NIST Technical Note 1483 

Measurements of Heat and Combustion Products in 

Reduced-Scale Ventilation-Limited Compartment Fires 

Matthew Bundy 

Anthony Hamins 

Erik L. Johnsson 

Sung Chan Kim 

Gwon Hyun Ko 

David B. Lenhert 

U.S. Department of Commerce 

Technology Administration 

Building and Fire Research Laboratory 

National Institute of Standards 

and Technology 

Gaithersburg, MD 20899 
ii

NIST Technical Note 1483

Measurements of Heat and Combustion Products in 

Reduced-Scale Ventilation-Limited Compartment Fires 

Matthew Bundy 

Anthony Hamins 

Erik L. Johnsson 

Sung Chan Kim 

Gwon Hyun Ko 

David B. Lenhert 

U.S. Department of Commerce 

Technology Administration 

Building and Fire Research Laboratory 

National Institute of Standards 

and Technology 

Gaithersburg, MD 20899 

July 2007 

 

U.S Department of Commerce 

Carlos M. Gutierrez, Secretary 

Technology Administration 

Robert Cresanti, Under Secretary 

of Commerce for Technology 

National Institute of Standards and Technology 

William Jeffery, Director
iii

Disclaimer: 

Certain commercial equipment, instruments, or materials are identified in this document. Such 

identification does not imply recommendation or endorsement by the National Institute of 

Standards and Technology, nor does it imply that the products identified are necessarily the best 

available for the purpose. 

National Institute of Standards and Technology Technical Note 1483 

Natl. Inst. Stand. Technol. Tech. Note 1483, 147 pages (July 2007) 

CODEN: NSPUE2 
iv

Table of Contents 

LIST OF FIGURES ...................................................................................................................... VI 

LIST OF TABLES..................................................................................................................... VIII 

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

1.1 Background and Relationship to Current Research.................................................................. 2 

1.2 Approach and Scope ................................................................................................................. 3 

2 EXPERIMENTAL METHOD..................................................................................................... 4 

2.1 Reduced Scale Enclosure.......................................................................................................... 4 

2.1.1 Design and construction..................................................................................................... 4 

2.1.2 Doorway Variations........................................................................................................... 6 

2.2 Burner Designs.......................................................................................................................... 6 

2.3 Experimental Conditions .......................................................................................................... 7 

2.4 Measurement Locations............................................................................................................ 8 

2.5 Measurement Instrumentation ................................................................................................ 11 

2.5.1 Heat Release Rate ............................................................................................................ 11 

2.5.2 Fuel Metering................................................................................................................... 12 

2.5.3 Gas Species ...................................................................................................................... 13 

2.5.4 Extractive Soot Measurement.......................................................................................... 16 

2.5.5 Temperatures.................................................................................................................... 17 

2.5.6 Pressures and Velocities .................................................................................................. 21 

2.5.7 Gas Chromatography ....................................................................................................... 22 

2.5.8 Optical Extinction Soot Measurements ........................................................................... 24 

2.5.9 Heat Fluxes ...................................................................................................................... 28 

2.6 Data Acquisition ..................................................................................................................... 29 

2.7 General Data Corrections........................................................................................................ 30 

3 RESULTS .................................................................................................................................. 32 

3.1 Heat Release Rate ................................................................................................................... 32 

3.2 Temperatures........................................................................................................................... 42 

3.3 Doorway Velocities and Temperatures................................................................................... 48 

3.4 Interior Gas Species and Soot................................................................................................. 51 

3.4.1 Gas Chromatography ....................................................................................................... 61 

3.5 Optical Soot ............................................................................................................................ 65 

3.6 Heat Fluxes ............................................................................................................................. 69 

4 DISCUSSION OF RESULTS ................................................................................................... 71 
v

4.1 Fuel and Fire Size Comparisons ............................................................................................. 71 

4.2 Doorway Ventilation Comparisons......................................................................................... 73 

4.3 Construction Material Comparisons ....................................................................................... 74 

4.4 Burner Type Comparisons ...................................................................................................... 80 

5 ANALYSIS OF COMPARTMENT CHEMISTRY.................................................................. 82 

5.1 Mixture Fraction ..................................................................................................................... 82 

5.2 Chemical Equilibrium........................................................................................................... 100 

5.3 Carbon Balance..................................................................................................................... 102 

5.4 Post-Compartment Product Yields ...................................................................................... 111 

5.5 Combustion Efficiency ......................................................................................................... 114 

5.6 Summary of Chemical Analysis ........................................................................................... 117 

6 SCALING DISCUSSION........................................................................................................ 118 

7 CONCLUSIONS...................................................................................................................... 120 

8 FUTURE WORK..................................................................................................................... 121 

9 REFERENCES ........................................................................................................................ 122 

10 ACKNOWLEDGEMENTS................................................................................................... 125 

11 APPENDICES ....................................................................................................................... 126 

A. Analysis of Thermocouple Temperature Measurement......................................................... 126 

B. Analysis of Probe Interactions ............................................................................................... 138 

C. Channel Lists.......................................................................................................................... 142 

D. Equipment List....................................................................................................................... 147 
vi

List of Figures 

Figure 1. Perspective views of the Reduced Scale Enclosure and upper layer gas sampling probes drawn to scale. Dimensions are given 

with respect to interior walls......................................................................................................................................................................... 5 

Figure 2. Dimensional drawing of burners used in the RSE experiments. Burners D and E were similar to Burner F, but the diameters were 

22 cm and 40 cm, respectively...................................................................................................................................................................... 7 

Figure 3. Photograph of extractive sampling probes and aspirated thermocouple at the front sample location (photo taken with rear wall 

removed). ...................................................................................................................................................................................................... 9 

Figure 4. Isometric view of Reduced Scale Enclosure showing the relative position of doorway measurement probes. .......................................... 9 

Figure 5. Isometric semi-transparent view of the RSE interior measurement probe locations and burner. .............................................................. 10 

Figure 6. Schematic drawing of 3 m hood and exhaust stack instrumented for calorimetry and light extinction measurements. ........................... 12 

Figure 7. Exhaust gas sampling system used for heat release rate measurement. ..................................................................................................... 13 

Figure 8. Schematic drawing of gas sampling system for series 1 tests #1 to test #6................................................................................................ 15 

Figure 9. Schematic drawing of gas sampling system for series 2 (test #6.5 to test #16). ........................................................................................ 15 

Figure 10. Schematic drawing of soot sampling system. ............................................................................................................................................ 17 

Figure 11. Detailed drawing of aspirated thermocouple using NACA design [34].................................................................................................... 18 

Figure 12. Schematic drawing of aspirated thermocouple measurement hardware.................................................................................................... 19 

Figure 13. Type K thermocouple on compartment floor surface near a 6 mm diameter total heat flux gauge.......................................................... 20 

Figure 14. Type K thermocouple attached to the outer surface of the compartment.................................................................................................. 21 

Figure 15. Schematic of the optical extinction measurement apparatus used during test #6. .................................................................................... 27 

Figure 16. Schematic of the optical extinction measurement apparatus used during test #13 and test #14 including a blow-up of the circled 

optical sampling region............................................................................................................................................................................... 27 

Figure 17. Heat release rate results for natural gas test #3 using Burner A. ............................................................................................................... 34 

Figure 19. Heat release rate results for methanol test #12 using spray burner C........................................................................................................ 35 

Figure 21. Heat release rate results for polystyrene test #16. A heptane spray was used to ignite 6 kg of polystyrene pellets................................. 36 

Figure 23. Heat release rate results for toluene pool fire (test #10). ........................................................................................................................... 37 

Figure 25. Heat release rate results for ethanol test #11.............................................................................................................................................. 38 

Figure 27. Heat release rate results for heptane test #15. ............................................................................................................................................ 39 

Figure 29. Steady state heat release rate results. Dashed line represents ideal or complete burning. ....................................................................... 40 

Figure 30. Gas temperature measurement results for 4 positions inside RSE. Natural gas test #3. ........................................................................... 43 

Figure 31. Gas temperature measurements at two front interior locations inside RSE during the heptane spray fire (test # 15). ............................ 44 

Figure 32. Steady state average temperature measurements at interior locations for repeated natural gas fires (tests #1, #2, #3). The lines in 

this figure are piecewise cubic polynomial fits to the data. ....................................................................................................................... 44 

Figure 33. Steady state average temperature results from aspirated thermocouple measurement at front gas sampling location. ........................... 45 

Figure 34 Steady state average temperature results from aspirated thermocouple measurement at rear gas sampling location.............................. 45 

Figure 35. Comparison of doorway centerline temperatures measured using aspirated and bare bead thermocouple at the same position. 

Steady state average values for all RSE tests. ............................................................................................................................................ 49 

Figure 36. Steady state doorway temperature measurements for natural gas test #3, HRR=265 kW. Error bars in this figure are 2 standard 

deviations. ................................................................................................................................................................................................... 49 

Figure 37. Doorway centerline velocity profiles for natural gas fire (test #3)............................................................................................................ 50 

Figure 38. Doorway centerline velocity profiles for heptane spray fire (test #15). .................................................................................................... 50 

Figure 39. Transient gas volume fractions and soot mass fractions for heptane test #15........................................................................................... 52 

Figure 40. Transient gas volume fractions and soot mass fractions for polystyrene test #16. ................................................................................... 53 

Figure 41. Steady state gas and soot species results for natural gas full door tests #1, #2, #3, and #65. The lines in this figure are piecewise 

cubic polynomial fits to the data................................................................................................................................................................. 53 

Figure 42. Steady state average oxygen volume fraction measurements at front sample probe location. ................................................................. 54 

Figure 43. Steady state average oxygen volume fraction measurements at rear sample probe location.................................................................... 54 

Figure 44. Steady state average carbon dioxide volume fraction measurements at front sample probe location. ..................................................... 55 

Figure 45. Steady state average carbon dioxide volume fraction measurements at rear sample probe location........................................................ 55 

Figure 46. Steady state average carbon monoxide volume fraction measurements at front sample probe location. ................................................. 56 

Figure 47. Steady state average carbon monoxide volume fraction measurements at rear sample probe location.................................................... 56 

Figure 48. Steady state average total hydrocarbon volume fraction measurements at front sample probe location.................................................. 57 

Figure 49. Steady state average total hydrocarbon volume fraction measurements at rear sample probe location. .................................................. 57 

Figure 50. Steady state gravimetric soot mass fraction measurements at front sample probe location. .................................................................... 58 

Figure 51. Steady state gravimetric soot mass fraction measurements at rear sample probe location. ...................................................................... 58 

Figure 52. Comparison of total hydrocarbons measured using the GC and the total hydrocarbon analyzer (THC Front), both expressed on a 

CH4 basis..................................................................................................................................................................................................... 64 

Figure 53. GC gas species composition measurements in the natural gas fire #65, HRR=429 kW........................................................................... 64 

Figure 54 GC measurements of methane, ethyne plus ethene, and benzene in the heptane pool fire (test #7) and heptane spray fire (test 

#15), front gas sample probe....................................................................................................................................................................... 65 

Figure 55. History of the soot mass concentration during natural gas fire (test #6). Mean values of the soot mass concentration were 

determined over the same periods as the gravimetric measurements as indicated by the dotted boxes. The fire heat release rate is 

also shown................................................................................................................................................................................................... 67 

Figure 56. Soot mass concentration in the doorway during natural gas fire (test #6) as a function of the HRR. ...................................................... 67 

Figure 57. History of the particle mass fraction in the doorway during test #13, polystyrene burning in the 20 cm diameter burner. The fire 

heat release rate is also shown. ................................................................................................................................................................... 68 

Figure 58. History of the soot mass concentration in the doorway during test #14, polystyrene burning in a 40 cm diameter burner. The fire 

heat release rate is also shown. ................................................................................................................................................................... 68 

Figure 59. Total heat flux gauge measurements at two locations on the interior floor of the RSE for test #15 using the spray burner with 

heptane fuel. ................................................................................................................................................................................................ 69 
vii

Figure 60 Steady state total heat flux measurement at front floor location. Note: Front heat flux gauge was partially blocked by debris for 

polystyrene test #16. ................................................................................................................................................................................... 70 

Figure 61. Steady state total heat flux measurement at rear floor location................................................................................................................. 70 

Figure 62. Photograph of soot agglomerate on the water inlet tube of the burner after toluene pool fire (test #10). ................................................ 72 

Figure 63 The CO volume fraction measured in the stack and at the front and rear of the compartment as a function of time during the 

burning of 6.0 kg of polystyrene pellets (Test 16). Photographs at various times show the fire appearance through the open 

doorway....................................................................................................................................................................................................... 73 

Figure 64. Photo of rear gas sample location and Marinite I construction before test #2. ......................................................................................... 77 

Figure 65. Photo of rear gas sample location and Marinite I construction after test #2 with natural gas................................................................... 77 

Figure 66. Photo inside enclosure after test #6 showing Marinite I condition. Large chunks of ceiling fell to the floor during this test................ 78 

Figure 67. Photo of rear gas sample location showing Kaowool M-board and furnace pin construction prior to test #6.5. ..................................... 78 

Figure 68. Photo of rear gas sample location after test #6.5 with natural gas using Kaowool M-Board. .................................................................. 79 

Figure 69. Photo of enclosure with rear wall removed after test #10 using toluene. Notice melted stainless steel furnace pins and bowed Mboard wall but very little cracking of walls. ............................................................................................................................................... 79 

Figure 70. Photo of rear lower aspirated probe (TR24A) after test #16 with polystyrene. Stainless steel furnace pins melted and dripped 

along the wall. ............................................................................................................................................................................................. 80 

Figure 71 The equivalence ratio as a function of mixture fraction for nonpremixed flames burning methane and n-heptane. ................................ 86 

Figure 72 The mass fraction vs. the mixture fraction calculated by the single-parameter mixture fraction model. .................................................. 86 

Figure 73 Mass fractions of front and rear compartment gas species as a function of mixture fraction for the natural gas fire tests #1-#3, and 

#6: (a) transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady. ................... 89 

Figure 74 Mass fractions of front and rear compartment gas species as a function of mixture fraction for the heptane fire tests #4, #5, #7 

and #15: (a) transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady. .......... 92 

Figure 75 Mass fractions of front and rear compartment gas species as a function of mixture fraction for the toluene fire test #10: (a) 

transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady................................ 93 

Figure 76 Mass fractions of front and rear compartment gas species as a function of mixture fraction for polystyrene fire tests #13, #14, and 

#16: (a) transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady. ................. 94 

Figure 77 Mass fractions of front and rear compartment gas species as a function of mixture fraction (with soot included) for the timeaveraged measurements when the HRR was quasi-steady in the experiments burning (a) heptane, (b) toluene, and (c) 

polystyrene. ................................................................................................................................................................................................. 95 

Figure 78 Comparison of mixture fractions calculated with and without soot using the time-averaged species measurements when the HRR 

was quasi-steady in the experiments burning (a) heptane, (b) toluene, and (c) polystyrene. .................................................................... 96 

Figure 79 Transient values of (a) CO2 Y and (b) YCO in the front and rear of the compartment as a function of the mixture fraction (without 

soot) with symbols colored to represent the local temperature for the heptane fire tests #4, #5, and #7.................................................. 97 

Figure 80 Mass fractions of front and rear compartment gas species as a function of mixture fraction for methanol fire tests #8 and #12: (a) 

transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady................................ 98 

Figure 81 Mass fractions of front and rear compartment gas species as a function of mixture fraction for ethanol tests #9 and #11: (a) 

transient measurements and (b) time-averaged measurements during the period when the HRR was quasi-steady................................ 99 

Figure 82 Comparison between the equilibrium calculations and the time-averaged species mass fractions measured in the front and rear of 

the compartment as a function of the mixture fraction during the period when the HRR was quasi-steady in the natural gas fire 

tests #1-#3, and #6. ................................................................................................................................................................................... 101 

Figure 83 The CO and soot yields as a function of the local equivalence ratio for the time-averaged measurements during the period when 

the HRR was quasi-steady in the heptane, toluene and polystyrene fires................................................................................................ 105 

Figure 84. The values of FCO and FSoot as a function of the local equivalence ratio during the period when the HRR was quasi-steady in the 

heptane, toluene and polystyrene fires...................................................................................................................................................... 106 

Figure 85 The CO and soot yields as a function of the local equivalence ratio during the period when the HRR was quasi-steady in the 

natural gas and ethanol fires. .................................................................................................................................................................... 107 

Figure 86. The CO yield as a function of the soot yield during the period when the HRR was quasi-steady in the heptane, toluene and 

polystyrene fires. Also shown is a line representing the results of Koylu [46]. ..................................................................................... 108 

Figure 87. The ratio of the CO to soot yield as a function of the local equivalence ration during the period when the HRR was quasi-steady 

in the heptane, toluene and polystyrene fires. Best fit lines to the data and a line representing the results of Koylu [46] are also 

shown. ....................................................................................................................................................................................................... 108 

Figure 88 The CO yield as a function of the soot yield for the same data shown in Fig. 85 during the period when the HRR was quasisteady in the natural gas and ethanol fires................................................................................................................................................ 109 

Figure 89 The ratio of the CO to soot yield as a function of the local equivalence ration for the same data shown in Fig. 85 during the 

period when the HRR was quasi-steady in the natural gas and ethanol fires. ......................................................................................... 109 

Figure 90 The CO2 yield in the exhaust stack as a function of the fire heat release rate during the periods when the HRR was quasi-steady 

for each of the fuels tested. ....................................................................................................................................................................... 112 

Figure 91. The CO yield in the exhaust stack as a function of the fire heat release rate during the periods when the HRR was quasi-steady 

for each of the fuels tested. ....................................................................................................................................................................... 112 

Figure 92. The soot yield in the exhaust stack as a function of the fire heat release rate during the periods when the HRR was quasi-steady 

for each of the fuels tested. ....................................................................................................................................................................... 113 

Figure 93. The total hydrocarbon yield in the exhaust stack as a function of the fire heat release rate during the periods when the HRR was 

quasi-steady for each of the fuels tested. .................................................................................................................................................. 113 

Figure 94. The combustion efficiency in the exhaust stack as a function of the fire HRR........................................................................................ 115 

Figure 95. The local combustion efficiency at the rear and front compartment sampling locations as a function of the fire heat release rate 

during the periods when the HRR was quasi-steady for three fuels. ....................................................................................................... 116 

Figure 96. The local combustion efficiency at the rear and front compartment sampling locations as a function of the oxygen volume 

fraction during the periods when the HRR was quasi-steady in the heptane fire. ................................................................................... 116 
viii

List of Tables 

Table 1. List of test numbers and key experimental conditions....................................................................................8

Table 2. Location of measurement probes inside of the enclosure. .............................................................................10

Table 3. Location of measurement probes in the enclosure doorway..........................................................................11

Table 4. Gas chromatograph settings and method parameters....................................................................................23

Table 5. List of retention times for identified compounds...........................................................................................23

Table 6. Components used in the optical extinction measurements. ...........................................................................26

Table 7. Tests in which optical light extinction measurements were conducted. ........................................................26

Table 8. Description of calorimetry measurement labels. ..........................................................................................32

Table 9. Composition (volume fraction %) and heating values of natural gas used in RSE tests...............................33

Table 10. Summary of time averaged steady-state results of HRR and exhaust stack species measurements. ..........41

Table 11. Description of interior gas temperature measurement labels......................................................................42

Table 12. Steady state rear gas temperatures and total heat flux to the floor inside the RSE. ....................................46

Table 13. Steady state front gas temperatures and total heat flux to the floor inside the RSE....................................47

Table 14. Description of interior gas species and soot measurement labels. ..............................................................51

Table 15. Summary of steady state rear gas and soot sample probe measurements. ...................................................59

Table 16. Summary of steady state front gas and soot sample probe measurements..................................................60

Table 17. Summary of GC sample results. Values in bold were identified as trace species. Blank spaces 

imply that the measurements were below the detection limits. ...................................................................62

Table 18. Summary of GC measurement uncertainty analysis results. Blank spaces imply that the 

measurements were below the detection limits. ..........................................................................................63

Table 19. Soot measurements during narrow doorway natural gas fire (test #6).........................................................66

Table 20. Description of interior total heat flux measurement labels. ........................................................................69

Table 21. Comparison of thermal, physical and construction properties of wall lining materials..............................76

Table 22. Stoichiometric value of the mixture fraction (Zst) for different fuels..........................................................84

Table 23. Average fractional soot, CO and CO/Soot ratio at the front and rear compartment measurement 

locations.....................................................................................................................................................104

Table 24. Time-averaged yields of soot, CO, and the ratio (yCO /ys) at the front and rear compartment 

measurement locations and in the exhaust stack for all fuel types. ...........................................................110

Table 25. List of non-dimensional scaling parameters for compartment fires and the range of values 

examined in this study. ..............................................................................................................................118

 
1

1 Introduction

A series of new reduced-scale compartment fire experiments were conducted, which included 

local measurements of temperature and species composition. The measurements are unique to 

the compartment fire literature. By design, the experiments provided a comprehensive and 

quantitative assessment of major and minor carbonaceous gaseous species and soot at two 

locations in the upper layer of fire in a 2/5 scale International Organization for Standards (ISO) 

9705 room. The enclosure defined in the international standard ISO 9705 “Full-scale room test 

for surface products” [1] is an important structure in which to conduct fire research. Many 

dozens of research projects and journal articles have focused on this enclosure and the standard 

describing its use. It is a common reference point for studies of many fire-related phenomena as 

well as fire modeling efforts. 

While some previous studies have considered the mixture fraction to analyze experimental 

compartment fire data, few have considered minor hydrocarbon species and none have 

considered soot. In tandem, accurate measurements of temperature at these same locations 

allowed analysis of thermal effects on species concentrations. A wide range of fuel types were 

considered, including aliphatic hydrocarbons (natural gas and heptane), aromatic hydrocarbons 

(toluene and polystyrene) and alcohols (methanol and ethanol). 

Field models, such as the National Institute of Standards and Technology (NIST) Fire Dynamics 

Simulator (FDS) [2], are widely used by fire protection engineers to predict fire growth and 

smoke transport for practical engineering applications. Field models numerically solve the 

conservation equations of mass, momentum and energy that govern low-speed, thermally-driven 

flows with an emphasis on smoke and heat transport from fires. All field models have strengths 

and weaknesses. Among the various assumptions used in the development of previous versions 

of FDS, all chemical species were tied to the mixture fraction state relations. A single mixture 

fraction variable cannot be used for the prediction of carbon monoxide and soot, and the yield of 

these species was prescribed in FDS 4, rather than predicted. In fact, the yield of these species is 

usually not constant, but a complex function of their time-temperature history. In practice, an 

engineer using FDS 4 would choose combustion product yields directly from literature values for 

well-ventilated burning, using data from a bench-scale apparatus [3]. Using this approach, the 

carbon monoxide (CO) volume fraction for pool fire burning in an under-ventilated compartment 

can be underestimated by as much as a factor of ten. A new version of FDS (version 5) is 

currently being tested which implements a predictive model of CO production. 

The experimental results provided in this report are the first step of a long-term NIST project to 

generate the data necessary to test our understanding of fire phenomena in enclosures and to 

guide the development and validation of field models by providing high quality experimental 

data. The experimental plan was designed in cooperation with developers of the NIST FDS 

model to assure that the measurements would be of maximum value. Advanced development of 

FDS and other field models is extremely important, since it will lead to improved accuracy in the 

prediction of underventilated burning, typical of fire conditions that occur in structures. 

Improving models for under-ventilated burning will foster improved prediction of important life 

safety and fire dynamic phenomena, including fire spread, backdraft, flashover, and egress 

(involving the presence of toxic gases and smoke), which are critically important for application 

of fire models for fire safety. In summary, the main objective of this project is to provide an 
2

improved understanding of the physics, chemistry, and structure of underventilated compartment 

fires, and to provide experimental measurements to guide the development of fire chemistry submodels. 

 1.1 Background and Relationship to Current Research

Experimental research on enclosure fires has been on-going in fire research laboratories and 

academic institutions over the last 50 years. The motivation has varied from applied 

investigations studying particular fire scenarios to more fundamental work with the goal of 

understanding toxic species production behavior in fires. Some of the fundamental research that 

tried to ascertain ventilation and upper-layer effects on enclosure fire chemistry was conducted 

in well-controlled hoods. Sometimes, the main objectives of this research was to generally 

develop and validate fire models or particular structural fire simulations, while much of the 

research was conducted to acquire a better understanding of complex enclosure fire dynamics 

with a focus on chemical and thermal conditions. This section provides an overview of some of 

the recent research efforts in enclosure fires and highlights some of the more pertinent 

experimental work. 

Research conducted at Harvard University and the California Institute of Technology in the 

1980s explored fires burning under an exhaust hood (false ceiling) to simulate the layer effect of 

an enclosure fire, e.g. [4,5]. The relative distance of the fire below the hood was adjusted to vary 

the entrainment of air into the plume before it entered the upper layer. These experiments 

focused on underventilated burning, pathways for air to enter the upper layer, and the validity of 

the concept of “global equivalence ratio” (GER) which is the fuel-to-air mass ratio normalized 

by the mass ratio required for stoichiometric burning. Some recent modeling work by Cleary 

and Kent [6], has also focused on experimental data from hoods. In a recent study, Brohez et al. 

explored the use of a bench-scale calorimeter to measure fire properties of materials burning in 

underventilated conditions [7,8]. 

Research at NIST by Bryner et al. further explored the global equivalence ratio concept and 

carbon monoxide production in a reduced (2/5) scale enclosure with natural gas as the principal 

fuel [9]. The results showed that the upper layer in enclosure fires is not homogeneous, and that 

CO can be produced in greater quantities than predicted by the GER concept, depending on 

temperatures and flow patterns developed within an enclosure. The current effort is meant to 

overlap some of the conditions explored by Bryner et al. and to repeat and fill gaps in the data. 

Pitts expanded the work to full-scale and other fuels such as heptane and wood. It was 

established that wood pyrolysis in the upper layer of an enclosure fire can produce high 

concentrations of CO directly without further oxidation to CO2 [10]. A subsequent study by 

Lattimer confirmed and expanded on this research [11]. 

Researchers at Virginia Tech investigated fires in a reduced-scale enclosure that directed the air 

inflow through slots in the floor connected to a duct where instrumentation was used to quantify 

air entrainment [12]. Several fuels were studied, and this configuration produced results 

consistent with GER predictions due to the more distinct, less dynamic nature of the gas layer 

structure. Later work used a more typical enclosure design and focused on transport of gas 

species outside the doorway and how it was affected by doorway geometry, soffit design, and 

hallway configuration [13]. More recently, Gann et al [14] conducted research on transport of 
3

toxic species in a full-scale enclosure with a corridor. These data were analyzed by 

Hirschler [15]. Researchers in Sweden conducted a study [16] of underventilated fires in an ISO 

9705 room with a window vent of varying height. Several polymer fuel types were included in 

this study and measurements of local equivalence ratio and toxic gas species were performed. 

Pitts [17] provides a comprehensive review of the application of the GER concept to predict CO 

concentration in building fires, using data from the Harvard and Cal Tech hood experiments 

[3,4], the Virginia Tech enclosure studies [11], and the NIST reduced-scale enclosure 

experiments [8,9]. Several CO formation mechanisms were identified, which were substantiated 

by detailed chemical kinetic modeling. While the GER concept is of limited utility for predicting 

the local CO concentration, important aspects of enclosure fire dynamics and chemistry are 

highlighted in this paper. 

Several recent experimental studies [18,19,20] have used very small scale enclosures (0.21 m3

, 

0.06 m3

, and 0.05 m3, respectively) while investigating underventilated burning of propane and 

heptane fires. These bench-scale studies described the structure and dynamics of underventilated 

burning including extinction, flame projection and flame stability. Another recent study [21,22] 

has used an intermediate-scale enclosure similar to that used for this paper, but a roof vent was 

added as well. 

Recently, NIST has conducted a number of high-profile case studies in which realistic-scale 

mock-ups of actual fire scenarios were recreated with the ultimate goal of improving building 

codes and standards. These studies included the World Trade Center disaster investigation [23] , 

the Rhode Island Station nightclub fire [24], and the Chicago Cook County Administration 

Building fire investigation [25]. The compartment fires in all of these studies burned real 

furnishings and became underventilated as the fire evolved. In addition, a series of large-scale 

compartment fire experiments were conducted to simulate an over-ventilated fire in a nuclear 

power plant cable room [26] to provide data for fire model validation. 

 1.2 Approach and Scope

The series of experiments reported on here was conducted in a reduced scale (2/5 ISO 9705 

room) enclosure (RSE). The experiments repeated and extended a part of the work of Bryner 

and coworkers [9]. Similar to Bryner’s experiments, natural gas served as a fuel; the burning of 

heptane, toluene, methanol, ethanol, and polystyrene was also investigated. In most experiments, 

the fuel was controlled and metered by flow valves or pumped into a pool burner or spray nozzle. 

Experiments were run to near-steady conditions. Multiple fire sizes were run consecutively to 

decrease the time required to approach steady-state. Ventilation was varied during some 

experiments by modifying the door opening. Two types of enclosure lining materials were 

investigated and compared. 

Temperature and species composition measurements were made at many of the same nominal 

locations as studied previously by Bryner [9]. Measurements included CO, CO2, temperature, 

heat fluxes, and dynamic pressures (used to obtain velocities). One emphasis of this series was 

to develop techniques for the measurement of hydrocarbons and soot. Hydrocarbons were 

measured with Flame Ionization Detector (FID) total hydrocarbon analyzers and gas 

chromatography (GC). The GC measurements were used to independently validate the total 
4

hydrocarbon measurements and to allow accurate determination of carbon mass distribution. 

The quantification of hydrocarbon species was needed to describe the chemical structure of 

underventilated fires. Soot samples were extracted from within the enclosure and measured 

gravimetrically. Optical soot measurements were performed at the doorway. 

The fuels included in this test series were selected to cover a wide range of combustion 

properties and to simulate fuels encountered in actual building fires. Gases, liquids and solids 

were selected for testing to cover a wide range of physical properties. Realistic materials 

represent complex multi-component fuels. In this study, all of the fuels selected were 

homogeneous single component fuels to simplify the analysis and attempt to find generalizable 

trends in the results. Real materials are often oxygenated. This includes many types of 

commodity materials including nylon (e.g., carpet), cellulose (e.g. paper and building products), 

polyester (e.g., fabric), epoxy (e.g., adhesives), polymethylmethacrylate (PMMA), and POM 

(polyoxometalate). In this study, alcohols were selected as a surrogate to represent the 

compartment fire chemistry in the burning of oxygenated fuels. 

In a real compartment fire, fuel sources are physically distributed throughout the compartment. 

In this study, a single location for the fuel was used to simplify the analysis of the experimental 

results for the purpose of model validation. Multiple fuel locations would have led to uncertainty 

in the specification of the location of the heat source, which is a critical boundary condition in a 

CFD fire calculation and, therefore, crucial information for model validation. 

In a real compartment fire, heat feedback and natural ventilation give rise to important aspects of 

the structure and dynamics of the fire, such as the temperature field and the spatial distribution of 

combustion products. This study deliberately set out to investigate representative fire conditions 

at two key locations in the upper layer of the compartment, which were selected based on a series 

of CFD fire modeling calculations. The upper layer locations were selected to provide two 

distinct conditions in the upper layer, one relatively close to the natural ventilation flow of fresh 

air through the doorway and the other relatively far from the doorway, on the far side of the fire 

source. The design calculations confirmed that these locations would provide a range of local 

conditions in terms of the combustion species equivalence ratio and the temperature that would 

be useful for the construction of a database for model development and validation. To enhance 

the range of conditions investigated and in an attempt to seek information on the relationship 

between the combustion products and generalizable local flame conditions, a broad range of fire 

heat release rates and a number of very different fuel types were selected for study. At the same 

time, the effect of compartment ventilation was changed to induce a range of mixing and 

compartment fire conditions. 

2 Experimental Method 

2.1 Reduced Scale Enclosure 

2.1.1 Design and construction 

Experiments were conducted using an enclosure, shown in Fig. 1, that is roughly a 2/5 scale 

replicate of the ISO 9705 room [1]. The steel frame for this enclosure was used extensively in 

the early 1990s to study carbon monoxide production in compartment fires. A detailed 
5

description of the design and construction of the Reduced Scale Enclosure (RSE) can be found in 

the original NIST report [9]. The original report described the “as designed” internal 

dimensions of the enclosure as 98 cm wide × 98 cm tall × 146 cm deep, however for this report, 

the “as constructed” internal dimensions were measured as 95 cm wide × 98 cm tall × 142 cm 

deep, with a pre-burn uncertainty of less than 1 cm on each dimension. The uncertainty in the 

internal dimensions increased as more fire experiments were conducted, e.g., during and after 

some of the fire tests, the walls were observed to deform in local areas by as much as 10 cm. 

142 cm

95 cm

29 cm

98 cm

81 cm

19 cm

10 cm

10 cm 29 cm

48 cm

Figure 1. Perspective views of the Reduced Scale Enclosure and upper layer gas sampling 

probes drawn to scale. Dimensions are given with respect to interior walls. 

The steel frame of the enclosure was lined with 2 layers of 1.27 cm thick insulation board. For 

the first six tests, a calcium silicate board (Marinite I) was used. For all other tests a rigid self 

supporting ceramic fiber (alumina and silica) board (Kaowool M-board) was used. The location 

of the retention bolts for the Marinite board were the same as the original test series [9]. Because 

the M-board sheet size was 122 cm × 91 cm, a single board would not span the length of the 

enclosure and joint seams were present. Stainless steel furnace pins were used to secure the 

M-board in both the original bolt locations (six retention points per wall and ceiling) and near the 
6

seams. The performance of the two different lining materials is discussed in Sec. 4.3. A 

comparison of the fire tests using the different lining materials showed no significant effect on 

the gas temperature and species measurements. 

2.1.2 Doorway Variations 

The standard doorway geometry (shown in Fig. 1) was 81 cm tall × 48 cm wide and centered 

horizontally on the 95 cm front wall. The bottom of the door was aligned with the inside floor. 

The inside floor was 43 cm above the laboratory floor. This configuration was used for all but 

two of the tests described here. The narrow door tests (listed in Table 1) were test #5 and test #6. 

The narrow doorway geometry was 81 cm tall × 24 cm wide. 

2.2 Burner Designs 

Four different burner designs, shown in Fig. 2, were used in this test series to accommodate the 

different fuels. A 13 cm square gravel-filled burner (Burner A) was used for the first three tests 

using natural gas. The area of this burner matched the area of the round burner used in the 

original test series [9]. The rim of the burner was 15 cm above the floor. Natural gas was 

delivered to the burner by a 1.3 cm tube that was fed through the floor and wrapped with 

Kaowool blanket insulation. The insulation nearly filled the space below the burner. The square 

geometry of burners A and B was chosen to match the rectangular grid used in FDS simulations. 

A 25 cm square liquid cooled burner (Burner B) was used for both natural gas and liquid fuels. 

The burner was designed to have a pool surface area that increased with the depth of the pool. 

The maximum depth of the pool was 6.5 cm and the burner walls were at a 24 degree angle with 

respect to the horizontal plane. This feature allows for different size steady pool fires with a 

single burner. Burner B was designed with a 2.5 cm vertical rim to prevent fuel from spilling out 

of the burner. This burner was filled with gravel for some tests with natural gas. Like burner A, 

the height of the rim was 15 cm from the floor and fuel was delivered by an insulated tube 

through the floor. 

A water-cooled downward spray burner (Burner C) was used for liquid fuels in tests #11, #12 

and #15. The nozzle was located 20 cm above the base of a 40 cm diameter round catch pan 

with a 12.5 cm rim. The spray was delivered using a 90 degree full-cone medium atomization 

(droplet diameter ≈ 250 μm) nozzle with a 1.40 mm orifice. The fuel delivery tube was fed 

through a hole in the ceiling and wrapped with Kaowool insulation of approximately 3 cm 

thickness. 

Polystyrene pellets were burned using round pans 22 cm, 40 cm and 60 cm in diameter (Burners 

D, E and F respectively). Each of the burners was centered on the floor. The pan size was 

increased for this fuel in order to reach under-ventilated conditions. A description of the test 

conditions including burner type can be found in Table 1. 
7

Figure 2. Dimensional drawing of burners used in the RSE experiments. Burners D and E were 

similar to Burner F, but the diameters were 22 cm and 40 cm, respectively. 

2.3 Experimental Conditions 

Experiments were conducted during two separate series. The test number (#), series and 

controlled test parameters are listed in Table 1. The two main differences between the series 1 

and series 2 were the wall lining material (see Sec. 2.1.1) and the gas sample conditioning 

systems (see Sec. 2.5.3). 

The fuels included in this test series are listed in Table 1 and included gases, liquids and solids at 

ambient temperature. The composition of natural gas used for these tests is described in Sec. 3.1. 

