NIST NCSTAR1-5B

Federal Building and Fire Safety Investigation of the

World Trade Center Disaster

Experiments and Modeling of

Structural Steel Elements

Exposed to Fire

Anthony Hamins

Alexander Maranghides

Kevin B. McGrattan

Erik L. Johnsson

Thomas J. Ohiemiller

Michelle K. Donnelly

Jiann C. Yang

George W. Mulholland

Kuldeep R. Prasad

Scott R. Kukuck

Robert L. Anieitner

Therese P. McAllister

National Insfilute of Standards and Technology • Technology Administration • U.S. Deportment of Commerce

NISTNCSTAR1-5B

Federal Building and Fire Safety Investigation of the

World Trade Center Disaster

Experiments and Modeling of

Structural Steel Elements

Exposed to Fire

Anthony Hamins

Alexander Maranghides

Kevin B. McGrattan

Erik L. Johnsson

Thomas J. Ohiemiller

Michelle K. Donnelly

Jiann C. Yang

George W. Mulholland

Kuldeep R. Prasad

Scott R. Kukuck*

Robert L, Anleitner

Therese P. McAllister

Building and Fire Research Laboratory

National Institute of Standards and Technology

*Scott R. Kukuck's current affiliation; U.S. Army Research Laboratory

Aberdeen Proving Grounds, Maryland

September 2005

U.S. Department of Commerce

Carlos M. Gutierrez, Secretary

Technology Administration Michelle O'Neill, Acting Under Secretary for Technology

National Institute of Standards and Technology

William Jeffrey, Director
Disclaimer No. 1

Certain commercial entities, equipment, products, or materials are identified in this document in order to describe a procedure or concept adequately or to trace the history of the procedures and practices used. Such identification is not intended to imply recommendation, endorsement, or implication that the entities, products, materials, or equipment are necessarily the best available for the purpose. Nor does such identification imply a finding of fault or negligence by the National Institute of Standards and Technology.

Disclaimer No. 2 The policy of NIST is to use the International System of Units (metric units) in all publications. In this document,

however, units are presented in metric units or the inch-pound system, whichever is prevalent in the discipline.

Disclaimer No. 3

Pursuant to section 7 of the National Construction Safety Team Act, the NIST Director has determined that certain evidence received by NIST in the course of this Investigation is "voluntarily provided safety-related information" that is "not directly related to the building failure being investigated" and that "disclosure of that information would inhibit the

voluntary provision of that type of information" (15 USC 7306c).

In addition, a substantial portion of the evidence collected by NIST in the course of the Investigation has been

provided to NIST under nondisclosure agreements.

Disclaimer No. 4

NIST takes no position as to whether the design or construction of a WTC building was compliant with any code

since, due to the destruction of the WTC buildings, NIST could not verify the actual (or as-built) construction, the

properties and condition of the materials used, or changes to the original construction made over the life of the

buildings. In addition, NIST could not verify the interpretations of codes used by applicable authorities in determining

compliance when implementing building codes. Where an Investigation report states whether a system was

designed or installed as required by a code provision, NIST has documentary or anecdotal evidence indicating

whether the requirement was met, or NIST has independently conducted tests or analyses indicating whether the requirement was met.

Use in Legal Proceedings

No part of any report resulting 'rom a NIST investigation into a structural failure or from an investigation under the

National Construction Safety Team Act may be used in any suit or action for damages arising out of any matter mentioned in such report (15 USC 281a; as amended by P.L. 107-231).

National Institute of Standards and Technology National Construction Safety Team Act Report 1-5B

Natl. Inst. Stand. Technol. Natl. Constr. Sfty. Tm. Act Rpt. 1-5B, 352 pages (September 2005)

CODEN: NSPUE2

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 2005

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

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

Mail: Stop SSOP, Washington, DC 20402-0001
Abstract

Reconstructing the fires and their impact on structural components in the World Trade Center (WTC)

buildings on September 11, 2001, requires extensive use of computational models. For the use of such

models to be a viable investigative tool, it is essential to know the accuracy with which they capture the physical phenomena of the fires and the concurrent heat transfer to the building structure. This report

documents a series of large-scale experiments that was conducted in the National Institute of Standards

and Technology (NIST) Large Fire Laboratory from March 10 to March 26, 2003. The experiments

represent one phase of an effort to ascertain the validity of the models for the NIST WTC Investigation.

The objective of the experiments was to assess the accuracy with which (1) the NIST Fire Dynamics

Simulator (FDS) fire model predicts the thermal environment in a burning compartment and (2) the NIST

Fire Structure Interface (FSI) model in combination with the ANSYS finite-element model predicts the temperature rise of structural steel components in a burning compartment. The experiments also had the

potential to improve input parameters in the modeling, if appropriate, and, in general, help to increase

understanding of the sequence of events that occurred in the WTC tower fires.

Within a steel-frame compartment (3 m by 7 m by 4 m) lined with calcium silicate board were placed four

steel components: two trusses, one thin-walled column, and a rod. The components either were

uninsulated or had fibrous sprayed fire-resistive material (SFRM) applied; two thicknesses were tested. The 2 MW and 3 MW fires were generated using liquid hydrocarbon fuels introduced by a two-nozzle

spray burner onto a 1 m by 2 m pan. The fuels were a commercial blend of heptane isomers and a mixture of the heptane blend with toluene. Six experiments were conducted.

Measurements were made of a large number of parameters, including the heat release rate; the radiative

and total heat flux at various targets in the compartment, including locations near the floor and the ceiling; the temperatures of insulated and uninsulated steel components; the concentrations of soot, oxygen,

carbon monoxide, and carbon dioxide; and the vertical profiles of gas-phase temperature. A video record was made from several orientations. Nearly 350 channels of data were acquired. Following the establishment of baseline signals from all of the measurement devices, the burner was ignited and burned

at a steady rate. The test continued until the surface temperature of any one of several steel components

present in the compartment reached approximately 600 °C.

For each test, a prediction of the thermal environment in the compartment was determined using the NIST

FDS fire model. A prediction of the rise of the steel temperature for the simulated themial environment was made using the NIST FSI. A comparison of the FDS predicfions and the experimental results showed that the predictions were within experimental and model uncertainty. The steel temperatures

predicted by the FSI and the ANSYS finite-element models compared favorably with the measurements

within experimental uncertainty and model sensitivity.

These experimental tests of the methodology using the FDS, the FSI, and the finite-element model to analyze the thermal envirormient in compartments experiencing fire and the associated temperature rise in

steel components provide the level of confidence necessary to apply the modeling methodology to the

investigation of the WTC disaster.

NISTNCSTAR 1-5B, WTC Investigation 111
Abstract

Keywords: Finite-element model, fire measurements, fire model, thermal-structural model, World Trade

Center. .

IV NISTNCSTAR 1-5B, WTC Investigation
Table of Contents

Abstract iii

List of Figures xi

List of Tables xxiii

List of Acronyms and Abbreviations xxv

Metric Conversion Table xxvii

Preface xxxi

Acknowledgments xli Executive Summary xliii

Chapter 1 Introduction 1

1 . 1 Modeling as an Investigative Tool 1

1 .2 Over\'iew of Models 1

1 .3 Overx'iew of Experiments 3

1 .4 Test of Model Accuracy 4

1 .5 Organization of This Report 5

1.6 References 6

Chapter 2 Experimental Configuration and General Observations 7

2. 1 Test Facility Description 7

2.2 Compartment Size and Construction 7

2.3 Compartment Contents 11

2.3.1 Structural Components 11

2.3.2 Instrumentation Supports 1

1

2.3.3 Air Deflecting Baffle 11

2.3.4 Natural Ventilation 1

1

2.4 Fire and Fuel Flow 12

2.5 Test Matrix 12

2.6 Instrumentation 13

2.7 Observations 15

2.8 Test Procedure 17

2.9 References 18

NIST NCSTAR 1-5B, WTC Investigation v
Table of Contents

Chapters

Fuel Flow and Heat Release Rate.. 19

3.1 Fuel Flow and Estimated Heat Release Rate 19

3.2 Oxygen Consumption Calorimetry 20

3.2. 1 Description of the Heat Release Rate Measurement Facility 20

3.2.2 Calibration and Measurement Uncertainty 2

1

3.3 Results and Discussion 22

3.4 References 23

Chapter 4

Gas Temperature 25

4.1 Instrumentation 25

4.2 Measurement Uncertainties 25

4.3 Measurement Results 3

1

4.3.1 Time Dependence 31

4.3.2 Effect of Heat Release Rate 31

4.3.3 Effect of Smoke Production 32

4.3.4 Repeatability 32

4.4 Layer Interface Height 36

4.4. 1 Calculation Uncertainties 36

4.4.2 Calculation Results 37

4.5 References 40

Chapter 5 Heat Flux Measurements 41

5.1 Objective 41

5.2 Description of the Heat Flux Measurement Devices 41

5.2.1 Location of Heat Flux Instrumentation 45

5.2.2 Coolant System for Heat Flux Gauges 50

5.3 Data Reduction for the Slug Calorimeters 51

5.4 Uncertainties in the Heat Flux Measurements 52

5.5 Discussion of Heat Flux Measurements 54

5.5.1 Repeatability 62

5.5.2 Effect of Fire Size 62

5.5.3 Effect of Soot 63

5.6 References 65

vi NIST NCSTAR 1-5B, WTC Investigation
Table of Contents

Chapter 6 Smoke, CO2, CO, and O2 Concentrations 67

6. 1 Instrumentation 67

6.2 Measurement Uncertainty for Gas Analyzer Results 69

6.3 Smoke and Gaseous Yields 7

1

6.4 Measurement Uncertainty for Soot, CO;, and CO Yields 72

6.5 Experimental Observations 74

6.6 Repeatability of Tests 75

6.7 Effect of Smoke Concentration 78

6.8 Effect of Heat Release Rate for a Low Sooting Fuel 79

6.9 Effect of the Fire Heat Release Rate on a High Sooting Fuel 81

6.10 References 83

Chapter 7 Velocity Measurements 85

7. 1 Instrumentation 85

7.1.1 Sampling Locations 86

7. 1 .2 Thermocouple Radiation Correction 87

7.2 Measurement Uncertainty 88

7.3 Discussion and Results 89

7.4 Radiation Correction 94

7.5 References 96

Chapter 8 Compartment Surface Temperature Measurements 97

8. 1 Instrumentation 97

8.2 Measurement Uncertainty 97

8.3 Discussion of Results 102

8.4 Repeatability 1 1

1

8.5 Effect of Heat Release Rate Ill

8.6 Effect of Fuel Type 113

8.7 References 1 13

Chapter 9 Temperatures of Steel Components 115

9.1 Objective 115

9.2 Experimental Description 115

NISTNCSTAR 7-56, WTC Investigation vii
Table of Contents

9.2.1 SFRM Thickness 123

9.3 Description of the Temperature Measurements 131

9.4 Measurement uncertainty 132

9.4. 1 Finite-Element Analysis 132

9.4.2 Simulation Results 134

9.5 Measurement Uncertainty of the SFRM Surface Temperature 136

9.6 Description of Experimental Results 136

9.7 Summary of Results 146

9.8 References 147

Chapter 10

Simulation of Compartment Fire Phenomena 149

10.1 Introduction 149

10.2 Description of the Calculations 149

10.3 Discussion of Model Sim.ulations and Measurements 151

10.3.1 Gas Temperatures 152

10.3.2 Heat Flux to Solid Objects 157

10.3.3 Compartment Surface Temperatures 165

10.3.4 Gas Species Concentrations 171

10.3.5 Inlet and Outlet Velocities 176

10.4 Sensitivity Analysis and Comparison to Measurements 178

10.4.1 Sensitivity of Gas Temperatures to Heat Release Rate 178

10.4.2 Sensitivity of Heat Flux to Solid Objects to Heat Release Rate 179

10.4.3 Sensitivity of Compartment Surface Temperature to Heat Release Rate 179

10.4.4 Sensitivity of Gas Species Concentration to Heat Release Rate 180

10.4.5 Sensitivity of Inlet and Outlet Velocities to Heat Release Rate 180

10.4.6 Model Sensitivity to Parameters other than Heat Release Rate 180

10.5 Coupling with Thermal-Strucmral Models 184

10.6 Summary 184

10.7 References 185

Chapter 11 Simulation of the Thermal Response of Structural Components 187

11.1 Introduction 187

11.2 Methodology 187

1 1.3 Comparison of Simulations with Measurements 194

viii NIST NCSTAR 1-5B, WTC Investigation
Table of Contents

11.3.1 Test 1 194

11.3.2 Test 5 203

1 1 .4 Model Sensitivity 212

1 1 .4. 1 Sensitivity to Heat Release Rate 2 1

7

1 1 .4.2 Sensitivity to Grid Size 222

1 1 .5 Sensitivity to the Thickness and Thermo-Physical Properties of the Sprayed Fire-Resistive

Material 223

1 1.6 Summary 224

11.7 References 225

Chapter 12

Summary of Results 227

Appendix A Selection of Fuels 231

Appendix B Instrument Locations 233

Appendix C Gas-Phase Temperature Measurements 243

Appendix D Velocity Measurements at the Compartment Inlet and Outlet 251

Appendix E Heat Flux Measurements 273

Appendix F Thermocouple Locations on Steel Components 285

Appendix G Temperature Histories of Bare and Coated Steel Components 289

NISTNCSTAR 7-56, WTC Investigation IX
Table of Contents

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X NISTNCSTAR 1-5B, WTC Investigation
List of Figures

Figure P-1. The eight projects in the federal building and fire safety investigation of the WTC

disaster xxxiii

Figure 2-1 . Compartment contents as seen through the access panel on the south wall 7

Figure 2-2. Experimental enclosure contents 8

Figure 2-3. Experimental enclosure contents viewed with access panels removed 9

Figure 2-4. Fuel pan, fuel nozzles, and instrumentation in the compartment 9

Figure 2-5. Schematic drawing of compartment contents 10

Figure 2-6. Instrumentation on the outside of the west wall of the compartment 14

Figure 2-7. Fire impinging on the bare steel trusses 10 s after ignition in Test 1 16

Figure 2-8. Smoke exiting the outlet vent into the exhaust hood during Test 2 16

Figure 2-9. Radiating hot gases exiting the outlet vent into the exhaust hood during Test 5 17

Figure 3-1 . The measured heat release rate ( O) as a function of time during Tests 1 through 6 23

Figure 4-1. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 1 28

Figure 4-2. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 2 29

Figure 4—3. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 3 29

Figure 4-4. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 4 30

Figure 4-5. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 5 30

Figure 4-6. 20 s average of the difference of bare-bead thermocouple temperatures minus

aspirated thermocouple temperatures as a function of time for Test 6 31

Figure 4-7. Aspirated thermocouple temperatures as a function of time for Test 1 33

Figure 4-8. Aspirated thermocouple temperatures as a function of time for Test 5 33

Figure 4-9. Comparison of Test 1 and Test 4 aspirated thermocouple temperatures at the upper

and middle locations as a fiinction of time, showing the effect of heat release rate 34

Figure 4-10. Comparison of Test 2 and Test 3 aspirated thermocouple temperatures at the upper

and middle locations as a function of time, showing the effect of heat release rate for a sooty fuel 34

NISTNCSTAR 1-5B, WTC Investigation xi
List of Figures

Figure 4-11. Comparison of Test 1 and Test 3 aspirated thermocouple temperatures at the upper

and middle locations as a function of time, showing the effect of soot for nearly the same heat release rate 35

Figure 4-12. Comparison of Test 5 and Test 6 aspirated thermocouple temperatures at the upper

and middle locations as a function of time, showing the repeatability of the

measurements for long tests with relatively large Q 35

Figure 4-13. West thermocouple tree temperatures as a function of time in Test 1 36

Figure 4-14. Calculated layer interface height as a function of time for Test 1 38

Figure 4-15. Temperature profiles for West and East thermocouple trees for Test 1 at t = 850 s 39

Figure 4-16. Temperature profiles for West and East thermocouple trees for Test 5 at t = 3,020 s 39

Figure 5-1 . Flux gauge holder showing gauge sensor surface (black disc) and soot purge tube. Surface of gauge holder and gauge sensor surface were set flush with the surrounding

Marinite cover on any monitored surface 43

Figure 5-2. Schematic of slug calorimeter 43

Figure 5-3. Front view of test compartment 46

Figure 5^. Top view of test compartment 47

Figure 5-5. Expanded sisde view of circled area in Figure 5-4 47

Figure 5-6. Locations of flux gauges and radiometers in the measurement station 48

Figure 5-7. Schematic showing the locations of column heat flux gauges 48

Figure 5-8. Cooling water system for radiometers and flux gauges 50

Figure 5-9. Total heat flux recorded by two flux gauges 15 cm above the floor, just downstream

of the fire pan in Test 5 55

Figure 5-10. Total heat flux recorded by the four flux gauges in the compartment ceiling in Test 5. The periodic spikes are due to the soot purge operation 56

Figure 5-11. Total heat flux recorded by the four flux gauges in structures adjacent to the vertical

column in Test 5. The periodic spikes are due to the soot purge operation on the two

higher gauges 58

Figure 5-12. Average total heat flux to the cylindrical surface of the slug calorimeter 1 in Test 5 58

Figure 5-13. Average total heat flux to the cylindrical surface of slug calorimeter 2 in Test 5 59

Figure 5-14. Comparison of the average total heat flux to the cylindrical

surface of slug calorimeter 3 in Tests 5 and 6 59

Figure 5-15. Comparison of the average total heat flux to the

cylindrical surface of slug calorimeter 4 in Tests 5 and 6 60

Figure 5-16. Comparison of the average total heat flux to the

cylindrical surface of slug calorimeter 5 in Tests 5 and 6 60

Figure 5-17. Heat fluxes recorded by the radiometers and total flux gauges in the flux

measurement station during Test 5. The periodic spikes are due to the soot purge

operation 62

xii NISTNCSTAR 1-5B, WTC Investigation
List of Figures

Figure 5-18. Effect of fire size in Tests 1 and 5 (2 MW and 3 MW, respectively) on total heat flux

at ceiling gauge HCe2, directly above the fire pan 64

Figure 5-19. Effect of fire size in Tests 1 and 5 (2 MW and 3 MW, respectively) on the recorded

total heat flux at slug calorimeter 1 , upstream of the fire plume 64

Figure 6—1. Schematic of apparatus for measuring the smoke yield and smoke concentration at location "Sm." The sampling tube extends from the inlet through the water jacket to the filter holder 68

Figure 6-2. Oxygen volume fraction (on a dry basis) sampled at location E

plotted as a function of time for Test 1 75

Figure 6-3. Comparison of the oxygen volume fraction (on a dry basis) for Tests 4 and 5 76

Figure 6-4. Comparison of CO: volume fraction (on a dry basis) for Tests 4 and 5. Resuhs are presented for the E sampling location and the Sm location, for which the gas is sampled twice for 1 20 s each 76

Figure 6-5. Ratio of volume fraction of CO2 minus the background value to the decrease in the

O2 volume fraction for Tests 4 and 5 at sample location E (on a dry basis) 77

Figure 6-6. Comparison of CO volume fraction (on a dry basis) for Tests 4 and 5 77

Figure 6-7. CO; and O; volume fraction (on a dry basis for low (Test 1) and high sooting

(Test 3) 78

Figure 6-8. Temperature as a function of time near the Sm probes

for low (Test 1 ) and high sooting (Test 3) 79

Figure 6-9. CO volume fraction (on a dry basis) as a function of time for low (Test 1) and high sooting (Test 3) 80

Figure 6-10. The oxygen volume fraction (on a dry basis) for a lower (Test 1

and a higher (Test 4) heat release rate 80

Figure 6-11. The CO: \'olume fraction (on a dry basis) for a lower (Test 1) and

a higher (Test 4) heat release rate 81

Figure 6-12. CO volume fraction (on a dry basis) for a lower (Test 1 ) and

a higher (Test 4) heat release rate 82

Figure 6-13. CO volume fraction (on a dry basis) for high sooting tests with a low (Test 3) and

a high (Test 2) heat release rate 82

Figure 7-1 . Inlet velocity measurements for Test 1 , inlet velocity probe 1 89

Figure 7-2. Outlet velocity at location 1 for Test 1 and Test 4 90

Figure 7-3. Outlet velocity at location 10 for Test 3 and Test 5 90

Figure 7-4. Outlet velocities for Test 5 at probe locations 1, 2, and 3 92

Figure 7-5. Outlet velocities for Test 5 at outlet locations 6, 7, and 8 92

Figure 7-6. Outlet velocities for probes located 15 cm above vent opening. Test 3. Outlets 8 and

1 0 are the probes toward the sides of the compartment, while outlets 5 and 3 are closer to the center of the compartment 93

NISTNCSTAR 1-5B. WTC Investigation xiii
List of Figures

Figure 7-7. Outlet velocities for probes located 61 cm above vent opening in Test 3. Outlets 6

and 9 are toward the side, while outlets 4 and 1 are closer to the compartment center 93

