 

6.2 Results and Discussion ........................................................................................................50 

7 Vent and Doorway Flows ...........................................................................................................55 

7.1 Doorway Flows....................................................................................................................55 

7.2 Velocities and Mass Flow Through the Doorway ...............................................................56 

7.3 Heat Loss Through the Doorway.........................................................................................61 

7.4 Ventilation Supply and Exhaust...........................................................................................62 

8 Gaseous Sampling.......................................................................................................................73 

8.1 Apparatus .............................................................................................................................73 

8.2 Calibration Procedures.........................................................................................................75 

8.3 Sample Transport Time and Analyzer Response Characterization .....................................75 

8.4 Estimated Expanded Uncertainty.........................................................................................75 

8.5 Results and Discussion ........................................................................................................76 

9 Gas Temperatures .......................................................................................................................95 

9.1 Measurement Uncertainties .................................................................................................95 
viii 

9.2 Hot Gas Layer Temperature...............................................................................................100 

9.3 Hot Gas Layer Depth .........................................................................................................106 

9.4 Gas Temperature at Targets B & E....................................................................................111 

9.5 Gas Temperature at Target G.............................................................................................115 

10 Heat Loss to the Enclosure .....................................................................................................117 

10.1 Introduction......................................................................................................................117 

10.2 Measurement Design .......................................................................................................117 

10.3 Instrumentation ................................................................................................................118 

10.4 Measurement Uncertainties .............................................................................................121 

10.5 Results..............................................................................................................................123 

10.6 Contact Resistance Between Marinite Boards.................................................................128 

11 Compartment Pressure............................................................................................................133 

12 Summary and Conclusions .....................................................................................................137 

13 Acknowledgements.................................................................................................................139 

14 References...............................................................................................................................141 

A Material and Optical Properties................................................................................................143 

A.1 Cables................................................................................................................................143 

A.2 Properties of Gypsum, Marinite, Feraloy, and the PVC Slab E........................................145 

B Ambient Humidity and Temperature during Testing ...............................................................147 

C Compartment Leakage Area.....................................................................................................149 

D Format of the Electronic Data..................................................................................................151 

E Inoperative Instruments during Test Series..............................................................................161 

F Response to Questions and Comments from July WPI international peer review Meeting ....165 

G Curve for Ventilation Supply Fan............................................................................................169 

H Junction Box ............................................................................................................................171 

I Fuel Flow Correction.................................................................................................................173 

J Documentation of Personal Communications...........................................................................175 
ix 

LIST OF FIGURES 

Figure 2-1: Photo of the test compartment being instrumented before testing............................... 4 

Figure 2-2: Compartment Isometric with thermocouple trees........................................................ 9 

Figure 2-3: Compartment Contents and Selected Instrumentation............................................... 10 

Figure 2-4: Plan View of Compartment Mid-Section................................................................... 11 

Figure 3-1: The calorimetric HRR data as a function of time during the calibration 

before Test 5 (time <0) and during Test 5 (time >0). ........................................................... 20 

Figure 3-2: The calorimetric HRR during the open door experiments (Tests 3, 5, 

9, 14, 15, and 18) as a function of time................................................................................. 22 

Figure 3-3: The measured HRR as a function of time for Tests 3 and 5. The solid 

line shows the idealized fire HRR using gas flow rate. ........................................................ 24 

Figure 3-4: The rate of heat loss to surfaces, through the doorway, accumulated 

by heating the gas in the compartment, and their sum as a function of time 

during Test 3. ........................................................................................................................ 26 

Figure 4-1: Photograph of two total (Gauges 4 and 6) and two radiative (Gauges 3 

and 5) heat flux gauges partially insulated............................................................................ 31 

Figure 4-2: Photograph of Gauges 9 and 10 and the vertical cable tray....................................... 31 

Figure 4-3: Cooling water system for total heat flux gauges........................................................ 32 

Figure 4-4: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 2........................ 34 

Figure 4-5: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 3........................ 35 

Figure 4-6: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 4........................ 35 

Figure 4-7: Radiative heat flux as a function of time for Gauge 10 in Test 2. The 

total heat flux Gauge, 9, was not functioning during this test. ............................................. 36 

Figure 4-8: Radiative and total heat flux as a function of time for Gauges 9 and 10 

in Test 3................................................................................................................................. 36 

Figure 4-9: Radiative and total heat flux as a function of time for Gauges 9 and 10 

in Test 4................................................................................................................................. 37 

Figure 4-10: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, 

and 8, in Test 2...................................................................................................................... 37 

Figure 4-11: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, 

and 8, in Test 3 (Gauge 7 failed after 1000 s into the test). .................................................. 38 

Figure 4-12: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, 

and 8, in Test 4...................................................................................................................... 38 

Figure 5-1: Smoke concentration and temperature as a function of time for Test 2. ................... 44 

Figure 5-2: Smoke concentration and temperature as a function of time for Test 3. ................... 44 

Figure 5-3: Smoke concentration and temperature as a function of time for Test 4. ................... 45 

Figure 5-4: Smoke concentration and temperature versus time for a 1 MW (Test 

2) and for a 2 MW (Test 13) blended heptane fire both with the door closed...................... 45 

Figure 5-5: Smoke concentration versus time for toluene (Test 17) and for heptane 

(Test 2) for a 1 MW test with door closed. ........................................................................... 46 

Figure 5-6: Comparison of smoke concentration and smoke temperature for two 

replicate, 1 MW tests with door closed (Test 2 and Test 8).................................................. 46 

Figure 6-1: Thermocouples on the PVC slab (1 cm x 1 cm x 5 cm long) measuring 

temperatures at the lower and upper surfaces and the slab center. ....................................... 47 
x 

Figure 6-2: Photograph looking towards the north wall with Target E present, a 

rectangular slab composed of PVC....................................................................................... 48 

Figure 6-3: Photograph and schematic cutaway of a target cable instrumented 

with bare-bead type K thermocouples on the surface and just under the outer 

insulation............................................................................................................................... 48 

Figure 6-4: The transient temperatures in slab E during Test 2.................................................... 51 

Figure 6-5: The transient temperature on Target B during Test 2. ............................................... 51 

Figure 6-6: The transient temperature on the vertical cable, Target G, during Test 

2............................................................................................................................................. 52 

Figure 6-7: The transient temperature at three locations in slab E during Test 3......................... 52 

Figure 7-1: Locations of instrumentation in the door opening. .................................................... 56 

Figure 7-2: Velocities through the doorway measured at the centerline. ..................................... 57 

Figure 7-3: Velocities through the doorway measured on the south side..................................... 58 

Figure 7-4: Velocities through the doorway measured on the north side..................................... 58 

Figure 7-5: Mass flow into the compartment through the open doorway for Test 3.................... 60 

Figure 7-6: Mass flow out of compartment through the upper portion of the 

doorway for Test 3. ............................................................................................................... 60 

Figure 7-7: Net mass flow through doorway for Test 3................................................................ 61 

Figure 7-8: The energy loss through the doorway during Test 9.................................................. 62 

Figure 7-9: Location of velocity measurements in the supply vent during 

calibration. ............................................................................................................................ 63 

Figure 7-10: Contour plot of supply vent velocities. .................................................................... 65 

Figure 7-11: Flow visualization of the supply duct flow field using smoke (left), 

and using string (right).......................................................................................................... 66 

Figure 7-12: Mass flow through supply vent for NRC Test 4, 1 MW, door closed. .................... 66 

Figure 7-13: Locations where exhaust velocities were measured. ............................................... 67 

Figure 7-14: Contour plot of the exhaust vent velocities.............................................................. 68 

Figure 7-15: Mass flow through exhaust vent for NRC Test 10, 1 MW, door 

closed. ................................................................................................................................... 70 

Figure 7-16: Energy flow through exhaust vent for NRC Test 10, 1 MW, door 

closed. ................................................................................................................................... 70 

Figure 7-17: Energy flow through the exhaust vent for Test 5, compartment door 

open....................................................................................................................................... 71 

Figure 8-1: Lower Level Gaseous System components and configuration. ................................. 74 

Figure 8-2: Test 7 Upper Layer O2. .............................................................................................. 77 

Figure 8-3: Test 7 Lower Layer O2............................................................................................... 77 

Figure 8-4: Test 7 Upper Layer CO2............................................................................................. 78 

Figure 8-5: Test 8 Upper Layer O2. .............................................................................................. 79 

Figure 8-6: Test 8 Upper Layer CO.............................................................................................. 79 

Figure 8-7: Test 8 Upper Layer CO2............................................................................................. 80 

Figure 8-8: Test 2 Lower Layer O2............................................................................................... 80 

Figure 8-9: Test 8 Lower Layer O2............................................................................................... 81 

Figure 8-10: Test 9 Upper Layer O2. ............................................................................................ 82 

Figure 8-11: Test 3 Upper Layer CO2........................................................................................... 82 

Figure 8-12: Test 14 Lower Layer O2........................................................................................... 83 

Figure 8-13: Test 10 Upper Layer O2. .......................................................................................... 84 
xi 

Figure 8-14: Test 4 Upper Layer CO2........................................................................................... 84 

Figure 8-15: Test 4 Lower Layer O2............................................................................................. 85 

Figure 8-16: Test 5 Upper Layer O2. ............................................................................................ 86 

Figure 8-17: Test 5 Upper Layer CO2........................................................................................... 86 

Figure 8-18: Test 13 Upper Layer O2. .......................................................................................... 87 

Figure 8-19: Test 13 Upper Layer CO.......................................................................................... 88 

Figure 8-20: Test 16 Upper Layer CO2......................................................................................... 88 

Figure 8-21: Test 16 Lower Layer O2........................................................................................... 89 

Figure 8-22: Test 18 Upper Layer O2. .......................................................................................... 90 

Figure 8-23: Test 15 Upper Layer CO2......................................................................................... 90 

Figure 8-24: Test 15 Upper Layer CO.......................................................................................... 91 

Figure 8-25: Test 15 Lower Level O2........................................................................................... 91 

Figure 8-26: Test 17 Upper Layer O2. .......................................................................................... 92 

Figure 8-27: Test 17 Upper Layer CO.......................................................................................... 93 

Figure 8-28: Test 17 Upper Layer CO2......................................................................................... 93 

Figure 8-29: Test 17 Lower Layer O2........................................................................................... 94 

Figure 9-1: Temperature comparison as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 

2.0 m on Tree 5 during Test 2............................................................................................... 97 

Figure 9-2: Temperature differences as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 

2.0 m on Tree 5 during Test 2............................................................................................... 97 

Figure 9-3: Temperature comparison as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z=1.05 m and 2.0 

m on Tree 5 during Test 3.................................................................................................... 98 

Figure 9-4: Temperature differences as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 

2.0 m on tree 5 during Test 3. .............................................................................................. 98 

Figure 9-5: Temperature comparison as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 

2.0 m on Tree 5 during Test 4............................................................................................... 99 

Figure 9-6: Temperature differences as a function of time for aspirated (TA5, 

TA6) and bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 

2.0 m on Tree 5 during Test 4............................................................................................... 99 

Figure 9-7: The expanded uncertainty spread around nominal values of hot gas 

layer temperature for all tests except Test 18. .................................................................... 101 

Figure 9-8: Bare bead thermocouple temperatures as a function of time for Tree 7 

during Test 2. ...................................................................................................................... 102 

Figure 9-9: Bare bead thermocouple temperatures as a function of time for Tree 7 

during Test 3. ...................................................................................................................... 103 

Figure 9-10: Bare bead thermocouple temperatures as a function of time for Tree 

7 during Test 4. ................................................................................................................... 103 

Figure 9-11: Bare bead thermocouple temperatures as a function of time for Tree 

7 during Test 5. ................................................................................................................... 104 

Figure 9-12: Hot gas layer temperature as a function of time for Tree 7 during 

Tests 2, 8, 3, 9, 4, 10, and 5 calculated using the two-layer reduction method. ................. 104 
xii 

Figure 9-13: Hot gas layer temperature as a function of time for Tree 7 during 

Tests 1, 7, 2, 8, 13 calculated using the two-layer reduction method................................. 105 

Figure 9-14: Hot gas layer temperature as a function of time for Tree 7 during 

Tests 13, 16, 2, and 17 calculated using the two-layer reduction method. ......................... 105 

Figure 9-15: Hot gas layer temperature as a function of time for Tree 7 during 

Tests 3, 14, 18, and 15 calculated using the two-layer reduction method. ......................... 106 

Figure 9-16: The expanded uncertainty spread around nominal values of the 

hot/cold gas layer interface height for all tests except Test 18. .......................................... 108 

Figure 9-17: Layer interface height as a function of time for Tree 7 during Tests 2, 

8, 3, 9, 4, 10, and 5 calculated using the two-layer reduction method................................ 109 

Figure 9-18: Layer interface height as a function of time for Tree 7 during Tests 1, 

7, 2, 8, and 13 calculated using the two-layer reduction method........................................ 109 

Figure 9-19: Layer interface height as a function of time for Tree 7 during Tests 

13, 16, 2, and 17 calculated using the two-layer reduction method.................................... 110 

Figure 9-20: Layer interface height as a function of time for Tree 7 during Tests 3, 

14, 18, and 15 calculated using the two-layer reduction method........................................ 110 

Figure 9-21: Temperature versus time for the thermocouple located at position 8 

on Tree 4 and also near Targets B and E for Tests 2, 8, 3, 9, 4, 10, and 5. ........................ 112 

Figure 9-22: Temperature versus time for the thermocouple located at position 8 

on Tree 4 and also near Targets B and E for Tests 1, 7, 2, 8, and 13. ................................ 113 

Figure 9-23: Temperature versus time for the thermocouple located at position 8 

on Tree 4 and also near Targets B and E for Tests 13, 16, 2, and 17 (labeled 

P1T13, P1T16, P1T2, and P1T17, respectively). ............................................................... 113 

Figure 9-24: Temperature versus time for the thermocouple located at position 8 

on Tree 4 and also near Targets B and E for Tests 3, 14, 15, and 18 (labeled 

P1T3, P1T14, P1T15, and P1T15b, respectively). ............................................................. 114 

Figure 9-25: Temperature versus time for the thermocouple located at position 5 

on Tree 2 and also near Target G for Tests 2, 8, 3, 9, 4, 10, and 5 (labeled 

P1T2, P1T8, P1T3, P1T9, P1T4, P1T10, and P1T5, respectively). ................................... 114 

Figure 9-26: Temperature versus time for the thermocouple located at position 5 

on Tree 2 and also near Target G for Tests 1, 7, 2, 8, 13 (labeled P1T1, P1T7, 

P1T2, P1T8, and P1T13, respectively). .............................................................................. 115 

Figure 9-27: Temperature versus time for the thermocouple located at position 5 

on Tree 2 and also near Target G for Tests 13, 16, 2, and 17 (labeled P1T13, 

P1T16, P1T2, and P1T17, respectively). ............................................................................ 116 

Figure 9-28: Temperature versus time for the thermocouple located at position 5 

on Tree 2 and also near Target G for Tests 3, 14, 15, and 18 (labeled P1T3, 

P1T14, P1T15, and P1T15b, respectively). ........................................................................ 116 

Figure 10-1: A Vatell model BF-04 heat flux gauge installed on the South wall 

along with a wall thermocouple.......................................................................................... 119 

Figure 10-2: North and south Wall Sensor Locations. ............................................................... 120 

Figure 10-3: Ceiling and floor Sensor Locations........................................................................ 121 

Figure 10-4: East and west wall sensor locations. ...................................................................... 121 

Figure 10-5: Net heat flux to the wall near N1 and N3 gauges versus time for 

Tests 2 to 5.......................................................................................................................... 124 
xiii 

Figure 10-6: Net heat loss rate to individual surfaces and the total sum for all 

surfaces versus time for Test 3............................................................................................ 125 

Figure 10-7: Net heat loss rate to individual surfaces and the total sum for all 

surfaces versus time for Test 5............................................................................................ 125 

Figure 10-8: Net heat loss rate to individual surfaces and the total sum for all 

surfaces versus time for Test 2............................................................................................ 126 

Figure 10-9: Net heat loss rate to individual surfaces and the total sum for all 

surfaces versus time for Test 4............................................................................................ 126 

Figure 10-10: Rate of heat loss to the enclosure surfaces, through the doorway, 

and their sum as a function of time for Test 3. ................................................................... 127 

Figure 10-11: Net heat flux comparison of uncooled gauges N3 and N4 with 

cooled gauge RG10 for Test 2 as a function of time. ......................................................... 127 

Figure 10-12: Net heat flux comparison of uncooled gauges N3 and N4 with 

cooled gauges FG9 and RG10 for Test 3 as a function of time.......................................... 128 

Figure 10-13: Interior installation of marinite thermocouples.................................................... 129 

Figure 10-14: Exterior installation of marinite thermocouples................................................... 130 

Figure 10-15: Peak marinite temperatures for Tests 2 to 4 at five measurement 

stations (N3, S3, W3, C1 and C2)....................................................................................... 131 

Figure 11-1: Differential pressure for Test 2, door closed, ventilation off................................. 134 

Figure 11-2: Differential pressure for Test 4, door closed, ventilation on. ................................ 135 

Figure A-1: Morphology of a 7/C #14 AWG Control Cable (Cable 1 in Table A1). ........................................................................................................................................ 144 

Figure A-2: Morphology of a 7/C #14 AWG Control Cable (Cable 2 in Table A1). ........................................................................................................................................ 144 

Figure A-3: Morphology of a 3/C #6 AWG Power Cable (Cable 3 in Table A-1). ................... 144 

Figure A-4: Morphology of a 3/C #6 AWG Power Cable (Cable 4 in Table A-1). ................... 144 

Figure G-1: The manufacturer supplied fan curve that relates the static pressure to 

the blowing volume for various values of the fan power.................................................... 170 

Figure H-1: Thermocouples in and on the surface of the junction box (30 cm x 30 

cm x 10 cm) measuring temperatures in the box center and on the middle of 

the bottom and west surfaces. ............................................................................................. 171 

Figure I-1: The measured fuel flow and nozzle temperature as a function of time 

during the 1MW closed door compartment experiment. .................................................... 174 
xiv 

LIST OF TABLES 

Table 2-1: Test Matrix and Experimental Conditions. ................................................................... 5 

Table 2-2: Observations Noted during the Test Series. .................................................................. 6 

Table 3-1: Duration of the Fuel Flow1.......................................................................................... 15 

Table 3-2: The Ambient Temperature Fuel Flow Calibration, Associated HRR, 

and its value Corrected for Compartment Effects................................................................. 16 

Table 3-3: Combustion Properties of the Test Fuels1 ................................................................... 16 

Table 3-4: Instruments for the 9 m x 12 m NIST Calorimeter System......................................... 18 

Table 3-5: Summary of the calorimetry calibration results. ......................................................... 21 

Table 3-6: Average value of the steady-state HRR. ..................................................................... 22 

Table 3-7: The Estimated and Measured HRRs1 and their Ratio. ................................................ 24 

Table 3-8: Estimated HRR during the Steady Burning Period in the Closed Door 

Tests. ..................................................................................................................................... 25 

Table 3-9: Estimated HRRs during the steady burning period for all of the tests 

and the duration of each of the phases of the fuel flow. ....................................................... 27 

Table 4-1: Locations of total heat flux gauges and radiometers................................................... 30 

Table 5-1: Standard Relative Uncertainties for a Mass Concentration of 0.1 g/m3...................... 41 

Table 5-2: Smoke and Temperature Results................................................................................. 43 

Table 6-1: Thermocouple Physical Location and Orientation on Targets B, E, and 

G............................................................................................................................................ 50 

Table 8-1: Gaseous sampling locations. ....................................................................................... 73 

Table 8-2: Analyzer Particulars. ................................................................................................... 74 

Table 8-3: Sample transport time and analyzer response. ............................................................ 75 

Table 8-4: Estimated expanded uncertainty (with an expansion factor of two). .......................... 76 

Table 9-1: Calculated differences between aspirated and bare bead thermocouple 

temperature measurements and expanded temperature uncertainties................................... 96 

Table 9-2: Hot gas layer temperature expanded uncertainty limits. ........................................... 100 

Table 9-3: Hot gas layer interface height expanded uncertainty limits. ..................................... 107 

Table 10-1: North and south wall sensor locations..................................................................... 120 

Table 10-2: Ceiling and floor sensor locations........................................................................... 120 

Table 10-3: East and west wall sensor locations. ....................................................................... 121 

Table 10-4: Overview uncertainties of net heat flux. ................................................................. 122 

Table 11-1: Maximum compartment pressure differential. ........................................................ 134 

Table A-1: Cable types used in the experiments. ....................................................................... 143 

Table A-2: Material and optical properties of the cable materials.............................................. 145 

Table A-3: Material and optical properties of marinite. ............................................................. 146 

Table A-4: Material and optical properties of feraloy and gypsum at ambient 

temperature. ........................................................................................................................ 146 

Table B-1: Humidity and temperature in the large fire lab......................................................... 147 

Table C-1: Measured Pressure, Leakage Area and Air Flow Rate............................................. 149 

Table C-2: Calculated C and n values, compartment equivalent leakage area and 

expanded uncertainty. ......................................................................................................... 150 

Table D-1: Instrument channel number, designation, units, and location. ................................. 152 
xv 

Table D-2: Instrument channel number, designation, units, and location for the 

second data file.................................................................................................................... 158 

Table E-1: List of inoperative instruments during the test series. .............................................. 161 

Table E-2: List of inoperative instruments during the test series. .............................................. 162 

Table E-3: List of inoperative instruments during the test series. .............................................. 163 
xvi 

ACRONYMS 

ASTM American Society for Testing and Materials 

AWG American Wire Gauge 

CSPE Chlorosulfinated Polyethylene 

FDS Fire Dynamics Simulator 

FHA Fire Hazards Analysis 

FRA Fire Risk Analysis 

HGL Hot Gas Layer 

HRR Heat Release Rate 

ICFMP International Collaborative Fire Model Project 

NFPA National Fire Protection Association 

NIST National Institute of Standards and Technology 

NPP Nuclear Power Plant 

NRC Nuclear Regulatory Commission 

PVC Polyvinyl Chloride 

RI/PB Risk-Informed/Performance-Based 

SDP Significance Determination Process 

TC Thermocouple 

WPI Worcester Polytechnic Institute 

XPE Crosslinked Polyethylene 
1 

1 INTRODUCTION 

This document contains experimental results, which are useful for comparison with fire models as part 

of Validation Exercise #3 of the International Collaborative Fire Model Project (ICFMP). The objective 

of the collaborative project is to share the knowledge and resources of various international 

organizations to evaluate and improve the state of the art of fire models for use in nuclear power plant 

fire safety and fire hazard analysis. A total of five benchmarking and validation exercises have been 

conducted in the ICFMP, of which this report documents exercise #3. NRC, in collaboration with 

NIST, conducted these experiments in summer 2003. A detailed description of the experimental 

procedures, configuration, and instrumentation is documented in this report. 

 

The test configuration and fire scenarios have been selected to examine the following effects: 

• Heat release rate (HRR) 

• Natural ventilation with an open door 

• Mechanical ventilation system operation 

• A combination of mechanical and natural ventilation 

• Distance between fire and target 

• Heating of cables and a PVC slab directly in the plume region 

The measurements include the following: 

• Duration of fuel flows and HRRs 

• Compartment pressure 

• Radiative and total heat flux at various targets in the compartment 

• Surface and core cable temperatures and PVC slab temperature 

• Hot gas layer (HGL) temperature, depth, soot density, and concentrations of oxygen, carbon 

monoxide and carbon dioxide 

• Gas temperature vertical profiles 

• Total heat loss to the walls, ceiling, and floor 

• Total heat loss from the door and mechanical exhaust vent 

Visual and infrared video records of the experiments were also acquired, but are not presented in this 

report. 

This report presents a portion of the data that was collected during the NRC/NIST test series. The 

experimental results were composed of 15 tests with more than 350 channels of data acquired per test at 

a sampling rate of 1 Hz. The test duration was typically 1600 s, yielding on the order of 107

 discrete 

pieces of data. Given limited resources, there is simply too much data to complete a comprehensive 

analysis of the results. Instead, representative measurement results and estimates of measurement 

uncertainty are provided in this report. The full set of data is available electronically from NIST. 
2 

This report includes a description of the measurement methods and experimental results for the tests 

conducted as part of the ICFMP Benchmarking Exercise #3. Section 2 of this report describes the 

experimental configuration including the experimental conditions, the fuel flow and details of the test 

compartment. The rest of the document focuses on the measurement methods and results. Section 3 

describes the fuel flow and HRR measurements. Section 4 discusses the heat flux measurement results. 

Section 5 describes the smoke concentration measurements. Section 6 discusses the target temperature 

measurements. Section 7 describes the vent and doorway flows. Section 8 discusses gaseous sampling 

results. The gas temperature measurements are described in Section 9 and the heat loss to the enclosure 

is discussed in Section 10. Section 11 describes the measurement of the compartment pressure. Section 

12 presents summary and conclusions. Section 13 acknowledges those that provided support for the 

experiments. The references cited in the text are listed in Section 14. 

 

A large amount of information is also provided in the appendices of this report. Appendix A of this 

report provides information on the structure, and the thermal and optical properties of the “marinite” 

(calcium silicate) boards that composed the compartment walls and ceiling, Feraloy that composed a 

junction box located on the ceiling, gypsum boards that comprised the compartment floor, and polyvinyl 

chloride (PVC) and crosslinked polyethylene (XPE) thermoplastic that comprised cable jacket materials 

of the electrical cables that acted as “targets” during the fire experiments. Appendix B provides 

information on the ambient humidity and temperature during the test dates. Appendix C describes 

measurements of the intrinsic leakage associated with the compartment, which was measured before 

initiation of the first experiment and a number of times during the test series. Appendix D provides a 

discussion of the format of the companion electronic data. Appendix E lists non-functioning instrument 

channels. Appendix F lists response to questions and comments on this report from a July 2003 ICFMP 

peer review meeting at Worcester Polytechnic Institute (WPI). Appendix G provides the pressure 

curves for the ventilation supply fan used in the experiments. Appendix H describes the metal junction 

box placed in the ceiling of the compartment. Appendix I describes the method by which a correction 

was applied to the estimated fuel flow during the experiments. Appendix J provides documentation of 

personal communications cited in the text. 
3 

2 EXPERIMENTAL CONFIGURATION AND 

CONDITIONS 

The test matrix and experimental conditions are presented in Table 2-1. The table lists the test 

number, the nominal peak HRR (Q& ), the type of cables present during the test, the fuel type, 

burner location, ventilation, and door condition. Observations made during the test series are 

listed in Table 2-2. The order of the tests is also listed. The observations highlight the melting of 

the cables and a small slab of PVC (Slab E, described in Section 6), as well as a number of 

instrumentation issues. 

2.1 Test Compartment 

Figure 2-1 is a photo of the compartment taken from the door in the west wall, looking east. The 

compartment was 7.04 m x 21.7 m x 3.82 m (23.1 ft x 6.60 ft x 12.5 ft) in dimension, designed to 

represent a realistic-scale compartment in a nuclear power plant. The total compartment volume 

was 582 m3 (20,500 ft3). Looking in from the 2.0 m by 2.0 m (6.6 ft by 6.6 ft) double door, 

Figure 2-1 shows the right (South), back (East), and left (North) walls. Walls and ceiling were 

covered with two layers of 12.7 mm (0.500 in) marinite boards, while the floor was covered with 

one layer of 12.7 mm (0.500 in) gypsum board on top of one layer of 18.3mm (0.720 in) 

plywood. The supply duct and horizontal cables are evident on the right of Figure 2-1, while the 

vertical cable tray and exhaust duct are on the left. Figures 2-2 – 2-4 are schematic drawings of 

the compartment in which the location of some of the compartment features are shown including 

the Targets (A-F), ventilation ducts, thermocouple trees, junction box, fire pan, and door. 

2.2 Compartment Contents 

The targets have been arranged to examine the following effects: 

1. Modeling one cable versus cables bundled in a cable tray. 

2. Modeling a cable as composed of a slab with uniform material versus a real cable geometry 

and composition. 

3. Heating characteristics of cables with a large diameter versus smaller cables 

4. Elevation of the target in the hot gas layer 

5. Vertical versus horizontal cable target 

6. Heating of a junction box on the ceiling 

The fire was located in the center of the compartment at floor level for most of the tests. For a 

number of tests the fire pan was moved away from the center of the compartment (see Table 2-

1). 

The compartment contained three control cables (A, B, C), a horizontal (Target D) and a vertical 

cable tray (Target G) with control cables, a solid PVC slab “target” (E), and a single power cable 
4 

(F). The control cables had seven conductors (7/C # 14 American Wire Gauge, AWG). The 

power cable had three conductors (3/C # 6 AWG). For the primary tests in the series, all the 

cables were constructed of XPE insulation and Hypalon jacket materials. For the replicate tests, 

all the cables were based on PVC jacket and PVC/Nylon insulation materials. Further 

information on the structure and the thermal and optical properties of the cables are discussed in 

Appendix A of this report. 

Figure 2-1: Photo of the test compartment being instrumented before testing. 

South
5 

Table 2-1: Test Matrix and Experimental Conditions. 

Test Nominal Peak 

Q& (MW) 

Cable 

Type 

Fuel; Burner Location Door Ventilation 

1 0.35 XPE1 Heptane3; Center Closed Off 

2 1 XPE Heptane; Center Closed Off 

3 1 XPE Heptane; Center Open Off 

4 1 XPE Heptane; Center Closed On 

5 1 XPE Heptane; Center Open On 

6 Not Conducted 

7 Replicate Test 1 PVC2 Heptane; Center Closed Off 

8 Replicate Test 2 XPE Heptane; Center Closed Off 

9 Replicate Test 3 XPE Heptane; Center Open Off 

10 Replicate Test 4 PVC Heptane; Center Closed On 

11 Replicate Test 5 Not Conducted 

12 Replicate Test 6 Not Conducted 

13 2 XPE Heptane; Center Closed Off 

14 1 XPE Heptane; 1.8 m from N wall 

on E-W centerline

Open Off 

15 1 PVC Heptane 

1.25 m from S wall 

on E-W centerline 

Open Off 

16 2 PVC Heptane; Center Closed On 

17 1 PVC Toluene, Center Closed Off 

18 1 XPE Heptane 

1.55 m from S wall 

1.50 m E of centerline 

Open Off 

1. XPE cable has crosslinked polyethylene insulation. 

2. PVC cable has a polyvinylchloride jacket 

3. Heptane is a commercial blend of heptane isomers. 
6 

Table 2-2: Observations Noted during the Test Series. 

Test Test Date Test 

Order 

Observations 

1 6/5/03 1 No observations noted. 

2 6/6/03 2 No observations noted. 

5 6/9/03 3 N2 purge flow for soot laser was low for first 5 min after 

ignition. 

Slab E appeared melted after test. 

4 6/10/03 4 Slab E not present. 

3 6/10/03 5 Slab E not present. 

8 6/11/03 6 Slab E not present 

Leakage around door during testing. 

Cable burned at Ts-10 and Tc-11 during test. 

9 6/12/03 7 Slab E not present. 

Cable C appears charred after test. 

13 6/12/03 8 Slab E not present. 

14 6/13/03 9 Wall flux gauges not functioning. 

Slab E not present. 

All cables melted during test. 

18 6/16/03 10 Smoke coming out of south wall-ceiling joint 17 min after 

ignition. Doused with water at 18 min 30 s. 

Flux Gauge #8 not functioning before test start. 

Slab E not present. 

7 6/18/03 11 New PVC cables and slab E installed before test. 

Biderectional Probes #13 & #14 not functioning properly 

before test start. 

10 6/19/03 12 Wall flux gauges (N6, S6, C4, C5, C8) not functioning before 

test start. 

Slab E and cables partially melted during test. 

16 6/20/03 13 Flux Gauge #8 not functioning before test start. 

15 6/20/03 14 Flux Gauge #7 erratic behavior noticed 730 s after ignition. 

vertical tray melted above 2 m. 

17 6/27/03 15 Fuel secured when loss of visibility completely obstructed the 

fire. 

Slab E not present. 

Flux Gauge #1 low water flow before test start. 

Flux Gauge #5 not working before test start. 

6 - - Not Conducted. 

11 - - Not Conducted. 