The heptane fuel was a blend of heptane isomers. The fuel referred to as ethanol was actually a 

blend of 90 % ethanol and 10 % methanol by volume. The polystyrene fuel was clear granulated 

(2.5 g / 100 granules) Dow Styron 666D general purpose resin with a manufacturer reported 

average molecular mass of 230.8 kg/mol. 

Burner A Burner B

Burner C Burner F

30 cm

20 cm

20 cm 

12.5 cm 

15 cm 

7 cm 

13 cm

25 cm

15 cm 

8 cm 

Gravel 

Pool 

Spray 

Fuel Pellets 
8

Table 1. List of test numbers and key experimental conditions. 

Test # Series Fuel Heat Release Rates* 

(kW) 

Door 

Vent Burner Wall Material 

1 1 Natural Gas 75 ,190, 75 Full A Marinite I 

2 1 Natural Gas 255, 395, 180, 115, 50 Full A Marinite I 

3 1 Natural Gas 265, 410, 180, 115, 75 Full A Marinite I 

4 1 Heptane 155, 270, 375 Full B Marinite I 

5 1 Heptane 140, 220 Narrow B Marinite I 

6 1 Natural Gas 75, 175, 270, 420, 80 Narrow B Marinite I 

6.5 2 Natural Gas 95, 425, 270, 180, 85 Full B M board 

7 2 Heptane 150, 245, 340 Full B M-board 

8 2 Methanol 15 Full B M-board 

9 2 Ethanol 20 Full B M-board 

10 2 Toluene 50, 140, 200, 295, 340 Full B M-board 

11 2 Ethanol 80, 145, 265, 335 Full C M-board 

12 2 Methanol 70, 140, 240, 305 Full C M-board 

13 2 Polystyrene 15 Full D M-board 

14 2 Polystyrene 70 Full E M-board 

15 2 Heptane 90, 160, 225, 300, 375, 85 Full C M-board 

16 2 Polystyrene 360, 310 Full F M-board 

* Nominal pseudo steady state heat release rate values from calorimetry measurements 

2.4 Measurement Locations 

Temperature, species volume fraction, soot mass fraction and velocity measurements were 

conducted at various locations in the compartment doorway and interior. A photograph of the 

front gas, soot, and temperature measurement probes is shown in Fig. 3. The sample probe for 

the gravimetric soot measurement is seen on the right of the image, and the aspirated 

thermocouple protrudes down through the ceiling. Figure 4 and Fig. 5 show the relative 

positions (drawn to scale) of the measurement probes in the doorway and inside the enclosure 

respectively. The reference point used to describe the positions within the enclosure is annotated 

in Fig. 5. 

The measurement locations inside the RSE are listed in Table 2 and the locations in the doorway 

are listed in Table 3. The column heading (data label) corresponding to these measurements 

locations are also listed in the measurement location tables. These data labels are referenced in 

the tables and figures in Section 3. A complete list of data column headings can be found in 

Appendix C. 
9

Figure 3. Photograph of extractive sampling probes and aspirated thermocouple at the front 

sample location (photo taken with rear wall removed). 

Soot Probes (2) 

Gas Sample 

Probe (2) 

Sample Aspirated 

Thermocouples (2) 

Doorway Aspirated 

Thermocouples (5) 

Doorway Velocity 

Probes (9) 

Interior Aspirated 

Thermocouples (4) 

Figure 4. Isometric view of Reduced Scale Enclosure showing the relative position of doorway 

measurement probes.

Aspirated thermocouple 

Gas sample probe Soot sample probe 
10

Origin 

(x=y=z=0) 

Burner B 

Front Floor Heat 

Flux Gauge 

x 

y 

z 

Soot Probes (2) 

Gas Sample 

Probe (2) 

Sample Aspirated 

Thermocouples (2) 

Interior Aspirated 

Thermocouples (4) 

Figure 5. Isometric semi-transparent view of the RSE interior measurement probe locations and 

burner. 

Table 2. Location of measurement probes inside of the enclosure.

Probe Description Data Label x (cm) y (cm) z (cm) 

Gas Sample Rear 1 O2Rear 29 113 88 

Gas Sample Front 7 O2Front 29 10 88 

Gravimetric Soot Sample Rear 6 SootRear 29 113 88 

Gravimetric Soot Sample Front 11 SootRear 29 10 88 

Aspirated Thermocouple 15 TRSampA 29 113 88 

Aspirated Thermocouple 16 TFSampA 29 10 88 

Aspirated Thermocouple 17 TR24A 75 122 24 

Aspirated Thermocouple 18 TR80A 75 122 80 

Aspirated Thermocouple 19 TF24A 75 20 24 

Aspirated Thermocouple 20 TF80A 75 20 80 

Total Heat Flux Gauge Rear 12 HFR 48 106 0 

Total Heat Flux Gauge Front 13 HFF 48 35 0 

Bare Bead Thermocouple 21 TFloorR 49 106 0 

Bare Bead Thermocouple 22 TFloorF 49 35 0 

Bare Bead Thermocouple 23 TCeilF 31 11 98 
11

Table 3. Location of measurement probes in the enclosure doorway. 

Probe Description Data Label x (cm) y (cm) z (cm)

Aspirated Thermocouple 31 TC70CA 48 -5 70 

Aspirated Thermocouple 29 TC70LA 32 -5 70 

Aspirated Thermocouple 36 TC50CA 48 -5 50 

Aspirated Thermocouple 39 TC30CA 48 -5 30 

Aspirated Thermocouple 33 TC30LA 32 -5 30 

Bi-Directional Velocity Probe 45 VD79L 32 -5 79 

Bi-Directional Velocity Probe 46 VD79C 48 -5 79 

Bi-Directional Velocity Probe 47 VD79R 64 -5 79 

Bi-Directional Velocity Probe 48 VD60C 48 -5 60 

Bi-Directional Velocity Probe 49 VD40C 48 -5 40 

Bi-Directional Velocity Probe 50 VD20L 32 -5 20 

Bi-Directional Velocity Probe 51 VD20C 48 -5 20 

Bi-Directional Velocity Probe 52 VD20R 64 -5 20 

Bi-Directional Velocity Probe 53 VD5C 48 -5 5 

2.5 Measurement Instrumentation 

 2.5.1 Heat Release Rate 

Heat Release Rate (HRR) measurements were conducted using the 3 m × 3 m calorimeter at the 

NIST Large Fire Research Laboratory (LFRL). The HRR measurement was based on the 

oxygen consumption calorimetry principle first proposed by Huggett [27]. This method assumes 

that a known amount of heat is released for each gram of oxygen consumed by a fire. The 

measurement of exhaust flow velocity and gas volume fractions (O2, CO2 and CO) were used to 

determine the HRR based on the formulation derived by Parker [28]. A detailed description of 

the methodology used for this measurement can be found in a previous report [29]. The 

experimental apparatus for the current measurements has been modified since the earlier report 

was written. In 2005, the 3 m × 3m square hood was installed in the LFRL. A schematic 

drawing of the 3 m hood is shown in Fig. 6. The exhaust flow rate, optical soot and extractive 

gas measurements were performed in a vertical section of the 48.3 cm diameter duct. A bidirectional probe located 9 cm from the edge of the duct was used to measure the exhaust flow 

velocity. Because of the non-uniform shape of the velocity profile, a flow calibration coefficient 

was used in the HRR calculation. The flow coefficient was determined using natural gas 

calibration performed before and after the test series. The flow calibration coefficients ± 2σ for 

these tests ranged from 0.85 ± 0.04 to 0.90 ± 0.05. 

The exhaust gas was sampled through a perforated tube across the duct downstream of the 

velocity probe. Figure 7 shows the exhaust gas sampling system. The main difference between 

this system and the one previously reported [29] is the method for removing water from the gas 

sample. The current system uses a Nafion dryer instead of a dry ice cold trap. Nafion is a 

copolymer of tetrafluoroethylene (Teflon) and perfluoro-3,6-dioxa-4-methyl-7-octene-sulfonic 

acid. A dew point meter was added to monitor the efficiency of the gas dryer. The dew point 

temperature meter measures the change in electrical impedance of a hygroscopic conductive 

polymer in the range of -80 oC to 20 oC. The delay time from the gas sample tube to the 

analyzers was 20 s. Measurements of exhaust soot and total hydrocarbons were performed, but 
12

were not included in the HRR calculation because in most cases they have negligible effect. The 

combined expanded relative uncertainty of the HRR measurements reported here was 14 %, 

based on a propagation of uncertainty analysis [29]. The exhaust mass flow rate was the largest 

component of uncertainty in the HRR measurement. A list of commercial equipment used for 

all of the measurements described in this report can be found in Appendix D. 

 2.5.2 Fuel Metering

Two different fuel delivery systems were used to control and measure the flow rate of fuel to the 

burners. The natural gas tests used a positive displacement flow meter with a standard relative 

uncertainty of 1 % to measure the fuel volume flow rate. Combined with measurements of the 

fuel temperature, pressure, and ideal heat of combustion, the ideal natural gas burner heat release 

rate was determined with a combined expanded uncertainty of 2.4 %. A gas chromatograph was 

used to measure the composition of the natural gas [30]. The net ideal heating value of the 

natural gas was determined using the composition measurements [31]. 

 

The liquid fuel delivery rate was measured using a dual rotor turbine flow meter with a 

manufacturer’s stated uncertainty of 0.1 % in the range from 0.06 L/min to 11 L/min. Although 

the liquid fuel volume flow rate was accurately measured, the fuel mass burning rate was not 

directly measured. In some cases, the amount of fuel (depth of liquid pool) in the burners was 

observed to vary with time, even though the fuel delivery rate was constant. Once the fuel 

delivery was stopped, the burnout of existing fuel could take several minutes. 1.2 m

RSE enclosure

Orifice plate 

(D = 0.305 m)

4.4 m

Flow vanes

Pressure 

transducer

Exhaust stack 

gas sampling 

system

Top-view

Bi-directional 

probe

Thermocouples

Smoke meter 

laser source

Laser 

detector

R = 15.9 cm

Smoke meter

Gas sampling tube

2.34 m

3 m x 3 m hood

27 cm

Thermocouples

Figure 6. Schematic drawing of 3 m hood and exhaust stack instrumented for calorimetry and 

light extinction measurements. 
13

(43 L/min)

Disposable

desiccant

tube

CO2=1.0 L/min

CO=1.0 L/min

O2=0.2 L/min

O2, bypass=2.8 L/min

Excess Sample Flow Enclosure:

Pump, Filter 

and 

Blow Back Air Heated teflon line (93 °C)

Dry purge 

air (or N2) in Wet purge 

air (or N2) out

Nafion tube dryer (2 m)

(5 L/min)

Dry flow 

to analyzers

Beaded

cold trap

(0 °C to 10 °C)

(~ 10 L/min)

3m X 3m hood

Exhaust stack

(D = 0.48 m)

Ball valve

3-way valve

Dew point meter 

Filter

Flow meter

Mass flow controller 

Pressure gauge

Pump

Ball valve

3-way valve

Dew point meter 

Filter

Flow meter

Mass flow controller 

Pressure gauge

Pump

Gas sampling

Calibration gases

N2(100%)

O

CO(0.28%)

2(20.95%)

CO2(2.80%)

CH4(0.0093%)

Calibration gases

N2(100%)

O

CO(0.28%)

2(20.95%)

CO2(2.80%)

CH4(0.0093%)

Figure 7. Exhaust gas sampling system used for heat release rate measurement. 

 2.5.3 Gas Species 

Gas species were continuously measured at two locations (front and rear) inside the RSE during 

each of the tests. Oxygen was measured using paramagnetic analyzers. The 10 % to 90 % 

response time (t10-90) of the oxygen analyzer was less than 12 s. Carbon monoxide and carbon 

dioxide were measured using non-dispersive infrared (NDIR) analyzers. The t10-90 response time 

for the CO2/CO analyzers was less than 5 s. Total hydrocarbons were measured using two flame 

ionization detectors (FID) having a t10-90 response time of less than 1 s.. A gas chromatograph 

(GC) was used intermittently during some of the tests at the front gas sampling location. The 

cycle time on the GC measurements was 20 min to 30 min. The dried sample gas dew point 

temperature was measured using a thin polymer sensor. Soot and temperature were also 

measured at these two locations (see Sec. 2.5.4 and Sec. 2.5.5). 

The two total hydrocarbon analyzers used in these experiments were designed to measure high 

volume fractions of hydrocarbons. The analyzers were factory calibrated for up to 50 % volume 

fraction of hydrocarbons as methane and were capable of measuring even higher concentrations. 

The primary span gas used for these tests was 20 % volume fractions of methane with a balance 

of nitrogen. A span gas of 1 % methane was also used to periodically check the linearity of the 

detector. The FID burner fuel used was 40 % hydrogen and 60 % nitrogen on a volumetric basis. 
14

Each hydrocarbon analyzer had an internal filter to prevent soot from accumulating in the 

plumbing and internal sample pump which could lead to less sensitivity due to hydrocarbon 

contamination and also deterioration of some components of the instrument. It was later 

determined that additional external filtration of soot was necessary to protect the analyzer and 

enable a sufficient time period for sampling soot-laden flows. The external filter could be 

replaced much more frequently and easily than the internal filter. 

 

Two liquid cooled probes were used to sample gas inside the enclosure at the front and rear 

locations. The 1 m long probes were constructed of 3 concentric stainless steel (type 304) tubes. 

Liquid coolant was forced through the inner shell and returned through the outer shell. This 

design allowed the cooling fluid to condition the entire length of the probe. The inner tube was 

lined with glass to reduce catalytic reactions. The inner diameter of the sample probe was 

4.0 mm. Two different gas sample configurations were used during the tests described here. For 

both configurations the front and rear gas sample systems were identical, except the GC 

measurement was conducted only at the front sample location. 

The first configuration (series 1 in Table 1), shown in Fig. 8, used a re-circulated temperature 

controlled bath of 50 % (by volume) ethylene glycol and 50 % water to cool the gas sample 

probes. The sample was drawn through a 3 m stainless steel sample line heated to 120 oC before 

the sample stream was split. Immediately after the heated line, 3 L/min of the sample went 

through a heated filter to the total hydrocarbon analyzer. The bypass stream of the total 

hydrocarbon analyzer was connected to the inlet of the gas chromatograph (not shown in Fig. 8). 

A two-stage water trap and filter was used to remove moisture and soot particles from the sample 

path going to the O2, CO2 and CO analyzers. The sample first passed through a filtered glass 

trap cooled in a wet ice bath and then passed through a beaded glass trap cooled with dry ice. 

The sample pump was located downstream of the drying traps. The gas analyzers were 

connected in parallel so the total flow rate to the rack gas was 2.3 L/min. A 5 way ball valve was 

connected to each analyzer to switch between the gas sample, zero calibration gas and span 

calibration gas. The zero and span gas volume fractions are shown in Fig. 8. A dew point 

transducer was connected to the sample gas line prior to the oxygen analyzer. The oxygen 

analyzer had separate inlet ports for zero and span gases. The expanded (k = 2) relative 

uncertainty of each of the span gas volume fractions was ± 1 %. 

A number of problems were encountered with the first sampling configuration. The coolant bath 

for the gas probes could not maintain a steady probe temperature. A vapor lock was formed in 

one of the probes and the loss of cooling caused the probe to melt and fail. The glass traps 

became clogged with ice after a period of time (creating a loss of sample flow) and there were 

intermittent leaks into the sample line through the trap seals. The internal filter in the total 

hydrocarbon analyzer became clogged with soot during some of the tests. After several tests, the 

gas sampling system was redesigned for improved performance. The redesigned gas sampling 

system (series 2 in Table 1) is shown in Fig. 9. 
15

Beaded

dry ice 

trap

Closed 

coolant 

bath 

(T=90°C)

Total hydrocarbon analyzer

FID, pump and bypass

Heated SS line (3.05 m)

Filtered

wet ice 

trap

Glass-lined 

liquid-cooled probe (1 m)

Gas rack

O2

CO/CO2

Calibration gases

RSE

enclosure

5-way valve

Dew point meter 

Filter

Flow meter

Pressure regulator

Pump

5-way valve

Dew point meter 

Filter

Flow meter

Pressure regulator

Pump

N2(100%)

O

CO(4.0%)

2(18.01%)

CO2(9.0%)

CH4(20%)

N2(100%)

O

CO(4.0%)

2(18.01%)

CO2(9.0%)

CH4(20%)

Vent

Figure 8. Schematic drawing of gas sampling system for series 1 tests #1 to test #6. 

Heated SS line (3.05 m)

Glass-lined 

liquid-cooled probe (1 m)

Gas rack

RSE

enclosure

Sample conditioning system Blow back air

CO2

O2

CO

Vent from FID

to GC

Total Hydrocarbon Analyzer

(1 L/min)

(3 L/min)

Calibration gases

Water 

(T = 55°C)

Drain Wet sample

Wet purge 

gas

(15 L/min)

Membrane 

dryer

Dry purge 

gas (N2)

Electric temperature 

controller (T = 60°C) Heated 

enclosure

Heater Dry sample 

to gas rack

Ball valve

3-way valve

5-way valve

Dew point meter 

Filter

Filter

Flow meter

Mass flow controller 

Pump

Ball valve

3-way valve

5-way valve

Dew point meter 

Filter

Filter

Flow meter

Mass flow controller 

Pump

Vent

Vent

FID

Bypass

N2(100%)

O

CO(4.0%)

2(18.01%)

CO2(9.0%)

CH4(20%)

N2(100%)

O

CO(4.0%)

2(18.01%)

CO2(9.0%)

CH4(20%)

N2(100%)

O

CO(4.0%)

2(18.01%)

CO2(9.0%)

CH4(20%)

To total 

hydrocarbon

analyzer

Metering

valve

Figure 9. Schematic drawing of gas sampling system for series 2 (test #6.5 to test #16). 
16

The sample probes shown in Figs. 3-5, 8, and 9 were cooled using house water heated to 55 oC at 

a flow rate of 1 L/min. The total hydrocarbon analyzers were placed in the gas racks with the 

other analyzers. The cold traps were replaced with a membrane type dryer. A bundle of Nafion 

tubes were purged with dry nitrogen to selectively remove moisture from the sample stream. The 

Nafion conditioner has no effect on most of the gas species of interest, however, polar organic 

compounds (i.e. ketones and alcohols) are trapped by the dryer. A large area filter was added 

between the heated line and gas dryer. Because the external filters and transfer lines after the gas 

dryer were not heated, there was a potential loss of high molecular mass hydrocarbons due to 

condensation. Due to limitations in the flow capacity of the dryer, the gas analyzers were 

connected in series. A mass flow controller set to 1 L/min was used to control the flow through 

the O2/CO2/CO analyzers. The flow to the hydrocarbon analyzer was split prior to the mass flow 

controller. A needle valve was used to set the total flow to 3 L/min (only a small fraction of this 

passed through the FID). The bypass flow from the hydrocarbon analyzer was connected to the 

injection port of the GC. 

 2.5.4 Extractive Soot Measurement 

A gravimetric sampling system (shown in Fig. 10) was used to measure soot mass fractions at 

the two sample locations within the enclosure. The design of the soot probe was similar to the 

gas sampling probes except the inner diameter of the sample tube was 6.4 mm. The soot 

sampling probes were conditioned with 65 oC water flowing at 1.0 L/min. A three way solenoid 

valve was used to rapidly switch from the bypass to sample flow. A sample gas mass flow rate of 

2.75 standard L/min (N2 @ 0 oC, 101.3 kPa) was drawn through the collection filter for a period 

of 60 s to 300 s. The collection filter was a 47 mm round Zeflour membrane filter with an aerosol 

retention efficiency of 99.99 % for 2 μm sized particles. A gas correction factor was applied to 

the mass flow rate measurement to account for the gas composition in the enclosure. The amount 

of time for sampling was determined by monitoring the pressure drop across the filter to ensure 

an optimal amount of filter loading. 

The collection filters (shown at the base of the probes in Fig. 10 below) and probe cleaning pads 

were conditioned in a desiccant drier before and after the tests. The conditioned filters were 

weighed using an analytic mass balance with an expanded uncertainty of 0.12 mg. After each 

soot sampling period, the probe was cleaned twice with gun cleaning pads. The total soot mass 

collected on both the filter and 2 cleaning pads was used in determining the soot mass fraction. 

Both the soot mass and sample mass flow rates were measured on a dry basis. For most of the 

tests conducted in this series between 10 mg and 200 mg of soot was collected during the 1 min 

to 5 min sample time. The extracted gas volume was corrected for the water removed by the 

method described in Sec. 2.7. The combined expanded relative uncertainty of the soot mass 

fraction measurement (for mass fraction measurements greater than 0.001 g/g) was in the range 

of 2 % to 5 % based on a propagation of uncertainty analysis. 
17

RSE

enclosure

Temperature 

Controlled 

Water bath 

(T=65 °C)

Water-cooled

sample probe

HEPA filter

(T=70 °C)

Bypass

intake

Standby probe

vent

LFL rack

( 2.75 L/min N2)

Beaded 

dry ice trap

Coiled wet 

ice trap

3-way valve

Filter

Mass flow controller 

Pressure gauge

Pump

Figure 10. Schematic drawing of soot sampling system. 

 2.5.5 Temperatures

Aspirated thermocouples 

A bare-bead thermocouple situated in a compartment fire typically experiences radiative 

exchange with walls, hot smoke, flames, and the surrounding environment with the effect that 

the measured temperature is not the true gas temperature. Accurate correction for these effects is 

complex, due to temporally and spatially varying local temperatures, velocities, and species. To 

reduce the effect of energy exchange on temperature measurement accuracy, aspirated 

thermocouple probes were used in addition to bare-bead thermocouples in this study. 

An aspirated thermocouple probe is a bare-bead thermocouple contained within a small 

cylindrical metal tube through which the sample gas flows. If the flow over the bead is at least 

5 m/s, a more accurate gas temperature measurement may be obtained [32]. According to 

Blevins [33], higher flows may be required depending on the thermal environment. Aspirated 

thermocouple probes may be shielded by a single cylindrical tube or by two or more concentric 

cylindrical tubes. In either case, the flow and thermal conditions and the detailed design of the 

assembly can impact measurement accuracy. Double-shielded aspirated thermocouple probes 

based on a design from National Advisory Committee for Aeronautics (NACA) were used in this 

study [34]. Figure 11 shows a drawing of an end-hole type NACA design aspirated 

thermocouple probe. Models with the entrance hole perpendicular to the probe axis were also 

used. 

Each aspirated thermocouple was connected to a set of wet-ice and dry-ice traps, a flowmeter, 

and a pump using 9.5 mm (3/8 in) outer diameter (OD) copper and polyethylene tubing. A 

schematic of this is shown in Fig. 12. The gas was filtered and dried with the traps to protect the 

flowmeters and pumps. Flows were set at 24 L/min for each aspirated probe. While the 

volumetric flows were set at the flow meters to be the same for all probes, since the ice traps 
18

cooled the hot gases, high temperature compartment gases produced much higher velocities at 

the bead compared to those produced by low temperature gases. The uniform setting of the cold 

volumetric flows kept the mass flows consistent across the probes. This velocity difference 

effect was not completely proportionate to the gas temperature differences since a higher flow 

would experience a greater pressure drop and flow resistance through the probe and tubing. Due 

to the large flows pumped through the aspirated thermocouple probes, the resulting temperature 

represents a volumetric average over a several centimeter diameter region at the end of the probe. 

For further discussion of the probe and gas interaction see Appendix B of this report. 

Figure 11. Detailed drawing of aspirated thermocouple using NACA design [34]. 
19

Each aspirated thermocouple probe was attached to the data acquisition system using K-type 

thermocouple wire and connectors. During each experiment, the flow meters and measured 

temperatures were monitored. These checks were performed in order to determine if any probe 

system became clogged so it could be unclogged with high pressure air. The difference in 

temperature signal between an inoperative probe and a properly flowing probe was obvious. A 

functioning aspirated thermocouple showed higher frequency temperature fluctuations due to the 

transient thermal environment and effective convective heat transfer while a non-functioning 

probe would not show rapidly fluctuating temperatures since the large mass of hot metal of the 

probe radiating to the bead and lack of convection would dampen any short fluctuations. A 

probe typically required about 1 min when activated to overcome accumulated heat and reach the 

true gas temperature. 

To evaluate measurement uncertainty and instrument time response, the present study performed 

a series of detailed flow and heat transfer calculations, focusing on double-shielded aspirated 

thermocouples and bare-bead thermocouples. A detailed description of the calculations and 

results can be found in Appendix A. 

Aspirated thermocouple

(end hole type)

Aspirated thermocouple 

(side hole type)

Data acquisition 

system

Filtered dry 

Ice trap

Coiled wet 

Ice Trap

Vent

Outer shield

Inner shield

TC bead

TC

Flow 

meter

Pump

Figure 12. Schematic drawing of aspirated thermocouple measurement hardware. 

Bare Bead Thermocouples 

Temperature measurements by thermocouples located on compartment surfaces are required to 

characterize the thermal environment created by the fire, as well as to provide information for the 

aspirated thermocouple uncertainty analysis. Surface temperature measurements on the external 

surfaces of the compartment are needed to check the overall enthalpy balance associated with the 
20

fire, which is important for validating predictive compartment fire models. Bare bead 

thermocouples were also used in the doorway adjacent to the bi-directional probes and aspirated 

thermocouples. 

 

The bare bead thermocouples were created by removing 1 cm to 2 cm of silica ceramic yarn 

insulation from chromel and alumel (type K) lead wires and spot welding them together. The 

thermocouple lead wire diameter was 0.51 mm (24 gauge). The mean bead diameter was 

approximately two times the wire diameter. 

Bare bead thermocouples were placed on the inside surface of the compartment at three 

locations: two on the floor and one on the ceiling. Table 2 in Section 2.4 lists the exact locations 

of the thermocouples, which were positioned on the floor, adjacent to the total heat flux gauges, 

and on the ceiling, almost directly above the front measurement station (which included the 

gravimetric soot and aspirated thermocouple probes). Figure 13 shows a 30 gauge Type K 

thermocouple on the compartment floor surface, which was held in place by spring loading to 

maintain its position near a 6 mm diameter total heat flux gauge. The screw/washer assembly 

about the thermocouple wire ensured that the thermocouple would not move. 

Bare bead thermocouples were also positioned on the external surface of the compartment at one 

to five locations on the rear wall, depending on the experiment. Figure 14 shows a type K 

thermocouple attached to the rear outer surface of the compartment and held in place by a 

washer/screw arrangement (with the thermocouple spring loaded to maintain its position). The 

expanded uncertainty associated with a type K thermocouple is approximately 4.4 °C [35]. 

Figure 13. Type K thermocouple on compartment floor surface near a 6 mm diameter total heat 

flux gauge. 
21

Figure 14. Type K thermocouple attached to the outer surface of the compartment. 

 2.5.6 Pressures and Velocities

Dynamic pressure was measured at 9 locations in the doorway of the enclosure in order to 

determine velocities in the doorway. The coordinates of the locations are contained in Table 3 

and shown in Fig. 4. The differential pressure transducers (0 V to 10 V output) had a maximum 

range of 133 Pa. The particular model information is contained in Appendix D. Each pressure 

transducer was mounted on a board, and the board was attached to one of the support legs of the 

exhaust hood. Insulating board was also used to shield the transducers from the thermal insult of 

the fire. Each transducer produced a voltage, Vbdp, related to the exposed differential pressure by 

the following equation: 13.332( ) ΔPbdp = Vbdp −Vbdp,zero , where the pressure difference is in pascals 

and the voltages are measured in volts. The zero voltage, Vbdp,zero,, condition is created when the 

positive and negative ports of the transducer are connected so there can be no pressure difference 

between them. 

The transducers were connected to 1.3 cm diameter bi-directional probes [36] with 6.4 mm 

diameter copper tubing. Probe leads were routed close to each other so each lead was exposed to 

the same levels of heating. This installation care minimized differential heating and any 

resulting non-flow induced pressure differences between the leads. 

Bi-directional probes enable the measurement of dynamic pressure which is the difference 

between the total pressure on the face where flow impinges and the static pressure on the 

downstream face of the probe. Using Bernoulli’s principle and including a calibration factor, 

velocity, v, can be obtained from the dynamic pressure and a local gas temperature through the 

following relation: C PbdpTbdp v = Δ where ΔPbdp is the measured pressure across the 

bidirectional probe, Tbdp is the temperature (in K) of the gases flowing past the probe and C is 

defined as:

bdp Pref MWgas

R

C

C 1 2 = . 
22

The calibration coefficient, Cbdp, for a bi-directional probe is equal to 1.08 ± 0.05 [36] when the 

local Reynolds number (defined by the probe diameter) is greater than 1000. R is the ideal gas 

constant and MWgas is the molecular mass of the gas. 

To generate the velocity from the differential pressure, the temperature near the bi-directional 

probe is required. Because aspirated thermocouples can intrude on the pressure measurement 

bare-bead thermocouples were used to measure temperature. See Appendix A for a discussion of 

the errors associated with this approach. The measured doorway velocities were in the range of 

-7 m/s (flow out of the enclosure) to +1.5 m/s (into the enclosure). The combined expanded 

(k = 2) uncertainty in the velocity measurement varied from ± 0.5 m/s to ± 2.3 m/s. The largest 

component of uncertainty in this measurement was the variation in the pressure signal. 

 2.5.7 Gas Chromatography

A gas chromatograph was used at discrete times during the RSE tests to identify and quantify the 

major hydrocarbon species for each fuel and fire size at the front gas sampling location. The 

majority of the stable intermediate species were identified and quantified with a Hewlett-Packard 

5890 gas chromatograph (GC) with flame ionization detector (FID). For chromatographic 

separation, a Restek Rt-QPLOT column (30 meter, 0.32 mm ID) was installed in the HP5890. 

Identification of the unknown species was accomplished by retention time matching. The 

quantification of the identified compounds was accomplished using the FID. The settings for the 

HP5890 are specified in Table 4. 

In order to identify and quantify the hydrocarbons, gas phase calibration standards were obtained 

from Scott Specialty Gas. These standards had a reported uncertainty of ±10 % of desired 

volume fractions. Several other sources of gas standards were available for only retention time 

matching of unknown compounds. Table 5 documents the retention times for the specific 

identified compounds. 

To quantify the identified compounds, either the calibration curve for the specific identified 

molecule or a calibration curve of a similar molecule was employed. If the molecular mass of 

the identified molecule and calibration curve molecule were different, then a correction factor 

based on the calibration molecule’s carbon number divided by the identified molecule’s carbon 

number was used. Employing this technique with two known compounds, the relative 

uncertainty associated with this correction was approximately 1 % for one carbon atom 

difference, and approximately 3 % for two carbon atom difference. As a result, all correction 

factors for unknown compounds were generally limited to a 1 or 2 carbon number difference. 
23

Table 4. Gas chromatograph settings and method parameters. 

GC Parameters Setting 

Inlet Parameters

 Inlet Mode Constant Pressure Rate 

 Inlet Pressure 124 kPa 

 Inlet Temperature 250 °C 

 Column Flow 2.3 mL/min 

 Split Ratio 10:1 

 Septum Purge 1.0 mL/min 

Oven Temperature Parameters 

 Initial Temperature 50 °C 

 Initial Time 1 min 

 Ramp #1 Rate 15 °C/min 

 Ramp #1 Temperature 80 °C 

 Ramp #1 Hold time 1 min 

 Ramp #2 Rate 20 °C/min 

 Ramp #2 Temperature 240 °C 

 Ramp #2 Hold time 5 min 

Auxiliary Parameters 

 Valve Oven Temperature 125 °C 

 FID Temperature 250 °C 

 

Valve Timing Parameters 

 0.10 min – Valve #A ON (Sample Injection) 

 1.10 min – Valve #A OFF 

Table 5. List of retention times for identified compounds. 

Species 

Elution 

Time 

(min) 

Species 

Elution 

Time 

(min) 

methane 2.41 n-butane 9.27 

ethene 3.60 1-butyne 9.34 

ethyne 3.62 2-butyne 10.22 

ethane 4.07 1-pentene 11.00 

propene 6.58 n-pentane 11.17 

propane 6.80 1-hexene 12.79 

propyne 7.12 n-hexane 12.93 

1-butane 9.05 benzene 13.74 

1,3-butadiene 9.18 
24

Analysis of variance (ANOVA) and regression analysis were employed in an effort to determine 

the errors associated with the quantification of the intermediate combustion species. The result 

of this analysis was the uncertainty of a single value, Sy, calculated from the calibration curve. 

The equations utilized for the analysis are show below: 

( )

xx

i regression

S

y y

m N

s

y s 2

2

1 1

β β

− = + + (1) 

where, 

( ) ( )

( )

2

2

−

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

−

= N

S

S

S

regression

xx

xy

yy

s (2) 

( )

N

y

s y yy

2

2 ∑ = ∑ + (3) 

( )

N

x y

s xy xy

∑ ∑ = ∑ + (4) 

( )

N

x

s x xx

2

2 ∑ = ∑ + (5) 

m = number of measurements of the unknown sample 

N = number of calibration curve points (typically, 3) 

y = FID area count of calibration species 

x = volume fractions of calibration species 

yi = FID area count of unknown species 

β = slope of linear least squares fit to calibration points 

The results of this analysis are included in all of the quantification graphs for each of the fuels 

examined in this program. However, it is important to note that this analysis does not include 

other sources of errors, such as sample extraction errors, possible decay of the samples prior to 

analysis, or errors resulting from carbon number correction factors. 

 2.5.8 Optical Extinction Soot Measurements 

Two optical configurations were used to measure soot using laser transmission. They are shown 

schematically in Figure 15 and Figure 16, and the optical components used in both 

configurations are listed in Table 6. Results are available for Tests 6, 13, and 14 only, mainly 

because the measurement proved to be very challenging. The method and instrumentation used 

during Test 6 differed from that of Tests 13 and 14, as the experimental technique was being 

optimized during the course of this study. Table 7 lists the fuel type used during the tests. The 

apparatus used in Test #6 (shown in Fig. 15) consisted of a laser light source, a chopper, a beam 

splitter, purged tubes, optical lenses and two detectors. A chopper and lock-in amplifier were 
25

used to minimize the background signal and amplify the original signal. The reference signal of 

the original laser source was monitored by splitting the laser beam. 

Figure 16 shows the apparatus used during tests #13 and tests #14. The laser light source was a 

9.7 mW 657 nm diode laser (continuous). A spherical lens with a 2.5 cm diameter and 40 cm 

focal length reduced the beam diameter. The optical path-length was defined by two thin (1 mm 

wall thickness) 2.5 cm diameter stainless steel tubes that were aligned and separated by a gap of 

distance LD. Each of these tubes was connected to commercially available “stackable lens tubes” 

that were used to align the laser and detector and to shield the detector from stray light. Low 

flowing nitrogen gas was used to prevent combustion products, including soot, from entering the 

tubes. A glass window was attached on the end of the stackable lens tubes to prevent backward 

flow of the nitrogen purge gas. The detector was a silicon photodiode sensitive from 350 nm to 

1100 nm. On the detector side of the optical assembly, a spherical lens (2.5 cm diameter and 1 m 

focal length) was positioned in the middle of stackable lens tubes to focus the laser light on the 

detector and prevent backward flow of the nitrogen purge gas. Several optical components were 

placed in front of the detector. A band pass filter (±10 nm window centered about 650 nm) acted 

to attenuate radiation other than that from the laser. A diffuser lens was used to expand the laser 

beam, a standard practice that reduces beam steering effects. To prevent detector saturation, a 

neutral density filter was used to attenuate the beam. 

The position of the optical apparatus used in tests #13 and #14 was different from that used in 

Test 6. The sampling position was located 1 cm below the doorway soffit (80 cm above the 

floor) and 1.4 cm from the plane defining the doorway as shown in Fig. 16. The nominal 

pathlength (LD) (the distance between the stainless steel purging tubes) in the optical soot 

measurement was 11.5 cm. The nitrogen purge flow rate was adjusted to keep combustion 

products and smoke out of the purge tubes. The purge was measured as 900 (±30) cm3

/min, 

equivalent to an average speed of 4 cm/s through each of the tubes. The purge flow was 

estimated to reduce the actual path-length of the optical measurement (L) by approximately 

1.0 cm. This value was observed to vary somewhat with fire size, such that the estimated 

expanded uncertainty of LD was about 1 cm, corresponding to a relative expanded uncertainty of 

9 %. 