Figure 7-8. Comparison of the measured temperature and corrected temperature

in Test 6 at outlet 9 94

Figure 7-9. Comparison of the component of temperature correction due to incoming

flux and the component of temperature correction due to radiative loss 95

Figure 7-10. Comparison of bare-bead temperature measurements with aspirated temperature

measurements in Test 5 at outlet 5 95

Figure 7-11. Comparison of the mean velocity computed using bare-bead

temperature measurements with the mean velocity computed using

aspirated temperature measurements in Test 5 at outlet 5 96

Figure 8-1 . A type K thermocouple attached to the Marinite surface 97

Figure 8-2. Simplified configuration used for the numerical analysis 98

Figure 8-3. Comparison of the measured and predicted ceiling surface thermocouple temperature

in Test 2. The measured gas temperature is also shown 100

Figure 8-4. The simulated thermocouple temperature minus the simulated Marinite surface

temperature for two values of the upper layer emittance in Test 2 100

Figure 8-5. Comparison of the measured and predicted ceiling surface thermocouple

temperatures for Test 5. The measured gas temperature is also shown 101

Figure 8-6. The simulated thermocouple temperature minus the simulated Marinite surface

temperature for two values of the upper layer emittance in Test 5 101

Figure 8-7. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 1 103

Figure 8-8. The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Test 1 103

Figure 8-9. The peak temperatures on the ceiling surface and sandwiched between Marinite

ceiling layers along the compartment centerline (x axis) during Test 1. 104

Figure 8-10. The measured ceiling surface temperatures as a function of time at seven locations

along the y axis at x = 456 cm during Test 5 105

Figure 8-1 1 . The measured peak ceiling surface temperatures as a function of location along the y

axis at two x locations (x = 456 cm and x = 612 cm) during Test 5 105

Figure 8-12. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 2 106

Figure 8-13. The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Test 2 106

Figure 8-14. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 3 107

xiv NISTNCSTAR 1-5B, WTC Investigation
List of Figures

Figure 8-15. The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Tests 107

Figure 8-16. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 4 108

Figure 8-17. The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Test 4 108

Figure 8-18. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 5 109

Figure 8-19. The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Test 5 109

Figure 8-20. The measured ceiling temperatures as a function of time at 15 locations along the

compartment centerline (x axis) during Test 6 110

Figure 8-2 1 . The measured temperatures between the two Marinite sheets in the ceiling along the

compartment centerline (x axis) as a function of time at several locations during

Test 6 110

Figure 8-22. Comparison of the temperatures on the ceiling surface and sandwiched between the

two Marinite ceiling layers 3,000 s and 3,080 s after fire initiation, respectively,

along the compartment centerline (x axis) during Tests 5 and 6 Ill

Figure 8-23. Comparison of the temperatures on the ceiling surface and sandwiched between the

two Marinite ceiling layers 830 s and 1,035 s after fire initiation, respectively, along

the compartment centerline (x axis) during Tests 1 and 5 112

Figure 8-24. Comparison of the temperatures on the ceiling surface and sandwiched between the

two Marinite ceiling layers 350 s and 660 s after fire initiation, respectively, along

the compartment centerline (x axis) during Tests 2 and 3 112

Figure 8-25. Comparison of the temperatures on the ceiling surface and sandwiched between the

two Marinite ceiling layers 830 s and 1.040 s after fire initiation, respecfively, along

the compartment centerline (x axis) during Tests 1 and 3 113

Figure 9-1. Bar dimensions 116

Figure 9-2. Tubular column dimensions 116

Figure 9-3. Truss dimensions 1 16

Figure 9^a. Thermocouple locations for bars 118

Figure 9-4b. SFRM measurement locations for bars 118

Figure 9-5a. Thermocouple locations for columns 119

Figure 9-5b. SFRM measurement locations for columns in Test 5 and 6 1 19

Figure 9-6a. Thermocouple locafions on trusses 120

Figure 9-6b. Thermocouple bead spot welded onto the bottom chord of a steel truss with the wire

running close to the steel surface 120

NISTNCSTAR 7-56, WTC Investigation xv
List of Figures

Figure 9-6c. Color photograph of a thermocouple wire supported on a wire tie, with the

thermocouple bead placed just under the SFRM surface 121

Figure 9-6d. Application of SFRM product to a tmss 121

Figure 9-6e. Steel components with SFRM product prior to testing 122

Figure 9-6f SFRM measurement locations for trusses 122

Figure 9-7a. Steel component locations for Tests 1, 2, and 3 129

Figure 9-7b. Steel component locations for Tests 5 and 6 129

Figure 9-7c. Steel components in test compartment for Tests 1, 2, and 3 130

Figure 9-7d. Insulated steel components in test compartment for Tests 5 and 6 131

Figure 9-8. Simplified configuration used for the numerical analysis 133

Figure 9-9. Calculated measurement error for thermocouples on uninsulated steel 135

Figure 9-10. Temperature history for bar A in Test 3 139

Figure 9-11. Temperature history for column in Test 3 140

Figure 9-12. Temperature history for truss A in Test 3 140

Figure 9-13. Temperature history for bar in Test 5 141

Figure 9-14. Temperature history for column in Test 5 141

Figure 9-15. Temperature history for truss A in Test 5 142

Figure 9-16. Temperature history for column in Test 5 143

Figure 9-17. Temperature history for truss A in Test 5 143

Figure 9-18. Temperature history for truss B in Test 5 144

Figure 9-19. View of truss and bar components before Test 6 144

Figure 9-20. Flame impingement on trusses and bar during Test 6 145

Figure 10-1 . Snapshot of numerical simulation of Test 1 showing the fire as seen from a cutaway

view looking towards the north wall 150

Figure 10-2. Snapshot of the numerical simulation of Test 1 showing the fire and a contour plot of

gas temperature along the centerline of the compartment as seen from a cutaway

view looking towards the north wall 150

Figure 10-3. Comparison of gas temperatures, Test 1, intake side of compartment. The pumps

connected to the aspirated thennocouples began to operate at 180 s 152

Figure 10-4. Comparison of gas temperatures. Test 1, exhaust side of compartment. The pumps

connected to the aspirated thermocouples began to operate at 180 s 153

Figure 10-5. Comparison of gas temperatures. Test 2, intake side of compartment 153

Figure 10-6. Comparison of gas temperatures. Test 2, exhaust side of compartment 154

Figure 10-7. Comparison of gas temperatures. Test 3, intake side of compartment 154

Figure 10-8. Comparison of gas temperatures. Test 3, exhaust side of compartment 155

XVI NISTNCSTAR 1-5B, WTC Investigation
List of Figures

Figure 10-9. Comparison of gas temperatures, Test 4, intake side of compartment 155

Figure 10-10. Comparison of gas temperatures. Test 4, exhaust side of compartment 156

Figure 10-1 1. Comparison of gas temperatures. Test 5, intake side of compartment 156

Figure 10-12. Comparison of gas temperatures. Test 5, exhaust side of compartment 157

Figure 10-13. Comparison of heat fluxes to column, Test 1. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 158

Figure 10-14. Comparison of heat fluxes to column. Test 1. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 158

Figure 10-15. Comparison of heat fluxes to column. Test 2. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 159

Figure 10-16. Comparison of heat fluxes to column. Test 2. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 159

Figure 10-17. Comparison of heat fluxes to column. Test 3. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 1 60

Figure 10-18. Comparison of heat fluxes to column. Test 3. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 160

Figure 10-19. Comparison of heat fluxes to column. Test 4. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 161

Figure 10-20. Comparison of heat fluxes to column. Test 4. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 161

Figure 10-21. Comparison of heat fluxes to column, Test 5. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 162

Figure 10-22. Comparison of heat fluxes to column, Test 5. The periodic spikes

in the measurements were due to the influence of a nitrogen purge 1 62

Figure 10-23. Comparison of heat fluxes to floor targets. Test 1 163

Figure 10-24. Comparison of heat fluxes to floor targets. Test 2 163

Figure 10-25. Comparison of heat fluxes to floor targets. Test 3 164

Figure 10-26. Comparison of heat fluxes to floor targets. Test 4 164

Figure 10-27. Comparison of heat fluxes to floor targets. Test 5 165

Figure 10-28. Comparison of ceiling temperatures. Test 1 166

Figure 10-29. Comparison of inner ceiling temperatures. Test 1 166

Figure 10-30. Comparison of ceiling temperatures, Test 2 167

Figure 10-3 1 . Comparison of inner ceiling temperatures. Test 2 167

Figure 10-32. Comparison of ceiling temperatures. Test 3 168

Figure 10-33. Comparison of inner ceiling temperatures. Test 3 168

Figure 10-34. Comparison of ceiling temperatures, Test 4 169

Figure 10-35. Comparison of inner ceiling temperatures. Test 4 169

Figure 10-36. Comparison of ceiling temperatures. Test 5 170

NISTNCSTAR 1-5B. WTC Investigation xvii
List of Figures

Figure 10-37. Comparison of inner ceiling temperatures, Test 5 170

Figure 10-38. Comparison of gas concentrations on a dry basis, Test 1 171

Figure 10-39. Comparison of gas concentrations on a dry basis. Test 1 172

Figure 10-40. Comparison of gas concentrations on a dry basis. Test 2 172

Figure 10-41. Comparison of gas concentrations on a dry basis, Test 2 173

Figure 10-42. Comparison of gas concentrations on a dry basis. Test 3 173

Figure 10-43. Comparison of gas concentrations on a dry basis. Test 3 174

Figure 10^4. Comparison of gas concentrations on a dry basis. Test 4 174

Figure 10-45. Comparison of gas concentrations on a dry basis. Test 4 175

Figure 10^6. Comparison of gas concentrations on a dry basis. Test 5 175

Figure 10-47. Comparison of gas concentrations on a dry basis. Test 5 176

Figure 10-48. Comparison of velocity near top of vents, inlet and outlet in Test 1 176

Figure 10-49. Comparison of velocity near center of vents, inlet and outlet in Test 1 177

Figure 10-50. Comparison of velocity near bottom of vents, inlet and outlet in Test 1 177

Figure 10-51. Coarse grid prediction of gas temperatures on the intake side of the compartment for Test 5 182

Figure 10-52. Coarse grid prediction of gas temperatures on the exhaust side of the compartment

for Test 5 182

Figure 1 1-1 . FDS simulation of the upper layer temperature on the plane 3.01 m above the floor

(top figure) and the layer interface height (bottom figure) after 500 s in Test 1 189

Figure 11-2. FDS simulation of the predicted absorption coefficients on the plane 3.01 m above

the floor (top figure) and the lower layer temperature (bottom figure) after 500 s in Test 1 190

Figure 11-3. Temperature dependent thermophysical properties for A242 steel used in the thermal

analysis 191

Figure 1 1-4. Temperature dependent thermophysical properties of the SFRM used in the

experiments and thermal analysis 191

Figure 11-5. Temperature dependent thermophysical properties for Marinite 191

Figure 1 1-6. Finite-element model for the 3.05 m ( 10.0 ft) long steel bar, showing the elements

distributed uniformly along the length of the bar 1 92

Figure 1 1-7. Cross section of bar showing the elements used to model the steel (in blue) and the

SFRM (in violet) 193

Figure 1 1-8. Typical temperature contours through the cross-section of the steel bar and the

SFRM after 2,000 s in Test 5 193

Figure 1 1-9. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations on bar A in Test 1 ..195

xvin NISTNCSTAR 1-5B, WTC Investigation
List of Figures

Figure 1 1-10. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations on bar B in Test 1 195

Figure 11-11. Perspective and top view of the fmite-element model of the steel column with SFRM. ... 196

Figure 11-12. Temperature contours plotted on the surface of the SFRM and the ceiling after 2,000 s in Test 5 196

Figure 1 1-13. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 3.69 m above the floor on the column in Test 1 197

Figure 1 1-14. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 2.13 m above the floor on the column in Test 1 198

Figure 11-15. Finite-element model of the insulated steel truss, the ceiling and the floor used in the

thermal analysis of Test 5 198

Figure 1 1-16. Finite-element representation of the insulated steel truss (blue), the SFRM (violet),

and the ceiling (red) used in the thermal analysis of Test 5 199

Figure 1 1-17. Temperature contours (in K) on the ceiling, floor, and the surface of the SFRM on

truss A after 1,000 s in Test 5 199

Figure 1 1-18. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 2.89 m above the floor on truss A in Test 1 200

Figure 1 1-19. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 3.29 m above the floor on truss A in Test 1 200

Figure 1 1-20. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 3.70 m above the floor on truss A in Test 1 201

Figure 11-21. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 2.89 m above the floor on truss B in Test 1 201

Figure 1 1-22. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 3.29 m above the floor on truss B in Test 1 202

Figure 1 1-23. Comparison of numerical simulations with measurements for the steel surface

temperature at four locations 3.70 m above the floor on truss B in Test 1 202

Figure 1 1-24. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations on the insulated bar in Test 5 203

Figure 1 1-25. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations on the insulated bar in Test 5 204

Figure 1 1-26. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 0.77 m above the floor on the insulated steel column

in Test 5 204

Figure 1 1-27. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 2.13 m above the floor on the insulated steel column

in Test 5 205

Figure 1 1-28. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 3.69 m above the floor on the insulated steel coluirm

in Test 5 205

NISTNCSTAR 1-5B. WTC Investigation xix
List of Figures

Figure 1 1-29. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 0.77 m above the floor on the insulated column in Test 5 206

Figure 1 1-30. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 2.13 m above the floor on the insulated column in Tests 206

Figure 11-31. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 3.69 m above the floor on the insulated steel column in Tests 207

Figure 1 1-32. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 3.70 m above the floor on truss A in Test 5 207

Figure 1 1-33. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 3.29 m above the floor on truss A in Test 5 208

Figure 1 1-34. Comparison of numerical simulations with measurements for the temperature of the

SFRM surface at four locations 2.89 m above the floor on truss A in Test 5 208

Figure 1 1-35. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 3.70 m above the floor on truss A in Test 5 209

Figure 1 1-36. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 3.29 m above the floor on truss A in Test 5 209

Figure 11-37. Comparison of numerical simulations with measurements for the temperature of the

steel surface at four locations 2.89 m above the floor on truss A in Test 5 210

Figure 11-38. FDS simulation of the upper layer temperature on a plane 3.01 m above the floor

after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure) and

a crude mesh with 50 cm resolution (bottom figure) 213

Figure 1 1-39. FDS simulation of the layer interface height after 500 s in Test 1 using a heat release

rate increased by 10 percent (top figure) and a crude mesh with 50 cm resolution

(bottom figure) 214

Figure 1 1-40. FDS simulation of the absorption coefficient on a plane 3.01 m above the floor after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure) and a crude mesh with 50 cm resolution (bottom figure) 215

Figure 1 1-41. FDS simulation of the lower layer temperature after 500 s in Test 1 using a heat

release rate increased by 10 percent (top figure) and a crude mesh with

50 cm resolution (bottom figure) 216

Figure 1 1-42. Numerical predictions of the steel surface temperature on bar A in Test 1 for an FDS

calculation performed with a heat release rate 10 percent above baseline 217

Figure 1 1-43. Numerical predictions of the steel surface temperature on bar B in Test 1 for an FDS

calculation perfonned with a heat release rate 10 percent above baseline 218

Figure 1 1-44. Numerical predictions of the steel surface temperature for the column in Test 1 at 3.69 m above the floor for an FDS calculation performed with a heat release rate 10 percent above baseline 218

Figure 1 1-45. Numerical predictions of steel surface temperature for the column in Test 1 at 2.13 m

above the floor for an FDS calculation performed with a heat release rate 10 percent

above baseline 219

XX NIST NCSTAR 1-5B, WTC Investigation
List of Figures

Figure 11^6. Numerical predictions of steel surface temperature at four locations 2.89 m above the

floor on truss A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline 219

Figure 1 1-47. Numerical predictions of steel surface temperature at four locations 3.29 m above the

floor on truss A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline 220

Figure 1 1-48. Numerical simulations of the steel surface temperature at four locations 3.70 m above

the floor on truss A in Test 1 for an FDS calculation performed with a heat release

rate 10 percent above basehne 220

Figure 1 1-49. Predicted numerical results for steel temperature for bar A in Test 1 . These results correspond to an FDS calculation performed with a crude mesh (50 cm resolution) 222

Figure 1 1-50. Predicted numerical results for steel temperature on bar B in Test 1 . These results correspond to an FDS calculation performed with a crude mesh (50 cm resolution) 223

Figure 1 1-51 . Numerical predictions of steel surface temperature for bar A in Test 5 assuming a SFRM thickness 25 percent below baseline 224

NISTNCSTAR 1-5B, WTC Investigation xxi
List of Figures

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xxn NISTNCSTAR 1-5B, WTC Investigation
List of Tables

Table P-1 . Federal building and fire safety investigation of the WTC disaster xxxii

Table P-2. Public meetings and briefings of the WTC Investigation xxxv

Table 2-1 . Test matrix 12

Table 2-2. Measurements 13

Table 2-3. Instrument characterization 15

Table 3-1 . Fuel flow and expected heat release rate 20

Table 3-2. Combustion properties of the test fuels 20

Table 3-3. Instruments used in the Calorimeter System 2

1

Table 3-4. Expanded relative uncertainty of the measured heat release rate 22

Table 3-5. The expected Q, the measured mean Q, and the experimental duration 22

Table 4—1 . Early (<60 s) mean difference: bare-bead minus aspirated at same location 26

Table 4-2. Late (>60 s) mean difference: bare-bead minus aspirated at same location 26

Table 4—3. Early (< 60 s) lower and upper expanded uncertainty limits 27

Table 4-4. Late (> 60 s) lower and upper expanded uncertainty limits 28

Table 4—5. West and East tree layer heights 37

Table 5-1 . Heat flux gauge location 49

Table 5-2. Estimated certainties in heat flux measurements 54

Table 6-1. Uncertainty components for gas analyzers 69

Table 6-2. Mean value and expanded uncertainty for gas and smoke volume fractions at the S,^ sampling location 70

Table 6-3. Major contributors to the uncertainty in the smoke yield 73

Table 6-4. Mean value and expanded uncertainty for the smoke and gas yields (on a dry basis)

at the Sm sampling location 73

Table 6-5. Mean value and expanded uncertainty for the smoke and

gas yields (on a dry basis) for free bums 81

Table 7-1 . Bidirectional probe measurement locations 87

NISTNCSTAR 1-5B. WTC Investigation xxiii
List of Tables

Table 7-2. Velocity measurement uncertainty 89

Table 8-1. Thermophysical properties of Marinite, Chromel, and Alumel 99

Table 9-1. Statistical summary of SFRM thickness for steel components in Tests 5 and 6 123

Table 9-2. SFRM thickness for bars in Tests 5 and 6 .124

Table 9-3. SFRM thickness for coluirms in Tests 5 and 6 125

Table 9-4. SFRM thickness for truss A in Test 5 125

Table 9-5. SFRM thickness for truss B in Test 5 126

Table 9-6. SFRM thickness for truss A in Test 6 126

Table 9-7. SFRM thickness for truss B in Test 6 127

Table 9-8. SFRM thickness measurements below allowable in Test 5 127

Table 9-9. SFRM thickness measurements below allowable in Test 6 128

Table 9-10. Themiophysical properties of steel, Chromel, and Alumel 133

Table 9-1 1 . Summary of peak temperatures recorded during Tests 1, 2, and 3 137

Table 9-12. Summary of peak temperatures recorded during Tests 5 and 6 138

Table 9-13. Thermal lag in insulated steel components during Tests 5 and 6 145

Table 11-1. Percentage difference between peak values of the measured and the simulated steel surface temperatures at the hottest and coolest locations on various uninsulated

components in Test 1 211

Table 11-2. Percentage difference between peak values of the measured and the simulated steel temperatures for the hottest and coolest locations on the steel surface of various

components with SFRM in Test 5 211

Table 1 1-3. Percentage difference in the peak value of temperatures between the baseline

simulations and those with a 10 percent higher heat release rate at the hottest and

coolest locations on the steel surface of various uninsulated components in Test 1 221

XXIV NISTNCSTAR 1-5B, WTC Investigation
List of Acronyms and Abbreviations

Acronyms

ASTM ASTM International

BFRL Building and Fire Research Laboratory

BPS Building Performance Study

CFD computational fluid dynamics

COV coefficient of variation

FDS Fire Dynamics Simulator fire model

FEMA Federal Emergency Management Agency

FSI Fire Structure Interface model

LES Large Eddy Simulation

NIST National Institute of Standards and Technology

T) A XTVXT T rAJN Yin J Port Authority of New York and New Jersey

P.L. Public Law

SEAoNY Structural Engineers Association of New York

SFRM sprayed fire-resistive material

use United States Code

WTC World Trade Center

Abbreviations

degrees Celsius

op degrees Fahrenheit

micrometer

CO2 carbon dioxide

CO carbon monoxide

cm centimeter

ft foot

ft^ cubic feet

hp horsepower

Hz hertz

NISTNCSTAR 1-5B. WTC Investigation XXV
List of Acronyms and Abbreviations

in. inch

J joule

kW kilowatt

K Kelvin

kg , kilogram

kPa kilopascal

lb pound

L liter m meter

-m^ square meters

irun millimeter

MW megawatt

min minute

O2 oxygen

Pa pascal

s second

TC thermocouples

XXVI NISTNCSTAR 1-5B, WTC Investigation
Metric Conversion Table

To convert from to Multiply by

AREA AND SECOxND MOMENT OF AREA

square foot (fr)

square inch (in.")

square inch (in.")

square yard (yd")

square meter (m")

square meter (m")

square centimeter (cm"

square meter (m")

9.290 304 E-02

6.451 6 E-04

6.4516 E+00

8.361 274 E-01

ENERGY (includes WORK)

kilovvan hour (kW • h)

quad (1015 BtuIT)

therm (U.S.)

ton of TNT (energy- equivalent)

watt hour (W • h)

watt second (W • s)

joule (J)

joule (J)

joule (J)

joule (J)

joule (J)

joule (J)