12 - - Not Conducted. 
7 

The compartment also contained the following: 

• A ladder type cable tray (Target D) was 0.3 m (1.0 ft) wide and 0.1 m (0.3 ft) deep. The tray 

contained two layers of control cables constructed of XPE insulation and Hypalon jacket 

materials in the primary tests, and three layers of control cables constructed of PVC jacket 

and PVC/Nylon insulation materials in the replicate tests. The center of the bottom of the 

tray was located 2.0 m (6.6 ft) from the right wall, 3.2 m (10.5 ft) above the floor. It was 10.0 

m (32.8 ft) long, extending from 5.85 m (19.2 ft) from the front wall to 5.85 m (19.2 ft) from 

the back wall. 

• The bottom of the center of a rectangular slab target (E) was located 1.25 m (4.10 ft) from the 

right wall, 2.7 m (8.9 ft) above the floor and centrally located between the front and back 

walls. The slab was composed of PVC and is detailed in Section 6 of this report. 

• The bottom of the center of the power cable (F) was located 0.5 m (1.6 ft) from the right 

wall, 2.2 m (7.2 ft) above the floor and extended 10.0 m (32.8 ft), from 5.85 m (19.2 ft) from 

the front wall to 5.85 m (19.2 ft) from the back wall. 

• The three control cables A, B and C were located at the same elevation and 0.1 m (0.3 ft) 

from the left edge of the power cable, slab target, and cable tray, respectively. They 

extended 10.0 m (32.8 ft), from 5.85 m (19.2 ft) from the front wall to 5.85 m (19.2 ft) from 

the back wall. 

• The ladder type vertical cable tray (G) was 0.3 m (1.0 ft) wide and 0.1 m (0.3 ft) deep. The 

tray contained one layer of control cables constructed of XPE (polyethylene) insulation and 

Hypalon jacket materials in the primary tests, and PVC (polyvinyl chloride) jacket and 

PVC/Nylon insulation materials in the replicate tests. The tray was located on the surface at 

the center of the north wall, extending from the floor to the ceiling. 

• The junction box was a WCB Junction Box. It was heavy-duty, dust-tight, weatherproof, rain 

and watertight, with nominal inside dimensions, 30 cm length x 30 cm width x 10 cm depth 

(1.0 ft x 1.0 ft x 0.3 ft). It had an approximate wall thickness of 0.7 cm (0.3 in). The 

junction box was mounted on the ceiling and located on the compartment centerline, with its 

center 17.7 m (58.1 ft) from the door. The box was made of Feraloy (see Appendix A for 

thermophysical properties and Appendix H for a more complete description). 

2.3 Compartment Openings and Ventilation 

A 2.00 m by 2.00 m (6.56 ft x 6.56 ft) door was present in the middle of the west wall (see 

Figure 2-2). The compartment was equipped with supply and exhaust forced ventilation, which 

was used during Tests 4, 5, 10, and 16, providing about 5 air changes per hour. The midpoint of 

the supply and exhaust vents was located 11.22 m (36.80 ft) from the door and 2.40 m (7.87 ft ) 

above the floor. An exhaust duct of equal area to the supply duct was positioned on the opposite 

wall at a comparable location. The vents were square (0.70 m x 0.70 m, or 2.3 ft x 2.3 ft) with an 

area of 0.50 m2 (5.4 ft2) each. The flow rates through the supply and exhaust ducts were 

measured in detail during breaks in the testing, in the absence of a fire and are discussed in detail 

in Section 7 of this report. During the tests, the flows were monitored with single bi-directional 

probes positioned in the ducts. 
8 

2.4 Fuel Delivery System and Fire 

The fuel system consisted of a fuel storage container and a Tuthill® sealless magnetic drive 

positive displacement gear pump. The pump was controlled by an alternating current (AC) Tech 

variable frequency driver. Fuel was delivered through one BETE® spray nozzle oriented 

downwards towards a 1.0 m by 2.0 m (3.3 ft by 6.6 ft) stainless steel pan positioned on the floor. 

A 90o

 spray angle WL 3⁄4 type nozzle was used for the 350 kW fire, and a similar angle WL-1 

type nozzle was used for the larger fires. The fire pan was located at the center of the 

compartment for most of the tests (Tests 1 - 13, 16, and 17). In Test 14, the fire was centered 1.8 

m (5.9 ft) from the North wall. In Test 15, the fire was centered 1.25 m (4.10 ft) from the South 

wall. In Test 18, the fire was centered 1.55 m (5.09 ft) from the South wall. Physically, the fuel 

pan was 2.0 m (6.6 ft) long x 1.0 m (3.3 ft) wide and 0.1 m (0.3 ft) deep. A single nozzle was 

used to spray liquid hydrocarbon fuels onto the pan. The test plan originally called for the use of 

two nozzles to provide the fuel spray. Experimental observation suggested that the fire was 

steadier with the use of a single nozzle. In addition, it was observed that the actual extent of the 

liquid pool was well-approximated by a 1.0 m (3.3 ft) circle about the center of the pan. The 

uncertainty in the location of the liquid fuel was about ±0.1 m (± 0.3 ft). 

2.5 Background Documentation 

A number of previous documents have provided information on the values of the fuel flow, 

measured fuel flow rates, supply and exhaust ventilation flows, ambient conditions, and 

compartment leakage needed to fully specify the boundary conditions for implementation of fire 

models. Much of that information is presented here for the sake of comprehensiveness. 

Knowledge of the optical and thermal properties of the “marinite” (calcium silicate) board, 

which composed the enclosure wall and ceiling material, is needed to fully document the 

experimental boundary conditions. Thermal property information including temperature 

dependent information on the specific heat (cp), the thermal diffusivity (α), and the thermal 

conductivity (K) of the marinite was determined using ASTM E1269 and ASTM E1461 [Taylor 

et al., 2003] and are listed in Appendix A. 

The spectrally integrated value of the emissivity (ε) for marinite is given in Appendix A. 

Analogous property information for the PVC and XPE cable jacket materials are also given in 

Appendix A. Material and Optical Properties of Feraloy, which composed a junction box located 

on the ceiling (see Figure 2-2), and gypsum, which comprised the compartment floor, are given 

in Appendix A.
9 

Figure 2-2: Compartment Isometric with thermocouple trees.

Fire Pan

TC Tree (#2)

TC Tree (#3)

TC Tree (#4)

TC Tree (#5)

TC Tree (#7)

TC Tree (#6)

Exhaust Vent

Supply Vent 

C D

 B

Vertical Ladder Tray (G) 

A F 

Ju

(x, y, z) = (0, 0, 0)

3.82 m

7.04 m

21.66 m

TC Tree (#1)

North 

 door 

2.00 m

2.00 m

Cable TypesControl cables Cable Trays Slab Power cable 

E 
10 

Figure 2-3: Compartment Contents and Selected Instrumentation 
11 

Figure 2-4: Plan View of Compartment Mid-Section

13 

3 HEAT RELEASE RATE 

The total HRR is the most critical parameter that characterizes a fire. The HRR drives the 

thermal environment including the radiative and convective heat transfer, the hydrodynamic 

flow, the rate of fire spread and growth, the amount of smoke and toxic products, and ultimately, 

the hazard associated with a fire. 

In this report, the HRR was measured two different ways. For all of the tests, it was estimated 

from the fuel flow. For the open door tests (Tests 3, 5, 9, 14, 15, and 18), it was measured using 

oxygen consumption calorimetry. As another consistency check, measurements of the 

convective and conductive heat losses from the compartment for one test (Test 3) were used to 

estimate the fire HRR. 

3.1 Heat Release Rate Based on the Fuel Flow 

The fuel system was designed to deliver a controlled amount of liquid fuel in the form of a spray. 

Table 3-2 includes the target fuel flow for each experiment, which was calibrated at ambient 

temperature. The fuel used in all but one of the tests was a commercially available blend of 

heptane isomers. In Test 17, toluene was used. A nominal HRR of 1 MW was the target for all 

tests, except Tests 1 and 7 (350 kW), and Tests 13 and 16 (2 MW). 

The fuel system has been previously described in detail [Hamins et al., 2003a; 2003b]. 

Typically, the transient fuel flow was composed of a nominal 3 min linear ramp-up from zero to 

a long steady burning period, followed by a 3 min linear ramp-down to zero flow. The exact 

duration of each of the periods is listed in Table 3-1. 

The HRR (Qf

& ) shown in Table 3-2 was estimated as the product of the mass flow rate of fuel 

( m& ), the heat of combustion of the fuel (Hc), and the fuel combustion efficiency (χa ): 

Qf

& = χa · m& · Hc (3.1)

Information on the global combustion properties of the hydrocarbon fuels tested here (including 

the heptane and the blend of commercially available heptane and toluene) were determined from 

a separate series of experiments, which measured the combustion efficiency, radiative fraction, 

and the yields of soot, CO and CO2 from a downward-oriented spray fire burning in a similar 

configuration to that used here, with the fire burning in the open [Hamins et al., 2003b]. For 

convenience, those results are summarized in Table 3-3. The heats of combustion of the test fuels 

are also listed in Table 3-3. 

The uncertainty in Qf

& (see Table 3-2) was dominated by the uncertainty in χa, and to a lesser 

extent by uncertainties associated with m& and Hc. The uncertainty in χa was described in detail 
14 

in Hamins et al. [2003b]. It was assumed that χa was the same in the compartment fires as in the 

open burn calibration experiments. The uncertainties in this report, unless otherwise stated, are 

expressed as the expanded relative uncertainty with an expansion factor equal to two (i.e., 2·σ, 

where σ is the standard deviation), which represents a 95 % confidence interval. 

The mass flow rate of fuel is equal to the product of the measured volumetric flow rate of fuel 

(V& ) and the fuel density (ρ): 

m& = ρ · V& (3.2) 

The estimated HRR assumes that the combustion efficiency in the compartment was the same as 

that measured during the free burn experiments, which were in the open. The fuel density (ρ) for 

heptane and toluene was 688.5 ± 0.4 kg/m3

 (43.00 ± 0.02 lb/ft3) and 871.2 ± 0.5 kg/m3 (54.40 ±

0.03 lb/ft3

), respectively [Hamins et al., 2003b]. The fuel flow (V& ) was calibrated at ambient 

temperature before each experiment. An experiment was conducted at the end of the test series to 

determine the effect of compartment heat-up on the fuel flow rate. Appendix I describes the 

experimental procedure and results. The results showed that the fuel flow during the steady 

burning period of the experiment was larger, by a factor of about 1.09, than the fuel flow 

calibrated at ambient temperature. The correction factor was applied to the HRR estimated from 

the ambient temperature fuel flow calibration, and the corrected HRR is shown in Table 3-2. 
15 

Table 3-1: Duration of the Fuel Flow1

Test Ramp-Up 

(min:sec) 

Steady State 

Duration (min:sec) 

Ramp-Down 

(min:sec) 

Total Duration 

(min:sec) 

1 4 2:28 20:02 2:30 25:00 

2 3:00 7:25 0:01 10:25 2 

3 2:58 20:01 3:03 5 26:03 

4 2:58 10:36 0:01 13:35 2

5 2:58 20:01 3:03 5 26:03 

7 2:09 20:03 2:08 24:20 

8 2:56 7:14 0:01 10:13 2

9 2:55 20:01 3:04 5 26:06 

10 2:56 10:50 0:01 13:47 2

13 2:57 3:07 0:01 6:05 2

14 2:56 20:05 3:06 5 26:07 

15 3:00 20:00 3:07 5 26:07 

16 2:57 3:25 0:01 6:23 2

17 3:01 1:31 0:01 4:33 3

18 2:58 20:02 3:07 5 26:07 

1. The temporal measurement uncertainty estimated as ± 2 s. 

2. The test was abruptly stopped when the oxygen concentration measurement in the lower layer was 

less than or equal to 15 %. 

3. The fuel flow was stopped after 273 s for safety reasons when loss of visibility completely 

obstructed the fire. 

4. There was no ramp-down for the tests with the door closed, except Test 1, as the fuel flow was 

rapidly shut for safety reasons when the measured oxygen concentration in the lower layer of the 

compartment was measured to be less than 15 % by volume (on a dry basis). 

5. Fuel dripping from the nozzle sometimes added as much 2 s to 7 s to the duration of the rampdown, but the fire size during that period was typically very small (less than 10 kW).
16 

Table 3-2: The Ambient Temperature Fuel Flow Calibration, Associated HRR, and its 

value Corrected for Compartment Effects. 

Test Fuel Fuel Flow (ml/min)1 Qf

& (kW) 2 Corrected 

Qf

& (kW) 3 

1 Heptane 697 360 390 

2 Heptane 2020 1040 1130 

3 Heptane 2040 1050 1140 

4 Heptane 2040 1050 1140 

5 Heptane 2040 1050 1140 

7 Heptane 676 350 380 

8 Heptane 2020 1040 1130 

9 Heptane 1990 1030 1120 

10 Heptane 2010 1040 1130 

13 Heptane 3950 2040 2220 

14 Heptane 2010 1040 1130 

15 Heptane 2010 1040 1130 

16 Heptane 3910 2020 2200 

17 Toluene 2200 1020 1110 

18 Heptane 2010 1040 1130 

1. measured at ambient temperature; the average expanded relative uncertainty was 2 %. 

2. the expanded relative uncertainty was 5 %. 

3. the expanded relative uncertainty was estimated as 10 %; a correction factor of 1.09 was 

applied to account for compartment effects on the fuel flow, including compartment heat-up.

Table 3-3: Combustion Properties of the Test Fuels1

Fuel Hc (kJ/g) 2 Combustion 

efficiency3

Radiative 

fraction3

Soot yield3

 CO yield3 CO2 yield3

Heptane 45.0 1.0 ± 0

0.14

4 0.44 ± 0.07 0.0149 ± .0033 <0.008 3.03 ± 0.12 

Toluene 40.3 0.76 ± 0.10 0.40 ± 0.09 0.195 ± 0.052 0.070 ± 0.017 2.52 ± 0.22 

1. Reproduced from Hamins et al. [2003b]. 

2. The expanded relative uncertainty is typically 5 % [Hamins et al., 2003b]. 

3. Hamins et al., 2003b. 

4. Physically, the combustion efficiency, χa ≤ 1, so that the uncertainty estimate is not symmetric. The 

measurement uncertainty of χa is dominated by uncertainty in the radiative flux, the mass burning rate (m& ), and 

Hc. 
17 

3.2 Heat Release Rate Based on Oxygen Consumption Calorimetry 

The HRR (Q& ) was measured using oxygen consumption calorimetry in the 9.0 m by 12.0 m (29 

ft by 39 ft) exhaust hood in the NIST Large Fire Laboratory. The west side of the compartment 

included a door, which was located adjacent to one edge of the exhaust hood. When the 

compartment door was open, the products of combustion from the fire filled the upper layer of 

the compartment and then flowed out of the door, where they were completely captured by the 

exhaust hood. Measurements of the duct mass flow rate and the concentration of various gas 

species were used to infer the HRR of the fire. This form of fire calorimetry was first suggested 

by Huggett [1980], who exploited the finding that the amount of heat evolved from most organic 

materials per unit mass of oxygen consumed in their complete combustion is nearly constant. 

Thus, the oxygen deficit in the duct flow (relative to ambient air) is a measure of the HRR in the 

flow. Huggett showed that for most common materials containing C, H, O, N, the heat release 

per unit mass of oxygen is constant to within ± 5 %; this sets a fundamental accuracy limit in this 

method for materials that are not chemically characterized. For fuels that are characterized, such 

as those used in the experiments described here, this parameter can be estimated with greater 

accuracy. The value of the heat of combustion per mass of oxygen consumed, which was used in 

the calorimetry measurements, is described in Hamins et al. [2003b]. 

3.2.1 Measurement Approach 

Bryant et al. [2003] describe the heat release measurement facility, instrumentation, calibration, 

measurement uncertainty and experimental procedures in detail. The calorimetry determination 

required about 40 measurements. Some of the instruments employed in the HRR measurement 

are listed in Table 3-4. Measurements of temperature and pressure occurred in the exhaust duct. 

Gas was sampled in the exhaust duct and transported to the instruments in a control room for 

measurements of oxygen, carbon dioxide, and carbon monoxide. Water vapor in the exhaust 

stream was trapped, and not measured. The relative humidity in the test bay area was measured. 

The computation of HRR (Q& ) was made following Bryant et al. [2003; 2004]. The relationship 

that describes the HRR determination [Bryant et al, 2004]) is reproduced here: 

o

O CO CO O

o

CO CO O CO

o

O

o

O

air

O

o

e H O

O

CO

HC CO HC

X X X X

X X X X X

with

X

M

m X M

X

X Q E E E

2 2 2

2 2 2 2

2

2 2

2

(1 )

(1 ) (1 )

(1 )

1 ( 1) ] 2

1 [ ( )

− − −

− − − − =

⎥

⎦

⎤ ⎢

⎣

⎡ −

+ −

− = − −

φ

φ α

φ φ & &

 (3.3) 

 EHC = heat of combustion of hydrocarbon fuel 

 ECO = heat of combustion of carbon monoxide 

 φ = oxygen depletion factor 

me & = mass flow rate in exhaust duct 

mair & = mass flow rate of air 
18 

 α = combustion products expansion factor 

 Mi = molecular weight of species i 

 Xi = exhaust-gas concentration measurement of species i 

o Xi = ambient-gas concentration measurement of species i 

Bryant et al. [2003; 2004] discuss the above equation, the calibration procedures, and the 

methods used to determine the HRR and its uncertainty for calorimetry in the NIST 6 m by 6 m 

(20 ft by 20 ft) exhaust hood. The instrumentation and uncertainty analysis is analogous to 

Bryant et al. [2003; 2004] for calorimetry in the 9.0 m by 12.0 m hood (29 ft by 39 ft), which 

was used here. The largest contribution to the measurement uncertainty is due to the 

determination of the mass flow in the exhaust duct [Bryant et al., 2004]. Parker [1982] and 

Janssens and Parker [1995] discuss the details of the HRR calculation based on the extent to 

which the duct gas flow is characterized. 

Table 3-4: Instruments for the 9 m x 12 m NIST Calorimeter System. 

Parameter 

Measured 

Instrument Type Make and 

 Model 

Ranges 

Oxygen Paramagnetic Servomex 540A 0-21 % v/v 

Carbon Dioxide Infrared extinction Siemens Ultramat 5E 0-5 % 

Carbon Monoxide Infrared extinction Siemens Ultramat 5E 0-3 % 

Relative Humidity Thin film 

capacitance 

Dickson THDx 0–95 % 

Temperature Thermocouple Omega K-type 0-1250 oC 

Exhaust Velocity Bi-directional probe MKS model 220 differential 

pressure transducer 

0-10 Pa 

3.2.2 Calibration and Measurement Uncertainty 

Calibration burns using natural gas were conducted on several occasions during the test series to 

assure the accuracy of the HRR measurement (Q& ) and to characterize its repeatability. Burning a 

substance such as a gaseous fuel at a controlled rate provides an independent measurement of 

HRR to compare to the measurement by oxygen consumption calorimetry. A natural gas burner 

with active flow control was employed [Bryant et al., 2003]. 

Calibration experiments were performed before Tests 3, 5, and 18. The burner was placed 

directly under the 9.0 m by 12.0 m (29 ft by 39 ft) exhaust hood. Figure 3-1 shows the 

calorimetry during the calibration immediately before and during Test 5. In the figure, Test 5 

began at time equal to zero (t = 0), while the calibration was conducted before the experiment for 

t < 0. The flow of natural gas to the burner was turned off during the experiment. The HRR 

expected from the natural gas burner was based on measurements of the mass flow of natural gas 

[Bryant et al., 2003; 2004] and Equations 3.1 and 3.2. During the calibration, the heat output of 

the burner was held constant for 2 min to 5 min at each setting. The expanded uncertainty of the 

natural gas HRR was estimated as 2 %, which was dominated by uncertainty in the caloric value 
19 

of the natural gas. The caloric value varied by approximately 1.3 % over a six-month period, as 

determined by examination of gas chromatographic concentration measurements provided by the 

natural gas supplier, the Washington Gas Company [Bundy, 2005]. 

The results of the calibration are summarized in Table 3-5, which lists Qf

& estimated from the 

flow of natural gas, the calorimetric measurement (Q& ), the background associated with the 

calorimetry, and the ratio of Q& from the calorimetry to that expected from the natural gas flow, 

Qf

& . The calorimetric HRR measurements were corrected for baseline drift (see Table 3-5), 

which was likely due to the instability of the oxygen analyzer. For example, note that the value 

of Q& in Figure 3-1 at t = -2000 s and at t = -500 s have different values. Table 3-5 lists the 

estimated value of the changing baseline for each of the calibration settings. Table 3-5 also 

shows that the ratio of the calorimetry Q& (with background corrected) to Qf

& of the natural gas 

flow varied from 1.04 to 1.18, with an average value of 1.11 and an estimated relative combined 

uncertainty of 15 %. 

In Table 3-5, the measured Q& was consistently higher than the expected HRR (Qf

& ) based on the 

natural gas mass flow rate by about 11 %. For this reason, the calorimetry results were reduced 

by a factor of 1.11. The reasons for the difference are not certain, but may be due at least in part 

to the nature of the flow field in the 9.0 m by 12.0 m exhaust hood (29 ft by 39 ft), in which the 

velocity profile does not correspond to fully developed pipe flow at the downstream 

measurement station. 
20 

0

500

1000

1500

-2000 -1500 -1000 -500 0 500 1000 1500 2000

Calorimetry

Gas Burner

Heat Release Rate (kW)

Time (s)

Figure 3-1: The calorimetric HRR data as a function of time during the calibration before 

Test 5 (time <0) and during Test 5 (time >0). 

3.2.3 Calorimetry Measurement Results 

Figure 3-2 shows the results of the oxygen depletion calorimetry measurement (Q& ) for the open 

door tests (Tests 3, 5, 9, 14, 15, and 18) as a function of time after ignition. The results show 

that the experiments were repeatable. The values of Q& were similar for all of the tests, except 

Test 5, which unlike the other tests had mechanical ventilation in addition to an open door. As 

expected, less of a delay in the HRR measurement occurred in the ramp-up during Test 5, as the 

ventilation exhaust was directly pulled from the compartment, rather than from the exhaust in the 

9.0 m by 12.0 m hood (29 ft by 39 ft), so that the mixing and delay associated with combustion 

products filling the upper layer of the compartment occurred over a different time-scale. The 

larger value of the HRR during the ramp-down is also evident. 
21 

Table 3-5: Summary of the calorimetry calibration results. 

Calorimetry 

Test 

 

Qf

& (kW) 

Gas Burner1

Q& (kW)2 Background (kW) 

Ratio3

(Calorimetry/ Burner) 

Test 3 741 

505 

932 

769 

533 

983 

-70 

-40 

-50 

1.13 

1.13 

1.11 

Test 5 

 

1190 

915 

714 

590 

1310 

1070 

831 

701 

70 

40 

45 

32 

1.04 

1.13 

1.10 

1.13 

Test 18 

 

1190 

978 

565 

1300 

1130 

659 

10 

50 

-10 

1.08 

1.10 

1.18 

Mean - - - 1.11 

1. Background was equal to zero for all gas burner data; the average relative expanded uncertainty was about 2 

%. 

2. the relative expanded instrument uncertainty was 14 %. 

3. the combined relative expanded uncertainty was 15 %.

The time-averaged values of Q& during the steady burn period of the open door tests are listed in 

Table 3-6. It is of interest to compare this value (1190 kW ± 15 %) to the mean of the corrected 

Qf

& for the same six tests (1130 kW ± 10 %), which are listed in Table 3-2. The mean values 

were not significantly different when uncertainty is considered. 

Calorimetric measurements were not conducted during Tests 1, 2, 7, 8, 13, and 17, because both 

the ventilation and the door were closed (see Table 2-1). Calorimetry was conducted during 

Tests 4, 10, and 16, by making measurements of the ventilation exhaust flow. In those 

experiments, the door was closed, but the ventilation was on. The calorimetric measurements 

conducted during Tests 4, 10, and 16 are not considered an accurate representation of the 

transient fire HRR, Q& . For typical experiments that measure the exhaust products of a fire 

burning in the open, oxygen depletion calorimetry can be thought of as a measure of the fire 

HRR with a measurement response time on the order of 15 s, which is principally limited by the 

oxygen analyzer time response [Bryant et al., 2003]. For the compartment experiments 

considered here, combustion products accumulated in the upper layer of the compartment, and 

the layer interface lowered until the products flowed, either through the open door as in Test 3, 

or through the ventilation exhaust duct as in Test 4. In either case, the volume of the test 

chamber affected the measurement of the fire HRR, as filling and mixing caused a timeaveraging and a delay in the calorimetry results. For the experiments with mechanical ventilation 

(Tests 4, 10, and 16), details of the mixing were unknown, and the calorimetric measurements 

should not be considered representative of the fire HRR. 
22 

0

300

600

900

1200

1500

0 500 1000 1500 2000

Test 3

Test 5

Test 9

Test 14

Test 15

Test 18

Heat Release Rate (kW)

Time (s)

Figure 3-2: The calorimetric HRR during the open door experiments (Tests 3, 5, 9, 14, 15, 

and 18) as a function of time. 

Table 3-6: Average value of the steady-state HRR. 

Test Average Q& (kW) during steady burning1

Test 3 1140 

Test 5 1160 

Test 9 1160 

Test 14 1230 

Test 15 1260 

Test 18 1150 

Mean 1190 

1. The average value for each test was determined from the data in 

Figure 3-2, during the steady burning period from 600 s to 1400 s; the 

combined relative expanded uncertainty is estimated as ± 15 %. 

To highlight the character of the calorimetry results, Figure 3-3 compares the measured HRR 

(Q& ) during Tests 3 and 5. The nature of the HRR profiles is very similar for these two tests. In 

Test 3, a steady Q& was observed after approximately 500 s. The delay is attributed to filling of 
23 

the large volume of the test compartment. The Q& in Test 4 was similar to that of Test 3 for the 

first 200 s of burning. In Test 5, Q& also achieved a steady value. In this test, mechanical 

ventilation was included, and the HRR was measured for the total effluent, passing through the 

door and through the exhaust vent. The forced ventilation system probably caused some amount 

of mixing in the compartment, but from Figure 3-3, the effect of this on the HRR appears to be 

limited to a HRR that more nearly matches the idealization as the ramp-up (and ramp-down) 

time is reduced. 

For the open door tests, it is presumed that the compartment volume did not affect the measured 

value of Q& during the steady burning period. While it may be possible to de-convolute the HRR 

measured in the exhaust duct in an attempt to quantify the instantaneous HRR during the rampup and ramp-down of Q& , the uncertainty in the calculation would probably undermine the utility 

of the results. The solid line in Figure 3-3 shows the idealized fire HRR within the compartment 

during Test 3. The curve was determined from the 50 s delay measured in the figure, and the 

information in Table 3-1, which includes a 180 s linear fuel ramp-up, 1201 s of steady burning, 

and a 180 s linear fuel ramp-down. The shape of the idealized curve appears to represent the 

Test 3 data fairly well, but is certainly different than the Test 4 measurement result, which was 

likely affected by mixing associated with the mechanical ventilation. 

Table 3-7 shows the measured HRRs for the open door tests (taken from Table 3-6). The table 

also shows the average values of the estimated HRR from the ambient temperature fuel flow 

calibrations (taken from Table 3-2). The ratio of these two HRR values yields a factor with a 

mean value of 1.14 for the six open door tests. The standard deviation of the results is 4 %, 

indicating a reasonable level of repeatability. To provide a source term for the HRR during the 

steady burning period for the tests not listed in Table 3-7, it is recommended that the same factor 

(1.14) be applied as a correction to the estimated HRRs during the steady burning period 

determined from the ambient temperature fuel flow measurements (see Table 3-2). The resulting 

values are listed in Table 3-8. This approach, based on the open door test results, inherently 

assumes that the combustion efficiency and the fuel flow rate did not significantly differ for 

experimental conditions different from the open door tests. The time dependent HRR for the tests 

listed in Table 3-8, including the ramp-up, the steady burning period and the ramp-down, can be 

estimated using the information in Table 3-1 in a manner analogous to the estimate of Q& for Test 

3 that is shown in Figure 3-3. 
24 

0

200

400

600

800

1000

1200

1400

0 500 1000 1500 2000

Test 3

Test 5

Heat Release Rate (kW)

Time (s)

Figure 3-3: The measured HRR as a function of time for Tests 3 and 5. The solid line 

shows the idealized fire HRR using gas flow rate. 

Table 3-7: The Estimated and Measured HRRs1 and their Ratio. 

Test 

Qf

& (kW) from ambient 

temperature fuel flow2

Q& (kW) from the Measured 

Calorimetric HRR3 Ratio4 

3 1050 1140 1.09 

5 1050 1160 1.11 

9 1030 1160 1.13 

14 1040 1230 1.18 

15 1040 1260 1.21 

18 1040 1150 1.11 

Mean 1040 1180 1.14 

1. Reproduced from Tables 3-2 and 3-6. 

2. Combined relative expanded uncertainty is 5 %. 

3. Combined relative expanded uncertainty is 15 %. 

4. Combined relative expanded uncertainty is 16 %.
25 

Table 3-8: Estimated HRR during the Steady Burning Period in the Closed Door Tests. 

Test Fuel Q& (kW)1

1 Heptane 410 

2 Heptane 1190 

4 Heptane 1320 

7 Heptane 440 

8 Heptane 1310 

10 Heptane 1310 

13 Heptane 2570 

16 Heptane 2550 

17 Toluene 1290 

1. The combined expanded relative uncertainty is estimated as 17 %. 

3.3 Energy Balance in the Compartment 

As another consistency check on the HRR results, measurements of the convective and 

conductive heat losses from the compartment for one test (Test 3) were used to estimate the fire 

HRR. This is discussed in more detail in Section 10 of this report. 

Conservation of energy holds that the fire HRR (Q(t) & ) heats the gases within the compartment 

(Q (t) g

& ), is transported through the doorway (Q (t) d

& ), through the mechanical ventilation duct 

(Q (t) v

& ), and is transferred to compartment surfaces (by radiation and convection) (Q (t) w

& ): 

Q(t) & = Q (t) g

& + Q (t) w

& + Q (t) d& + Q (t) v

& (3.4) 

Measurements were used to track these time-varying enthalpy components, and are discussed in 

Sections 7, 9, and 10 of this report. The value of Q (t) g

& was estimated from thermocouple 

temperature measurements made at a series of heights on the thermocouple trees. A single door 

to the compartment was open during 6 of the 15 experiments. In those tests, Q (t) d

& was 

characterized using an array of bidirectional probes and associated thermocouples. The value of 

Q (t) w

& was also measured using a series of sensors placed at select locations on the compartment 

walls, floor, and ceiling. 

Figure 3-4 considers the energy balance in the compartment during Test 3 as a function of time 

as represented by the energy loss to the compartment surfaces (Qw

& ), through the doorway (Qd& ), 

heating the gas phase (Qg

& ), and the sum of these three terms. For comparison, the corrected 

calorimetric HRR (Q& ) is also shown. Equation 3.3 states that the sum of the heat losses should 

be equal to the total HRR. Early in the experiment (t<300 s), the calorimetric results lagged the 

sum of the enthalpy heat loss terms as the hot upper layer grew and obtained a quasi-steady 
26 

temperature. The time lag was due to the measurement configuration in which the hot gases first 

filled the compartment and only then began to spill into the exhaust hood, where the HRR was 

measured. 

The sum of the heat loss terms during the steady burning period (from about 400 s to about 1400 

s) slowly increased. The mean value during the steady burning period was calculated as 

1210 kW, which is within experimental uncertainty of the time-averaged value of the 

calorimetric measurement (1140 kW) during the steady burning period and even closer to the 

mean value of the calorimetric measurements (1180 kW; Table 3-7) for the open door tests. The 

uncertainty in the sum of the heat losses was estimated as 12 %, while the uncertainty in the 

calorimetry was 15%. The agreement between these two curves in Figure 3-4 was generally 

within the overlapping uncertainty limits. 

0

200

400

600

800

1000

1200

1400

0 500 1000 1500 2000

Surface Heat Loss

Doorway Heat Loss

Gas Heating 

Sum

Calorimetry

Rate of Enthalpy Change (kW)

Time (s)

Figure 3-4: The rate of heat loss to surfaces, through the doorway, accumulated by heating 

the gas in the compartment, and their sum as a function of time during Test 3. 