The determination of soot density (ms) from an optical extinction measurement is based on 

Bouguer’s law, in which the light extinction coefficient (K) is defined in terms of the attenuated 

intensity (I), and the reference intensity (Io) of monochromatic light passing through a 

homogeneous smoke path of distance (L ): 

L

I I K o ln( / ) = − (6) 

S

s

K

m

σ = (7) 

where σs is the mass specific extinction coefficient. The recommended value of σs for flame 

generated smoke in over-ventilated fires is 8.7 m2

/g ± 5.4 % (standard relative uncertainty). The 

expanded uncertainty of the ln(Io/I) term during a period of 6 min before the fire test was on the 

order of 0.2 %. The laser drift, which was about 5 %, dominated the uncertainty in the baseline. 
26

The expanded combined uncertainty of the soot volume fraction is computed as the square root 

of the sum of the individual standard uncertainties [u(xi)] associated with each of the terms that 

influence the soot mass measurement. The combined uncertainty of the optical soot density 

measurement is calculated as follows: 

( ) 2

2

2

2

2

2

ln( / ( ) ( ) ln( / )2 s

s

s s

o

o

s

m u

m

u L

L

m

u I I

I I

m U s σ

σ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

∂

∂ ⎟ + ⎠

⎞ ⎜

⎝

⎛

∂

∂

+ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

∂

∂ = (8) 

An estimate of the combined relative expanded uncertainty was about 20 %, for both 

experimental configurations. 

 Table 6. Components used in the optical extinction measurements. 

Component Test 6 Tests 13 and 14 

Spherical lens before detector Present Present 

Spherical lens after laser Not Present Present 

Beam splitter Present Not Present 

Neutral density filter Not Present Present 

Reference detector Present Not Present 

Aperture in the detector part Not Present Present 

Lock-in amplifier Present Not Present 

Probe separation distance (LD) >> 24 cm (≈ 6 m) 11.5 cm ± 0.1 cm 

Probe location 

5 cm below door soffit; 

Across entire doorway; 

1.4 cm beyond compartment 

1 cm below door soffit; 

about center of doorway; 

1.4 cm beyond compartment 

Table 7. Tests in which optical light extinction measurements were conducted.

Test 1 Fuel type Burner 2 Fire size, Qmax (kW) Path-length, L (cm) 

6 Natural Gas B (25 cm square) 75 to 400 24.0 ± 1.5 

13 Polystyrene D (20 cm round) 18 10.5 ± 1.0 

14 Polystyrene E (40 cm round) 80 10.5 ± 1.0 

1. measurements not made in other tests, because the fire adversely affected the detection system. 

2. see Fig. 2.
27

Figure 15. Schematic of the optical extinction measurement apparatus used during test #6. 

Figure 16. Schematic of the optical extinction measurement apparatus used during test #13 and 

test #14 including a blow-up of the circled optical sampling region. 
28

 2.5.9 Heat Fluxes 

Total heat flux was measured at two locations during each experiment. The heat flux gauges 

were 6.4 mm diameter Schmidt-Boelter type, water cooled gauges with embedded type-K 

thermocouples. The particular model information is contained in Appendix D. The nominal 

range for the gauges was 150 kW/m2

. Schmidt-Boelter gauges measure a temperature difference 

across a thin insulating material using a thermopile to generate a voltage from the small 

temperature difference. These gauges typically generate voltages much less than 100 mV even 

for heat fluxes near their maximum range. 

Each gauge was inserted in the floor flush with the upper surface and facing vertically upward. 

The rear gauge was located on the centerline of the enclosure and approximately 3⁄4 of the way 

toward the rear from the front. The front gauge was located on the centerline of the enclosure 

and approximately 1⁄4 of the way toward the rear from the front. The exact location coordinates 

for the gauges are listed in Table 2. The condition of the installed gauges was checked 

periodically. If significant soot accumulated on a gauge, it was brushed off. If a gauge was no 

longer flush with the surface of the floor, a note was made, but there was no attempt to move the 

gauge since the gauges were very difficult to access and attempting to do so could have impacted 

the integrity of the floor. 

Heat fluxes as high as 250 kW/m2

 were observed. These heat fluxes are beyond the stated range 

of the gauges. According to the manufacturer, the calibrations remain linear and valid beyond 

the stated range as long as the materials do not degrade and change the sensitivity of the gauge. 

After the first six experiments the heat flux gauges were checked for changes to their 

calibrations. Each gauge’s responsivity was found to remain within 3 % of the factory 

calibration. 

The main sources of uncertainty related to the total heat flux measurements are: the calibration, 

soot and dust deposition, and shifting of the gauge surface below the floor. These sources will 

be described and the total uncertainty estimated for the reported measurements. A model of 

uncertainty for heat flux gauge measurements in fire environments can be found in the study by 

Bryant et al. [37]. 

The total heat flux gauge calibration from the manufacturer was used to convert millivolt 

readings to kW/m2

. This calibration was performed using cooling water at 23 °C ± 3 °C. The 

cooling water in the Large Fire Laboratory was found to be within the same range. The 

manufacturer reported a ±3 % expanded uncertainty in the responsivity (the slope in 

kW/m2

/mV). Calibrations at the NIST facility have varied within the 3 % range of the nominal 

manufacturer’s calibration. A recent round-robin study of heat flux gauge calibration 

consistency [38] sent the same heat flux gauges to multiple laboratories around the world and 

found that while several calibrations fell within the 3 % range, if some outlier data were 

included, then the uncertainty rose to around 8 %. For this current project, an uncertainty of 

±6 % for gauge calibration was chosen as fairly conservative since the NIST calibration was 

within the 3 % range in the round-robin study. 

While the cooling water was supplied at approximately 23 °C, the fire heated the water such that 

the gauge temperature typically rose to between 40 °C and 60 °C, and less frequently to 100 °C. 
29

For the fires where the water temperatures increased to between 40 °C and 100 °C, the heat 

fluxes were on the order of 100 kW/m2

 to 300 kW/m2 which represent blackbody temperatures in 

the 950 °C to 1300 °C range. The most extreme combination (affecting uncertainty) of cooling 

water and environment temperature would be a 75 °C increase in cooling water in a 950 °C 

environment. This combination would only have about a 0.5 % effect on the measured heat flux. 

The effect was determined by calculating the ratio of the T4

 difference between 950 °C and the 

25 °C cooling water with 950 °C and the 100 °C cooling water. This is a simplified comparison 

which assumes everything else is equal, but generates an approximation of the magnitude of the 

cooling water effect under specified conditions. 

Heat flux uncertainty due to soot and dust deposition is difficult to quantify. For many tests, 

such as those burning methanol, ethanol, and natural gas, there was little to no contact with soot 

or combustion products. Also, even for the sootier fuels at low heat release rates, the lower layer 

remained as air with little opportunity for soot-laden gases to contact the gauges. For those 

experiments with sooty fuels and underventilated conditions (>200 kW HRR), combustion 

products including soot sometimes impinged on the floor. For these periods of time, it was 

estimated that the soot coating on the gauge would add an additional uncertainty of ±10 % due to 

variations in surface emissivity, and soot agglomerates shadowing the surface of the gauge. 

The physical shifting of the gauge surface below the floor could have impact on a heat flux 

measurement if the solid angle viewable by the gauge was significantly diminished. Since the 

gauge is not sensitive either in calibration or application to radiation at angles close to the plane 

of the gauge surface due to reflection, and the radiation approaching from the lowest angles is 

generally from the coolest regions of the enclosure, the gauge would have to be below the 

surface of the floor by a few millimeters or more for there to be a significant impact on its 

measurement. Neither gauge was ever observed to be shifted by that amount in the course of 

testing. 

 2.6 Data Acquisition 

Data acquisition (DAQ) for this series of experiments was divided into two systems. One DAQ 

system was dedicated to fuel flows, oxygen depletion calorimetry, and the constituent 

measurements required to calculate heat release rate using that method. The other DAQ system 

was used to record signals from all other measurements. Each DAQ system used National 

Instruments hardware and was controlled with LabVIEW software. The calorimetry DAQ 

system has been previously described in detail [29]. 

For this series of experiments, the channel list contained in Appendix C was used to program the 

DAQ system. The types of measurements included: gas analyzers, dew point readers, heat flux 

gauges, pressure transducers, and thermocouples. These measurements were recorded on the 

DAQ hardware as voltages with 200 samples recorded every second. Each second, the average 

value for each channel was then converted to meaningful physical units. Two event marking 

channels were used to note the time of important events such as ignition, fuel flow change, or 

extinguishment. These event marker channels, which are in both DAQ programs, were 

especially useful in synchronization of the two data sets. 
30

The DAQ system for measurements not related to heat release rate, called MIDAS (Modular Insitu Data Acquisition System), had a structure and hardware components that differ from the 

HRR/fuel DAQ system. The MIDAS system utilized a fiberoptic extension of the computer’s 

PCI bus. A series of fiberoptic cables connected the main computer in the control room to three 

experiment stations in the main bay of the laboratory. Each station has its own DAQ card, 

multiplexing hardware, and terminal blocks for voltage or thermocouple inputs. The East 

MIDAS Station was the station used for this series of experiments. 

There were four comma-delimited spreadsheet files produced for each experiment. One file 

(with the -raw suffix) contained all of the raw voltages and temperatures recorded. A second file 

(-adj suffix) contained values with converted units calculated from the raw voltages. A third file 

(-ZS suffix) contained calibration data for each instrument that is calibrated at the beginning of 

an experiment. Finally, a fourth file (-sd) contained standard deviations for selected instruments 

based on the 200 raw voltages averaged each second. 

The data acquisition hardware had 16 bit precision, with stated accuracies of the data acquisition 

board and multiplexing module equal to 0.014 % and 0.015 % of the reading. These 

uncertainties were orders of magnitude lower than those from other sources in all of the 

measurements reported here. 

 2.7 General Data Corrections

A Matlab script file was created for post-processing all data files generated during the test series. 

This program was used to make corrections to the data, generate plots, and save results to ASCII 

text files for archival purposes. The program was also used to compute time averaged values and 

uncertainties for examining trends in the data. An input file was used to allow batch processing 

of the raw data files. The input file contained the parameters needed for the heat release 

calculation (this file was also read by the DAQ program during the data collection process). 

Additional parameters were added to the end of the standard HRR input file to account for the 

gravimetric soot measurements and to record the time windows when channels had known 

missing or corrupted data. 

The first step in data reduction was to inspect the data files and lab notebooks for erroneous data 

resulting from open channels, loss of sample flow, or some other instrument or data acquisition 

malfunction. Because data were collected on two separate computers, the series were 

synchronized to a common reference time. The ignition time was marked using a virtual event 

channel on each computer and defined as time zero for the reduced data. The gas analyzer 

measurements from inside the RSE and exhaust hood measurements were shifted in time to 

account for the sample flow transfer (delay) time. There was no adjustment for instrument 

response time in the data reduction. . The t10-90 response time of the instruments used in this 

study varied from 1 s to 12 s. 

Corrections to the heat release rate measurements were applied to account for the exhaust flow 

calibration factor and drift in the oxygen analyzer. The exhaust flow rate data was smoothed 

over a 10 s window to reduce noise due to turbulent flow in the duct. This smoothing was of the 

same order as the response time of the exhaust gas sample oxygen measurement. 
31

Since the gases sampled from the RSE were dried before entering the detectors, an estimate of 

the water removed must be made in order to correct the measurements to the in situ wet volume 

fraction. The general combustion reaction assuming all the fuel is reacted and that the soot can 

be represented as pure carbon is: 

CxHyOz + aO2 → bCO2 + cCO + dCH4 + eC + fH2O (9) 

The molecular yield of water can be related to the combustion product yields using the known 

hydrogen/carbon (y/x) ratio of the fuel: 

( ) b c d e 2d

2x

y f = + + + − (10) 

If the yield of soot is small compared to the other products, the water volume fraction, XH2O, can 

be estimated from Eq.11. 

 

( ) H O XCO ,wet XCO,wet 2x

y X 2 2 = + (11) 

The relationships for wet CO2 and CO are given by the following: 

( ) CO ,dry CO,dry

CO,dry

CO,wet

X X

2x

y 1

X

X

2 + +

= (12) 

( ) CO ,dry CO,dry

CO ,dry

CO ,wet

X X

2x

y 1

X

X

2

2

2

+ +

= (13) 

Other gas volume fraction measurements performed on a dry basis were corrected using the 

following relationship: 

( ) Xspec,wet = Xspec,dry 1− XH2O (14) 

The total hydrocarbons can contribute to the formation of water, however the gas composition 

measurements confirmed that when total hydrocarbons were present in significant quantities, 

they were in the form of unburned fuel (methane in the natural gas tests). Unburned fuel does not 

contribute to the formation of water. Therefore, the resulting relative error in the water volume 

fraction estimation due to neglecting hydrocarbons was always less than 3 %. The error in the 

water volume fraction estimate due to neglecting soot was as much as 10 % for the highly 

sooting fuels. However, since the soot measurements were sparse, we chose to report the results 

on a consistent basis. A more accurate estimate of water volume fraction could be made for the 

short time windows where soot was collected. Hydrogen gas was not quantified for these tests, 

but could also affect the estimation of water. 
32

Unless otherwise noted, all uncertainty results reported here represent the combined expanded 

(coverage factor k = 2) uncertainty resulting from a propagation of uncertainty analysis. The 

uncertainty values are represented by error bars on the steady state average values presented in 

Sec. 3. 

3 Results 

 3.1 Heat Release Rate 

The heat release rate measurement was used to characterize the size of the fire and also to help 

determine (along with heat flux data) when the fire conditions had reached steady state. As the 

fire becomes underventilated burning can take place outside of the enclosure. The HRR 

measurement represents the total burning inside and outside of the enclosure. Table 8 shows a 

description of the measurement labels used in the table column headings and figure legends in 

this section. These labels are identical to the column headings in the reduced data files. 

Table 8. Description of calorimetry measurement labels. 

Measurement Label Description 

54 HRRcal Heat Release Rate from Calorimeter, kW

55 HRRburner Heat Release Rate from Burner (gas, pool or spray), kW

56 StackMFR Exhaust hood mass flow rate, kg/s

57 Tstack Exhaust hood temperature (near bi-directional probe), °C

58 O2stack Exhaust O2 volume fraction (dry)

59 CO2stack Exhaust CO2 volume fraction (dry)

60 COstack Exhaust CO volume fraction (dry)

61 THCstack Exhaust total hydrocarbons volume fraction (dry)

62 MSstack Exhaust soot mass concentration (wet), mg/m3

Figure 17 shows the heat release rate results for one of the natural gas experiments (test #3). For 

this test, the measured HRR and enthalpy input match closely indicating that the combustion 

efficiency was close to 1. The composition and heating values for the natural gas used during all 

of the tests is shown in Table 9. The heating value is defined at the standard conditions of 300 K 

and 101.3 kPa. The propane and nitrogen levels were unusually high on the first test day; 

however the variation in heating value was less than 1.5 % for all of the tests with natural gas. 

Figure 18 shows two photographs of test #3 (natural gas) looking into the doorway at the 

nominal fire sizes of 75 kW (left) and 400 kW (right). The intensity and transparency of the 

flames provided visual evidence of the lack of soot produced by the fire. The image of the 

400 kW fire clearly shows flames exiting the doorway, indicating the fire is underventilated. The 

75 kW fire in the reduced-scale enclosure would scale up to a 485 kW fire in the ISO 9705 room, 

and the 400 kW fire would correspond to a 2.6 MW fire in the ISO room (see Sec. 6). 

The heat release results for the methanol spray fire (burner C) test #12 are shown in Fig. 19. The 

dashed line in this figure represents the ideal heat release rate of the fuel based on the delivered 

fuel flow rate. The liquid fuel height in the catch pan varied during the test, and once the fuel 

flow was stopped the burnout time was more than 20 min (see Fig. 19 from 3400 s to 4900 s). 

This accumulation effect explains why the heat release rate of the fire measured by calorimetry 
33

was significantly different than the set burner heat release rate. Although the HRR measurement 

had a larger uncertainty than the fuel flow rate measurement, in some cases it was a more reliable 

measurement of the actual fuel burning rate. Photographic images of three different methanol 

fires are shown in Fig. 20. The image on the left shows a 50 kW methanol pool fire burning 

outside of the enclosure in a 40 cm diameter pan prior to test #12. The middle and right images 

show methanol spray fires from test #12 at HRR’s of 70 kW and 300 kW, respectively. 

Table 9. Composition (volume fraction %) and heating values of natural gas used in RSE tests. 

Test 

Name 

Methane Ethane Propane i-Butane n-Butane i-Pentane n-Pentane C6+ Nitrogen CO2 MWng (g/mol) 

HOCng (MJ/m^3) 

RSE1NG 86.92 3.84 3.57 0.13 0.15 0.05 0.03 0.07 4.40 0.84 18.56 34.07 

RSE2NG 93.45 3.79 0.72 0.12 0.15 0.04 0.03 0.07 0.79 0.84 17.32 33.78 

RSE3NG 93.39 3.71 0.71 0.12 0.15 0.05 0.04 0.07 0.83 0.93 17.34 33.73 

RSE6NG 93.95 3.07 0.70 0.12 0.15 0.05 0.04 0.08 0.83 1.02 17.28 33.54 

RSE65NG 93.46 3.76 0.86 0.12 0.16 0.04 0.03 0.05 0.78 0.74 17.31 33.86 

Figure 21 shows the heat release rate results for test #16 with polystyrene. Six kilograms of 

polystyrene pellets were ignited using a heptane spray fire which remained on for a period of 

30 s. Figure 22 shows images of the fire at two different heat release rates. The left side of the 

figure shows the fire approximately 6.5 min after ignition when the HRR was 170 kW. Large 

amounts of black soot can be seen exiting the doorway. The right side of the figure shows the 

fire 8.5 min after ignition, when the fire size was 340 kW, approaching its peak HRR. 

Figure 23 shows the heat release rate results for a toluene pool fire (test #10). The pool (burner 

B in Fig. 2) was half-filled with fuel when ignited at t = 0 s. As the thermal environment of the 

compartment changed, the fuel flow rate and the water flow rate to the liquid-cooled burner were 

continually adjusted to achieve near-steady burning. Figure 24 shows images of the toluene fires 

at heat release rates of 60 kW (left side) and 200 kW (right side). 

The HRR results for the ethanol pool fire (test #11) are shown in Fig. 25. As with the methanol 

fire, the discrepancy between the measured HRR and the enthalpy input was primarily due to 

fuel accumulation in the pool. Images of test #11 with ethanol are shown in Fig. 26. 

The HRR results for the heptane spray fire (test #15) are shown in Fig. 27. Unlike the tests with 

alcohol fuels, there was no evidence of fuel accumulation in the burner during test #15. 

Inspection of the HRR curve at approximately 3200 s after ignition shows that when the fuel 

delivery was reduced (from 500 kW to 90 kW), there was no lag in the HRR measurement. This 

suggests that incomplete combustion (rather than accumulation) is responsible for the significant 

differences in the measured and complete burning rates of the fuel. Images of the 160 kW and 

370 kW heptane fires are shown in Fig. 28. 
34

0 1000 2000 3000 4000 5000

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

54 HRRcal (kW)

55 HRRburner (kW)

Figure 17. Heat release rate results for natural gas test #3 using Burner A. 

Figure 18. Photograph of test #3 natural gas, HRR = 75 kW (left), HRR = 400 kW (right). (kW) 
35

0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

54 HRRcal (kW)

55 HRRburner (kW)

Figure 19. Heat release rate results for methanol test #12 using spray burner C. 

Figure 20. Photographs of methanol test #12: Open burn, HRR = 50 kW (left), t = 645 s, 

HRR = 72 kW (middle), t = 2885 s, HRR = 305 kW (right). (kW) 
36

-200 0 200 400 600 800 1000 1200 1400

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

yy p j

54 HRRcal (kW)

55 HRRburner (kW)

Figure 21. Heat release rate results for polystyrene test #16. A heptane spray was used to ignite 

6 kg of polystyrene pellets.

Figure 22. Photographs of polystyrene test #16: t = 389 s, HRR = 170 kW (left), 

 t = 515 s, HRR = 340 kW (right). (kW) 
37

0 1000 2000 3000 4000 5000 6000

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

p j

54 HRRcal (kW)

55 HRRburner (kW)

Figure 23. Heat release rate results for toluene pool fire (test #10). 

Figure 24. Photographs of toluene pool fire (test #10): t = 1279 s, HRR = 60 kW (left), 

 t = 3843 s, HRR = 200 kW (right). (kW) 
38

0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

p j

54 HRRcal (kW)

55 HRRburner (kW)

Figure 25. Heat release rate results for ethanol test #11.

Figure 26. Photographs of ethanol test #11: t = 1709 s, HRR = 145 kW (left), 

t = 3328 s, HRR = 328 kW (right). (kW) 
39

0 500 1000 1500 2000 2500 3000 3500 4000

0

100

200

300

400

500

Time From Ignition (s)

Heat Release Rate

p pj

54 HRRcal (kW)

55 HRRburner (kW)

Figure 27. Heat release rate results for heptane test #15.

Figure 28. Photographs of heptane test #15: t = 1000 s, HRR = 160 kW (left), 

t = 2870 s, HRR = 370 kW (right). (kW) 
40

A summary of HRR results is shown in Fig. 29. The measured heat release rate of the fire using 

oxygen calorimetry is plotted as a function of the ideal heat release rate predicted by the set fuel 

delivery rate. The dashed line on this plot represents a combustion efficiency of 1 (complete 

combustion). As expected, the combustion efficiency of the cleaner burning fuels (natural gas, 

methanol and ethanol) was closer to 1.0 than the highly sooting fuels (toluene and heptane). For 

most of the fuels, the global combustion efficiency decreased as the fire became more 

underventilated. Further discussion of the combustion efficiency can be found in Sec. 5.5 of this 

report. A summary of the time averaged steady-state results of HRR and exhaust stack species 

measurements is given in Table 10. The averaging period for each row of data in this table is 

given in the column labeled “SS Window”. A description of the remaining columns are given in 

Table 8. Total hydrocarbons in the exhaust stack were not measured in the exhaust stack during 

the first set of tests. 

0 100 200 300 400 500 600

0

50

100

150

200

250

300

350

400

450

500

Ideal (Burner) Heat Release Rate (kW)

Measured Heat Release Rate (kW) 

Natural Gas

Heptane, pool

Heptane, spray

Toluene, pool

Methanol, spray

Ethanol, spray

 Figure 29. Steady state heat release rate results. Dashed line represents ideal or complete 

burning. 
41

Table 10. Summary of time averaged steady-state results of HRR and exhaust stack species 

measurements. 

start (s) stop (s) Mean U Mean U Mean U Mean U Mean U Mean U Mean U

500 1924 74.6 9.1 75.7 2.4 20.75 0.21 1897 98 32 15

2220 3384 186.4 22.0 181.8 5.7 20.43 0.21 3780.0 194.4 59.2 26.5

3600 3944 77.0 10.0 75.9 1.7 20.74 0.21 1921.1 78.5 32.2 11.0

400 1594 256.9 30.3 268.1 4.6 20.14 0.20 5133.1 191.3 23.3 5.9

1840 3229 394.8 46.2 396.6 6.1 19.65 0.20 7939.8 203.3 31.4 5.0

3500 4454 179.1 21.4 179.2 3.4 20.40 0.20 3666.0 108.9 26.2 5.5

4550 5004 115.3 14.3 114.7 2.5 20.59 0.21 2588.1 93.3 18.1 2.7

5200 5579 47.8 6.3 47.7 1.3 20.78 0.21 1487.3 70.0 23.4 5.6

475 1139 264.8 31.9 266.1 4.1 20.12 0.20 5213.2 138.8 20.7 8.7

1300 2224 407.7 49.5 397.6 6.1 19.61 0.20 8068.6 258.7 14.9 5.0

2555 3449 179.4 22.0 178.1 2.9 20.39 0.20 3692.6 94.7 20.9 5.1

3645 4019 115.9 14.3 113.1 2.0 20.59 0.21 2576.8 110.3 9.2 2.9

4390 5249 73.8 9.1 74.4 1.8 20.72 0.21 1846.6 67.3 4.0 3.2

1375 2149 153.3 19.0 20.49 0.21 3761.3 208.9 67.2 21.1 43.6 21.6

2850 3334 268.7 33.0 273.9 5.3 20.14 0.21 6112.6 227.9 132.2 28.6 291.9 87.3

4090 5489 374.9 45.3 425.1 8.8 19.83 0.22 7456.2 480.7 230.3 77.6 864.6 216.0

1245 1799 140.5 17.3 20.55 0.21 3245.6 169.5 40.2 16.4 71.0 24.7

2340 2969 221.2 27.2 20.33 0.21 4304.1 168.0 152.0 43.1 409.8 105.6

660 1539 73.6 9.2 75.2 1.5 20.77 0.21 1756.1 75.2 2.1 2.4

2515 4179 173.7 20.2 179.3 5.1 20.48 0.21 3295.0 123.5 50.0 15.4

4425 4944 272.1 32.5 268.6 4.2 20.19 0.21 4922.9 164.7 21.9 3.4

5090 5724 417.5 49.5 399.3 9.0 19.72 0.20 7550.3 228.9 18.8 2.7

6090 6549 80.5 10.1 75.0 1.6 20.76 0.21 1708.8 59.4 1.7 2.2

285 679 96.5 11.8 76.1 8.8 20.71 0.21 2178.2 76.6 176.6 26.9 7.2 5.1

920 1204 423.9 51.9 399.4 6.1 19.64 0.20 8246.5 384.4 111.4 56.9 58.2 12.4

1600 2329 272.8 32.2 269.2 12.8 20.15 0.20 5170.4 118.6 45.2 13.8 20.6 9.2

2540 2804 181.3 22.3 178.7 10.8 20.45 0.21 3510.4 88.9 31.5 8.1 8.5 5.6

2980 3259 85.5 10.7 74.1 10.6 20.75 0.21 1842.6 107.4 55.8 19.3 4.2 5.1

1200 1669 147.7 18.6 20.50 0.21 3714.7 182.5 12.4 2.9 2.2 5.1 52.0 21.9

2105 2664 246.2 32.2 20.15 0.22 5888.7 421.3 71.5 23.4 11.8 7.2 274.0 90.6

3040 3709 341.4 41.1 404.5 7.4 19.83 0.21 7465.3 292.3 247.6 47.6 24.2 10.4 820.8 167.3

8 Methanol (pool) 1439 2009 17.2 2.2 22.5 71.4 20.94 0.21 825.9 58.8 0.0 2.0 1.0 5.0

9 Ethanol (pool) 1300 2019 19.3 2.4 5.8 19.6 20.94 0.21 881.6 48.3 0.0 2.0 1.0 5.0

1400 1884 48.9 6.8 48.8 2.3 20.84 0.21 1623.4 156.2 70.8 37.8 42.2 6.3 115.7 32.5

2805 3154 137.6 18.1 179.3 3.3 20.51 0.21 4151.4 255.2 95.2 10.1 9.6 5.7 306.5 56.1

3600 4224 202.2 25.0 270.4 4.6 20.26 0.21 6059.0 352.3 41.4 9.7 5.4 5.4 458.4 105.1

4435 5044 295.4 37.6 399.9 6.3 19.93 0.21 8512.8 533.6 47.1 22.9 2.0 5.0 486.2 180.3

5120 5394 338.8 43.5 498.7 7.8 19.93 0.29 8267.7 1495.5 107.2 38.8 1.3 5.1 682.8 217.2

550 1039 82.5 11.4 74.0 2.0 20.72 0.21 2412.6 302.7 117.4 99.9 3.8 5.3

1400 1714 143.7 18.1 181.0 3.1 20.53 0.21 3805.4 141.8 12.3 11.6 2.0 5.0

2175 2849 263.1 32.1 268.8 4.5 20.11 0.21 6721.1 312.7 19.6 5.2 4.5 5.2

2940 4200 335.3 41.2 398.3 6.6 19.84 0.22 8531.0 728.5 54.1 7.9 19.1 6.8

300 724 71.8 9.3 85.0 2.1 20.77 0.21 1977.7 106.0 0.7 2.2 1.0 5.0

1145 1609 142.6 17.9 181.0 3.1 20.54 0.21 3585.3 209.5 1.5 2.3 1.0 5.0

1949 2669 239.8 28.8 270.5 4.3 20.21 0.21 5890.9 376.1 35.7 7.4 1.0 5.0

2760 3299 306.5 37.1 400.6 6.6 19.96 0.21 7612.1 470.1 87.4 15.8 2.0 5.0

13 Polystyrene 710 1344 14.9 2.0 0.0 1.0 20.93 0.21 845.0 57.5 17.5 2.9 9.5 5.1 47.5 23.0

14 Polystyrene 870 1724 67.3 8.8 0.0 1.0 20.76 0.21 2119.2 180.2 80.3 10.6 36.6 7.2 304.3 173.6

280 759 87.7 12.2 82.0 22.8 20.69 0.21 2406.0 198.3 6.4 2.4 2.0 5.0 31.4 21.8

950 1259 160.2 19.7 179.8 5.9 20.43 0.21 4149.7 124.1 20.0 4.0 2.6 5.1 111.9 25.9

1475 1999 227.3 28.1 271.8 4.4 20.20 0.20 5485.7 212.1 63.0 17.2 9.1 6.3 328.0 96.0

2200 2764 300.8 39.1 401.2 6.2 19.92 0.21 6789.6 256.9 243.8 31.6 41.1 15.8 912.8 166.1

2790 3169 377.3 50.6 501.5 7.7 19.63 0.20 8412.1 317.1 384.1 35.9 89.4 22.5 1155.3 162.4

3390 3734 83.3 11.9 84.7 5.1 20.69 0.21 2379.4 159.0 9.7 3.3 2.0 5.0 45.1 24.9

545 649 358.1 45.7 0.0 1.0 19.73 0.20 9108.1 214.6 251.1 30.3 3.0 5.0 1258.4 228.3

715 769 308.7 40.0 0.0 1.0 19.86 0.20 8814.7 213.8 78.2 38.4 2.0 5.0 537.1 183.6

12 Methanol (spray)

15 Heptane (spray)

16 Polystyrene

7 Heptane (pool)

10 Toluene (pool)

11 Ethanol (spray)

5 Heptane (pool)

6 Natural Gas

6.5 Natural Gas

2 Natural Gas

3 Natural Gas

4 Heptane (pool)

THC stack 

(mol/mol)*1e6

MS stack 

(mg/m3

)

1 Natural Gas

HRRburner 

(kW)

O2 stack 

(mol/mol) %

CO2 stack 

(mol/mol)*1e6

CO stack 

Test # Fuel (mol/mol)*1e6 SS Window HRRcal 

(kW)
42

 3.2 Temperatures

The time history of the interior gas temperature is shown at 4 locations during the natural gas 

test (#3) in Fig. 30. This figure shows the difference in temperature between the front and rear 

gas sample locations and the temporal variation of temperature in the compartment. Refer to 

Table 2 and Fig. 4 for exact locations of the temperature probes. The measurement labels for the 

figures and tables in this section are described in Table 11. The aspirated thermocouple pumps 

remained on for the duration of this test. The general trend for these tests was higher 

temperatures in the upper layer at the front sample location than at the rear, however the 

magnitude of this difference was a function of fuel type and fire size. For all of the natural gas 

fires, the front and rear gas temperatures in the lower layer of the enclosure (24 cm from the 

floor) were not significantly different. The front to rear variation in lower layer temperature was 

more pronounced for the other fuels. 

Table 11. Description of interior gas temperature measurement labels.

Measurement 

Label Description 

15 TRSampA (C) Aspirated thermocouple at rear sample location (88 cm above floor)

16 TFSampA (C) Aspirated thermocouple at front sample location (88 cm above floor)

17 TR24A (C) Aspirated thermocouple at lower rear location (24 cm above floor)

18 TR80A (C) Aspirated thermocouple at upper rear location (80 cm above floor)

19 TF24A (C) Aspirated thermocouple at lower front location (24 cm above floor)

20 TF80A (C) Aspirated thermocouple at upper front location (80 cm above floor)

For a number of the tests, the aspiration flow pumps were run intermittently. This was done to 

observe the effect of the aspiration flow on the temperature and other measurements at nearby 

locations, as well as to conserve the water traps and filters. It was determined that turning on or 

off the suction flow to the aspirated thermocouples had no measurable effect on the observed gas 

species volume fractions. This is an important result since the separation distance between the 

probes was less than 3 cm. Further analysis of probe interactions can be found in Appendix B. 

The thermocouple response to cycling the aspiration flow is show in Fig. 31. This figure shows 

the front aspirated thermocouple measurements at two different heights in the enclosure for 

test #15 using heptane and the spray burner. The HRR was approximately 220 kW during this 

time window. The upper series in Fig. 31 is the aspirated thermocouple at the front sample 

location (TFSampA in Table 2), 10 cm below the ceiling. Because this probe was in a region 

with high soot mass fraction, the aspiration had little effect on the average temperature results. 

The faster time response (due to high convective heat transfer while aspirating) increases the 

measurement variation. The high soot mass fraction caused little difference to the probe’s 

average temperature with and without aspiration because the probe’s optical view of any cooler 

temperature radiative heat sinks was minimal due to soot blockage. 

The aspirated temperature probe located inside the front of the enclosure at a height of 24 cm 

above the floor (TF24A) is also shown. The measured temperature at this location decreased by 

more than 500 o

C when the aspiration was applied. This result shows the important application 

of aspirated thermocouples to the thermal characterization of flashed-over compartment fires. 

Although differences between bare-beads and aspirated thermocouples are usually much greater 
43

in the lower layer, the upper layer differences may be as much as 100 °C to 200 °C depending, 

primarily, on a fuel’s sooting characteristics. 

A detailed analysis of the aspirated thermocouple time response and uncertainty can be found in 

Appendix A of this report. All of the temperature results reported here represent the temperature 

of the thermocouple bead, not the true temperature of the gas. 

Average temperatures were calculated over pseudo-steady periods for all of the tests. 

Temperature measurements collected when the aspiration pumps were off or during a transient 

period were not included in the averages. A summary of the rear gas temperature measurements 

with combined expanded uncertainty (U) are listed in Table 12. The front gas temperatures are 

given in Table 13. The average temperatures are plotted as a function of HRR for the natural gas 

full-door tests in Fig. 32. These temperature measurements demonstrate the reproducibility of the 

measurements over a number of different days. Figure 33 and Fig. 34 show the steady 

temperatures at the front and rear gas sample location for all of the fuels included in this study. 

In general, the soot producing fires (heptane, toluene, polystyrene) produced hotter gas 

temperatures inside the enclosure than the cleaner fires (natural gas, alcohols) at the same 

measured HRR. 

0 1000 2000 3000 4000 5000

0

200

400

600

800

1000

1200

1400

Time From Ignition (s)

Gas Temperature (C)

p j

15 TRSampA (C)

16 TFSampA (C)

17 TR24A (C)

19 TF24A (C)

Figure 30. Gas temperature measurement results for 4 positions inside RSE. Natural gas test #3. 
44

5250 5300 5350 5400 5450 5500 5550

200

300

400

500

600

700

800

900

1000

1100

1200

Time (s)

Temperature (oC)

Figure 31. Gas temperature measurements at two front interior locations inside RSE during the 

heptane spray fire (test # 15). 

0 50 100 150 200 250 300 350 400 450

0

200

400

600

800

1000

1200

Heat Release Rate (kW)

Temperature (oC)

Figure 32. Steady state average temperature measurements at interior locations for repeated 

natural gas fires (tests #1, #2, #3). The lines in this figure are piecewise cubic polynomial fits to 

the data. 