3.6 E+06

1.055 056E+18

1.054 804 E+08

4.184 E+09

3.6 E+03

1.0 E+00

FORCE

dyne (dyn)

kilogram-force (kgf)

kilopond (kilogram-force) (kp)

kip (1 kip=l,000 Ibf) kipd kip=1.0001bf)

pound-force (ibf)

newton (N)

newlon (N)

newton (N)

newton (N)

kilonewton (kN)

newion (N)

l.OE-05

9.806 65 E+00

9.806 65 E+00

4.448 222 E+03

4.448 222 E+00

4.448 222 E+00

FORCE DIV IDED BY LENGTH

pound-force per foot (Ibf/ft)

pound-force per inch (Ibf/in.)

newton per meter (N/m)

newton per meter (N/m)

1.459 390 E+01

1.751 268 E+02

HEAT FLOW R.ATE

calorieth per minute (caith/min)

calorieth per second (calth/s)

kilocalorieth per minute (kcaith/min)

kilocalorieth per second (kcalth/s)

watt (W)

watt (W)

watt (W)

watt (W)

6.973 333 E-02

4.184 E+00

6.973 333 E+01

4.184 E+03

NISTNCSTAR 1-5B. WTC Investigation XXVll
Metric Conversion Table

To convert from to Multiply by

LENGTH

foot (ft)

inch (in)

inch (in.)

micron (m)

yard (yd)

meter (m)

meter (m)

centimeter (cm)

meter (m)

meter (m)

3.048 E-01

2.54 E-02

2.54 E+00

1 .0 E-06

9.144 E-01

MASS and MOMENT OF INERTIA

kilogram-force second squared

per meter (kgf • s"/m) kilogram (kg) 9.806 65 E+00

pound foot squared (lb ft') kilogram meter squared (kg • m") 4.214 011 E-02

pound inch squared (lb • \n.') kilogram meter squared (kg m") 2.926 397 E-04

ton, metric (t) kilogram (kg) 1 .0 E+03

ton, short (2,000 lb) kilogram (kg) 9.071 847E+02

MASS DIVIDED BY AREA

pound per square foot ( lb/ft"

)

kilogram per square meter (kg/m") 4.882 428 E+00

pound per square inch

{nol pound force) (lb/in.") kilogram per square meter (kg/m") 7.030 696 E+02

MASS DIVIDED BY LENGTH

pound per foot (lb/ft) kilogram per meter (kg/m) 1.488 164 E+00

pound per inch (lb/in.) kilogram per meter (kg/m) 1.785 797 E+01

pound per yard (lb/yd) kilogram per meter (kg/m) 4.960 546 E-01

PRESSURE or STRESS (FORCE DIVIDED BY AREA)

kilogram-force per square centimeter (kgf/cm') pascal (Pa) 9.806 65 E+04

kilogram-force per square meter (kgf/m") pascal (Pa) 9.806 65 E+00

kilogram-force per square millimeter (kgf/mm") pascal (Pa) 9.806 65 E+06

kip per square inch (ksi) (kip/in.') pascal (Pa) 6.894 757 E+06

kip per square inch (ksi) (kip/in.") kilopascal (kPa) 6.894 757 E+03

pound-force per square foot (Ibf/fr) pascal (Pa) 4.788 026 E+01

pound-force per square inch (psi) (Ibf/in.") pascal (Pa) 6.894 757 E+03

pound-force per square inch (psi) (Ibf/in.") kilopascal (kPa) 6.894 757 E+00

psi (pound-force per square inch) (Ibf/in.") pascal (Pa) 6.894 757 E+03

psi (pound-force per square inch) (Ibf/in.') kilopascal (kPa) 6.894 757 E+00

XXVlll NISTNCSTAR 1-5B, WTC Investigation
Metric Conversion Table

To convert from

TEMPERATURE

degree Celsius (°C)

degree centigrade

degree Fahrenheit (°F)

degree Fahrenheit (°F)

kelvin (K)

to

kelvin (K)

degree Celsius (°C)

degree Celsius (°C)

kelvin (K)

degree Celsius (°C)

Multiply by

T/K = t/°C + 273.15

t/°C a t /deg. cent.

t/°C = (t/°F- 32)/1.8

T/K = (t/°F + 459.67)/!.

t/°C = T/K 2 273.15

TEMPERATURE INTERV AL

degree Celsius (X) kelvin (K) 1 .0 E+00

degree centigrade degree Celsius (°C) 1.0 E+00

degree Fahrenheit (°F) degree Celsius (°C) 5.555 556 E-01

degree Fahrenheit (°F) kelvin (K) 5.555 556 E-01

degree Rankine (°R) kelvin (K) 5.555 556 E-01

VELOCITY (includes SPEED)

foot per second (ft/s) meter per second (m/s) 3.048 E-01

inch per second (in./s) meter per second (m/s) 2.54 E-02

kilometer per hour (km/h) meter per second (m/s) 2.777 778 E-01

mile per hour (mi h) kilometer per hour (km/h) 1.609 344 E+00

mile per minute (mi min) meter per second (m/s) 2.682 24 E+01

VOLUME (includes CAPACITY)

cubic foot (ft'') cubic meter (m^)

cubic inch (in.'' ) cubic meter (m')

cubic yard (yd^) cubic meter (m'')

gallon (U.S.) (gal) cubic meter (m')

gallon (U.S.) (gal) liter (L)

liter (L) cubic meter (m^)

ounce (U.S. fluid) (fl oz) cubic meter (m'')

ounce (U.S. fluid) (fl oz) milliliter (mL)

2.831 685 E-02

1.638 706 E-05

7.645 549 E-01

3.785 412 E-03

3.785 412 E+00

1.0 E-03

2.957 353 E-05

2.957 353 E+01

NISTNCSTAR 7-56, WTC Investigation XXIX
Metric Conversion Table

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XXX NISTNCSTAR 1-5B, WTC Investigation
Preface

Genesis of This Investigation

Immediately following the terrorist attack on the World Trade Center (WTC) on September 11, 2001, the

Federal Emergency Management Agency (FEMA) and the American Society of Civil Engineers began

planning a building perfonnance study of the disaster. The week of October 7, as soon as the rescue and

search efforts ceased, the Building Performance Study Team went to the site and began its assessment. This was to be a brief effort, as the study team consisted of experts who largely volunteered their time

away from their other professional commitments. The Building Performance Study Team issued its report in May 2002, fulfilling its goal "to determine probable failure mechanisms and to identify areas of

future investigation that could lead to practical measures for improving the damage resistance of buildings

against such unforeseen events."

On August 21, 2002, with funding from the U.S. Congress through FEMA, the National Institute of

Standards and Technology (NIST) announced its building and fire safety investigation of the WTC

disaster. On October 1, 2002, the National Construction Safety Team Act (Pubhc Law 107-231), was

signed into law. The NIST WTC Investigation was conducted under the authority of the National

Construction Safety Team Act.

The goals of the investigation of the WTC disaster were:

• To investigate the building construction, the materials used, and the technical conditions that contributed to the outcome of the WTC disaster.

• To serve as the basis for: - Improvements in the way buildings are designed, constructed, maintained, and used; - Improved tools and guidance for industry and safety officials; - Recommended revisions to current codes, standards, and practices; and - Improved public safety.

The specific objectives were:

1 . Determine why and how WTC 1 and WTC 2 collapsed following the initial impacts of the

aircraft and why and how WTC 7 collapsed;

2. Determine why the injuries and fatalities were so high or low depending on location,

including all technical aspects of fire protection, occupant behavior, evacuation, and

emergency response;

3. Determine what procedures and practices were used in the design, construction, operation,

and maintenance of WTC 1, 2, and 7; and

4. Identify, as specifically as possible, areas in current building and fire codes, standards, and

practices that warrant revision.

NIST NCSTAR X-X. WTC Investigation xxvn
Preface

NIST is a nonregulatory agency of the U.S. Department of Commerce's Technology Administration. The

purpose of NIST investigations is to improve the safety and structural integrity of buildings in the United

States, and the focus is on fact finding. NIST investigative teams are authorized to assess building

perfonnance and emergency response and evacuation procedures in the wake of any building failure that has resulted in substantial loss of life or that posed significant potential of substantial loss of life. NIST

does not have the statutory authority to make findings of fault nor negligence by individuals or organizations. Further, no part of any report resulting from a NIST investigation into a building failure or from an investigation under the National Constmction Safety Team Act may be used in any suit or action

for damages arising out of any matter mentioned in such report (15 USC 281a, as amended by Public

Law 107-231).

Organization of the Investigation

The National Constmction Safety Team for this Investigation, appointed by the then NIST Director, Dr. Arden L. Bement, Jr., was led by Dr. S. Shyam Sunder. Dr. William L. Grosshandler served as Associate Lead Investigator, Mr. Stephen A. Cauffman served as Program Manager for Administration,

and Mr. Harold E. Nelson sei'ved on the team as a private sector expert. The Investigation included eight

interdependent projects whose leaders comprised the remainder of the team. A detailed description of

each of these eight projects is available at http://wtc.nist.gov. The purpose of each project is summarized

in Table P-1, and the key interdependencies among the projects are illustrated in Fig. P-1.

Table P-1. Federal building and fire safety investigation of the WTC disaster.

Technical Area and Project Leader Project Purpose

Analysis of Building and Fire Codes and

Practices; Project Leaders: Dr. H. S. Lew

and Mr. Richard W. Bukowski

Document and analyze the code provisions, procedures, and

practices used in the design, construction, operation, and

maintenance of the stnictural, passive fire protection, and

emergency access and evacuation systems of WTC 1, 2, and 7. Baseline Structural Performance and

Aircraft Impact Damage Analysis; Project

Leader: Dr. Fahim H. Sadek

Analyze the baseline perfonnance of WTC 1 and WTC 2 under

design, service, and abnormal loads, and aircraft impact damage on

the structural, fire protection, and egress systems.

Mechanical and Metallurgical Analysis of

Structural Steel; Project Leader: Dr. Frank

W. Gayle

Detennine and analyze the mechanical and metallurgical properties

and quality of steel, weldments, and connections from steel recovered from WTC I, 2, and 7. Investigation of Active Fire Protection

Systems; Project Leader: Dr. David

D. Evans; Dr. William Grosshandler

Investigate the performance of the active fire protection systems in WTC 1, 2, and 7 and their role in fire control, emergency response, and fate of occupants and responders.

Reconstruction of Thennal and Tenability

Environment; Project Leader: Dr. Richard

G. Gann

Reconstruct the time-evolving temperature, thermal environment,

and smoke movement in WTC 1, 2, and 7 for use in evaluating the

structural performance of the buildings and behavior and fate of

occupants and responders.

Structural Fire Response and Collapse

Analysis; Project Leaders: Dr. John

L. Gross and Dr. Therese P. McAllister

Analyze the response of the WTC towers to fires with and without

aircraft damage, the response of WTC 7 in fires, the performance

of composite steel-trussed floor systems, and detennine the most

probable structural collapse sequence for WTC I, 2. and 7. Occupant Behavior, Egress, and Emergency

Communications; Project Leader: Mr. Jason D. Averill

Analyze the behavior and fate of occupants and responders, both

those who survived and those who did not, and the performance of

the evacuation system.

Emergency Response Technologies and

Guidelines; Project Leader: Mr. J. Randall

Lawson

Document the activities of the emergency responders from the time

of the terrorist attacks on WTC 1 and WTC 2 until the collapse of

WTC 7, including practices followed and technologies used.

xxviii NIST NCSTARX-X, WTC Investigation
Preface

NIST WTC Investigation Projects

Figure P-1. The eight projects in the federal building and fire safety

investigation of the WTC disaster.

National Construction Safety Team Advisory Committee

The NIST Director also established an advisory committee as mandated under the National Construction

Safety Team Act. The initial members of the committee were appointed following a public solicitation. These were:

Paul Fitzgerald, Executive Vice President (retired) FM Global, National Construction Safety

Team Advisory Committee Chair

John Barsom, President, Barsom Consulting, Ltd.

John Bryan, Professor Emeritus, University of Maryland

David Collins, President, The Preview Group, Inc.

Glenn Corbett, Professor, John Jay College of Criminal Justice

Philip DiNenno, President, Hughes Associates, Inc.

NIST NCSTAR X-X. WTC Investigation XXIX
Preface

• Robert Hanson, Professor Emeritus, University of Michigan

• Charles Thornton, Co-Chairman and Managing Principal, The Thomton-Tomasetti Group,

Inc.

• Kathleen Tiemey, Director, Natural Hazards Research and Applications Information Center,

University of Colorado at Boulder

• Forman Williams, Director, Center for Energy Research, University of California at San

Diego

This National Construction Safety Team Advisory Cominittee provided technical advice during the

Investigation and commentary on drafts of the Investigation reports prior to their public release. NIST

has benefited from the work of many people in the preparation of these reports, including the National

Construction Safety Team Advisory Committee. The content of the reports and recominendations,

however, are solely the responsibility of NIST. ^

Public Outreach

During the course of this Investigation, NIST held public briefings and meetings (listed in Table P-2) to

solicit input from the public, present preliminary findings, and obtain comments on the direction and

progress of the Investigation from the public and the Advisory Committee.

NIST maintained a publicly accessible Web site during this Investigation at http://wtc.nist.gov. The site contained extensive information on the background and progress of the Investigation.

NIST's WTC Public-Private Response Plan

The collapse of the WTC buildings has led to broad reexamination of how tall buildings are designed,

constructed, maintained, and used, especially with regard to major events such as fires, natural disasters,

and terrorist attacks. Reflecting the enhanced interest in effecting necessary change, NIST, with support

from Congress and the Administration, has put in place a program, the goal of which is to develop and

implement the standards, technology, and practices needed for cost-effective improvements to the safety

and security of buildings and building occupants, including evacuation, emergency response procedures,

and threat mitigation.

The strategy to meet this goal is a three-part NIST-led public-private response program that includes:

• A federal building and fire safety investigation to study the most probable factors that contributed to post-aircraft impact collapse of the WTC towers and the 47-story WTC 7

building, and the associated evacuation and emergency response experience.

• A research and development (R&D) program to (a) facilitate the implementation of

recommendations resulting from the WTC Investigation, and (b) provide the technical basis

for cost-effective improvements to national building and fire codes, standards, and practices

that enhance the safety of buildings, their occupants, and emergency responders.

XXX NIST NCSTAR X-X, WTC Investigation
Preface

Table P-2. Public meetings and briefings of the WTC Investigation.

Date Location Principal Agenda

June 24. 2002 New York City. NY Public meeting: Public comments on the Draft Plan for the pending WTC Investigation.

August 21. 2002 Gaithersburg, MD Media briefing announcing the formal start of the Investigation.

December 9. 2002 Washington, DC Media briefing on release of the Public Update and NIST request

for photographs and videos. April 8. 2003 New York City, NY Joint public forum with Columbia University on first-person interviews. Apnl 29-30. 2003 Gaithersburg. MD NCST Advisory Committee meeting on plan for and progress on WTC Investigation with a public comment session. May 7. 2003 New York City, NY Media briefing on release of May 2003 Progress Report.

August 26-27. 2003 Gaithersburg. MD NCST Advisory Committee meeting on status of the WTC

investigation with a public comment session. September 17,2003 New York City, NY Media and public briefing on initiation of first-person data

collection projects.

December 2-3. 2003 Gaithersburg. MD NCST Advisory Committee meeting on status and initial results and release of the Public Update with a public comment session. Februar\' 12. 2004 New York City, NY Public meeting on progress and preliminary findings with public

comments on issues to be considered in formulating final recommendations.

June 18.2004 New York City. NY Media/public briefing on release of June 2004 Progress Report. June 22-23, 2004 Gaithersburg, MD NCST Advisory Committee meeting on the status of and

preliminary findings from the WTC Investigation with a public

comment session. August 24. 2004 Northbrook, IL Public viewing of standard fire resistance test of WTC floor system at Underwriters Laboratories, Inc. October 19-20, 2004 Gaithersburg, MD NCST Advisory Committee meeting on status and near complete

set of preliminary findings with a public comment session. November "'004 Gailhpr^hiirfT N/ir) Nr^ST AHvisnrv (^ommittep Hi*;rii*s<sinn nn draft annual rpnort to Congress, a public comment session, and a closed session to discuss pre-draft recommendations for WTC Investigation.

April 5, 2005 New York City, NY Media and public briefing on release of the probable collapse

sequence for the WTC towers and draft reports for the projects on

codes and practices, evacuation, and emergency response. June 23, 2005 New York City, NY Media and public briefing on release of all draft reports for the WTC towers and draft recommendations for public comment.

September 12-13.

2005

Gaithersburg. MD NCST Advisory Committee meeting on disposition of public

comments and update to draft reports for the WTC towers. September 13-15.

2005

Gaithersburg, MD WTC Technical Conference for stakeholders and technical community for dissemination of findings and recommendations

and opportunity for public to make technical comments.

• A dissemination and technical assistance program (DTAP) to (a) engage leaders of the

construction and building community in ensuring timely adoption and widespread use of

proposed changes to practices, standards, and codes resulting from the WTC Investigation

and the R&D program, and (b) provide practical guidance and tools to better prepare facility owners, contractors, architects, engineers, emergency responders, and regulatory authorities

to respond to future disasters. The desired outcomes are to make buildings, occupants, and first responders safer in future disaster events.

NIST NCSTAR X-X, WTC Investigation XXXl
Preface

National Construction Safety Team Reports on the WTC Investigation

A final report on the collapse of the WTC towers is being issued as NIST NCSTAR 1 . A companion

report on the collapse of WTC 7 is being issued as NIST NCSTAR lA. The present report is one of a set that provides more detailed documentation of the Investigation findings and the means by which these

technical results were achieved. As such, it is part of the archival record of this Investigation. The titles of the full set of Investigation publications are:

NIST (National Institute of Standards and Technology). 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Final Report on the Collapse ofthe World Trade

Center Towers. NIST NCSTAR 1. Gaithersburg, MD, September.

NIST (National Institute of Standards and Technology). 2006. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Final Report on the Collapse of World Trade Center 7. NIST NCSTAR 1 A. Gaithersburg, MD.

Lew, H. S., R. W. Bukowski, and N. J. Carino. 2005. Federal Building and Fire Safety Investigation of

the World Trade Center Disaster: Design, Construction, and Maintenance ofStructural and Life Safety

Systems. NIST NCSTAR 1-1. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Fanella, D. A., A. T. Derecho, and S. K. Ghosh. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Design and Construction ofStructural Systems.

NIST NCSTAR 1-lA. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Ghosh, S. K., and X. Liang. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Comparison ofBuilding Code Structural Requirements. NIST

NCSTAR 1-lB. National Institute of Standards and Technology. Gaithersburg, MD, September.

Fanella, D. A., A. T. Derecho, and S. K. Ghosh. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Maintenance and Modifications to Structural

Systems. NIST NCSTAR 1-lC. National Institute of Standards and Technology. Gaithersburg,

MD, September.

Grill, R. A., and D. A. Johnson. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Fire Protection and Life Safety Provisions Applied to the Design and

Construction of World Trade Center I, 2, and 7 and Post-Construction Provisions Applied after Occupancy. NIST NCSTAR 1- ID. National Institute of Standards and Technology. Gaithersburg,

MD, September.

Razza, J. C, and R. A. Grill. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Comparison of Codes, Standards, and Practices in Use at the Time ofthe

Design and Construction of World Trade Center 1, 2, and 7. NIST NCSTAR 1-lE. National

Institute of Standards and Technology. Gaithersburg, MD, September.

Grill, R. A., D. A. Johnson, and D. A. Fanella. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Comparison of the 1968 and Current (2003) New

xxxii NIST NCSTAR X-X, WTC Investigation
Preface

York City Building Code Provisiom. NIST NCSTAR 1-lF. National Institute of Standards and

Technology. Gaithersburg, MD, September.

Grill, R. A., and D. A. Johnson. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Amendments to the Fire Protection and Life Safety Provisions of the New

York City Building Code by Local Laws Adopted While World Trade Center 1, 2, and 7 Were in Use. NIST NCSTAR 1-lG. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Grill, R. A., and D. A. Johnson. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Post-Construction Modifications to Fire Protection and Life Safety Systems

of World Trade Center 1 and 2. NIST NCSTAR 1-lH. National Institute of Standards and

Technology. Gaithersburg, MD, September.

Grill, R. A.. D. A. Johnson, and D. A. Fanella. 2005. Federal Building and Fire Safety Investigation

ofthe World Trade Center Disaster: Post-Construction Modifications to Fire Protection, Life

Safety, and Structural Systems of World Trade Center 7. NIST NCSTAR 1-1 1. National Institute of

Standards and Technology. Gaithersburg, MD, September.

Grill. R. A., and D. A. Johnson. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Design, Installation, and Operation ofFuel System for Emergency Power in World Trade Center 7. NIST NCSTAR 1-1 J. National Institute of Standards and Technology.

Gaithersburg, MD, September.

Sadek, F. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center Disaster:

Baseline Structural Performance andAircraft Impact Damage Analysis ofthe World Trade Center

Towers. NIST NCSTAR 1-2. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Faschan, W. J., and R. B. Garlock. 2005. Federal Building and Fire Safety Investigation ofthe

World Trade Center Disaster: Reference Structural Models and Baseline Performance Analysis of

the World Trade Center Towers. NIST NCSTAR 1-2A. National Institute of Standards and

Technology. Gaithersburg, MD, September.

Kirkpatrick, S. W., R. T. Bocchieri, F. Sadek, R. A. MacNeill, S. Holmes, B. D. Peterson,

R. W. Cilke, C. Navarro. 2005. Federal Building and Fire Safety Investigation ofthe World Trade

Center Disaster: Analysis ofAircraft Impacts into the World Trade Center Towers, NIST

NCSTAR 1-2B. National Institute of Standards and Technology. Gaithersburg, MD, September.