3.4 Summary 

Two methods were used to estimate the fire HRR: a fuel flow rate measurement corrected for 

compartment heating effects and an oxygen calorimetry measurement. The two approaches give 

estimates of the HRR during the steady burning period (see Tables 3-2 and 3-9) that are within 

about 4 % of each other, which is insignificant if measurement uncertainties are considered. A 

third approach was demonstrated for Test 3, which involved measurements of the heat losses 

from the compartment. These three approaches led to results that were consistent, within 

experimental uncertainty. 
27 

Because the determination of the compartment heating effects on the fuel flow calibration was 

measured only once and for essentially only one test condition, the method involving oxygen 

calorimetry that was used during the six open door tests may be considered a more reliable way 

to estimate the fire HRR. This approach was taken. For each of the open door tests (Tests 3, 5, 9, 

14, 15, and 18), the values of the recommended HRRs during the steady burning period were 

based on the product of the mean correction factor (with a value of 1.14; see Table 3-7) and the 

HRR associated with the ambient temperature fuel flow calibration (from Table 3-2 and Table 3-

7). The same approach was taken for the closed door tests (Tests 1, 2, 4, 7, 8, 10, 13, 16, and 

17), and the recommended HRRs during the steady burning period were taken from Table 3-8. 

Table 3-9 summarizes the recommended HRRs during the steady burning period for all of the 

tests. The table also lists the duration (from Table 3-1) of the ramp-up, the steady burning 

period, and the ramp-down. 

Table 3-9: Estimated HRRs during the steady burning period for all of the tests and the 

duration of each of the phases of the fuel flow. 

Test Fuel Q& (kW) Ramp-Up3

(min:sec) 

Steady Burning 

Period3

(min:sec) 

Ramp-Down3

(min:sec) 

1 Heptane 4101 2:28 20:02 2:30 

2 Heptane 11901 3:00 7:25 0:01 

3 Heptane 11902 2:58 20:01 3:03 

4 Heptane 12001 2:58 10:36 0:01 

5 Heptane 11902 2:58 20:01 3:03 

7 Heptane 4001 2:09 20:03 2:08 

8 Heptane 11901 2:56 7:14 0:01 

9 Heptane 11702 2:55 20:01 3:04 

10 Heptane 11901 2:56 10:50 0:01 

13 Heptane 23301 2:57 3:07 0:01 

14 Heptane 11802 2:56 20:05 3:06 

15 Heptane 11802 3:00 20:00 3:07 

16 Heptane 23001 2:57 3:25 0:01 

17 Toluene 11601 3:01 1:31 0:01 

18 Heptane 11802 2:58 20:02 3:07 

1. The combined expanded relative uncertainty is estimated as 17 %; see Table 3-8. 

2. The combined expanded relative uncertainty is estimated as 17 %; the product of the mean correction 

factor of 1.14 (see Table 3-7) and the ambient temperature fuel flow calibration (see Table 3-2). 

3. Taken from Table 3-1.

29 

4 HEAT FLUX MEASUREMENTS 

The objective of these measurements was to provide information on the incident heat flux to the 

cable trays and cables. Because of the physical size of the measurement devices, it was not 

possible to place them on or in the test cable trays and cables. Instead, the sensors were placed 

within close proximity, on the order of 10.0 cm (3.9 in) from the targets. Both radiative and total 

heat flux were measured. The difference between these results allows estimation of the 

convective heat flux. 

4.1 Heat Flux Measurement Devices 

Five Schmidt-Boelter total heat flux gauges and five ellipsoidal radiometers were used to 

measure the total and radiative heat fluxes near the cable trays and cables. A Schmidt-Boelter 

flux gauge uses a water-cooled thermopile as a sensor whose surface temperature is uniform and 

close to that of the cooling water used. The gauge is used to measure the combined convective 

and radiative heat fluxes to the sensor surface. The gauges (Medtherm Model GTW-15SB-6-60-

40-484K) had a field of view of 180°. They are rated at 150 kW/m2, and have a time response of 

approximately 0.1 s to 0.2 s. The water used to cool the gauges was heated to 75 °C ± 3 °C (167 

°F ± 6 °F) to prevent water vapor condensation on the gauge surface during fire tests using the 

cooling system described below. Soot deposition on the gauge surfaces was blown off using 

compressed air before each test. The five total heat flux gauges used here were designated as 

Total Heat Flux Gauge 2, Total Heat Flux Gauge 4, Total Heat Flux Gauge 6, Total Heat Flux 

Gauge 8, and Total Heat Flux Gauge 9. 

The radiometers, which measure the radiative flux incident on the sensor surface, have a 

Schmidt-Boelter sensor at the base of an ellipsoidal cavity. The entrance to the cavity (≈ 3 mm 

(0.12 in) in diameter) receives radiation incident over a 160o field of view. The radiometers were 

cooled using room temperature water with a flow rate of at least 1.0 L/min (0.3 Gal/min). The 

sensors were N2-purged to prevent soot and water vapor from entering the cavity and subsequent 

deposition and condensation on the sensor surface. Three radiometers (Medtherm Model 64EP15SB-6-60-20544K, designated as Radiative Heat Flux Gauge 3, Radiative Heat Flux Gauge 5 

and Radiative Heat Flux Gauge 7) were rated to 150 kW/ m2

, and two radiometers (Medtherm 

Model 64EP-3-20544, designated as Radiative Heat Flux Gauge 1 and Radiative Heat Flux 

Gauge 10) at 30 kW/ m2

. 

The total flux gauges were re-calibrated at NIST using 75 °C ± 3 °C (167 °F ± 6 °F) cooling 

water. The radiometers could not be re-calibrated with the NIST radiant source due to focusing 

problems associated with the source; therefore, the manufacturer’s calibration was used. 

Manufacturer calibration results for the total heat flux gauges are typically within 5 % of the 

NIST calibration results, giving confidence in the given calibrations. 
30 

The gauges and radiometers were mounted in pairs and placed adjacent to each other. The pairs 

were mounted, using steel wires, nearly flush to one side of a steel L-bracket through two drilled 

holes. Figure 4-1 is a photograph that shows the relative location of Gauges 3-6. The photo was 

taken from below the gauges, which were at nearly the same height above the floor. Figure 4-2 

shows Gauges 9 and 10, which were located adjacent to the vertical cable tray. Signal and 

thermocouple wires from the gauges and cooling water supply and return lines were thermally 

protected using a thermal ceramic blanket and then wrapped with aluminum foil. A closed flow 

water-cooling system was used to circulate hot water through the total heat flux gauges (see 

Figure 4-3). Table 4-1 summarizes the location, orientation, and designation of the total heat 

flux gauges and radiometers used in this test series. 

Table 4-1: Locations of total heat flux gauges and radiometers. 

Designation Type X* 

(m) 

Y* 

(m) 

Z* 

(m) 

Orientation 

Rad. Gauge 1 Radiative 10.81 0.44 2.05 Pointing downward 

Total Gauge 2 Total 10.87 0.44 2.05 Pointing downward 

Rad. Gauge 3 Radiative 10.81 1.25 2.52 Pointing downward 

Total Gauge 4 Total 10.87 1.25 2.52 Pointing downward 

Rad. Gauge 5 Radiative 10.81 1.40 2.54 Pointing North Wall 

Total Gauge 6 Total 10.87 1.40 2.54 Pointing North Wall 

Rad. Gauge 7 Radiative 10.81 2.00 3.04 Pointing downward 

Total Gauge 8 Total 10.87 2.00 3.04 Pointing downward 

Total Gauge 9 Total 10.81 6.85 1.75 Pointing South Wall 

Rad. Gauge 10 Radiative 10.87 6.85 1.76 Pointing South Wall 

* the uncertainty in the distance is ± 0.02 m. 

X = distance from West wall 

Y = distance from South Wall 

Z = distance from floor 

The pairing of total heat flux gauge and radiometer is highlighted.

4.2 Uncertainty Analysis 

The main sources of uncertainty pertaining to the total heat flux are: (1) the uncertainty of A/D 

conversion, (2) uncertainty in the calibration, and (3) uncertainty due to soot deposition on the 

sensing surface of the gauge. The uncertainty in A/D conversion is inherent to the data 

acquisition system. It is system specific and is associated with the digitization of the analog 

signals from the gauge. This type of uncertainty was taken as negligible. The calibration from 

the manufacturer was obtained using water at 23 oC ± 3 oC (73 °F ± 6 °F). The relative expanded 

uncertainty reported by the manufacturer is ± 3 % of responsivity (the slope of the calibration 

curve) with a coverage factor of 2. 
31 

Figure 4-1: Photograph of two total (Gauges 4 and 6) and two radiative (Gauges 3 and 5) 

heat flux gauges partially insulated. 

Figure 4-2: Photograph of Gauges 9 and 10 and the vertical cable tray. 

#3 

#4

#5 

#6 

Vertical cable tray 

#9 #10
32 

filter

flow meters

water lines to 

 radiometers

and flux gages

 heating

elements

manifold

 return lines

from radiometers

and flux gages

water heater water pump

Figure 4-3: Cooling water system for total heat flux gauges. 

A linear calibration curve is given as 

Y = a + b X (4.1) 

where X is the measured signal in mV and Y is the calculated heat flux in kW/m2

. If the input Xs 

are uncorrelated and the NIST standard reference gauge used in the re-calibration has negligible 

uncertainty in the conversion factor (from mV to kW/m2

), the combined standard uncertainty uc

of a predicted value Y from the linear calibration curve can be readily obtained by applying the 

law of propagation of uncertainty. 

2 2 2

c a b u = u + X u (4.2) 

where ua and ub are the standard uncertainties of the intercept and slope of the calibration curve 

respectively. The combined standard uncertainty associated with each measurement was 

estimated to be less than 0.1 kW/m2

 for all the results reported here. 

For the total flux gauges, the uncertainty due to soot deposition is difficult to quantify. The 

amount of soot deposition depends on many parameters, such as the location of the gauge, the 

flow field and temperature fields near the gauge, the duration of a test, and the soot volume 

fraction. 

There is no effect of soot deposition on the purged radiometers. The radiometer has two main 

sources of uncertainty including (1) a 160o (less than 180o) field of view and (2) its calibration. 

If a uniform radiation field is assumed, a radiometer with a 160o field of view would receive 3 % 

less incident radiation than a hemispherical (180o) field of view. The calibration curve provided 
33 

by the manufacturer was used. According to the manufacturer’s specification, the radiometer has 

a relative expanded uncertainty of ± 3 % in responsivity with a coverage factor of 2. 

4.3 Results and Discussion 

Since the same general trends were observed for all the tests under similar experimental 

conditions, the discussion that follows focuses on Tests 2, 3, and 4. These three tests were 

selected for detailed discussion because they represent the three major experimental conditions 

(enclosure closed, enclosure opened, and enclosure closed with ventilation on) used in the test 

matrix. Experimental data is available electronically from NIST. 

Figures 4-4 – 4-6 show the heat flux measurements from Gauges 3, 4, 5, and 6 for Tests 2, 3, and 

4. Three similar prominent features are noted in the respective figures: 

(1) Upon ignition, the heat fluxes slowly increased with time as the fire developed. 

(2) After the fire reached quasi-steady state burning, the heat fluxes remained relatively constant. 

(3) Upon termination of the fuel flow, the fire diminished in size and subsequently extinguished, 

and the heat fluxes decreased with time. Note that the heat fluxes did not return to their 

initial values before ignition over the duration of the measurement because the enclosure was 

still at elevated temperature after fire-out. 

In Figure 4-4, the spikes from the pair of Gauges [5 and 6] are probably due to electronic noise. 

After the fire in Test 2 was fully developed, the radiative fluxes accounted for approximately 50 

% and 70 % of the total heat flux for pairs [3 and 4] and [5 and 6], respectively. For Test 3, the 

radiative fluxes constituted approximately 60 % and 60 % of the total heat flux for the same 

corresponding pairs of gauges. For Test 4, the radiative fluxes constituted approximately 50 % 

and 70 %, respectively. It is important to recognize that the convective component in the total 

heat flux measurement is based on the convective heat transfer to a “cold surface” total heat flux 

gauge maintained at 75 °C ± 3 °C (167 °F ± 6 °F). The measured convective component is thus 

enhanced. During a fire test, the temperature of a target surface is not constant and is rising thus 

reducing the temperature gradient between the surface and the free stream. It is imperative to 

compare the measurements with the model predictions on the same basis. Model simulations 

should consider the flux gauge temperature. 

Figures 4-7 – 4-10 show the temporal variation of the radiative and total heat fluxes for pair [10 

and 9] in Tests 2, 3, and 4 respectively. The main features of these figures remain similar to 

those discussed in Figures 4-4 – 4-6, namely, the heat fluxes increased during the development 

phase of the fire, reached a quasi-steady state, and decreased upon the termination of the tests. In 

Figure 4-7, Gauge 9 was not functioning during Test 2 due to a bad connection between the 

signal wires and the computer interface. 

Some peculiar behavior is noted in Figures 4-8 and 4-9. In Test 3, the door of the enclosure was 

open. For the first 200 s as the fire developed, the radiative flux was measured to be somewhat 

higher than the total heat flux. The difference was within the uncertainties associated with the 

measurements. However, the radiative flux was less than the total flux during the quasi-steady 
34 

burning stage. In Test 4, similar behavior was observed. In this test, the ventilation system was 

on. In Test 4, the radiative flux was comparable to the total heat flux within experimental 

uncertainties as the fire developed and also during the quasi-steady burning stage. This effect 

could have been due to increased soot deposition on the sensor surface of the flux gauge, which 

would artificially indicate a lower than normal total heat flux measurement. The location of the 

gauge and possibly the ventilation system could also be contributing factors to these 

observations. Although post-test examination of Gauge 9 could not conclusively support the 

hypothesis of increased soot deposition, a thin soot layer was observed on the sensor requiring 

cleaning with a small burst of compressed air. 

Figures 4-10 – 4-12 show the results of heat flux measurements in Tests 2, 3, and 4 for Gauges 1, 

2, 7, and 8 respectively. Again, features similar to Figures 4-4 – 4-6 are noted in these plots. 

The spikes at ≈ 600 s in Figure 4-10 were likely due to electronic interferences. 

0 200 400 600 800 1000 1200

0

2

4

6

8

10

Heat Flux (kW/m2

)

Time (s)

 3

 4

 5

 6

Figure 4-4: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 2. 
35 

0 500 1000 1500 2000

0

2

4

6

8

10

Heat Flux (kW/m2

)

Time (s)

 3

 4

 5

 6

Figure 4-5: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 3. 

0 200 400 600 800 1000 1200

0

2

4

6

8

Heat Flux (kW/m2

)

Time (s)

 3

 4

 5

 6

Figure 4-6: Heat fluxes as a function of time for Gauges 3, 4, 5, and 6 in Test 4. 
36 

0 200 400 600 800 1000 1200

0

1

2

3

4

5

6

7

8

9

Heat Flux (kW/m2

)

Time (s)

 9

 10

Figure 4-7: Radiative heat flux as a function of time for Gauge 10 in Test 2. The total heat 

flux Gauge, 9, was not functioning during this test. 

0 500 1000 1500 2000

0

1

2

3

4

5

6

7

8

Heat Flux (kW/m2

)

Time (s)

 9

 10

Figure 4-8: Radiative and total heat flux as a function of time for Gauges 9 and 10 in Test 

3. 
37 

0 200 400 600 800 1000 1200

0

1

2

3

4

5

6

7

8

Heat Flux (kW/m2

)

Time (s)

 9

 10

Figure 4-9: Radiative and total heat flux as a function of time for Gauges 9 and 10 in Test 

4. 

0 200 400 600 800 1000 1200

0

2

4

6

8

10

12

14

16

Heat Flux (kW/m2

)

Time (s)

 1

 2

 7

 8

Figure 4-10: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, and 8, in 

Test 2. 
38 

0 500 1000 1500 2000

0

2

4

6

8

10

12

14

16

Heat Flux (kW/m2

)

Time (s)

 1

 2

 7

 8

Figure 4-11: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, and 8, in 

Test 3 (Gauge 7 failed after 1000 s into the test). 

0 200 400 600 800 1000 1200

0

2

4

6

8

10

12

14

16

Heat Flux (kW/m2

)

Time (s)

 1

 2

 7

 8

Figure 4-12: Radiative and total heat flux as a function of time for Gauges 1, 2, 7, and 8, in 

Test 4. 
39 

5 SMOKE CONCENTRATION 

5.1 Principle of Operation 

The smoke concentration measurement was based on the strong light extinction characteristic of 

flame-generated smoke. This measurement was based on Bouguer’s Law, which relates the ratio 

of the transmitted (I) and incident (I0) intensities to the mass concentration of smoke Ms, the path 

length through the smoke, L, and the specific extinction coefficient φs via the following 

expression: 

exp( )

0

M L

I

I = −φ s s (5.1)

The ability to infer mass concentration from a light extinction measurement is made possible by 

the discovery that φs is nearly universal for post-flame smoke produced from over-ventilated 

fires. Previous studies indicate little change in the smoke yield as the air becomes vitiated. 

Mulholland and Croarkin [2000] found that φs has an average value of 8.7 m2

/g (42,000 ft2/lb) at 

a wavelength of 632.8 nm. The basic qualitative ideas that support this universality are that soot 

from all flames is basically carbon in the form of agglomerates with primary sphere sizes much 

smaller than the wavelength of light and a fractal dimension less than two. For these conditions 

the light absorption cross section is proportional to the mass and is the dominant contribution to 

the light-extinction coefficient. There was a smaller contribution from the light scattering crosssection, which depends on the agglomerate size. 

Solving Eq. 1 for the mass concentration, one obtains the following result: 

L

I I

M

s

s φ

ln( / ) 0 = (5.2) 

5.2 Design 

The basic instrument components were a 0.5 mW randomly polarized He-Ne laser and a silicon 

detector with a small temperature coefficient and a uniform response over its 1.0 cm × 1.0 cm 

area (0.4 in x 0.4 in). Positioning the detector 40 cm (15.8 in) from the smoke end of the purge 

tube minimized the effect of scattered light from reaching the detector (acceptance angle for 

system is approximately 0.7 o

). The laser and detector were outside of the enclosure and the 

beam was directed through the south-east corner of the enclosure 20 cm (7.9 in) from the ceiling. 

The laser end of the optics was located 1.0 m (3.3 ft) west of the corner and the detector end was 

located 0.94 m (3.1 ft) north of the corner. Key design features included attachment of the optics 
40 

to a rigid mounting bar and the use of purge air through a steel pipe to prevent smoke deposition 

on the window and to set the path length. The support bar was water-cooled and both the 

support bar and the steel pipe were insulated to minimize movement of the optics during the fire 

test. The standard instrument configuration had a 0.80 m (2.6 ft) path length for the heptane 

fires. Tubing attachments allowed for a 0.29 m (0.95 ft) path length for use with the very sooty 

toluene fire. With these path lengths, the smoke concentration range was approximately 0.02 

g/m3

 (1.2×10-6 lb/ft3) to 0.4 g/m3 (2.5×10-6 lb/ft3) for the 0.80 m (2.62 ft) path length and about 

0.05 g/m3

 (3.1×10-6 lb/ft3) to 2.0 g/m3 (1.3×10-4 lb/ft3) for the 0.29 m (0.95 ft) path length. 

Several steps were taken to verify the instrument performance for each test. The transmitted 

laser beam intensity was monitored before and after the test to assess the amount of drift in the 

laser intensity and whether a significant amount of soot was depositing on the optical surfaces. 

A neutral density filter reduced the light intensity by a factor of about 2, which was near the peak 

attenuation for most of the tests. The filter was positioned in the laser beam before each test and 

the average reduction was 2.12 with a standard deviation of 0.06. The laser beam was also 

blocked and the magnitude of the detector output decreased from a nominal 4.4 V to a signal of 

at most 2.0×10-5. This indicated that the leakage of light to the detector was negligible. 

5.3 Uncertainty 

The major sources of uncertainty were the mass specific extinction coefficient, the path length, 

and the intensity ratio. The flow of purge air was selected as a compromise between keeping the 

smoke out of the pipe and blowing the smoke out of the measurement volume. The estimated 

standard uncertainty in the path length was 0.02 m (0.06ft). The combined standard uncertainty 

for the universal specific extinction coefficient σs was 0.47 m2

/g (2,300 ft2/lb) [Mulholland and 

Croarkin, 2000]. The primary source of uncertainty in the intensity measurement was abrupt 

small changes in the laser intensity of about 0.2 volts as measured by the silicon detector 

compared to a nominal signal of 4.4 volts. For eight tests the smoke intensity was recorded after 

the smoke had been cleared from the room at the end of the tests. Such measurements assessed 

the effects of laser drift, smoke deposition, and heating effects on the alignment. The ratio of the 

initial to final voltage was computed and the mean of the eight tests was 1.000 with a standard 

deviation of 0.043. A slightly more conservative value of 0.05 was used in computing the 

standard uncertainty in ln (I0/I). Table 5-1 lists the relative standard uncertainties based on a 

mass concentration of 0.10 g/m3 (6.2×10-6 lb/ft3

). Using the law of propagation of uncertainties, 

the relative combined uncertainty was determined and then the expanded uncertainty with a 

coverage factor of two was computed as 0.18. For a mass concentration of 0.10 g/m3 (6.2×10-6

lb/ft3

), the resulting expanded uncertainty was 0.018 g/m3 (1.10×10-6 lb/ft3). 

The calculated uncertainty given above was based on a uniform concentration over the path 

length. The actual experimental concentration varies over the path length and was changing with 

respect to time. To obtain the best comparison with experiment, the path length averaged 

concentration should be computed. 
41 

Table 5-1: Standard Relative Uncertainties for a Mass Concentration of 0.1 g/m3. 

Quantity Nominal value Standard relative uncertainty 

Specific extinction 

coefficient, m2

/g 

8.7 m2

/g 0.053 

Path length, m 0.8 m 0.025 

ln (I0/I) 0.70 0.07 

5.4 Temperature Uncertainty 

A Type K thermocouple was used with a standard uncertainty of 2 oC (4 °F) based on the 

manufacturer’s specification. The expanded uncertainty (95 % confidence interval) with a 

coverage factor of two is 4 o

C (8 °F). 

The thermocouple measurement was a single point measurement made approximately at the 

midpoint of the optical path of the smoke meter. The grid cell including the thermocouple 

should be used for comparing with the thermocouple output rather than the path averaged value 

used for the smoke. 

The response time for the 24 gauge thermocouples (0.05 cm, or 0.02 in diameter) used in this 

study was about 1 s [OMEGA, 2000]. The time response was defined as the time for the 

temperature difference between the heated thermocouple and the cooler ambient to decrease by a 

factor 1/e. For the smoke results, the response time is less than 1 s. This was one reason that the 

smoke measurement fluctuates more than the temperature measurement. The fluctuations in the 

smoke concentration data were computed as a standard deviation based on 10 measurements and 

are found to be about 2 % of the mean concentration. The corresponding fluctuation for the 

temperature measurements were about 1 % of the mean temperature (in Celsius) for temperatures 

in the range 150 o

C to 300 oC (from 302 °F to 572 °F). 

5.5 Results 

Smoke data were collected on all 15 tests and summary data are provided on all these tests. We 

focus our discussion on the effects of the ventilation conditions, fire size, and fuel type on the 

smoke results. The ventilation was varied in three tests involving a 1 MW blended heptane spray 

burner located in the center of the enclosure. In Test 2 the door was closed, in Test 3 the door 

was open, and Test 4 the door was closed and ventilation was on. The temperature of the purge 

tube at the end nearest the laser was continuously monitored during the tests. Figures 5-1 – 5-3 

show that the tube was heated by no more than 20 o

C (36 °F) above ambient. The highest smoke 

concentration (0.124 g/m3

; 7.70×10-6 lb/ft3) and highest smoke temperature (240 oC; 469 °F) 

were both observed during Tests 4, when the door was closed. The smoke temperature was 

measured about 3.0 cm (1.2 in) off the axis of the laser beam near the midpoint of the optical 

path. 
42 

During Test 2, the smoke concentration increased to its peak value about 7 min after ignition, 

decreased by about 30 %, and then increased again about the time the flame went out (Figure 5-

1). This oscillatory behavior may result from first a weakening of the plume dynamics for the 

vitiated burning and then a cooling of the gases when the flame is extinguished resulting in an 

increased smoke concentration. Key results from this test and others are summarized in Table 5-

2 including peak HRR and the peak temperature. 

In Test 3, the smoke concentration increased rapidly to a relatively constant value of about 0.11 

and then abruptly decreased by about 20 % when the flame extinguished (Figure 5-2). Again, a 

small amount of oscillatory behavior was observed. The pattern was different for the case of 

ventilation with the door closed. The smoke concentration rose to a peak value and then 

decreased well before the flame extinguished. Figure 5-3 shows the results for Test 4, which had 

a somewhat smaller peak soot concentration than in Tests 2 or 3. 

Two different fire sizes were used: 1 MW and 2 MW. Test 2 was carried out with a 1 MW HRR 

and Test 13 was carried out with a 2 MW HRR. The tests were identical in all other respects. 

The peak smoke concentration is about twice as high for the larger fire as one would expect 

(Figure 5-4). An oscillatory behavior of smoke concentration versus time is observed for the 2 

MW fire as it was for the smaller fire. The fire was extinguished at about 360 s for Test 13 as 

the oxygen concentration decreased below 15 % O2 (by volume) in the lower layer; the 

corresponding time for Test 2 was 597 s. 

Test 17 was carried out using toluene. The test was identical to Test 2 in terms of heat release 

and every other parameter except the smoking tendency. Previous free burn tests for these two 

fuels indicated that the smoke yield was about 8 times larger for the mixture. As seen in Figure 

5-5, the peak concentration for Test 17 is about a factor of 8 larger than Test 2. This result is 

consistent with the smoke yield results. No oscillatory time dependence is observed for the 

heptane/toluene mixture, though in this test, the burner was on for only 273 s compared to 625 

for the heptane blend. The path length of the light extinction instrument was reduced from 80 

cm (2.6 ft) to 29 cm (0.95 ft) for the much more concentrated smoke. There appears to be a 

higher fluctuation level for the smoke data for the toluene fire. Part of the increased fluctuation 

was a result of the much shorter path length. 

There were four replicate tests for Tests 1-4. In all cases, the smoke concentrations and the 

smoke temperatures were of similar magnitude and followed the same trends for the repeat tests. 

Figure 5-6 shows that the rather complex time dependence for the smoke concentration observed 

for Test 2 with the closed door is also observed in Test 8, which is carried out for the same 

conditions as Test 2. The peak smoke concentration for Test 2 is 0.124 g/m3

 (7.70×10-6 lb/ft3) 

compared to 0.105 g/m3

 (6.50×10-6 lb/ft3) for Test 8 (See Table 5-2). The expanded uncertainty 

for the peak concentrations in Tests 2 and 8 overlap. The repeatability of the peak smoke 

concentration (3 % difference) and peak temperature (less than 1 % difference) are better for the 

open door experiments than for the closed door experiments with an average 25 % difference for 

smoke concentration and 6 % for temperature. For the closed door experiments, the peak smoke 

concentration and temperature are expected to be more variable than for the open door, because 

of a decreasing oxygen concentration in the lower layer and because of a more complicated flow 
43 

especially for the case with the cross ventilation flow from the inlet. Even with this variability, 

the fact that the difference in the smoke concentrations for repeat tests is smaller than the 

expanded uncertainty in all cases but one, Test 4/Test 10 with the cross ventilation, enhances the 

confidence in the smoke measurement. A summary of the peak concentrations, temperatures, 

and the associated times are listed in Table 5-2. 

Table 5-2: Smoke and Temperature Results. 

Test Burn 

Time, s 

Peak 

Smoke1

Conc., g/m3

Peak 

Time, s

Expanded 

Uncert., g/m3

Peak 

Temp.1, o

C 

Peak 

Time, s 

1 1500 0.045 1620 0.015 155 1375 

2 625 0.124 489 0.020 243 600 

3 1563 0.116 1010 0.020 236 1411 

4 815 0.082 499 0.017 213 832 

5 1563 0.089 1385 0.018 198 1400 

7(1)2 1332 0.060 1499 0.016 144 1360 

8(2) 613 0.105 481 0.019 228 563 

9(3) 1566 0.119 1010 0.020 235 1416 

10(4) 827 0.061 493 0.016 202 812 

13 305 0.228 405 0.030 287 346 

14 1567 0.090 1619 0.018 232 1418 

15 1567 0.130 1792 0.021 240 1409 

16 383 0.151 330 0.023 254 370 

173 273 1.007 275 0.178 160 272 

18 1567 0.114 1241 0.020 241 261 

1 The peak smoke concentration and temperature are averaged over a five s interval. 

2 The #’s in () indicate a repeat test. 

3 The path length was reduced from 80 cm to 29 cm for this measurement.
44 

0

0.03

0.06

0.09

0.12

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200

Smoke Concentration, g/m3

Temperature, o C

Time, s

Smoke temperature

Purge tube temperature

Figure 5-1: Smoke concentration and temperature as a function of time for Test 2. 

0

0.03

0.06

0.09

0.12

0

50

100

150

200

250

300

0 500 1000 1500 2000

Smoke Concentration, g/m3

Temperature, o C

Time, s

Smoke temperature

Purge tube temperature

Figure 5-2: Smoke concentration and temperature as a function of time for Test 3. 
45 

-0.025

0

0.025

0.05

0.075

0.1

0

50

100

150

200

250

0 400 800 1200

Smoke Concentration, g/m3

Temperature, o C

Time, s

Smoke temperature

Purge tube temperature

Figure 5-3: Smoke concentration and temperature as a function of time for Test 4. 

-0.05

0

0.05

0.1

0.15

0.2

0.25

0.3

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200

Smoke Concentration, g/m3

Smoke Temperature, o

C

Time, s

Conc., test 2

Conc., test 13

Temp., test 2

Temp., test 13

Figure 5-4: Smoke concentration and temperature versus time for a 1 MW (Test 2) and for 

a 2 MW (Test 13) blended heptane fire both with the door closed. 
46 

0

0.3

0.6

0.9

1.2

0

0.03

0.06

0.09

0.12

0 200 400 600 800 1000

Smoke Concentration (Test 17), g/m3

Smoke Concentration (Test 2), g/m3

Time, s

Figure 5-5: Smoke concentration versus time for toluene (Test 17) and for heptane (Test 2) 

for a 1 MW test with door closed. 

-0.03

0

0.03

0.06

0.09

0.12

0.15

0

50

100

150

200

250

0 200 400 600 800 1000 1200

Smoke Concentration, g/m3

Smoke Temperature, o

C

Time, s

Conc., test 2

Conc., test 8

Temp., test 2

Temp., test 8

Figure 5-6: Comparison of smoke concentration and smoke temperature for two replicate, 

1 MW tests with door closed (Test 2 and Test 8). 
47 

6 TARGET TEMPERATURES 

6.1 Description of Measurements 

The surface and core temperatures of Targets A-G were measured by type K thermocouples. 

Target E was a small rectangular slab (1 cm x 1 cm x 5 cm (0.4 in x 0.4 in x 2 in) long, with 

uncertainty equal to ±0.001 m (±0.003 ft)) as shown in Figures 6-1 and 6-2, which was 

manufactured at NIST using material from the outer jacket of the three conductor PVC cable. 

Three thermocouples were positioned on the slab to measure the temperature at the lower and 

upper surfaces and the slab center. Thermocouple E-TC-17 was a type K thermocouple with a 

0.16 cm (0.06 in.) diameter stainless steel sheath, with a time response estimated as 

approximately 4 s [OMEGA, 2000]. The two type K thermocouples on the surface of Target E 

were composed of insulated bare bead 24 gauge wire, with an estimated time of 1 s. 

E-TS-16’ Thermocouple 

 E-TC-17 

 E-TS-16 

PVC slab 

5 cm 

1 cm 

1 cm 

Figure 6-1: Thermocouples on the PVC slab (1 cm x 1 cm x 5 cm long) measuring 

temperatures at the lower and upper surfaces and the slab center. 
48 

Figure 6-2: Photograph looking towards the north wall with Target E present, a 

rectangular slab composed of PVC. 