15 TRSampA

16 TFSampA

18 TR80A

20 TF80A

17 TR24A

19 TF24A

Aspiration 

flow on 

Aspiration 

flow off 

TF24A 

TFSampA 
45

0

200

400

600

800

1000

1200

1400

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample Temperature (°C)

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 33. Steady state average temperature results from aspirated thermocouple measurement 

at front gas sampling location. note: Front sample thermocouple failed during toluene fire 

(test #10) and polystyrene fire (test #16) 

0

200

400

600

800

1000

1200

1400

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample Temperature (°C)

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

 

Figure 34 Steady state average temperature results from aspirated thermocouple measurement at 

rear gas sampling location. 
46

Table 12. Steady state rear gas temperatures and total heat flux to the floor inside the RSE. 

start (s) stop (s) Mean U Mean U Mean U Mean U Mean U

500 1924 75 11 17.1 4.1 536 27 546 32 80 28

2220 3384 186 23 57.9 10.8 717 53 719 46 211 25

3600 3944 77 11 30.2 3.6 609 20 589 29 140 16

400 1594 257 32 57.6 18.0 747 58 740 69 225 44

1840 3229 395 49 75.6 6.3 777 56 759 37 268 11

3500 4454 179 23 62.2 4.6 708 46 691 40 228 11

4550 5004 115 15 43.3 4.1 674 11 715 25 178 12

5200 5579 48 7 18.1 2.9 511 37 513 19 105 22

475 1139 265 33 37.8 9.0 737 53 747 50 232 30

1300 2224 408 55 55.3 7.8 769 34 803 33 277 17

2555 3449 179 25 50.0 3.4 701 19 803 48 237 10

3645 4019 116 14 35.3 2.8 712 23 715 26 182 14

4390 5249 74 10 20.2 2.1 594 16 598 18 121 15

1375 2149 153 21 52.9 7.0 703 22 707 26 234 22

2850 3334 269 34 96.9 44.1 942 82 939 61 430 66

4090 5489 375 54 235.0 35.4 1186 50 1171 43 845 274

1245 1799 140 18 67.6 9.9 789 36 783 33 441 64

2340 2969 221 29 220.7 34.2 1219 101 1248 90 1254 170

660 1539 74 11 24.3 4.3 638 28 644 30 186 28

2515 4179 174 22 70.6 5.9 722 80 781 44 390 23

4425 4944 272 33 76.5 5.2 614 44 698 36 418 15

5090 5724 417 50 73.9 4.8 572 25 654 28 420 14

6090 6549 80 11 42.8 5.0 742 42 738 35 285 22

285 679 97 12 23.8 3.0

920 1204 424 51 83.1 12.0

1600 2329 273 33 87.9 7.4

2540 2804 181 22 77.4 6.9

2980 3259 85 11 29.5 4.3

1200 1669 148 20 54.1 6.5 741 21 718 16 210 14

2105 2664 246 38 133.0 41.5 913 59 921 54 379 34

3040 3709 341 43 204.8 37.0 1139 45 1125 32 642 94

8 Methanol (pool) 1439 2009 17 3 2.4 0.5 188 6 185 8 67 5

9 Ethanol (pool) 1300 2019 19 3 2.7 0.4 210 9 203 9 75 5

1400 1884 49 9 14.6 2.4

2805 3154 138 20 65.7 7.0 741 21 763 21 240 20

3600 4224 202 27 116.1 15.7 890 53 919 45 354 36

4435 5044 295 43 257.4 37.3 1150 58 1254 56 970 89

5120 5394 339 44 336.2 42.9

550 1039 83 15 19.9 2.5 566 26 546 20 126 6

1400 1714 144 18 44.6 5.0 748 44 729 36 221 15

2175 2849 263 34 80.5 8.1 848 44 307 13

2940 4200 335 51 80.9 5.9 856 45 305 10

300 724 72 10 13.4 2.4 499 26 490 27 94 11

1145 1609 143 19 39.5 5.7 769 64 736 51 192 20

1949 2669 240 30 62.4 4.3 864 95 842 105 259 13

2760 3299 306 38 61.1 4.2 801 50 803 45 260 13

13 Polystyrene 710 1344 15 3 3.5 0.9 171 23 183 26 112 19

14 Polystyrene 870 1724 67 12 23.6 5.1 467 30 467 26 130 13

280 759 88 16 30.8 6.5 575 18 571 15 155 7

950 1259 160 19 76.9 8.0 775 27 784 21 325 14

1475 1999 227 30 134.3 19.2 954 83 999 106 497 85

2200 2764 301 46 143.0 20.5 921 140 1005 137 468 42

2790 3169 377 58 135.7 11.8 858 129 901 127 452 27

3390 3734 83 15 46.7 11.0 627 22 628 17 219 22

545 649 358 42 104.2 20.5 1169 84 1240 130 1256 137

715 769 309 35 169.9 17.5 1263 36 1160 45 16 Polystyrene

15 Heptane (spray)

1 Natural Gas

12 Methanol (spray)

10 Toluene (pool)

11 Ethanol (spray)

TR24A 

(C)

HRR 

(kW)

HFR 

(kW/m2

)

TR80A 

(C) SS Window TRSampA 

Test # Fuel (C)

6.5 Natural Gas

5 Heptane (pool)

6 Natural Gas

4 Heptane (pool)

2 Natural Gas

3 Natural Gas

7 Heptane (pool)
47

Table 13. Steady state front gas temperatures and total heat flux to the floor inside the RSE. 

start (s) stop (s) Mean U Mean U Mean U Mean U Mean U

500 1924 75 11 19.7 4.6 645 53 604 48 85 35

2220 3384 186 23 69.4 10.1 1104 206 979 136 223 21

3600 3944 77 11 28.0 6.2 733 53 631 34 137 20

400 1594 257 32 80.6 16.2 1066 101 1075 108 235 38

1840 3229 395 49 94.8 7.6 988 84 1025 110 274 23

3500 4454 179 23 79.6 5.7 1032 173 1043 85 226 23

4550 5004 115 15 52.7 5.2 796 39 832 61 173 17

5200 5579 48 7 20.5 3.2 520 46 503 23 101 21

475 1139 265 33 80.6 12.8 1055 99 1024 80 232 32

1300 2224 408 55 95.8 7.8 1013 119 1016 140 269 27

2555 3449 179 25 80.4 5.7 1104 175 1038 142 226 28

3645 4019 116 14 52.8 4.2 924 91 823 62 179 19

4390 5249 74 10 28.3 3.0 670 40 633 27 113 16

1375 2149 153 21 51.3 5.1 767 39 847 44 284 12

2850 3334 269 34 98.4 11.8 1044 54 1090 77 468 58

4090 5489 375 54 168.1 21.9 1173 21 1210 33 677 97

1245 1799 140 18 80.8 12.0 903 71 991 70 569 63

2340 2969 221 29 119.2 16.5 1043 66 1149 50 1037 144

660 1539 74 11 9.8 1.6 739 48 714 44 210 30

2515 4179 174 22 43.2 7.6 1035 71 1032 84 423 14

4425 4944 272 33 51.5 3.6 966 45 978 45 456 14

5090 5724 417 50 50.2 3.3 860 79 829 63 450 10

6090 6549 80 11 28.9 4.1 830 73 804 43 302 24

285 679 97 12 24.7 2.7

920 1204 424 51 88.8 8.9

1600 2329 273 33 90.9 6.6

2540 2804 181 22 77.2 5.5

2980 3259 85 11 29.1 3.8

1200 1669 148 20 57.8 6.4 857 39 831 28 276 19

2105 2664 246 38 118.5 12.2 1029 35 1091 41 423 40

3040 3709 341 43 166.7 13.4 1145 22 1230 56 594 63

8 Methanol (pool) 1439 2009 17 3 2.2 0.4 80 5 181 8 62 5

9 Ethanol (pool) 1300 2019 19 3 2.6 0.4 91 6 198 8 76 6

1400 1884 49 9 13.6 2.4

2805 3154 138 20 59.2 6.6 859 27 297 24

3600 4224 202 27 101.9 10.3 1067 54 421 50

4435 5044 295 43 174.8 17.1 1241 38 643 75

5120 5394 339 44 210.0 22.2

550 1039 83 15 17.6 3.3 605 46 541 27 134 11

1400 1714 144 18 45.3 5.9 951 108 851 70 224 39

2175 2849 263 34 86.8 8.9 1085 77 308 23

2940 4200 335 51 82.3 6.4 1013 59 288 25

300 724 72 10 11.7 2.6 533 40 493 41 100 14

1145 1609 143 19 40.0 5.9 993 95 850 109 200 22

1949 2669 240 30 64.9 4.9 1171 94 1134 109 252 32

2760 3299 306 38 61.5 4.1 1040 113 1144 122 244 33

13 Polystyrene 710 1344 15 3 2.3 0.6 161 17 167 21 88 17

14 Polystyrene 870 1724 67 12 17.2 4.5 460 28 454 26 139 19

280 759 88 16 26.1 7.1 601 29 569 26 157 11

950 1259 160 19 74.2 8.9 864 50 870 22 309 21

1475 1999 227 30 128.5 13.7 1040 47 1121 43 346 71

2200 2764 301 46 145.0 12.8 1140 38 1197 44 382 95

2790 3169 377 58 144.0 11.7 1169 36 405 81

3390 3734 83 15 38.2 8.4 629 21 161 57

545 649 358 42 71.2 14.7 1195 56 968 111

715 769 309 35 90.7 6.7 1244 29 947 64

2 Natural Gas

3 Natural Gas

1 Natural Gas

7 Heptane (pool)

Test # Fuel

6.5 Natural Gas

5 Heptane (pool)

TF24A 

(C)

HRR 

(kW)

HFF 

(kW/m2)

TF80A 

(C) SS Window TFSampA (C)

10 Toluene (pool)

11 Ethanol (spray)

4 Heptane (pool)

6 Natural Gas

16 Polystyrene

15 Heptane (spray)

12 Methanol (spray)
48

 3.3 Doorway Velocities and Temperatures

Doorway velocity measurements were performed in order to establish the ventilation conditions 

of the compartment fires. Velocity probes and thermocouples were placed at five different 

heights in the doorway. Aspirated thermocouples and bare bead thermocouples were placed 

half-way between the velocity probes at three different heights to get a first order estimate of the 

error in temperature with the bare bead thermocouples due to radiation. See Fig. 4 and Table 3 

for exact locations of the doorway probes. 

Figure 35 compares the bare bead and aspirated thermocouple temperature measurements at 

three different heights in the doorway for all 17 fire tests. The results show that at heights of 

50 cm and 70 cm in the doorway the bare bead temperate measurements agree reasonably well 

with the aspirated temperature measurements. This was because these points were above the 

neutral plane where the velocities were relatively high and the large opacity of the hot gases in 

the upper layer reduced the radiation losses. However, at 30 cm, the temperature probes are 

below the neutral plane where the velocity is lower and the gas is optically thin cool room air. 

This thermocouple is exposed to radiation from the hot upper layer resulting in a significant 

overestimate of the gas temperature. The results of this analysis are that the error in velocity due 

to using the bare-bead temperature is less than 10 % above the neutral plane and less than 30 % 

below the neutral plane. Figure 36 show the doorway temperature measurements (not corrected 

for radiation) for a steady natural gas fire at 265 kW during test #3. The error bars in this figure 

represent twice the standard deviation of the measurement and does not include uncertainty due 

to radiation. 

The steady state velocity profiles along the vertical centerline of the doorway for test #3 using 

natural gas are shown in Fig. 37. As expected, the magnitude of the velocity measurement 

increases and the location of the neutral plane moves downward as the fire size increases. 

Figure 38 shows the steady doorway velocity profiles for test #15 using the heptane spray burner. 

The lines shown in both of these figures are drawn to highlight trends in the data and do not 

represent a physical model. 

The mass flow rate of air into the enclosure is often used to define a global equivalence ratio. 

This value can be found by integrating the product of velocity and density over the area below 

the neutral plane. Although this calculation was attempted, the uncertainty in velocity, 

temperature, and location of neutral plane (due to sparse data) prevented a meaningful 

determination of mass flow rate. Future work is planned to better quantify the doorway mass 

flow. 
49

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

Temperarture of Apirated Thermocouple, o

C

Temperarture of Bare Bead Thermocouple, o

C

z = 30 cm

z = 50 cm

z = 70 cm

Figure 35. Comparison of doorway centerline temperatures measured using aspirated and bare 

bead thermocouple at the same position. Steady state average values for all RSE tests. 

0

10

20

30

40

50

60

70

80

0 100 200 300 400 500 600 700 800 900 1000 1100

Measured Temperature (C)

Vertical Position in Doorway (cm)

Bare Bead Thermocouple

Aspirated Thermocouple

Figure 36. Steady state doorway temperature measurements for natural gas test #3, 

HRR=265 kW. Error bars in this figure are 2 standard deviations. 
50

0

10

20

30

40

50

60

70

80

-7 -6 -5 -4 -3 -2 -1 0 1 2

Velocity Normal to Door (m/s)

Vertical Position in Doorway (cm)

264.78

407.71

179.42

115.91

73.75

HRR (kW)

Figure 37. Doorway centerline velocity profiles for natural gas fire (test #3). 

0

10

20

30

40

50

60

70

80

-7 -6 -5 -4 -3 -2 -1 0 1 2

Velocity Normal to Door (m/s)

Vertical Position in Doorway (cm)

87.72

160.16

227.34

300.75

377.35

HRR (kW)

Figure 38. Doorway centerline velocity profiles for heptane spray fire (test #15). 
51

 3.4 Interior Gas Species and Soot 

The measurement labels for the figures and tables in this section are described in Table 14. The 

time history of the gas and soot species measurements at the front and rear probe locations for 

the heptane spray fire (test #15) is shown in Fig. 39. The mean measured HRR value over a 

given steady time window is annotated on this figure. Several observations of this plot are 

noteworthy. When the fire was underventilated (between 2000 s and 3000 s) there was a 

mirroring of the random temporal variations in the CO and CO2 volume fractions. Also evident 

in Fig. 39 was a strong positive correlation between CO and total hydrocarbons. The measured 

gas species volume fractions became less uniform (front to back) as the fire size increased. The 

uniformity of the upper gas layer was dependant on fire size and fuel type. The transient gas 

volume fractions and soot mass fractions for polystyrene (test #16) are shown in Fig. 40. Very 

high soot mass fractions (nearly 11 % at the front sample location) were observed during this 

test. The total hydrocarbons measured less than 0.1 % for the entire test. This was a surprising 

result since the oxygen in the upper layer was completely depleted and flames were observed 

exiting the doorway (see photograph in Fig. 22). More work is needed to understand this result. 

Table 14. Description of interior gas species and soot measurement labels.

Measurement Label Description 

2 O2Rear Rear O2 volume fraction corrected for water (wet)

3 CO2Rear Rear CO2 volume fraction corrected for water (wet)

4 CORear Rear CO volume fraction corrected for water (wet)

5 THCRear Rear Total Hydrocarbons volume fraction corrected for water (wet)

6 SootRear Rear Soot Mass fraction corrected for water (wet)

7 O2Front Front O2 volume fraction corrected for water (wet)

8 CO2Front Front CO2 volume fraction corrected for water (wet)

9 COFront Front CO volume fraction corrected for water (wet)

10 THCFront Front Total Hydrocarbons volume fraction corrected for water (wet)

11 SootFront Front Soot Mass fraction (g/g), corrected for water (wet)

Figure 41 shows the time averaged species volume fractions as a function of heat release rate for 

all of the natural gas tests with the full-door configuration (tests #1, #2, #3 and #6.5). The trend 

lines are included in this figure to help visualize general trends in the data, but do not have a 

theoretical basis. The figure demonstrates the excellent reproducibility of the gas species 

measurements and lack of sensitivity of the results to the two different wall lining materials and 

burners (see Table 1) used in this study. 

Figure 42 shows the oxygen volume fraction at the front sample location as a function of heat 

release rate for the six different fuel types included in this study. Oxygen was depleted for fires 

larger than about 280 kW. There was some small difference for the various fuel types, with the 

natural gas fires exhibiting oxygen depletion for slightly smaller values (about 260 kW). The 

rear sample oxygen measurements are summarized in Fig. 43. For the natural gas fires, the 

oxygen was depleted in the rear of the enclosure at a lower HRR (≈ 180 kW) than in the front 

(≈ 250 kW), however this result could not be generalized for all fuels. For example, the toluene 

and heptane fires displayed significant amounts oxygen at the rear sample location at HRR’s 

where oxygen was completely depleted at the front sample location. This was an unexpected 
52

result since flames were observed exiting the doorway (see right side of Fig. 28); however it was 

not inconsistent with the other species measurements that showed locally lean conditions at the 

upper rear sample location. In addition, preliminary FDS modeling results showed a similar 

structure in the compartment. The condition at which oxygen is depleted inside of the 

compartment is a critical point in the characterization of compartment fire chemistry, indicating 

when the fire becomes underventilated. In terms of the global equivalence ratio (GER) concept 

(see Section 1.1 of this report), it is the point at which the GER value is equal to 1.0. Based on 

the results shown in Figure 42, oxygen depletion could be expected in an ISO 9705 enclosure at 

about 1800 kW, based on the ventilation scaling relation given in Table 25. 

Figure 44 through Fig. 51 shows the steady gas and soot sample results for all of the different 

fuels at the front and rear locations. The steady state carbon monoxide results are summarized in 

Fig. 46 (front) and Fig. 47 (rear). As expected, the measured CO values were significantly 

increased after the fire reached a ventilation limited regime (as indicted by depleted oxygen at 

the sample locations). The results of the total hydrocarbon volume fraction measurements are 

shown in Fig. 48 and Fig. 49. Figure 50 and Fig. 51 show the results of the gravimetric soot 

mass fraction measurements. The species volume fraction results are examined further in Sec. 5 

of this report. 

The time averaged values for all of the gas and soot species volume fractions are listed in 

Table 15 and Table 16 for the front and rear sample locations, respectively. The values listed in 

these tables were used to generate the plots shown in Fig. 41 through Fig. 51. The combined 

expanded uncertainties listed in these tables represent the absolute percentage (not relative to the 

mean fractional value). 

0 500 1000 1500 2000 2500 3000 3500 4000

0

0.05

0.1

0.15

0.2

Time From Ignition (s)

Species Concentration (wet)

Figure 39. Transient gas volume fractions and soot mass fractions for heptane test #15.

2 O2Rear

3 CO2Rear

4 CORear 

5 THCRear

6 SootRear (g/g)

7 O2Front 

8 CO2Front

9 COFront

10 THCFront

11 SootFront (g/g)

160 kW 

88 kW 

227 kW 

301 kW 

377 kW 

83 kW 

Gas Volume Fraction, wet Soot Mass Fraction, wet 
53

-200 0 200 400 600 800 1000 1200 1400

0

0.05

0.1

0.15

0.2

Time From Ignition (s)

Species Concentration (wet)

Figure 40. Transient gas volume fractions and soot mass fractions for polystyrene test #16.

0 50 100 150 200 250 300 350 400 450

0

2

4

6

8

10

12

14

16

Heat Release Rate (kW)

Species Concentration, Volume % (wet)

Figure 41. Steady state gas and soot species results for natural gas full door tests #1, #2, #3, and 

#65. The lines in this figure are piecewise cubic polynomial fits to the data. 

2 O2Rear

7 O2Front

3 CO2Rear

8 CO2Front

4 CORear

9 COFront

5 THCRear

10 THCFront

6 SootRear (g/g)

11 SootFront (g/g)

2 O2Rear

3 CO2Rear

4 CORear 

5 THCRear

6 SootRear (g/g)

7 O2Front 

8 CO2Front

9 COFront

10 THCFront

11 SootFront (g/g)

Gas Volume Fraction, wet, % Soot Mass Fraction, wet, % 

Gas Volume Fraction, wet Soot Mass Fraction, wet 
54

0%

2%

4%

6%

8%

10%

12%

14%

16%

18%

20%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample Oxygen Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 42. Steady state average oxygen volume fraction measurements at front sample probe 

location. 

0%

2%

4%

6%

8%

10%

12%

14%

16%

18%

20%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample Oxygen Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 43. Steady state average oxygen volume fraction measurements at rear sample probe 

location. 
55

0%

2%

4%

6%

8%

10%

12%

14%

16%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample CO2 Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 44. Steady state average carbon dioxide volume fraction measurements at front sample 

probe location. 

0%

2%

4%

6%

8%

10%

12%

14%

16%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample CO2 Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptanes (pool burner B)

Heptanes (narrow door)

Heptanes (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 45. Steady state average carbon dioxide volume fraction measurements at rear sample 

probe location.
56

0%

1%

2%

3%

4%

5%

6%

7%

8%

9%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample CO Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 46. Steady state average carbon monoxide volume fraction measurements at front 

sample probe location.

0%

1%

2%

3%

4%

5%

6%

7%

8%

9%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample CO Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 47. Steady state average carbon monoxide volume fraction measurements at rear sample 

probe location.
57

0%

5%

10%

15%

20%

25%

30%

35%

40%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample THC Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 48. Steady state average total hydrocarbon volume fraction measurements at front 

sample probe location.

0%

5%

10%

15%

20%

25%

30%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample THC Volume Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 49. Steady state average total hydrocarbon volume fraction measurements at rear sample 

probe location. 
58

0%

1%

2%

3%

4%

5%

6%

7%

8%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Sample Soot Mass Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 50. Steady state gravimetric soot mass fraction measurements at front sample probe 

location.

0%

2%

4%

6%

8%

10%

12%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Sample Soot Mass Fraction %, wet

Natural Gas

Natural Gas (narrow door)

Heptane (pool burner B)

Heptane (narrow door)

Heptane (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 51. Steady state gravimetric soot mass fraction measurements at rear sample probe 

location.
59

Table 15. Summary of steady state rear gas and soot sample probe measurements.

start (s) stop (s) Mean U Mean U Mean U Mean U Mean U Mean U

500 1924 75 11 9.86 0.77 5.20 0.38 0.11 0.04 0.01 0.01 0.14 0.01

2220 3384 186 23 0.12 0.12 8.88 0.28 1.22 0.42 2.16 2.14 0.17 0.01

3600 3944 77 11 10.69 1.37 4.77 0.70 0.06 0.03 0.01 0.01

400 1594 257 32 0.05 0.12 8.12 0.61 1.63 0.40 4.60 1.94 0.14 0.01

1840 3229 395 49 0.06 0.15 7.17 0.48 2.47 0.31 9.30 2.33 0.17 0.01

3500 4454 179 23 0.17 0.52 8.87 0.20 0.75 0.27 1.23 2.42 0.08 0.01

4550 5004 115 15 4.91 1.06 7.41 0.49 0.02 0.02 0.02 0.04

5200 5579 48 7 13.76 0.43 3.34 0.22 0.05 0.03 0.01 0.01

475 1139 265 33 0.12 0.10 7.93 0.22 1.69 0.21 6.62 2.11 0.15 0.01

1300 2224 408 55 0.06 0.10 7.06 0.28 2.40 0.22 5.68 2.85 0.17 0.01

2555 3449 179 25 0.31 0.25 8.79 0.20 0.77 0.25 1.80 3.87 0.08 0.01

3645 4019 116 14 5.64 1.26 7.07 0.56 0.01 0.03 0.03 0.28

4390 5249 74 10 10.99 0.69 4.59 0.32 0.01 0.02 0.00 0.01 0.01 0.01

1375 2149 153 21 7.37 1.13 8.28 0.69 0.06 0.04 0.02 0.12 0.11 0.01

2850 3334 269 34 7.38 2.26 8.52 1.38 0.05 0.07 0.05 0.13 0.17 0.01

4090 5489 375 54 3.20 1.87 9.80 0.93 0.64 0.51 0.29 1.34 1.01 0.02

1245 1799 140 18 3.18 0.90 10.21 0.48 0.24 0.09 0.15 0.24 0.25 0.01

2340 2969 221 29 0.16 0.57 9.17 0.40 2.11 0.31 0.21 0.29 1.09 0.02

660 1539 74 11 6.44 1.31 6.67 0.60 0.02 0.03 0.02 0.19 0.01 0.01

2515 4179 174 22 0.16 0.04 8.00 0.44 1.49 0.37 5.92 2.71

4425 4944 272 33 0.14 0.04 7.00 0.23 2.13 0.15 13.47 1.96 0.17 0.01

5090 5724 417 50 0.09 0.05 6.11 0.31 2.35 0.12 23.99 3.63 0.17 0.01

6090 6549 80 11 6.52 2.16 6.67 0.97 0.00 0.02 0.51 0.36

285 679 97 12 10.01 1.08 5.42 0.61 0.19 0.06 0.01 0.09

920 1204 424 51 0.03 0.04 7.62 0.41 2.17 0.14 7.31 0.80 0.19 0.01

1600 2329 273 33 0.10 0.13 8.33 0.23 1.37 0.24 2.82 0.83

2540 2804 181 22 1.99 1.42 8.51 0.52 0.36 0.18 0.46 0.28

2980 3259 85 11 11.48 0.78 4.44 0.39 0.07 0.07 0.02 0.01

1200 1669 148 20 9.18 0.67 6.92 0.38 0.03 0.02 0.01 0.08 0.11 0.01

2105 2664 246 38 8.36 2.24 7.24 1.00 0.14 0.21 0.07 0.10 0.23 0.01

3040 3709 341 43 3.46 1.94 9.08 0.53 0.99 0.66 0.38 0.27 1.42 0.03

8 Methanol (pool) 1439 2009 17 3 18.97 0.27 1.08 0.10 0.00 0.02 0.00 0.01

9 Ethanol (pool) 1300 2019 19 3 18.82 0.21 1.23 0.06 0.00 0.02 0.00 0.01

1400 1884 49 9 15.74 0.63 3.58 0.41 0.21 0.09 0.18 0.05

2805 3154 138 20 11.24 0.55 6.86 0.46 0.16 0.03 0.01 0.01 0.69 0.01

3600 4224 202 27 9.30 0.88 8.43 0.67 0.04 0.03 0.01 0.01 0.47 0.01

4435 5044 295 43 3.74 1.35 12.33 0.86 0.18 0.11 0.01 0.01 1.33 0.03

5120 5394 339 44 2.19 0.76 12.76 0.39 0.65 0.40 0.01 0.01

550 1039 83 15 12.57 1.35 4.92 0.79 0.29 0.27 0.01 0.01

1400 1714 144 18 8.62 1.19 7.27 0.71 0.02 0.03 0.01 0.13 0.01 0.01

2175 2849 263 34 1.39 1.34 11.06 0.62 0.55 0.37 0.26 0.21 0.03 0.01

2940 4200 335 51 0.00 0.03 9.80 0.43 3.60 0.67 2.64 0.57 0.05 0.01

300 724 72 10 14.30 0.37 3.63 0.18 0.00 0.02 0.00 0.01

1145 1609 143 19 7.61 1.30 7.35 0.70 0.02 0.03 0.00 0.01 0.00 0.01

1949 2669 240 30 0.07 0.18 9.61 0.76 2.97 1.34 0.05 0.03 0.00 0.01

2760 3299 306 38 0.00 0.02 6.70 0.38 7.75 0.51 0.18 0.02 0.00 0.01

13 Polystyrene 710 1344 15 3 19.03 0.26 1.34 0.12 0.07 0.02 0.07 0.01 0.14 0.01

14 Polystyrene 870 1724 67 12 15.20 0.72 4.02 0.51 0.21 0.03 0.12 0.01 0.59 0.01

280 759 88 16 12.32 0.73 5.07 0.41 0.02 0.02 0.00 0.01

950 1259 160 19 9.72 0.87 6.54 0.50 0.04 0.02 0.01 0.01 0.20 0.01

1475 1999 227 30 7.93 2.75 7.57 1.37 0.06 0.13 0.03 0.08 0.34 0.01

2200 2764 301 46 0.93 1.22 10.66 0.52 0.76 0.44 0.43 0.35 1.24 0.03

2790 3169 377 58 0.24 0.49 9.93 0.76 1.70 0.85 1.47 0.92 2.80 0.06

3390 3734 83 15 12.76 0.66 4.75 0.38 0.04 0.03 0.03 0.01

545 649 358 42 0.16 0.13 9.36 0.76 3.93 0.64 0.08 0.04 10.72 0.21

715 769 309 35 0.12 0.03 10.89 0.79 2.81 0.51 0.04 0.01 8.07 0.16

Ethanol (spray)

16 Polystyrene

15 Heptane (spray)

4 Heptane (pool)

THC Rear 

(%)

HRR 

(kW)

O2 Rear 

(%)

12 Methanol (spray)

10 Toluene (pool)

11

Soot Rear 

g/g (%)

CO Rear 

(%)

7 Heptane (pool)

Test # Fuel

6.5 Natural Gas

5 Heptane (pool)

6

SS Window

CO2 Rear 

(%)

Natural Gas

1 Natural Gas

2 Natural Gas

3 Natural Gas
60

Table 16. Summary of steady state front gas and soot sample probe measurements. 

start (s) stop (s) Mean U Mean U Mean U Mean U Mean U Mean U

500 1924 75 11 12.45 0.63 3.98 0.29 0.07 0.03 0.02 0.18 0.18 0.01

2220 3384 186 23 3.14 1.06 7.54 0.41 0.87 0.32 1.27 1.83 0.25 0.01

3600 3944 77 11 13.07 0.65 3.81 0.30 0.06 0.03 0.03 0.28

400 1594 257 32 0.20 0.38 7.27 0.35 2.12 0.40 6.44 2.29 0.17 0.01

1840 3229 395 49 0.22 0.34 5.92 0.33 3.28 0.25 0.00 0.00

3500 4454 179 23 3.34 0.86 7.12 0.30 0.85 0.25 0.00 0.00 0.09 0.01

4550 5004 115 15 8.12 0.74 5.66 0.32 0.14 0.07 0.00 0.00

5200 5579 48 7 15.01 0.32 2.68 0.14 0.04 0.02 0.00 0.00

475 1139 265 33 0.60 0.35 7.28 0.22 1.97 0.27 5.91 2.72 0.17 0.01

1300 2224 408 55 0.12 0.10 6.08 0.33 3.19 0.20 13.45 5.67 0.00 0.00

2555 3449 179 25 2.87 0.65 7.42 0.17 0.90 0.21 2.59 3.37 0.09 0.01

3645 4019 116 14 7.98 0.58 5.84 0.23 0.13 0.05 0.18 0.33

4390 5249 74 10 12.92 0.49 3.70 0.21 0.02 0.02 0.03 0.39 0.01 0.01

1375 2149 153 21 8.21 0.97 7.63 0.56 0.05 0.03 0.07 0.14 0.12 0.01

2850 3334 269 34 0.83 1.86 11.12 1.22 1.63 1.16 1.26 2.85 0.87 0.02

4090 5489 375 54 0.23 0.34 9.94 0.61 1.89 0.58 0.82 1.46 2.52 0.05

1245 1799 140 18 2.30 0.78 10.94 0.44 0.15 0.09 0.92 2.72 0.29 0.01

2340 2969 221 29 1.56 1.91 10.37 1.14 0.87 0.69 1.13 1.54 2.80 0.07

660 1539 74 11 8.64 0.76 5.66 0.34 0.02 0.03 0.04 0.27 0.01 0.01

2515 4179 174 22 0.36 0.25 7.73 0.44 1.65 0.46 8.99 4.63

4425 4944 272 33 0.10 0.13 6.81 0.57 2.31 0.39 20.79 8.83 0.23 0.01

5090 5724 417 50 0.04 0.04 6.34 0.41 2.45 0.18 29.89 8.65 0.18 0.01

6090 6549 80 11 7.92 0.72 5.99 0.32 0.02 0.04 0.10 0.08

285 679 97 12 11.13 0.62 4.71 0.31 0.10 0.03 0.01 0.01

920 1204 424 51 0.13 0.10 5.97 0.33 2.94 0.20 12.47 0.95 0.20 0.01

1600 2329 273 33 0.51 0.15 6.80 0.14 2.41 0.13 5.90 0.51

2540 2804 181 22 2.59 0.16 7.39 0.11 1.10 0.09 1.55 0.25

2980 3259 85 11 12.29 0.64 4.02 0.29 0.03 0.03 0.05 0.02

1200 1669 148 20 9.10 0.59 6.92 0.33 0.05 0.03 0.03 0.02 0.13 0.01

2105 2664 246 38 3.09 1.08 9.68 0.33 0.65 0.27 0.40 0.16 0.72 0.02

3040 3709 341 43 0.71 0.36 9.45 0.37 1.97 0.39 0.64 0.20 2.06 0.04

8 Methanol (pool) 1439 2009 17 3 19.15 0.27 1.01 0.10 0.00 0.02 0.00 0.01

9 Ethanol (pool) 1300 2019 19 3 19.02 0.21 1.15 0.06 0.00 0.02 0.00 0.01

1400 1884 49 9 16.46 0.62 3.08 0.43 0.20 0.10 0.17 0.03

2805 3154 138 20 8.06 0.80 9.05 0.39 0.20 0.03 0.05 0.01 0.96 0.02

3600 4224 202 27 3.55 0.87 12.32 0.60 0.15 0.05 0.06 0.03 1.14 0.02

4435 5044 295 43 1.21 0.55 13.43 0.39 0.52 0.13 0.03 0.02 2.37 0.05

5120 5394 339 44 0.28 0.38 13.10 0.47 1.36 0.49 0.01 0.01

550 1039 83 15 14.05 0.69 4.03 0.39 0.12 0.09 0.01 0.01

1400 1714 144 18 8.07 0.86 7.42 0.40 0.20 0.15 0.08 0.08 0.01 0.01

2175 2849 263 34 0.98 0.44 9.05 0.30 3.64 0.64 2.27 0.50 0.09 0.01

2940 4200 335 51 0.05 0.08 6.86 0.28 7.60 0.47 7.39 0.67 0.10 0.01

300 724 72 10 15.18 0.36 3.17 0.17 0.01 0.02 0.00 0.01

1145 1609 143 19 7.70 0.96 7.01 0.38 0.40 0.23 0.01 0.01 0.00 0.01

1949 2669 240 30 1.11 0.69 7.68 0.57 4.83 1.39 0.09 0.03 0.00 0.01

2760 3299 306 38 0.25 0.29 5.75 0.96 8.52 1.96 0.23 0.08 0.00 0.01

13 Polystyrene 710 1344 15 3 19.22 0.25 1.24 0.11 0.07 0.02 0.07 0.01 0.15 0.01

14 Polystyrene 870 1724 67 12 15.66 0.60 3.70 0.39 0.19 0.03 0.14 0.01 0.53 0.01

280 759 88 16 13.32 0.55 4.47 0.31 0.01 0.02 0.01 0.01

950 1259 160 19 7.65 0.48 7.69 0.28 0.08 0.03 0.06 0.04 0.21 0.01

1475 1999 227 30 3.46 1.05 9.58 0.46 0.45 0.21 0.26 0.15 0.82 0.02

2200 2764 301 46 0.32 0.19 8.42 1.25 2.82 1.01 1.83 2.00 2.38 0.05

2790 3169 377 58 0.02 0.08 7.00 1.08 3.96 0.80 2.91 1.35 4.58 0.09

3390 3734 83 15 13.38 0.51 4.23 0.23 0.05 0.05 0.04 0.03

545 649 358 42 0.14 0.10 10.32 0.54 3.13 0.39 0.05 0.03 7.17 0.14

715 769 309 35 0.26 0.13 12.24 1.26 1.76 0.89 0.03 0.02 2.96 0.06

SS Window

CO2 Rear 

(%)

Natural Gas

1 Natural Gas

2 Natural Gas

3 Natural Gas

THC Rear 

(%)

HRR 

(kW)

O2 Rear 

(%)

Soot Rear 

g/g (%)

CO Rear 

(%)

7 Heptane (pool)

Test # Fuel

6.5

10 Toluene (pool)

11 Ethanol (spray)

4 Heptane (pool)

Natural Gas

5 Heptane (pool)

6

16 Polystyrene

15 Heptane (spray)

12 Methanol (spray)
61

 3.4.1 Gas Chromatography 

The results of the gas chromatography (GC) analysis are given in Table 17 below. All gas 

volume fraction values listed in this table and in the following figures are on a wet basis. The 

measurements listed in the table were multiplied by a factor of 106

. The GC was connected to the 

front probe gas sample line using the outlet of the bypass line on the total hydrocarbon analyzer. 

The front probe location is given in Table 2 and shown in Fig. 4. 

The injection times were manually recorded relative to the fire ignition time. Because the 

turnaround time for the GC sampling was greater than 20 min, a limited number of samples were 

acquired. For comparison with the GC results, the HRR, the total hydrocarbons (FID), and the 

carbon monoxide volume fraction were averaged over a 30 s window ending at the injection 

time. The total hydrocarbon volume fractions determined using the GC (“Total HC, from GC” 

in Table 17) were calculated for comparison with the total hydrocarbon analyzer results. This 

value was determined by summing each of the measured hydrocarbon volume fractions after 

converting to an equivalent methane basis. The results of this comparison are illustrated in 

Fig. 52. 

For most of the GC samples acquired, the two methods agreed within experimental uncertainty. 

This confirms that the compounds occurring in significant amounts were correctly accounted for 

in the GC measurement. The most notable exception to this is the GC sample acquired during 

the natural gas test #3 at t = 2140 s, where the results differed by a factor of 5 for unknown 

reasons. For the polystyrene tests, the total hydrocarbons measurements from the GC were 

significantly lower than from the total HC analyzer. Since the GC method did not include 

compounds above C6, the presence of very large hydrocarbons could explain the differences, 

although this is an unlikely explanation. Future work will revisit this issue. The values in 

Table 17 that are in bold type are below the quantifiable detection limits of the method (the 

uncertainty was greater than the measured value), but were included to show that they were 

identified in trace amounts. The results of the uncertainty analysis of the GC measurements 

(described in Sec. 2.5.7) are show in Table 18. The values represent the combined expanded 

uncertainty with a coverage factor of 2. 