Gayle, F. W., R. J. Fields, W. E. Luecke, S. W. Banovic, T. Foecke, C. N. McCowan, T. A. Siewert, and

J. D. McColskey. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: Mechanical and Metallurgical Analysis ofStructural Steel. NIST NCSTAR 1-3. National

Institute of Standards and Technology. Gaithersburg, MD, September.

Luecke, W. E., T. A. Siewert, and F. W. Gayle. 2005. Federal Building andFire Safety

Investigation of the World Trade Center Disaster: Contemporaneous Structural Steel

Specifications. NIST Special Publication 1-3A. National Instimte of Standards and Technology.

Gaithersburg, MD, September.

NIST NCSTAR X-X. WTC Investigation xxxni
Preface

Banovic, S. W. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: Steel Inventoiy and Identification. NIST NCSTAR 1-3B. National Institute of Standards

and Technology. Gaithersburg, MD, September.

Banovic, S. W., and T. Foecke. 2005. Federal Building and Fire Safety Investigation of the World

Trade Center Disaster: Damage and Failure Modes ofStructural Steel Components. NIST

NCSTAR 1-3C. National Institute of Standards and Technology. Gaithersburg, MD, September.

Luecke, W. E., J. D. McColskey, C. N. McCowan, S. W. Banovic, R. J. Fields, T. Foecke,

T. A. Siewert, and F. W. Gayle. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Mechanical Properties ofStructural Steels. NIST NCSTAR 1-3D.

National Institute of Standards and Technology. Gaithersburg, MD, September.

Banovic, S. W., C. N. McCowan, and W. E. Luecke. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Physical Properties ofStructural Steels. NIST

NCSTAR 1-3E. National Institute of Standards and Technology. Gaithersburg, MD, September.

Evans, D. D., R. D. Peacock, E. D. Kuligowski, W. S. Dols, and W. L. Grosshandler. 2005. Federal

Building and Fire Safety Investigation ofthe World Trade Center Disaster: Active Fire Protection

Systems. NIST NCSTAR 1-4. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Kuligowski, E. D., D. D. Evans, and R. D. Peacock. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Post-Construction Fires Prior to September 11,

2001. NIST NCSTAR 1-4A. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Hopkins, M., J. Schoenrock, and E. Budnick. 2005. Federal Building and Fire Safety Investigation

ofthe World Trade Center Disaster: Fire Suppression Systems. NIST NCSTAR 1-4B. National

Institute of Standards and Technology. Gaithersburg, MD, September.

Keough, R. J., and R. A. Grill. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Fire Alarm Systems. NIST NCSTAR I-4C. National Institute of Standards

and Technology. Gaithersburg, MD, September.

Fen-eira, M. J., and S. M. Strege. 2005. Federal Building and Fire Safety Investigation ofthe

World Trade Center Disaster: Smoke Management Systems. NIST NCSTAR I-4D. National

Institute of Standards and Technology. Gaithersburg, MD, September.

Gann, R. G., A. Hamins, K. B. McGrattan, G. W. Mulholland, H. E. Nelson, T. J. Ohlemiller,

W. M. Pitts, and K. R. Prasad. 2005. Federal Building and Fire Safety Investigation ofthe World Trade

Center Disaster: Reconstruction ofthe Fires in the World Trade Center Towers. NIST NCSTAR 1-5. National Institute of Standards and Technology. Gaithersburg, MD, September.

Pitts, W. M., K. M. Butler, and V. Junker. 2005. Federal Building and Fire Safety Investigation of

the World Trade Center Disaster: Visual Evidence, Damage Estimates, and Timeline Analysis.

NIST NCSTAR 1-5 A. National Institute of Standards and Technology. Gaithersburg, MD,

September.

XXXIV NIST NCSTAR X-X, WTC Investigation
Preface

Hamins, A., A. Maranghides, K. B. McGrattan, E. Johnsson, T. J. Ohlemiller, M. Donnelly,

J. Yang. G. Mulholland, K. R. Prasad, S. Kukuck, R. Anleitner and T. McAllister. 2005. Federal

Building and Fire Safety Investigation ofthe World Trade Center Disaster: Experiments and

Modeling ofStructural Steel Elements Exposed to Fire. NIST NCSTAR 1-5B. National Institute of

Standards and Technology. Gaithersburg, MD, September.

Ohlemiller, T. J., G. W. Mulholland, A. Maranghides, J. J. Filliben, and R. G. Gann. 2005. Federal

Building and Fire Safety Investigation ofthe World Trade Center Disaster: Fire Tests ofSingle

Office Workstations. NIST NCSTAR 1-5C. National Institute of Standards and Technology.

Gaithersburg, MD, September.

Gann. R. G., M. A. Riley, J. M. Repp, A. S. Whittaker, A. M. Reinhom, and P. A. Hough. 2005.

Federal Building and Fire Safety Investigation ofthe World Trade Center Disaster: Reaction of

Ceiling Tile Systems to Shocks. NIST NCSTAR 1-5D. National Institute of Standards and

Technology. Gaithersburg, MD, September.

Hamins. A., A. Maranghides, K. B. McGrattan, T. J. Ohlemiller, and R. Anleitner. 2005. Federal

Building and Fire Safety Investigation ofthe World Trade Center Disaster: Experiments and

Modeling ofMultiple Workstations Burning in a Compartment. NIST NCSTAR 1-5E. National

Institute of Standards and Technology. Gaithersburg, MD, September.

McGrattan, K. B., C. Bouldin, and G. Forney. 2005. Federal Building and Fire Safety

Investigation ofthe World Trade Center Disaster: Computer Simulation ofthe Fires in the World

Trade Center Towers. NIST NCSTAR 1-5F. National Institute of Standards and Technology.

Gaithersburg, MD, September.

Prasad, K. R., and H. R. Baum. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Fire Structure Interface and Thermal Response ofthe World Trade Center

Towers. NIST NCSTAR 1-5G. National Institute of Standards and Technology. Gaithersburg,

MD, September.

Gross, J. L., and T. McAllister. 2005. Federal Building and Fire Safety Investigation ofthe World Trade

Center Disaster: Structural Fire Response and Probable Collapse Sequence ofthe World Trade Center

Towers. NIST NCSTAR 1-6. National Institute of Standards and Technology. Gaithersburg, MD,

September.

Carino, N. J., M. A. Stames, J. L. Gross, J. C. Yang, S. Kukuck, K. R. Prasad, and R. W. Bukowski.

2005. Federal Building and Fire Safety Investigation ofthe World Trade Center Disaster: Passive

Fire Protection. NIST NCSTAR 1-6A. National Instimte of Standards and Technology.

Gaithersburg, MD, September.

Gross, J., F. Hervey, M. Izydorek, J. Mammoser, and J. Treadway. 2005. Federal Building and

Fire Safety Investigation ofthe World Trade Center Disaster: Fire Resistance Tests ofFloor Truss

Systems. NIST NCSTAR 1-6B. National Institute of Standards and Technology. Gaithersburg,

MD, September.

Zarghamee, M. S., S. Bolourchi, D. W. Eggers, O. O. Erbay, F. W. Kan, Y. Kitane, A. A. Liepins,

M. Mudlock, W. I. Naguib, R. P. Ojdrovic, A. T. Sarawit, P. R Barrett, J. L. Gross, and

NIST NCSTAR X-X, WTC Investigation XXXV
Preface

T. P. McAllister. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: Component, Connection, and Subsystem Structural Analysis. NISTNCSTAR 1-6C.

National Institute of Standards and Technology. Gaithersburg, MD, September.

Zarghamee, M. S., Y. Kitane, O. O. Erbay, T. P. McAllister, and J. L. Gross. 2005. Federal

Building and Fire Safety Investigation ofthe World Trade Center Disaster: Global Structural

Analysis ofthe Response ofthe World Trade Center Towers to Impact Damage and Fire. NIST

NCSTAR 1 -6D. National Institute of Standards and Technology. Gaithersburg, MD, September.

McAllister, T., R. W. Bukowski, R. G. Gann, J. L. Gross, K. B. McGrattan, H. E. Nelson, L. Phan,

W. M. Pitts, K. R. Prasad, F. Sadek. 2006. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: Structural Fire Response and Probable Collapse Sequence of World Trade

Center?. (Provisional). NIST NCSTAR 1-6E. National Institute of Standards and Technology.

Gaithersburg, MD.

Gilsanz, R., V. Arbitrio, C. Anders, D. Chlebus, K. Ezzeldin, W. Guo, P. Moloney, A. Montalva,

J. Oh, K. Rubenacker. 2006. Federal Building and Fire Safet}> Investigation ofthe World Trade

Center Disaster: Structural Analysis ofthe Response of World Trade Center 7 to Debris Damage

and Fire. (Provisional). NIST NCSTAR 1-6F. National Institute of Standards and Technology.

Gaithersburg, MD.

Kim, W. 2006. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: Analysis ofSeptember 11, 2001, Seismogram Data. (Provisional). NIST NCSTAR 1-6G.

National Institute of Standards and Technology. Gaithersburg, MD.

Nelson, K. 2006. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: The Con Ed Substation in World Trade Center 7. (Provisional). NISTNCSTAR 1-6H.

National Institute of Standards and Technology. Gaithersburg, MD.

Averill, J. D., D. S. Mileti, R. D. Peacock, E. D. Kuligowski, N. Groner, G. Proulx, P. A. Reneke, and

H. E. Nelson. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center Disaster:

Occupant Behavior, Egress, and Emergency Communication. NIST NCSTAR 1-7. National Institute of

Standards and Technology. Gaithersburg, MD, September.

Fahy, R., and G. Proulx. 2005. Federal Building and Fire Safety Investigation ofthe World Trade

Center Disaster: Analysis ofPublished Accounts ofthe World Trade Center Evacuation. NIST

NCSTAR 1-7A. National Institute of Standards and Technology. Gaithersburg, MD, September.

Zmud, J. 2005. Federal Building and Fire Safety Investigation ofthe World Trade Center

Disaster: Technical Documentation for Sun'ey Administration. NISTNCSTAR 1-7B. National

Institute of Standards and Technology. Gaithersburg, MD, September.

Lawson, J. R., and R. L. Vettori. 2005. Federal Building and Fire Safety Investigation ofthe World

Trade Center Disaster: The Emergency Response Operations. NIST NCSTAR 1-8. National Institute of

Standards and Technology. Gaithersburg, MD, September.

xxxvi NIST NCSTAR X-X. WTC Investigation
Acknowledgments

The authors are grateful to National Institute of Standards and Technology staff including Lauren

DeLauter, Edward Hnetkovsky, Jack Lee, Roy McLane, Jay McElroy, and Mike Selepak for their

assistance conducting the experiments, to Tyler Smith for assistance with data analysis, and to Dr. Rodney Br>'ant for helpful discussions.

NISTNCSTAR 7-56, WTC Investigation xli
Acknowledgments

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xlii NISTNCSTAR 1-5B, WTC Investigation
Executive Summary

This report represents part of the reconstruction of the thermal and tenabihty environment of the World

Trade Center (WTC) Investigation, the goal of which was to evaluate the structural performance of the

buildings. As part of this endeavor, three computational models were used to reconstruct the timeevoh ing temperature, thermal environment, and smoke movement in WTC 1, 2, and 7 and to assess the

effect of the fire on the temperature rise of the insulated structural steel components.

The models applied to the analysis of the WTC fires were ( 1 ) the National Institute of Standards and

Technology (NIST) Fire Dynamic Simulator (FDS), which is a computational model of fire behavior and

its effects; (2) the NIST Fire-Structure Interface (FSI), which applies the fire model (FDS) results to determine boundary conditions for computational finite-element models of the temperature behavior of

the insulated structural components; and (3) the commercially available ANSYS finite-element thermal

model. The FDS is the first large-domain computational fluid dynamics (CFD) fire model that predicts

and visualizes the spread, growth and suppression of a fire based on the underlying scientific principles

governing fluid motion. The FDS code numerically solves the modeled conservation equations of mass, momentum and energy that govern low-speed, thermally driven flows with an emphasis on smoke and

heat transport from fires. The companion software package, called Smokeview, graphically presents the

results of the FDS three-dimensional time-dependent simulation. Smokeview animates in three

dimensions the smoke particulates, heat fluxes, temperatures and fluid velocities within a building. Users

of the FDS/Smokeview software package can view the simulated results from any angle and from inside or outside the computational boundaries. The FSI calculates the transfer of radiant and convective heat

from a CFD fire model such as FDS to a coupled, transient, three-dimensional finite-element model for the thermal response of structural components such as ANSYS. The structural components may be

simple (e.g., uninsulated steel) or complex (e.g., steel insulated with a fire protective coating).

For application to the WTC Investigation, each of these models needs experimental data (a) to guide

adaptation development of the models for this specific purpose and (b) with which to ascertain the

accuracy of the model predictions. Ideally, the uncertainty in the agreement with experiments would be

much smaller than the effect of varying the unknowns in the actual conditions on September 11, 2001.

To gain confidence in the accuracy of these models as applied to the WTC investigation, two large-scale

test series were conducted to assess the accuracy of the numerical model. A number of small-scale

experiments were conducted to provide input data on the different materials used in the experiment. The

two large-scale test series involved fires in compartments with approximately the same height as a floor in the WTC towers. The results of the first series of tests are described in this report. In these experiments, a

liquid fuel spray burner was used to generate a fixed amount of energy in a compartment that was fitted with various targets and obstructions including steel objects such as columns, trusses and rods, and the

fire conditions were over-ventilated, that is, the rate of burning was not limited by the availability of

oxygen. In the experiments described in this report, the fuels used were well-characterized hydrocarbons

or mixture of hydrocarbons. This facilitated experimental control of the fire heat release. The heat release

rate drives fire effects including radiative and convective heat transfer within the compartment and the

temperature rise of the gases and the contents of the compartment and its surfaces.

NISTNCSTAR 1-5B, WTC Investigation xliii
Executive Summary

A second series of fire tests (NIST NCSTAR 1-5E') was conducted in a larger compartment in which

muhiple office workstations similar to those present in the WTC were burned under conditions similar to those experienced on September 1 1, 2001 . The second series of experiments investigated the accuracy of

the fire model for under-ventilated conditions. In those experiments, actual furniture was used as the fuel. The experiments were designed to recreate very different aspects of the fire model for application to the

WTC fires including issues associated with limited ventilation, fire spread and growth on real furnishings,

and the effects of debris and jet fuel on the heat release rate of the fires.

Both series of experiments were conducted at the NIST Large Fire Laboratory. An objective of the first series of experiments, which is reported here, was to assess the accuracy with which the NIST FDS

predicted the thermal environment in a burning compartment and to establish a data set to validate the

prediction of the temperature rise of structural steel components. Within a steel-frame compartment

(nominally 3 m by 7 m by 4 m high) lined with calcium silicate boards were placed several steel components including two trusses, one thin-walled hollow column, and simple rods. The components

were tested in both an uninsulated (unprotected bare) state and with an insulating coating of fireproofing

or fibrous sprayed fire-resistive material (SFRM), which was applied by the manufacturer. Two

thicknesses of SFRM were tested. A significant variation in the thickness of the SFRM applied to the

steel components was measured, with the 1.91 cm specified thickness averaging 2.2 cm to 2.7 cm with a coefficient of variation (COV; equal to the ratio of the standard deviation to the mean value of the

thickness) from 0.17 to 0.27 and the 3.81 cm specified thickness averaging 4.05 cm to 4.10 cm with a COV of 0.07 to 0.20. The lower COV values were measured for the columns; the higher COV values

occurred for the bars and angles, which are more difficult shapes for uniform application of the SFRM

product. The SFRM measurements may be considered as a lower bound estimate of the variance that would be expected for field applications.

The fires were generated using liquid hydrocarbon fuels introduced by a two-nozzle spray burner onto a

1 m by 2 m pan. The fire size tested was between 1 .9 MW and 3.4 MW to assure that the structural components were immersed in flames and hot gases. The fuels used were a commercial blend of heptane

isomers and a mixture of the heptane blend with toluene. These fuels were selected to cover the expected

range of soot loading in fires generated by typical office furnishings. Six fire experiments were

conducted.

Nearly 350 instruments were used to measure a number of important variables, including the heat release

rate of the fire, the radiative and total heat flux at various targets in the compartment at locations near the

floor and the ceiling, the temperatures of insulated and uninsulated steel components, the concentrations

of oxygen, carbon monoxide and carbon dioxide, and the vertical profiles of gas-phase temperature. The

overall fire behavior was documented using four video cameras viewing fire behavior inside the

compartment from different orientations.

Following the establishment of baseline signals from all the measurement devices, the burner was ignited

and continued burning at a steady rate. The fire was allowed to bum until the temperature at any steel surface approached approximately 600 °C, which is a temperature at which steel mechanical properties

begin to degrade significantly. Data were acquired for several minutes after the fuel flow was stopped to observe cooling of the compartment and its contents.

This reference is to one of the companion documents from this Investigation. A hst of these documents appears in the Preface

to this report.

xliv NIST NCSTAR 1-5B, WTC Investigation
Executive Summary

The measurements led to findings regarding the thermal behavior of the compartment and its contents,

including the steel components, as they were exposed to fires of various size, fuel sootiness, and

proximit\' to the fire plume. The effect of the SFRM on the rate of temperature rise of the steel components was also measured. Noteworthy results included the following:

• Ceiling surface temperatures obtained values as high as 700 °C to 800 °C and 800 °C to 900 °C after 15 min and 50 inin, respectively, of exposure to a 3 MW heat release rate fire with peak temperatures observ'ed to occur above the fuel pan.

• The value of the heat flux about the compartment was dependent on location and test conditions. For example, at the floor in front of the fire pan, the heat flux varied from

20 kW/m' to 90 kW/m'. In the upper layer near the ceiling, the heat flux was measured to be

as high as 90 kW/m".

• The experimental results showed that the SFRM significantly delayed the rise in the steel temperature. The highest temperatures obtained at the steel surface were significantly lower

than the temperature at the outside face of the insulation material.

• The surface temperature of uninsulated steel trusses that were subject to impingement or proximity of flames from fires with a nominal 3 MW heat release rate reached 600 °C within

15 min. In comparison, the insulated steel trusses with 1.91 cm thick SFRM protection

reached 100 °C to 200 °C and the truss with 3.81 cm SFRM reached 50 °C to 100 °C after 15

min of exposure to a 3 MW heat release rate fire. The steel surface of an insulated steel truss with 1.91 cm (nominal) and 3.81 cm (nominal) SFRM obtained a maximum temperature of

680 °C and 415 °C, respectively, after 50 min of exposure to a 3 MW fire.

• For the insulated steel components, the steel surface temperatures obtained a maximum

sometime after the fire was stopped due to lag of the thermal wave through the SFRM. A

thermal lag of between 100 s and 800 s for the nominal 1.91 cm SFRM thickness and 1,000 s to over 1.900 s for the nominal 3.81 cm SFRM thickness was observed, depending on the

values of the temperature and the temperature gradients within the SFRM and the steel.

FDS, FSI, and ANSYS were used to simulate the thermal environment in the compartment and the time

varying steel temperature distribution. The numerical accuracy of the fire dynamic predictions was

determined through quantitative comparison with the data. Similar analysis was performed to validate the

finite-element modeling of the thermal behavior of the bare and insulated steel components. A

comparison of predictions with measurements showed that:

• FDS predictions of the upper layer gas temperature increase agreed with measurements to within 10 percent, which was fairly accurate considering model sensitivity to input

parameters such as the heat release rate. Since the heat flux to the walls and objects within

the upper layer is highly dependent on the upper layer temperature, these predictions were

also equally accurate.

• The sootier-burning fuel led to similar temperature rise in the ceiling and the steel above the

fire plume, while the measured and predicted flux to the lower layer was smaller.

NISTNCSTAR 1-5B. WTC Investigation xlv
Executive Summary

• FDS predicted asymmetry of the fire plume caused by obstructions to uniform flow through

the compartment, but did not accurately predict the extent of the asymmetry, which adversely

impacted the FSI prediction of structural components very close to the fire. Away from the

fire, small differences in the exact flame position were not as important.

• The predictions of the steel surface temperature increases for the uninsulated and the SFRM

insulated steel components using FSI in combination with ANSYS agreed with measurements

to within 8 percent and 20 percent on average, respectively. Quantitative disagreements

between the models and the measurements were attributed to model sensitivity to specific

input parameters that had experimental uncertainty. For the bare steel, a series of calculations

showed that the differences between the predictions and the measurements could be attributed

primarily to sensitivity of the fire model to the fire heat release rate. For the insulated steel, the differences between the predictions and the measurements depended mainly on the

uncertainty of the SFRM coating thickness.