Targets A-D and F were mounted horizontally. Target G was mounted vertically in a cable tray 

that contained control cables. For Test 7, 10, and 18, Targets A-D, F, and G were power and 

control cables constructed of nylon/PVC insulation and PVC jacket material. Target B was 

located at the same elevation and 0.1 m (3.9 in) from the left edge of the power cable, slab target, 

and cable tray respectively. The thermocouples were oriented such that they were at the bottom 

of the horizontal cables. Appendix A of this report describes the structure of the cable materials 

in detail. Appendix A also contains optical and thermophysical data for the XPE and PVC cable 

jackets, as well as the “marinate” calcium silicate board wall material. 

outer XPE or PVC

insulation

≈1.7 cm O.D. 

2 Thermocouples 

Figure 6-3: Photograph and schematic cutaway of a target cable instrumented with barebead type K thermocouples on the surface and just under the outer insulation. 

Type K thermocouples were set onto the cable surface and just inside the outer cable jacket 

material (see Figure 6-3) at various locations as specified in Appendix D. For placement on the 

cable surface, the insulation was first sliced. The thermocouple was then pushed into the surface 

at the location of the cut such that the thermocouple bead was partially embedded into the 

insulation. The same approach was used for mounting the thermocouples on the surface of 

Target E. 
49 

It should be noted that the cable morphology is complicated and the cables are not symmetrical 

about the cable center as seen in the detailed schematic diagram in Appendix A. Furthermore, 

the exact position of the thermocouples relative to various morphological features of the cables 

was not known. Measurement of temperatures inside and on the cable surface may be affected 

by the exact thermocouple placement relative to the cable structure. This effect is not quantified 

and further study may be needed to understand the impact of cable morphology on cable thermal 

behavior 

Two Target E slabs were fabricated using PVC cable jacket material. Target E was located in a 

high temperature region of the compartment and was initially present during Tests 2 and 10, but 

was damaged during these tests. Target E was undamaged during Tests 1 and 7. Target E was not 

present during Tests 3-6, 8, 9, and 11-18. 

For Tests 1-5, 8, 9, 13, 14 and 15, Targets A-D, F, and G were power and control cables 

constructed of XPE (flame retarded crosslinked polyethylene) insulation and Hypalon jacket 

material. Hypalon is a registered DuPont trademark for chlorosulfinated polyethylene (CSPE). 

Table 6-1 lists the designation, measurement location, and orientation of a number of 

thermocouples. All of the thermocouples on the vertical cable were positioned such that they 

faced the south wall of the compartment, whereas all of the thermocouples on the horizontal 

cables were positioned such that they faced downward. The thermocouples were type K and 24 

gauge with bead diameters of 1.05 mm ±0.04 mm (0.04 in ±0.002 in). 

For a variety of reasons (including primarily burning of cables), some of the thermocouple 

signals did not operate properly during a number of the tests. Appendix E of this report lists the 

inoperative channels, which are excluded from the companion electronic data set. 
50 

Table 6-1: Thermocouple Physical Location and Orientation on Targets B, E, and G. 

Designation Type Orientation 

A-TS-6,18,29 

A-TC-19 

on cable surface 

under surface 

facing downward 

facing downward 

B-TS-4,14,27 

B-TC-15 

on cable surface 

under surface 

facing downward 

facing downward 

C-TS-1,8,10,22,24 

C-TC-2,9,11,23,25 

on cable surface 

under surface 

facing downward 

facing downward 

D-TS-3,12,26 

D-TC-13 

on cable surface 

under surface 

facing downward 

facing downward 

E-TS-5, 16 

E-TC 17 

on top & bottom 

surfaces slab center 

facing upwards & downwards 

in center 

F-TS-7,20,30 

F-TC-21 

on cable surface 

under surface 

facing downward 

facing downward 

Vertical-TS-31,32,33,35,36

Vertical-TC-34 

on cable surface 

under surface 

facing South Wall 

facing South Wall 

* the uncertainty is estimated as ± 0.02 m (0.06 ft). 

** Target G: cable in vertical cable tray

6.2 Results and Discussion 

The transient thermocouple measurements on slab E during Test 2 are shown in Figure 6-4. The 

expanded instrument uncertainty (with a coverage factor of 2) in the thermocouple measurements 

is estimated as approximately 4 o

C (7 °F ) based on manufacturer literature. This does not 

include uncharacterized uncertainty associated with surface contact, thermocouple placement, 

etc. 

The results show that for the first 400 s after ignition the top of the slab heated the fastest, 

followed by the bottom and then the slab center. After 400 s, the thermocouple in the slab center 

exhibited a higher temperature than the slab surface, which is not expected. The softening 

temperature of the PVC used here was observed to be approximately 180 o

C (356 °F) [Harris, 

2003], which may explain the temperature trends after 400 s when polymer melting may have 

caused the observed slab temperatures. The results for Tests 3 and 4 were similar to those in 

Test 2 (see Figure 6-7 for Test 3; Test 4 not shown). The gas temperature near Slab E is 

discussed in detail in Section 9.4 of this report. 
51 

0 200 400 600 800 1000 1200

0

50

100

150

200

250

300

Temperature (C)

Time (s)

 Cable E Bottom

 Cable E Center

 Cable E Top

Figure 6-4: The transient temperatures in slab E during Test 2. 

0 200 400 600 800 1000 1200

0

50

100

150

200

250

300

Temperature (C)

Time (s)

 Cable B Surface

 Cable B Insulation

Figure 6-5: The transient temperature on Target B during Test 2. 
52 

0 200 400 600 800 1000 1200

0

50

100

150

200

250

300

Temperature (C)

Time (s)

 Cable G Surface

 Cable G Insulation

Figure 6-6: The transient temperature on the vertical cable, Target G, during Test 2. 

Test 3, E Cable Temperatures 

Time (s)

0 500 1000 1500 2000

Temperature (C)

0

50

100

150

200

250

300

Channel 237, TCEs5 (top)

Channel 249, TCEc17 (center)

Bad Channel 248, TCEs16 (bottom)

Figure 6-7: The transient temperature at three locations in slab E during Test 3. 
53 

The transient thermocouple measurements on the vertical cable (target G) during Test 2 are 

shown in Figure 6-6. The surface temperature steadily rises from ambient at ignition to nearly 

130 o

C (266 °F). As expected, the surface temperature was consistently higher than the core 

temperature. The results for Tests 3 and 4 were qualitatively similar (not shown). 

The transient thermocouple measurements on the horizontal cable (target B) during Test 2 are 

shown in Figure 6-5. The temperatures were somewhat higher than those on the vertical cable, 

which was considerably further from the fire source. The surface temperature peaked at nearly 

200 o

C (392 °F) and the core temperature was nearly 190 oC (374 °F). The temperature results 

were qualitatively similar in Tests 3 and 4 (not shown). 

The gas temperature near the targets is discussed in Section 9 of this report. 

55 

7 VENT AND DOORWAY FLOWS 

7.1 Doorway Flows 

In this section, a description of the measured mass and enthalpy flows through the compartment 

doorway is given. The doorway was open during Tests 3, 5, 9, 14, 15 and 18. To determine the 

mass and enthalpy flows into and out of the compartment, measurements of both the temperature 

field and the gas velocity field were conducted. 

To determine the velocity profile in the doorway, 14 bidirectional probes were placed in three 

vertical arrays in the door opening. The probe locations are indicated by the dark circles in 

Figure 7-1 as viewed from the outside of the compartment. The instrument locations were 

essentially symmetric. One probe (PD-12) in the lower north portion of the doorway did not 

function and is not shown in Figure 7-1. 

The gas temperatures in the doorway were measured using type-K bare bead thermocouples. A 

bare bead thermocouple was positioned 1 cm above the center of each of the bidirectional 

probes. To confirm the thermocouple measurements, a limited number of aspirated 

thermocouples were placed in the doorway. The locations of the three aspirated thermocouples 

are indicated by the “+” symbols in Figure 7-1. The results of the aspirated thermocouple 

measurements confirm that radiative flux from the fire was not affecting the bare bead 

thermocouple readings. 

The bidirectional probes were connected to MKS model 220 Baratron differential pressure 

transducers, each with a maximum differential pressure measurement capability of 133 Pa (3 

lb/ft2

). The transducers were calibrated using a Microtector Hook Gauge type manometer. The 

differential pressure, Δp , measured by each probe, and the temperature, T, at the corresponding 

probe location were used to compute the gas velocity, v, by the equation: 

ρ

1 (2 p)

K

v

Δ = , where RT

M Pabsolute ( ) ρ = . (7.1) 

In this computation, the density, ρ, was computed as a function of temperature using the ideal gas 

law. The gas was assumed to be pure air with a molecular weight (M) of 0.029 kg/mol. Here 

Pabsolute is the absolute barometric pressure and R is the universal gas constant. The K factor was 

taken as 1.08, as reported by McCaffrey and Heskestad [1976]. The bidirectional probes 

measure the flow in either direction, into or out of the compartment. A positive value denotes 

flow into the compartment and a negative value signifies flow out of the compartment. 
56 

80 cm 80 cm

40 cm

40 cm

40 cm

40 cm

20 cm

20 cm 20 cm

20 cm

PD6

PD7

PD8

PD2

PD1

PD3

PD9

PD13

PD14

PD11

PD4

PD15 PD10 PD5

TAD1

TAD2

TAD3

Figure 7-1: Locations of instrumentation in the door opening. 

7.2 Velocities and Mass Flow Through the Doorway 

A plot of the velocity profile measured using the bidirectional probes at the doorway opening for 

Test 3 is shown in Figure 7-2. The plot shows the probes located along the centerline. The 

lowest two probes recorded velocities into the compartment equal to approximately 1.0 m/s (3.3 

ft/s) during the fire. The uppermost probe measured a velocity equal to nearly -2.0 m/s (-6.6 

ft/s)(out of the compartment), where as the probe second from the top was equal to 

approximately 1.0 m/s (3.3 ft/s) during the test. The probe located in the center fluctuated 

slightly above zero indicating a small flow into the compartment at this location. The combined 

expanded uncertainty of the velocity measurements is ± 0.3 m/s (± 0.98 ft/s). The uncertainty is 

due to both scatter in the data and uncertainty in the calibration. 
57 

0 500 1000 1500 2000

-3

-2

-1

0

1

2

3

Velocity (m/s)

Time (s)

 PD 6

 PD 7

 PD 8

 PD 9

 PD 10

Figure 7-2: Velocities through the doorway measured at the centerline. 

Figure 7-3 shows the velocity profile measured on the south side of the doorway for bidirectional 

probes PD1 through PD5 in Test 3. As with the velocities measured at the centerline, the lower 

probes measured gas flow into the compartment while the upper probes measured flow out of the 

compartment. The probe located at the halfway height of the door opening was slightly negative, 

indicating the flow was going out of the compartment at this location. The velocities measured 

on the north side were similar to those measured at the south side, as shown in Figure 7-4. For 

the north side, there were only four bidirectional probes measuring the velocities. There was no 

probe at the location 60.0 cm (23.6 in) from the floor (see Figure 7-1). 
58 

0 500 1000 1500 2000

-3

-2

-1

0

1

2

3

Velocity (m/s)

Time (s)

 PD 1

 PD 2

 PD 3

 PD 4

 PD 5

Figure 7-3: Velocities through the doorway measured on the south side. 

0 500 1000 1500 2000

-3

-2

-1

0

1

2

3

Velocity (m/s)

Time (s)

 PD 11

 PD 13

 PD 14

 PD 15

Figure 7-4: Velocities through the doorway measured on the north side. 
59 

The bidirectional probes were used to measure the velocity of the gas flow through the opening, 

which was then used to compute the mass flow through the doorway. The 4.0 m2 (43 ft2

) 

cross-sectional area of the doorway was divided into 14 sections, each containing a bidirectional 

probe and a thermocouple. The instrument measurements were taken as representative of flow in 

each section. The net mass flow through the doorway was computed as the summation of the 

mass flows through each of the sections by using the equation: ∑ m& = ρvA , where A is the 

area of the section corresponding to each velocity measurement. As with the velocity 

calculations, the density was computed assuming pure air using the ideal gas law. 

During the test series, bidirectional probe PD13 experienced a failure and did not record velocity 

data for Tests 9, 15, or 18. No velocity column data is given for this location in the data files. In 

order to compute the mass flow through the doorway, the velocity data from Test 3 for PD13 was 

substituted into the mass calculations as the approximate velocity at this location. For each of 

the other bidirectional probes, the velocities measured at a particular location were very similar 

from test to test (excluding Test 5 when the ventilation was operating). Because the results were 

so repeatable, using the data from Test 3 to replace the missing data in the mass calculations 

appears to be a good approximation. 

Figure 7-5 shows the mass flow into the compartment through the doorway for Test 3. The flow 

into the compartment was predominantly through the lower sections of the doorway. The mass 

flow out of the compartment is shown in Figure 7-6. Figure 7-7 shows the net mass flow 

through the doorway for Test 3. The combined expanded uncertainty in the mass flow 

measurements is ± 0.60 kg/s (± 1.3 lb/s) and is predominantly due to fluctuations in the velocity 

measurements. Before the fuel was ignited, there was an initial net flow of approximately 

0.50 kg/s (1.1 lb/s) out of the compartment through the doorway. This was due to the close 

proximity of the doorway to the large overhead hood drawing air out of the test facility. The 

zero differential reading on the compartment pressure transducer during this time indicates that 

any mass flow out of the doorway was being replenished by flow into the compartment through 

leakage areas. When the fire was ignited, there was an initial net mass flow out of the 

compartment as the fire grew and the expanding gases pushed air out of the compartment, as 

seen in Figure 7-7. The net mass flow out of the compartment reached nearly 2.0 kg/s (4.4 lb/s). 

About 300 s into the test, the conditions equilibrated and the mass flow leaving the compartment 

was approximately equal to the mass flow entering. This equilibrium continued until the fire was 

extinguished at 1500 s. The cooling of the compartment increased the gas density and drew a 

burst of cool air into the compartment (further cooling the compartment). 
60 

-1

0

1

2

3

4

0 500 1000 1500 2000

NRCP1T3

Total Mass Flowrate (kg/s)

Time (s)

Figure 7-5: Mass flow into the compartment through the open doorway for Test 3. 

-2.5

-2

-1.5

-1

-0.5

0

0 500 1000 1500 2000

NRCP1T3

Total Mass Flowrate (kg/s)

Time (s)

Figure 7-6: Mass flow out of compartment through the upper portion of the doorway for 

Test 3. 
61 

-3

-2

-1

0

1

2

3

0 500 1000 1500 2000

NRCP1T3

Net Mass Flowrate (kg/s)

Time (s)

Figure 7-7: Net mass flow through doorway for Test 3. 

7.3 Heat Loss Through the Doorway 

The temperature and velocity measurements acquired at the door opening were used to compute 

the total heat loss through the doorway (Qd

& ): 

Q mC T vAC T

p p

&

d = ∑ & Δ = ∑ρ Δ (7.2) 

The exiting gas was assumed to be pure air, and the density, ρ, was computed using the ideal gas 

law. The heat capacity, Cp, was likewise computed for pure air as a function of temperature 

using the DIPPR 85/NIST equation for ideal gas heat capacity. The temperature difference, ΔT, 

is the difference of temperatures on either side of the doorway. The energy transfer through the 

doorway was computed for each section about the velocity and temperature measurement 

location. The summation over the entire doorway opening was computed as the net energy 

transfer through the doorway. A graph of the resulting thermal energy convected through the 

doorway for Test 9 is shown in Figure 7-8. The negative total energy indicates a heat flow out of 

the compartment. The results indicate that as much as one-third of the heat exited through the 

doorway in the 1 MW fire in Test 9. The rest of the energy went to heat the walls during the test 

(see Section 10). The combined expanded uncertainty in the energy flux measurements is 

± 40 kW. 
62 

-500

-400

-300

-200

-100

0

0 500 1000 1500 2000

NRC Test 9

Total Energy (kW)

Time (s)

Figure 7-8: The energy loss through the doorway during Test 9. 

7.4 Ventilation Supply and Exhaust 

A forced air ventilation system supplied air to the compartment through one vent and exhausted 

it through another. The two 0.70 m by 0.70 m (2.3 ft by 2.3 ft) vents were located directly 

opposite one another on the north and south walls of the compartment. The center of each vent 

opening was located 2.4 m (7.9 ft) from the compartment floor at a distance of 11.22 m (36.80 ft) 

from the west wall. A blower was attached to the vent on the south wall to feed the supply air to 

the compartment. The exhaust vent on the north wall led to the building exhaust hood. The 

ventilations system was operated during Tests 4, 5, 10 and 16. 

To determine the air flow into and out of the compartment, bidirectional probes were used to 

measure the pressure differential in the vents, Δp , which was then used to compute the velocity. 

As with the doorway probes, the ventilation probes were connected to MKS model 220 Baratron 

differential pressure transducers, with a maximum differential of 133 Pa (3 lb/ft2

). The velocity 

of the air moving through the vents was computed by the equation: 

ρ

1 (2 p)

K

v

Δ = , where RT

M Pabsolute ( ) ρ = . (7.3) 

As before, the K factor was taken as 1.08 and the density, ρ, was computed as a function of 

temperature using the ideal gas law. The supply ventilation gas was assumed to be pure air, (M = 

0.029 kg/mol). The exhaust gas, which actually included a variety of combustion products, was 

also approximated as pure air for the sake of this calculation. On the supply side, the 
63 

temperature of the air entering the supply fan outside of the compartment was used in the 

calculation. For the exhaust gas, an aspirated thermocouple at the location of the bidirectional 

probe was used to measure the temperature of the gas. 

The air flow supplied into the compartment by the ventilation system was computed from a 

57-point calibration measurement of the inlet velocities. The cross-sectional area of the supply 

vent was divided into 57 smaller areas, which were selected based on the flow gradient. The 

velocity in each of the small areas was measured using the bidirectional probe to create a 

velocity profile across the vent. Figure 7-9 shows the 57 velocity measurement locations on the 

supply side. More measurement points were concentrated in the lower section of the vent, where 

the velocity gradient was the greatest, than in the upper section, where the velocity was steady 

over larger areas. A contour diagram of the measured velocities across the duct is shown in 

Figure 7-10. The data between the measured points was interpolated to create a continuous plot. 

The greatest velocities occurred near the outer edges of the lower portion of the duct, between 10 

cm (3.9 in) and 30 cm (11.8 in) high. There were large areas where the velocities were nearly 

zero. 

The volumetric flow rate for each area was computed as the product of the velocity and the area. 

The total volumetric flow rate through the vent was then found as a summation of the flow 

measurements. The supply ventilation volumetric flow was computed to be 1.06 m3

/s (37.4 ft3/s) 

with a combined expanded uncertainty of 0.22 m3

/s (7.8 ft3/s). This corresponded to a mass flow 

rate of 1.3 kg/s ± 0.2 kg/s (2.9 lb/s ± 0.4 lb/s). The flow calibration measurements were 

performed both before and after the series of tests, and the results varied by only 3 %. 

5 10 20 35 50 60 65

5

10

20

15

30

40

50

60

0

2 1/2

0

Figure 7-9: Location of velocity measurements in the supply vent during calibration. 
64 

The air supplied by the ventilation system was not uniform when it exited the supply vent, as 

evidenced by the contour plot of the measured velocities (Figure 7-10). Most of the air was 

blown out of the bottom third of the supply duct. A visualization of the flow field was 

performed to determine the path of the air exiting the supply duct. Smoke was added to the 

supply airflow and the observed results are shown in Figure 7-11(a). Arrows indicating the path 

of the smoke show that the flow exited the supply vent moving upward at an angle of 

approximately 35 ̊ to the horizontal plane. The smoke visualization also indicated that the bulk 

of the flow exited from the mid to lower part of the duct, confirming the velocity measurements, 

which showed no flow exiting the top of the duct. To confirm that the upward motion of the gas 

was due to direction from the duct, and not the temperature of the smoke, the flow visualization 

was performed a second time with room temperature supply air, using string attached to a grid 

across the duct exit. A photograph of this is shown in Figure 7-11(b). The angle at which the 

string was deflected from the horizontal was measured. The angle varied between 25 ̊ and 50 ̊ in 

the areas where the velocity was not near zero. 

During the fire tests that had ventilation, the supply velocity was monitored with a single 

bidirectional probe in the inlet duct. The probe was placed at the center of the duct (left to right), 

on the plane of the compartment wall. For the first two tests with ventilation, Tests 4 and 5, the 

probe was positioned 30 cm (11.8 in) above the bottom edge of the duct. For tests 10 and 16, the 

probe was moved to 15 cm (5.9 in) above the bottom edge of the duct. The mean air velocity 

measured at the probe location prior to the start of the test was set to correspond to the total 

volumetric flow for the supply duct of 1.06 m3

/s (37.4 ft3/s). The volumetric flow during the test 

was computed by scaling the flow to match the change in velocity as measured by the 

bidirectional probe. Using the volumetric flow and the air density, the mass flow was then 

computed from the equation m& = ρV , where V is the volumetric flow of the duct, and ρ the 

density of the supply air. A graph of the supply vent mass flow for Test 4 is shown in Figure 7-

12. During the fire, the supply mass flow drops from an initial rate of 1.3 kg/ s ± 0.2 kg/s (2.9 

lb/s ± 0.4 lb/s) to an average mass flow of 1.0 kg/s ± 0.2 kg/s (2.2 lb/s ± 0.4 lb/s). 
65 

X Position (cm)

10 20 30 40 50 60

Y Position (cm)

0

10

20

30

40

50

60

Figure 7-10: Contour plot of supply vent velocities. 

-2 m/s

0 m/s

2 m/s

4 m/s

6 m/s

8 m/s
66 

Figure 7-11: Flow visualization of the supply duct flow field using smoke (left), and using 

string (right). 

0.8

1

1.2

1.4

1.6

1.8

0 200 400 600 800 1000 1200

NRC Test 4

Supply Mass Flow (kg/s)

Time (s)

Fuel flow stopped 

at 815 s 

Figure 7-12: Mass flow through supply vent for NRC Test 4, 1 MW, door closed. 
67 

The exhaust flow exiting the compartment when there was no test in progress was determined in 

a similar way as the supply flow, by using a bidirectional probe to measure the gas exit velocity 

along the plane compartment wall. The exhaust velocity was fairly uniform across the 

cross-sectional area of the duct, so the velocity was measured at only nine locations, near the 

center of the duct. These measurement locations are shown in Figure 7-13. The velocity was 

slightly higher towards the top of the duct, as seen in the contour plot of the exhaust velocity 

measurements, Figure 7-14. The contour plot shows only the center of the duct, starting and 

ending with the measured points and including the interpolated velocities between the 

measurements. To compute the total volumetric flow rate of the exhaust vent, the cross-sectional 

area of the duct was divided into nine areas, each containing one of the velocity measurements, 

and a summation over the areas was computed. The volumetric flow rate through the exhaust 

vent was measured to be 1.03 m3

/s (36.4 ft3/s) with a combined expanded uncertainty of 

0.20 m3

/s (7.1 ft3/s). 

During the tests, the bidirectional probe was located in the center of the exhaust exit duct, as 

shown in Figure 7-13. As in the supply duct, the bidirectional probe measured the change in 

velocity, which was then used to compute the total volumetric flow rate of the gases exiting the 

duct. The mass flow rate was then computed as the product of the volumetric flow rate and the 

density. The same density that was computed for the velocity calculation was used in the mass 

flow calculation. This meant that the combustion products in the exhaust gas were neglected and 

the density was computed for pure air, as a function of temperature, using the ideal gas law. 

20 cm

50 cm

50 cm

35 cm

0 cm 35 cm

20 cm

0 cm

Location of probe 

 during testing.

Figure 7-13: Locations where exhaust velocities were measured. 
68 

Figure 7-14: Contour plot of the exhaust vent velocities. 

The velocity monitored by the bidirectional probe located in the center of the exhaust vent 

remained fairly constant throughout the testing series when ever there was no fire present. The 

initial mass flow for all of the tests was 1.2 kg/s (2.7 lb/s) with a combined expanded uncertainty 

of 0.2 kg/s (0.4 lb/s). In the presence of a fire, the velocity, and thus the mass flow rate, 

immediately increased by about 50 % and then decreased to a steady state slightly higher than 

the original flow throughout the remainder of the test. An example of the exhaust mass flow 

measured during a test is shown for Test 10 in Figure 7-15. When the fire starts, there is a quick 
69 

increase in the mass flow through the exhaust vent. After reaching its peak 120 s into the test, 

the mass flow rate decreases to an average flow of 1.4 kg/s ± 0.2 kg/s (3.1 lb/s ± 0.4 lb/s) for the 

remainder of the burn time. The fuel flow was shut off in Test 10 after about 830 s. At that 

time, the pressure in the compartment dropped rapidly, causing a rapid drop in the mass flow 

through the exhaust duct. Similar results were measured for the other tests with ventilation. 

The single point measurements during the fire tests indicated that the flow field changed from its 

ambient values. The measured supply volume flow rate decreased during testing. For the 

exhaust, the measured volume flow rate increased during testing. The uncertainties during the 

fires should be considered substantially higher than the uncertainties in the ambient 

measurements. The ventilation system affected the compartment pressure, HGL temperature, and 

the surface temperature of various cable targets. The cable surface TCs were just outside of the 

direct path of the supply fan. In the absence of a fire, blowing was observed to flow upwards at 

about a 35° angle. 

The energy flow through the exhaust vent (Qv

& ) was calculated from the equation: 

Q mC T

p

&

v = ∑ & Δ (7.4) 

where m& is the exhaust mass flow rate, Cp is the heat capacity computed for air using the 

DIPPR 85/NIST equation for ideal gas heat capacity, and ΔT is the temperature increase. As 

with the density calculation, the exhaust gas is approximated as pure air for the purpose of 

computing the heat capacity. The temperature difference, ΔT, dominates the energy calculation. 

The temperature of the gas exiting the exhaust duct was measured by four bare-bead 

thermocouples spaced vertically along the centerline of the duct. Measurements showed a 

temperature gradient in the duct, with temperatures at the upper location reaching up to 100 ̊ C 

more than those measured at the lower location. The aspirated thermocouple located at the 

center of the duct measured the same temperature as the bare bead thermocouple at the same 

location, confirming that the thermocouples were not affected by radiation from the fire. The 

exhaust duct was divided into four area sections, each containing one of the thermocouples. The 

total energy flow through the exhaust vent was computed as the summation of the energy flow 

through each area as a function of the temperature. The combined expanded uncertainty in the 

exhaust vent energy measurements is ± 80 kW. The large uncertainty is due to uncertainties in 

the temperature over the area of the duct, since only four measurements were taken during 

testing. 

An example of the energy flow through the exhaust vent is shown for Test 10 in Figure 7-16. 

These results are for a nominally 1 MW fire centered in the compartment with the ventilation 

system operating and the compartment door closed. After an initial rapid increase, the energy 

flow continues to increase throughout the test. The fuel flow was shut off in Test 10 after about 

830 s. At that time, the pressure in the compartment dropped rapidly, causing a rapid drop in the 

energy flow through the exhaust duct. Test 5 was also a nominally 1 MW fire, but for this test 

the compartment door was open while the ventilation system was operating and some of the 

energy exited through the door opening. A plot of this is shown in Figure 7-17. Here, the energy 

flow through the exhaust vent does not continually increase, but instead it fluctuates as some of 

the hot gases are diverted to the door opening. 
70 

0

0.5

1

1.5

2

2.5

0 200 400 600 800 1000 1200

Test 10

Exhaust Mass Flow (kg/s)

Time (s)

Figure 7-15: Mass flow through exhaust vent for NRC Test 10, 1 MW, door closed. 

0

100

200

300

400

500

0 200 400 600 800 1000 1200

NRC Test 10

Energy Flow Exhaust (kW)

Time (s)

Figure 7-16: Energy flow through exhaust vent for NRC Test 10, 1 MW, door closed. 
71 

0

100

200

300

400

500

0 500 1000 1500 2000

NRC Test 5

Energy Flow Exhaust (kW)

Time (s)

Figure 7-17: Energy flow through the exhaust vent for Test 5, compartment door open. 

73 

8 GASEOUS SAMPLING 

Extractive gas concentration measurements were made at two locations in the experimental 

enclosure. Oxygen (O2), carbon monoxide (CO) and carbon dioxide (CO2) were measured high 

in the compartment and O2 was also measured low. Both sampling locations are listed in Table 

8-1. 

Table 8-1: Gaseous sampling locations. 

Measurement Distance from West 

wall (m) 

Distance from South 

Wall (m) 

Height from Floor 

(m) 

O2, CO, CO2 6.85 3.48 3.22 

O2 6.85 3.48 0.50 

8.1 Apparatus 

The upper level continuous gas sampling system was made of stainless steel tubing in the 

compartment and copper and plastic tubing outside the enclosure. The gas sample once removed 

from the compartment went through two moisture and soot traps in series filled with 8 

micrometer fiber glass Corning #3950. A GAST Model DOA-P09-FB pump was used to pull 

the sample from the compartment and send it to the flow panel. Only a small fraction of the 

sample was directed to the analyzers, most of the flow being diverted to the excess flow 

discharge. The lower level O2 sampling was similar; the only difference being that only one 

analyzer and one trap were employed, see Figure 8-1. The particulars of the analyzers used are 

listed in Table 8-2. The CO and CO2 analyzers both have auto-ranging features, allowing the 

analyzers to switch ranges as the concentration of the measured gas changes. The constants 

listed in Table 8-2 were used in the data acquisition system because these analyzers were setup in 

their auto ranging mode. 
74 

Table 8-2: Analyzer Particulars. 

O2 – 1: SERVOMEX 4100 Gas Purity Analyzer S/N 10410261 (upper level) 

Range % O2 v/v Constant 

1 0-25 % n/a 

O2 – 2: SERVOMEX Oxygen Analyzer Model 540A (lower level) 

Range % O2 v/v Constant 

1 0-25 % n/a 

CO – 3: SIEMENS Ultramat 5E S/N E9-955 

Range % CO v/v Constant 

1 0-3 % 0.623 

2 0-5 % 1.033 

3 0-10 % 2.066 

4 0-20 % 4.0 

CO2 – 4: SIEMENS Ultramat 5E S/N E9-959 

Range % CO2 v/v Constant 

1 0-5 % 1.036 

2 0-10 % 2.069 

3 0-30 % 6.217 

4 0-50 % 4.0 

Sample from enclosure

Dry Ice 

Analyzer Flow Panel

Sampling 

Pump 

Moisture and Soot Extracting Trap 

Excess flow discharge

Figure 8-1: Lower Level Gaseous System components and configuration. 
75 

8.2 Calibration Procedures 

Every analyzer was calibrated prior to each test. Nitrogen was used to zero all analyzers. A 

7.0018 % CO, 18.0012 % CO2, Nitrogen balance standard was used to span the CO and CO2

analyzers. The standard was made by Matheson Gas Products, NIST Traceable Certificate 

#107304, Lot #102-26-02403, expiration date 6/19/05. Ambient air was used to span both 

oxygen analyzers. 

8.3 Sample Transport Time and Analyzer Response Characterization 

A series of experiments were conducted to determine the sample transport times and analyzer 95 

% response times (T95). Nitrogen was used to determine these parameters for the oxygen 

analyzers. Nitrogen was introduced, not under pressure, at the sampling location and the oxygen 

analyzer response was recorded. Post test analysis revealed the 

transport times and T95 values. This experiment was conducted for each oxygen analyzer. A 

similar experiment was conducted for the CO and CO2 analyzers. In that case the CO and CO2

standard was introduced, again not under pressure, at the sampling point and system response 

documented and characterized. Table 8-3 lists the transport time and T95 values for each 

analyzer. Uncertainty is estimated as ±1.0 s. 

The reporting limit for carbon monoxide was determined by examining baseline CO data from 

multiple tests and estimating the signal noise. A noise value of 0.002 % CO was determined. A 

value of 5-times the signal to noise (S/N), or 0.01 % CO, is used as the reporting limit. 

Table 8-3: Sample transport time and analyzer response. 

Analyzer Transport Time (sec)* Analyzer 95 % Response 

Time (T95) 

O2 –1 (upper) 5.5 ±1.0 25.6 ±1.0 

CO2 – 4 7.9 ±1.0 17.8 ±1.0 

CO – 3 7.8 ±1.0 17.3 ±1.0 

O2 – 2 (lower) 9.3 ±1.0 28.7 ±1.0 

* The time that the gas was introduced to the sample line is defined as 0 s. 