The GC results for the fire tests using natural gas showed, to a large degree, that the 

hydrocarbons measured at the sample location were simply unburned fuel. Figure 53 shows the 

composition of hydrocarbon species for a typical natural gas test. For the liquid and solid fuels 

included in this study, a large number of intermediate hydrocarbon species were quantified. In all 

cases, methane was the largest measured component of hydrocarbon species. Even though the 

GC measured species only as large as C6, the similarity between the total hydrocarbon results 

from the GC and the total hydrocarbon analyzer (see Fig. 52) provides evidence that there were 

no species of significant quantity missed by the GC analysis. These results imply that in the 

upper layer of these compartment fires methane was the most abundant hydrocarbon species, 

higher in concentration than the parent fuel in all cases. For example, Fig. 54 shows the GC 

measurements of the most abundant species (methane, ethyne plus ethene, and benzene) for 

several fire sizes, burning heptane as the fuel. Other species were below the GC detection limits 

(see blank spaces in Table 17). For the fires burning the two alcohol fuels, there is some 

uncertainty with regard to the total hydrocarbon and GC results, since the Nafion filter is known 

to absorb polar organic compounds such as alcohols. 
62

Table 17. Summary of GC sample results. Values in bold were identified as trace species. Blank 

spaces imply that the measurements were below the detection limits. Test # Fuel GC Injection Time (s) HRR (kW) Total HC, FID, (mol/mol)*1e6 Total HC, from GC, (mol/mol)*1e6 Carbon Monoxide, (mol/mol)*1e6 Methane, (mol/mol)*1e6 Ethyne + Ethene, (mol/mol)*1e6 Ethane, (mol/mol)*1e6

Propene, (mol/mol)*1e6Propane, (mol/mol)*1e6Propyne, (mol/mol)*1e61,3-butadiene, (mol/mol)*1e6n-Pentane, (mol/mol)*1e6Benzene, (mol/mol)*1e6

1 Natural Gas 2940 184.3 11881 16590 7785 12058 397 1456 138

970 255.8 21089 19687 16993 423 1164 154

3130 391.4 51188 32685 42152 3330 58 37 62 11 320

4330 175.7 8642 9203 6685 734 29 12 66

910 272.1 45673 30372 19356 23849 2432 83 19 35 4 219

2140 405.5 116515 26511 32587 21920 1733 50 21 32 6 140

3375 176.3 29487 11240 10055 8301 1121 35 18 8 91

1065 429.1 131001 137260 29794 134908 11770 905 216 51 365 64 10 752

2310 272.5 60342 56684 24537 52668 6373 138 50 184 20 496

1150 145.6 300 155 396 -2 70 5

2460 222.9 2967 2599 4363 793 756 5 3 100

3675 341.5 5563 5064 20178 2944 1165 6 89

2840 131.5 580 469 2015 8 30 71

4055 199.5 488 410 1257 53 34 53

1500 142.2 601 825 2054 270 252 9

2740 268.5 20298 20618 33146 8332 5378 19 11 48 5 216

3950 348.0 72784 80041 74807 33342 20725 156 76 134 58 10 679

2025 236.8 870 805 47196 706 50

3255 309.7 1923 2066 77652 1801 132

13 Polystyrene 1110 14.8 745 74 686 -23 48

14 Polystyrene 1155 63.2 1351 748 1777 -4 114 1 86

1565 228.1 4236 4021 7004 1744 995 7 134

3090 396.6 31937 26328 43786 13656 4954 6 43 12 1044

16 Polystyrene 620 355.4 502 309 32227 268 21

3 Natural Gas

2 Natural Gas

6.5 Natural Gas

7 Heptane

10 Toluene

11 Ethanol

12 Methanol

15 Heptane

note: negative values listed in this table are a result of measurement uncertainty associated with the baseline drift 

correction of the quantification method and are not physically meaningful. 
63

Table 18. Summary of GC measurement uncertainty analysis results. Blank spaces imply that 

the measurements were below the detection limits. Test # Fuel GC Injection Time (s) Methane, (mol/mol)*1e6 Ethyne + Ethene, (mol/mol)*1e6 Ethane, (mol/mol)*1e6 Propene, (mol/mol)*1e6 Propane, (mol/mol)*1e6 Propyne, (mol/mol)*1e6

1,3-butadiene, (mol/mol)*1e6n-Pentane, (mol/mol)*1e6Benzene, (mol/mol)*1e6

1 Natural Gas 2940 123 16 64 73

970 152 16 57 72

3130 342 40 67 14 14 36 67

4330 103 15 68 15 76

910 199 28 66 15 14 36 70

2140 185 20 67 15 14 36 73

3375 108 15 67 15 36 75

1065 1114 162 53 13 65 12 35 75 61

2310 427 83 64 14 13 36 63

1150 104 18 78

2460 102 14 15 36 74

3675 100 15 15 75

2840 104 18 75

4055 104 18 76

1500 103 17 78

2740 108 69 68 15 14 36 70

3950 271 293 64 14 14 35 75 61

2025 102 18 

3255 100 18

13 Polystyrene 1110 104 18

14 Polystyrene 1155 104 18 36 75

1565 101 15 15 73

3090 132 63 68 15 14 36 62

16 Polystyrene 620 103 18

12 Methanol

15 Heptane

10 Toluene

11 Ethanol

6.5 Natural Gas

7 Heptane

3 Natural Gas

2 Natural Gas
64

100

1000

10000

100000

1000000

184272405176429273146223342228397131200142268348237310

1563

355

Heat Release Rate (kW)

Wet Volume Concentration, ppm

Total HC (FID)

Total HC (GC)

Natural Gas Heptane Toluene Ethanol Methanol Polystyrene

Figure 52. Comparison of total hydrocarbons measured using the GC and the total hydrocarbon 

analyzer (THC Front), both expressed on a CH4 basis. 

1

10

100

1000

10000

100000

1000000

Methane Ethyne Ethane Propene Propane Propyne 1,3-

butadiene

n-Pentane Benzene

Hydrocabon Species

Wet Volume Concentration, ppm

Figure 53. GC gas species composition measurements in the natural gas fire #65, 

HRR=429 kW. Total Hydrocarbon Volume Fraction *106 Gas Species Volume Fraction *106
65

0

2000

4000

6000

8000

10000

12000

14000

223 342 228 397

HRR (kW)

Wet Volume Concentration, ppm

Methane

Ethyne + Ethene

Benzene

Figure 54 GC measurements of methane, ethyne plus ethene, and benzene in the heptane pool 

fire (test #7) and heptane spray fire (test #15), front gas sample probe. 

 3.5 Optical Soot

Figure 55 shows the time variation of the optical measurement of soot mass concentration for 

test #6, a natural gas fire with a nominal heat release rate that varied from 75 kW to 400 kW. 

Table 19 presents the average optically measured soot mass concentrations in the doorway for 

the four fire sizes considered in test #6. The optical measurement averages were determined 

during the same time periods as the gravimetric measurements. Figure 56 shows the measured 

soot concentration varied from 47 mg/m3

 to 380 mg/m3, following the trend in fire size. 

Figure 57 shows the variation with time of the soot mass concentration for the 20 cm diameter 

polystyrene fire (test #13). For the first 6 min after ignition, the soot mass concentration linearly 

increased as the fire size increased until a steady state was reached. At 500 s after ignition, the 

soot mass concentration was almost constant, with an average value of (700 ± 90) mg/m3

(between 600 s and 1200 s after ignition). As expected, the polystyrene fire was highly sooting, 

with the soot concentration significantly larger than that of natural gas. 

Figure 58 shows the time variation of the soot mass concentration for the 40 cm diameter 

polystyrene fire (test #14). The mass concentration of soot increased as the fire HRR increased. 

After ten minutes, the mass concentration of soot stopped increasing, and a near-steady value 

was reached. The soot mass concentration was approximately (1500 ± 400) mg/m3

 (between 

1115 s and 1230 s after ignition). As expected, the doorway of the 40 cm polystyrene pan fire 

(test #14) had significantly higher soot concentrations than that of the natural gas fire (tests #6) 

and the 22 cm polystyrene pan fire (test #13). 

test #7 test #15 

Gas Species Volume Fraction *106
66

Table 19 also shows the gravimetric measurements made inside the compartment at the front and 

rear locations. Both sets of data show that as the fire became larger and the compartment 

became underventilated, higher soot mass concentrations were present. A comparison of the soot 

mass concentrations at the front and rear of the compartment with the doorway shows that the 

amount inside of the compartment was larger by a factor of 2 to 3 for all but the smallest fire 

sizes. As was true for the larger natural gas fires in test #6, the observed soot mass 

concentrations inside of the compartment in tests #13 and 14 (1140 mg/m3 ± 140 mg/m3

 and 

2700 mg/m3 ± 140 mg/m3

, respectively) were much (a factor of 1.6 to 1.8) larger than in the 

doorway (700 mg/m3 ± 90 mg/m3

 and 1500 mg/m3 ± 400 mg/m3, respectively) during the same 

sampling periods in these tests. The larger concentrations within the compartment may be due to 

air entrainment in the doorway, leading to soot oxidation. 

Table 19. Soot measurements during narrow doorway natural gas fire (test #6).

Soot (mg/m3 Gravimetric ) 

Sampling 

Location 

HRR 

(kW) Gravimetric 

(Compartment)

Optical 

(Doorway) 

rear 35 ± 62 % 

front 

75 

 31 ± 62 % 

47 ± 13 % 

rear 428 ± 7 % 

front 180 

390 ± 7 % 147 ± 17 % 

rear 763 ± 7 % 

front 270 

702 ± 7 % 270 ± 18 % 

rear 766 ± 7 % 

front 

400 

582 ± 7 % 380 ± 17 % 
67

Figure 55. History of the soot mass concentration during natural gas fire (test #6). Mean values 

of the soot mass concentration were determined over the same periods as the gravimetric 

measurements as indicated by the dotted boxes. The fire heat release rate is also shown. 

0

100

200

300

400

0 100 200 300 400

Soot Mass Concentration (mg/m3

)

Heat Release Rate (kW)

Figure 56. Soot mass concentration in the doorway during natural gas fire (test #6) as a function 

of the HRR. 
68

Figure 57. History of the particle mass fraction in the doorway during test #13, polystyrene 

burning in the 20 cm diameter burner. The fire heat release rate is also shown. 

Figure 58. History of the soot mass concentration in the doorway during test #14, polystyrene 

burning in a 40 cm diameter burner. The fire heat release rate is also shown. 

Soot concentration 

Soot concentration 

Heat Release Rate 

Heat Release Rate 
69

 3.6 Heat Fluxes

The heat flux measurements to the floor of the enclosure help to characterize the thermal 

environment within the enclosure and the transient nature of the interior burning. In an enclosure 

with a distributed fuel source the heat flux to the floor can be used to predict the onset of 

flashover. The total heat flux measurement labels are described in Table 20 and their locations 

are listed in Table 2 of Sec. 2.4. Figure 59 shows the front and rear heat flux results for test #15 

with heptane fuel. The front and rear steady state average floor heat flux values for all of the tests 

are summarized in Table 12 and Table 13 and shown graphically in Fig. 60 and Fig. 61. In 

general the heat flux levels were significantly higher for the fuels with high soot yields. In 

addition, for the clean burning fuels the heat flux was fairly constant above heat release rates of 

200 kW. Heat flux levels in excess of 200 kW/m2

 were measured for heptane and toluene fires. 

Although these value are possible based on the measured temperature of the upper layer, they are 

well beyond the calibrated range of the transducer. Furthermore, all of the heat flux 

measurements are somewhat artificial since the gauge temperatures were held constant while the 

floor temperatures were observed to increase significantly. The actual net heat flux to the floor 

(that has been heated by the fire) would be significantly less than the measured heat flux to the 

water-cooled gauge. 

Table 20. Description of interior total heat flux measurement labels.

Measurement Label Description 

12 HFR (kW/m2) Total heat flux to floor at rear of enclosure

13 HFF (kW/m2) Total heat flux to floor at front of enclosure

0 500 1000 1500 2000 2500 3000 3500 4000

0

20

40

60

80

100

120

140

160

180

Time From Ignition (s)

Total Heat Flux to Floor (kW/m2

)

12 HFR (kW/m2)

13 HFF (kW/m2)

Figure 59. Total heat flux gauge measurements at two locations on the interior floor of the RSE 

for test #15 using the spray burner with heptane fuel. 

160 kW 

88 kW 

227 kW 

301 kW 

377 kW 

83 kW 
70

0

50

100

150

200

250

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Front Floor Total Heat Flux (kW/m2)

Natural Gas

Natural Gas (narrow door)

Heptanes (pool burner B)

Heptanes (narrow door)

Heptanes (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 60 Steady state total heat flux measurement at front floor location. note: front heat flux 

gauge was partially blocked by debris for polystyrene test #16. 

0

50

100

150

200

250

300

350

400

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Rear Floor Total Heat Flux (kW/m2)

Natural Gas

Natural Gas (narrow door)

Heptanes (pool burner B)

Heptanes (narrow door)

Heptanes (spray burner C)

Methanol

Ethanol

Toluene

Polystyrene

Figure 61. Steady state total heat flux measurement at rear floor location. 
71

4 Discussion of Results 

 4.1 Fuel and Fire Size Comparisons

Because of the precise metering used, natural gas was the most well controlled fuel used in this 

study. For these reasons, natural gas was used to observe the effects of changes in burner area, 

wall material construction, and doorway ventilation. Natural gas produced the highest quantities 

of total hydrocarbons in this test series. The transition to underventilated burning occurred at a 

lower HRR for natural gas than for other fuels in this study. Depletion of O2 and increased CO 

in the upper layer and flames exiting the doorway are indicators of underventilated burning. The 

reason for this transition occurring at a lower HRR for natural gas can be explained by 

differences in fuel stoichiometry (natural gas fuel requires 1 % to 6 % more oxygen for each unit 

of energy released than the condensed fuels studied) and differences in compartment fire 

structure. 

Like natural gas, liquid heptane fuel was used to study the effects of different burners, door sizes 

and wall material configurations. Heptane (C7H16) fuel (a blend of heptane isomers) was 

selected to represent a moderately sooting liquid alkane. The transition to underventilated 

burning occurred at a greater HRR for heptane than natural gas. Once the enclosure was 

underventilated, the CO and THC volume fractions were consistently lower for heptane than 

natural gas, but the soot concentrations were considerably higher. In general, the gas 

temperatures and floor heat fluxes were greater and more uniform from front to back for the 

heptane fires compared to natural gas fires. 

Liquid toluene (C7H8 - an aromatic hydrocarbon) was included to represent a fuel with a very 

high soot yield. The toluene fire produced very high temperatures, heat fluxes, and soot 

concentrations. Temperatures greater than 1500 K were measured, and some of the stainless 

steel wall pins and aspirated thermocouple probes were melted and destroyed during these tests. 

Test #10 with toluene appeared to be slightly over-ventilated at the largest fire size, made evident 

by the presence of oxygen at the rear sample location. However, at the front sample location, 

oxygen was almost completely depleted and flames were observed outside the doorway. The 

toluene fires yielded relatively low THC and CO volume fractions compared to the other fuels. 

One feature that was unique to the toluene test was the formation of a large soot agglomerate 

(≈10 cm) on the water cooling inlet tube adjacent to the burner (see Fig. 62). 

Liquid methanol (CH3OH) and liquid ethanol (C2H5OH) were chosen as clean burning fuels with 

high combustion efficiency, very low soot yields, and relatively low heats of combustion. The 

temperatures measured inside the enclosure during the tests with alcohol fuels were similar in 

magnitude to the tests with natural gas. As expected, there was no measurable amount of soot 

produced by the methanol fire and the ethanol fire produced soot concentrations similar to the 

natural gas fires. Extremely high volume fractions of CO were observed for the alcohol fuels 

after the fires reached ventilation-limited conditions. CO volume fractions greater than 8 % were 

measured in the front of the enclosure for both of these fuels. 
72

Figure 62. Photograph of soot agglomerate on the water inlet tube of the burner after toluene 

pool fire (test #10). 

Polystyrene pellets were burned in a 60 cm diameter pan (test #16). This test differed from all of 

the other tests, as it involved a solid material in which natural feedback from the fire controlled 

the mass burning rate. A spray of heptane was ignited and was used to initiate burning of the 

polystyrene pellets. 

The character of the measurements during Test 16 are highlighted in Fig. 63 which shows the 

measured CO volume fraction in the exhaust stack and at the front and rear of the upper layer in 

the compartment as a function of time. The data have been time shifted to account for transport 

delays in the exhaust and sample flows. The figure includes photographs at various times during 

the experiment which show the appearance of the fire as seen through the open doorway. At 

both sample locations within the compartment, the CO volume fractions increased as a function 

of time, reached a maximum, and then decreased to near-zero. In the stack, the CO volume 

fraction measurements show that the peak at 600 s which was observed inside of the 

compartment was also reflected in the stack. There were two additional peaks in the stack that 

were not as pronounced within the compartment. 

The photo evidence provides some insight into the CO measurements in the compartment. At 

380 s after ignition, the fire was rather small and flames were restricted to locations immediately 

above the pan. This began to change by 430 s, when the glow of hot gases can be seen in the 

upper reaches of the compartment until at 600 s, the entire doorway appeared to be filled with 

flame and smoke rolled out from the top of the doorway. While the doorway was still luminous 

at 825 s, there were few flames rolling from the doorway. Instead, there was a general glow that 

was observed. This suddenly changed and by 866 s, the fire had become rather small and the 

flames were observed to exist only immediately over the pan. At the same time, a steady flow of 

smoke was transported out from the top part of the compartment. 

10 cm 

Soot 

agglomerate 
73

Observation of the video record does not provide insight into the various extra peaks in the CO 

measurements in the stack. The reasons for the extra peaks are unclear, but the CO 

measurements are not an anomaly since the measured soot mass concentration in the exhaust 

stack tracks the CO peaks observed in the stack. Interestingly, the total hydrocarbon volume 

fraction has a large peak at 350 s that precedes the first large CO and soot peaks at 380 s. The 

simultaneous occurrence of the CO and soot peaks is not unexpected. The timing of these 

concentration peaks are likely related to the compartment fire dynamics. Additional 

experiments are being planned to investigate this further. 

Figure 63 The CO volume fraction measured in the stack and at the front and rear of the 

compartment as a function of time during the burning of 6.0 kg of polystyrene pellets (Test 16). 

Photographs at various times show the fire appearance through the open doorway.

 4.2 Doorway Ventilation Comparisons

Most of the tests conducted during this series used a 48 cm wide by 81 cm high door vent. 

Test #5 (heptane) and test #6 (natural gas) were performed using a 24 cm wide by 81 cm high 

door vent. The most obvious effect of the narrow doorway configuration was that the transition 

to ventilation limited burning occurred at a lower HRR than the full doorway configuration. This 

0.0%

0.5%

1.0%

1.5%

2.0%

2.5%

3.0%

3.5%

4.0%

4.5%

-200 0 200 400 600 800 1000 1200 1400

Time from Ignition (s)

CO Volume Fraction

0.000%

0.005%

0.010%

0.015%

0.020%

0.025%

0.030%

Exhasut Stack CO Volume Fraction

4 CORear (mol/mol)

9 COFront (mol/mol)

60 COstack (mol/mol)

600 s

825 s

866 s

430 s

380 s 
74

transition is evident by the reduced oxygen and increased carbon monoxide volume fractions as 

well as the appearance of flames outside the doorway. The magnitude and front to rear variation 

of the gas temperatures and species volume fractions were similar for the narrow (test #6) and 

full door tests using natural gas. Test #5 with heptane showed an interesting reversal of the 

internal structure of the fires. In contrast to the full door tests, the narrow doorway heptane test 

showed higher temperatures and CO volume fractions at the rear sample location. Because the 

narrow doorway testing was limited to only two fuels, further work is needed to generalize the 

effects of doorway geometry on the compartment fire structure. 

 4.3 Construction Material Comparisons

In Section 2.1.1, the two different construction materials, Marinite I and Kaowool M-board, were 

discussed along with the variation in the techniques used to line the enclosure with each material. 

There were concerns that the different composition of the materials as well as the possibility of 

leaks through additional seams (M-board) would cause differences in fire behavior. It was 

conceivable that the higher organic fraction of the Marinite could lead to additional heat release 

rate and gaseous products of combustion and pyrolysis. It was already known that Marinite 

experiences significant shrinkage when exposed to 1000 °C temperatures for tens of minutes due 

to baking off of organics and water. 

In order to allay any concerns and reveal any noteworthy effects, experiments with the same 

fuels and HRRs were repeated for the enclosure lined with each of the two construction 

materials. For natural gas, the experiments with Marinite were test #1, test #2, and test #3 while 

test #6.5 used M-board. For heptane (using the same burner), the experiment with Marinite was 

test #4 and that with M-board was test #7. 

One potential impact of construction material was the evolution of combustible organic 

components. Kaowool M-board has from 3 % to 6 % organic components. Taking into account 

the 0.23 m3

 volume of the board used to construct the innermost lining, the reported density of 

272 kg/m3

, a generic heat of combustion of 35 MJ/kg, and estimating the organic component at 

4.5 %, there was about 100 MJ of energy available in the M-board layer. In order for this to 

impact the calorimetry, the organic material would have to vaporize and enter into the interior of 

the enclosure where it could be oxidized by enclosure air or by external entrained air as the 

doorway plume rose into the hood. 

To see if the organic components of the M-board added to the HRR, the ratios of the 

calorimetrically measured HRRs to those estimated from metering of the fuel were calculated for 

test #6.5 which was the first test conducted with M-board. This ratio was 1.27 for the initial fire 

size of 76 kW. For the second fire size of 399 kW, the ratio dropped to 1.06. It’s worth noting 

that 27 % of 76 kW and 6 % of 399 kW both result in 20 kW to 25 kW higher calorimeter HRRs 

than the HRRs calculated from the fuel. The third and fourth fire sizes, 269 kW and 179 kW, 

respectively, resulted in ratios of about 1.01. The last fire size of 74 kW had a ratio of 1.15 

representing a calorimeter HRR 11 kW higher than the fuel metered HRR. Since the uncertainty 

on the 3 m calorimeter HRR is on the order of 15 %, these differences between calorimeter and 

fuel metered HRRs are significant, but not easily quantifiable. The conclusion from a review of 

this data is that there was likely additional heat release from evolved organic components of the 

construction material that may have been on the order of 20 kW to 25 kW. The effect was more 
75

significant for early HRRs during the first “bake out” test and diminished by about half by the 

end of the test. 

Another potential impact of the construction materials is on enclosure temperatures. Gas 

temperatures could be affected by additional HRR from bake-out of organics and by the thermal 

conductivity and emissivity of the wall material. Surface temperatures could also be affected by 

the insulating properties, heat capacities, and emissivities of the materials. Table 21 lists each 

material’s properties including density, thermal conductivity, specific heat, and emissivity. The 

M-board had significantly lower thermal conductivity and density than Marinite I. Aspirated 

thermocouples were not deployed for long enough periods in test #6.5 to be useful so only the 

gas temperatures from heptane fire tests #4 and #7 may be compared. Nominal upper layer gas 

temperatures in the rear were consistently about 25 °C higher for the M-board than for the 

Marinite I. Nominal upper layer gas temperatures in the front varied from zero to about 80 °C 

higher for the M-board. No measurable effect was observed on the lower layer gas temperatures. 

The observed correlation between upper layer gas temperature and wall material was relatively 

weak given the relatively large uncertainties in the temperature measurements. 

Examination of the total heat flux to the floor for tests with natural gas and heptane showed the 

nominal heat fluxes for M-board construction were substantially higher (20 kW/m2

 to 30 kW/m2

for HRRs greater than 200 kW) than for Marinite I construction except for the front heat fluxes 

for the natural gas tests which were about the same. The larger heat fluxes and temperatures 

observed for the test using M-board are likely due to better insulating properties compared to 

Marinite I. 

Finally, since mixing and reaction rates and products are influenced by the thermal environment, 

material differences have some potential to affect gas species measurements. However, an 

inspection of O2, CO, CO2 did not reveal any consistent differences related to construction 

material. 

While the impact of construction materials on the enclosure fire dynamics and measurements 

was of concern, the primary reasons for trying alternate materials were durability and ease of 

construction. Table 21 lists some differences between the two construction materials with 

regards to these and other issues. Of the many differences, the most important is that the Mboard survived longer than Marinite I with no cracking and a tolerable amount of warping when 

it was supported by a sufficient number of furnace pins, especially near the seams. The main 

differences in the materials are in shrinkage and water content. 

Figure 64 through Fig. 69 photographically show some of the differences listed in the table. 

Figure 64 shows the Marinite I construction before test #2 (natural gas) after experiencing just 

one fire with a max HRR of 180 kW. The fit is still good, there is no apparent warping, and 

cracking is minimal with some appearing in the upper right corner. Figure 65 shows a similar 

view after test #2, but significant cracking has occurred along with some sagging and warping. 

Figure 66 shows another view after test #6 (natural gas) with even more severe cracking and 

sagging and a gap from a fallen piece of inner lining. 
76

Figure 67 through Fig. 69 show photographs of M-board construction. Figure 67 shows a closeup view of the furnace pins and tight seams before the enclosure was exposed to a fire. 

Figure 68 shows a similar view after the first fire, test #6.5. There is a small amount of warping 

at the wall seam. Figure 69 shows the whole enclosure interior with the back removed. 

Extensive warping and sagging occurred, but not much cracking. The warping and sagging were 

remedied for the construction of the next enclosure lining by increasing the number of furnace 

pins, especially near the seams. Figure 70 depicts an aspirated thermocouple probe and furnace 

pins that have been highly degraded by extremely high heat fluxes and temperatures. The pins 

still held, but were extremely fragile and easily broken in this state. 

Table 21. Comparison of thermal, physical and construction properties of wall lining materials.

Propery Marinite I Kaowool M-board 

Sheet size 1.2 m × 2.4 m (4 ft × 8 ft) 0.9 m × 1.2 m (3 ft × 4 ft) 

Number of 

internal seams 

12 18 

Fastener type Screws, washers Furnace pins 

Fastener durability Survive well Survive with some warping 

Cracking Some after first test, noticeable 

increases with each test 

No cracking 

Warping Increases with each test up to 

sagging of several cm 

Was significant near seams, not 

significant with additional fasteners 

Fragility Still stiff and strong after multiple 

tests although less securely attached 

and susceptible to falling loose 

Fairly soft and fragile even when 

new, should not be handled after 

exposed to fires 

Shrinkage @ 

982 o

C 

13.2 % 2.2 % 

Thermal 

Conductivity, k 

0.11 W/m·K to 0.13 W/m·K (24 °C 

to 538 °C) 

0.06 W/m·K to 0.22 W/m·K 

(0.20 W/ m·K at 1000 °C) 

Density, ρ 737 kg/m3 272 kg/m3

Specific Heat, cp 1172 J/kg·K to 1424 J/kg·K (93 °C to 

425 °C) 

Not specified 

Emissivity, ε 0.74±0.4 0.95 

Organic Content 4 % to 8 % 3 % to 6 % 

Water Content 3 % of dry weight < 0.5 % 

Ease of Cutting Slow, requires saw, produces a lot of 

dust 

Can be cut with a utility knife. 
77

Figure 64. Photo of rear gas sample location and Marinite I construction before test #2. 

Figure 65. Photo of rear gas sample location and Marinite I construction after test #2 with 

natural gas. 
78

Figure 66. Photo inside enclosure after test #6 showing Marinite I condition. Large chunks of 

ceiling fell to the floor during this test.

Figure 67. Photo of rear gas sample location showing Kaowool M-board and furnace pin 

construction prior to test #6.5. 
79

Figure 68. Photo of rear gas sample location after test #6.5 with natural gas using Kaowool MBoard.

Figure 69. Photo of enclosure with rear wall removed after test #10 using toluene. Notice 

melted stainless steel furnace pins and bowed M-board wall but very little cracking of walls. 
80

Figure 70. Photo of rear lower aspirated probe (TR24A) after test #16 with polystyrene. 

Stainless steel furnace pins melted and dripped along the wall. 

 4.4 Burner Type Comparisons

A number of different burner designs were used during this test series for both practical and 

technical reasons. Natural gas was delivered using square gravel-filled burners (A and B), shown 

in Fig. 2. Although the surface areas of these two burners varied by almost a factor of 4, and 

only burner B was water cooled, there was no measurable difference in the fire conditions inside 

the enclosure as a function of HRR. 

During the original test series in the 1990s, heptane was burned in a round pan and the depth of 

the pool was controlled. It was found that steady underventilated burning could not easily be 

attained using this method. In the test described here, liquid fuels were delivered using a square 

water-cooled pool burner with inclined walls (burner B) and a spray nozzle directed into a round 

catch pan (burner C). Burner B performed well in attaining steady burning for the heptane and 

toluene fuels (test #7 and test #10). In these tests, a large amount of radiation from the upper 

layer was imposed on the fuel surface. This additional heat flux was necessary to reach 

underventilated conditions since open burning tests with heptane using burner B resulted in a 

maximum HRR of only about 80 kW (the fuel surface area was 625 cm2

). During test #8 and 

test #9 attempts were made to achieve underventilated fires using methanol and ethanol liquid 

fuels with burner B. These attempts failed because of lack of re-radiation from the upper layer to 

increase the burning rate necessary for ventilation limited conditions. This was due to the 

combined effect of low heat of combustion and lack of soot to act as a radiation source. The 

maximum HRR was 20 kW and the maximum heat flux to the floor was 3 kW/m2

 for the alcohol 
81

fuels using burner B. The spray burner (C) allowed a greater burning area, enabling the 

attainment of underventilated fire conditions for the alcohol fuels in test #11 and test #12. 

Test #15 was conducted with heptane fuel to compare the spray burner performance with the 

results of test #4 and test #7 using the heptane pool burner B. Significant differences were 

observed in the species volume fractions, temperatures and heat fluxes at the same HRR using 

different burners. The spray burner C appeared to be less efficient both locally at the interior 

sampling locations and globally in the exhaust stack. During the 375 kW steady heptane fire in 

test #15, the rear sample location CO, THC and soot measurements were between 3 and 5 times 

greater than the same fire size using the pool burner in test #4. The same trend was observed in 

the exhaust stack species measurements. The interior gas temperatures and heat flux to the floor 

were lower for the spray burner test. One possible explanation for the inefficiency of the spray 

burner is that the fuel was injected higher in the compartment where the differences in 

temperature and oxygen volume fraction could lead to differences in mixing, impacting the 

compartment fire dynamics. This effect should be considered when making comparisons of 

different fuel types in the results presented here. Specifically, the CO volume fractions observed 

in test #11 and test #12 using the alcohols fuels could have been enhanced by the configuration 

of the fuel delivery (burner C). 
82

5 Analysis of Compartment Chemistry

The dynamics and chemistry of the processes occurring in a compartment fire are very complex. 

In an attempt to better understand the chemistry of the compartment fire experiments, gas species 

and soot measurements were made at two locations in the hot upper layer of the compartment 

and far downstream in the exhaust stack, where the temperatures had cooled, typically to less 

than 200 o

C. 

5.1 Mixture Fraction 

It is useful to consider the compartment fire composition measurements in terms of the mixture 

fraction. The use of mixture fraction to analyze flame data was first used by Bilger [39] and later 

modified by Peters [40] and others. The mixture fraction approach has been widely used to 

represent the chemistry in turbulent flame models and fire field models, and has been used to 

analyze the structure of laminar counterflowing and coflowing hydrocarbon and alcohol flames 

[41,42]. 

Pool fires and compartment fires differ from simple laminar flames, as they are typically 

transient and turbulent by nature. Yet, application of the mixture fraction concept to these 

complex combustion situations can provide additional insight into the structure of the fire. The 

mixture fraction approach allows evaluation of a set of species measurements in terms of selfconsistency, and at the same time facilitates rapid assessment of the overall behavior of a 

combustion system. Floyd et al [43] applied the mixture fraction approach to evaluate the 

species composition at various locations in compartment fires. Pitts [17] measured the local 

equivalence ratio at various locations in compartment fires, investigating the possibility of a 

correlation for CO. Since there is a one-to-one correspondence between mixture fraction and 

equivalence ratio, the approach used here is similar to that used previously by Pitts [17] and 

other experimentalists, with the difference that soot is considered in the analysis of mixture 

fraction and local equivalence ratio. 

Sivathanu and Faeth [44] considered the relationship between soot and mixture fraction in an 

effort to improve the understanding associated with radiative emissions from fires. Their 

measurements [44] clearly showed that soot did not correlate well with mixture fraction in 

laminar hydrocarbon diffusion flames. Their data suggest, however, a relationship between soot 

volume fraction and temperature in the fuel rich regions of turbulent hydrocarbon diffusion 

flames. 

In this study, the mixture fraction was used to evaluate the species composition at various 

locations in the hot upper layer of the compartment for a number of reasons. First, the analysis 

provides a check on the quality of the data and provides insight into the chemistry of 

compartment fires. Second, the significance of the inclusion of soot as part of the mixture 

fraction analysis was investigated. The intent of this part of the study is to determine if the 

inclusion of soot adds coherence to the mixture fraction approach. Finally, the importance of 

measurement uncertainty is highlighted, and its value is quantified as part of the mixture fraction 

analysis. 
83

Definition of Mixture Fraction 

The mixture fraction is a non-dimensional quantity representing the mass fraction of a species, at 

a particular location, that was originally part of the fuel stream. The mixture fraction based on 

carbon containing species is defined as follows: 

Soot

F

Soot

co

F

co

co

F

F co x MW

MW Y

x MW

MW Y

x MW

MW Z = Y + Y + +

2

2 (15) 

where MWi is the molecular mass of chemical species i, Yi is the mass fraction of that species, x 

is the number of carbon atoms in the parent fuel molecule (CxH yOz ), MWF is the molecular 

mass of the parent fuel, MWCO is 28 g/mol, MWSoot is taken as 12 g/mol (assuming that soot can 

be approximated as pure carbon), and MWCO2 is 44 g/mol. Alternative definitions of mixture 

fraction yield results similar to those shown below. 

In the experiments reported here, the measurement of total unburned hydrocarbons was made 

using the total hydrocarbon analyzer, reported on an equivalent methane basis. The subscript F in 

the first term of Eq. 15 can be thought of as referring to total hydrocarbons (THC). 

 

In the fire literature, soot is typically not considered in Eq. 15. Here, it is included formally. But 

in the analysis given below, this term is initially neglected, because it is small. Its inclusion is 

important for highly sooting conditions, as will be shown in the results section below. 

 

The mass fraction, Yi, of each species i is determined from the measured volume fraction, Xi, by 

the following expression: 

Yi XiMWi MWtot = / (16) 

MWtot represents the average molecular mass of all gas species and is a function of the local 

composition. 

i

i

MWtot = ∑XiMW (17) 

The state relations can be derived by considering the idealized reaction of a hydrocarbon fuel, 

rewritten here in an expanded form of Eq. 9: 

2 2 2

2 2 2

min(1, )( / 2)H O max(0, 1)( / 4 / 2)O ( / 4)3.76N

C H O ( / 4 / 2)(O 3.76N ) max(0,1 )C H O min(1, ) CO

y x y z x y

x y z x x y z x y z

+ + − + − + +

+ + − + → − +

η η η

η η η (18) 

where the function max(α,β) returns the larger of the two parameters, α or β, and the function 

min(α,β) returns the smaller of the two parameters, α or β. Here, η is a parameter ranging from 

zero (all fuel and zero oxygen) to infinity (all oxygen and zero fuel) and becomes unity for 

stoichiometric conditions. The definition of η shows that it is the reciprocal of the local fuel 

equivalence ratio, φ . 

η

η φ 1

/( / 4 / 2)( 3.76 )

/ ( / 4 / 2)( 3.76 )

( / )

( / )

2 2

2 2 = + − +

+ − + = =

F O N

F O N

st MW x y z MW MW

MW x y z MW MW

F A

F A (19) 
84

where F/A is the fuel-air ratio and the subscript st refers to stoichiometric conditions. The 

idealized mass fractions of products are obtained from the right side of the Eq. 18. For 

stoichiometric conditions, YF = YCO = 0 , and Eq. 15 leads to: 

2

2

CO

F

st CO

x MW

MW Z = Y (20) 

The value of the stoichiometric mixture fraction for the fuels considered in this report is shown 

in Table 22. Its value varied from about 0.0554 for natural gas to 0.1346 for methanol. 

Table 22. Stoichiometric value of the mixture fraction (Zst) for different fuels. 

Fuel Chemical Formula Zst

Methane CH4 0.0552 

Natural Gas 0.93 CH4 + 0.04 C2H6 + 0.01 C3H8 + 0.01 CO2 +... * 0.0554 ± 0.0002 **

n-Heptane C7H16 0.0622 

Toluene C7H8 0.0694 

Polystyrene (C8H8) n 0.0705 

Methanol CH3OH 0.1346 

Ethanol C2H5OH 0.1006 

* typical composition; actual composition varies day to day. 