In summary, measurements in a series of specially designed large-scale fire experiments were compared

with predictions from the FDS/FSI/ANSYS methodology to quantify the accuracy of the calculation

results and to detennine which parameters the models were most sensitive. The models provided a reasonable characterization of the thennal enviromnent for compartments experiencing a fire as well as the associated temperature rise in steel components, providing the level of confidence necessary to apply

the methodology to the investigation of the WTC disaster.

xlvi NISTNCSTAR 1-5B, WTC Investigation
Chapter 1 Introduction

1.1 MODELING AS AN INVESTIGATIVE TOOL

Reconstructing the fires and their impact on structural components in the World Trade Center (WTC)

buildings on September 1 1, 2001, required extensive use of computational models. WTC buildings 1, 2, and 7 each sur\ ived for a different amount of time from the start of the disaster. Each time to failure resulted from complex interactions among the features of the building construction, the initial damage to the buildings, and the ensuing fires. Accurate computational modeling was critical to the Investigation

because of this complexity and because so httle direct evidence was available. While the Investigation

team, led by the National Institute of Standards and Technology, compiled and catalogued an extensive

array of photographic and videographic records, these were limited to perspectives external to the

building. They provided little information on the physical damage and the progress of the fires inside the

buildings, other than near the immediate vicinity of the windows and other openings. Analysis of the

recovered WTC steel provides some indication of the maximum temperatures experienced. However, this analysis cannot identify the times at which peak temperatures were reached or the time-varying history of

the thermal environment throughout the buildings. Information on the timeline of the thermal

environment in the buildings, combined with characterization of the building damage from the incident

aircraft (WTC 1 and WTC 2) or from debris from the collapsing towers (WTC 7), was central to assessing

the plausibility of various possible explanations for the collapse of the three buildings.

Computer simulation of the fires included consideration of such factors as:

• The initial building damage (exterior and interior);

• The nature, quantity and distribution of the combustibles and debris; and

• Changes in the ventilation as the fires progressed.

Additional modeling of the transfer of heat from the fire to the structural components complemented these

modeling considerations. Combined, the simulations then produced a complete thennal history, filling in the gaps left by the paucity of information. Simulations that captured the essence of the photographic

evidence and the occupant accounts were then presumed to be the more probable replications of the actual events.

For models to be a viable investigative tool, it is essential to know the accuracy with which they captured

the physical phenomena of the fires and the concurrent heat transfer to the building structure. This report

documents one phase of the effort to ascertain the validity of the models for this Investigation.

Simulating the effects of a fire on the structural integrity of a building requires a coupling between the gas-phase enthalpy released by the fire and the transfer of that enthalpy to the solid-phase load bearing

1.2 OVERVIEW OF MODELS

NISTNCSTAR 1-5B, WTC Investigation 1
Chapter 1

structure, which is vulnerable to thermal effects. The methodology involves the implementation of the

two pieces of software that are described below. The comiection between the two is established primarily

through the interaction of radiative and convective heat transfer between the solid- and gas-phases, with

the conduction of heat tlirough the structural components. This process is made difficult in large,

geometrically complex buildings by the wide disparity in length and time scales that must be accounted

for in the simulations. A methodology for achieving this coupling was developed and applied to the

investigation of the WTC collapse (Prasad and Baum 2005).

The building fires are modeled using the National Institute of Standards and Technology (NIST) Fire

Dynamics Simulator/Smokeview (McGrattan 2004) software package. The Fire Dynamics Simulator

(FDS) is one of the first large-domain computational fiuid dynamics (CFD) fire codes that predicts and

visualizes the spread, growth and suppression of a fire based on the underlying scientific principles

governing fluid motion. Smokeview is an OpenGL graphics-based software application for visualizing

information from large calculations. Since the purpose of computing is insight, not numbers, Smokeview

facilitates an understanding of calculation output through visualization of the results. Smokeview

animates FDS results in a variety of user selected modes, including the trajectory of smoke and slices of

the temperature and velocity fields within the three-dimensional domain.

The code numerically solves the modeled conservation equations of mass, momentum, and energy that govern low-speed, thermally driven flows, with an emphasis on smoke and heat transport from fires. The formulation of the equations and the numerical algorithm are contained in a companion document,

called Fire Dynamics Simulator (Version 4) - Technical Reference Guide (McGrattan 2004). The

models and their documentation are available to the public free of charge on the World Wide Web

(http://www.fire.nist.gov/fds/). This site gives links to a setup program for PC installation. It also

contains documentation for Smokeview and FDS, sample calculations, software updates, and links for requesting feedback about the software.

The core algorithm is an explicit predictor-coiTCCtor scheme, second order accurate in space and time.

Turbulence is treated by means of the Smagorinsky fonn of Large Eddy Simulation (LES). For most

applications, FDS uses a mixture fraction combustion model. Radiative heat transfer is included in the

model via the solution of the radiation transport equation for a non-scattering gray gas, and in some

limited cases using a wide band model. The equation is solved using a technique similar to finite volume

methods for convective transport and uses approximately 100 discrete angles. All solid surfaces are assigned thermal boundary conditions, plus information about the burning behavior of the material. More

detailed information regarding the FDS model is discussed in Chapter 10 of this report and in the

technical reference guide (McGrattan 2004) and references therein.

The companion software package, called Smokeview, graphically presents the results of the FDS threedimensional time-dependent simulation. Smokeview is a software tool designed to visualize numerical

predictions generated by FDS. FDS and Smokeview are used in concert to respectively model and

visualize fire phenomena. Smokeview visualizes both dynamic and static data. Dynamic data are visualized by animating particle flow (showing location and "values" of tracer particles), 2D contour

slices (both within the domain and on solid surfaces), and 3D level surfaces. 2D contour sHces can also

be drawn with colored vectors that use velocity data to show flow direction, speed, and value. Static data

are visualized similarly by drawing 2D contours, vector plots, and 3D level surfaces anywhere within a simulation scene at a fixed time.

2 NISTNCSTAR 1-5B, WTC Investigation
Introduction

Normally, Smokeview is used in a post-processing step to visualize FDS data after a calculation has been

completed. Smokeview may also be used during a calculation to monitor a simulation's progress and

before a calculation to visualize the calculation set-up, which may include walls, vents, sprinklers, smoke

detectors, etc. In order to set up correct FDS input files more quickly, Smokeview is typically used to edit or create blockages by specifying the size, location, and/or material properties. More detailed information

regarding Smokeview can be found in the Smokeview User's Guide (Forney and McGrattan 2004).

The Fire Structure Interface (FSI) (NIST NCSTAR 1-5G') involves calculating heat transfer to sub-grid

scale structural components through consideration of both convection and radiation. Radiative heat

transfer, which dominate the heat transfer process, is calculated using a simple radiative transport model

that assumes that the compartment is locally divided into a hot, soot laden upper layer and a cool

relatively clear lower layer. The properties of the two layers are extracted from temporal averages of the

results obtained from the FDS (NIST NCSTAR 1-5F). Explicit formulae for the heat flux are obtained as a fijnction of temperature, hot layer depth, soot concentration and orientation of each structural element.

These formulae are used to generate realistic thermal boundary conditions for a coupled transient three- dimensional finite-element code. Convective heat transfer is calculated based on simplifying assumptions

and heat transfer correlations available in the literature. The code is used to generate solutions for the

heating of complex structural assemblies.

The thermal response of the steel structure and any associated fire protective coating is modeled using the

ANSYS (2003) finite-element software. ANSYS uses iterative, direct, and eigenvalue matrix solvers to analyze the thermal field in complex structural geometries. The output of the FDS calculation is used by

the FSI to develop thermal boundary conditions for the ANSYS calculation. The FSI calculates the

transfer of radiant and convective heat from the FDS fire model, which acts as the input to ANSYS for

calculation of the thermal response of structural components. The structural components may be bare

steel or insulation-coated steel. The FSI results are applied in a time dependent fashion at each location on a structural component.

In summary, the objectives of this study are to (1) assess the accuracy of FDS predictions of the thermal

environment in a burning compartment through comparison with a series of measurements, and (2) assess

the accuracy of the numerically predicted temperature profiles for the structural steel components and the

fire-proofing through comparison w ith experimental data. In this manner, the efficacy of the proposed

coupled methodology is tested through a series of calculations and their comparison with experimental

measurements.

1 .3 OVERVIEW OF EXPERIMENTS

A series of large-scale experiments was conducted in the NIST Large Fire Laboratory from March 13 to March 26, 2003. Within a steel-frame compartment (3 m by 7 m by 4 m) lined with calcium silicate board (referred to as Marinite) were placed several steel components: two trusses, one thin-walled tubular

column, and a simple rod. The components were either left bare or had a fibrous sprayed fire-resistive material (SFRM) applied in two nominal thicknesses (17 mm and 34 mm). The fire was generated using

liquid hydrocarbon fuels introduced by a two-nozzle spray burner onto a 1 m by 2 m pan. Two fuels were

This reference is to one of tiie companion documents from this Investigation. A hst of these documents appears in the Preface

to this report.

NIST NCSTAR 1-5B, WTC Investigation 3
Chapter 1

used: either a commercial blend of heptane isomers or a mixture of the commercial heptane with toluene

in the proportions of 60 percent heptane and 40 percent toluene (by mass). The latter produced a significantly sootier fire. The openings in the room through which fresh air entered were located 1 m

above the floor. The openings through which heat and combustion products were emitted were located 2

m above the floor on the other end of the enclosure.

The experimental configuration and fire scenarios were selected to examine the effects of (1) fire size or

heat release rate, (2) distance between fire and a structural element, (3) smokiness of the fire, (4) the

impact of various thicknesses of the SFRM, and (5) fire immersion on the thermal response of vertical

and horizontally mounted structural steel components.

The cross sectional thicknesses of the steel and the fibrous SFRM were selected based on representative

values used in the WTC. The fire heat release rates were selected such that steel temperatures of 600 °C

would be obtained in approximately 60 min of steady burning. The temperature 600 °C was regarded as

significant because steel mechanical properties significantly degrade at 600 °C. For steel temperatures

between 600 °C and 700 °C, the yield strength is reduced by approximately 40 percent and 1 0 percent,

respectively, of the room temperature value, and the modulus of elasticity (stiffness) is reduced to approximately 75 percent and 60 percent, respectively, of the room temperature value (refer to NIST

NCSTAR 1-3D). The structural elements were unloaded in these tests. The 60 min duration was

regarded as important because collapse of the WTC buildings occurred on this time frame.

The experimental configuration included natural ventilation flowing from one side of the compartment to the other. Such a situation may have occurred in WTC 1 with ventilation through open windows on one

side of the building leading to an inner area in the building where chimney-like flow may have occurred.

The video record suggests the plausibility of this flow pattern as smoke was observed to flow out from

near the top of WTC 1 as it burned (NIST NCSTAR 1-5 A). Regardless of the actual WTC ventilation,

the over-ventilated fire scenario selected for testing here is a worst-case condition in tenns of the heat

release rate and in this sense is a good test of the model. The efficacy of the model under conditions of

under-ventilated burning had been investigated with real fuels in a separate series of experiments (Hamins

etal. 2003).

1 .4 TEST OF MODEL ACCURACY

The experiments reported here were designed to provide the data necessary to test the accuracy of the fire and thermal-structural simulations that were applied to model the WTC disaster. The experiments

allowed comparison of measurements and predictions of a number of critical parameters, which facilitated a check of the accuracy of the predictions and ensured that the models were capturing properly the

physics of fire dynamics and thermal-structural behavior.

For these reasons, the compartment was instrumented to measure key transient fire dynamic effects such

as total and radiative heat flux to targets within the compartment and the upper layer compartment

temperatures and species concentrations, as well as the temperatures of bare and insulated steel elements

situated in the upper layer of the compartment. Additional qualitative comparisons of the extent and

shape of the flames within the compartment were considered. The large number of measurements was

intended to provide information beyond single location comparisons with the models. The

instrumentation for the tests measured nearly 350 channels of data, including over 100 thermocouples that

4 NIST NCSTAR 1-5B, WTC Investigation
Introduction

measured the temperature at various locations in the test chamber. Approximately 40 measurements were made to determine the global heat release rate of the fire. Thermocouples were placed on the surfaces of

the walls and ceiling, within the walls, on the surface of the steel components, and at the surface of the

SFRM. Heat flux gauges were placed strategically around the compartment to measure the transport of

radiant energy at various targets in the compartment including locations near the floor and the ceiling.

The values of heat flux and temperature are useful for testing how well the fire model is tracking the

pathways of enthalpy transport associated with the fire, which is one of the most important operations of

the fire model. It is particularly important in the case of the WTC investigation because the fire model is being applied to simulate the time varying boundary conditions on the structural components in the

building.

Inaccuracy in the fire model calculations propagate as inaccuracies in the calculation of temperatures of

the building structure. The fire model can then be thought of as providing boundary conditions for modehng the thermal behavior of the structure. After the experiments were simulated using the FDS, the

time var\'ing computed upper and lower layer temperatures, local absorption coefficients, and depth of the

hot layer were used to compute the heat flux incident on the structural components. The ANSYS finite- element program computed the local temperatures of the structural components for comparison with

values of the temperature measured on the steel and fire resistive coating.

Other important measurements were made during the experiments. The concentrations of smoke, oxygen,

carbon monoxide (CO), and carbon dioxide (CO2) were measured in the upper layer of the compartment.

Since the calculation of heat release rate is dependent on the reaction of oxygen to form CO2.

measurements of this gas specie in the upper layer provided a consistency check on the efficacy of the

model calculation. The velocity at the inlets and outlets of the compartment was coupled with burning of

the fiiel and the rise of temperature in the compartment. Velocity measurements made at a limited

number of locations documented the flows through the inlets and exits of the compartment. Finally, a video record of the experiments was completed to qualitatively compare the model simulations with

obser\ ed fire beha\ ior.

1.5 ORGANIZATION OF THIS REPORT

A description of the instrumentation, procedure, uncertainties, and results are given in detail in the

following chapters of this report. Chapter 2 presents the overall experimental set up. Chapters 3 through 9 describe the measurements of the fuel flow and the heat release rates (Chapter 3), the gas

temperatures (Chapter 4), heat fluxes (Chapter 5), the smoke and gas species concentrations (Chapter 6), velocity measurements at the compartment openings (Chapter 7), compartment surface temperatures

(Chapter 8). and the steel component temperatures (Chapter 9). Chapter 10 provides a description of the

fire simulations, which were conducted using the NIST FDS, and compares the simulation results to the

experimental measurements. Chapter 1 1 provides a description of the simulations used to predict the

temperature rise in the structural steel, and the simulations of the time varying temperature fields in the

structural components are compared to the experimental measurements. Chapter 13 summarizes the

results of this study.

To increase readability, a significant portion of the experimental results are presented in a series of

appendixes. Appendix A discusses the rationale for the selection of the fuel mixture. Appendix B is a

detailed list of the instrumentation, including the location of each measurement device. Appendix C

NISTNCSTAR 1-5B. WTC Investigation 5
Chapter 1

documents the temperatures measured at the east and west thermocouple trees in each of the experiments,

analogous to the measurements presented in Chapter 4. Appendix D documents the bidirectional probe

velocity data obtained at the compartment inlets and outlets in each of the experiments, analogous to the

measurements presented in Chapter 7. Appendix E contains heat flux measurement data supplementary

to that presented in Chapter 5. Appendix F contains information on the location of some of the

thermocouples mounted on the surfaces of the SFRM and the steel components. Appendix G documents

the thermocouple temperature data measured on the surface of the SFRM and the steel components in each of the experiments, analogous to the measurements presented in Chapter 9.

1.6 REFERENCES

ANSYS, Inc. 2003. ANSYS Release 8.0 Documentation.

Forney, G.P., and K. McGrattan. 2004. User 's Guidefor Smokeview Version 4 - A Toolfor Visualizing

Fire Dynamics Simulation Data. NIST Special Publication 1017, National Institute of Standards and

Technology, August.

Hamins, A., A. Maranghides, and G. MulhoUand. 2003. 77?^ Global Combustion Behavior of 1 MWto 3 MWHydrocarbon Spray Fires Burning in an Open Environment, NISTIR 7013, National Institute of Standards and Technology, Gaithersburg, MD, June.

McGrattan, K.B., ed. 2004. Fire Dynamics Simulator (Version 4), Technical Reference Guide. NIST

Special Pubhcation 1018. National Institute of Standards and Technology, Gaithersburg, MD, July.

Prasad, K., and H.R. Baum. 2005. '"Coupled Fire Dynamics and Thermal Response of Complex Building

Structures," Proc. Combustion Institute 30, 2255-2262.

6 NISTNCSTAR 1-5B, WTC Investigation
Chapter 2

Experimental Configuration and General Observations

2.1 test facility description

Experiments were conducted in the National Institute of Standards and Technology (NIST) Large Fire

Laboratory', which is 27 m (90 ft) by 37 m (120 ft) in size. Fire products flow into a large exhaust hood

for measurement of the heat release rate and are then exhausted from the building. The hood for these

experiments is approximately 6 m (20 ft) by 6 m (20 ft) in size and has a maximum design capacity of

approximately 10 MW. A computerized data acquisition system was used to record the instrumentation

signals.

2.2 compartment SIZE AND CONSTRUCTION

The compartment dimensions were 7.04 m (23.1 ft) long by 3.60 m (1 1.8 ft) wide and 3.82 m (12.5 ft) high. The total compartment volume was 99 m"^ (3,420 ft^). The compartment was built inside an

exoskeleton of steel tubing and beams. Two removable panels provided access to load materials from

the compartment. The walls and ceiling were covered with 25 mm of calcium silicate board, Marinite I, manufactured by BNZ Materials, Inc. The floor was covered with 25 mm of gypsum, with Marinite I placed under the fire pan (described in Sec. 2.4). The short west wall of the compartment had the supply

air vents and was defined as the front. The compartment was located so that the exhaust vents were under

the 6 m by 6 m exhaust hood. Figures 2-1 through 2-5 illustrate the compartment layout and its contents.

Source: NIST

Figure 2-1. Compartment contents as seen through the

access panel on the south wall.

NISTNCSTAR 7-56, WTC Investigation 7
Chapter 2

Source: NIST. Figure 2-2. Experimental enclosure contents.

8 NISTNCSTAR 1-5B, WTC Investigation
Experimental Configuration and General Observations

^

aspirated

thermocouple

^

floor heat

flux gauges

Source: NIST

Figure 2-4. Fuel pan, fuel nozzles, and instrumentation in the compartment.

NISTNCSTAR 1-5B, WTC Investigation 9
Chapter 2

3.68 m

Inlets WEST

(0,0;

PLAN VIEW

7.04 m NORTH

Baffle Column

Trusses

It Y

•X

Inlet Vents

Fire Pan

SusDended bars

Fire Pan

Access panels

1.2 SECTION A-A

3.68 m

Truss suDport

Baffle •

(0,0)

Truss support

i

Outlets

-Suspended

bar

3.82 m

NORTH

Column

Column

support

Figure 2-5. Schematic drawing of compartment contents.

10 NISTNCSTAR 1-5B, WTC Investigation
Experimental Configuration and General Obsen/ations

2.3 COMPARTMENT CONTENTS

The compartment contained structural components (targets), instrumentation, and their supports.

Additionally, the compartment contained a fire pan and an air deflecting baffle. Figures 2-1 through 2-A

illustrate the compartment contents relative to the layout. The location of the compartment contents,

including the instrumentation, are referred to a (x,y,z) coordinate system with reference origin on the inner wall of the air inlet (y = 0), along the centerline of the compartment (x = 0), and at floor level

(z = 0) as shown in Fig. 2-5. The following describes the particulars of the compartment contents.

2.3.1 Structural Components

The structural components used in these tests were two steel trusses, two rods, and one column. The

particulars of the steel components along with the specifics of their instrumentation are listed in Chapter 9. Two truss supports were fabricated to support the trusses. The steel rods rested on the lower

chord of the steel trusses. The steel truss supports were made out of 0. 1 m steel tubes covered with

25 mm to 37 mm of Marinite.

2.3.2 instrumentation Supports

The supports housed total heat flux gauges and radiometers. Chapter 5 describes these gauges and their location in detail. Two total flux gauges and two radiometers were placed within a 0.15 m diameter

insulated stainless steel cylinder (see Fig. 5^). The cylinder rested on a 0.1 m steel tube. All

instrumentation wire leads and cooling lines were run through the steel tube station support to the outside

of the compartment. The station support was covered with 25 mm to 37 mm of Marinite. Two other flux

gauges were placed in a 0. 1 m steel tube to measure total fluxes in two orientations near the upper

portions of the column. The instrumentation support adjacent to the upper part of the column was

inserted through the ceiling and covered with 25 mm of Marinite. An additional instrumentation support,

fabricated of 0. 1 m steel tube and covered with 25 mm of Marinite, housed two other flux gauges. This

instrumentation support was located near the base of the column.

2.3.3 Air Deflecting Baffle

A vertical baffle 1 .00 m high, 2.05 m wide, and 0.15 m deep was placed 2 m from the west wall of the

compartment. Baffle placement and dimensions were selected to limit flame tilt and prevent flames from

extending through the outlet vents beyond the compartment. The baffle design was based on FDS

simulations of the compartment, its obstructions, and the fire induced flowfield.

2.3.4 Natural Ventilation

Ventilation in the compartment was induced by the fire. There was no forced ventilation. The

compartment was equipped with eight openings, four low for supply and four high for exhaust air. The

vents were located on the front and back walls. The bottom of the supply vents was 0.3 m above the floor

and 0.7 m tall. The two outer vents were 0.5 m wide and the two inner vents were 0.7 m wide. The top of

the four exhaust vents were 1 m below the ceiling and had similar dimensions to the supply vents.

NISTNCSTAR 1-5B. WTC Investigation 11
Chapter 2

2.4 FIRE AND FUEL FLOW

The fire was designed to generate a constant heat release rate up until the steel elements approached

600 °C. This was accomplished by using a hydrocarbon spray fire. A blend of heptane isomers

(designated here as "heptane") was selected as the baseline fuel. A mixture composed of 60 percent

heptane and 40 percent toluene by mass was selected as the second fuel because it is representative of

fires that yield elevated amounts of soot. Appendix A describes the rationale for selection of the

proportions in the fuel blend and shows that the two fuels span the range of sooting expected during the

burning of office furniture. The density, the heat of combustion, and the carbon to hydrogen ratio for the

fuels were measured for the commercial heptane and n-heptane (Hamins et al. 2003). The heat of

combustion and the carbon to hydrogen ratio for the commercial heptane was within 1 percent (by mass)

of the value for n-heptane (Galbraith Labs 2003). This was within experimental uncertainty for the

measurements. The properties of the fuel mixture were calculated from its components. The fuel mixture was prepared before use in Tests 3 and 4 by measuring heptane and toluene into a large storage container

using a calibrated scale.