8.4 Estimated Expanded Uncertainty 

Table 8-4 lists the estimated expanded uncertainties for all four continuous sampling 

measurements. These uncertainties values were estimated based on calibration, instrument drift, 

sampling, and analyzer response. Confidence level is approximately 95 %. 
76 

Table 8-4: Estimated expanded uncertainty (with an expansion factor of two). 

Analyzer Estimated Expanded Uncertainty 

(volume fraction) 

O2 –1 (upper) ±0.01 

CO2 – 2 ±0.0025 

CO – 3 ±0.0001 

O2 – 4 (lower) ±0.01 

8.5 Results and Discussion 

During Test 1, a sampling anomaly occurred in both sampling systems. At some point following 

the fuel off event, flow in both the upper layer and lower layer systems stopped as evidenced by 

data from the gas analyzers. For this reason, data are not reported for Test 1 after the fuel was 

secured at the 25 min (1500 s) mark. 

During replicate Tests 1 and 7, the upper level oxygen decreased steadily from ambient volume 

fractions at fire ignition to 14.9 % and 15.0 %, respectively, near the end of the fuel ramp down 

stage. Carbon monoxide in the upper layer remained at or below the 0.01 % reporting limit for 

both tests. During both tests, the upper layer CO2 volume fractions peaked at approximately 4 %. 

During Tests 1 and 7, the lower level O2 volume fractions decreased from 20.95 % to a minimum 

of roughly 16.9 % shortly after fuel shut off. Figures 8-2 – 8-4 are representative of the gas 

volume fractions during Tests 1 and 7. 
77 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200 1500 1800

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-2: Test 7 Upper Layer O2. 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200 1500 1800

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-3: Test 7 Lower Layer O2. 
78 

0

0.02

0.04

0.06

0.08

0.1

0 300 600 900 1200 1500 1800

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-4: Test 7 Upper Layer CO2. 

During Test 2, a sampling anomaly occurred in the upper layer gaseous sampling system. Carbon 

dioxide likely was pulled into the sampling line at the moisture and soot trap. For this reason, 

data collected from the upper layer sampling system after 640 s following fire ignition are not 

reported. 

During replicate Tests 2 and 8, the upper layer oxygen volume fraction decreased from ambient 

values at fire initiation to approximately 12 % shortly after the fuel was secured. Figure 8-5 

below is representative of the upper layer O2 volume fraction. The carbon monoxide volume 

faction in the upper layer peaked at 0.01 % during Test 2 and 0.02 % during Test 8 shortly after 

fuel was secured as represented in Figure 8-6. Similarly, carbon dioxide during both tests rose 

steadily from ambient at fire initiation to approximately 6 % shortly after fuel is secured. See 

Figure 8-7. 

Lower layer O2 volume fraction decreased from near ambient at fire initiation to minimum 

volume fractions near 13.7 % during Tests 2 and 8. Figure 8-8 and Figure 8-9 present the lower 

level O2 volume fractions during Tests 2 and 8. 
79 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-5: Test 8 Upper Layer O2. 

0

0.0001

0.0002

0.0003

0.0004

0.0005

0 300 600 900

Volume Fraction CO

Test Time (s)

Gaseous Sampling CO -3

Figure 8-6: Test 8 Upper Layer CO. 
80 

0

0.02

0.04

0.06

0.08

0.1

0 300 600 900

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-7: Test 8 Upper Layer CO2. 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-8: Test 2 Lower Layer O2. 
81 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-9: Test 8 Lower Layer O2. 

Tests 3 and 9 were replicate tests while Test 14 differed only in that it employed an off-center 

fire location. The oxygen, carbon monoxide and carbon dioxide volume fractions for all three 

tests were similar. 

The upper layer O2 volume fraction, from near ambient at fire initiation, decreased to a volume 

faction near 16 %, as shown in Figure 8-10, when fuel ramp down was initiated. Carbon 

monoxide volume fraction in the upper layer did not exceed 0.01 % for any of the three tests. 

The CO2 volume fraction in the upper level rose from ambient at fire initiation to a steady value 

of approximately 3.1 % during all three tests. The CO2 levels began to decreasing when the fuel 

ramp down commenced as displayed in Figure 8-11. 

The O2 volume fraction in the lower level changed little during the three tests. From near 

ambient at fire initiation, lower level oxygen measurements decreased to a minimum of 

approximately 20.9 % within 5 min to 6 min of fire initiation. The lower level O2 volume 

fraction then gradually returned to near ambient as shown in Figure 8-12. 
82 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200 1500 1800

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-10: Test 9 Upper Layer O2. 

0

0.005

0.01

0.015

0.02

0.025

0.03

0.035

-300 0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-11: Test 3 Upper Layer CO2. 
83 

0.206

0.2065

0.207

0.2075

0.208

0.2085

0.209

0.2095

0.21

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-12: Test 14 Lower Layer O2. 

Replicate Tests 4 and 10 exhibited upper layer oxygen volume fraction starting near ambient at 

fire initiation and gradually decreasing to approximately 13.6 % s after fuel was secured. The O2

volume fraction then slowly increased for the remainder of each test. Figure 8-13 presents the 

upper layer O2 volume fraction during Test 10. During both tests, carbon monoxide in the upper 

layer did not increase by more than 0.01 %. Carbon dioxide rose from near ambient to 

approximately 4.9 % just after fuel was secured. The CO2 levels then declined steadily as shown 

in Figure 8-14. 

In the lower layer, O2 decreased from ambient reaching a minimum volume fraction of 

approximately 14.7 % roughly 1 min after fuel was secured. As shown in Figure 8-15, the O2

volume fraction in the lower level then gradually increased for the remainder of the test. 
84 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-13: Test 10 Upper Layer O2. 

0

0.02

0.04

0.06

0.08

0.1

0 300 600 900 1200

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-14: Test 4 Upper Layer CO2. 
85 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-15: Test 4 Lower Layer O2. 

During Test 5, the upper layer oxygen volume fraction decreased from ambient at fire initiation 

to a steady concentration of approximately 18.3 %. The O2 levels began to slowly return to 

ambient during the fuel ramp down. Carbon monoxide in the upper layer was below 0.01 % at all 

times. Carbon dioxide in the upper layer increased from ambient to a steady concentration near 

1.6 %. The CO2 volume fraction began returning towards ambient during the fuel ramp down. In 

the lower layer, the O2 volume fraction did not decrease by more than 0.05 % at any time. The 

O2 and CO2 volume fractions in the upper layer are shown in Figure 8-16 and Figure 8-17. 
86 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-16: Test 5 Upper Layer O2. 

0

0.02

0.04

0.06

0.08

0.1

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-17: Test 5 Upper Layer CO2. 
87 

Replicate Tests 13 and 16 exhibited upper layer oxygen volume fractions that decreased from 

near ambient at fuel initiation to minimum values of 11.3 % and 12.2 %, respectively, after fuel 

was secured. The upper layer O2 volume fractions during Test 13 are plotted in Figure 8-18 

below. Carbon monoxide increased from ambient at fuel initiation to approximately 0.05 % and 

0.02 % during Tests 13 and 16, respectively, after fuel was secured. Figure 8-19 below is 

representative of upper layer CO during both tests. Similarly, carbon dioxide increased from 

ambient and reached a maximum volume fraction of over 6 % during both tests once fuel flow 

was stopped. See Figure 8-20. 

In the course of both tests, lower level O2 decreased from ambient levels to approximately 13.7 

% shortly after fuel flow was stopped as shown in Figure 8-21. For the remainder of the tests, the 

lower level O2 volume fraction gradually increased towards ambient levels. 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-18: Test 13 Upper Layer O2. 
88 

0

0.0001

0.0002

0.0003

0.0004

0.0005

0 300 600

Volume Fraction CO

Test Time (s)

Gaseous Sampling CO -3

Figure 8-19: Test 13 Upper Layer CO. 

0

0.02

0.04

0.06

0.08

0.1

0 300 600

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-20: Test 16 Upper Layer CO2. 
89 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-21: Test 16 Lower Layer O2. 

The upper layer O2 concentrations decreased for Tests 15 and 18 from ambient oxygen volume 

fraction at fire initiation to near steady state volume fractions of approximately 16.3 % during 

both tests. Carbon monoxide in the upper layer during both tests reached maximum volume 

fractions of 0.015 % during fuel ramp down. The maximum volume fraction of carbon dioxide 

was roughly 3.1 %, during the fuel ramp down period. Oxygen, carbon monoxide and carbon 

dioxide then gradually returned towards ambient values once the fuel was off as shown in 

Figures 8-22 – 8-24. 

 

The lower layer O2 data from Test 18 suggest that sample flow in the lower layer sampling 

system stopped approximately 6 min after the fire was initiated. The oxygen had decreased to 

roughly 20.8 % at that time. Lower level O2 during Test 15 as shown in Figure 8-25, decreased 

to approximately 20.8 % and began to return to ambient once fuel was secured. 
90 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-22: Test 18 Upper Layer O2. 

0

0.02

0.04

0.06

0.08

0.1

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-23: Test 15 Upper Layer CO2. 
91 

0

0.0001

0.0002

0.0003

0.0004

0.0005

0 300 600 900 1200 1500 1800 2100

Volume Fraction CO

Test Time (s)

Gaseous Sampling CO -3

Figure 8-24: Test 15 Upper Layer CO. 

0.2

0.202

0.204

0.206

0.208

0.21

0 300 600 900 1200 1500 1800 2100

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-25: Test 15 Lower Level O2. 
92 

During Test 17, oxygen in the upper layer decreased progressively from ambient to 

approximately 18 % shortly after fuel was secured. Carbon monoxide and carbon dioxide 

steadily increased from ambient values to maximum values of 0.1 % and 2.3 % respectively. 

Oxygen, carbon monoxide and carbon dioxide began to return to ambient concentrations once 

fuel was secured. Oxygen in the lower layer decreased from ambient to about 18.3 % roughly 1.5 

mins after fuel flow was stopped. The lower level O2 then gradually increased towards ambient. 

Gas concentration data from Test 17 are presented in Figures 8-26 – 8-29. 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -1

Figure 8-26: Test 17 Upper Layer O2. 
93 

0

0.0002

0.0004

0.0006

0.0008

0.001

0 300 600

Volume Fraction CO

Test Time (s)

Gaseous Sampling CO -3

Figure 8-27: Test 17 Upper Layer CO. 

0

0.02

0.04

0.06

0.08

0.1

0 300 600

Test Time (s)

Volume Fraction CO2

Gaseous Sampling CO2 -4

Figure 8-28: Test 17 Upper Layer CO2. 
94 

0.1

0.12

0.14

0.16

0.18

0.2

0.22

0 300 600

Test Time (s)

Volume Fraction O2

Gaseous Sampling O2 -2

Figure 8-29: Test 17 Lower Layer O2. 
95 

9 GAS TEMPERATURES 

Gas temperatures were measured using seven thermocouple trees, with each tree holding ten type 

K thermocouples. The thermocouple locations are denoted in Appendix D. 

9.1 Measurement Uncertainties 

A discussion of uncertainties is necessary before describing the results of the gas temperature 

measurements and their derived results. The uncertainties in measured temperature are primarily 

due to radiative heating and cooling of the thermocouple bead that causes it to respond to 

phenomena other than the surrounding gas temperature. Since the thermal environment 

surrounding a given thermocouple was very difficult to characterize, aspirated thermocouples, 

which do not suffer radiative exchange problems [McCaffrey and Heskestad, 1976], were used to 

check the magnitude of the needed temperature correction at a limited number of locations. 

The aspirated probes used a double-shielded design, which overcomes radiation effects on the 

bead much more effectively than a single-shielded model [Pitts et al., 1998]. The materials were 

either 304 SS or Inconel. The outer diameter of the outer shield is 0.95 cm (0.375 in). The 

probes available extended either 1.8 m (5.9 ft), 1.2 m (3.9 ft), or 7.0 m (23 ft) into the enclosure. 

To generate sufficient velocity (approximately 1.0 m/s (3.3 ft/s)) and convective heat transfer 

over the bead, gases were pulled through the probes at a minimum of 24 L/min (0.85 ft3

/min). 

The results for two bare-bead and aspirated thermocouple pairs (T5-3, TA5; T5-8, TA6) are 

considered here. Thermocouples T5-3 and T5-8 were part of Thermocouple Tree 5, which was 

located approximately 3.0 m (9.8 ft) from the center of the fire pan for Tests 1 to 13, 16, and 17, 

approximately 4.80 m (15.7 ft) from the pan center for Test 14, 1.0 m (3.3 ft) from the pan center 

for Test 18, and 0.7 m (2.3 ft) from the pan center for Test 15. The Appendix lists the exact 

locations for Thermocouple Tree 5 and the aspirated thermocouples (TA5 and TA8). Each pair 

of the two types of thermocouples was positioned within 4 cm (1.6 in) of each other. The 

differences between bare bead and aspirated thermocouple temperatures varied during each test, 

and from test to test. Table 9-1 summarizes these variations as represented by the mean and 

standard deviation of the differences that occurred during each of the tests. The magnitudes of 

the differences at these locations show that the difference between the bare-bead thermocouple 

temperature measurements and the aspirated temperature measurements were generally less than 

10 o

C (18 °F) for the lower layer and less than 5 oC (9 °F) in the hot gas layer, with the exception 

of Tests 15 and 18 where the burner was closer to the thermocouple tree resulting in much larger 

radiative effects. 

Also included in Table 9-1 are the resulting expanded temperature uncertainty limits for the 

upper and lower compartment layers. The uncertainty limits are asymmetric and consider 

radiation effects as well as the inherent uncertainty associated with thermocouple calibration. 
96 

The uncertainties are considered conservative for thermocouple measurements more than 3.0 m 

(9.8 ft) from the fire since the comparisons were made at locations near the fire plume where 

radiative effects are greatest. For thermocouples further away from the fire plume, radiation 

effects would be less significant. The results suggest that the thermocouple measurements were 

accurate within 10 o

C (18 °F), except when the fire was within 1.0 m (3.3 ft) of the 

thermocouples as occurred in Tests 15 and 18, where radiation effects are significant. 

Figures 9-1 to 9-6 are pairs of plots showing the absolute and relative differences between the 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m (3.44 ft) and 2.0 m (6.6 ft) on Tree 5 

and the nearby aspirated thermocouples (TA5, TA6) for tests 2, 3, and 4. The results of analysis 

of all the tests are contained in Table 9-1. 

Table 9-1: Calculated differences between aspirated and bare bead thermocouple 

temperature measurements and expanded temperature uncertainties. 

Differences (oC) Expanded Uncertainty Limits (oC) 

Test No. T5-3 - TA5 T5-8 - TA6 Lower Layer Hot Gas Layer 

Mean S.D. Mean S.D. Low High Low High 

1 -3.1 1.8 -3.5 0.8 -2 8 0 7 

2 1 6 -0.4 2.6 -14 12 -6 7 

3 14.9 5.2 1 3.1 -27 -3 -9 7 

4 0 4.8 1.5 2.9 -11 11 -9 6 

5 0.7 5.9 -1.6 2 -14 13 -4 7 

7 -2.3 0.9 -2.9 1.6 -1 6 -2 8 

8 -2.2 3.2 -2.2 2 -6 10 -3 8 

9 4.9 10.5 -1 2.4 -27 18 -5 7 

10 0.5 4.3 6.2 3.1 -11 10 -14 1 

13 -1.5 5.2 -1.5 3.5 -10 13 -7 10 

14 -1.1 2.4 -5.3 5.3 -5 7 -7 17 

15 75.3 98.5 -0.5 6.6 -274 123 -14 15 

16 -1.6 3.7 -1.5 3 -7 10 -6 9 

17 -2.5 2.8 -2.1 2.4 -5 10 -4 8 

18 18.5 28.5 -2.6 2.4 -77 40 -4 9 
97 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200

Test 2 T5-3

T5-8

TA5

TA6

Temperature (oC)

Time (s)

Figure 9-1: Temperature comparison as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 2.0 m on Tree 5 during 

Test 2. 

-15

-10

-5

0

5

10

15

0 200 400 600 800 1000 1200

Test 2 T5-3 - TA5

T5-8 - TA6

Temperature (oC)

Time (s)

Figure 9-2: Temperature differences as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 2.0 m on Tree 5 during 

Test 2. 
98 

0

50

100

150

200

250

300

0 500 1000 1500 2000

Test 3 T5-3

T5-8

TA5

TA6

Temperature (oC)

Time (s)

Figure 9-3: Temperature comparison as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z=1.05 m and 2.0 m on Tree 5 during 

Test 3. 

-10

-5

0

5

10

15

20

25

30

0 500 1000 1500 2000

Test 3 T5-3 - TA5

T5-8 - TA6

Temperature (oC)

Time (s)

Figure 9-4: Temperature differences as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 2.0 m on tree 5 during 

Test 3. 
99 

0

50

100

150

200

0 200 400 600 800 1000 1200 1400

Test 4 T5-3

T5-8

TA5

TA6

Temperature (oC)

Time (s)

Figure 9-5: Temperature comparison as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 2.0 m on Tree 5 during 

Test 4. 

-12

-8

-4

0

4

8

12

0 200 400 600 800 1000 1200 1400

Test 4 T5-3 - TA5

T5-8 - TA6

Temperature (oC)

Time (s)

Figure 9-6: Temperature differences as a function of time for aspirated (TA5, TA6) and 

bare-bead thermocouples (T5-3, T5-8) located at Z= 1.05 m and 2.0 m on Tree 5 during 

Test 4. 
100 

9.2 Hot Gas Layer Temperature 

Estimates of the mean temperature of the hot upper gas layer and the lower layer are reported 

here, based on the temperature measurements made at Tree 7. Tree 7 consisted of 10 

thermocouples spaced 35 cm (13.8 in) apart starting 35 cm (13.8 in) from the floor and ending 

near the ceiling. The type K thermocouples used were constructed of 24 gauge wires. The beads 

were located on individual horizontal “branches” approximately 6.0 cm (2.4 in) from the vertical 

centerline of the tree. The tree was suspended from the ceiling and attached to the floor with a 

screw to provide tension for maintaining vertical alignment. Tree 7 was approximately 6.0 m 

(19.7 ft) from the fire, so uncertainties due to radiation effects can be expected to be significantly 

smaller than those for Tree 5 (located about 3.0 m, or 9.8 ft, from the fire) discussed above. 

The estimated expanded uncertainties for the hot gas layer temperatures are contained in Table 9-

2. The uncertainties were calculated by the following procedure. Times were chosen for each 

test when the fire was most steady or nearing the time of extinguishment. Those times were: 

1300 s for tests 1, 3, 5, 7, 9, 14, 15, and 18; 600 s for tests 2 and 8; 800 s for tests 4 and 10; 300 s 

for tests 13 and 16; and 250 s for test 17. At those times, the appropriate expanded temperature 

uncertainties listed in Table 9-1 were added to the Tree 7 temperatures to obtain the 2 standard 

deviation limits around the nominal measurement. The layer interfaces were determined from 

the layer height calculations described in the next section. Four combinations of variations were 

then used: lowest limit for both lower and hot gas layers, lowest limit for lower layer and highest 

limit for hot gas layer, highest limit for lower layer and lowest limit for hot gas layer, and highest 

limit for both layers. The layer interface heights and temperatures were recalculated for these 

modified temperature profiles. The spread of the four new heights and temperatures about the 

nominal, unmodified values determined typical lower and upper expanded uncertainty bounds to 

be used for the hot gas layer temperatures and layer heights. A plot of the nominal and modified 

results for layer temperature is shown in Figure 9-7. 

Table 9-2: Hot gas layer temperature expanded uncertainty limits. 

Uncertainty Limits (oC) Uncertainty Limits (oC) Test No. 

Lower Upper 

Test No.

Lower Upper 

1 -0.2 7 10 -14 2 

2 -7 8 13 -7 10 

3 -9 7 14 -7 17 

4 -9 7 15 -6 10 

5 -4 7 16 -6 9 

7 -2 8 17 -4 8 

8 -3 8 18 -7 6 

9 -6 8 - - - 
101 

120

140

160

180

200

220

240

260

280

1 3 5 7 9 11 13 15 17

Nominal Value

Uncertainty Spread

Temperature (oC)

Test No.

Figure 9-7: The expanded uncertainty spread around nominal values of hot gas layer 

temperature for all tests except Test 18. 

Figures 9-8 to 9-11 show the temperature versus time traces of all 10 thermocouples on Tree 7 

for Tests 2 - 5. Figures 9-12 to 9-15 show the average hot gas layer temperatures for all of the 

tests versus time calculated using the two-layer reduction method. Figure 9-12 compares the 

first and repeat tests of center burner 1 MW fires under the four combinations of door 

open/closed and ventilation on/off. The repeatability of the layer temperature for these tests is 

excellent. The door closed with the ventilation off condition produced the highest temperatures 

(240 oC; 460 °F)) and at the fastest rate. The door closed with the ventilation on condition 

produced temperatures typical of the door open tests with the ventilation off (210 oC; 410 °F)), 

although less than the door closed/ventilation off condition. The door open conditions produced 

parallel layer temperatures with the ventilation off condition higher by about 40 oC (70 °F) by the 

end of the tests. 

Figure 9-13 compares the hot gas layer temperature for the three fire sizes (350 kW, 1 MW, and 

2 MW) for the condition of the door closed and the ventilation off. The temperatures and rate of 

increase of temperature are highest for the largest fire and lowest for the smallest fire. The 

repeatability of the layer temperature for these tests is also excellent. 

Figure 9-14 compares the hot gas layer temperature for two 2 MW heptane blend fires with the 

door closed and with the ventilation on and off and for two 1 MW fires with the door closed and 

ventilation off for the heptane blend and toluene. The 2 MW fire with the ventilation on had 

about 40 oC (70 °F)) lower temperatures, but it paralleled the ventilation off curve. The 1 MW 
102 

toluene fire had about 30 oC (50 °F) lower temperatures, but it paralleled the heptane blend fire 

curve. 

Figure 9-15 compares the hot gas layer temperatures for four tests with the door open and 

ventilation off for 1 MW heptane blend fires with the centered burner and three alternate burner 

positions described in the test descriptions. All of the temperature-time curves were very similar 

with continued increases for the duration, and there was no significant effect of burner position. 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200

Test 2 133 T7-10

132 T7-9

131 T7-8

130 T7-7

129 T7-6

128 T7-5

127 T7-4

126 T7-3

125 T7-2

Temperature (

124 T7-1

oC)

Time (s)

Figure 9-8: Bare bead thermocouple temperatures as a function of time for Tree 7 during 

Test 2. 
103 

0

50

100

150

200

250

300

0 500 1000 1500 2000

Test 3 133 T7-10

132 T7-9

131 T7-8

130 T7-7

129 T7-6

128 T7-5

127 T7-4

126 T7-3

125 T7-2

Temperature (

124 T7-1

oC)

Time (s)

Figure 9-9: Bare bead thermocouple temperatures as a function of time for Tree 7 during 

Test 3. 

0

50

100

150

200

250

0 200 400 600 800 1000 1200 1400

Test 4 133 T7-10

132 T7-9

131 T7-8

130 T7-7

129 T7-6

128 T7-5

127 T7-4

126 T7-3

125 T7-2

Temperature (

124 T7-1

oC)

Time (s)

Figure 9-10: Bare bead thermocouple temperatures as a function of time for Tree 7 during 

Test 4. 
104 

0

50

100

150

200

250

0 500 1000 1500 2000

Test 5 133 T7-10

132 T7-9

131 T7-8

130 T7-7

129 T7-6

128 T7-5

127 T7-4

126 T7-3

125 T7-2

Temperature (

124 T7-1

oC)

Time (s)

Figure 9-11: Bare bead thermocouple temperatures as a function of time for Tree 7 during 

Test 5. 

0

40

80

120

160

200

240

0 200 400 600 800 1000 1200 1400 1600

P1T2, Door Closed/Vent Off

P1T8, Door Closed/Vent Off

P1T3, Door Open/Vent Off

P1T9, Door Open/Vent Off

P1T4, Door Closed/Vent On

P1T10, Door Closed/Vent On

P1T5, Door Open/Vent On

Layer Temperature (oC)

Time (s)

Figure 9-12: Hot gas layer temperature as a function of time for Tree 7 during Tests 2, 8, 3, 

9, 4, 10, and 5 calculated using the two-layer reduction method. 
105 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200 1400

P1T1, 350 kW

P1T7, 350 kW

P1T2, 1 MW

P1T8, 1 MW

P1T13, 2 MW

Layer Temperature (oC)

Time (s)

Figure 9-13: Hot gas layer temperature as a function of time for Tree 7 during Tests 1, 7, 2, 

8, 13 calculated using the two-layer reduction method. 

0

50

100

150

200

250

300

0 100 200 300 400 500

P1T13, 2 MW/Vent Off/Heptane

P1T16, 2 MW/Vent On/Heptane

P1T2, 1 MW/Vent Off/Heptane

P1T17, 1 MW/Vent Off/Toluene

Layer Temperature (oC)

Time (s)

Figure 9-14: Hot gas layer temperature as a function of time for Tree 7 during Tests 13, 16, 

2, and 17 calculated using the two-layer reduction method. 
106 

0

50

100

150

200

250

0 200 400 600 800 1000 1200 1400 1600

P1T3, Burner Centered

P1T14, Burner 1.78 m from N

P1T18, Burner 1.55 m from S

P1T15, Burner 1.25 m from S

Layer Temperature (oC)

Time (s)

Figure 9-15: Hot gas layer temperature as a function of time for Tree 7 during Tests 3, 14, 

18, and 15 calculated using the two-layer reduction method. 

9.3 Hot Gas Layer Depth 

The results reported here for the hot gas layer depth are from an analysis of the temperature 

profile of thermocouple Tree 7 using the two-layer reduction method. The two-layer reduction 

model is a one-dimensional analytical method based on the zone model concept. It uses 

conservation of mass and energy and temperature profiles to define the height at which the upper 

and lower masses are equal. The equations used were: 

∫ − + = =

H

H z Tup z Tlow T z dz I

0

( int) int ( ) 1 (9.1) 

2

0

int int

( )

1 1 1 ( ) dz I

T T z

z

T

H z

H

up low

− + = = ∫ (9.2) 

I I T T H

T I I H

z

low low

low

2

( )

2

1 2

2

1 2

int

+ −

− = (9.3) 

where H is the ceiling height (m), Zint is the layer interface height (m), Tup is the hot layer 

average temperature (K), Tlow is the lower layer average temperature (K), T(z) is the temperature 

as a function of height (K), z is the height (m), I1 is the term 1, and I2 is the term 2. This method 

defines the lower layer temperature by the temperature of the lowest thermocouple. 
107 

The uncertainties in the calculation are based on the uncertainties in the bare-bead 

thermocouples. The method of estimating the uncertainties is explained in the previous section 

on the layer temperature. Table 9-3 lists the estimated expanded uncertainty limits for each test. 

Figure 9-16 shows a plot of the spread of the modified profile interface heights about the 

unmodified heights for each test except 15. 

Table 9-3: Hot gas layer interface height expanded uncertainty limits. 

Uncertainty Limits (m) Uncertainty Limits (m) Test No. 

Lower Upper 

Test No.

Lower Upper 

1 -0.01 0.03 10 -0.03 0.27 

2 -0.08 0.15 13 -0.07 0.16 

3 -0.02 0.01 14 -0.03 0.04 

4 0 0.08 15 -0.02 0.07 

5 -0.01 0.02 16 -0.03 0.07 

7 -0.07 0.10 17 -0.05 0.10 

8 -0.02 0.04 18 N/A N/A 

9 -0.01 0.03 

Figure 9-17 compares the first and repeat tests of the centered burner 1 MW fires for the four 

combinations of door open/closed and ventilation on/off. The repeatability of the layer height 

for the three pairs of similar tests is excellent. The results for the condition of the door closed 

with the ventilation on and off produced similar curves and had deeper layers than those with the 

door open. The door closed tests experienced oscillating layer heights. The algorithm used to 

estimate layer height from thermocouple measurements was developed for scenarios in which 

two distinct layers were present, and has not been validated for cases when there is no ventilation 

and/or the doors are closed. The assumption that there is a cooler lower layer may not hold for 

the tests with the door closed (Tests 1, 2, 4, 7, 8, 10, 13, 16, and 17). In tests 4 and 10, the fire 

was turned off for safety reasons after about 800 s, with the ventilation staying on, resulting in 

severe oscillations shown in Figure 9-17. The door open tests (Tests 3, 5, 9, 14, 15, 18) showed 

no significant differences in location of the interface height or temporal behavior, even between 

ventilation on and off. For the door open tests, the height was very steady after the first few 

minutes. 

Figure 9-18 compares the hot/cold gas interface height for the three fire sizes (350 kW, 1 MW, 

and 2 MW) for the condition of the door closed and the ventilation off. The two-layer reduction 

calculation yields results that are not physical for these closed-door tests, with the layer height 

about 1 m off the floor. The 350 kW fires were extinguished at about 1400 s, the 1 MW fires at 

about 600 s and the 2 MW fire at about 350 s. It is interesting to note that the larger fires’ 

interface heights dropped at faster rates than the smaller fires. 
108 

Figure 9-19 compares the hot/cold gas interface height for two 2 MW heptane blend fires with 

the door closed and with the ventilation on and off and for two 1 MW fires with the door closed 

and ventilation off for the heptane blend and for toluene. Again, as seen in Figure 9-18, the twolayer reduction calculation yields results that are not physical for these closed-door tests, with the 

layer height about 1.0 m (3.3 ft) off the floor. 

Figure 9-20 compares the hot/cold gas interface heights for four tests with the door open and 

ventilation off for 1 MW heptane blend fires with the centered burner and three alternate burner 

positions described in the test description. All of the temperature-time curves were very similar 

with steady values around 1.3 m (4.3 ft), and there was no significant effect of burner position. 

0.8

0.9

1.0

1.1

1.2

1.3

1.4

1 3 5 7 9 11 13 15 17

Nominal Value

Uncertainty Spread

Layer Height (m)

Test No.

Figure 9-16: The expanded uncertainty spread around nominal values of the hot/cold gas 

layer interface height for all tests except Test 18. 
109 

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 200 400 600 800 1000 1200 1400 1600

P1T2, Door Closed/Vent Off

P1T8, Door Closed/Vent Off

P1T3, Door Open/Vent Off

P1T9, Door Open/Vent Off

P1T4, Door Closed/Vent On

P1T10, Door Closed/Vent On

P1T5, Door Open/Vent On

Layer Height (m)

Time (s)

Figure 9-17: Layer interface height as a function of time for Tree 7 during Tests 2, 8, 3, 9, 

4, 10, and 5 calculated using the two-layer reduction method. 

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 200 400 600 800 1000 1200 1400

P1T1, 350 kW

P1T7, 350 kW

P1T2, 1 MW

P1T8, 1 MW

P1T13, 2 MW

Layer Height (m)

Time (s)

Figure 9-18: Layer interface height as a function of time for Tree 7 during Tests 1, 7, 2, 8, 

and 13 calculated using the two-layer reduction method. 
110 

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 100 200 300 400 500

P1T13, 2 MW/Vent Off/Heptane

P1T16, 2 MW/Vent On/Heptane

P1T2, 1 MW/Vent Off/Heptane

P1T17, 1 MW/Vent Off/Toluene

Layer Height (m)

Time (s)

Figure 9-19: Layer interface height as a function of time for Tree 7 during Tests 13, 16, 2, 

and 17 calculated using the two-layer reduction method. 

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 200 400 600 800 1000 1200 1400 1600

P1T3, Burner Centered

P1T14, Burner 1.78 m from N

P1T18, Burner 1.55 m from S

P1T15, Burner 1.25 m from S

Layer Height (m)

Time (s)

Figure 9-20: Layer interface height as a function of time for Tree 7 during Tests 3, 14, 18, 

and 15 calculated using the two-layer reduction method. 
111 

9.4 Gas Temperature at Targets B & E 

The gas temperature nearest Targets B & E is best represented by the thermocouple at position 8 

on thermocouple Tree 4. Tree 4 was very close to the supply vent. Position 8 was located 2.8 m 

(9.2 ft) above the floor. In Table 9-1, the expanded uncertainties for the hot gas layer 

temperature are associated with this temperature measurement. 