** average value based on measured natural gas composition. 

A mixture fraction calculation for a methane-air flame is presented here as an example. For 

methane, Eq. 18 becomes: 

2 2

4 2 2 4 2 2

2max(0, 1)O 7.52 N

CH 2 (O 3.76N ) max(0,1 )CH min(1, )CO 2min(1, )H O

η η

η η η η

+ − +

+ + → − + +

 (21) 

The traditional mixture fraction model holds that the mass fraction, Yi, of products can be 

determined through the right side of Eq. 21 as follows: 

N N tot

O O tot

H O H O tot

CO CO tot

CH CH tot

Y MW MW

Y MW MW

Y MW MW

Y MW MW

Y MW MW

2 2

2 2

2 2

2 2

4 4

7.52

2max(0, 1)

2min(1, )

min(1, )

max(0,1 )

η

η

η

η

η

=

= −

=

=

= −

 (22) 

where YCO and YSoot in Eq. 18 were taken as zero for this mixture fraction model calculation. The 

molecular mass of the mixture is a function of the local composition and can be calculated from 

the reactant concentrations: 

2 ( 3.76 ) MWtot MWCH4 MWO2 MWN2 = + η + . (23) 

SinceYCO and YSoot are assumed to be equal to zero and F CH4 Y = Y the mixture fraction defined in 

Eq. 15 can be rewritten as: 

2

4

4 2

CO

CH

CH CO

x MW

MW

Z = Y + Y (24) 
85

Using Eqs. 22 and 23, Eq. 24 can be rewritten as: 

2 ( 3.76 ) 4 2 2

4 4

CH O N

CH

tot

CH

MW MW MW

MW

MW

MW

Z

+ + = = η

 , (25) 

and 

st

O N

CH F A

Z

Z

MW MW

MW

Z

Z ( ) (1 )

2( 3.76 )

(1 )

2 2

4 − = +

− η = . (26) 

Figure 71 presents the relationship between the mixture fraction and the equivalence ratio (1/η) 

as delineated in Eq. 26 for the methane-air system. Under stoichiometric conditions (η = 1), the 

mixture fraction is 0.0552 for a methane-air flame as listed in Table 22. In Fig. 71, natural gas is 

treated as if it were methane. The figure shows that the mixture fraction compresses a large 

range of equivalence ratio values. Figure 72 shows the relationship between the mass fraction 

and the mixture fraction for most of the major species in the methane-air system, when YCO is 

taken as zero. 

Mixture Fraction Uncertainty 

The uncertainty in the mixture fraction is propagated through Eq. 15 and is based on the 

measurement uncertainty of the species concentrations. The positive square root of the estimated 

variance, ( ) UZ Yi , is obtained from 

2

1

2

2

( ) Yi

N

i i

Z i U

Y

Z U Y ∑=

⎥

⎦

⎤ ⎢

⎣

⎡

∂

∂ = (27) 

where Yi U is an estimate of the combined expanded measurement uncertainty of the measured 

mass fraction, Yi, of species i. 
86

mixture fraction

fuel equivalence ratio

0 0.25 0.5 0.75 1 10-1

100

101

102

103

methane

heptane

0.0552 for methane

0.0622 for heptane

Figure 71 The equivalence ratio as a function of mixture fraction for nonpremixed flames 

burning methane and n-heptane.

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

CH4

CO2

O2

H2O

N2 /2

Figure 72 The mass fraction vs. the mixture fraction calculated by the single-parameter mixture 

fraction model.
87

Species Composition Results in terms of Mixture Fraction 

In this section, the time-varying species measurements are presented as a function of the mixture 

fraction. The results are organized in terms of fuel type, since the fuel type establishes the basis 

for the correlation (see Eqs. 9 and 15). 

The species data are considered in terms of the species mass fraction (Yi), which is plotted as a 

function of the local mixture fraction (Z), based on the fuel mass. Measurements from the front 

and rear of the compartment, for all fire conditions (i.e., heat release rate, burner type) and all 

times during the experiment are plotted on a single graph in terms of mixture fraction. The mass 

fractions of H2O and N2 were not measured in the experiments; the values of these species in this 

report (and shown in Fig. 73 through Fig. 96) are estimated from the stoichiometric relation (Eqs. 

9-11). The mass fractions of the unburned hydrocarbons (THC) in each plot were taken from the 

hydrocarbon analyzer measurements. The total hydrocarbons (THC) results were normalized in 

terms of the equivalent fuel molecule for each fuel type. 

The lines in Fig. 73 through Fig. 77, Fig. 80, and Fig. 81) represent, respectively, complete 

stoichiometric combustion and the hypothetical case when only CO2 is produced (no CO or soot; 

see Fig. 72). In some cases, because of the number of data points, the theoretical lines are 

somewhat obscured. The lines on the average steady-state measurement results are easier to 

distinguish as the plots are less crowded. In those plots, the propagated uncertainty is also 

presented. Soot was not measured at all times, but only during the periods when the fire heat 

release rate was quasi- steady. Thus, soot is shown only on the plots labeled “(b)” and is 

presented with the time-averaged gas species results only. The data labeled “THC” in the figures 

represents the total unburned hydrocarbons measured with the FID detector on the total 

hydrocarbon analyzer. 

Natural Gas 

Figure 73 presents all of the gas species measurements taken during all of the natural gas 

experiments (tests #1 - #3, and #6) in both the front and rear of the compartment as a function of 

mixture fraction. Figure 73a shows all of the transient measurements for all of the natural gas 

tests with the full-door configuration (tests #1 - #3 and #6). Figure 73b shows the time-averaged 

steady-state measurements (and represents the same data as shown in Fig. 41). At any single 

location, the mixture fraction can vary from lean to rich, due to the dynamics of the fire. The 

stoichiometric mixture fraction (Zst) is a useful reference point for consideration of fire chemistry 

(see Table 22; Zst = 0.0544). For fuel lean conditions (Z < Zst), the measured mass fractions of 

methane and carbon monoxide are near zero. As the mixture fraction increases, the mass 

fraction of oxygen decreases, and the carbon dioxide and water vapor mass fractions increase. 

For mixture fraction values greater than stoichiometric, the oxygen mass fraction approaches 

zero, whereas the fraction of unburned fuel increases approximately linearly. Under these 

conditions, the generation of carbon monoxide is observed and YCO attains a maximum value of 

about 0.04 g/g. 

As seen in the figure, the hypothetical lines show reasonable agreement with the measurements 

for fuel lean and near-stoichiometric conditions. As the mixture fraction increases beyond 

stoichiometric, however, the difference between the hypothetical lines and the measurements 

becomes considerable. The value of YCO is not negligible for fuel rich conditions. As a result, the 
88

hypothetical lines over-predict the CO2 mass fraction by about 10 % for mixture fraction values 

for Zst < Z < 0.2. As expected, the plots show that the simple traditional mixture faction 

approach does not correlate the experimental results for CO. This behavior is also observed in 

laminar flames, which is attributed to finite rate chemistry effects associated with slow CO 

chemistry [41]. Other approaches to predict CO, possibly using variables that are functions of 

mixture fraction, will need to be considered to improve predictions of its concentration. 

The vertical and horizontal error bars in Fig. 73b represent the combined expanded uncertainty 

of the mass fractions of gas species and the mixture fraction, respectively. For Z < 0.2, the 

uncertainties are relatively small, and the mixture fraction correlations show reasonable 

agreement with the experiments. For Z > 0.2 (in Fig. 73b), the uncertainties in the mixture 

fraction and the mass fraction of gas species, especially unburned fuel and nitrogen, become 

large, and the maximum relative error of mixture fraction reaches values as large as 15 %. 

Typically, the mixture fraction model is within experimental uncertainty for the conditions 

considered in this study, when natural gas was the fuel. The results for H2O are the exception to 

this. Direct measurement of the H2O concentration would be helpful in this regard. Since the H2

concentration can be on the order of a few percent in a hydrocarbon diffusion flame [42], its 

measurement would also be of interest.. 
89

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0.5 0.6 0

0.1

0.2

0.3

0.4

THC

CH4

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0.5 0.6 0

0.1

0.2

0.3

0.4

THC

CH4

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 73 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for the natural gas fire tests #1-#3, and #6: (a) transient measurements and (b) timeaveraged measurements during the period when the HRR was quasi-steady. 
90

Condensed-Phase Hydrocarbon Fuels 

Figure 74 to Fig. 76 show the mass fraction as a function of mixture fraction for the fires burning 

heptane (tests #4, #5, #7, #15), toluene (test #10) and polystyrene (tests #13, #14, #16), 

respectively. For small values of mixture fraction (Z << Zst), the species mass fractions agree 

with the mixture fraction model for all three fuels. Figure 74 shows, however, for nearstoichiometric conditions, YCO is non-zero, which leads to CO2 Y lower than predicted by the 

mixture fraction model. Figure 74b and Fig. 75b show that the measurement uncertainty was 

relatively small for lean mixture fractions. For large values of the mixture fraction, the variance 

of the species mass fraction results was relatively broad. This is particularly true for the transient 

results, but also for the time-averaged results. It is interesting to note that the figures show that 

the local conditions are not fuel-rich for any of the conditions investigated during the toluene and 

the polystyrene tests. Although there was a high concentration of mass fraction results about 

near-stoichiometric conditions, the negligible amounts of hydrocarbons measured during these 

tests led to mixture fraction values which were less than stoichiometric in value. 

Species concentration results in the fire literature, such as those presented in Fig. 74 to Fig. 76, 

are typically reported without consideration of soot in the definition of mixture fraction. It is 

correct to include soot in Eq. 15 as the conserved scalar approach is based on the idea that 

elemental mass is neither created nor destroyed in a fire. The appearance of the plots 

qualitatively change when soot is considered. Figure 77a, Fig. 77b, and Fig. 77c replot the data 

shown in Fig. 74b, Fig. 75b, and Fig. 76b, respectively, and show that the inclusion of soot 

reduces the scatter in the mass fractions for large values of Z, while otherwise leaving the plots 

unchanged. Inclusion of soot stretches the value of Z proportional to the measured soot mass 

fraction in a non-linear manner as illustrated in Fig. 78 using the heptane, toluene and 

polystyrene results. This is because YSoot is negligible for lean conditions, whereas it is 

significant for large values of Z, taking on values as large as 0.1. Neglecting soot for the fires 

burning natural gas and alcohol fuels is reasonable, whereas considering it for heavily sooting 

fires is necessary. The scatter in the mass fractions was reduced for these fuels when soot was 

considered in the definition of Z (see Eq. 15). In Fig. 77a, Fig. 77b, and Fig. 77c, the sum of the 

soot and total hydrocarbons (THC) appears to closely follow the mixture fraction model results. 

The results plotted in this way are particularly convincing in Fig. 77a, where the independent 

results for soot and THC do not follow the state relationship model, but their sum does. 

Interestingly, Fig. 77 shows that there was no significant amount of THC measured in the upper 

layer of the compartment in the toluene or polystyrene fires. The carbon in the upper layer of 

these fires is primarily in the form of CO, CO2 or soot. Examination of Fig. 52 presents the same 

data, which reaffirms that the total HC measurements were relatively small in the polystyrene 

and toluene fires, and that the unburned hydrocarbons did not represent a significant fraction of 

the carbon in the upper layer. 

It is also of interest to examine the results in which soot was not used in the mixture fraction 

definition. This is the typical manner of representing the data in the combustion and fire 

literature. The value of neglecting soot in the mixture fraction definition is seen in the 

comparison of Fig. 74 to Fig. 76 with Fig. 77. For conditions when Z < Zst, the results that 

neglect soot in the mixture fraction definition more closely track the state relationship model 

than the results that consider soot in the mixture fraction definition. The reasons for this are 

unclear, but this result may be important for understanding the success of mixture fraction 
91

correlations without consideration of soot. The comparison of Fig. 74 to Fig. 76 with Fig. 77 

illustrates the importance of including soot in the mixture fraction definition, particularly for 

conditions when Z>Zst. Further investigation of these results may prove useful in the 

development of predictive capabilities for compartment fire species. 

Since CO chemistry is relatively slow compared to many flame processes, it may be reasonable 

to consider the effect of the local temperature on the state relationship model. Figure 79a and 

Fig. 79b shows the transient values of YCO and CO2 Y in the front and rear of the compartment 

(same data as Fig. 74a) as a function of mixture fraction with symbols colored to represent the 

local temperature in the heptane experiments (tests #4, #5, and #7). The results show that while 

there are some general trends associated with temperature, the values of YCO are quite scattered 

and do not systematically correlate with temperature.

Alcohol Fuels 

Figure 80 and Fig. 81 present the mass fractions of products for the methanol (tests #8, #12) and 

ethanol fires (tests #9, #11), respectively. Figure 80 shows the gas species mass fractions at the 

front and rear of the compartment as a function of mixture fraction for the methanol fire tests #8 

and #12. Figure 80a and Fig. 81a shows all of the transient measurements for the methanol and 

ethanol tests, respectively and Fig. 80b and Fig. 81b shows the averaged quasi-steady 

measurements for the methanol and ethanol tests, respectively. The soot measurements are 

plotted with the quasi-steady results, but were not measured and are not shown as part of the 

transient results. Compared to the mixture fraction model, the concentrations of water vapor are 

too large, while the unburned fuel and carbon dioxide fall below the predictions. For fires 

burning both of the alcohol fuels, the value of YCO is as high as 10 %, which is over two times its 

value in the natural gas and heptane fires. 
92

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0.5 0

0.1

0.2

0.3

0.4

THC

C7H16

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0.5 0

0.1

0.2

0.3

0.4

THC

C7H16

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 74 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for the heptane fire tests #4, #5, #7 and #15: (a) transient measurements and (b) timeaveraged measurements during the period when the HRR was quasi-steady. 
93

. mixture fraction mass fraction

0 0.025 0.05 0.075 0.1 0

0.1

0.2

0.3

0.4

THC

C7H8

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.025 0.05 0.075 0.1 0

0.1

0.2

0.3

0.4

THC

C7H8

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 75 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for the toluene fire test #10: (a) transient measurements and (b) time-averaged 

measurements during the period when the HRR was quasi-steady. 
94

mixture fraction

mass fraction

0 0.03 0.06 0.09 0.12 0

0.1

0.2

0.3

0.4

THC

(C8H8)n

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.03 0.06 0.09 0.12 0

0.1

0.2

0.3

0.4

THC

(C8H8)n

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 76 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for polystyrene fire tests #13, #14, and #16: (a) transient measurements and (b) timeaveraged measurements during the period when the HRR was quasi-steady. 
95

mixture fraction

mass fraction

0 0.1 0.2 0.3 0

0.1

0.2

0.3

0.4

THC

C7H16

CO2

CO2

O2

O2

H2O

H2O

N2/ 2

N2/ 2

CO

THC+soot

soot

 mixture fraction mass fraction

0 0.03 0.06 0.09 0.12 0.15 0

0.1

0.2

0.3

0.4

THC

C7H8

CO2

CO2

O2

O2

H2O

H2O

N2/ 2

N2/ 2

CO

THC+soot

soot

 (a) (b) 

mixture fraction

mass fraction

0 0.1 0.2 0.3 0

0.1

0.2

0.3

0.4

THC

(C8H8)n

CO2

CO2

O2

O2

H2O

H2O

N2/ 2

N2/ 2

CO

THC+soot

soot

(c) 

Figure 77 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction (with soot included) for the time-averaged measurements when the HRR was quasisteady in the experiments burning (a) heptane, (b) toluene, and (c) polystyrene. 
96

Z (without soot)

Z (with soot)

0 0.05 0.1 0.15 0.2 0.25 0

0.05

0.1

0.15

0.2

0.25

Heptane

Toluene

Polystyrene

Figure 78 Comparison of mixture fractions calculated with and without soot using the timeaveraged species measurements when the HRR was quasi-steady in the experiments burning 

(a) heptane, (b) toluene, and (c) polystyrene. 

.
97

mixture fraction

mass fraction of CO2

0 0.1 0.2 0.3 0.4 0

0.05

0.1

0.15

0.2

1200

1100

1000

900

800

700

600

500

400

300

200

100

Temp(oC)

(a) 

mixture fraction

mass fraction of CO

0 0.1 0.2 0.3 0.4 0

0.01

0.02

0.03

0.04

1200

1100

1000

900

800

700

600

500

400

300

200

100

Temp(oC)

(b) 

Figure 79 Transient values of (a) CO2 Y and (b) YCO in the front and rear of the compartment as a 

function of the mixture fraction (without soot) with symbols colored to represent the local 

temperature for the heptane fire tests #4, #5, and #7. 
98

mixture fraction

mass fraction

0 0.05 0.1 0.15 0.2 0.25 0.3 0

0.1

0.2

0.3

0.4

THC

CH3OH

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.1 0.2 0.3 0

0.1

0.2

0.3

0.4

THC

CH3OH

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 80 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for methanol fire tests #8 and #12: (a) transient measurements and (b) time-averaged 

measurements during the period when the HRR was quasi-steady. 
99

mixture fraction

mass fraction

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0

0.1

0.2

0.3

0.4

THC

C2H5OH

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(a) 

mixture fraction

mass fraction

0 0.1 0.2 0.3 0.4 0

0.1

0.2

0.3

0.4

THC

C2H5OH

CO2

CO2

O2

O2

H2O

H2O

N2 /2

N2 /2

CO

(b) 

Figure 81 Mass fractions of front and rear compartment gas species as a function of mixture 

fraction for ethanol tests #9 and #11: (a) transient measurements and (b) time-averaged 

measurements during the period when the HRR was quasi-steady. 
100

5.2 Chemical Equilibrium 

This section is complementary to the previous section, investigating aspects of the species 

concentrations in the compartment fire. Here, the measurements are used to test the hypothesis 

that local chemical equilibrium exists within the compartment. As an example, results are 

presented for natural gas. Although this type of analysis has been previously considered (e.g., 

Sivathanu and Faeth, [41] ), it is of interest to confirm that the measurements conducted in this 

study are consistent with previous results. 

The time-averaged data taken during the period when the HRR was quasi-steady shown in 

Fig. 73 are considered in this analysis. Local chemical equilibrium was calculated using the 

measured species (unburned hydrocarbons, CO2, CO, O2, N2, H2O, and soot) and temperature 

data as input. The STANJAN software [45] was used to calculate the products of local 

thermodynamic equilibrium. Product species that were considered included CH4, C2H6, C3H8, 

CO2, CO, O2, N2, H2O, soot and H2. Figure 82a and Fig. 82b compare the equilibrium results 

(filled symbols) with the measured mass fractions (open symbols) for the major species, e.g. 

CH4, CO2, CO, O2, H2O, and soot. Because of the dependency of equilibrium on temperature, the 

calculated results show significant scatter at larger mixture fractions. Like the mixture fraction 

correlation, the calculations show good agreement with the measurements for fuel lean 

conditions, but significant differences exist for fuel rich conditions - consistent with previous 

results [41]. In particular, predictions of CO are entirely inaccurate using this approach. In the 

following sections, compartment chemistry is investigated using other methods. 
101

mixture fraction

mass fraction

0 0.05 0.1 0.15 0.2 0.25 0.3 0

0.05

0.1

0.15

0.2

0.25

0.3

CH4

CO2

Measurements

Equilibrium cal.

CO

THC

CO2

CO

(a) 

mixture fraction

mass fraction

0 0.05 0.1 0.15 0.2 0.25 0.3 0

0.05

0.1

0.15

0.2

0.25

0.3

O2

O2

H2O

H2O

Measurements

Equilibrium cal.

soot

soot

(b) 

Figure 82 Comparison between the equilibrium calculations and the time-averaged species mass 

fractions measured in the front and rear of the compartment as a function of the mixture fraction 

during the period when the HRR was quasi-steady in the natural gas fire tests #1-#3, and #6.
102

5.3 Carbon Balance

The compartment measurements show that elemental carbon was partitioned among soot and 

three principal gaseous species (CO2, CO, and CH4) in the upper layer of the compartment. 

Other hydrocarbons were measured in only trace quantities compared to methane. The 

fractional mass-based amount of carbon that existed in the form of carbon monoxide (FCO) or 

carbonaceous soot (Fsoot) is related to the mass fractions of carbon containing species at each 

measurement location as: 

;

28

12

44

12

16

12

4 2 CH CO CO soot

soot

soot

Y Y Y Y

Y F

+ + +

= 

CH CO CO soot

CO

CO

Y Y Y Y

Y

F

+ + +

=

28

12

44

12

16

12

28

12

4 2

 (28) 

Typically, composition results are presented in the form of product yields or generation rates 

(defined below), rather than simply fractional mass-based amounts. There are some advantages, 

however, to examining the data in this form, as the values of Fi are bounded from 0 to 1. In the 

results presented for the compartment data, the value of Xs, which is a representation of the 

amount of carbonaceous soot is defined as: 

= ∑

i

i

c

soot

s MW

Y

MW

Y

X (29) 

which comes directly from algebraic manipulation of Eqs. 16 and 17, and the facts that 

∑ Xi

MWi is a constant, and ∑ Xi = 1. 

Table 23 lists Fsoot and FCO based on averages of the quasi-steady species measurements at the 

front and rear locations in the heptane, toluene and polystyrene fires. For convenience, the fire 

heat release rate (HRR), the local equivalence ratio (φ) and the ratio (FCO/Fsoot) are also included 

in the table. The value of Fsoot was different for the different fuels, tending to increase with the 

local equivalence ratio (or mixture faction.) The Fsoot was largest for the polystyrene fires (see 

Table 23), reaching a value of 0.66. The Fsoot in the other fires was also large, taking on values 

as large as 0.29 in one of the toluene fires and 0.45 in one of the heptane fires. The value of FCO

was as large as 0.03 to 0.1 for the polystyrene fires, 0.03 to 0.04 for the toluene fires, and 0.02 to 

0.14 for the heptane fires.

Table 23 lists value of FCO /Fsoot, which depends on fuel type, and physical location. Its value 

was less than 0.4, except in one case. 

Measurements by Koylu et al. [46] and Santoro and co-workers [47] showed that there is a linear 

relation in the emission of soot and CO from buoyant turbulent diffusion flames burning various 

hydrocarbon fuels (acetylene, propene, etc.). Measurements in the fuel lean (overfire) plume 

region of hydrocarbon fires showed that the soot and CO generation factors (ηS and ηCO) tended 

to increase with flame residence time, until a near-constant value was reached after long times 

(compared to the smoke point). Koylu et al. [46] reported that the ratio of the CO and soot 

generation factors for a range of fuel types was such that, ηCO /ηS = 0.34 ± 0.09. The generation 

rate was defined as the mass of soot (or gas species) produced per unit mass of fuel carbon 

consumed. This is slightly different than the soot (or gas species) yield (yCO and y S), which is 

based on the mass of all elements (not just carbon) in the fuel stream. The ratios of the yields 
103

and the generation rates, however, are equal, and their value can be determined at any location 

from the ratio of the mass fractions of CO and soot: 

ηCO / ηS = CO soot y / y = YCO /Ysoot = (7/3) FCO/Fsoot (30) 

The constant value (7/3) in Eq. 30 is the ratio of the total CO mass to the mass of carbon. 

Table 24 lists CO y , soot y , and the ratio CO soot y / y based on the time-averaged species 

measurements at the front and rear compartment locations and the stack when the heat release 

rate was quasi-steady during each of the fires (natural gas, methanol, ethanol, heptane, toluene 

and polystyrene). The fire heat release rate (HRR) and the local equivalence ratio (φ) are also 

listed. Much of the same data was used as in Table 23. 

Figure 83a and Fig. 83b show the yield of CO and soot as a function of the local equivalence 

ratio for the heptane, toluene and polystyrene fires. Straight lines seem to be reasonable fits to 

the data, intersecting the intercept at a non-zero value. Figure 84 is analogous to Fig. 83, with 

the parameters Fsoot and FCO considered in lieu of CO y and soot y . The trends and values of the 

data shown in the graphs are very similar in appearance, consistent with the data presented in

Table 23 and Fsoot and FCO is a reasonable way to represent the results and understand that in 

the toluene and polystyrene fires, almost twice as much carbon exists in the form of soot as 

compared to CO. The heptane data is much more scattered, making any sort of authoritative 

generality less convincing. In the richest polystyrene fire (φ ≈ 2.5), almost 80 % of the carbon 

exists in the form of CO or soot, with relatively little carbon in the form of CO2. 

Figure 85a and b shows the yield of CO and soot as a function of the local equivalence ratio for 

the natural gas and ethanol fires. The values of CO y and soot y were relatively low, as compared 

to the results for the smoky fuels presented in Fig. 83. 

Figure 86 shows the CO yield as a function of the soot yield for the same quasi-steady data 

shown in Fig. 83 for the heptane, toluene and polystyrene fires. Also shown is a line 

representing the results of Koylu et al [46]. Koylu reported about 30 % scatter in the ratio of the 

yields of CO to soot, which is considerably smaller than that seen in the figure. Nevertheless, 

more data are needed to examine this relationship in the upper layer of compartment fires. It is 

interesting to note that Tewarson et al [48] reported that the ratio of the CO and soot generation 

efficiencies from small fires burning polymers varied, depending on the exact fuel type and the 

amount of ventilation. 

Figure 87 shows the ratio of the CO yield to the soot yield as a function of the local equivalence 

ratio for the same quasi-steady data shown in Fig. 83 and Fig. 86 for the heptane, toluene and 

polystyrene fires. Best fit lines for each fuel type are different, and highlight the trends in the 

CO soot y / y data. 

Figure 88 shows the CO yield as a function of the soot yield for the same quasi-steady data 

shown in Fig. 85 for the natural gas and ethanol fires. Results for methanol are not shown, 

because the measured soot yield was zero. Also shown in the figure is a line representing the 
104

results reported by Koylu et al [46] for rather heavy hydrocarbons, including propane, and 

propylene. The Koylu results do not agree with the current set of data for these non-smoky fuels. 

Figure 89 shows the ratio of the CO yield to the soot yield as a function of the local equivalence 

ratio for the same quasi-steady data shown in Fig. 85 for the natural gas and ethanol fires. Both 

sets of data could be fit by straight lines. The figure shows that a horizontal line could 

adequately represent the natural gas results for rich conditions (φ>1). The ethanol results, 

however, show a finite slope, which is unique among the fuels tested (also see Fig. 87). 

Experiments over a range of ventilation conditions in full-scale are planned to further investigate 

the consistency and repeatability of the trends in the CO and soot measurement results. 

Table 23. Average fractional soot, CO and CO/Soot ratio at the front and rear compartment 

measurement locations. 

Rear Front 

Fuel HRR φlocal FCO Fsoot FCO Fsoot / φlocal FCO Fsoot FCO Fsoot /

140 0.957 0.047 0.049 0.959 1.463 0.019 0.039 0.49 

148 0.563 0.01 0.035 0.280 0.581 0.015 0.041 0.36 

153 0.686 0.016 0.031 0.532 0.660 0.013 0.034 0.37 

160 0.547 0.013 0.066 0.205 0.677 0.022 0.058 0.38 

221 1.214 0.321 0.168 1.911 2.024 0.078 0.263 0.30 

227 0.669 0.039 0.094 0.421 1.080 0.101 0.142 0.72 

246 0.650 0.017 0.064 0.267 1.169 0.076 0.116 0.65 

269 0.730 0.012 0.043 0.288 1.922 0.161 0.089 1.8 

301 1.366 0.103 0.170 0.603 2.407 0.22 0.195 1.13 

341 1.249 0.143 0.211 0.678 1.636 0.22 0.236 0.93 

375 1.163 0.101 0.162 0.622 1.855 0.187 0.257 0.73 

heptane 

377 2.281 0.138 0.238 0.578 3.469 0.216 0.267 0.81 

138 0.540 0.044 0.190 0.230 0.741 0.04 0.195 0.21 

toluene 202 0.599 0.009 0.117 0.077 0.981 0.022 0.179 0.13 

295 0.989 0.026 0.206 0.127 1.244 0.061 0.291 0.21 

15 0.137 0.074 0.149 0.498 0.133 0.071 0.157 0.45 

67 0.394 0.075 0.216 0.347 0.374 0.071 0.205 0.35 

309 2.017 0.195 0.583 0.335 1.307 0.191 0.337 0.57 polystyrene 

358 2.391 0.231 0.650 0.355 1.884 0.234 0.555 0.42 
105

local equivalence ratio

yCO

01234 0

0.1

0.2

0.3

0.4

0.5

heptane

toluene

polystyrene

(a) 

local equivalence ratio

ysoot

01234 0

0.2

0.4

0.6

0.8

1

heptane

toluene

polystyrene

(b) 

Figure 83 The CO and soot yields as a function of the local equivalence ratio for the timeaveraged measurements during the period when the HRR was quasi-steady in the heptane, 

toluene and polystyrene fires.
106

local equivalence ratio

Fsoot

01234 0

0.2

0.4

0.6

0.8

1

heptane

toluene

polystyrene

(a) 

local equivalence ratio

FCO

01234 0

0.05

0.1

0.15

0.2

heptane

toluene

polystyrene

(b) 

Figure 84. The values of FCO and FSoot as a function of the local equivalence ratio during the 

period when the HRR was quasi-steady in the heptane, toluene and polystyrene fires. 
107

local equivalence ratio

01234567 0

0.1

0.2

0.3

0.4

0.5

0.6

Natural gas

Methanol

Ethanol

yCO

(a) 

local equivalence ratio

ysoot

01234567 0

0.02

0.04

0.06

0.08

0.1

Natural gas

Ethanol

(b) 

Figure 85 The CO and soot yields as a function of the local equivalence ratio during the period 

when the HRR was quasi-steady in the natural gas and ethanol fires.
108

ysoot

yCO

10-2 10-1 100 10 10 1

-3

10-2

10-1

100

101

heptane

toluene

polystyrene

yCO ysoot = 0.34 ± 0.09

Figure 86. The CO yield as a function of the soot yield during the period when the HRR was 

quasi-steady in the heptane, toluene and polystyrene fires. Also shown is a line representing the 

results of Koylu [46].

= 0.34 ± 0.09 CO soot y y

local equivalence ratio

yCO/ysoot

01234 0

0.5

1

1.5

2

2.5

heptane

toluene

polystyrene

Ave.=0.673

Ave.=0.162

Ave.=0.415

Figure 87. The ratio of the CO to soot yield as a function of the local equivalence ration during 

the period when the HRR was quasi-steady in the heptane, toluene and polystyrene fires. Best fit 

lines to the data and a line representing the results of Koylu [46] are also shown.
109

y

y

10-4 10-3 10-2 10-1 10 10 0

-4

10-3

10-2

10-1

100

101

Natural gas

Ethanol

soot

CO

yCO ysoot = 0.34 ± 0.09

Figure 88 The CO yield as a function of the soot yield for the same data shown in Fig. 85 

during the period when the HRR was quasi-steady in the natural gas and ethanol fires. 

local equivalence ratio

y /y

01234567 -20

0

20

40

60

80

100

Natural gas

Ethanol

soot CO

yCO ysoot = 0.34 ± 0.09

Figure 89 The ratio of the CO to soot yield as a function of the local equivalence ration for the 

same data shown in Fig. 85 during the period when the HRR was quasi-steady in the natural gas 

and ethanol fires.
110

 Table 24. Time-averaged yields of soot, CO, and the ratio (yCO /ys) at the front and rear 

compartment measurement locations and in the exhaust stack for all fuel types. 

HRR Rear Front Stack 

Fuel [kW] φlocal yco ys yco/ys φlocal yco ys yco/ys yco ys yco/ys

74 0.686 0.005 0.002 2.126 0.579 0.005 0.003 1.933 - - - 

74 0.460 0.004 0.002 1.802 0.373 0.009 0.005 1.955 - - - 

75 0.559 0.035 0.043 0.817 0.436 0.027 0.073 0.373 - - - 

85 0.452 0.026 - - 0.407 0.015 - - 0.067 0.000 - 

97 0.565 0.058 - - 0.481 0.036 - - 0.161 0.000 - 

179 1.239 0.117 0.011 10.35 1.188 0.141 0.014 10.14 - - - 

181 0.980 0.068 - - 1.070 0.192 - - 0.018 0.000 

186 1.368 0.169 0.023 7.396 1.069 0.148 0.041 3.588 - - 

265 1.891 0.179 0.015 12.30 1.753 0.221 0.018 12.01 - - - 

272 2.849 0.162 0.012 13.58 4.164 0.133 0.012 11.51 - - - 

273 1.372 0.191 - - 1.720 0.279 - - 0.016 0.000 - 

417 4.681 0.125 0.008 15.45 6.125 0.109 0.007 16.07 - - - 

Natural gas 

424 2.025 0.217 0.018 11.86 2.679 0.235 0.015 16.03 0.024 0.000 

83 0.383 0.015 - - 0.349 0.022 - - 0.009 0.044 0.212 

88 0.394 0.007 - - 0.347 0.005 - - 0.006 0.029 0.212 

140 0.957 0.043 0.044 0.957 1.463 0.023 0.047 0.491 - - - 

148 0.563 0.008 0.030 0.280 0.581 0.013 0.036 0.359 0.007 0.030 0.234 

153 0.686 0.014 0.026 0.531 0.660 0.011 0.031 0.366 - - - 

160 0.547 0.011 0.056 0.204 0.677 0.020 0.051 0.382 0.010 0.056 0.174 

221 1.216 0.294 0.156 1.887 2.038 0.088 0.305 0.289 - - - 

227 0.670 0.015 0.081 0.183 1.081 0.072 0.136 0.531 0.021 0.118 0.176 

246 0.651 0.035 0.057 0.610 1.170 0.102 0.118 0.864 0.022 0.095 0.237 

269 0.730 0.011 0.037 0.287 1.923 0.198 0.111 1.783 - - - 

301 1.369 0.101 0.170 0.593 2.417 0.290 0.264 1.100 0.052 0.230 0.228 

341 1.253 0.139 0.209 0.667 1.644 0.226 0.248 0.912 0.051 0.197 0.257 

375 1.165 0.096 0.156 0.613 1.866 0.196 0.277 0.708 - - - 

Heptane 

377 2.294 0.165 0.294 0.562 3.508 0.300 0.389 0.772 0.064 0.240 0.265 

49 0.312 0.115 - - 0.275 0.124 - - 0.098 0.148 0.662 

138 0.541 0.040 0.176 0.229 0.742 0.037 0.185 0.203 0.045 0.146 0.305 

202 0.600 0.008 0.110 0.076 0.983 0.021 0.169 0.124 0.013 0.154 0.084 

295 0.991 0.024 0.191 0.125 1.252 0.056 0.273 0.204 0.011 0.127 0.084 

Toluene 

339 0.861 0.104 - - 0.936 0.200 - - 0.023 0.171 0.136 

15 0.137 0.089 0.179 0.497 0.133 0.087 0.194 0.452 0.072 0.176 0.410 

67 0.395 0.080 0.232 0.345 0.374 0.078 0.226 0.343 0.070 0.253 0.277 

309 2.117 0.173 0.561 0.308 1.319 0.176 0.320 0.550 0.017 0.133 0.128 Polystyrene 

358 2.572 0.200 0.631 0.317 1.962 0.210 0.536 0.391 0.044 0.249 0.178 

143 0.592 0.002 0.000 - 0.598 0.048 0.000 - 0.000 0.000 - 

Methanol 240 1.137 0.206 0.000 - 1.151 0.336 0.000 - 0.006 0.000 - 

306 1.414 0.464 0.000 - 1.411 0.515 0.000 - 0.011 0.000 - 

144 0.565 0.003 0.001 2.288 0.610 0.031 0.002 17.71 0.004 0.000 - 

Ethanol 263 1.014 0.056 0.003 20.68 1.739 0.292 0.007 40.68 0.004 0.000 - 

335 1.927 0.272 0.004 67.13 3.850 0.418 0.005 78.05 0.008 0.000 - 
111

5.4 Post-Compartment Product Yields 

It is useful to consider the product yields downstream from the fire compartment in the exhaust 

stack, and to compare these results to conditions in the compartment. These comparisons 

highlight the effects of a compartment on fire chemistry. 

The yields of CO2, CO, soot, and hydrocarbons determined from the measurements made in the 

exhaust hood during the quasi-steady burning periods for each of the fuels tested are shown in 

Figure 90 through Fig. 93 as a function of the fire heat release rate. The largest yield in the 

stack was for CO2 (125 % to 300 %), followed by soot (< 25 %), CO (< 10 %), and total 

hydrocarbons (< 3.5 %). The CO2 yield in the stack shown in Fig. 90 was related to the 

stoichiometry (Eq. 9) and the combustion efficiency. The CO yield in the stack shown in Fig. 91 

appeared to be a function of fuel type and fire size. Some fuels exhibited high CO yields for 

lower heat release rates (toluene, polystyrene, ethanol, natural gas) and other fuels exhibited 

relatively higher CO yields (heptane, methanol) for higher heat release rates. The results for the 

soot yield shown in Fig. 92 were also dependent on the fuel type and the HRR, whereas some 

fuels (natural gas and methanol) produced absolutely negligible amounts of soot. The soot yield 

was as large as 15 % to 25 % for the heptane, toluene and polystyrene fires. 