The liquid fuel was sprayed tlirough two Bete WLl Vi atomizing nozzles onto the fire pan. The fire pan was located on the floor of the enclosure, 3.5 m from the front of the compartment (air inlet side) as seen

in Figs. 2-1 and 2-4. The fire pan was constructed out of 3 mm stainless steel. The pan measured 2 m by

1 m by 0. 1 m deep. The flow through the nozzles was calibrated before each test with a stopwatch and a graduated cylinder.

2.5 TEST MATRIX

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

the date the test was conducted, the nominal fire heat release rate (Q), the fuel type, the nominal

thickness of the SFRM on the steel, and the approximate test duration. The experimental series consisted

of six tests. The nominal heat release rate was 2 MW in Tests 1 through 3 and 3 MW in Tests 4 through 6. In Chapter 3, Table 3-5 lists the actual heat release rate, which was measured. The fuel mixture was used

in Tests 3 and 4; otherwise heptane was used as the fijel. The duration of the fuel flow was nominally

15 min for Tests 1, 3, and 4, and 50 min for Tests 5 and 6. Test 2 was stopped after 6 min because a number of the thermocouples on the steel mistakenly read over 600 °C (the preset limit for the test). The

components were tested in an unprotected (bare) state in Tests 1, 2, and 3 and with fibrous SFRM in Tests 4, 5, and 6. The SFRM was typically applied in two nominal thicknesses ( 1 7 mm or 34 mm) on

different steel elements to test the effectiveness of various insulation thicknesses.

Table 2-1. Test matrix.

Nominal Q SFRM Nominal

Test Date (MW) Fuel Thickness (mm) Duration (min)

1 3/13/03 2.0 Heptane None 14.3 2 3/17/03 2.4 Heptane/toluene'' None 6.2

3 3/18/03 2.0 Heptane/toluene^ None 16.3 4 3/20/03 3.2 Heptane 17 and 34 14.0

5 3/24/03 3.0 Heptane 17 and 34 51.2

6 3/26/03 3.0 Heptane 17 50.3

a. The fuel mixture was composed of 60 percent heptane and 40 percent toluene by mass.

12 NISTNCSTAR 1-5B, WTC Investigation
Experimental Configuration and General Obsen/ations

2.6 INSTRUMENTATION

More than 350 channels of data were acquired for each of the six experiments. Table 2-2 summarizes the

instrumentation, measurement technique, measurement locations, and the chapters of this report where

they are discussed. Appendix B is a detailed list of the instrumentation, including the location of each

measurement. Data were acquired with a sampling rate of 1 Hz. The test duration was approximately

15 min for Tests 1. 3, and 4; 6 min for Test 2; and 50 min for the last two tests, yielding approximately

1 million discrete pieces of data.

Table 2-2. Measurements.

(Chapter) Measurement Method Location

(3) Heat release rate Oxygen consumption calorimetry Exhaust duct

(3) Fuel flow Calibration before test Fuel pan

(4) Gas temperature and location of upper/lower layer

1. Bare-bead thermocouples

2. Aspirated thermocouples

Thermocouple tree at two

locations

(5) Heat flux to steel and SFRM 1 . Total and radiative flux gauges

2. Slug calorimeters with

thermocouples

Attached and embedded on

surfaces

(6) Upper layer soot Gravimetric extractive sampling with

heated sample line Upper layer

(6) Upper layer 0;, CO:, and CO

volume fractions (on a dry basis)

Nondispersive IR (CO, CO:)

Paramagnetic (0:)

Upper layer

(7) Velocity/temperature 1. Thennocouples/aspirated

thermocouples

2. Bidirectional probes

At inlets and near structural components

(8) Temperature of compartment

surfaces

Bare bead thermocouples Compartment ceiling;

between Marinite layers.

(9) Surface temperature of steel and the SFRM

Bare bead thermocouples Steel surface and SFRM

Video record Digital and Hi-8 video cameras Inside and outside views

Many of the measurements involved thermocouples measuring temperature on the surfaces of the walls

and ceiling, within the walls, and on the surface of steel components and its SFRM coating.

Approximately 40 independent measurements were involved with oxygen consumption calorimetry,

which was used to determine the heat release rate. Heat flux gauges were placed strategically around the

compartment to measure the transport of radiant energy. Twelve Schmidt-Bolter total heat flux gauges,

two ellipsoidal radiometers, and five slug calorimeters were used to measure the total and radiative heat

fiux onto and near the steel components. A comparison of the fluxes between the ellipsoidal and

Schmidt-Bolter measurements allowed differentiation of radiative and convective heat flux. Slug

calorimeters, which are a large piece of brass with embedded thermocouples, were used to detemiine the

net heat flux onto their surface. The slug calorimeter design was based on calculations of its heat-up time,

the experimental duration, and geometric considerations in terms of the fire experiments. Thermocouples were used throughout the fire enclosure to characterize the temperature of the walls, ceiling, and the gas

phase. Temperature measurements in the ceiling above the flux meters using bare-bead thermocouples

gave information on the location of the fire plume. Two thermocouple trees, each composed of 15 Type

NISTNCSTAR 7-56, WTC Investigation 13
Chapter 2

K thermocouples aligned vertically, were positioned within the enclosure. Surface temperature

measurements were made in the ceiling and walls by embedding the bead of type K themiocouples into

the Marinite. Bidirectional probes and thermocouples were used to measure flow velocities near the inlets and outlets. A number of aspirated thermocouples were used to check the validity of the bare bead

thermocouple measurements.

Thermocouple and flux gauge signal wires, and tubes for the flux gauge water cooling lines, the fuel flow,

aspirated thermocouples, and gas sampling were passed into the compartment through small holes drilled

through the floor and walls near the instrumentation. Small gaps around the wires and tubing in the holes were filled with SFRM. Since the large access panel to the compartment comprised most of the south

wall of the compartment, signal wire and necessary hardware were passed mainly through the north wall. Signal wire for the bidirectional probes and thermocouples positioned at the middle of the inlet and outlet vents was insulated with a combination of Kaowool and an outer sheet of aluminum foil. Figure 2-6

shows the outside of the west wall with the signal wires and tubing running on the outer wall of the

compartment.

Source: NIST. Figure 2-6. Instrumentation on the outside of the west wall of the compartment.

Chapters 3 to 9 of this report describe each of the measurements in detail, including the instrumentation

that was used, an estimate of uncertainty, and the experimental results. Table 2-3 summarizes some of

this information and lists the character of the instrumentation, including typical values of the expanded

measurement uncertainty (with an expansion factor of 2), the measurement time response, and estimates

of the spatial resolution of the measurement. The time response varied from 0.2 s for the flux gauges to 100 s for the soot gravimetric measurement. The spatial resolution of the measurements varied from

14 NISTNCSTAR 1-5B. WTC Investigation
Experimental Configuration and General Obsen/ations

sub-centimeter to 25 cm for the slug calorimeter. These properties characterize and limit the utility of a measurement and are important considerations in tenns of comparison of models with measurements.

Table 2-3. Instrument characterization

Instrumentation

Typical Uncertainty

(Expanded; k = 2)

Nominal

Time

Response (s)

Spatial

Resolution

(cm)

Heat release rate via oxygen

consumption calorimetry

11 %

(for ^ > 0.8 MW)

15 Global measure

Radiative heat flux: ellipsoidal radiometer

10% = 0.2 0.3

Total heat flux: SchmidtBoelter gauge

6% «0.2 0.6 to 2

I otal neat ilux: slug

calorimeter

8% to 10%for^ = 2MW

16% to 20%for^ = 3 MW

10 25

Velocity: bidirectional probe 15 % inlet 10 % outlet

1 2

Gas temperature: bare bead

thermocouple < 76 °C (dependent on upper layer

temperature, which varied with time, location and test)

1 0.5

Gas temperature: aspirated

thermocouple

7 °C 1 5

Temperature of steel surface -9 °C to + 1 °C 1 0.3 Temperature of SFRM

surface

-7 % to + 5 °C at steady state 1 0.1

Temperature of Marinite

surface ^ 50 °C at steady state 1 0.3

Soot gravimetric 15 % 100; timeaveraged

=5

CO \ ia nondispersive IR^ 7 % «10 =5

CO; via nondispersive IR^ 2% ==10 =5

0: via paramagnetic'' 5 % = 10 »5

a. Gas volume percentages are reported on a dry basis.

2.7 OBSERVATIONS

Figure 2-7 is a frame extracted from the digital video record of the fire impinging on the bare steel trusses 10 s after ignition in Test 1 . The camera viewed the compartment through a window located 20 cm above

the floor on the north side of the west wall of the compartment. The baffle, instrument support, column,

and exit vents were visible. Figures 2-8 and 2-9 are photographs of the smoke and hot gases exiting the

outlet vent on the east wall of the compartment during Tests 2 and 5, respectively. The combustion

products contained copious amounts of soot in Test 2 as seen in Fig. 2-8 when the fuel was composed of

the heptane/toluene fuel mixture. This was also the case in Test 3, but did not occur in Test 5 as seen in

Fig. 2-9, which shows the glowing hot gas products exiting the outlet vent. Tests 1 and 4 through 6 used

NISTNCSTAR 1-5B, WTC Investigation 15
Chapter 2

heptane as the fuel, and heptane is known to produce less soot than toluene (see discussion in Appendix A). The measurement of the soot concentration in the upper layer of the compartment is consistent with these observations (see Chapter 6).

Source: NIST. Figure 2-7. Fire impinging on the bare steel trusses 10 s after ignition in Test 1.

Source: NIST. Figure 2-8. Smoke exiting the outlet vent into the exhaust hood during Test 2.

16 NISTNCSTAR 1-5B, WTC Investigation
Experimental Configuration and General Observations

Source: NIST. Figure 2-9. Radiating hot gases exiting the outlet vent into the

exhaust hood during Test 5.

During Test 4, the Type K Omega GG Glass Braid wire thermocouple extension cables failed, leading to erroneous thermocouple readings. This was likely due to the opening-up of the Kaowool thermal

insulation protective blanket around the thermocouple wires and subsequent shorting. To prevent reoccurrence of this problem during Tests 5 and 6, the extensions were water cooled and double wrapped

with Kaowool insulation. Each layer of Kaowool insulation was secured to the thermocouples using steel wire. Additionally, the thermocouple extensions were rerouted through the cooler west side of the

compartment. After Tests 5 and 6. the thermocouple extensions were visually inspected and were

obserx ed to be undamaged.

2.8 TEST PROCEDURE

The list below identifies the procedure sequence performed to prepare and conduct the experiments:

• Checked that all fixed instrumentation was functioning

• Installed steel components (column, trusses, and rods)

• Photographically documented steel components

• Connected and checked steel member instrumentation

• Calibrated liquid spray burner fuel flow

NISTNCSTAR 1-5B. WTC Investigation 17
Chapter 2

• Conducted final check of compartment instrumentation

• Installed experimental enclosure access panels

• Initiated radiometer and calorimeter cooling

• Calibrated heat release rate oxygen calorimeter using natural gas burner

• Conducted safety brief

• Collected background/baseline data

• Ignited spray burner

• Monitored steel temperatures

• Secured spray burner when steel temperatures exceeded 600 °C

• Allowed compartment to cool to ambient

• Photographically documented steel components

• Removed experimental enclosure access panels

• Disconnected steel member instrumentation and removed steel components

After each experiment, at least one day was required to remove the steel components, inspect the

compartment, patch cracks in the Marinite, check instrumentation, and install new steel components.

2.9 REFERENCES

Galbraith Labs. 2003. ¥Aioxvi\\e,TN, Report of Test Results, March.

Hamins, A., A. Maranghides, and G. MulhoUand. 2003. The Global Combustion Behavior of 1 MWto 3 MWHydrocarbon Spray Fires Burning in an Open Environment, NISTIR 7013, National Institute of Standards and Technology, Gaithersburg, MD, June.

18 NISTNCSTAR 1-5B, WTC Investigation
Chapter 3

Fuel Flow and Heat Release Rate

The heat release rate of a fire is a critical experimental parameter as it represents the amount of energy per

unit time that is generated by the fire. The value of the heat release rate drives the thermal environment,

including the radiative and convective heat transfer, the fluid flow, the thermal insult onto secondary

items such as structural components, the rate of fire spread and growth, the volume of smoke and toxic

products, and ultimately, the hazard associated with a fire. This chapter is broken into several parts that address the heat release rate determined from the fuel flow and oxygen consumption calorimetry and

measurement uncertainty. This chapter also contains a discussion of the measurement results.

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

fuels were heptane and a mixture of heptane and toluene. The nominal heat release rate was 2 MW in Tests 1 through 3 and 3 MW in Tests 4 through 6. The fuel flow was rapidly ramped up from zero to the

steady value in less than 5 s. The mean flow during the steady period is listed in Table 3-1 . At the end of

the bum. the fuel flow was rapidly shut, but data collection continued for at least 1 ,000 s.

The heat release rate ( Q) was estimated as the product of the mass flow rate of fuel ( ih ), the heat of

combustion of the fuel (He), and the fuel combustion efficiency (Xa ):

A companion series of experiments was conducted to measure the global combustion properties of the

hydrocarbon fuels used here (a commercial blend of heptane isomers, and a mixture of the heptane isomer

blend and toluene) (Hamins et al. 2003). The measurements included the combustion efficiency, radiative

fraction, and the yields of soot, CO, and CO: from a downward-oriented spray fire burning in a configuration similar to that used here, with the spray fire burning in the open (Hamins et al. 2003). This

information is often needed as input for fire models. FDS, for example, uses information on the yields of

soot and CO as input. For convenience those results are summarized in Table 3-2, which hsts the

measured combustion properties of the test fuels, including the combustion efficiency, and the yields of

soot, CO, and CO2. Measurements of the fuel density and the heat of combustion per mass of oxygen

consumed, which was used in the calorimetry measurements, are described in Hamins et al. (2003).

The mass flow rate of fuel is equal to the product of the measured volumetric flow rate of fuel ( V ) and

the fuel density (p):

The estimated heat release rate assumes that the combustion efficiency in the compartment was the same

as that measured during the free bum experiments, which were in the open. The uncertainty in Q

(see Table 3-2) is dominated by the uncertainty in Xa, and to a lesser extent by uncertainties associated

with m and He. The uncertainty in Xa was described in detail in Hamins et al. (2003). The uncertainties

3.1 fuel flow and estimated heat release rate

^ = Xa W He (3-1)

m ^ p V (3-2)

NISTNCSTAR 7-56, WTC Investigation 19
Chapter 3

in Table 3-2, are expressed as the expanded relative uncertainty with an expansion factor equal to two

(i.e., 2-a), which represents a 95 percent confidence inter\'al.

Table 3-1. Fuel flow and expected heat release rate.

Test Fuel

Fuel Flow

(L/niin)

Idealized 0

(MW)

{^nmhii^tinn

Efficiency

Expected Q

(MW)

1 Heptane 3.80 ±0.04 1.96 ±0.02 1.0 ±0.05 1.96 ±0.10

2 Heptane/toluene^ 4.8010.12 2.61 ±0.06 0.89 ±0.05 2.33 ±0.14

3 Heptane/toluene'' 3.97 ±0.07 2.1610.04 0.89 ± 0.05 1.93±0.10

4 Heptane 5.65 ±0.14 2.91 ±0.07 1.0 ±0.05 2.91 ±0.16

5 Heptane 5.58 ±0.15 2.88 ±0.08 1.0 ±0.05 2.88 ±0.16

6 Heptane 5.54 ±0.23 2.86 ±0.12 1.0 ±0.05 2.86 ±0.19

a. The fuel mixture was composed of 60 percent heptane and 40 percent toluene by mass.

Table 3-2. Combustion properties of the t est fuels.

Fuel" CombustionEfficiency

Radiative

Fraction Soot Yield CO Yield CO2 Yield

Heptane"^ 1.01 ±0.14^^ 0.44 ± 0.06 0.0149 ± .0033 <0.008 3.03 ±0.12

Heptane/toluene'^ 0.90 ±0.13 0.39± 0.08 0.1 14 ±0.022 0.042 ±0.016 2.70 ±0.15

a. Data from Hamins et al. 2003.

b. A commercial blend of heptane isomers; not n-heptane.

c. By definition X., < 1 . The uncertainty in the determination of %^ is dominated by uncertainty in the measurement of the heat

release rate.

d. The fuel mixture was composed of 60 percent heptane and 40 percent toluene by mass.

3.2 OXYGEN CONSUMPTION CALORIMETRY

3.2.1 Description of the Heat Release Rate Measurement Facility

Much of this section of the report is based on Bryant et al. (2003). The fire heat release rate ( Q) was

measured using oxygen consumption calorimetry in the standard manner through measurement of exhaust

products in the 6 m by 6 m hood. This form of fire calorimetry was used by Huggett (1980), who showed

that the oxygen deficit in the duct flow (relative to ambient air), is a measure of the heat release rate of a

fire. Parker (1982) and Janssens and Parker (1995) discuss the details of the heat release rate calculation.

The combustion products from the compartment flowed into the exhaust duct where the gas flow was

characterized. Bryant et al. (2003) give a detailed description of the experimental methodology, hardware,

instrumentation, calibration, experimental procedures, and uncertainty. The facility was designed to provide quantitative results with minimum uncertainty. A large number of instruments were employed in the heat release rate measurement (see Table 3-3), including the measurement of temperature and

pressure in the exhaust duct. Gas was sampled in the exhaust duct and transported to the instruments in a control room for measurements of the oxygen, carbon dioxide, and carbon monoxide concentrations.

Relative humidity was measured in the test bay area, and a correction was made for ambient humidity.

The computation of heat release rate ( Q ) was made following Bryant et al. (2003) shown in Eq. 3-3.

20 NISTNCSTAR 1-5B, WTC Investigation
Fuel Flow and Heat Release Rate

m

M ....

with

2 Xo/\ + <t){a - 1)

~ ^o. ~ ^co-, ~ ^co )^o.

(3-3)

where:

Ehc ^ heat of combustion of hydrocarbon fliel

Eco ^ heat of combustion of carbon monoxide

^ = oxygen depletion factor

rh^ - mass flow rate in exhaust duct

m^-^ - mass flow rate of air

a = combustion products expansion factor

Mi - molecular weight of species i

Xi - exhaust-gas concentration measurement of species i

X° = ambient-gas concentration measurement of species i

Table 3-3. Instruments used in the Calorimeter System.

Parameter

Measured

Instrument

Type

Make and

Model Ranges

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

Carbon dioxide Infrared extinction Siemens Ultramat 5E 0 % - 5 % v/v

Carbon monoxide Infrared extinction Siemens Ultramat 5E 0 % - 3 % v/v

Relative humidity Thin fdm capacitance Dickson THDx 0 % - 95 %

Temperature Thermocouple Omega K-type 0°C- 1,250 °C

Exhaust velocity Bidirectional probe MKS model 220 differential pressure transducer OPa- 10 Pa

Key: v/v=volume percent.

3.2.2 Calibration and Measurement Uncertainty

A calibration burner was used to establish that the Heat Release Rate Facility provided repeatable results within 2 percent for fire sizes of 1 MW and larger (Bryant et al. 2003). Table 3^, which is reproduced

from Bryant et al. (2003), summarizes a comprehensive study to determine the expanded relative

uncertainty for the measured peak heat release rate in the 6 m exhaust hood that was used in the

experiments. A propagation of uncertainty analysis showed that the uncertainty was dominated by

NISTNCSTAR 1-5B, WTC Investigation 21
Chapter 3

uncertainty in the measurement of the oxygen concentration, the exhaust mass flux, and the heat of

combustion of the fuel (Biyant et al. 2003). The expanded uncertainty (with a coverage factor k equal 2)

corresponds to a level of confidence of 94.5 percent. The term "relative" means that the uncertainty is expressed as a ratio of the uncertainty to the measured value and represents two times the standard

deviation of the uncertainty.

Table 3-4. Expanded relative uncertainty of the measured heat release rate. Heat Release Rate (MW) Expanded Relative Uncertainty"

0.10-0.20 ±0.235

0.20-0.40 + 0.187

0.40-0.80 ±0.110

0.80-3.00 ±0.107

a. From Bryant et al. (2003).

Calibration bums using natural gas were conducted several times during the test series to assure accurate measurement of the heat release rate ( Q ). The calibration burner used natural gas flowing at a measured

and controlled rate. This provided an independent measurement of the heat release rate to compare to the

measurement by oxygen consumption calorimetry. The heat output of the burner was held constant for

3 min to 5 min at each setting. The measured Q was typically within 10 percent of the expected heat

release rate based on the natural gas mass flow rate.

The calorimetry measurement response time was on the order of 15 s (Bryant et al. 2003). The volume of

the test compartment also affected the time response of the measurement as filling and mixing caused a time-averaging and a lag in the calorimetry results.

3.3 RESULTS AND DISCUSSION

Table 3-5 summarizes the heat release rate measurement results and includes the mean measured heat

release rate ( Q ), its expanded uncertainty, and the experimental duration. For comparison to the

measured Q , the expected Q (taken from Table 3-1 ) is also listed. The mean heat release rate was

computed from data acquired over the period encompassing 100 s after burner ignition to 50 s prior to the

burner being extinguished. The relative expanded uncertainty of the mean heat release measurement was

1 1 percent (see Table 3-4).