Figure 9-21 compares the first and repeat tests of the centered burner 1 MW fires for the four 

combinations of door open/closed and ventilation on/off. The repeatability of the temperature 

for these tests is generally less than 5 o

C to 10 oC (9 °F to 18 °F). Unlike the layer behavior 

where the door had the strongest effect, the ventilation condition seems to be more important for 

these results. The highest temperatures were with the ventilation off and the door closed, which 

also produced a much faster rate of increase. The tests with the ventilation on were very similar 

in both the open and closed door tests, with the closed door condition producing a slightly higher 

temperature (+20 oC, or +36 °F) by the end of the experiments. Figure 9-22 compares the hot gas 

layer temperature for the three fire sizes (350 kW, 1 MW, and 2 MW) for the condition of the 

door closed and the ventilation off. The temperatures and rate of increase of temperature are 

highest for the largest fire and lowest for the smallest fire. The repeatability of the layer 

temperature for these tests was generally less than 10 o

C (18 °F). 

Figure 9-23 compares the hot gas layer temperature for two 2 MW heptane blend fires with the 

door closed and with the ventilation on and off and for two 1 MW fires with the door closed and 

ventilation off for heptane blend and toluene. The 2 MW fire with the ventilation on had about 

90 oC (160 °F) lower temperatures by the end of the experiments. At 250 s in the 1 MW toluene 

fire, the temperature of Tree 4-8 was about 75 oC (167 °F), while the temperature in the heptane 

blend fire was about 160 oC (320 °F). 

Figure 9-24 compares the hot gas layer temperatures for four tests with the door open and 

ventilation off for 1 MW heptane blend fires with the centered burner and three alternate burner 

positions described in the test description. Tests 3, 14, and 15 had very similar temperature-time 

curves with continued increases for the duration of the experiments. There was no significant 

effect of burner position. Test 15 had a much higher temperature, which may be accounted for 

by the close proximity of the burner to the thermocouple tree. 

Figure 9-25 compares the thermocouple Tree 2-5 during the first test and the repeat tests with a 1 

MW fire burning in the center of the compartment for the four combinations of door open/closed 

and ventilation on/off. The repeatability of the temperature for these tests was generally less 

than 5 o

C (9 °F). The behavior of the Tree 2-5 thermocouple is different than that of Tree 4-8 

and is very similar to the results for the hot gas layer average temperatures. Like the layer 

temperatures, but unlike TC4-8, the door condition is more important than the ventilation 

condition. The door closed with the ventilation off condition produced the highest temperatures 

(230 oC; 446 °F) and at the fastest rate. The door closed with the ventilation on condition 

produced temperatures typical of the door open tests with the ventilation off (215 oC; 419 °F), 

although less than the door closed/ventilation off condition. The door open conditions produced 
112 

parallel temperatures with the ventilation off condition, higher by about 25 oC (45 °F) by the 

ends of the tests. 

0

50

100

150

200

250

0 200 400 600 800 1000 1200 1400 1600

P1T2, Door Closed/Vent Off

P1T8, Door Closed/Vent Off

P1T3, Door Open/Vent Off

P1T9, Door Open/Vent Off

P1T4, Door Closed/Vent On

P1T10, Door Closed/Vent On

P1T5, Door Open/Vent On

Temperature (oC)

Time (s)

Figure 9-21: Temperature versus time for the thermocouple located at position 8 on Tree 4 

and also near Targets B and E for Tests 2, 8, 3, 9, 4, 10, and 5. 
113 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200 1400

P1T1, 350 kW

P1T7, 350 kW

P1T2, 1 MW

P1T8, 1 MW

P1T13, 2 MW

Temperature (oC)

Time (s)

Figure 9-22: Temperature versus time for the thermocouple located at position 8 on Tree 4 

and also near Targets B and E for Tests 1, 7, 2, 8, and 13. 

0

50

100

150

200

250

300

350

400

0 100 200 300 400 500

P1T13, 2 MW/Vent Off/Heptane

P1T16, 2 MW/Vent On/Heptane

P1T2, 1 MW/Vent Off/Heptane

P1T17, 1 MW/Vent Off/Toluene

Temperature (oC)

Time (s)

Figure 9-23: Temperature versus time for the thermocouple located at position 8 on Tree 4 

and also near Targets B and E for Tests 13, 16, 2, and 17 (labeled P1T13, P1T16, P1T2, and 

P1T17, respectively). 
114 

0

200

400

600

800

0 200 400 600 800 1000 1200 1400 1600

P1T3, Burner Centered

P1T14, Burner 1.78 m from N

P1T15, Burner 1.55 m from S

P1T15b, Burner 1.25 m from S

Temperature (oC)

Time (s)

Figure 9-24: Temperature versus time for the thermocouple located at position 8 on Tree 4 

and also near Targets B and E for Tests 3, 14, 15, and 18 (labeled P1T3, P1T14, P1T15, and 

P1T15b, respectively). 

0

40

80

120

160

200

240

0 200 400 600 800 1000 1200 1400 1600

P1T2, Door Closed/Vent Off

P1T8, Door Closed/Vent Off

P1T3, Door Open/Vent Off

P1T9, Door Open/Vent Off

P1T4, Door Closed/Vent On

P1T10, Door Closed/Vent On

P1T5, Door Open/Vent On

Temperature (oC)

Time (s)

Figure 9-25: Temperature versus time for the thermocouple located at position 5 on Tree 2 

and also near Target G for Tests 2, 8, 3, 9, 4, 10, and 5 (labeled P1T2, P1T8, P1T3, P1T9, 

P1T4, P1T10, and P1T5, respectively). 
115 

9.5 Gas Temperature at Target G 

The gas temperature nearest Target G is best represented by the thermocouple at position 5 on 

thermocouple Tree 2. Position 5 was located 1.75 m (5.74 ft) above the floor. In Table 9-1, the 

expanded uncertainties for the hot gas layer temperature are associated with this temperature 

measurement. 

Figure 9-26 compares the Tree 2-5 temperature for the three fire sizes (350 kW, 1 MW, and 2 

MW) for the condition of the door closed and the ventilation off. The temperatures and rate of 

increase of temperature are highest for the largest fire and lowest for the smallest fire. The 

repeatability of the temperature for the two pairs of similar tests is also excellent. 

Figure 9-27 compares the hot gas layer temperature for two 2 MW heptane blend fires with the 

door closed and with the ventilation on and off and for two 1 MW fires with the door closed and 

ventilation off for heptane blend and toluene. The 2 MW fire time-temperature curves were very 

similar with no significant effect of ventilation condition. The 1 MW toluene fire and 1 MW 

heptane blend fire time-temperature curves were also very similar with no significant effect of 

fuel type. 

Figure 9-28 compares the hot gas layer temperatures for four tests with the door open and 

ventilation off for 1 MW heptane blend fires with the centered burner and three alternate burner 

positions described in the test descriptions. All of the temperature-time curves were parallel with 

some differences related to location. The maximum temperatures reached were 265 oC (509 °F) 

for Test 14, 220 oC (428 °F) for Test 3, 210 oC (410 °F) for Test 18, and 200 oC (392 °F) for Test 

15. 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200 1400

P1T1, 350 kW

P1T7, 350 kW

P1T2, 1 MW

P1T8, 1 MW

P1T13, 2 MW

Temperature (oC)

Time (s)

Figure 9-26: Temperature versus time for the thermocouple located at position 5 on Tree 2 

and also near Target G for Tests 1, 7, 2, 8, 13 (labeled P1T1, P1T7, P1T2, P1T8, and 

P1T13, respectively). 
116 

0

50

100

150

200

250

300

350

400

0 100 200 300 400 500

P1T13, 2 MW/Vent Off/Heptane

P1T16, 2 MW/Vent On/Heptane

P1T2, 1 MW/Vent Off/Heptane

P1T17, 1 MW/Vent Off/Toluene

Temperature (oC)

Time (s)

Figure 9-27: Temperature versus time for the thermocouple located at position 5 on Tree 2 

and also near Target G for Tests 13, 16, 2, and 17 (labeled P1T13, P1T16, P1T2, and 

P1T17, respectively). 

0

50

100

150

200

250

300

0 200 400 600 800 1000 1200 1400 1600

P1T3, Burner Centered

P1T14, Burner 1.78 m from N

P1T15, Burner 1.55 m from S

P1T15b, Burner 1.25 m from S

Temperature (oC)

Time (s)

Figure 9-28: Temperature versus time for the thermocouple located at position 5 on Tree 2 

and also near Target G for Tests 3, 14, 15, and 18 (labeled P1T3, P1T14, P1T15, and 

P1T15b, respectively). 
117 

10 HEAT LOSS TO THE ENCLOSURE 

10.1 Introduction 

In order to evaluate the fire models, it is important to have measurements with estimated 

uncertainties of the heat lost to the walls, ceiling, and floor of the enclosure as well as through 

the vents. The resulting energy balance is important when evaluating how the models handle 

energy transport. Measurements and analysis were performed to generate this energy balance for 

the test scenarios that incorporated a 1 MW fire in the center burner location with and without 

door vents, but without the mechanical vents. The result of this testing and analysis are a 

valuable, previously unavailable validation tool. 

10.2 Measurement Design 

In order to estimate the total heat loss in the test enclosure, measurements of heat flux to the 

interior surfaces were required at numerous locations. The primary design choices to accomplish 

this were the kind of measurements to make, the sensors to use, the number of sensors necessary, 

and their location. 

The number of heat flux measurement locations was limited by resource availability. The 

measurement approach was to optimize the placement and sensor type through analysis using 

computational modeling. 

To determine how many measurements were necessary, an evaluation of sensor density versus 

uncertainty level was conducted. The NIST Fire Dynamics Simulator (FDS) was used to model 

the enclosure with a 1 MW fire located in the center with the door open [McGrattan et al., 2001]. 

Among many outputs, the model predicted the temperature and radiation fields on the interior 

surfaces of the enclosure. The sensor placement was analyzed by testing the effect of the 

number and placement of nodes used in the calculation relative to the calculated value using all 

of the nodes. The accuracy of the calculation was sensitive to the distribution pattern as well as 

the number of nodes used. A non-uniform approach worked best with more nodes used in 

regions of high flux gradients and fewer nodes used where fluxes were more uniform. This 

finding was translated from the model prediction to physical test design by generating patterns of 

sensor locations that provided greater concentrations of measurements to characterize small areas 

of high flux and more sparse measurements far from the burner to characterize larger areas of 

more uniform, lower fluxes. 

The uncooled gauges used here measured the net heat flux. The gauges were sensitive to total 

heat flux including both radiation and convection. The net heat flux takes into account the 

incident total heat flux less re-radiation to the environment from the gauge and is a function of 

the temperature and emissivity of the gauge. The net radiation to the wall (not the sensor) is 
118 

equal to the net flux to the gauge adjusted by the difference between the net flux to the gauge and 

to the adjacent wall material. This depends on the differences between their temperatures and 

emissivities. The emissivities of the gauges and wall materials were measured or provided by 

the manufacturers. For the two types of gauges used here, the emissivities averaged about 0.94 

over the infrared spectrum. The temperatures of the gauges and surfaces near the gauges were 

measured to estimate the net flux to the compartment walls. 

To estimate the net flux to the compartment walls, the temperatures of the gauge as well as the 

compartment surface near the gauge were measured in case their temperatures differed. The 

gauge temperature was generally lower than the adjacent surface temperature by a few degrees to 

tens of degrees. There are a few possible explanations for this. While the emissivities of the 

gauge and wall were approximately the same at ambient temperature, they may have varied as 

the temperature increased in the compartment. The gauge also has a finite mass with the 

thermocouple actually embedded within the gauge. This causes an additional heat capacity and 

time lag not experienced on a bare surface. While the heat flux gauges were attached with a 

conductive paste, the contact and level of conduction may not have allowed sufficient thermal 

communication between the gauge and wall. The quality of the surface thermocouple 

installation may have also affected the surface temperature measurement. Finally, near the fire 

and especially in the lower layer, the surface temperatures could have been overestimated by 

thermocouples by some amount due to radiative heating of the thermocouple bead.

For the calculations of heat loss to the enclosure, (which were performed for Tests 2 – 5), the 

gauge net heat flux data was transformed to compartment surface net heat fluxes by correcting 

for the temperature difference between the compartment surface and the gauge according to the 

following equation: 

qwall &′′ = qgauge &′′ - σ (εwallTwall

4

-εgaugeTgauge

4

) 10.1 

where qwall &′′ is the net heat flux to the compartment surface, qgauge &′′ is the net heat flux to the 

gauge, σ is the Stefan- Boltzmann constant and ε is the emissivity (of the marinite and the 

gauge). For all tests, the data in the associated data files represents the net heat flux to the gauge 

with no temperature adjustment. 

10.3 Instrumentation 

Preliminary modeling of the test fires predicted heat fluxes to average about 2 kW/m2 on the 

compartment walls with some regions reaching several times that and much of the enclosure 

receiving less than 1 kW/m2

. Modeling also predicted surface temperatures greater than 150 oC 

(302 °F) on relatively small areas near the burner, while the remaining surface temperatures were 

predicted to be less than 150 oC (302 °F). Because of the temperature variations, different types 

of sensors were selected for the high and low temperature regions. Low cost, uncooled heat flux 

sensors were used in regions expected to remain below 150 oC (302 °F). More expensive, hightemperature heat flux gauges were used in the higher flux regions near the fire. 
119 

The lower temperature gauge was a Vatell model BF-04 with an embedded K-type thermocouple 

and utility up to 150 oC (302 °F). The higher temperature gauge was ITI Model C Polyamide 

Heat Flux Transducer with an embedded type-K thermocouple usable to 300 oC (572 °F). The 

gauges were mounted with Omega Engineering’s Omegatherm thermally conductive high 

temperature grease, model number OT-201. Four 6-32 screws and washers (0.79 cm OD, 0.37 in 

ID) were used to hold the corners against the grease and prevent slipping from vertical surfaces 

or motion due to temperature cycling. Each heat flux gauge had an associated, embedded 

thermocouple to determine the surface temperature of the gauge and also one located 

approximately 5 cm (2 in) away on the enclosure surface for comparison. 

The following tables, 10-1 to 10-3, describe the type and location, and area represented by the 

sensor groups. Figures 10-2 to 10-4 show schematics of the approximate sensor locations and 

coverage areas on component surfaces. Sensor Package L consisted of a lower temperature 

Vatell gauge and two type-K thermocouples. Sensor Package H consisted of a higher 

temperature ITI gauge and two type-K thermocouples. Both large side walls had the same sensor 

distribution pattern, both smaller end walls also had matching patterns, and the ceiling and floor 

had matching patterns. The total number of sensor packages was 36 of which 8 were for higher 

temperature. 

Figure 10-1: A Vatell model BF-04 heat flux gauge installed on the South wall along with a 

wall thermocouple. 
120 

Table 10-1: North and south wall sensor locations. 

Sensor Package Number North South 

Location (X m,Z m) Area Represented 

(% of surface, m2

) 

1 L L 3.91, 1.49 19.5, 15.7 

2 L L 3.91, 3.72 16.8, 13.5 

3 L L 9.73, 1.86 13.7, 11.1 

4 H L 12.15, 1.86 13.7, 11.1 

5 L L 17.79, 1.49 19.5, 15.7 

6 L L 17.79, 3.72 16.8, 13.5 

Totals 5 L, 1 H 6 L, 0 H 100.0, 80.7 (per wall) 

X

Z3.72 m

21.7 m

1

2

5

6

3 4

Figure 10-2: North and south Wall Sensor Locations. 

Table 10-2: Ceiling and floor sensor locations. 

Sensor Package Number Floor Ceiling 

Location (X m,Y m) Area Represented 

(% of surface, m2

) 

1 L L 3.04, 3.58 35.0, 54.3 

2 L H 9.11, 1.99 5.0, 7.8 

3 L H 9.11, 5.96 5.0, 7.8 

4 L H 10.85, 2.38 5.0, 7.8 

5 H H 10.85, 5.16 5.0, 7.8 

6 L H 13.02, 1.99 5.0, 7.8 

7 L H 13.02, 5.96 5.0, 7.8 

8 L L 18.66, 3.58 35.0, 54.3 

Totals 7 L, 1 H 2 L, 6 H 100.0, 155.2 (per surface) 
121 

X

Y 7.15 m

21.7 m

1

2

5

4 6

3 7

8

Figure 10-3: Ceiling and floor Sensor Locations. 

Table 10-3: East and west wall sensor locations. 

Number Sensor Package Location (Y m,Z m) Area Represented 

(% of surface, m2

) 

1 L 1.59, 1.12 25.0, 6.7 

2 L 1.59, 2.42 25.0, 6.7 

3 L 5.76, 1.12 25.0, 6.7 

4 L 5.76, 2.42 25.0, 6.7 

 4 L, 0 H (each wall) 100.0, 26.6 (per wall) 

Y

Z 3.72 m

7.15 m

1 3

2 4

Figure 10-4: East and west wall sensor locations. 

10.4 Measurement Uncertainties 

The number of sensors used and their location pattern contributed to the uncertainty estimate for 

the flux to the compartment surfaces. The model predictions were used to estimate uncertainty 

levels produced by the particular sensor pattern used. Also, uncertainties associated with 

emissivities and temperature measurements were included. 

All of the heat flux gauges were provided with factory calibrations with uncertainties of ±3 %. 

The calibration of one of each model was checked at NIST to determine if there was a significant 

difference between the factory and NIST calibrations. To check the calibration, each gauge was 

mounted on a 30 cm (11.8 in) square piece of the enclosure wall material, Marinite, and placed at 

predetermined positions in front of a radiant panel. The heat fluxes at the positions in front of 
122 

the panel were previously measured with a cooled reference gauge. Since the gauges’ output 

changed during heating, the gauge temperature as well as the flux output was measured. 

Table 10-4 lists the sources of uncertainty in the heat loss measurements. The first source, 

factory calibration of the gauges, is the uncertainty reported by the manufacturers. The 

“additional calibration” uncertainty in Table 10-4 resulted from the differences experienced 

during the check, described above, of the factory calibration with the gauges mounted in a 

configuration similar to the actual test situation. The surface temperature uncertainty arises 

because the gauge flux output is corrected to account for the difference between the gauge 

temperature and the wall temperature. This has some uncertainty due to radiation effects and 

uncertainty inherent to both thermocouples. The active area uncertainty is related to the fact that 

the screws and washers used to mount the gauges covered up small portions of the corners of 

each gauge which could lead to more or less flux depending on how much the metal conducted 

heat to a gauge. This blockage varied slightly with each installation so a maximum uncertainty 

based on the relative geometry of the gauge and washers was assigned to it. The emissivity 

uncertainty was determined from the manufacturer’s measurement for the lower temperature 

gauge and from a NIST measurement for the higher temperature gauge. Finally, the table shows 

the typical uncertainty due to the difference from the model average flux per region and the 

average flux using the model values at only the actual sensor locations. 

The total uncertainty resulting from adding these component uncertainties in quadrature is 11 %. 

The difference between the factory calibration and the calibration for a gauge mounted on 

Marinite (per radiant panel experiments) is the driving uncertainty source. The uncertainty 

components such as the surface temperature are generally much less than ±8 %. The uncertainty 

due to using the model to design the sensor layout is unknown. Because a simulation of Test 3, 

with the open door and no ventilation, was used to determine the locations of the sensors, the 

heat loss results for Tests 2, 4, and 5, which had closed doors and/or the ventilation on, may be 

less robust than those for Test 3. The uncertainty for gauge data during Tests 7, 10, 13, 14, 15, 17 

and 18 was greater than for the other tests (and is difficult to quantify) since most of the gauges 

experienced higher than their stated temperature limits during the 2 MW tests (Tests 13 and 16), 

although they were protected by a cover. 

Table 10-4: Overview uncertainties of net heat flux. 

Source of Uncertainty Expanded Relative Uncertainty 

(2 standard deviations) 

Factory Calibration ±3 % 

Additional Calibration ±8 % 

Surface Temperature Uncertainty typically ±5 % 

Active Area ±2 % 

Emissivity ±4 % 

Sensor Average Difference from Model Varied per sensor location, averaged ±2 % 

Total Uncertainty 11 % 
123 

10.5 Results 

The only tests for which heat losses to the enclosure are reported are Tests 2, 3, 4, and 5. For 

other tests, the instrumentation was inoperative, which was likely due to overheating of the 

gauges. After 1150 s in Test 3, some adjustment to the data was required. From 1150 s to 1300 

s, symmetric sensor (S3, N4) data were substituted for the failed sensor (N3, S4) data. After 

1300 s, both of a pair (S3, N3) of symmetric sensors failed so the values were retained constant 

at their last operating values until the time of extinguishment. Therefore, for Test 3, the heat 

fluxes for the north and south walls and the total heat loss were less accurate after t=1150 s. As 

was described in the uncertainty section, the most confidence in the sensor layout resides with 

Test 3 which had the door open, but the ventilation off. 

Figure 10-5 is a comparison of net heat fluxes to the wall (see Eq. 10.1) near sensors N1 and N3 

on the North wall for Tests 2 to 5. Sensor N1 was located in a low flux region and sensor N3 

was in a high flux region. The high flux region sensor, N3, experienced very similar fluxes in 

Tests 2 and 4 and in Tests 3 and 5. Each of those test pairs had the same doorway condition. 

The low flux region sensor, N1, experienced different fluxes for each test condition. In general 

for all of the tests, the N3 sensor experienced about two times the flux as the N1 sensor. 

Figure 10-6 shows the net heat fluxes on each surface as well as their sum. The symmetric walls 

(north & south, east & west) experienced nearly identical fluxes as would be expected. The 

ceiling flux was not significantly different than the flux to the floor. As the upper layer heated, 

the temperature difference with the plume decreased and the reradiation from the ceiling was 

larger so the net flux decreased. The floor flux increased during the fire, probably because the 

whole upper layer grew hotter causing more radiative heating to the relatively cool floor. Test 5 

is the test most similar to Test 3 in that the door was open, although the ventilation was on. 

Figure 10-7 shows the plot of the component and total net heat fluxes. Test 5 shows very similar 

behavior to Test 3 with similar symmetry and trends. 

Test 2 had the door closed and the ventilation off. Figure 10-8 shows the corresponding net heat 

flux plot for Test 2. Test 4 had the door closed and the ventilation on. Figure 10-9 shows its net 

heat flux plot. For both Tests 2 and 4, the symmetry between north and south and east and west 

is still apparent. The net heat flux trends with time are similar between Tests 2 and 4, but 

somewhat different from those in tests 3 and 5. In Tests 2 and 4, the ceiling flux is always higher 

than the floor flux and both ceiling and floor fluxes tend to increase throughout. 

Figure 10-10 represents an accounting of the energy balance in the compartment. The figure 

shows the total energy loss to compartment surfaces, through the doorway, and accumulated in 

the upper layer, the sum of these three terms, and the calorimetric HRR (see Section 3). The term 

accounting for heating of gases in the compartment represented less than 4 % of the total 

enthalpy [Hamins et al., 2005] - a relatively small fraction. For early times (t < 250 s), the 

calorimetry results lag behind the heat loss to the walls as the compartment volume fills before 

combustion products are transported to the exhaust hood where the calorimetry measurements 

are made. The total heat loss rate and the calorimetry HRR should, in theory, be equal, apart 

from the time lag. The uncertainty in the total heat loss is estimated to be 12 %, while the 

uncertainty in the calorimetry is 15%. The agreement between the two curves is within the 
124 

overlapping uncertainty limits. For other tests, the agreement is also within the overlapping 

uncertainties. 

Two of the uncooled heat flux gauges, N3 and N4, were in the vicinity of two cooled gauges, 

FG9 and RG10. The uncooled gauges were physically separated from the cooled gauges by 

about 1.1 m (3.6 ft), which is not insignificant near the fire plume in terms of heat flux. Figures 

10-11 and 10-12 compare the heat fluxes to these different types of sensors for Tests 2 and 3. A 

flux adjustment was applied to the uncooled gauges’ output to account for the 75 oC (167 °F) 

temperature of the cooled gauges. The differences are greater than can be accounted for by the 

uncertainties and are probably due to the different environments at the three locations. 

-1

0

1

2

3

4

0 500 1000 1500 2000

N1 Test 2

N3 Test 2

N1 Test 3

N3 Test 3

N1 Test 4

N3 Test 4

N1 Test 5

N3 Test 5

Net Heat Flux (kW/m2

)

Time (s)

Figure 10-5: Net heat flux to the wall near N1 and N3 gauges versus time for Tests 2 to 5. 
125 

0

200

400

600

800

1000

0 500 1000 1500 2000

North Wall

South Wall

East Wall

West Wall

Floor

Ceiling

Total Surfaces

Net Heat Loss Rate (kW)

Time (s)

Figure 10-6: Net heat loss rate to individual surfaces and the total sum for all surfaces 

versus time for Test 3. 

0

200

400

600

800

0 500 1000 1500 2000

North Wall

South Wall

East Wall

West Wall

Floor

Ceiling

Total Surfaces

Net Heat Loss Rate (kW)

Time (s)

Figure 10-7: Net heat loss rate to individual surfaces and the total sum for all surfaces 

versus time for Test 5. 
126 

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

North Wall

South Wall

East Wall

West Wall

Floor

Ceiling

Total Surfaces

Net Heat Loss Rate (kW)

Time (s)

Figure 10-8: Net heat loss rate to individual surfaces and the total sum for all surfaces 

versus time for Test 2. 

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200 1400

North Wall

South Wall

East Wall

West Wall

Floor

Ceiling

Total Surfaces

Time (s)

Figure 10-9: Net heat loss rate to individual surfaces and the total sum for all surfaces 

versus time for Test 4. 
127 

0

200

400

600

800

1000

1200

1400

0 500 1000 1500 2000

Surface Heat Loss

Doorway Heat Loss

Gas Heating 

Sum

Calorimetry

Rate of Enthalpy Change (kW)

Time (s)

Figure 10-10: Rate of heat loss to the enclosure surfaces, through the doorway, and their 

sum as a function of time for Test 3. 

-1

0

1

2

3

4

5

6

7

0 200 400 600 800 1000 1200

Test 2

N3

N4

RG10

Net Heat Flux (kW/m2

)

Time (s)

Figure 10-11: Net heat flux comparison of uncooled gauges N3 and N4 with cooled gauge 

RG10 for Test 2 as a function of time. 
128 

0

2

4

6

8

0 500 1000 1500 2000

Test 3

N3

N4

FG9

RG10

Net Heat Flux (kW/m2

)

Time (s)

Figure 10-12: Net heat flux comparison of uncooled gauges N3 and N4 with cooled gauges 

FG9 and RG10 for Test 3 as a function of time. 

10.6 Contact Resistance Between Marinite Boards 

As part of the evaluation of the heat loss to the enclosure boundaries, a study was conducted of 

the contact resistance between the two 12.7 mm (0.500 in) marinite boards constituting the walls 

and ceiling. One purpose for this study was to determine the feasibility of using a simple 

conduction model for the heat transfer through the walls based on the properties of the marinite 

and measured temperatures on the inside and outside surfaces. A second objective was to 

determine the importance of the contact resistance and how much it varied at different locations. 

Five locations were selected for making these measurements. The locations were near the heat 

loss flux gauges designated N4, S3, W3, C1, and C5. Three of these are wall positions and two 

are ceiling positions. Adjacent to the gauge locations, single layers of doubly thick marinite 

board were inserted in holes cut in the 2-layer surfaces. These square sections were outfitted 

with thermocouples on the inside and outside at the centers. Also, thermocouples were installed 

on the outside of the surfaces in the 2-layer regions near the gauges and opposite the near-gauge 

inside temperature measurement locations. With these measurements, the conduction of heat 

through single and 2-layer boards of the same material and local conditions could be compared. 

The following are the details of the implementation of these measurements. The square marinite 

board inserts were 30 cm (11.8 in) on a side and 25.4 mm (1.00 in) thick. They replaced two 

marinite layers of 12.7 mm (0.500 in) thickness. The outer 12.7 mm (0.500 in) of the perimeter 
129 

of the inserts was milled down to 12.7 mm (0.500 in) thick and the outer board hole was 25.4 

mm (1.00 in) smaller on each side to create an inset arrangement and flange-like attachment area 

for screws which also provided a much tighter seal than an insert cut straight through. The edges 

of the square inserts were located between 10 cm (3.9 in) and 15 cm (5.9 in) from the gauge 

location and also within about 10 cm (3.9 in) of the interior surface temperature measurement. 

Variations in these relative locations were necessary in order to accommodate the structural 

members of the enclosure. The exact x, y, and z positions of all the thermocouples are detailed 

in the list of instrumentation (see Appendix D). The thermocouples were type K and 24 gauge. 

Some thermocouple leads were equipped with high temperature insulation while others were not. 

The insulation characteristics are described elsewhere. Each thermocouple was attached by 

spring-loading it (bending it near the end so the bead pushed against the surface) and also by 

lightly peening the bead (tapping it into the surface). Figures 10-13 and 10-14 are photos of an 

interior and exterior installation, respectively. 

Figure 10-13: Interior installation of marinite thermocouples. 
130 

Figure 10-14: Exterior installation of marinite thermocouples. 

The uncertainty in the thermocouple measurements was approximately 2 °C (4 °F) for the 

manufacturer’s calibration. The uncertainty could also have been affected by radiative heating 

(or cooling), surface contact quality, or other installation and environmental parameters. A 

conservative uncertainty estimate would use the values used in Table 9-1 for gas temperature 

measurements. 

Figure 10-15 shows the peak temperatures inside and outside the two thicknesses of marinite 

board at five measurement stations (N3, S3, W3, C1 and C2) for Tests 2 - 4. The data show that 

significant differences were sometimes present between the interior temperatures. The reasons 

for the differences may be due to loss of contact between thermocouple bead and compartment 

surface. For the C-5 installation, however, the differences were probably due to actual 

temperature gradients associated with non-uniform heating of the ceiling by the fire plume. The 

exterior temperatures in Figure 10-15 were often within 5 oC to 10 o

C (9 °F to 18 °F) which is 

near the uncertainty limits of the measurements and make differentiating the results for the two 

thicknesses unreasonable. This type of data could be used in an analysis of heat loss to the 

compartment surfaces, although the uncertainty in such an analysis would be large due to 

inconsistencies in the data. 
131 

0

50

100

150

200

250

300

350

400

N3-2 N3-4 N3-3 S3-2 S3-4 S3-3 W3-2 W3-4 W3-3 C1-2 C1-4 C1-3 C5-2 C5-4 C5-3

Inside two 1.3 cm layers

Inside one 2.5 cm layer

Outside two 1.3 cm layers

Outside one 2.5 cm layer

Temperature (oC)

Location-Test

Figure 10-15: Peak marinite temperatures for Tests 2 to 4 at five measurement stations 

(N3, S3, W3, C1 and C2). 

133 

11 COMPARTMENT PRESSURE

The compartment pressure was determined by measuring the differential pressure between the 

compartment and the ambient environment outside the experimental enclosure. A temperature 

controlled pressure transducer with a maximum reading of 1.3 kPa (27.2 lb/ft2

) was used to 

measure the differential. The high side of the pressure transducer was located inside the 

compartment, 10.85 m (35.59 ft) from the west wall, 0.1 m (0.3 ft) from the south wall and 0.1 m 

(0.3 ft) from the floor, as specified in the test plan. The low pressure side was open to the high 

bay area. A positive value indicated that the pressure was higher inside the compartment than in 

the high bay area. The pressure transducer was calibrated using a Microtector Hook Gage type 

manometer. The combined expanded uncertainty in the compartment pressure measurements 

was ± 40 Pa (± 0.8 lb/ft2

). 

Table 11-1 summarizes the maximum increase in compartment pressure measured during tests 

when the door was closed. As expected, the smaller fires resulted in smaller pressure differentials 

than the larger fires. Operating the ventilation system also resulted in lower pressure 

differentials compared to similar tests in which the ventilation was off. For tests in which the 

doorway was open (Tests 3, 5, 9, 14, 15, 18; see Table 2-1), the measured pressure difference 

between the compartment and the test area was much smaller than the uncertainty. 

Figure 11-1 shows the increase in compartment pressure for Test 2, in which the door was closed 

and the ventilation system was not operating. These results are typical for all of the tests 

conducted with the compartment door closed and no ventilation operating. The pressure inside 

the compartment increased immediately after the fire began, reaching a peak pressure differential 

within the first 200 s of the test and then decreased over the test duration. Figure 11-2 shows a 

plot of the pressure differential for Test 4, which represents typical results for tests when the 

doors were closed but the ventilation system was operating. For tests when the ventilation 

system was on, the pressure inside the compartment increased for approximately 100 s before the 

pressure difference returned to near-zero. 
134 

Table 11-1: Maximum compartment pressure differential. 