The local product species yields in the rear and front of the compartment were discussed 

previously. Table 24 lists the CO and soot yields in the smoke-laden (heptane, toluene and 

polystyrene) and non-smoky fires (natural gas and ethanol) as a function of the fire size, both in 

the compartment and in the stack. The data in the Table are also plotted in Fig. 83 and Fig. 85. 

For the smoky fires, the yield of soot was almost always larger in the compartment than in the 

stack. For example, the yield of soot in the toluene fire varied from 0.15 to 0.17 in the stack as 

compared to values as high as 0.27 in the compartment. The polystyrene fire was the smokiest 

and the yield of soot varied from 0.13 to 0.25 in the stack as compared to 0.18 to 0.61 in the 

compartment. The yield of CO was generally, but not always larger in the compartment than in 

the stack, and typically a factor of 1.5 to 7 times larger depending on the fuel type. The yield of 

CO in the polystyrene fire varied from 0.02 to 0.07 in the stack as compared to 0.08 to 0.26 in 

the compartment. The yield of CO in the ethanol fire varied from approximately 0.004 to 0.07 in 

the stack as compared to 0.003 to 0.42 in the compartment. 
112

1.00

1.50

2.00

2.50

3.00

3.50

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Carbon Dioxide Yield (g/g)

Natural Gas Methanol Ethanol Heptane Toluene Polystyrene

Figure 90 The CO2 yield in the exhaust stack as a function of the fire heat release rate during the 

periods when the HRR was quasi-steady for each of the fuels tested. 

0.00

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0.09

0.10

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Carbon Monoxide Yield (g/g)

Natural Gas Methanol Ethanol Heptane Toluene Polystyrene

Figure 91. The CO yield in the exhaust stack as a function of the fire heat release rate during the 

periods when the HRR was quasi-steady for each of the fuels tested. 
113

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Soot Yield (g/g)

Natural Gas Methanol Ethanol Heptane Toluene Polystyrene

Figure 92. The soot yield in the exhaust stack as a function of the fire heat release rate during the 

periods when the HRR was quasi-steady for each of the fuels tested. 

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0.035

0.040

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Total Hydrocabon (as methane) Yield (g/g)

Natural Gas Methanol Ethanol Heptane Toluene Polystyrene

Figure 93. The total hydrocarbon yield in the exhaust stack as a function of the fire heat release 

rate during the periods when the HRR was quasi-steady for each of the fuels tested. 
114

5.5 Combustion Efficiency 

To better understand the compartment chemistry, it is of interest to determine the combustion 

efficiency both in the exhaust stack and at various locations in the upper layer of the 

compartment. The combustion efficiency (χa) is a global representation of the fractional amount 

of heat released by the fire as compared to complete combustion. It is defined as: 

c,ideal

c

H

H

Δ

Δ

χ a = (31) 

where ΔHc,ideal is the net heat of complete combustion based on the conversion of all carbon and 

hydrogen in the fuel to CO2 and H2O (assumed to remain in the vapor phase) and ΔHc is the net 

heat of combustion, which is the actual heat released in a chemical reaction. The value of χa is 

bounded by 0 % and 100 %. 

 

Using the nomenclature defined in Eq. 9 for the stoichiometry of the combustion reaction, the 

value of ΔHc,ideal is: 

o o o

c,ideal f CO , fHH O, fHfuel,( , , ) 2

H H 2 2 gas gas gas liquid solid

y Δ = xΔ + Δ − Δ (32) 

where o ΔfHi,state is the heat of formation of species i at a given state. The heats of formation of 

CO2 and H2O are given in Ref. [49]. The value of ΔHc , the net heat of combustion is given by: 

o o

o o o o

f H O, f fuel,( , , )

c f CO , f CO, f C, f CH ,

f H H

H b H c H e H H

2

2 4

gas gas liquid solid

gas gas solid d gas

+ Δ − Δ

Δ = Δ + Δ + Δ + Δ

 (33) 

where the coefficients b – f represent the amount of molecular products in the general 

combustion reaction defined in Eq. 9, and the heats of formation of CO, soot, and CH4 are given 

by Ref. [49] . The molecular product yield coefficients are given in Eq. 34 below. 

( ) ( )

( ) ( ) 2d

2 , fX X X X

X , d

X X X X

X

e

X X X X

X , c

X X X X

X

b

2 4

4

2 4

2 4 2 4

2

CO CO C CH

CH

CO CO C CH

C

CO CO C CH

CO

CO CO C CH

CO

= − + + + = + + + =

+ + + = + + + =

x x y

x x

 (34) 

Figure 94 shows the combustion efficiency and its uncertainty in the exhaust stack using 

measurements made during the steady burning periods as a function of the fire heat release rate 

for each of the fuels tested. The value of χa was nearly 100 % for all conditions in the methanol, 

ethanol and natural gas fires, whereas its value was smaller in the fires with the largest soot 

yields (heptane, toluene and polystyrene). The value of χa in the exhaust stack was as low as 

75 % during the polystyrene fire. At the same time, the soot yield was nearly 0.25. The 

expanded relative uncertainty of χa varied from 0.3 % to 6 %. If χa was determined through the 

ratio of the measured heat release rate to the measured mass delivery rate, then the expanded 

relative uncertainty was larger than 15 % and was not a function of the value of χa. 

Figure 95 shows the combustion efficiency in the rear and front compartment sampling locations 

as a function of the measured fire heat release rate during the steady burning periods for the three 
115

smokiest fuels (heptane, toluene and polystyrene). There was a large amount of scatter in the 

results. In general, the value of χa tended to decrease with increasing values of the HRR. During 

one of the larger polystyrene fires, the value of χa was as low as 45 % in the compartment. For 

the same fire, the value of χa was much larger in the exhaust stack, reaching almost 80 %. The 

value of χa for the heptane fires varied between 50 % and 100 %. The scatter in the heptane 

results was rather large. As expected, the value of χa tended to be higher when the oxygen 

volume fraction was larger as seen in Fig. 96. A comparison of Fig. 94 and Fig. 95 shows that 

the value of χa inside the compartment was typically, but not always, equal to or smaller than the 

value of χa in the stack, consistent with the idea that incomplete products of combustion continue 

to oxidize once they exit the compartment and are exposed to air. 

70%

75%

80%

85%

90%

95%

100%

105%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Combustion Efficiency

Natural Gas Methanol Ethanol Heptane Toluene Polystyrene

Figure 94. The combustion efficiency in the exhaust stack as a function of the fire HRR.
116

40%

50%

60%

70%

80%

90%

100%

110%

0 50 100 150 200 250 300 350 400 450

Heat Release Rate (kW)

Combustion Efficiency

Heptane Toluene Polystyrene

Figure 95. The local combustion efficiency at the rear and front compartment sampling locations 

as a function of the fire heat release rate during the periods when the HRR was quasi-steady for 

three fuels. 

0%

20%

40%

60%

80%

100%

120%

0% 2% 4% 6% 8% 10% 12%

Oxygen Volume Fraction

Combustion Efficiency

Figure 96. The local combustion efficiency at the rear and front compartment sampling locations 

as a function of the oxygen volume fraction during the periods when the HRR was quasi-steady 

in the heptane fire.
117

5.6 Summary of Chemical Analysis 

The mass fractions of gas species, such as hydrocarbons, carbon dioxide, carbon monoxide, 

oxygen, water vapor, and nitrogen, were considered as a function of the mixture fraction for each 

fuel type. The measurements were compared with the state relations based on the mixture 

fraction correlation model. Measurements in the upper layer of the compartment showed that the 

CO volume fraction was a function of fuel type, fire size, ventilation opening and specific 

sample location. In addition, the gas chromatographic results showed that hydrocarbons in the 

upper layer were composed mainly of CH4. 

Consistent with previous studies, plotting the local composition as a function of the mixture 

fraction (or equivalence ratio) collapses hundreds of individual species measurements from an 

assortment of compartment conditions, with varying heat release rates, ventilation openings, and 

spatial locations, into a few coherent lines or bands. The results show that CO is not wellcorrelated with mixture fraction. Recognizing that CO kinetics are relatively slow, it was shown 

that consideration of temperature effects through analyses based on complete local thermal 

equilibrium and modification of the mixture fraction model to account for direct temperature 

correlation were not particularly fruitful. That these simple methods were not successful does not 

mean that improved accuracy cannot be achieved. 

The results presented here demonstrate that it is important to consider soot as part of the mixture 

fraction analysis of compartment fires. This is particularly true in the upper layer of smoky fires 

in which about half (or more) of the fuel carbon may exist in the form of carbonaceous soot. 

Inclusion of soot in the analysis allows identification of fuel rich or underventilated conditions, 

conditions that otherwise would be considered lean or overventilated. 

The combined mass of carbon in the form of soot and total hydrocarbons seem to be correlated 

by the soot-based mixture fraction for the conditions tested in this study. The CO and soot yields 

were found to be a function of fuel type and local equivalence ratio in the upper layer of the 

compartment for the smoky heptane, toluene and polystyrene fires. In addition, the ratio of CO 

yield to soot yield was found to be independent of local equivalence ratio for the smoky fuels. 

Additional experiments are needed to test this finding for other conditions and for full-scale 

compartment fires. In summary, mixture fraction is a useful initial way to describe the chemical 

structure of compartment fires. 
118

6 Scaling Discussion 

The applicability of the experimental results reported here to other compartment fire scenarios 

can be considered in terms of a number of normalized parameters traditionally used in fire 

modeling applications. Use of normalized parameters facilitates comparison of results from 

scenarios of different scales by normalizing key physical characteristics of the scenario. A 

number of different forms of scaling may be considered, depending on the fire phenomena of 

interest [50]. Table 25 lists three normalized parameters that may be used to compare fire 

scenarios with the experiments reported here. The ranges of values for the normalized parameters 

examined in this study are listed in the table. The table is intended to provide guidance when 

evaluating the applicability of the data set reported here. For any given fire scenario, more than 

one normalized parameter may be necessary for determining applicability of the validation 

results, depending on the parameters of interest. In this sense, the Table should be considered 

illustrative, not exhaustive. 

Table 25. List of non-dimensional scaling parameters for compartment fires and the range of 

values examined in this study. 

Parameter Normalized Representation Range of Values Fuel Type 

Heat Release 

Rate 2

*

d

c T gDD

Q Q

∞ p ∞

=

ρ

&

1.6 to 44.2 

0.7 to 8.0 

1.0 to 7.2 

1.6 to 5.5 

1.5 to 5.5 

1.6 to 5.0 

Natural Gas 

Heptane 

Toluene 

Polystyrene 

Methanol 

Ethanol 

Ventilation 

O o o

F O where m A h

r

m m

2

0.23 , /

2

2 = & = & & φ

0.10 to 1.71 

0.16 to 0.83 

0.09 to 0.96 

0.03 to 0.69 

0.04 to 0.75 

0.01 to 0.77

Natural Gas 

Heptane 

Toluene 

Polystyrene 

Methanol 

Ethanol

Compartment 

height 

2 / 5

*

* , where ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

=

∞c T∞ g

Q D

D

H

ρ p

&

1.4 to 3.5 

1.5 to 2.7 

1.6 to 3.4 

1.6 to 5.2 

1.6 to 5.0 

Natural Gas 

Heptane 

Toluene 

Polystyrene 

Methanol 

Ethanol

The most important parameter of any fire experiment is the heat release rate, as its magnitude 

drives changes in the thermal environment of the compartment or space of interest. A 

normalized quantity that relates the heat release rate to the diameter of the fire, D, is the first 

entry in Table 25, commonly known as * Qd , where Q& is the heat release rate (kW), ρ ∞ is the 

ambient density (kg/m3

), T∞ is the ambient temperature (K), cp is the specific heat (kJ/kg-K), 

and g is the acceleration of gravity (m/s2

). A large value of * Qd represents a fire with a relatively 

large value of energy output power compared to its physical diameter, like an oil well blowout 

fire. A low value of * Qd represents a fire with a relatively small value of energy output compared 

to its diameter, like a smoldering fire. Many typical accidental fire scenarios have * Q d values on 

the order of 1. The physical diameter of a realistic fire may not be well-defined and may not 
119

actually matter when assessing the “size” of a fire. Instead, a characteristic diameter, D* is 

considered in the definition of * Qd as noted in the table. The range of values of * Qd varied as a 

function of fuel type. In this study, * Qd took on values as small as 0.7 and as large as 44 as seen 

in Table 25. 

The second entry in the table is the global equivalence ratio (φ), which is associated with the 

overall fire-induced ventilation and compartment stoichiometry. An estimate of the maximum 

achievable steady-state oxygen supply is given by: mO Ao ho 2

0.23

2

& = , where O2 m& is an empirical 

correlation for the mass flow rate of oxygen (kg/s), Ao and ho are the area and height of the 

doorway opening (m2

), 0.23 is the mass fraction of oxygen in air. The parameter r in the table is 

the mass-based stoichiometric ratio of fuel to air required for complete combustion. The value of 

φ is useful in characterizing whether a given compartment fire is limited in size by its fuel supply 

or by its oxygen supply. The correlation for oxygen entrainment is valid for flashover conditions 

only, that is for values of φ >1. 

In all of the experiments performed as part of this study, the fuel mass flow rate was either 

controlled (for the gaseous fuels) or measured (for the condensed fuels), whereas the oxygen 

supply was naturally controlled by the size of the compartment doorway and the fire heat release 

rate. The range of values of φ varied as a function of fuel type, taking on values as small as 0.01 

as seen in Table 25. The value of φ was less than 1.0 for almost all of the experimental 

conditions, except for natural gas when φ was as large as 1.7. This implies that conditions inside 

of the compartment were nearly always over-ventilated. There is strong evidence that the 

estimate for air mass flow is over-predicted using this approach. For each of the fuels listed in 

the table, flames were observed outside of the doorway (see photos in section 3 of this report), 

oxygen volume fractions were near zero and increased CO production was measured in the upper 

layer for the largest fires sizes. These are all strong indicators of underventilated burning. This 

may be due to an inaccurate assumption of the incoming air mass flow rate, or an invalid 

assumption in the GER model that all the incoming air enters the mixed upper layer. Pitts [51] 

proposed that a large fraction of the incoming air is entrained into the out-flowing gases and 

never reaches the reaction zone. 

The third entry in the table is the compartment height, H, normalized by D*. The parameter 

H/D* relates Q& to its physical dimensions and indicates the relative importance of the fire plume 

to other features of the fire-driven flow, such as the ceiling jet or doorway flow. The range of 

values of H/D* varied as a function of fuel type, taking on values as small as 1.4 and as large as 

5.2 for conditions examined in this study, as seen in Table 25.
120

7 Conclusions

This reports details the test methods and experimental results from a series of fire tests in the 

reduced scale enclosure. The following list describes the main findings of this work: 

• New measurements of total hydrocarbons, and soot were successfully performed. These 

new measurement provide a more complete data set for validating and improving predictive 

fire models. 

• The performance of various burner designs, wall materials and sample conditioning methods 

were evaluated in order to aid in the planning of future full-scale fire experiments. 

• A detailed analysis of thermocouple temperature measurements provided an estimate of 

radiative error and response time in the thermal environment of a fully developed enclosure 

fire (see Appendix A). 

• CFD modeling was successfully used to examine the effects of probe interactions (see 

Appendix B) and help design the experiments. 

• The gas species composition measurements showed that methane was the most abundant 

hydrocarbon species in the upper layer for all of the fuels and fire conditions tested, and was 

higher in concentration than the parent fuel in all cases. 

• No significant amount of hydrocarbons was measured in the upper layer of the compartment 

in the toluene or polystyrene fires. 

• As much as 60 % of the carbon in a polystyrene fire was present in the form of soot, which 

was more than twice the amount from any of the other hydrocarbons tested. 

• The examination of compartment fire species data in terms of mixture fraction provided a 

rapid assessment of the overall chemical structure of the combustion environment. The 

measured data plotted as a function of mixture fraction confirmed the linear state 

relationship model for complete combustion when the fire was well ventilated. 

• The results show that it is useful to consider soot as part of a mixture fraction analysis of 

compartment fires. The compartment gas species measurements showed that the CO 

concentration was a complex function of fuel type, fire size, ventilation and compartment 

sample location. The measured data plotted as a function of mixture fraction was consistent 

with previous studies that show the linear state relationship model is not valid for predicting 

CO in underventilated fires. In addition, a systematic correlation with temperature was not 

found. The results suggest that more research is needed to unravel the complexities of 

compartment fire chemistry. 

• The species measurements showed that the soot yield was a function of fuel type and local 

equivalence ratio in the upper layer, whereas the ratio of the CO yield to the soot yield was 

independent of local equivalence ratio, and was not dissimilar to the results reported 

previously in Ref. [46]. 

• The mass fraction of CO in the upper layer of a reduced-scale compartment fire does not 

systematically correlate with the local temperature or mixture fraction, nor is it in local 

thermodynamic equilibrium. 
121

• The yields of CO and soot in the exhaust stack were usually, but not always, lower than the 

local yields in the upper layer of a compartment, for a wide range of fuels and heat release 

rates. 

• The post-compartment conditions generally had larger combustion efficiency values than 

local conditions in the upper layer of the reduced-scale compartment fires, for a wide range 

of fuels and heat release rates. 

• The ratio of the yields of CO to soot in the upper layer of the reduced-scale compartment 

were relatively constant for each of the fuels types, for the fires burning heptane, toluene 

and polystyrene; in addition, the values were comparable to previous measurements [46] in 

the plumes of over-ventilated turbulent hydrocarbon fires. 

• The yields of CO and soot in the upper layer of the reduced-scale compartment fires were 

related to the local equivalence ratio for fires burning ethanol, heptane, toluene and 

polystyrene, over a wide range of fire sizes, but significant scatter precludes use of the data 

for CO prediction; the relationship was non-linear in fires burning methanol and natural gas. 

8 Future Work 

The experiments described in this report are part of an ongoing project to explore the thermal and 

chemical phenomena associated with, primarily, underventilated compartment fires. In the next 

phase of this project fire measurements will be conducted using a full-scale ISO 9705 

compartment with quantified uncertainties. Based on the second scaling rule shown in Table 25 

and the results shown in Fig. 42, oxygen depletion could be expected in the front of an ISO 9705 

enclosure for fires larger than about 1800 kW. Since investigation of underventilated burning is 

important, it may be advantageous to consider using a narrow doorway to reduce the fire size 

requirements to obtain conditions of interest in the experimental enclosure. The following tasks 

are planned to further this goal: 

• Investigate the effects of distributed fuel sources on the structure of full-scale enclosure 

fires. 

• Explore the effects of larger global equivalence ratios (primarily through decreased 

ventilation). 

• Investigate scaling effects on the chemical and thermal structure of the fires. 

• Explore the utility and durability of high-temperature-resistant blankets for full-scale 

enclosure construction. 

• Improve the high temperature accuracy and time response of aspirated thermocouple 

probes through improved design. 

• Extend gas species measurements to include quantification of hydrogen and water. 

• Instrument the doorway to differentiate the heat release within and external to the 

compartment and to enable determination of the flows of enthalpy, mass and elemental 

carbon in and out of the compartment. 
122

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10 Acknowledgements

A number of individuals were instrumental in the completion of this work. The fabrication work 

and execution of these tests would not have been possible without the expert assistance and 

enthusiasm of the LFL staff: Laurean DeLauter, Edward Hnetkovsky and Jack Lee. Mike 

Selepak did exemplary work in carefully performing the labor intensive extractive soot 

measurements. Jay McElroy performed the video documentation and instrument installation. 

The authors are grateful to the following NIST staff for extremely helpful discussions that were a 

significant contribution to this study: Kevin McGratten, Jason Floyd, Rodney Bryant, Bill Pitts, 

Alex Maranghides, Bill Grosshandler, and Andrew Lock. 

This work is dedicated in memory of Jack Lee (1953-2007). Jack was a technician in the large 

fire laboratory and will be greatly missed. He provided 35 years of loyal service to NIST and 

countless contributions to fire safety research. 
126

11 Appendices 

A. Analysis of Thermocouple Temperature Measurement 

To estimate components of measurement uncertainty and instrument time response, the present 

study performed a series of detailed flow and heat transfer calculations, focusing on double 

shielded aspirated thermocouples and bare bead thermocouples. The character of the 

calculations is summarized in Table A1. There were two categories of calculations. The first 

category involved three-dimensional (3D) computational fluid dynamic (CFD) simulations, 

while the second involved an algebraic solution of the simplified energy balance equation. Two 

types of CFD modeling were performed. The first considered the realistic geometry in order to 

understand details of the flow field associated with the aspirated thermocouple. The second 

considered a simplified geometry, focusing on details of the heat transfer process. Ideally, the 

CFD simulation would be able to consider details of the flow and heat transfer in the actual 

geometry, including turbulence and conjugate heat transfer with conduction, convection and 

radiation for the real geometry, but this would be prohibitively expensive computationally. For 

this reason, the CFD calculations were split into detailed flow calculations for the actual 

geometry, and detailed conjugate heat transfer calculations for a simplified geometry. The 3D 

CFD calculation results are compared with algebraic solutions of a simplified energy balance 

model. The results predicted by the algebraic model were in agreement with the CFD model over a broad 

temperature range, despite its many assumptions and idealizations. A parametric study was conducted to 

quantify the thermocouple errors for various gas temperature and surrounding conditions. The CFD 

solutions were transient, allowing determination of instrument time response. This is in contrast 

to the analytical solution, which presented a steady-state solution only. 

Table A1. Summary of the Numerical Simulations. 

Simple energy 

balance model 

3D CFD model for 

Simple Geometry 

3D CFD model for 

Real Geometry 

Physics 

Heat transfer based 

on specified 

aspiration rate 

Flow and heat 

transfer 

Turbulence 

associated with 

aspiration flow 

Solid Phase Heat 

Transfer Thermally Thin Considered Not Considered 

Radiative Heat 

Transfer Surface radiation Surface-Surface Radiation Not Considered 

Convective Heat 

Transfer Empirical correlation Surface Energy balance 

Surface Energy 

Balance 

Surface Emissivity 0.8 0.8 Not considered 

Surrounding 

Temp. 300 K to 1200 K 300 K to 1200 K Adiabatic 

Gas Temp. 300 K to 1200 K 300 K to 1200 K 300 K, 1200 K 

Solution Type Steady Solution Transient solution Transient Solution 

Turbulence Empirical correlation Standard K-ε model Standard K-ε model 

External Gas 

Velocity 1 m/s 1 m/s 1 m/s 

No. Grids None 25000 1500000 
127

3D Flow Modeling for a Realistic Geometry 

The 3D flow modeling considered the realistic geometry of an open-ended double shielded 

aspirated thermocouple as seen in Fig. A1. The computational configuration included an inner 

and outer shields, a bare type K thermocouple taken to be cylindrical, and an extended domain. 

The computational domain was divided into approximately 1.5 million cells of a tetrahedral type 

mesh using the ICEM-CFD [2], which is a commercial CAD and grid generation program. 

Figures A2 and A3 show the computational grids for the open-end and open-side aspirated 

thermocouple geometries. Both types were used in the RSE compartment fire experiments. The 

flow fields were calculated using a commercially available CFD package to model the flow for 

given operating conditions [3]. The code is based on the finite volume method on a collocated 

grid. A non-staggered grid system was used for storage of discrete velocities and pressures. The 

standard k-ε turbulence model and incompressible ideal gas assumption were applied to solve the 

Reynolds stress term and the density change, respectively. The governing equations were 

discretized by the 2nd order upwind scheme in space and the SIMPLE algorithm (Semi-Implicit 

Method for Pressure Linked Equation) with under-relaxation used to iteratively solve the 

momentum equation in discretized form. The implicit solver was used to capture the main 

features of the unsteady motion around the aspirated thermocouples. In this study, CFD 

calculations investigated the flow field in the inner and outer aspiration tubes with an emphasis 

on characterizing any differences. That information was then used as boundary conditions for 

the CFD and analytic heat transfer analyses. 

Figure A1. Schematic of a double shielded end-open aspirated thermocouple. 
128

(a) Perspective view 

(b) Front view 

Figure A2. Computational grids for end-open double shield aspirated thermocouples. 
129

(a) Perspective view 

(b) Front view 

Figure A3. Computational grids for side-open double shield aspirated thermocouples. 
130

Conjugate Heat Transfer Modeling for a Simplified Geometry 

A detailed 3D heat transfer calculation (including conduction, convection, and radiation) was 

performed to estimate the effectiveness of bare-bead and double-shielded aspirated 

thermocouples for a simplified geometry. Figure A4 presents a schematic of the two 

configurations. The double-shielded calculation assumed that the flow domain consisted of an 

external flow, an annular flow between the outer and inner cylinders, and an inner flow within 

the inner cylinder. In the bare-bead calculation, the thermocouples were directly exposed to the 

external gas and the surroundings without a shield or an applied aspiration flow. The incoming 

gas temperature was assumed to be uniform. The incoming flow velocity induced by aspiration 

was imposed as a boundary condition, and was determined from the detailed flow calculations 

described above. The material properties of the shield and thermocouple bead were taken to be 

steel and nickel, respectively, and are listed in Table A2. This is a reasonable approximation as 

K-type thermocouples are composed of more than 90 % nickel [4]. 

Table A2. Material Properties [5]. 

Material Density (kg/m3

) Specific heat (J/kg·K) Thermal conductivity (W/m·K) 

Nickel 8900 460 91.7 

Steel 8030 502 16.3 

Figure A4. Schematic of bare bead and open-end double shielded aspirated thermocouples. 
131

Radiative heat transfer was computed using a surface to surface radiation model in which the 

energy exchange between the two surfaces depended on the view factor, which is a geometric 

function involving the size, distance, and orientation of surfaces. The surfaces were taken as gray 

and diffuse. The calculations were performed assuming a constant emissivity (ε) of 0.8 for all 

metal surfaces, using the value recommended by Blevins [1]. The external flow velocity was 

assumed to be 1 m/s in all cases. 

The representative thermocouple temperature (Tb) was calculated using a volume weighted 

average as follows: 

∫ ∑=

= ⋅ = ⋅

N

i

b i i

b b

b T dV

V

T dv

V

T

1

,

1 1 (A1) 

The thermocouple effectiveness was defined in terms of the percentage error (E) between the 

incoming gas flow temperature (Tg in Fig. A4) and the representative temperature at the 

thermocouple bead (Tb in Fig. A4): 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ − = ×

g

b g

T

T T

E 100 (A2) 

Steady-State Solution of the Algebraic Energy Balance Equation 

An analytic solution of the steady state energy balance for aspirated thermocouples was 

previously reported [1]. That calculation, however, did not consider conductive heat transfer 

through the solid shield. Also, the temperature difference between the inner and outer surfaces 

was assumed to be zero. In the case of the double shielded aspirated thermocouple, the inner and 

outer gas velocities were assumed to be equal and the external flow was assumed to be parallel to 

the probe axis. 

Results 

Figure A5 shows the calculated pressure and velocity fields of an end-open double shielded 

aspirated thermocouple at ambient temperature with an aspiration flow rate of 24 L/min. The 

entrance area of the inner cylinder of the aspirated thermocouple was larger than the exit area 

through its eight inner holes (see Figs. A1 and A2), causing a relatively high stagnation pressure 

inside the inner shield. This adverse pressure gradient forced the aspiration flow to pass through 

the outer passage with a differentially high velocity as compared to the inner cylinder. The flow 

blockage effect was characterized as a ratio of the average velocities between the inner and outer 

annular passages. It influenced the rate of convective heat transfer in the aspirated thermocouple, 

impacting the effectiveness of the aspirated thermocouple. Figure A6 presents the calculated 

maximum values of the velocity of the inner shield (Ui), the outer annular passage (Uo), and the 

maximum value of the ratio of these velocities (ζ) as a function of the total aspiration flow. The 

velocity ratio (ζ) is defined as: 

i

o

U

U

ς = (A3) 

For nominal operating conditions (24 L/min at STP), the maximum velocity in the inner shield 

was found to be about 3 times larger than that within the annular passage. Figure A6 shows that 

this ratio increased with aspiration flow. 
132

(a) Pressure field 

(b) Velocity magnitude 

Figure A5. Calculated pressure and velocity fields in end-open double shielded aspirated 

thermocouples at ambient temperature with an aspiration flow rate of 24 L/min. 
133

0

5

10

15

20

25

30

35

40

0 5 10 15 20 25 30 35 40

Volume flow rate [liter/min]

axial velocity component along the tube [m/s]

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Maximum velocity ratio

Max. velocity of inner shield

Max. velocity of outer annular passage

Maximum velocity ratio

Figure A6. Comparison of the maximum axial velocity for end-open double shielded aspirated 

thermocouples ( Tgas = 300 K ). 

Figure A7 shows the calculated thermocouple error (see Eq. A2) for a bare bead thermocouple as 

a function of the surrounding (To in Fig. A4) temperatures for both the simple energy balance 

model and the 3D CFD model for an incoming gas temperature of either 300 K or 900 K. The 

external gas velocity was taken as 1 m/s directed toward the open end of the aspirated 

thermocouple. The results show one large error regime of thermocouple measurement associated 

with low values of the gas temperature and high surrounding temperatures. The results show that 

the error was relatively small for moderately-high gas temperatures, regardless of the 

surrounding temperature. Despite its many assumptions, the solution to the simple energy 

balance model was essentially in agreement with the 3D CFD calculation results for the bare 

bead thermocouple. 

Figure A8 show maps of the calculated thermocouple error (Eq. A2) for an open-end doubleshielded aspirated thermocouple as a function of the thermocouple bead (Tb) and the surrounding 

(To) temperatures determined using the 3D CFD model. The map shows two regimes of 

significant error for the thermocouple temperature measurements. The first occurs for relatively 

low temperature surroundings in which the gas temperature is systematically under-predicted, 

notably for higher gas temperatures. The temperature error predicted by the CFD model in this 

regime varied with temperature, but was as large as -280 K for gas temperatures of 1200 K, 

corresponding to an error of about 25 %. The other significant-error regime occurs for large 

surrounding temperatures, in which the gas temperature is over-predicted. The maximum error in 

this regime was as large as 140 K for a high surrounding temperature (To = 1500 K) and a 

thermocouple temperature of about 900 K. Calculations using the algebraic model show that 

similar trends result, but the magnitude of the differences was much smaller. There is less 

confidence in the algebraic model, due to its many assumptions. 
134

Figure A9 shows the calculated thermocouple response to a hypothetical step change in the gas 

temperature with an invariant surrounding temperature using the CFD model. The initial gas 

temperature and surrounding temperature were identical. Results are shown for initial 

temperatures of 1200 K and an aspiration rate of 24 L/min. The temperature rose or fell towards 

the final temperature over tens of seconds, the exact time depending on the relatively size of the 

step change. Results from several calculations similar to those presented in Fig. A9 are 

summarized in Fig. A10, which shows the time to reach a quasi-steady state condition after the 

hypothetical step change for various surrounding temperatures. Steady state was defined as the 

time when the thermocouple temperature variation was less than 1 K/s. Under this condition, the 

calculated results are within 10 K of the asymptotic calculation results. 

The time to steady state varied with incoming gas temperatures, but was largest for high 

incoming gas temperatures. For example, for the case of a surrounding temperature of 300 K, 

the time to steady state took a maximum value of about 50 s for an incoming gas temperature of 

1100 K. 

0

50

100

150

200

250

300

300 400 500 600 700 800 900 1000 1100 1200

Surrounding temperature [K]

percent error in measuring temperature [%]

Simple energy balance model for Tg=300K

Simple Energy balance model for Tg=900K

3D CFD model for Tg=300K

3D CFD model for Tg=900K

Figure A7. A comparison of the predicted percent error in measuring temperature between the 

simple energy balance model and the 3D CFD model for bare bead thermocouples with an 

external gas velocity of 1 m/s. 
135

Figure A8. Map of the thermocouple uncertainty for a double shield aspirated thermocouple as 

a function of the measured (Tb) and surrounding (To) temperatures calculated using the CFD 

model. 

300

500

700

900

1100

1300

1500

-20 0 20 40 60 80 100

Time [s]

Thermocouple temperature [K]

initial 

condition

Tg=300K

Tg=500K

Tg=700

K

Tg=900

Tg=1100K

Tg=1300K

Tg=1500

Figure A9. Calculated time history of the temperature of a double shielded aspirated 

thermocouple for various incoming gas temperatures (Tg) with a surrounding temperature of 

1200 K and an aspiration flow of 24 L/min). 

Tg-Tb
136

Figure A10. Comparison of calculated time to reach steady state for a double shielded aspirated 

thermocouple with a 24 L/min aspiration flow rate as a function of the incoming gas temperature 

for surrounding temperatures of 300 K, 600 K, and 1200 K. 

Summary 

The present study investigated the flow and heat transfer characteristics of aspirated 

thermocouples using a simple energy balance model and a 3D CFD model. The calculations 

quantify the systematic measurement error, providing an estimate of the measurement 

uncertainty. At the same time, the model provides information on the time to achieve steadystate, which should be carefully considered in terms of the development of an experimental 

procedure and in the interpretation of the results. 

 

Despite the application of additional assumptions and idealizations, calculations using the 

previously developed algebraic energy balance model generally showed good agreement with the 

results of the 3D CFD model. The algebraic model can be useful, particularly in parametric 

studies used to evaluate thermocouple measurement error. Consistent with previous findings, 

calculations show that use of the double shield aspirated thermocouple can greatly reduce the 

thermocouple error especially for low gas temperatures. The results, however, can still be biased 

by hundreds of degrees, depending on the conditions. Figure A8 provides information on the 

magnitude of the measurement error for a given value of the surrounding temperature. Since the 

surrounding temperature takes on multiple values and is a complex function of thermocouple 

location and fire conditions, a representation of the surrounding temperature can be made based 

on estimated temperature averages. Precise determination of the surrounding temperature is 

impossible, and engineering judgment is a key component of the uncertainty analysis. 
137

The results of the CFD model allow determination of the transient response of the double shield 

aspirated thermocouple, which is helpful in the interpretation of measurement results and 

possibly for design of the experiment itself. The calculated time response of the aspirated 

thermocouples suggests that while the measurements provide an adequate representation of the 

average local temperature, they do not provide an accurate representation of the magnitude of the 

temperature excursions. Turning the aspiration flow off during the experiment may bias the 

results if the time response is not taken into account. As a rule of thumb, for incoming gas 

temperatures less than 900 K, an aspirated thermocouple should be in place for at least 1⁄2 min 

before the data should be considered acceptable. For incoming gas temperatures greater than 

900 K, an aspirated thermocouple should be in place for about 1 min before the data should be 

considered acceptable, unless the thermocouple is sampling from an optically thick upper layer, 

when 1⁄2 min should be adequate. 

References 

[1] Blevins, L.G., Behavior of bare and aspirated thermocouples in compartment fires, Proc. 33rd

National Heat Transfer Conference, New Mexico, August 15-17, 1999. 

[2] ICEM-CFD Version 4.0, User’s Manual. 

[3] FLUENT Inc., FLUENT Version 6.0 User’s Guide. 

[4] Burns, G. W. and Scroger, M. G., The calibration of thermocouples and thermocouple 

materials, NIST Special Publication 250-35, 1989. 

[5] Incropera, F. P and DeWitt, D. P., Introduction to Heat Transfer, 2nd edition, Willy, 1993. 
138

B. Analysis of Probe Interactions

There were five different types of measurement probes at various positions in the fire 

compartment. They included bi-directional probes for velocity, bare-bead and aspirated 

thermocouples for temperature, gas sampling tubes for gas species and soot sampling probes for 

the gravimetric soot measurement. These probes affect the flow field through geometric 

obstructions and create a flow sink due to the extraction of combustion gas. The effects of the 

probes on the flow field were not avoidable in the measurement system, but should be minimized 

during the test. Among the measurement probes, the suction type probes, such as the aspirated 

thermocouples, gas and soot sampling probes could significantly influence the flow field. If the 

suction type probes are operated too close to each other, there could be an interaction associated 

with the suction flow rate. 

Before the tests were conducted, numerical simulations were used to investigate the effect of the 

sampling probe on the flow field for various operating conditions. The interactions between the 

aspirated thermocouples and the gas sampling probes were investigated because of the high 

suction flow rate. The experimental suction flow rate of the soot probe was similar to the gas 

sampling probe (Q ≈ 3 L/min @ 300 K). 

Figure B1shows the configuration of the two sampling probes perpendicular to incoming gas 

flow. The flow field near the entrance of the sampling probe was examined for the typical 

operating conditions of the probes (Q1 = 30 L/min for the aspirated thermocouple, Q2 = 3 L/min 

for the gas sampling). In the calculations the incoming gas had a constant velocity (Ug) and gas 

temperature (Tg). Figure B2 shows the velocity field near the sampling probe for an incoming 

gas velocity of 1 m/s, and a gas temperature of 900 K. The influence of the sampling flow rate 

was restricted to the probe entrance. Generally, the gas velocities induced by the fire were much 

higher than the suction velocities of the sampling probes, and the effect of the probe suction on 

the flow field was negligible. 