Table 3-5. The expected Q, the measured mean Q, and the experimental duration.

Test Fuel Expected" Q (kW) Measured Q (kW)

Experimental

Duration (s)

1 Heptane 1,960 ± 100 1,970 ±220 860 ±4

2 Heptane/toluene'' 2,330 ± 140 2,410 ±260 370 ±4

3 Heptane/toluene^ 1,930 ± 100 1,990 ±220 980 ±4

4 Heptane 2,910 ± 160 3.240 ± 360 840 ±4

5 Heptane 2.880 ± 160 3,000 ± 330 3,070 ±4

6 Heptane 2;860± 190 2,980 ±330 3.020 ±4

a. The expected Q is taken from Table 3-1.

b. The fuel mixture was composed of 60 percent heptane and 40 percent toluene by mass.

22 NISTNCSTAR 1-5B, WTC Investigation
Fuel Flow and Heat Release Rate

Figure 3-1 shows the heat release rate as a function of time for Tests 1 through 6. The noise or range of

scatter in the data is on the order of 200 kW for all of the tests, which is on the same order of magnitude

as the uncertainty (see Table 3-5). The results show that Q was relatively steady for Tests 2 and 3, while

O continuously increased by approximately 7 percent and 6 percent, during Tests 1 and 4, respectively. In Test 6, Q increased by approximately 5 percent over the first 600 s and then became nearly steady. In Test 5, O increased by approximately 1 percent over the first 600 s, became steady, and then decreased

by approximately 2 percent over the last 1,200 s. The experimental duration is listed in Table 3-5. At the completion of the test, the fuel flow valve was secured. The fuel flow continued for approximately 10 s to 20 s. as the line pressure equilibrated. Even after this time, a finite value of Q was measured in the

exhaust hood due to continued flow of hot combustion products exiting the compartment into the

exhaust hood.

Time (s)

Figure 3-1. The measured heat release rate (^) as a function of

time during Tests 1 through 6. A comparison of the expected and measured heat release rates in Table 3-5 shows agreement within

experimental uncertainty. This suggests that the free bum combustion efficiency measurements for the

fuel mixture are an adequate representation of the fires burning in the compartment. The value of ^ is a very important parameter as it impacts the dynamics of the fire within the compartment and influences the temperature and other parameters in the upper layer. As FDS uses Q as input, the uncertainty in Q can be linked to uncertainty in various important parameters in the FDS output (see Chapter 10).

3.4 REFERENCES

Bryant, R., T. Ohlemiller, E. Johnsson, A. Hamins, B. Grove, W.F. Guthrie, A. Maringhides, and

G. Mulholland. 2003. The NIST 3 Megawatt Quantitative Heat Release Rate Facility - Procedures

NISTNCSTAR 7-56, WTC Investigation 23
Chapter 3

and Guidance, NIST Special Report 1007. National Institute of Standards and Technology,

Gaithersburg, MD, September.

Galbraith Labs. 2003. Knoxville, TN, Report of Test Results, March.

Hamins, A., A. Maranghides, and G. Mulholland. 2003. The Global Combustion Behavior of 1 MW to 3 MWHydrocarbon Spray Fires Burning in an Open Environment, NISTIR 7013. National Institute of Standards and Technology, Gaithersburg, MD, June.

Huggett, C. 1980. "Estimation of the rate of heat release by means of oxygen consumption," Journal of

Fire and Materials, 12, 61-65.

Janssens, M., and W. J. Parker. 1995. "Oxygen consumption calorimetry," Heat Release in Fires,

(Babrauskas, V., Grayson, S. J., Eds., E & FN Spon), pp. 31-59, London, UK.

Parker, W. J. 1982. Calculations ofthe Heat Release Rate by Oxygen Consumptionfor Various

Applications, NBSIR 8 1 -2427. National Bureau of Standards, Gaithersburg, MD

24 NISTNCSTAR 1-5B, WTC Investigation
Chapter 4

Gas Temperature

4.1 INSTRUMENTATION

Most gas temperatures were measured on two vertical thermocouple "trees" referred to as the "East" and

"West" trees with (x, y) coordinates (6.04 m, 0.52 m) and (0.95 m, 0.51 m), respectively (in Chapter 5, see Figs. 5-3 and 5^). Appendix B lists the (x,y,z) coordinates of each of the thennocouples. Type K

thermocouples were used and were constructed from 24 gauge wires with glass braid insulation. The

beads were located on individual horizontal "branches" approximately 6 cm from the vertical tree support. Each tree was suspended from the ceiling and attached to the floor with a screw to provide

tension for maintaining vertical alignment.

Three additional gas temperature measurements were made near each tree using aspirated thermocouples.

These were located at nearly the same (x, y) locations as the trees and within a few centimeters of the

bare-bead thermocouples at heights (z) approximately 0.34 m (low), 2.15 m (mid), and 3.65 m (high)

above the compartment floor. The purpose of the aspirated thennocouple measurements was to generate

measurements more accurate than the bare-bead thermocouples in the highly radiating environment. The

aspirated thermocouples provided a basis for uncertainty estimates for the bare-bead thermocouple

temperature measurements. The aspirated probes used a double-shielded design, which overcomes

radiation effects on the bead much more effectively than a single-shield model (Pitts et al. 1998). The

materials were either 304 SS or Inconel. The outer diameter of the outer shield was 0.95 cm. The probes

extended 1 .27 m into the enclosure from the north wall (53 cm from the compartment centerline, i.e., x -

0.53 m). To generate sufficient velocity (greater than 4 m/s) and convective heat transfer over the bead,

gases were pulled through the probes at a fixed flow of 24 L/min. This flow pulled room temperature gas over the bead at approximately 5 m/s. The gas velocity at the bead was at least 5 m/s and increased in proportion to the ratio of the gas temperature at the end of the probe to the gas temperature at the

flowmeter (near ambient).

4.2 MEASUREMENT UNCERTAINTIES

A discussion of uncertainties is necessary before describing the results of the gas temperature

measurements and the layer height calculations derived from the temperatures. The uncertainties in the

temperatures measured by the bare bead themiocouples in the East and West trees were highly dependent on the radiative and convective environment in which they were located. Whereas a radiative

environment tends to cause a bead to have a temperature different from the local gas temperature,

convection from the local gas velocity tends to drive a bead's temperature closer to that of the gas. Radiation heat transfer has a T^ dependence (Thigh'' - Tiow"*) calculated from the high and low temperatures

(in Kelvin). Upper layer thermocouples tend to underestimate gas temperatures because any view of

relatively cool surfaces acts as a radiative heat sink, lowering a bead's temperature. Lower layer

thermocouples tend to overestimate gas temperatures because they can view very hot surfaces and gases

that radiatively heat a bead. Since the thermal environment surrounding a given thermocouple was very

difficult to characterize, aspirated thermocouples, which do not suffer radiafive exchange problems

NISTNCSTAR 1-5B. WTC Investigation 25
Chapter 4

(McCaffrey and Heskestad 1 976), were used to check the magnitude of the needed temperature correction

at a limited number of locations and to estimate the uncertainties of all of the thermocouples.

Three aspirated thermocouples were located next to the West and East thermocouple trees at the same

heights as the low, middle, and high bare-bead thermocouple locations. The temperature differences

measured by the bare bead and aspirated thermocouples varied during each test and from test to test. Tables 4-1 and 4-2 list the heights associated with the themiocouple pairs on each tree and summarize

these temperature variations as represented by the mean and standard deviation of the differences that occurred during each of the tests. Because of the large differences at the beginning of the tests relative to

later in the tests, a demarcation was made between the transient and quasi-steady periods for each test. Table 4-1 lists the statistics during the first 60 s after ignition for each test. Table 4-2 summarizes the

statistics from 60 s to extinguishment. An exception was the East tree middle pair of thennocouples in Tests 4, 5, and 6. At this location and for these tests, the division between transient and quasi-steady was

set at 240 s because the rise to quasi-steady temperatures was relatively slow. Also, transient statistics were not available for Test 1 as the aspirated thennocouple pumps were not activated until t = 177 s. To

correct a bare-bead thermocouple temperature at or near an aspirated thermocouple, the mean difference

value from the bare-bead temperature should be subtracted.

Table 4-1. Early (<60 s)^ mean difference: bare-bead minus aspirated at same location.

West Tree East Tree

Height 3.65 m 2.16 m 0.33 m 3.65 Ml 2.16 m 0.34 m

mean s.d. mean s.d. Mean s.d. mean s.d. mean s.d. mean s.d. Test No. (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C)

1 NA NA NA NA NA NA NA NA NA NA NA NA

2 52 46 46 31 1 1 48 56 56 38 16 6

3 44 36 46 27 1 1 34 43 52 33 17 6

4 30 59 67 50 5 2 132 68 183 105 30 10

5 74 57 48 39 4 2 166 89 199 110 33 10

6 75 56 54 38 6 2 181 80 190 115 38 10

a. For Tests 4, 5, and 6, the East Tree difference at Z = 2. 16 m was calculated for t<240 s (early times) and t>240 s (late times). Key: NA. not available.

Table 4-2. Late (>60 s)^ mean difference: bare-bead minus aspirated at same location.

West Tree East Tree

Height (Z) 3.65 m 2.16 m 0.33 m 3.65 III 2.16 m 0.34 m

mean s.d. mean s.d. mean s.d. mean s.d. mean s.d. mean s.d. Test No. (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C) (°C)

1 -16 8 -10 4 2 1 -13 7 2 5 6 3

2 -13 5 -5 4 2 1 -15 5 1 5 11 3

3 -12 5 -4 4 5 2 -17 5 -1 3 16 4 4 -41 17 -5 4 26 12 4 11 5 17 35 6

5 -7 6 -1 4 40 15 9 12 -43 32 53 13

6 -3 6 5 3 38 13 8 15 -53 26 36 9

a. For Tests 4, 5, and 6, the East Tree difference at Z = 2. 16 m was calculated for t<240 s (early times) and t>240 s (late times).

26 NI ST NCSTAR 1-5B, WTC Investigation
Gas Temperature

One consistent pattern in the differences is that during the early periods, the mean difference was always

positive which means that the bare-bead temperature was always higher than the corresponding aspirated

temperamre because of early radiative heating with slowly increasing real gas temperatures. Also, during

the transient periods, the East tree experienced much greater mean differences and fluctuations (seen in the standard deviations) than did the West tree. During the quasi-steady periods, all of the high and

middle height differences on the West tree were negative, which means that the real gas temperatures were higher than the measured bare-bead temperatures. On the East tree, these locations were not

generally consistent. During the quasi-steady periods at low positions on both trees, all differences were

positive, which indicates consistent radiative heating of the lower-layer thermocouples above the real gas

temperamres.

Tables 4-3 and 4-4 detail the resulting expanded temperature uncertainty limits (relative to the nominal

temperatures) for regions surrounding the six locations. The uncertainties were calculated using the means and standard deviations from Tables 4-1 and 4-2 as well as the inherent uncertainties (k = 2a) in each thermocouple of about 5 °C and the approximately 5 °C uncertainty in the extent to which aspiration causes the aspirated thermocouple to approach the real gas temperature (Blevins and Pitts 1999). The

expanded measurement uncertainty based on these two effects is equal to 7 °C. The uncertainty limits are asymmetric and consider radiation effects as well as the inherent uncertainty associated with

thermocouple calibration. To generate the absolute expanded (k = 2a) uncertainty limits for a nominal, or uncorrected, temperature, add the lower and upper relative uncertainty limits to the nominal temperature.

Figures 4-1 through 4-6 are plots showing the differences (with a 20 s running average) between the

bare-bead and aspirated thermocouples on the West and East thermocouple trees for the six experiments.

Table 4-3. Early (< 60 s)^ lower and upper expanded uncertainty limits

Regions

\\ est Tree East Tree

Upper

2.90 to 3.82 m

Middle

L25to2.90iii

Lower

0 to L25 m

Upper

2.90 to 3.82 m

Middle

1.25 to 2.90 m

Lower

0 to L25 m

L'nc Limits Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C) Test

1 NA NA NA NA NA NA NA NA NA NA NA NA

2 -144 41 -108 17 -10 8 -161 64 -132 20 -30 -1

-> J) -116 28 -100 8 -10 8 -120 53 -117 14 -32 -1

4 -148 89 -166 33 -15 4 -268 4 -393 28 -51 -9

5 -189 41 -125 30 -14 5 -344 13 -419 21 -55 -11

6 -186 37 -131 23 -16 3 -342 -21 -420 40 -59 -16

a. For Tests 4, 5, and 6, the East Tree difference at Z

times). Key: NA, not available.

.16 m was calculated for t<240 s (early times) and t>240 s (late

NISTNCSTAR 7-56, WTC Investigation 27
Chapter 4

Table 4-4. Late (> 60 s)^ lower and upper expanded uncertainty limits

Regions

West Tree East Tree

Upper

2.90 to 3.82 m

Middle

1.25 to 2.90 m

Lower

0 to 1.25 m

Upper

2.90 to 3.82 m

Middle

1.25 to 2.90 m

Lower

Oto 1.25 m

Unc

Limits Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C)

Lower

(°C)

Upper

(°C) Test

1 -2 34 -2 23 -11 7 -4 29 -16 12 -17 5

2 0 27 -8 17 -11 7 1 29 -14 12 -21 0

3 -1 25 -8 16 -15 5 4 30 -10 12 -28 -3

4 7 76 -7 17 -52 0 -28 20 -41 30 -50 -20

5 -8 21 -10 12 -72 -8 -33 16 -22 107 -80 -26

6 -12 18 -16 7 -65 -10 -40 24 0 106 -56 -15

a. For Tests 4, 5, and 6, the East Tree difference at Z = 2.16 m was calculated for t<240 s (early times) and t>240 s (late times).

10

o

l-c OhS

oH

10

-20

-30

-[ 1—I

\

1—

r

-t4 ^ te f

West High

West Mid

West Low —f£as! High

East Mid

- East Low Test 1

J 1 \ I I 1 I I L J I I I L. 240 320 400 480 560 640 720 800 880

Time (s)

Figure 4-1. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 1.

28 NISTNCSTAR 1-5B, WTC Investigation
Gas Temperature

0 50 100 150 200 250 300 350

Time (s)

Figure 4-2. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 2.

East Mid

- East Low

200 400 600 800

Time (s)

Figure 4-3. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 3.

NISTNCSTAR 1-5B, WTC Investigation 29
Chapter 4

T I \ I I r 1 I I I I r ~i—I—I

\—

r

u

-4-> a

sH

80

40

-40

-80

West High - —— East: High.

West Mid " East Mid

West Low " " •

1 .() w

0 100 200 300 400 500 600

Time (s)

700 800

Figure 4-4. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 4.

U

D

o

Ohg

H

40

- \

-40

-80

-120

T 1 1 r "T 1 r—

r

n 1 r

* . % -a ^

' ' 'I

3^

~ West High •-»—

East High

- West Mid —" East Mid,

West Low ~ Fast \.v>v%-

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 500 1000 1500

Test 5

J ^ ^ L 2000 2500 3000

Time (s)

Figure 4-5. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 5.

30 NISTNCSTAR 1-5B, WTC Investigation
Gas Temperature

Time (s)

Figure 4-6. 20 s average of the difference of bare-bead thermocouple temperatures

minus aspirated thermocouple temperatures as a function of time for Test 6.

4.3 MEASUREMENT RESULTS

Figures 4-7 and 4—8 are plots of all of the aspirated thermocouple temperatures versus time for Tests 1 and 5, respectively. The upper and middle height temperatures are shown for the other tests in subsequent

comparison plots. Note that at the West location, low position temperatures remain less than 30 °C for Test 1 and less than 60 °C for Test 5. The low position temperatures at the East location remain near 100 °C for Test 1 and 200 °C for Test 5. Since the temperatures at the low positions on each tree are consistently not of the magnitude of interest for this study, further plots focus on the middle and high

position results. Figures C-1 through in Appendix C document analogous measurements from

Tests 2, 3. 4, and 6. respectively.

4.3.1 Time Dependence

The plots in Figs. 4-7 and 4-8 show time dependence in that the temperatures continued to rise throughout each experiment. The temperatures during Test 5, which was allowed to proceed much longer

than Test 1 because of the insulated steel, approached steady values by the end of the experiment. High

and middle position Test 1 temperatures were still rising at about 10 °C/min at the end of the experiment.

4.3.2 Effect of Heat Release Rate

Figure 4-9 is a plot of the upper and middle position aspirated thermocouple temperatures versus time for Tests 1 and 4. The experiments were identical except that the Test 1 heat release rate ( Q ) was 2 MW

whereas in Test 2, Q was 3 MW. The West tree high position temperature was 205 °C greater for Test 4

than for Test 1 . The differences for West middle. East high, and East middle positions were 185 °C,

NISTNCSTAR 1-5B. WTC Investigation 31
Chapter 4

245 °C, and 230 °C, respectively, over the minute prior to 800 s after ignition (an interval in common late

in the fires over which the temperatures were averaged). The magnitude order of the temperatures was

different between Tests 1 and 4 with the West middle position hotter than the East middle position for Test 1, but cooler for Test 4.

Figure 4-10 is a plot of the upper and middle aspirated thermocouple temperatures versus time for Tests 2

and 3. Again, the experiments were identical except that the Q in Test 2 was 2.4 MW, whereas the Q in Test 3 was 2 MW. The West tree high position temperature was 80 °C greater for Test 2 than for Test 3. The differences for West middle. East high, and East middle positions were 75 °C, 95 °C, and 65 °C,

respectively, over the minute prior to 380 s after ignition. The relative order of the temperatures was the same for both experiments. Temperatures at the East tree were generally higher than those at the West

tree at the highest position, but at the middle position, the West temperatures were higher.

4.3.3 Effect of Smoke Production

Figure 4-11 is a plot of the upper and middle aspirated thermocouple temperatures as a function of time

for Tests 1 and 3. The experiments were identical except that the Test 1 fuel was heptane whereas the Test 3 fuel was a mixture of heptane and toluene. The high and middle position temperatures in Test 1 were consistently 15 °C to 25 °C higher than for the same locations in Test 3 over the minute prior to 880 s after ignition. This is consistent with the expectation that the smokier, and therefore, less efficient fuel would release less heat energy, and that the higher flame emissivity due to the soil would lose more

heat via radiation (as opposed to the soot trapping the radiation, which could lead to higher temperatures).

4.3.4 Repeatability

The aspirated thennocouple temperatures for Tests 5 and 6 were compared to determine the repeatability

of the tests and whether the results were within the expected uncertainty. Figure 4-12 shows the middle

and high position temperatures versus time for the two trees. The average difference between

temperatures at the same time from ignition ranged from 1 °C to 5 °C for these positions with Test 5 generally hotter than Test 6. It is not shown on the figure, but one position was not as repeatable as the

rest. The average difference for the lowest height on the East tree was 2 1 °C over the whole test. The

divergence occurred after 800 s. Prior to 800 s, the agreement was within 4 °C, but after 800 s, the

difference averaged 27 °C. There is no obvious explanation for this since the other five positions agreed

well within the expanded measurement uncertainty of about ±7 °C for each aspirated thermocouple.

Thermocouple failure is not a likely explanation for this result. Unexpected leaning of the fire is a possible reason, although this is not clearly demonstrated by the video record.

32 NISTNCSTAR 7-58, WTC Investigation
Gas Temperature

500

400

u

o

300 -

§ 200

100

0

-1—I—I

\—

I—I—

_ Test 1

Aspirated

1—I—I—I—

I—I—I—I—I—

I—i—i—I—I—I—I—I—I—I—

]—I

\—

Pumps On *;vRv':^' f

O-. i-'g

/J — West High

e-- West Mid

A - - West Low

East High

East Mid

East Low

,-<

- /J •

A

. 4' - k.

M'- A- -A- - A - A- A- -ZV -A - A- ^- -A -A - A- A- -Zy,-A - A- A- -b:.

-I \ I I I I I I

I T I

I I I I I I I I I I

I I I r I I I I J I I I \ I I I I I L 0 100 200 300 400 500 600 700 800

Time (s)

Figure 4-7. Aspirated thermocouple temperatures as a function of time for Test 1.

c

H

800

700

600

500

400

300

200

100

0

e— West High

e West Mid

A - - West Low

East High

East Mid

:-A"A--A~-A---A---A--A--^A--"-A--A---A'-A.-.A'--A"--A---A>''A'-A---^-'^

__j I I I I I I I 1 I I I I I I I I I I I I I I I I I \ I I 0 500 1000 1500 2000 2500 3000

Time (s)

Figure 4-8. Aspirated thermocouple temperatures as a function of time for Test 5.

NISTNCSTAR 1-5B, WTC Investigation 33
Chapter 4

I L I I I I I I I I ! I I I I 1 I I I 1 I I I I ! I I I I I I I I I

I '- I

•

200 300 400 500 600 700 800 900

Time (s)

Figure 4-9. Comparison of Test 1 and Test 4 aspirated thermocouple temperatures at the

upper and middle locations as a function of time, showing the effect of heat release rate.

600

Time (s)

Figure 4-10. Comparison of Test 2 and Test 3 aspirated thermocouple temperatures at the upper and middle locations as a function of time, showing the effect of heat release

rate for a sooty fuel.

34 NISTNCSTAR 1-5B, WTC Investigation
Gas Temperature

- Tests 1 and 3

Time (s)

Figure 4-11. Comparison of Test 1 and Test 3 aspirated thermocouple temperatures at the upper and middle locations as a function of time, showing the effect of soot for

nearly the same heat release rate.