Test 

Number 

Nominal Heat 

Release (MW) Ventilation System Maximum Compartment 

Pressure Increase (Pa)* 

1 0.35 Off 59.7 

2 1 Off 293.1 

4 1 On 59.8 

7 0.35 Off 48.3 

8 1 Off 193.7 

10 1 On 51.7 

13 2 Off 243.2 

16 2 On 114.8 

17 1 Off 198.1 

*combined expanded uncertainty = ± 40 Pa 

0

50

100

150

200

250

300

0 200 400 600 800 1000

Test 2

Compartment Pressure Differential (Pa)

Time (s)

Figure 11-1: Differential pressure for Test 2, door closed, ventilation off. 
135 

-20

-10

0

10

20

30

40

50

60

0 200 400 600 800 1000 1200

Test 4

Compartment Pressure Differential (Pa)

Time (s)

Figure 11-2: Differential pressure for Test 4, door closed, ventilation on. 

137 

12 SUMMARY AND CONCLUSIONS 

Measurements from the ICFMP Benchmarking and Validation Exercise #3 are reported here. 

The accompanying electronic data set includes thirty files, two corresponding to each of the 

fifteen tests, in tab-delimited column format. The finite instrument response time for several of 

the measurements such as the HRR, gas concentrations, and the thermocouple temperature 

measurements, suggest that computer simulations should be time-averaged for appropriate 

comparison with the data. 

139 

13 ACKNOWLEDGEMENTS 

This work was co-sponsored by the Nuclear Regulatory Commission (NRC) and NIST. The 

authors are indebted to Rodney Bryant, Edward Hnetovsky, Jack Lee, Mike Selepak, Roy 

McLane, Jay McElroy, Tyler Smith, and Lauren DeLauter for providing excellent technical 

assistance and to Kevin McGrattan and Tom Ohlemiller for helpful discussions regarding 

experimental design. The authors are grateful to Jason Dreisbach (NRC Project Manager) and 

Monideep Dey (NRC) for providing support for many aspects of the experiments and this report. 

141 

14 REFERENCES 

Bryant, R., Ohlemiller, T., Johnsson, E., Hamins, A., Grove, B., Guthrie, W.F., Maringhides, A., 

and Mulholland, G., The NIST 3 Megawatt Quantitative Heat Release Rate Facility, NIST 

Special Publication 1007, National Institute of Standards and Technology. Gaithersburg, MD, 

December 2003. 

Bryant, R., Ohlemiller, T., Johnsson, E., Hamins, A., Grove, B., Guthrie, W.F., Maringhides, A., 

and Mulholland, G., The NIST 3 Megawatt Quantitative Heat Release Rate Facility – 

Description and Procedures, NISTIR 7052, National Institute of Standards and Technology. 

Gaithersburg, MD, September 2004. 

Bundy, M., NIST, Gaithersburg, MD, personal communication, 2005. (see Appendix J). 

DIPPR Data Compilation of Pure Compound Properties Database, V9.02, NIST Standard 

Reference Data Program #11, National Institute of Standards and Technology, Gaithersburg MD 

(1995). 

Hamins, A., Donnelly, M., Johnsson, R., and Maranghides, A., Report of Test on Experimental 

Conditions and Preliminary Results for the International Fire Model Benchmarking and 

Validation Exercise #3, NISTIR 7024, June 2003a (see http://techconf.llnl.gov). 

Hamins, A., Maranghides, M., and Mulholland, G., “The Global Combustion Behavior of 1 MW 

to 3 MW Hydrocarbon Spray Fires Burning in an Open Environment,” National Institute of 

Standards and Technology (NIST) Internal report NISTIR 7013, Gaithersburg, MD, June 2003b. 

Hamins, A., Johnsson, E., and Donnelly, M., “Energy Balance in a Large Compartment Fire,” 

Proceedings of the 8th International Association of Fire Safety Science (IAFSS), p. 1616, 2005. 

Hanssen, L., Report of Optical Test Data, Optical Physics Division, National Institute of 

Standards and Technology (NIST), Gaithersburg, MD, March 2003. 

Harris, R., personal communication, Fire Research Division, National Institute of Standards and 

Technology (NIST), 2003. (see Appendix J). 

He, Y. P.; Fernando, A.; Luo, M. C., “Determination of interface height from measured 

parameter profile in enclosure fire experiment,” Fire Safety Journal, 31:19-38, 1998. 

Huggett, C., “Estimation of the Rate of Heat Release by Means of Oxygen Consumption,” 

Journal of Fire and Materials, 12: 61-65 (1980). 
142 

Janssens M., Parker, W. J., “Oxygen Consumption Calorimetry,” Heat Release in Fires, 

Babrauskas, V., Grayson, S. J., Eds. (E & FN Spon.), pp. 31-59 London, UK, 1995. 

McCaffrey, B.J., and Heskestad, G., “Robust Bidirectional Low-Velocity Probe for Flame and 

Fire Application,” Combust. Flame, 26:125-127 (1976). 

McGrattan, K. B.; Baum, H. R.; Rehm, R. G.; Hamins, A.; Forney, G. P.; Floyd, J. E.; Hostikka, 

S., “Fire Dynamics Simulator (Version 2) – Technical Reference Guide,” National Institute of 

Standards and Technology (NIST) Internal Report NISTIR 6783, Gaithersburg MD, 44 p. 

November 2001. 

Mulholland, G. and Croarkin, C., “Specific Extinction Coefficient of Flame Generated Smoke,” 

Fire and Materials, 24: 227-230, (2000). 

OMEGA, “Comparison of Time Constant vs. Overall Outside Diameter of Bare Thermocouple 

Wires or Grounded Junction Thermocouples in Air,” The Temperature Handbook, Technical 

Reference Section, p.Z51, Omega Engineering, Inc., Stamford, CT, 2000. 

Parker, W. J., “Calculations of the Heat Release Rate by Oxygen Consumption for Various 

Applications,” National Bureau of Standards (NBS) Internal report NBSIR 81-2427, 1982. 

Pitts, W. M., Braun, E., Peacock, R. D., Mitler, H. E., Johnsson, E. L., Reneke, P. A., and 

Blevins, L. G., “Temperature Uncertainties for Bare-Bead and Aspirated Thermocouple 

Measurements in Fire Environments,” National Institute of Standards and Technology (NIST) 

Internal Report NISTIR 6242, October 1998. 

Reid R.C., Prausnitz, J.M., and Poling, B.E., The Properties of Gases and Liquids, 4th Ed. 

McGraw Hill, NY, 1987. 

SFPE Handbook of Fire Protection Engineering, NFPA. Quincy MA, 1990. 

Taylor, B.N., and Kuyatt, C.E., “Guidelines for Evaluating and Expressing the Uncertainty of 

NIST Measurement Results,” National Institute of Standards and Technology (NIST) Technical 

Note 1297, 1994. 

Taylor, R.E., Groot, H., and Ferrier, J., “Thermophysical Properties of PVC, PE and Marinite,” 

Report TPRL 2958, April 2003. 
143 

A MATERIAL AND OPTICAL PROPERTIES 

A.1 Cables 

Four types of cables were used during the experiments. They are schematically shown in Figures 

A-1 – A-4 and described in Table A-1. The cables all have a complex morphology and are 

composed of an outer insulating jacket and three or seven copper conductors each with an 

insulating jacket. The jacket and insulation material used in the cables are listed in the table. 

The power cables have a ground wire and many tiny 0.180 mm (0.007 in) (outer diameter) bristle 

in portions of the interstitial space that provide additional insulation. To better understand heat 

transfer within the electrical cables, it is beneficial to have knowledge of their optical and 

thermal properties. 

Table A-1: Cable types used in the experiments. 

# Conductors AWG Insulator Jacket Ground Nominal O.D. (cm) 

1 7 14 XPE1 Hypalon2 N 1 

2 7 14 PVC/Nylon PVC N 1.3 ± 0.1 

3 3 6 XPE Hypalon Y 1.9± 0.1 

4 3 6 PVC/Nylon PVC Y 1.6 ± 0.1 

1 flame retarded crosslinked polyethylene 

2 Hypalon is a registered DuPont trademark for chlorosulfinated polyethylene (CSPE)

Thermal property information including the specific heat (cp), the thermal diffusivity (α), the 

thermal conductivity (K) of the PVC and XPE cable insulation has been determined using ASTM 

E1269 and ASTM E14611

 and are reproduced in Table A-2 below. 

Table A-2 also includes the spectrally integrated value of the emissivity (ε). This was determined 

by normalizing the measured ambient temperature spectral hemispherical reflectance (from 

1.5 μm to 19.2 μm)2

 with the Planck function at 1200 K (1700 °F). The spectral range of the data 

(1.5 μm to 19.2 μm) covers a major fraction (~95 %) of the intensity of a 1200 K (1700 °F) 

blackbody. 

 

1

 Taylor, R.E., Groot, H., and Ferrier, J., “Thermophysical Properties of PVC, PE and Marinite,” Report TPRL 

2958, April 2003. 

2

 Hanssen, L., Report of Optical Test Data, Optical Physics Division, National Institute of Standards and 

Technology (NIST), Gaithersburg, MD, March 2003. 
144 

 plastic liner Copper conductors (0.18 cm O.D.) 

outer Hypalon jacket 

1.00 cm O.D. 

 0.06 cm inner XPE jacket 

Figure A-1: Morphology of a 7/C #14 AWG Control Cable (Cable 1 in Table A-1). 

0.14 cm ±0.01 outer jacket 

 Copper conductors (0.18 cm O.D.) 

outer PVC jacket 

1.27 cm O.D. 

0.1 cm inner PVC/Nylon jacket 

Figure A-2: Morphology of a 7/C #14 AWG Control Cable (Cable 2 in Table A-1). 

Copper ground (0.4 cm O.D.) 

 Copper conductors (0.51 cm O.D.) 

outer Hypalon jacket 

1.91 cm O.D. 

0.25 cm outer jacket 0.13 cm inner XLP jacket 

~50 x 0.018 cm diameter tubes 

0.005 cm diameter 

clear plastic liner 

Figure A-3: Morphology of a 3/C #6 AWG Power Cable (Cable 3 in Table A-1). 

Copper ground (0.4 cm O.D.) 

 Copper conductors (0.51 cm O.D.) 

outer PVC jacket 

1.65 cm O.D. 

0.14 cm outer jacket 0.095 cm inner PVC jacket 

~50 x 0.018 cm diameter tubes 

Figure A-4: Morphology of a 3/C #6 AWG Power Cable (Cable 4 in Table A-1). 
145 

Table A-2: Material and optical properties of the cable materials. 

Properties of the PVC Cable 

T (oC) k (W/m K) * α (m2

/s) * cp(J/kg K) * ε ** 

23 0.192 1.08 x 10 -7 1289 0.95±0.01

50 0.175 9.4 x 10 –8 1353 - 

75 0.172 8.9 x 10 –8 1407 - 

100 0.147 7.3 x 10 –8 1469 - 

125 0.141 6.7 x 10 –8 1530 - 

150 0.134 6.2 x 10 –8 1586 - 

Properties of the XLP Cable 

T (oC) k (W/m K) * α (m2

/s) * cp(J/kg K) * ε ** 

23 0.235 1.23 x 10 -7 1390 0.95±0.01

50 0.232 1.14 x 10 –7 1476 - 

75 0.223 1.06 x 10 –7 1526 - 

100 0.210 9.8 x 10 –8 1560 - 

125 0.190 8.7 x 10 –8 1585 - 

150 0.192 8.7 x 10 –8 1607 - 

* Taylor, R.E., Groot, H., and Ferrier, J., Thermophysical Properties of PVC, PE and Marinite, Report 

TPRL 2958, April 2003. 

** Hanssen, L., Report of Optical Test Data, Optical Physics Division, National Institute of Standards and 

Technology (NIST), Gaithersburg, MD, March 2003.

A.2 Properties of Gypsum, Marinite, Feraloy, and the PVC Slab E 

The compartment walls and ceiling were composed of marinite, while the floor was composed of 

gypsum. The temperature dependent thermal properties and room temperature emissivity of 

marinite are listed in Table A-3, which were determined in the same manner as the marinite (see 

Table A-2). The properties of gypsum are listed in Table A-4. The PVC slab (Target E) was 

made of PVC insulation from the cables used in this study. 
146 

Table A-3: Material and optical properties of marinite. 

T (oC) K (W/m K) * α (m2

/s) * cp(J/kg K) * ε ** 

23 0.111 2.13 x 10 –7 778 0.74±0.04

50 0.114 2.15 x 10 –7 795 - 

100 0.126 2.17x 10 –7 871 - 

200 0.140 2.17 x 10 –7 965 - 

300 0.153 2.18 x 10 –7 1047 - 

400 0.160 2.21 x 10 –7 1082 - 

500 0.175 2.26x 10 –7 1160 - 

600 0.190 2.36x 10 –7 1205 - 

650 0.198 2.42 x 10 –7 1223 - 

* Taylor, R.E., Groot, H., and Ferrier, J., Thermophysical Properties of PVC, PE and Marinite, Report TPRL 

2958, April 2003. 

** Hanssen, L., Report of Optical Test Data, Optical Physics Division, National Institute of Standards and 

Technology (NIST), Gaithersburg, MD, March 2003.

Table A-4: Material and optical properties of feraloy and gypsum at ambient temperature. 

Material K (W/m K) ρ (kg/m3

) cp (J/kg K) ε

Feraloy* 78.2** 787** 456** - 

Gypsum 0.16 + 790 + 900 + 0.9 +

* assumed to have properties similar to iron. 

** Smithells Metals Reference Book. 7th Ed., Ed.: E.A. Brandes and G.B. Brook, Elsevier, 1998. 

+ from the CFAST database ( http://fast.nist.gov/ )
147 

B AMBIENT HUMIDITY AND TEMPERATURE 

DURING TESTING 

A Dickson Humidity and Temperature Logger was used to monitor ambient conditions before 

each test. Manufacturer instrument accuracy is listed at ± 5 % for humidity and ± 1 o

C (2 °F) in 

temperature. The measurement results and expanded uncertainty are listed in Table B-1. 

Table B-1: Humidity and temperature in the large fire lab. 

Test Number Ambient Temperature 

(

o

C) 

Humidity (%) 

Test 1 22± 2 32± 10 

Test 2 26± 2 36 ± 10 

Test 3 30± 2 34 ± 10 

Test 4 27± 2 44 ± 10 

Test 5 28± 2 37± 10 

Test 7 24± 2 58± 10 

Test 8 25± 2 63± 10 

Test 9 27± 2 62± 10 

Test 10 27± 2 63± 10 

Test 13 31± 2 52± 10 

Test 14 28± 2 61± 10 

Test 15 18± 2 95± 10 

Test 16 26± 2 55± 10 

Test 17 29± 2 45± 10 

Test 18 27± 2 40± 10 

149 

C COMPARTMENT LEAKAGE AREA

Experimental enclosure leakage measurements were performed before tests 1, 2, 8 and 13. The 

measurements were conducted with a commercial INFLILTEC Model E-3 Blower Door and E3-

DM4 Digital Blower Door Micro-manometer. Sensor accuracy is rated as 1% of the measured 

pressure according to manufacturer literature and 3 % for the leakage area. The measurements 

consist of using a fan to blow a controlled amount of air into a compartment and then measure 

the pressure rise. Two fan flow rates are used, creating two values of pressure differential. The 

digital manometer measures the pressure and the Equivalent Leakage Area (Aequiv) at that 

pressure is then calculated. 

The experimental procedure involved closing both the compartment door and the supply and 

exhaust ventilation ducts. The instrument frame was installed in an access opening and the fan 

was installed in the frame. A digital manometer was connected. Readings were taken for low and 

high fan flow, which were then converted to an Equivalent Leakage Area (Aequiv). The 

calculation involved solving for the two unknowns, the flow coefficient (C) and the flow 

exponent (n) of the enclosure, from the two measurements and the following equations. 

Equation C.1 is used to solve for the constant n, then Eq. C.2 is used to solve for the constant C 

using the existing experimental Pressure and leakage area values. The equivalent leakage area at 

50 Pa (1 lb/ft2

) is then determined from Eq. C.2. 

Aequiv1 / Aequiv2= (Comp. Pressure1/Comp. Pressure 2) n-0.5 (C.1) 

Aequiv = C (Comp. Pressure) n-0.5 (C.2) 

The measured values for Pressure and Leakage area are listed in Table C-1. Table C-2 lists the C 

and n coefficient values together with the calculated leakage area at 50 Pa (1 lb/ft2

). 

Table C-1: Measured Pressure, Leakage Area and Air Flow Rate. 

Test Pressure (Pa) Leakage Area (cm2) 

1 118 710 

1 210 800 

2 115 710 

2 215 826 

7 107 903 

7 192 1084 

8 114 832 

8 200 987 

13 105 1019 

13 180 1110 
150 

Table C-2: Calculated C and n values, compartment equivalent leakage area and expanded 

uncertainty. 

Test Flow Coefficient 

C

Flow Exponent 

n

Equivalent Leakage 

Area (cm2

) at 50 Pa 

1 263 0.708 593±42 

2 223 0.744 580 ±40 

7 210 0.812 712±50 

8 198 0.803 645±45 

13 450 0.657 833±58 
151 

D FORMAT OF THE ELECTRONIC DATA 

The experimental data associated with this report is available on a companion set of thirty 

electronic files; fifteen pairs of files, two for each test. The data are electronically stored in tabdelimited column format. The columns are labeled and ordered in an identical manner for each 

of the fifteen tests. Some columns in the files do not contain data, either because no measurement 

was made or because an instrument was inoperative. In the case when a measurement was not 

made (e.g., when the ventilation and doorway were closed, no calorimetry HRR measurement 

was made), the symbol “999” is inserted in the first row of that column and the rest of the rows 

are filled with zeros. If an instrument was not functioning, the data columns are treated in the 

same manner. Non-functioning instruments are specified in Appendix E of this report. 

The data were acquired at 1 Hz, but are presented as average values over a 10 s interval. Time 

averaging the data over too long a period may cause loss of legitimate information. The selection 

of a 10 s averaging period is a compromise, but appears to be reasonable considering the time 

response of the instrumentation used in the experiments. 

Tables D-1 and D-2 describe the electronic data files including a list of the assigned designation 

for each measurement and its units. The tables also contain information on the measurement 

type, the (X,Y,Z) location of the instrument, and expressly for NIST use - the original NIST data 

channel number. For each data file, the data begins at time equal to –60 s, which provides 1 min 

of representative background information. 

The format of the first file for each test is shown in Table D-1 and is comprised of 247 columns 

of data. The first column of data represents time during the experiment. Time equal to zero is the 

time of the initiation of the fuel spray. Ignition of the fuel spray typically occurred within a few 

seconds of time zero, but on occasion took as long as 10 s to 15 s. Columns 2 to 37 represent 

uncooled heat flux gauges mounted on the surface of the compartment. Columns 38 to 107 

represent thermocouples mounted on the seven thermocouple trees. Columns 108 to 179 

represent wall-mounted thermocouples, either mounted directly on the compartment surface 

(designated by a “-2” suffix) or on the surface-mounted flux gauges (designated by a “-1” 

suffix). Columns 180 to 217 represent thermocouples mounted on and inside the insulation of the 

horizontal and vertical cables and the junction box. Columns 218 and 219 represent 

thermocouples near the ceiling mounted aspirated thermocouples. Columns 220 to 224 represent 

gas and smoke concentration measurements. Columns 225 to 234 represent heat flux to the target 

gauges near the cables. Column 235 is the differential pressure between inside of the 

compartment and the environment outside of the compartment. Columns 236 to 254 represent 

calculated information on the energy and average mass flow through the compartment openings 

including the supply and exhaust ventilation ducts and the door, heat losses to the walls, the HRR 

measured by calorimetry, and the layer height and average temperatures. 
152 

The format of the second file for each test is shown in Table D-2. The format follows the first 

file. The first column of data represents time during the experiment. Columns 2 to 28 represent 

the thermocouples associated with the bi-directional probes in the doorway, the supply duct and 

on the ceiling. Columns 29 to 32 represent thermocouples in the exhaust. Columns 33 to 36 

represent aspirated thermocouples in the doorway and the exhaust. Columns 37 to 51 represent 

the thermocouples mounted near, on, and outside the 2.5 cm (1.0 in) thick, 30.0 cm x 30.0 cm 

(11.8 in x 11.8 in) wide single layer of marinite installed at five locations throughout the 

compartment. Columns 52 to 69 represent velocities measured by bi-directional probes in the 

doorway, supply, exhaust and near the ceiling. 

Table D-1: Instrument channel number, designation, units, and location. 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

Time 1 1 s - - - 

Wall Flux Gauges1

N,S,E,W,Ceiling,Floor kW/m2

 

North U-1 2 32 kW/m2

 3.91 7.04 1.49 

North U-2 3 33 kW/m2

 3.91 7. 04 3.82 

North U-3 4 34 kW/m2

 9.73 7. 04 1.86 

North U-4 5 35 kW/m2

 12.15 7. 04 1.86 

North U-5 6 36 kW/m2

 17.79 7. 04 1.49 

North U-6 7 37 kW/m2

 17.79 7. 04 3.82 

South U-1 8 38 kW/m2

 3.91 0 1.49 

South U-2 9 39 kW/m2

 3.91 0 3.82 

South U-3 10 40 kW/m2

 9.73 0 1.86 

South U-4 11 41 kW/m2

 12.15 0 1.86 

South U-5 12 42 kW/m2

 17.79 0 1.49 

South U-6 13 43 kW/m2

 17.79 0 3.82 

East U-1 14 337 kW/m2

 21.66 1.59 1.12 

East U-2 15 338 kW/m2

 21.66 1.59 2.42 

East U-3 16 339 kW/m2

 21.66 5.76 1.12 

East U-4 17 340 kW/m2

 21.66 5.76 2.42 

West U-1 18 44 kW/m2

 0 1.59 1.12 

West U-2 19 45 kW/m2

 0 1.59 2.42 

West U-3 20 46 kW/m2

 0 5.76 1.12 

West U-4 21 47 kW/m2

 0 5.76 2.42 

Floor U-1 22 48 kW/m2

 3.04 3.58 0 

Floor U-2 23 49 kW/m2

 9.11 1.99 0 

Floor U-3 24 50 kW/m2

 9.11 5.96 0 

Floor U-4 25 51 kW/m2

 10.85 2.38 0 

Floor U-5 26 52 kW/m2

 10.85 5.16 0.10 

Floor U-6 27 53 kW/m2

 13.02 1.99 0 

Floor U-7 28 54 kW/m2

 13.02 5.96 0 

Floor U-8 29 55 kW/m2

 18.66 3.58 0 

Ceiling U-1 30 56 kW/m2

 3.04 3.58 3.82 
153 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

Ceiling C-2 31 57 kW/m2

 9.11 1.99 3.82 

Ceiling C-3 32 58 kW/m2

 9.11 5.96 3.82 

Ceiling C-4 33 59 kW/m2

 10.85 2.38 3.82 

Ceiling C-5 34 60 kW/m2

 10.85 5.16 3.82 

Ceiling C-6 35 61 kW/m2

 13.02 1.99 3.82 

Ceiling C-7 36 62 kW/m2

 13.02 5.96 3.82 

Ceiling U-8 37 63 kW/m2

 18.66 3.58 3.82 

Thermocouple Trees o

C 

Tree 1-1 38 64 o

C 5 3.58 0.35 

Tree 1-2 39 65 o

C 5 3.58 0.70 

Tree 1-3 40 66 o

C 5 3.58 1.05 

Tree 1-4 41 67 o

C 5 3.58 1.40 

Tree 1-5 42 68 o

C 5 3.58 1.75 

Tree 1-6 43 69 o

C 5 3.58 2.10 

Tree 1-7 44 70 o

C 5 3.58 2.45 

Tree 1-8 45 71 o

C 5 3.58 2.80 

Tree 1-9 46 72 o

C 5 3.58 3.15 

Tree 1-10 47 73 o

C 5 3.58 3.50 

Tree 2-1 48 74 o

C 10.85 6.85 0.35 

Tree 2-2 49 75 o

C 10.85 6.85 0.70 

Tree 2-3 50 76 o

C 10.85 6.85 1.05 

Tree 2-4 51 77 o

C 10.85 6.85 1.40 

Tree 2-5 52 78 o

C 10.85 6.85 1.75 

Tree 2-6 53 79 o

C 10.85 6.85 2.10 

Tree 2-7 54 80 o

C 10.85 6.85 2.45 

Tree 2-8 55 81 o

C 10.85 6.85 2.80 

Tree 2-9 56 82 o

C 10.85 6.85 3.15 

Tree 2-10 57 83 o

C 10.85 6.85 3.50 

Tree 3-1 58 84 o

C 10.85 2.2 0.35 

Tree 3-2 59 85 o

C 10.85 2.2 0.70 

Tree 3-3 60 86 o

C 10.85 2.2 1.05 

Tree 3-4 61 87 o

C 10.85 2.2 1.40 

Tree 3-5 62 88 o

C 10.85 2.2 1.75 

Tree 3-6 63 89 o

C 10.85 2.2 2.10 

Tree 3-7 64 90 o

C 10.85 2.2 2.45 

Tree 3-8 65 91 o

C 10.85 2.2 2.80 

Tree 3-9 66 92 o

C 10.85 2.2 3.15 

Tree 3-10 67 93 o

C 10.85 2.2 3.50 

Tree 4-1 68 94 o

C 10.85 1.35 0.35 

Tree 4-2 69 95 o

C 10.85 1.35 0.70 

Tree 4-3 70 96 o

C 10.85 1.35 1.05 

Tree 4-4 71 97 o

C 10.85 1.35 1.40 

Tree 4-5 72 98 o

C 10.85 1.35 1.75 
154 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

Tree 4-6 73 99 o

C 10.85 1.35 2.10 

Tree 4-7 74 100 o

C 10.85 1.35 2.45 

Tree 4-8 75 101 o

C 10.85 1.35 2.80 

Tree 4-9 76 102 o

C 10.85 1.35 3.15 

Tree 4-10 77 103 o

C 10.85 1.35 3.50 

Tree 5-1 78 104 o

C 10.85 0.55 0.35 

Tree 5-2 79 105 o

C 10.85 0.55 0.70 

Tree 5-3 80 106 o

C 10.85 0.55 1.05 

Tree 5-4 81 107 o

C 10.85 0.55 1.40 

Tree 5-5 82 108 o

C 10.85 0.55 1.75 

Tree 5-6 83 109 o

C 10.85 0.55 2.10 

Tree 5-7 84 110 o

C 10.85 0.55 2.45 

Tree 5-8 85 111 o

C 10.85 0.55 2.80 

Tree 5-9 86 112 o

C 10.85 0.55 3.15 

Tree 5-10 87 113 o

C 10.85 0.55 3.50 

Tree 6-1 88 114 o

C 11.95 3.58 0.35 

Tree 6-2 89 115 o

C 11.95 3.58 0.70 

Tree 6-3 90 116 o

C 11.95 3.58 1.05 

Tree 6-4 91 117 o

C 11.95 3.58 1.40 

Tree 6-5 92 118 o

C 11.95 3.58 1.75 

Tree 6-6 93 119 o

C 11.95 3.58 2.10 

Tree 6-7 94 120 o

C 11.95 3.58 2.45 

Tree 6-8 95 121 o

C 11.95 3.58 2.80 

Tree 6-9 96 122 o

C 11.95 3.58 3.15 

Tree 6-10 97 123 o

C 11.95 3.58 3.50 

Tree 7-1 98 124 o

C 16.7 3.58 0.35 

Tree 7-2 99 125 o

C 16.7 3.58 0.70 

Tree 7-3 100 126 o

C 16.7 3.58 1.05 

Tree 7-4 101 127 o

C 16.7 3.58 1.40 

Tree 7-5 102 128 o

C 16.7 3.58 1.75 

Tree 7-6 103 129 o

C 16.7 3.58 2.10 

Tree 7-7 104 130 o

C 16.7 3.58 2.45 

Tree 7-8 105 131 o

C 16.7 3.58 2.80 

Tree 7-9 106 132 o

C 16.7 3.58 3.15 

Tree 7-10 107 133 o

C 16.7 3.58 3.50 

Wall TCs2 

 

TC North U-1-1 108 134 o

C 3.91 7.04 1.49 

TC North U-1-2 109 135 o

C 3.85 7.04 1.49 

TC North U-2-1 110 136 o

C 3.91 7.04 3.70 

TC North U-2-2 111 137 o

C 3.86 7.04 3.71 

TC North U-3-1 112 138 o

C 9.52 7.04 1.86 

TC North U-3-2 113 139 o

C 9.48 7.04 1.86 

TC North U-4-1 114 140 o

C 12.11 7.04 1.86 
155 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

TC North U-4-2 115 141 o

C 12.07 7.04 1.88 

TC North U-5-1 116 142 o

C 17.74 7.04 1.49 

TC North U-5-2 117 143 o

C 17.69 7.04 1.49 

TC North U-6-1 118 144 o

C 17.74 7.04 3.69 

TC North U-6-2 119 145 o

C 17.69 7.04 3.69 

TC South U-1-1 120 146 o

C 3.92 0 1.49 

TC South U-1-2 121 147 o

C 3.86 0 1.49 

TC South U-2-1 122 148 o

C 3.92 0 3.82 

TC South U-2-2 123 149 o

C 3.86 0 3.82 

TC South U-3-1 124 150 o

C 9.54 0 1.86 

TC South U-3-2 125 151 o

C 9.54 0 1.86 

TC South U-4-1 126 152 o

C 12.12 0 1.86 

TC South U-4-2 127 154 o

C 12.08 0 1.86 

TC South U-5-1 128 155 o

C 17.74 0 1.50 

TC South U-5-2 129 156 o

C 17.69 0 1.50 

TC South U-6-1 130 157 o

C 17.74 0 3.70 

TC South U-6-2 131 158 o

C 17.74 0 3.70 

TC East U-1-1 132 159 o

C 21.66 1.60 1.12 

TC East U-1-2 133 160 o

C 21.66 1.52 1.12 

TC East U-2-1 134 161 o

C 21.66 1.60 2.42 

TC East U-2-2 135 162 o

C 21.66 1.52 2.40 

TC East U-3-1 136 163 o

C 21.66 5.75 1.12 

TC East U-3-2 137 164 o

C 21.66 5.68 1.13 

TC East U-4-1 138 165 o

C 21.66 5.76 2.42 

TC East U-4-2 139 166 o

C 21.66 5.70 2.42 

TC West U-1-1 140 167 o

C 0 1.59 1.12 

TC West U-1-2 141 168 o

C 0 1.59 1.12 

TC West U-2-1 142 169 o

C 0 1.59 2.42 

TC West U-2-2 143 170 o

C 0 1.59 2.42 

TC West U-3-1 144 171 o

C 0 5.76 1.12 

TC West U-3-2 145 172 o

C 0 5.70 1.12 

TC West U-4-1 146 173 o

C 0 5.77 2.42 

TC West U-4-2 147 174 o

C 0 5.70 2.42 

TC Floor U-1-1 148 175 o

C 3.03 3.51 0 

TC Floor U-1-2 149 176 o

C 3.08 3.51 0 

TC Floor U-2-1 150 177 o

C 9.08 1.99 0 

TC Floor U-2-2 151 178 o

C 9.08 1.94 0 

TC Floor U-3-1 152 179 o

C 9.11 5.97 0 

TC Floor U-3-2 153 180 o

C 9.06 5.97 0 

TC Floor U-4-1 154 181 o

C 10.84 2.39 0 

TC Floor U-4-2 155 182 o

C 10.86 2.38 0 

TC Floor C-5-1 156 183 o

C 10.89 5.20 0.01 

TC Floor C-5-2 157 184 o

C 10.93 5.20 0.01 

TC Floor U-6-1 158 185 o

C 13.09 1.99 0 
156 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

TC Floor U-6-2 159 186 o

C 13.13 1.99 0 

TC Floor U-7-1 160 187 o

C 13.00 5.98 0 

TC Floor U-7-2 161 188 o

C 13.00 5.92 0 

TC Floor U-8-1 162 189 o

C 18.63 3.58 0 

TC Floor U-8-2 163 190 o

C 18.63 3.54 0 

TC Ceiling U-1-1 164 191 o

C 3.06 3.52 3.82 

TC Ceiling U-1-2 165 192 o

C 3.04 3.60 3.82 

TC Ceiling C-2-1 166 193 o

C 9.03 1.97 3.82 

TC Ceiling C-2-2 167 194 o

C 8.99 2.00 3.82 

TC Ceiling C-3-1 168 195 o

C 9.08 5.96 3.82 

TC Ceiling C-3-2 169 196 o

C 9.03 5.97 3.82 

TC Ceiling C-4-1 170 197 o

C 10.85 2.38 3.82 

TC Ceiling C-4-2 171 198 o

C 10.79 2.38 3.82 

TC Ceiling C-5-1 172 199 o

C 10.83 5.20 3.82 

TC Ceiling C-5-2 173 200 o

C 10.79 5.20 3.82 

TC Ceiling C-6-1 174 201 o

C 13.01 2.03 3.82 

TC Ceiling C-6-2 175 202 o

C 13.00 2.07 3.82 

TC Ceiling C-7-1 176 203 o

C 12.88 6.00 3.82 

TC Ceiling C-7-2 177 204 o

C 12.84 5.98 3.82 

TC Ceiling U-8-1 178 205 o

C 18.63 3.58 3.82 

TC Ceiling U-8-2 179 206 o

C 18.71 3.54 3.82 

Cable TCs 

C - Ts-1 180 233 o

C 6.85 2.25 3.2 

C - Tc-2 181 234 o

C 6.85 2.25 3.2 

D - Ts-3 182 235 o

C 6.85 2.03 3.22 

B - Ts-4 183 236 o

C 6.85 1.4 2.7 

A - Ts-6 184 238 o

C 6.85 0.65 2.19 

F - Ts-7 185 239 o

C 6.85 0.5 2.175 

C - Ts-8 186 240 o

C 8.85 2.25 3.2 

C - Tc-9 187 241 o

C 8.85 2.25 3.2 

C - Ts-10 188 242 o

C 10.83 2.27 3.2 

C - Tc-11 189 243 o

C 10.83 2.27 3.2 

D - Ts-12 190 244 o

C 10.83 2.03 3.22 

D - Tc-13 191 245 o

C 10.83 2.03 3.22 

B - Ts-14 192 246 o

C 10.83 1.4 2.7 

B - Tc-15 193 247 o

C 10.83 1.4 2.7 

E - Ts-16 bottom 194 248 o

C 10.83 1.25 2.7 

E - Tc-17 middle 195 249 o

C 10.83 1.25 2.7 

E - Ts-16’ top 196 237 o

C 10.83 1.25 2.7 

A - Ts-18 197 250 o

C 10.83 0.65 2.19 

A - Tc-19 198 251 o

C 10.83 0.65 2.19 

F - Ts-20 199 252 o

C 10.83 0.5 2.175 

F - Tc-21 200 253 o

C 10.83 0.5 2.175 
157 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

C - Ts-22 201 254 o

C 12.85 2 3.2 

C - Tc-23 202 255 o

C 12.85 2 3.2 

C - Ts-24 203 256 o

C 14.85 2.25 3.2 

C - Tc-25 204 257 o

C 14.85 2.25 3.2 

D - Ts-26 205 258 o

C 14.85 2.03 3.22 

B - Ts-27 206 259 o

C 14.85 1.4 2.7 

A - Ts-29 207 261 o

C 14.85 0.65 2.2 

F - Ts-30 208 262 o

C 14.85 0.5 2.2 

Vertical Cable Ts-31 209 263 o

C 10.79 7.04 0.35 

Vertical Cable Ts-32 210 264 o

C 10.79 7.04 0.7 

Vertical Cable Ts-33 211 265 o

C 10.79 7.04 1.75 

Vertical Cable Tc-34 212 266 o

C 10.79 7.04 1.75 

Vertical Cable Ts-35 213 267 o

C 10.79 7.04 2.45 

Vertical Cable Ts-36 214 268 o

C 10.79 7.04 3.15 

Junction Box TC-37 

(inside) 215 269 o

C 17.66 3.44 3.76 

Junction Box TS-38 

(bottom) 216 270 o

C 17.66 3.44 3.70 

Junction Box TS-39 

(west) 217 271 o

C 17.51 3.44 3.76 

Aspirated TCs 

 

ATC -5 218 276 o

C 10.85 0.55 1.05 

ATC -6 219 277 o

C 10.85 0.55 2.76 

Gas/Smoke Conc. 