Figure B3 shows the streamlines at the center-plane of the probe. The distance from the probe 

entrance to the aspirated flow field disturbance was on the order of one diameter. Figure B4 and 

Fig. B5 represent the static pressure and stream-wise velocity (Ux) profiles along the center axis 

of the probes for the 1 m/s gas velocity and 900 K temperature. The static pressure and streamwise velocity do not vary significantly in this configuration. Also, the velocity profiles of the 

aspirated thermocouple and the gas sampling probe have nearly the same magnitude. If one 

probe affects the flow field of the other probe, the velocity profiles would not match. The 

incoming gas flow dominates the entire flow field except at the entrance of the probes because 

the flow rate of the sampling probes is relatively small compared to the main gas flow. 

Therefore, the probe interactions were negligible in the main plume of the fire induced flow, but 

in a low velocity region such as stagnation or recirculation zone, the suction from one probe can 

affect the flow field and the measurement volume being interrogated by other probes. 
139

Figure B1. Schematic of the calculations of the sampling probes interaction. 

Figure B2. Velocity field across cross-section of the probes position.
140

Figure B3. Streamlines near the aspirated thermocouple.

-1

-0.8

-0.6

-0.4

-0.2

0

0.2

0.4

0.6

0.8

1

-0.01 0 0.01 0.02 0.03 0.04 0.05

distance from the entrance of probe [m]

static pressure [pa]

Aspirated TC

Gas sampling

probe

Figure B4. Static pressure profiles along the center axis of the probes. 
141

Figure B5. U-velocity profile along the center axis of the probes. 
142

 C. Channel Lists Series: Measurements x x0=inside left wall (looking at door) y y0=inside front (door) wall z z0=floor Location Description Overall Channel Number Abbr. MIDAS Station Module: Channels Mod. Mod. Ch. No. Conv.U nits Wire Gain Main Channels O2 at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 0 O2Rear East 1:0-31 1 0 Vol Fr Cu 1 Bad channel 1 Bad 1 1 CO2 at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 2 CO2Rear East 1 2 Vol Fr Cu 1 CO at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 3 CORear East 1 3 Vol Fr Cu 1 UH at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 4 UHRear East 1 4 Vol Fr Cu 1 Rear Gas Rack Dewpoint outside RSE 5 DPRear East 1 5 °C Cu 1 Rear Sampling location gas temperature 1-near probe outside RSE 6 TGasRear1 East 1 6 °C TC 100 Rear Sampling location gas temperature 2-near UH outside RSE 7 TGasRear2 East 1 7 °C TC 100 Rear Sampling location water inlet temperature outside RSE 8 TH2OInRear East 1 8 °C TC 100 Rear Sampling location water outlet temperature outside RSE 9 TH2OOutRear East 1 9 °C TC 100 O2 at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 10 O2Front East 1 10 Vol Fr Cu 1

CO2 at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 11 CO2Front East 1 11 Vol Fr Cu 1CO at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 12 COFront East 1 12 Vol Fr Cu 1UH at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 13 UHFront East 1 13 Vol Fr Cu 1Front Gas Rack Dewpoint outside RSE 14 DPFront East 1 14 °C Cu 1

Front Sampling location gas temperature 1- near probe outside RSE 15 TGasFront1 East 1 15 °C TC 100

Front Sampling location gas temperature 2- near UH outside RSE 16 TGasFront2 East 1 16 °C TC 100Front Sampling location water inlet temperature outside RSE 17 TH2OInFront East 1 17 °C TC 100

Front Sampling location water outlet temperature outside RSE 18 TH2OOutFront East 1 18 °C TC 100Optical Soot (I/Io in adj file) 48 -5 75 6 cm from Soffit, doorway CL 28 OptSoot East 1 28 V Cu 1

Optical Soot Laser Reference Signal 48 -5 75 6 cm from Soffit, doorway CL 29 OptSootRef East 1 29 V Cu 1Optical Soot Apparatus Temperature outside RSE 30 TOptSoot East 1 30 °C TC 100Gravimetric Soot Probe 1 Filter Temperature - rear outside RSE 31 TGravSoot1 East 1 31 °C TC 100

Gravimetric Soot Probe 2 Filter Temperature -rear outside RSE 32 TGravSoot2 East 2:0-31 2 0 °C TC 100Gravimetric Soot Probe 1 Solenoid Signal -rear outside RSE 33 VGravSoot1 East 2 1 V Cu 1Gravimetric Soot Probe 2 Solenoid Signal -rear outside RSE 34 VGravSoot2 East 2 2 V Cu 1Gravimetric Soot Probe 1 Mass Flow -rear outside RSE 35 SootMF1 East 2 3 V Cu 1Gravimetric Soot Probe 2 Mass Flow-front outside RSE 36 SootMF2 East 2 4 V Cu 1Total Heat Flux Gauge Front (SN=131836) 48 35 0 In floor on room CL, 35 cm from Front 37 HFF East 2 5 kW/m2 Cu 100

Total Heat Flux Gauge Rear (SN=131835) 48 106 0 In floor on room CL, 35 cm from Rear 38 HFR East 2 6 kW/m2 Cu 100Doorway Pressure 79 cm Left (+out) 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 39 PD79L East 2 7 Pa Cu 1Doorway Pressure 79 cm CL (+out) 48 -5 79 79 cm from floor, doorway CL 40 PD79C East 2 8 Pa Cu 1

Doorway Pressure 79 cm Right (+out) 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 41 PD79R East 2 9 Pa Cu 1

Doorway Pressure 60 cm (+out) 48 -5 60 60 cm from floor, doorway CL 42 PD60C East 2 10 Pa Cu 1

Doorway Pressure 40 cm (+in) 48 -5 40 40 cm from floor, doorway CL 43 PD40C East 2 11 Pa Cu 1Doorway Pressure 20 cm Left (+in) 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 44 PD20L East 2 12 Pa Cu 1

Doorway Pressure 20 cm CL (+in) 48 -5 20 20 cm from floor, doorway CL 45 PD20C East 2 13 Pa Cu 1

Doorway Pressure 20 cm Right (+in) 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 46 PD20R East 2 14 Pa Cu 1

Doorway Pressure 5 cm (+in) 48 -5 5 5 cm from floor, doorway CL 47 PD5C East 2 15 Pa Cu 1

Doorway Temperature 79 cm Left BB 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 48 TD79LBB East 2 16 °C TC 100

Doorway Temperature 79 cm CL BB 48 -5 79 79 cm from floor, doorway CL 49 TD79CBB East 2 17 °C TC 100

Doorway Temperature 79 cm Right BB 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 50 TD79RBB East 2 18 °C TC 100

Doorway Temperature 70 cm Left Asp 32 -5 70 70 cm from floor, doorway 8 cm from Left side of dr 51 TD70LA East 2 19 °C TC 100Doorway Temperature 70 cm Left BB 32 -5 70 70 cm from floor, doorway 8 cm from Left side of dr 52 TD70LBB East 2 20 °C TC 100Doorway Temperature 70 cm CL Asp 48 -5 70 70 cm from floor, doorway CL 53 TD70CA East 2 21 °C TC 100

Doorway Temperature 70 cm CL BB 48 -5 70 70 cm from floor, doorway CL 54 TD70CBB East 2 22 °C TC 100Doorway Temperature 30 cm Left Asp 32 -5 30 30 cm from floor, doorway 8 cm from Left side of dr 55 TD30LA East 2 23 °C TC 100

Doorway Temperature 30 cm Left BB 32 -5 30 30 cm from floor, doorway 8 cm from Left side of dr 56 TD30LBB East 2 24 °C TC 100Doorway Temperature 60 cm CL BB 48 -5 60 60 cm from floor, doorway CL 57 TD60CBB East 2 25 °C TC 100

Doorway Temperature 50 cm CL Asp 48 -5 50 50 cm from floor, doorway CL 58 TD50CA East 2 26 °C TC 100Doorway Temperature 50 cm CL BB 48 -5 50 50 cm from floor, doorway CL 59 TD50CBB East 2 27 °C TC 100Doorway Temperature 40 cm CL BB 48 -5 40 40 cm from floor, doorway CL 60 TD40CBB East 2 28 °C TC 100Doorway Temperature 30 cm CL Asp 48 -5 30 30 cm from floor, doorway CL 61 TD30CA East 2 29 °C TC 100Doorway Temperature 30 cm CL BB 48 -5 30 30 cm from floor, doorway CL 62 TD30CBB East 2 30 °C TC 100

RSEmax=(95,142,98)

LFL MIDAS Hookup Sheet Instrument and Channel Description Underventilated RSE Revision Date: 3/29/2006
143

Series:

Measurements

x x0=inside 

left wall 

(looking at 

door)

y y0=inside 

front (door) 

wall

z z0=floor

Location Description

Overall 

Channel 

Number

Abbr. MIDAS Station Module: Channels Mod. Mod. Ch. No. Conv.U nits Wire Gain

Main Channels

Doorway Temperature 20 cm Left BB 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 63 TD20LBB East 2 31 °C TC 100

Doorway Temperature 20 cm CL BB 48 -5 20 20 cm from floor, doorway CL 64 TD20CBB East 4:0-31 4 0 °C TC 100

Doorway Temperature 20 cm Right BB 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 65 TD20RBB East 4 1 °C TC 100

Doorway Temperature 5 cm CL BB 48 -5 5 5 cm from floor, doorway CL 66 TD5CBB East 4 2 °C TC 100Rear Temperature at Sampling Location Asp 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 67 TRSampA East 4 3 °C TC 100Rear Temperature at Sampling Location BB 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 68 TRSampBB East 4 4 °C TC 100Rear Temperature low Asp 75 122 24 24 cm from Floor, 20 cm from Right, 20 cm from Rear 69 TR24A East 4 5 °C TC 100Rear Temperature low BB 75 122 24 24 cm from Floor, 20 cm from Right, 20 cm from Rear 70 TR24BB East 4 6 °C TC 100Rear Temperature high Asp 75 122 80 80 cm from Floor, 20 cm from Right, 20 cm from Rear 71 TR80A East 4 7 °C TC 100Rear Temperature high BB 75 122 80 80 cm from Floor, 20 cm from Right, 20 cm from Rear 72 TR80BB East 4 8 °C TC 100Front Temperature at Sampling Location Asp 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 73 TFSampA East 4 9 °C TC 100Front Temperature at Sampling Location BB 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 74 TFSampBB East 4 10 °C TC 100Front Temperature low Asp 75 20 24 24 cm from Floor, 20 cm from Right, 20 cm from Front 75 TF24A East 4 11 °C TC 100Front Temperature low BB 75 20 24 24 cm from Floor, 20 cm from Right, 20 cm from Front 76 TF24BB East 4 12 °C TC 100Front Temperature high Asp 75 20 80 80 cm from Floor, 20 cm from Right, 20 cm from Front 77 TF80A East 4 13 °C TC 100Front Temperature high BB 75 20 80 80 cm from Floor, 20 cm from Right, 20 cm from Front 78 TF80BB East 4 14 °C TC 100Gas Burner Temperature Center 48 71 15 79 TNGasC East 4 15 °C TC 100

Gas Burner Temperature Front 48 65 15 80 TNGasF East 4 16 °C TC 100Gas Burner Temperature Rear 48 77 15 81 TNGasR East 4 17 °C TC 100Liquid Burner Temperature 1 cm 48 71 9 1 cm from burner bottom 82 TLiq1 East 4 18 °C TC 100Liquid Burner Temperature 2 cm 48 71 10 2 cm from burner bottom 83 TLiq2 East 4 19 °C TC 100Liquid Burner Temperature 3 cm 48 71 11 3 cm from burner bottom 84 TLiq3 East 4 20 °C TC 100Liquid Burner Temperature 4 cm 48 71 12 4 cm from burner bottom 85 TLiq4 East 4 21 °C TC 100Liquid Burner Temperature 5 cm 48 71 13 5 cm from burner bottom 86 TLiq5 East 4 22 °C TC 100Liquid Burner Coolant Temperature Inlet outside RSE 87 TCoolIn East 4 23 °C TC 100

Liquid Burner Coolant Temperature Outlet outside RSE 88 TCoolOut East 4 24 °C TC 100Temperature of Total Heat Flux Gauge Front (SN=131836) 48 35 0 In floor on room CL, 35 cm from Front 89 THFF East 4 25 °C TC 100Temperature of Total Heat Flux Gauge Rear (SN=131835) 48 106 0 In floor on room CL, 35 cm from Rear 90 THFR East 4 26 °C TC 100

Tamb At MIDAS East station 91 Tamb East 4 27 °C TC 100Gravimetric Soot Probe 3 Filter Temperature -front outside RSE 92 TGravSoot3 East 4 28 °C TC 100Gravimetric Soot Probe 4 Filter Temperature-front outside RSE 93 TGravSoot4 East 4 29 °C TC 100

Gravimetric Soot Probe 3 Solenoid Signal-front outside RSE 94 VGravSoot3 East 4 30 V Cu 1Gravimetric Soot Probe 4 Solenoid Signal-front outside RSE 95 VGravSoot4 East 4 31 V Cu 1Surface Temperature Back Wall Exterior Upper Left 20? 144? 87? 10 cm down, 20 cm from left (facing in doorway) 96 TWallUL East 3:0-6 3 0 °C TC 100

Surface Temperature Back Wall Exterior Upper Right 76? 144? 87? 10 cm down, 20 cm from right (facing in doorway) 97 TWallUR East 3 1 °C TC 100

Surface Temperature Back Wall Exterior Lower Left 20? 144? 10? 10 cm up, 20 cm from left (facing in doorway) 98 TWallLL East 3 2 °C TC 100

Surface Temperature Back Wall Exterior Lower Right 76? 144? 10? 10 cm up, 20 cm from right (facing in doorway) 99 TWallLR East 3 3 °C TC 100

Surface Temperature Inside Floor 49? 35? 0 near front HF gauge 100 TFloor East 3 4 °C TC 100

Surface Temperature Inside Ceiling 31? 11? 98? near front sampling probe @ ceiling 101 TCeil East 3 5 °C TC 100

Liquid Fuel Temperature @ Flowmeter outside RSE 102 TAmbFuel East 3 6 °C TC 100

Created Channels

Event Marker 1 103 Event1

Event Marker 2 104 Event2Doorway Velocity 79 cm Left 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 105 VD79L B 0 m/s

Doorway Velocity 79 cm CL 48 -5 79 79 cm from floor, doorway CL 106 VD79C B 1 m/s

Doorway Velocity 79 cm Right 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 107 VD79R B 2 m/s

Doorway Velocity 60 cm 48 -5 60 60 cm from floor, doorway CL 108 VD60C B 3 m/s

Doorway Velocity 40 cm 48 -5 40 40 cm from floor, doorway CL 109 VD40C B 4 m/s

Doorway Velocity 20 cm Left 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 110 VD20L B 5 m/s

Doorway Velocity 20 cm CL 48 -5 20 20 cm from floor, doorway CL 111 VD20C B 6 m/s

Doorway Velocity 20 cm Right 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 112 VD20R B 7 m/s

Doorway Velocity 5 cm 48 -5 5 5 cm from floor, doorway CL 113 VD5C B 8 m/s

Soot Volume Fraction from Optical Measurement 48 -5 75 6 cm from Soffit, doorway CL 114 Fv C 0 n/a

RSEmax=(95,142,98)

LFL MIDAS Hookup Sheet Instrument and Channel Description Underventilated RSE Revision Date: 3/29/2006
144

Series:

Measurements

x x0=inside 

left wall 

(looking at 

door)

y y0=inside 

front (door) 

wall

z z0=floor

Location Description

Overall 

Channel 

Number

Abbr. MIDAS Station Module: Channels Mod. Mod. Ch. No. Conv.U nits Wire Gain

Main Channels

O2 at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 0 O2Rear East 1:0-31 1 0 Vol Fr Cu 1

Bad channel 1 Bad 1 1CO2 at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 2 CO2Rear East 1 2 Vol Fr Cu 1CO at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 3 CORear East 1 3 Vol Fr Cu 1

UH at Rear Sampling location 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 4 UHRear East 1 4 Vol Fr Cu 1Rear Gas Rack Dewpoint outside RSE 5 DPRear East 1 5 °C Cu 1Rear Sampling location gas temperature 1-near probe outside RSE 6 TGasRear1 East 1 6 °C TC 100Rear Sampling location gas temperature 2-near UH outside RSE 7 TGasRear2 East 1 7 °C TC 100Rear Sampling location water inlet temperature outside RSE 8 TH2OInRear East 1 8 °C TC 100

Rear Sampling location water outlet temperature outside RSE 9 TH2OOutRear East 1 9 °C TC 100O2 at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 10 O2Front East 1 10 Vol Fr Cu 1

CO2 at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 11 CO2Front East 1 11 Vol Fr Cu 1

CO at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 12 COFront East 1 12 Vol Fr Cu 1

UH at Front Sampling location 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 13 UHFront East 1 13 Vol Fr Cu 1Front Gas Rack Dewpoint outside RSE 14 DPFront East 1 14 °C Cu 1

Front Sampling location gas temperature 1- near probe outside RSE 15 TGasFront1 East 1 15 °C TC 100Front Sampling location gas temperature 2- near UH outside RSE 16 TGasFront2 East 1 16 °C TC 100Front Sampling location water inlet temperature outside RSE 17 TH2OInFront East 1 17 °C TC 100

Front Sampling location water outlet temperature outside RSE 18 TH2OOutFront East 1 18 °C TC 100Liquid Burner Temperature 1 cm 48 71 9 1 cm from burner bottom 19 TLiq1 East 1 19 °C TC 100

Liquid Burner Temperature 2 cm 48 71 10 2 cm from burner bottom 20 TLiq2 East 1 20 °C TC 100Liquid Burner Temperature 3 cm 48 71 11 3 cm from burner bottom 21 TLiq3 East 1 21 °C TC 100Liquid Burner Temperature 4 cm 48 71 12 4 cm from burner bottom 22 TLiq4 East 1 22 °C TC 100Liquid Burner Temperature 5 cm 48 71 13 5 cm from burner bottom 23 TLiq5 East 1 23 °C TC 100Liquid Burner Coolant Temperature Inlet outside RSE 24 TCoolIn East 1 24 °C TC 100

Liquid Burner Coolant Temperature Outlet outside RSE 25 TCoolOut East 1 25 °C TC 100Temperature in space under floor under RSE 26 TAmbFuel East 1 26 °C TC 100Tamb At MIDAS East station 27 Tamb East 1 27 °C TC 100Optical Soot (I/Io in adj file) 48 -5 75 6 cm from Soffit, doorway CL 28 OptSoot East 1 28 V Cu 1

Optical Soot Laser Reference Signal 48 -5 75 6 cm from Soffit, doorway CL 29 OptSootRef East 1 29 V Cu 1Optical Soot Apparatus Temperature outside RSE 30 TOptSoot East 1 30 °C TC 100

Gravimetric Soot Probe 1 Filter Temperature - rear outside RSE 31 TGravSoot1 East 1 31 °C TC 100

Gravimetric Soot Probe 2 Filter Temperature -rear outside RSE 32 TGravSoot2 East 2:0-31 2 0 °C TC 100Gravimetric Soot Probe 1 Solenoid Signal -rear outside RSE 33 VGravSoot1 East 2 1 V Cu 1

Gravimetric Soot Probe 2 Solenoid Signal -rear outside RSE 34 VGravSoot2 East 2 2 V Cu 1Gravimetric Soot Probe 1 Mass Flow -rear outside RSE 35 SootMF1 East 2 3 V Cu 1

Gravimetric Soot Probe 2 Mass Flow-front outside RSE 36 SootMF2 East 2 4 V Cu 1Total Heat Flux Gauge Front (SN=131836) 48 35 0 In floor on room CL, 35 cm from Front 37 HFF East 2 5 kW/m2 Cu 100

Total Heat Flux Gauge Rear (SN=131835) 48 106 0 In floor on room CL, 35 cm from Rear 38 HFR East 2 6 kW/m2 Cu 100

Doorway Pressure 79 cm Left (+out) 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 39 PD79L East 2 7 Pa Cu 1

Doorway Pressure 79 cm CL (+out) 48 -5 79 79 cm from floor, doorway CL 40 PD79C East 2 8 Pa Cu 1Doorway Pressure 79 cm Right (+out) 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 41 PD79R East 2 9 Pa Cu 1Doorway Pressure 60 cm (+out) 48 -5 60 60 cm from floor, doorway CL 42 PD60C East 2 10 Pa Cu 1Doorway Pressure 40 cm (+in) 48 -5 40 40 cm from floor, doorway CL 43 PD40C East 2 11 Pa Cu 1Doorway Pressure 20 cm Left (+in) 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 44 PD20L East 2 12 Pa Cu 1Doorway Pressure 20 cm CL (+in) 48 -5 20 20 cm from floor, doorway CL 45 PD20C East 2 13 Pa Cu 1Doorway Pressure 20 cm Right (+in) 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 46 PD20R East 2 14 Pa Cu 1Doorway Pressure 5 cm (+in) 48 -5 5 5 cm from floor, doorway CL 47 PD5C East 2 15 Pa Cu 1Doorway Temperature 79 cm Left BB 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 48 TD79LBB East 2 16 °C TC 100Doorway Temperature 79 cm CL BB 48 -5 79 79 cm from floor, doorway CL 49 TD79CBB East 2 17 °C TC 100Doorway Temperature 79 cm Right BB 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 50 TD79RBB East 2 18 °C TC 100

RSEmax=(95,142,98)

LFL MIDAS Hookup Sheet Instrument and Channel Description Underventilated RSE Revision Date: 8/14/2006
145

Series:

Measurements

x x0=inside 

left wall 

(looking at 

door)

y y0=inside 

front (door) 

wall

z z0=floor

Location Description

Overall 

Channel 

Number

Abbr. MIDAS Station Module: Channels Mod. Mod. Ch. No. Conv.U nits Wire Gain

Main Channels

Doorway Temperature 70 cm Left Asp 32 -5 70 70 cm from floor, doorway 8 cm from Left side of dr 51 TD70LA East 2 19 °C TC 100Doorway Temperature 70 cm Left BB 32 -5 70 70 cm from floor, doorway 8 cm from Left side of dr 52 TD70LBB East 2 20 °C TC 100Doorway Temperature 70 cm CL Asp 48 -5 70 70 cm from floor, doorway CL 53 TD70CA East 2 21 °C TC 100Doorway Temperature 70 cm CL BB 48 -5 70 70 cm from floor, doorway CL 54 TD70CBB East 2 22 °C TC 100Doorway Temperature 30 cm Left Asp 32 -5 30 30 cm from floor, doorway 8 cm from Left side of dr 55 TD30LA East 2 23 °C TC 100Doorway Temperature 30 cm Left BB 32 -5 30 30 cm from floor, doorway 8 cm from Left side of dr 56 TD30LBB East 2 24 °C TC 100Doorway Temperature 60 cm CL BB 48 -5 60 60 cm from floor, doorway CL 57 TD60CBB East 2 25 °C TC 100

Doorway Temperature 50 cm CL Asp 48 -5 50 50 cm from floor, doorway CL 58 TD50CA East 2 26 °C TC 100Doorway Temperature 50 cm CL BB 48 -5 50 50 cm from floor, doorway CL 59 TD50CBB East 2 27 °C TC 100Doorway Temperature 40 cm CL BB 48 -5 40 40 cm from floor, doorway CL 60 TD40CBB East 2 28 °C TC 100Doorway Temperature 30 cm CL Asp 48 -5 30 30 cm from floor, doorway CL 61 TD30CA East 2 29 °C TC 100Doorway Temperature 30 cm CL BB 48 -5 30 30 cm from floor, doorway CL 62 TD30CBB East 2 30 °C TC 100Doorway Temperature 20 cm Left BB 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 63 TD20LBB East 2 31 °C TC 100Doorway Temperature 20 cm CL BB 48 -5 20 20 cm from floor, doorway CL 64 TD20CBB East 4:0-20 4 0 °C TC 100Doorway Temperature 20 cm Right BB 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 65 TD20RBB East 4 1 °C TC 100Doorway Temperature 5 cm CL BB 48 -5 5 5 cm from floor, doorway CL 66 TD5CBB East 4 2 °C TC 100Rear Temperature at Sampling Location Asp 29 113 88 10 cm from C, 29 cm from Left, 29 cm from Rear 67 TRSampA East 4 3 °C TC 100Rear Temperature low Asp 75 122 24 24 cm from Floor, 20 cm from Right, 20 cm from Rear 68 TR24A East 4 4 °C TC 100Rear Temperature high Asp 75 122 80 80 cm from Floor, 20 cm from Right, 20 cm from Rear 69 TR80A East 4 5 °C TC 100Front Temperature at Sampling Location Asp 29 10 88 10 cm from C, 29 cm from Left, 10 cm from Front 70 TFSampA East 4 6 °C TC 100Front Temperature low Asp 75 20 24 24 cm from Floor, 20 cm from Right, 20 cm from Front 71 TF24A East 4 7 °C TC 100Front Temperature high Asp 75 20 80 80 cm from Floor, 20 cm from Right, 20 cm from Front 72 TF80A East 4 8 °C TC 100Temperature of Total Heat Flux Gauge Front (SN=131836) 48 35 0 In floor on room CL, 35 cm from Front 73 THFF East 4 9 °C TC 100Temperature of Total Heat Flux Gauge Rear (SN=131835) 48 106 0 In floor on room CL, 35 cm from Rear 74 THFR East 4 10 °C TC 100Gravimetric Soot Probe 3 Filter Temperature -front outside RSE 75 TGravSoot3 East 4 11 °C TC 100

Gravimetric Soot Probe 4 Filter Temperature-front outside RSE 76 TGravSoot4 East 4 12 °C TC 100Gravimetric Soot Probe 3 Solenoid Signal-front outside RSE 77 VGravSoot3 East 4 13 V Cu 1Gravimetric Soot Probe 4 Solenoid Signal-front outside RSE 78 VGravSoot4 East 4 14 V Cu 1Surface Temperature Back Wall Exterior Upper Left 20? 144? 87? 10 cm down, 20 cm from left (facing in doorway) 79 TWallUL East 4 15 °C TC 100

Surface Temperature Back Wall Exterior Upper Right 76? 144? 87? 10 cm down, 20 cm from right (facing in doorway) 80 TWallUR East 4 16 °C TC 100Surface Temperature Back Wall Exterior Lower Left 20? 144? 10? 10 cm up, 20 cm from left (facing in doorway) 81 TWallLL East 4 17 °C TC 100Surface Temperature Back Wall Exterior Lower Right 76? 144? 10? 10 cm up, 20 cm from right (facing in doorway) 82 TWallLR East 4 18 °C TC 100Surface Temperature Inside Floor Front 49? 35? 0 near front HF gauge 83 TFloorF East 4 19 °C TC 100Surface Temperature Inside Ceiling 31? 11? 98? near front sampling probe @ ceiling 84 TCeil East 4 20 °C TC 100Surface Temperature Inside Floor Rear near rear HF gauge 85 TFloorR East 4 21 °C TC 100

Created Channels

Event Marker 1 86 Event1Event Marker 2 87 Event2Doorway Velocity 79 cm Left 32 -5 79 79 cm from floor, doorway 8 cm from Left side of dr 88 VD79L B 0 m/sDoorway Velocity 79 cm CL 48 -5 79 79 cm from floor, doorway CL 89 VD79C B 1 m/sDoorway Velocity 79 cm Right 64 -5 79 79 cm from floor, doorway 8 cm from Right side of dr 90 VD79R B 2 m/sDoorway Velocity 60 cm 48 -5 60 60 cm from floor, doorway CL 91 VD60C B 3 m/sDoorway Velocity 40 cm 48 -5 40 40 cm from floor, doorway CL 92 VD40C B 4 m/sDoorway Velocity 20 cm Left 32 -5 20 20 cm from floor, doorway 8 cm from Left side of dr 93 VD20L B 5 m/sDoorway Velocity 20 cm CL 48 -5 20 20 cm from floor, doorway CL 94 VD20C B 6 m/sDoorway Velocity 20 cm Right 64 -5 20 20 cm from floor, doorway 8 cm from Right side of dr 95 VD20R B 7 m/sDoorway Velocity 5 cm 48 -5 5 5 cm from floor, doorway CL 96 VD5C B 8 m/sSoot Volume Fraction from Optical Measurement 48 -5 75 6 cm from Soffit, doorway CL 97 Fv C 0 n/a

RSEmax=(95,142,98)

LFL MIDAS Hookup Sheet Instrument and Channel Description Underventilated RSE Revision Date: 8/14/2006
146

Description of data columns (channels) in reduced ASCII data files. 

Channel Name (units) Description of Measurement

1 Time From Ignition (s) NA NA NA Time relative to ignition defined by Event 1

2 O2Rear (mol/mol) 29 113 88 Rear O2 volume fraction corrected for water (wet)

3 CO2Rear (mol/mol) 29 113 88 Rear CO2 volume fraction corrected for water (wet)

4 CORear (mol/mol) 29 113 88 Rear CO volume fraction corrected for water (wet)

5 THCRear (mol/mol) 29 113 88 Rear Total Hydrocarbons volume fraction corrected for water (wet)

6 SootRear (g/g) 29 113 88 Rear Soot Mass fraction corrected for water (wet)

7 O2Front (mol/mol) 29 10 88 Front O2 volume fraction corrected for water (wet)

8 CO2Front (mol/mol) 29 10 88 Front CO2 volume fraction corrected for water (wet)

9 COFront (mol/mol) 29 10 88 Front CO volume fraction corrected for water (wet)

10 THCFront (mol/mol) 29 10 88 Front Total Hydrocarbons volume fraction corrected for water (wet)

11 SootFront (g/g) 29 10 88 Front Soot Mass fraction corrected for water (wet)

12 HFR (kW/m2) 48 106 0 Total Heat Flux at Rear Floor

13 HFF (kW/m2) 48 35 0 Total Heat Flux at Front Floor

14 TambCal (C) NA NA NA Ambient Temperature in LFL

15 TRSampA (C) 29 113 88 Aspirated thermocouple at rear sample location

16 TFSampA (C) 29 10 88 Aspirated thermocouple at front sample location

17 TR24A (C) 75 122 24 Aspirated thermocouple in RSE

18 TR80A (C) 75 122 80 Aspirated thermocouple in RSE

19 TF24A (C) 75 20 24 Aspirated thermocouple in RSE

20 TF80A (C) 75 20 80 Aspirated thermocouple in RSE

21 TFloorR (C) 49 106 0 Bare-bead thermocouple near heat flux gauge

22 TFloorF (C) 49 35 0 Bare-bead thermocouple near heat flux gauge

23 TCeilF (C) 31 11 98 Bare-bead thermocouple on ceiling

24 TWallUR (C) 19 144 87 Bare-bead thermocouple on outside rear wall

25 TUFloor (C) Bare-bead thermocouple in space below floor

26 TD79LBB (C) 32 -5 79 Bare-bead thermocouple in doorway

27 TD79CBB (C) 48 -5 79 Bare-bead thermocouple in doorway

28 TD79RBB (C) 64 -5 79 Bare-bead thermocouple in doorway

29 TD70LA (C) 32 -5 70 Aspirated thermocouple in doorway

30 TD70LBB (C) 32 -5 70 Bare-bead thermocouple in doorway

31 TD70CA (C) 48 -5 70 Aspirated thermocouple in doorway

32 TD70CBB (C) 48 -5 70 Bare-bead thermocouple in doorway

33 TD30LA (C) 32 -5 30 Aspirated thermocouple in doorway

34 TD30LBB (C) 32 -5 30 Bare-bead thermocouple in doorway

35 TD60CBB (C) 48 -5 60 Bare-bead thermocouple in doorway

36 TD50CA (C) 48 -5 50 Aspirated thermocouple in doorway

37 TD50CBB (C) 48 -5 50 Bare-bead thermocouple in doorway

38 TD40CBB (C) 48 -5 40 Bare-bead thermocouple in doorway

39 TD30CA (C) 48 -5 30 Aspirated thermocouple in doorway

40 TD30CBB (C) 48 -5 30 Bare-bead thermocouple in doorway

41 TD20LBB (C) 32 -5 20 Bare-bead thermocouple in doorway

42 TD20CBB (C) 48 -5 20 Bare-bead thermocouple in doorway

43 TD20RBB (C) 64 -5 20 Bare-bead thermocouple in doorway

44 TD5CBB (C) 48 -5 5 Bare-bead thermocouple in doorway

45 VD79L (m/s) 32 -5 79 Bi-directional probe velocity in doorway

46 VD79C (m/s) 48 -5 79 Bi-directional probe velocity in doorway

47 VD79R (m/s) 64 -5 79 Bi-directional probe velocity in doorway

48 VD60C (m/s) 48 -5 60 Bi-directional probe velocity in doorway

49 VD40C (m/s) 48 -5 40 Bi-directional probe velocity in doorway

50 VD20L (m/s) 32 -5 20 Bi-directional probe velocity in doorway

51 VD20C (m/s) 48 -5 20 Bi-directional probe velocity in doorway

52 VD20R (m/s) 64 -5 20 Bi-directional probe velocity in doorway

53 VD5C (m/s) 48 -5 5 Bi-directional probe velocity in doorway

54 HRRcal (kW) Heat Release Rate from Calorimeter

55 HRRburner (kW) Heat Release Rate from Burner (gas, pool or spray)

56 StackMFR (kg/s) Exhaust hood mass flow rate

57 Tstack (C) Exhaust hood temperature (near bi-directional probe)

58 O2stack (mol/mol) Exhaust O2 volume fraction (dry)

59 CO2stack (mol/mol) Exhaust CO2 volume fraction (dry)

60 COstack (mol/mol) Exhaust CO volume fraction (dry)

61 THCstack (mol/mol) Exhaust total hydrocarbons volume fraction (dry)

62 MSstack (mg/m3) Exhaust soot mass concentration (wet)

63 H2ORear (mol/mol) 29 113 88 Rear water volume fraction (calculated from CO/CO2)

64 H2OFront (mol/mol) 29 10 88 Front water volume fraction (calculated from CO/CO2)

Position (cm) 

RSEmax=(95,142,98)
147

 D. Equipment List

Description Manufacturer Model Serial# NIST# 

Oxygen analyzer for HRR Servomex 540A 549709 

CO2/CO analyzer for HRR Seimens Ultramat 6 615207 

Total HC analyzer for HRR Rosemount 400A 569041 

Mass flow controller for HRR MKS 1179A53C 000346712 

Dew Point Transmitter for HRR Vaisala DMT242 A4850006 

Sample dryer for HRR PermaPure PD-200T-72SS 973-0905-6 

Micro GC for natural gas Agilent 3000A 623489 

Sample pump for HRR Gast MOA-P122-AA 4Z026 

Liquid fuel turbine flow meter Exact Flow 

Natural gas flow meter Instromet IRMA 15M-125 319396 605032 

Total heat flux gauge (HF Front) Medtherm 16-0.75-10-4-12-

36-20679k 131836 

Total heat flux gauge (HF Rear) Medtherm 16-0.75-10-4-12-

36-20679k 131835 

Oxygen Analyzer (O2Rear) Servomex 4100 393063 623487 

Oxygen Analyzer (O2Front) Servomex 4100 393064 623488 

CO2/CO Analyzer (Rear) Seimens Ultramat 6E NI-L00197 600671 

CO2/CO Analyzer (Front) Seimens Ultramat 6E 609425 

Total HC analyzer (Rear) Baseline-Mocon 8800 H 625764 

Total HC analyzer (Front) Baseline-Mocon 8800 H 623892 

Dew point meter (Rear) Vaisala DMT242 B074008 

Dew point meter (Front) Vaisala DMT242 B074009 

Mass flow controller (Rear&Front) MKS M100B53C 

MFC power supply MKS 247D 

Gas Chromatograph (Front) HP 5890A 2843A20868 541824 

GC column Restek Rt-QPLOT 19718 

Pressure Transducer for Velocity MKS 220DD 

Flow meter (spot check flows) Bios Dry Cal DCLT 20K 

Sample pump for aspirated TCs 

and gas sample tests #1-6 Gast DOA-P703-FB 

Gas conditioning system 

(Rear&Front) PermaPure MG-2812 rental 

Glass-lined stainless steel tubing Grace Davison 3149 

Soot sample MFC MKS M100B53CCS1BV 021407828 

Soot sample MFC MKS M100B53CCS1BV 021407829 

MFC power supply MKS 247D 000763015 

Soot sample filter Pall P5PJ047 

Soot sample cleaning pad Hoppe’s 1203 

Soot sample filter holder Gelman Sciences 2220 

Soot sample 3-way solenoid 

valves Parker 04F30C2208AAF4C05 

Soot sample pumps Gast MOA-P122-AA