O2 –1 220 331 Vol. Frac. 6.85 3.52 3.20 

O2 –2 221 332 Vol. Frac. 6.85 3.52 0.50 

CO –3 222 333 Vol. Frac. 6.85 3.52 3.20 

CO2 –4 223 335 Vol. Frac. 6.85 3.52 3.20 

Smoke Obs./Conc. 224 354 mg/m3

 21.11 0.5 3.57 

Rad and Total Flux 

Gauges 

Cable Rad Gauge 1 225 341 kW/m2

 10.81 0.44 2.05 

Cable Total Flux 2 226 342 kW/m2

 10.87 0.44 2.05 

Cable Rad Gauge 3 227 343 kW/m2

 10.81 1.25 2.52 

Cable Total Flux 4 228 344 kW/m2

 10.87 1.25 2.52 

Cable Rad Gauge 5 229 345 kW/m2

 10.81 1.40 2.54 

Cable Total Flux 6 230 346 kW/m2

 10.87 1.40 2.54 

Cable Rad Gauge 7 231 347 kW/m2

 10.81 2.00 3.04 

Cable Total Flux 8 232 348 kW/m2

 10.87 2.00 3.04 

Cable Total Flux 9 233 349 kW/m2

 10.81 6.85 1.75 

Cable Rad Gauge 10 234 350 kW/m2

 10.87 6.85 1.76 

Comp Pressure 

 

CompP 235 351 Pa 10.85 0.10 0.10 
158 

Location 

Designation Column 

NIST 

data 

Channel Units X (m) Y (m) Z (m) 

Calculated 

Parameters 

Mass Flow In Door 236 - kg/s - - - 

Mass Flow Out Door 237 - kg/s - - - 

Net Mass Flow thru 

Door 238 

- kg/s - - - 

Energy Flow through 

Door 239 

- kW 

- - - 

Mass Flow through 

Supply Vent 240 

- kg/s 

- - - 

Mass Flow through 

Exhaust Vent 241 

- kg/s 

- - - 

Energy Flow through 

Exhaust Vent 242 

- kW 

- - - 

Loss North Wall 243 - kW - - - 

Loss South Wall 244 - kW - - - 

Loss East Wall 245 - kW - - - 

Loss West Wall 246 - kW - - - 

Loss Floor Wall 247 - kW - - - 

Loss Ceiling Wall 248 - kW - - - 

Total Loss All Surfaces 249 - kW - - - 

Sum of Energy Losses 

through 

Door+Vent+Surfaces 250 

- kW 

- - - 

Heat release rate,Q&  251 - kW - - - 

Layer Height (from 

floor) 252 

- m 

- - - 

Lower Layer Temp 253 - o

C - - - 

Upper Layer Temp 254 - o

C - - - 

1. all gauges were uncooled. The letters C and U in the designation have no meaning. Net heat flux to the 

gauge is reported. 

2. “-1” refers to the thermocouple incorporated in the heat flux gauge, and “-2” refers to the thermocouple on 

the interior surface of the compartment.

Table D-2: Instrument channel number, designation, units, and location for the second 

data file. 

Location 

Measurement Column 

NIST data

Channel Units X (m) Y (m) Z (m) 

Time 1 1 s - - - 

Bidirectional Probe TCs1 

 TC Door –1 2 207 o

C 0.10 2.72 0.21 

 TC Door –2 3 208 o

C 0.10 2.72 0.61 

 TC Door –3 4 209 o

C 0.10 2.72 1.01 

 TC Door –4 5 210 o

C 0.10 2.72 1.21 
159 

Location 

Measurement Column 

NIST data

Channel Units X (m) Y (m) Z (m) 

 TC Door –5 6 211 o

C 0.10 2.72 1.41 

 TC Door –6 7 212 o

C 0.10 2.72 1.61 

 TC Door –7 8 213 o

C 0.10 2.72 1.81 

 TC Door –8 9 214 o

C 0.10 2.72 1.91 

 TC Door –9 10 215 o

C 0.10 3.52 0.21 

 TC Door –10 11 216 o

C 0.10 3.52 0.61 

 TC Door –11 12 217 o

C 0.10 3.52 1.01 

 TC Door –12 13 218 o

C 0.10 3.52 1.21 

 TC Door –13 14 219 o

C 0.10 3.52 1.41 

 TC Door –14 15 220 o

C 0.10 3.52 1.61 

 TC Door –15 16 221 o

C 0.10 3.52 1.81 

 TC Door –16 17 222 o

C 0.10 3.52 1.91 

 TC Door –17 18 223 o

C 0.10 4.32 0.21 

 TC Door –18 19 224 o

C 0.10 4.32 0.61 

 TC Door –19 20 225 o

C 0.10 4.32 1.01 

 TC Door –20 21 226 o

C 0.10 4.32 1.21 

 TC Door –21 22 227 o

C 0.10 4.32 1.41 

 TC Door –22 23 228 o

C 0.10 4.32 1.61 

 TC Door –23 24 229 o

C 0.10 4.32 1.81 

 TC Door –24 25 230 o

C 0.10 4.32 1.91 

TC Supply Vent –252 26 231 o

C 11.22 0 2.41 

TC C4 gas Bidir 27 285 o

C 10.85 2.38 3.82 

TC C5 gas Bidir 28 286 o

C 10.85 5.27 3.82 

Exhaust TCs 

Exhaust TC 1 29 367 o

C 11.22 7.04 2.20 

Exhaust TC 2 30 368 o

C 11.22 7.04 2.40 

Exhaust TC 3 31 369 o

C 11.22 7.04 2.55 

Exhaust TC 4 32 370 o

C 11.22 7.04 2.65 

Aspirated TCs 

 

ATC Door -1 33 272 o

C 0.10 3.61 0.20 

ATC Door -2 34 273 o

C 0.10 3.61 1.00 

ATC Door -3 35 274 o

C 0.10 3.61 1.80 

ATC Exhaust –43 36 275 o

C 11.22 7.27 2.39 

Marinite Wall TCs 

TC N U outside 1⁄2 “ 37 278 o

C 9.55 7.07 1.86 

TC N U inside 1” 38 279 o

C 9.95 7.04 1.86 

TC N outside 1” 39 280 o

C 9.95 7.07 1.86 

TC S outside 1⁄2 “ 40 371 o

C 9.48 -0.03 1.86 

TC S inside 1” 41 372 o

C 9.74 0 1.86 

TC S outside 1” 42 373 o

C 9.74 -0.03 1.86 

TC W outside 1⁄2 “ 43 287 o

C -0.03 5.71 2.43 

TC W inside 1” 44 288 o

C 0 6.03 2.43 

TC W outside 1” 45 289 o

C -0.03 6.03 2.43 

TC C outside 1⁄2 “ 46 290 o

C 3.10 3.52 3.85 
160 

Location 

Measurement Column 

NIST data

Channel Units X (m) Y (m) Z (m) 

TC C inside 1” 47 291 o

C 2.80 3.58 3.82 

TC C outside 1” 48 292 o

C 2.80 3.58 3.85 

TC C outside 1⁄2 “ 49 293 o

C 11.45 5.18 3.85 

TC C outside 1” 50 294 o

C 11.66 5.24 3.85 

TC C inside 1” 51 295 o

C 11.66 5.24 3.82 

Bidirectional Probes 

BP Door -1 52 305 m/s 0.10 2.72 0.19 

BP Door -2 53 307 m/s 0.10 2.72 0.58 

BP Door –3 54 308 m/s 0.10 2.72 1.00 

BP Door –4 55 309 m/s 0.10 2.72 1.39 

BP Door –5 56 310 m/s 0.10 2.72 1.80 

BP Door –6 57 311 m/s 0.10 3.52 0.19 

BP Door –7 58 312 m/s 0.10 3.52 0.6 

BP Door –8 59 313 m/s 0.10 3.52 1.00 

BP Door –9 60 314 m/s 0.10 3.52 1.4 

BP Door –10 61 315 m/s 0.10 3.52 1.79 

BP Door –11 62 316 m/s 0.10 4.32 0.21 

BP Door –13 63 318 m/s 0.10 4.32 1.01 

BP Door –14 64 319 m/s 0.10 4.32 1.40 

BP Door –15 65 320 m/s 0.10 4.32 1.80 

BP Supply Vent -16 66 321 m/s 11.22 0 2.354 

BP Exhaust Vent 67 317 m/s 11.22 7.04 2.40 

Ceiling C-4-V 68 323 m/s 10.85 2.38 3.82 

Ceiling C-5-V 69 324 m/s 10.85 5.27 3.82 

1. thermocouples mounted near bi-directional probes (PC4 & PC5) on ceiling in middle of compartment 

2. Although specified in the original test plan, a bi-directional probe was not put in the plane of the exhaust 

duct. Instead, several probes were positioned downstream in the duct. 

3. Although specified in the original test plan, an aspirated thermocouple was not put in the plane of the 

supply duct. 

4. The Z location of the bi-directional probe was changed for Tests 10 and 16 to 2.20 m. For Tests 4 and 5 the 

Z location was 2.35 m.
161 

E INOPERATIVE INSTRUMENTS DURING TEST 

SERIES 

Table 2-2 lists some observations noted during the experiments, including among other things, 

the status of the data acquisition system, the instrumentation, and the measurements. A number 

of instruments did not function properly or displayed signals that were suspect (signal dropout, 

noise, odd behavior, etc.). The accompanying electronic data files do not include data that are 

clearly not meaningful. This information is documented in Tables E-1 - E-3 for each test. 

Table E-1: List of inoperative instruments during the test series. 

Inoperative Instrument 

Test 

Order 

Test 

No. Slab E TCs Cable TCs Tree TCs 

1 1 D-Ts-12 ** 4-3 *, 4-9 ** 

2 2 4-9 ** 

3 5 4-9 ** 

4 4 E-Ts-16,16’,17 ** 4-9 **, 3-10 * 

5 3 E-Ts-16,16’,17 ** 4-9 ** 

6 8 E-Ts-16,16’,17 ** 4-9 ** 

7 9 E-Ts-16,16’,17 ** 4-9 **, 5-9* 

8 13 E-Ts-16,16’,17 ** C-Tc-11, C-Ts-10 4-9 **, 5-9* 

9 14 E-Ts-16,16’,17 ** C-Tc-11, C-Ts-10 4-9 **, 5-9* 

10 18 E-Ts-16,16’,17 ** A-Ts-18, Vertical Cables Ts-35; 

Ts-36 *** 

4-9 **, 

TC Door-17 * 

11 7 4-9 ** 

12 10 4-9 ** 

13 16 E-Ts-16,16’,17 ** *** 4-9 ** 

14 15 E-Ts-16,16’,17 ** *** 4-all*,5-9* 

15 17 E-Ts-16,16’,17 ** *** 4-9 ** 

* intermittent signal dropout and/or odd behavior over all or part of the test. 

** eliminated from data set – inoperative. 

*** included in data set, but portions of the data may be suspect due to cable melting.
162 

Table E-2: List of inoperative instruments during the test series. 

Inoperative Instrument * 

Test 

Order 

Test 

No. 

Flux 

Gauge 

Wall Flux Gauges* and Associated TCs Additional 

Wall Flux 

Gauge TCs 

1 1 8, 9 North U-2, North U-4 TC South U-3-1 

2 2 9 North U-2 

3 5 North U-2 

4 4 North U-2 

5 3 7 North U-2, North U-3 (t>1176 s)**, South U-3 (t>1280s), 

South U-4 (t>1113 s)** 

TC East U-1-1 

(996<t<1068, 

t>1750)** 

6 8 North U-2, North U-3, South U-3, South U-4 TC East U-1-1; 

Ceiling C-2-1 ** 

7 9 North U-2, North U-3, South U-3, South U-4, East U-2, East 

U-4, West U-2, West U-4, Ceiling U-1, Ceiling U-8, Floor U4, Floor U-5 

TC East U-1-1 

8 13 All Gauges*** 

9 14 North U-2, North U-3, North U-6 (t>1185)**, South U-3, 

South U-4, East U-4, West U-4, Ceiling C-4, Ceiling C-5, 

Floor U-4, Floor U-5 

TC East U-1-1 

10 18 5, 7 North U-2, North U-3, North U-6, South U-3, South U-4, 

South U-6 (t>1054)**, East U-2, East U-4, West U-4, Ceiling 

U-1, Ceiling C-4, Ceiling U-8 (t>1288)**, Floor U-4, Floor 

U-5 

TC East U-1-1 

11 7 North U-6, South U-6, East U-2, East U-4, Ceiling C-4, 

Ceiling C-5, Ceiling C-6, Ceiling U-8, Floor U-6, Floor U-7 

TC East U-1-1 

12 10 North U-6, South U-6, Ceiling C-4, Ceiling C-5, Ceiling C-6, 

Ceiling U-8, Floor U-6, Floor U-7 

TC East U-1-1 

13 16 All Gauges*** 

14 15 7 North U-1 (171<t<190)**, North U-3, North U-6, South U-2 

(t>1302)**, South U-4 (t>760)**, South U-6, East U-4, West 

U-4, Ceiling U-1, Ceiling C-4, Ceiling C-5, Ceiling U-8, 

Floor U-4 (t>602), Floor U-5, Floor U-6, Floor U-7 

TC East U-1-1, 

TC South U-4-1 

(t>508)** 

15 17 5 North U-3, North U-6, South U-2, South U-4, South U-6, East 

U-4, West U-4, Ceiling U-1, Ceiling C-4, Ceiling C-5, Ceiling 

U-8, Floor U-4, Floor U-5, Floor U-6, Floor U-7 

TC East U-1-1 

* default situation: inoperative instrument eliminated from data set. 

** intermittent signal dropout and/or odd behavior during all or part of the test; thus, a portion of the data 

was eliminated 

*** gauges inoperative and/or covered to protect from overheating; all data eliminated.
163 

Table E-3: List of inoperative instruments during the test series. 

Test Inoperative Instrument 

Order 

Test 

No. Bi-directional probes Gas Analyzer 

1 1 CO * 

2 2 CO ** 

3 5 BP Door –12 * CO * 

4 4 CO * 

5 3 BP Door –12 * CO * 

6 8 CO (90<t<120)** 

7 9 BP Door –12; 13; 15 * CO * 

8 13 

9 14 BP Door –12 * CO * 

14 18 BP Door –12; 13 * CO **; O2-2 (t>350)** 

11 7 CO * 

12 10 CO * 

15 16 CO ** 

13 15 BP Door –12; 13 * 

10 17 

* eliminated from data set – inoperative. 

** intermittent signal dropout and/or odd behavior during all or part of the test.

165 

F RESPONSE TO QUESTIONS AND COMMENTS 

FROM JULY WPI INTERNATIONAL PEER REVIEW 

MEETING 

The following questions or comments regarding the test configuration and preliminary report of 

results for Benchmark Exercise # 3 were made during the meeting of peers in the International 

Collaborative Project on Fire Modelling at Worcester Polytechnic University (WPI) in July 

2003. The response to the questions was provided by NIST. 

1. What is the standard sampling frequency of the measurements? 

 1 Hz; see Section 1 of this report. 

2. The location of slab E and its TCs are not consistent in the specification document and 

addenda, and needs to be clarified. 

The designation of the thermocouples in Figure 6-1 has been modified. The thermocouple 

locations are listed in Appendix D of this report. 

3. Were the emissivity measurements taken only at room temperature? 

Yes; see Appendix A of this report. 

4. The equations in the test report that describe the relationship between compartment 

pressure and leakage area are not sufficient to reduce the measured data to values that 

can be used as input for fire models. 

Appendix C of this report includes a relation between compartment pressure and leakage 

area. 

5. The design of the supply ventilation system should have accounted for the length of 

duct necessary between the fan and the vent in order to ensure straight flow out of the 

vent into the compartment. 

The rule of thumb for straightening flow from a simple open duct is 10 to 12 diameters. To 

straighten an intensely swirling flow, flow strengtheners of some sort would be needed to 

condition the flow and the number of diameters downstream from the strengtheners would 

likely be significantly greater than 12 diameters. Space limitations in the NIST Large Fire 

Laboratory precluded using this approach for the 0.7 m x 0. 7 m (2.3 ft x 2.3 ft) supply duct. 

6. The radiative fraction for heptane needs to be clarified; different reports and addenda 

have identified two values. 

The radiative fraction for the spray fire burning in the open was measured to be 0.45, 

whereas the radiative fraction for a heptane pool fire burning in the open is 0.3 according to 

previous studies. 
166 

7. The footnote to the test matrix for the planned time should be clarified. 

The duration of the fuel flow is stipulated in Table 3-1 of this report. 

8. The observation of the melting of Slab E after Test 2 should be noted in the test report. 

Table 2-2 of this report highlights the status of Slab E during the test series. 

9. The types of heat flux gauges used for the wall and cables should be discussed. The 

discussion should include whether the gauge is measuring net or incident heat flux. The 

reference temperature for measurements of convective heat flux, i.e., Ts in Qconv. = 

h(Tg-Ts) should be specified. In general the discussion of heat flux measurements 

should be comprehensive, accurate, and clear. 

A detailed discussion of the gauges is given in Sections 4 and 10 of this report. The reference 

temperature for the cooled total heat flux gauges on the Targets was 75 o

C (167 °F). The 

reference temperature for the surface total heat flux gauges was the gauge temperature, which 

was measured as it changed during the experiment.

10. The 20 % discrepancy between the heat release measured using fuel mass flow and 

calorimetry should be investigated further. 

This has been completed. An experiment was conducted to determine the effect of fuel line 

heat-up on the fuel flow. The corrected fuel flow was then used to estimate the HRR. This 

value was compared to the HRR determined from calorimetry. The difference between the 

HRR estimated by the fuel flow and that measured using oxygen consumption calorimetry 

differs for each test. For Test 4, for example, the difference was approximately 100 kW, 

which was well-within the combined experimental uncertainties of the two measurements. 

11. Spray fires used in the test series may not be the most suitable fire source for model 

validation because of the difficulty of accurately measuring the heat release rate. Pool 

fires in a pan, which can be mounted on a weighing device, may be a more appropriate 

fuel source since the heat release rate can be more accurately measured. 

The above statement is subject to discussion. There are measurement issues associated with 

pool fires as well as spray fires. A pool fire burning down in a pan (and mounted on a 

weighing device) as suggested in the question will lead to a non-steady fuel burning rate. 

Also, water vapor from the gas phase can be transported by diffusion into the fuel pan, 

leading to significant systematic errors in the mass weight loss measurement for some fuel 

types. Depending on the pool size and the fuel type, the local mass flux of fuel on the surface 

of the pan can vary significantly. Planning for future experimentation should revisit this 

issue. 

12. The data used for the heat release rate calorimetry should be included in the data set 

for the tests. 

The calorimetry determination incorporates over thirty measurements including gas 

temperatures, velocities, and gas concentrations. The data used for the HRR calorimetry is 

not included in the accompanying data set as it would be of little utility without an 

understanding of the measurement system, procedures and the software required to process 
167 

the information. A complete discussion of the measurement and its uncertainty is given in 

Bryant et al [2003]. The calorimetry data can be made available upon special request. 

13. Is the specific heat of air or gas used for calculating the heat released through the door? 

The specific heat of air was used. 

14. Does the uncertainty (standard deviation) stated for the measurements include 

systematic and random errors? 

The analysis typically includes both systematic and non-systematic contributions to 

measurement uncertainty as appropriate for each particular measurement and as described in 

the text. 

15. Can the fan curves for the supply system be provided? 

The manufacturer fan curve that relates the static pressure to the blowing volume for various 

values of the fan power is included in Appendix G of this report. The fan had a 1⁄2 hp motor. 

The static pressure represents the pressure difference at the fan between ambient and the 

compartment pressure. This pressure was measured as described in Section 11 of this report. 

16. It is important to measure the temperature of the insulation near the internal copper 

conductors of the cables since this temperature has been found to be a good indicator of 

cable dysfunction. 

As specified in the original Test Plan, which was open to comment, thermocouples were 

placed just inside the outer cable insulation (see Section 6 in this report). The sequence of 

the tests and the point when the thermoset cables were replaced with thermoplastic 

cables should be indicated. 

Table 1-2 shows the test sequence and notes that the cables were changed before Test 7. 

Table 1-1 lists the cable type present during each of the tests. 

17. It would have been useful to measure mass in addition to extinction coefficient to 

confirm the smoke measurements. 

The limited resources were expended to make more important measurements. 

18. The calculation of an interface height using a 2-layer method when two distinctive 

layers do not exist may not have any physical meaning, especially when the 

compartment is filled with hot gas. 

We agree. The 2-layer method calculation was computed as defined and is subject to 

interpretation. 

19. There is a discrepancy between the TS 3-3 and TN 3-3 temperature, which are located 

in symmetrical positions on the wall. 

The difference varies from test to test. Figure 10-15 (section 10) presents differences in the 

peak temperatures of TS 3-3 and TN 3-3 during Tests 2-4. The reasons for the differences 

are not clear, but may be influenced by an asymmetric flow field induced by ventilation, or 

asymmetric heat flux due to the off-center fire pan position used in many of the tests. 
168 

20. Is there any measurement or indicator of how much soot was deposited on the walls? 

No, the quantity of soot deposited on the walls was not quantified. Some observations, 

however, were made. The walls were black after the toluene experiment and were darkened 

after the heptane tests. Observations suggest that only small amounts of soot were deposited 

on the walls. 
169 

G CURVE FOR VENTILATION SUPPLY FAN 

The manufacturer fan curves that relate the static pressure to the blowing volume for various 

values of the fan power is found below. The fan had a 1⁄2 hp motor. The static pressure represents 

the pressure difference at the fan between ambient and the compartment pressure, which is 

discussed in Section 11 of this report. 
170 

Figure G-1: The manufacturer supplied fan curve that relates the static pressure to the 

blowing volume for various values of the fan power. 
171 

H JUNCTION BOX 

An electrical junction box made of Feraloy was present in the ceiling of the compartment during 

the tests. The junction box was mounted on the ceiling and located on the compartment 

centerline, with its center 17.7 meters (58.1 feet) from the door. A drawing of the box is shown 

in Figure H-1. Feraloy is a proprietary manufactured product that is a gray-iron alloy. Its 

properties are somewhat similar to iron and are listed in Table A-4. The box was 30.0 cm x 30.0 

cm x 10.0 cm (11.8 in x 11.8 in x 3.9 in) in dimension with walls that were 7 mm (0.3 in) thick. 

Three type K thermocouples were placed in the box center and on the middle of the bottom and 

west surfaces. 

W est 

 

 

 TC -37 

 

 TS-39

10 cm 

 30 cm

 30 cm 

 TS-38 Bottom

Figure H-1: Thermocouples in and on the surface of the junction box (30 cm x 30 cm x 10 

cm) measuring temperatures in the box center and on the middle of the bottom and west 

surfaces. 

173 

I FUEL FLOW CORRECTION 

An experiment was conducted at the end of the test series to determine the effect of compartment 

heat-up on the fuel flow rate. It was hypothesized that the fuel flow during the experiment was 

not equal to the fuel flow during the ambient temperature fuel flow calibration, which was 

conducted before the tests. An experiment was conducted to test this hypothesis. 

The experiment was essentially a repeat of Test 2, which was nominally a 1 MW heptane fire 

with the door closed and the ventilation off. The only difference between the experiment and 

Test 2 was that the compartment was shorter by 3.66 meters (12.0 feet) or about 15 %. Test 2 

was representative of the fire scenarios in the test series, as eleven of the fifteen tests used the 

same nominal heptane fuel flow rate. The compartment door was closed and the ventilation off, 

since that scenario yielded the highest floor temperatures for a given fire size in the lower part of 

the compartment and therefore represented a worst-case scenario in terms of impact on the fuel 

flow. 

An in-line flow meter was placed downstream of the fuel pump, located just outside of the test 

compartment. Thermocouples were placed on the fuel line and at a location just upstream of the 

nozzle (by adding an in-line tee) to monitor the fuel temperature as it flowed through the 1.2 cm 

(0.50 in.) outer diameter copper tubing. A video camera was used to record the fire behavior. 

Figure I-1 shows the experimental results for the flow rate and the fuel temperature at two 

locations: just upstream of the nozzle and on the fuel line in the middle of the compartment. At 

early times in the experiment, the flow rate ramped-up from zero. An irregularity occurred at 

about 200 s, when the flow fell to near zero for a brief period. At that time, the video record 

shows that the spray fire nearly stopped. Figure I-1 shows that the temperature of the fuel near 

the nozzle at that time was greater than the boiling point of heptane, which is about 100 o

C (212 

°F). An interpretation of these events was that the fuel near the nozzle reached temperatures that 

exceeded the boiling point, and that a vapor lock of the flow occurred, essentially stopping the 

fuel flow. While this type of fire behavior was observed during this particular experiment, it was 

not observed during the test series. This may have been due to extra heat transfer to the fuel line 

caused by the addition of the thermocouple and the stainless steel tee just upstream of the nozzle. 

At 220 s, the fuel flow increased to about 1.8 L/min (0.06 ft3

/min), which was somewhat less 

than the target fuel flow of 1.93 L/min (0.07 ft3

/min). At the same time, the fuel temperature near 

the nozzle decreased below the fuel boiling point and maintained nearly the same value for the 

remainder of the experiment. Figure I-1 also shows that the fuel flow reached a fairly steady 

value for t > 300 s, with the average value during that period equal to 1.9 L/min (0.07 ft3

/min), 

which was greater than the target fuel flow by a factor of about 1.09. 

The temperature of the fuel flowing near the nozzle almost certainly played a key role in the rate 

of the fuel flow. As the temperature near the fuel nozzle decreased, the fuel flow tended to 
174 

increase, and vice- versa. The measurement of the temperature on the fuel line near the middle 

of the compartment, increased nearly monotonically, attaining a maximum temperature of 65 o

C 

(149 °F). The fuel temperature next to the fuel nozzle was always higher. As the temperature of 

a hydrocarbon liquid increases, the viscosity decreases. The viscosity of heptane decreases by 

approximately a factor of two as the temperature increases from 20 o

C to 90 oC (68 °F to 194 

°F)3

. For flow through a pipe for the Reynolds numbers estimated in the fuel line, a decrease in 

the fuel viscosity would lead to a reduced friction factor, a reduced pressure drop, and an 

increased flow rate. 

 

Although the exact values of the temperature and the heat flux in the compartment differed for 

the open door and closed door tests, and for the tests with different fire sizes, the compartment 

temperature increased in all of the tests. While it is not possible to model the change in flow 

during the various tests, because not enough information about the local transient fuel 

temperature or fuel pump behavior is available, the increased fuel flow measured in the 

compartment during the supplementary experiment is regarded here as representative. 

Table 3-2 shows the effect of the fuel flow correction on the estimated HRR if the correction 

factor is applied to all of the tests. 

0.00

0.50

1.00

1.50

2.00

2.50

0

50

100

150

200

250

0 100 200 300 400

Fuel LineTemperature (C) Fuel (L/min)

Nozzle Temperature (C)

Fuel Flow (L/min)

Nozzle Temperature (C)

Time (s)

Figure I-1: The measured fuel flow and nozzle temperature as a function of time during the 

1MW closed door compartment experiment. 

 

3

 Reid R.C., Prausnitz, J.M., and Poling, B.E., The Properties of Gases and Liquids, 4th Ed. McGraw Hill, NY, 1987. 
175 

J DOCUMENTATION OF PERSONAL 

COMMUNICATIONS 

Bundy, Matthew, NIST, Gaithersburg, MD. Conversation at NIST, May 2005, on determination 

of the HRR calibration of the calorimetric measurement in the NIST Large Fire Laboratory. 

Examination of gas chromatographic concentration measurements provided by the natural gas 

supplier, the Washington Gas Company, the caloric value of the natural gas was found to vary by 

approximately 1.3 % over a six-month period. 

Harris, Richard, NIST, Gaithersburg, MD. Conversation at NIST in March 2003. Mr. Harris 

manufactured the PVC slab, and observed that the softening temperature of the PVC was 

approximately 180 o

C (356 °F).