0 500 1000 1500 2000 2500 3000

Time (s)

Figure 4-12. Comparison of Test 5 and Test 6 aspirated thermocouple temperatures at

the upper and middle locations as a function of time, showing the repeatability of the

measurements for long tests with relatively large Q.

NISTNCSTAR 1-5B. WTC Investigation 35
Chapter 4

4.4 LAYER INTERFACE HEIGHT

4.4.1 Calculation Uncertainties

The results reported for the hot gas layer interface height are from an analysis of each thermocouple tree's temperature profile using the two-layer reduction method. Figure 4-13 shows the temperatures

from the West thermocouple tree in Test 1 as an example of such a profile. Figures C-5 through C-15

in Appendix C document analogous measurements from the East and West thermocouple trees for Tests 1 - 6. The two-layer reduction model is a one-dimensional analytical method based on the zone model concept. It uses conservation of mass and energy and temperature profiles to define the height at which the upper and lower masses are equal. It can only be applied when the temperatures can be broken

into two distinct regions. The equations used in this method are the following:

n^T{z)dz = {H t)Tti[i + Z int Tlow (4-1)

1 dz = (H -Z mi) 11+ Zml

Jr(-) Tio (4-2)

Tio.jhli-H-)

h + lzTio.- -IT/onH

(4-3)

500 1 i r

J I L \ I 1 1 I I L

e ^ TTRWl

TTRW2

y TTR.W,-

X TTRW4

-H- TTRW5

«- - riRW"- • TTRW8

TTRW9

-A---- TTRWIO

TTRWU

^ TTRWl 2 R

-Ti— TTRWl 3

<!- - TTRW14

L> TTRWl 6

1200 2000 2400

Time (s)

Figure 4-13. West thermocouple tree temperatures as a function of time in Test 1.

36 NISTNCSTAR 1-5B, WTC Investigation
Gas Temperature

where //is the ceiling height (m), Z,„, is the layer interface height (m), T„p is the hot layer average

temperature (K). T/,,,, is the lower layer average temperature (K), T(z) is the temperature as a function of

height (K), and r is the height (m).

The layer interface height uncertainties were calculated by the following procedure. Typical temperature

profiles were chosen for each test within 2 min of extinguishment. Those times were 850 s, 380 s, 950 s, 800 s. 3,020 s, and 3,000 s for Tests 1 to 6, respectively. At those times, the appropriate expanded

temperature uncertainties listed in Table 4-4 were added to the tree temperatures to obtain the two

standard deviation limits around the nominal measurements. Since each tree had three regions of

uncertainty' based on the three aspirated thermocouple temperatures, eight possible combinations of

highest and lowest temperature were generated. These eight temperature profiles and the corrected

profile (each temperature in the center of its range of uncertainty) were applied using the two-layer

reduction equations (Eqs. 4-1 through 4-3). This generated a set of 9 layer heights for each tree. Two

times the standard deviation of these nine values was considered the uncertainty for this calculation.

The estimated expanded uncertainties for the layer interface height are contained in Table 4-5. The West

tree uncertainties for all of the tests are clustered around 1 percent of the values calculated using the

nominal temperature profiles. For the East tree, the uncertainties are significantly higher than for the

West, but are still quite low. They range between 2 percent and 6 percent.

Table 4—5. West and East tree layer heights.

Test

W est Tree Layer Interface Height East Tree Layer Interface Height

End of Test Uncorrected

(m)

Expanded

Uncertainty

(m)

Expanded

0//o Uncertainty

End of Test Uncorrected

(m)

Expanded

Uncertainty

(m)

Expanded

%

Uncertainty

1 1.1(S 0.009 0.8 1.48 0.058 3.9

1.U2 0.010 1.0 1.20 0.030 2.5

~t 1.04 0.009 0.8 1.23 0.030 2.4 4 0.99 0.011 1.1 1.17 0.052 4.5

5 1.17 0.009 0.8 1.15 0.069 6.0

6 1.1 1 0.009 0.8 0.97 0.053 5.4

4.4.2 Calculation Results

Figure 4-14 shows the results of the two-layer reduction method on the themiocouple tree data from

Test 1 . For this figure and the following two, the height is plotted on the y-axis since it allows the reader

to better visualize the temperature profile. The independent parameter is time, not height or temperature.

There is very little trend with time and the fluctuations are not very large. The implication of the plot is that the layer interface is over 1 m below the top of the exit vent. This seems counterintuitive since the smoke layer should not be easily maintained far below the soffit.

NISTNCSTAR 1-5B, WTC Investigation 37
Chapter 4

Figure 4-15 shows the temperature profiles with height for Test 1 at t = 850 s. It suggests that a strong

thermal transition begins between 1 m and 1 .5 m and ends between 2.1m and 2.4 m. This indicates that a mixed transitional layer begins near the bottom of the exit vent and ends near the top with lesser temperature gradients above and below. Figure 4-16 shows results for Test 5, which has a temperature

profile similar in structure to Test 1, but with higher temperatures.

The two-layer reduction method apparently generates the lowest height at which the thermal gradients

increase dramatically from the relatively cool lower layer. The Table 4-5 list of layer heights can still be

used to compare the relative positions of this onset of the upper layer between East and West locations

and different tests. The visible upper smoke layer interface was most likely higher than the tabulated

values.

3.5

3.0

2.5

o

(U

(U

2.0

1.5

1.0

0.5

0.0

—i r Test 1 West Tree B East Tree

Top of Exhaust Vent

0 200 400 600 800

Time (s)

Figure 4-14. Calculated layer interface height as a function of time for Test 1.

38 NISTNCSTAR 1-5B, WTC Investigation
Gas Temperature

200 300

Temperature (°C)

Figure 4-15. Temperature profiles for West and East thermocouple trees

for Test 1 at t = 850 s.

3.5

3 _ 2.5

1 2 = 1.5

1

0.5

0

n—i—I

i—I—I—I—I—

r

— West Bare-Bead • West Aspirated —— East Bare-Bead

East Aspirated

0 100 200 300 400 500 600 700 800

Temperature (°C)

Figure 4-16. Temperature profiles for West and East thermocouple trees

for Test 5 at t = 3,020 s.

NISTNCSTAR 7-56, WTC Investigation 39
Chapter 4

4.5 REFERENCES

Blevins, L. G., and W. M. Pitts. 1999. Fire Safety Journal, 33, 239-259, and National Instimte of

Standards and Technology, Gaithersburg, MD, NISTIR 63 10, April. Available at http://fire.nist.gov/bfrlpubs/fire99/PDF/f99022.pdf.

He, Y. P., A. Fernando, and M. C. Luo. 1998. Fire Safety Journal, 31, 19-38.

McCaffrey, B.J., and G. Heskestad. 1976. Combust. Flame, 26, 125-127.

Pitts, W. M., E. Braun, R. D. Peacock, H. E. Mitler, E. L. Johnsson, P. A. Reneke, and L. G. Blevins.

1 998. Temperature Uncertainties for Bare-Bead and Aspirated Thermocouple Measurements in Fire

Environments, National Institute of Standards and Technology, Gaithersburg, MD, NISTIR 6242;

October. Available at http://fire.nist.gov/bfrlpubs/fire03/PDF/fD3052.pdf

40 NISTNCSTAR 1-5B, WTC Investigation
Chapter 5

Heat Flux Measurements

5.1 OBJECTIVE

The heat flux measurements were used to assess the accuracy of the National Institute of Standards and

Technology (NIST) Fire Dynamic Simulator (FDS) model. The heat flux incident on the surfaces of the

structural components is key to their temperature behavior. However, it was not deemed practical to place flux measuring devices on or in the structural components as the instruments themselves are inherently disruptive.' Instead, the flux devices were distributed throughout the compartment, in separate

enclosures for mounting and lead protection, to characterize the heat flux field in which the structural components were immersed and to provide an extensive basis for model comparison. The model, with its accuracy determined, can then be used to predict the heat flux that falls on the surface of structural

Twelve Schmidt-Boeher total heat flux gauges, two ellipsoidal radiometers, and five slug calorimeters were used to measure the total and radiative heat fluxes at various locations throughout the test compartment.

A Schmidt-Boelter flux gauge is a calibrated, water-cooled thermopile whose sensor surface face

temperamre is uniform and close to that of the cooling water used. These qualities make it preferable to flux gauges of the Gardon design (metal foil sensor with a single central thermocouple) in handling mixed

convective and radiative heat fluxes. Schmidt-Boelter gauges have a field of view of 180 degrees and

record the sum of the radiant and convective fluxes (total heat flux) to the sensor surface. The gauges were rated at 150 kW/m". Schmidt-Boeher sensors have a time response of approximately 0.1 s to 0.2 s.

The convective component of the total heat flux is enhanced by the fact that the sensor face of the gauge

remains near the water coolant temperature (whereas the surface temperature of the objects in the

enclosure is rising considerably). Such a gauge thus provides a "cold wall" heat flux. The FDS model

can predict the convective component on this basis, so this is not a drawback to the goal of model-versusexperiment comparisons (see caveat noted below). Here, the cooling water was elevated in temperature

to preclude water vapor condensation on the gauge face. Typically, the gauge temperature was in the

70 °C to 80 °C range though it exceeded this upper limit in the longer experiments. This was one source of uncertainty in the flux gauge readings, as discussed below.

The typical heat flux gauge is a 1 .6 cm diameter by 2.5 cm long water-cooled metal cylinder with a pair of water leads,

electrical leads and. in some cases, a gas purge lead, as well. Inserting the gauge into a structural element alters the local temperature. The bundle of leads, with its necessary insulation, cannot run along the structural element without blocking heat

transfer to the element's surface. Even if the leads run normal to the element surface, they still present significant shadows,

blocking radiative transfer.

components.

5.2 DESCRIPTION OF THE HEAT FLUX MEASUREMENT DEVICES

NISTNCSTAR 1-5B, WTC Investigation 41
Chapter 5

The relatively cold sensor surface of a Schmidt-Boelter gauge is potentially a drawback in another

way—it is a natural deposition site for fire-generated soot, pushed toward it by thennophoresis.

Preliminary experiments were done with this type of gauge suspended above a 0.4 m diameter pool fire. Two such gauges were mounted side-by-side in a plate that provided both a means of covering one of the

gauges and of blowing any soot deposition off of the surface of the other gauge (by means of a nearly

tangential blast of high pressure nitrogen). Experiments were done as a function of height above the pool

surface and of fuel type (see Chapter 3). The heptane/toluene fuel mixture provided a particularly

challenging soot deposition environment, especially one diameter above the pool surface; the indicated

flux level from a clean gauge would immediately decay downward as soot deposited at a rapid rate. A

nitrogen purge was found to be capable of restoring this flux reading to the same level as the adjacent

gauge read the moment its cover (which kept its face clean) was removed. The same purge flow and

physical configuration were used for all of the total flux gauges in the hot upper layer for the tests reported here, (see Fig. 5-1 and also below for gauge locations.) The purge duty cycle was varied during

the experiments. For most tests it was 3 s of purging every 120 s; in the early experiments it was as frequent as every 30 s. The gauges showed no sign of signal decay due to soot accumulation in any test. Post-test inspection showed a very light soot layer on the sensor face that was wiped clean before being

used in any subsequent test.

Radiometers measure only the radiation portion of the incident heat flux on a surface in which they are imbedded. Each consists of a Schmidt-Boelter sensor at the base of a shallow cavity that is ellipsoidal in

shape. An opening at one end accepts radiation incident over a 1 60 degree wide field of view. All

radiation incident on the approximately 3 mm diameter hole at the open end of the cavity is absorbed by

the sensor at the other end. Soot deposition on the reflective cavity walls could modify this behavior, so

the cavity was purged with a weak nitrogen flow to keep soot from entering. The temperature of this gas

purge proved to be difficult to fully control in some of the experiments, and this was a source of

uncertainty in the radiometer readings (see below). The radiometers were also cooled with heated water

in the same temperature range as above, though given the nitrogen purge, heating was unnecessary to prevent water vapor condensation in this case. The gauges were rated at 150 kW/m'.

Medtherm Inc. supplied all of the flux gauges and radiometers. The total flux gauges were re-calibrated at NIST using 75 °C cooling water. The radiometers could not be calibrated with the NIST radiant source due to its partially focused nature, so the factory calibration was used with the uncertainty estimated as

5 percent.

The slug calorimeters used here were an adaptation of a technique that has long been used to infer heat

fluxes (e.g., in ASTM E 457 [2002]). Here, the slugs were 18 kg (40 lb) cylinders of nickel 200

(> 99 percent nickel), 10.2 cm (4.00 in.) in diameter by 25.4 cm (10.0 in.) long (see Fig. 5-2). The large mass was chosen to extend the useful measurement time since the slugs must lag the local temperature in

the technique used here. Nickel was chosen for its high temperature inertness and relatively high thermal

conductivity. The cylinders were insulated on the ends (1.3 cm of ceramic wool insulation). The cylinder

surface can be thought of as the sensor. It was coated with a black paint designed for high temperature

applications (Medtherm flat black) with a known absorptivity of 0.92. The cylinders were suspended

from the compartment ceiling by means of 1 .3 cm diameter threaded rods to minimize shadowing of the

cylindrical surface. These rods attached to brackets on each cylinder end; each bracket was in turn attached to the end of the cylinder only via a pair of '/4-20 stainless steel screws to minimize thermal

conduction to/from the bracket to the cylinder. The cylinders were oriented with their axes parallel to the

42 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

long axes of the trusses. In this sense, they averaged their incident flux distribution in much the same manner as does a unit length of the longitudinal element of a truss in the compartment." Note, however,

that they do not necessarily see the same fluxes as the trusses because the radiation field in the

compartment is not uniform.

'iiiimiliilll^^ i n-n iTtunrni

Source: NIST

Figure 5-1. Flux gauge holder showing gauge sensor surface (black disc) and soot

purge tube. Surface of gauge holder and gauge sensor surface were set flush

with the surrounding Marinite cover on any monitored surface.

Clearance hole for

Dimension in cm I

Figure 5-2. Schematic of slug calorimeter.

^ Being more massive and having a lesser surface area per unit length than an element of the truss, they heat up more slowly.

NISTNCSTAR 1-5B, WTC Investigation 43
Chapter 5

A metal cylinder exposed to a constant heat flux on its cylindrical surface necessarily has a temperature

gradient between the outer surface and the cylinder center, even as its average temperature goes up at a steady rate in response to the flux (Carslaw and Jaeger 1959). However, the average temperature of the

cylinder is very near the mass mean radius of 0.707 (radius). Thus, all cylinder temperature

measurements were done at this radius. Since the slugs used here were placed into a complex, threedimensional heat flux environment and nickel has a finite thermal conductivity, temperature variations were anticipated. ANSYS solutions for the internal temperature field resulting from a heat flux field with

a 2:1 asymmetry in the flux on opposite sides of the cylinder indicated an approximately 20 °C difference

in temperature around the mass mean radius. Thus, four thermocouples (3 mm diameter stainless steel sheathed, type K) were embedded in each cylinder (two from each end, through the ends) as shown in Fig. 5-2. The time-varying average heat flux over the cylindrical surface of each slug was calculated

from the instantaneous average of the four thermocouple signals using a heat balance as discussed below.

Since a slug remains below the local gas temperature, it too is a thennophoretic target for soot deposition,

at least until the slug surface is hot enough to oxidize soot (see below). The same can be said of the

uninsulated steel trusses in the fire compartment—since they have a greater surface area per unit volume,

they heat faster, gathering somewhat less soot. The point is that soot deposition is, in fact, a natural part

of exposure of structural components to a fire and it potentially affects the heat flux reaching the surface

of the element on which it deposits. Soot has an absorptivity of 0.95 or greater, enhancing radiation

absorption. The slugs started the first test at 0.92 absorptivity due to their black paint, and though it was

wiped off after each test, they probably retained enough to stay very near 0.95 even at the start of the next

three tests (the last two tests are discussed below). Thus, their radiation absorption was little affected by

soot deposition. It can probably be safely assumed that the bare steel tmsses acquired enough soot on

their surfaces to quickly become highly absorbent early in any of the tests. There is a second potential

effect of soot absorption that could be more of a problem to heat transfer analysis: a sufficient layer of

soot becomes a significant thermal insulator on its own (this is how it affects heat flux sensor readings).

All of the slugs were inspected at the end of each test for soot deposition. When present, it was estimated

to be 0.1 mm to 0.2 mm thick, which was not expected to have an appreciable effect on the heat transfer

rate to the slug (or to the uninsulated trusses, which could have had comparable build-ups). The two

tests using toluene/heptane mixtures yielded much sootier fires, however. At the end of those

experiments the slugs had a layer estimated at 0.5 mm, which possibly had some retarding effect on heat

transfer to these units and to the bare steel trusses used in these tests for late times in the experiments.

In the last two experiments, which were extended in duration, the slugs reached a sufficient temperature

to bum off the soot (estimated at > 600 °C in the oxidizing environment of the upper layer) and the

organic part of the black paint (they were not recoated). As a result, after each of these experiments, the

slug surfaces were dark gray, not black. At the beginning of each test, soot deposition would have rapidly

blackened them, as discussed above.

It is not possible to be more quantitative about this issue since the thermal properties of these very thin soot layers are not known. The layers appeared to have a complex stRicture, making quantification even more difficult. There appeared to be a denser inner layer covered by an outer layer that included a large proportion of agglomerated soot particles, which lowered the density of this outer layer.

44 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

5.2.1 Location of Heat Flux Instrumentation

Heat flux measurement devices were placed in the ceiling, in the floor, in locations next to a test column

and in a special "flux measurement station" approximately above the fire. The exact locations are indicated below.

The heat flux measurement station, which incorporated two total heat flux gauges (1.6 cm diameter, with

nitrogen purge to prevent soot deposition, rated to 150 kW/m") and two ellipsoidal radiometers discussed

above, was placed at the mid-height of the trusses and just downstream of the downstream end of the fu-e pan (see Figs. 5-3 and 5-4 for the location). The station was made of a 310 stainless steel cylindrical

shell with an outer diameter of 15.2 cm. a length of 50.8 cm, and a thickness of 1.3 cm. One flux gauge

and one radiometer were mounted flush on the top of the tube facing the ceihng of the compartment,

while the pair on the bottom faced straight down. A schematic cross section of the station is shown in Fig. 5-5; not shown is the ceramic wool insulation that filled the interstices inside the shell. The station was anchored on an L-shaped steel box beam support, 10.2 cm by 10.2 cm square, and 0.3 cm thick (see

Fig. 5-6) insulated with Marinite up to 3.8 cm in thickness. Output signal and thermocouple wires from

the gauges, cooling water to the gauges, return water lines, and nitrogen purge lines were bundled

together (with all the electrical lead wires placed in the core of the bundle), wrapped with ceramic wool,

and run from the station through the L-shaped insulated steel support to the outside of the test compartment. This arrangement prov ed adequate to protect the lead wires until the last two extended

duration tests. Thermocouple lead wires, which carried information on the water temperature in the flux

gauge interiors, began to short during these tests as the shell of the flux measurement station reached

750 °C. Flux gauge signal wires also failed later in the test.

Total heat flux gauges (6 mm diameter, rated at 100 kW/m") were mounted 15 cm off the floor flush

within stainless steel blocks on top of a 10.2 cm by 10.2 cm square, 0.3 cm thick steel box beam insulated

with 2.54 cm thick Marinite (flush with the top of the gauge-holding blocks and the Marinite). Signal

wires from the gauges and the cooling water lines were placed inside the insulated steel box beam in a manner similar to the station support described above. Since soot deposition was negligible at these two

locations, the floor flux gauges were not purged with nitrogen. The locations of the two gauges,

designated as HFl and HF2. are shown in Figs. 5-3 and 5^. Note that they have a large view factor with

respect to the fire plume.

Two larger total heat flux gauges (1.6 cm diameter, rated at 150 kWW), one facing the north-side wall

and one facing the fire pan, were also mounted within stainless steel blocks on a similar insulated steel support box beam adjacent to the upper part of the structural column element under test. Two other total flux gauges (0.6 cm diameter, rated at 100 kW/m') with the same orientation as the upper ones, were mounted on an L-shaped insulated steel support adjacent to the lower part of the column. Only the upper

two of these four gauges were purged with nitrogen to prevent soot accumulation. The locations of these

four gauges are shown in Figure 5-7 (designated as HCoHW, HCoHF, HCoLW, and HcoLF. Table 5-1

hsts the location and the orientation of the active surface of the heat flux gauges used in the experiment.

NISTNCSTAR 7-56, WTC Investigation 45
Chapter 5

NISTNCSTAR 1-5B, WTC Investigation 46
Heat Flux Measurements

I: 1TSL3

Fire par.

TSO-.ffTr'i H

•-iOe 7BNORTH WALL

rllM.2

TSL5 TC tree

16 5

SOUTH WALL

# Slug Station radiometer

Ceiling flux gage Station flux gage I Roor flux gage Column flux gage TC tree * Fuel spray nozzle

Note: All units in centimeters. An expanded view of circled area is given in Figure 5-5. Figure 5-4. Top view of test compartment.

H2RU H2FU

Figure 5-5. Expanded sisde view of circled area in Figure 5-4.

NISTNCSTAR 7-56, WTC Investigation 47
Chapter 5

1

- 10

station

- 50.8 I

H2RU

Ll5

30 5 _!_

Cross section view @ station location

' = station flux gage IP station radiometer

Figure 5-6. Locations of flux gauges and radiometers in the measurement station.

360 H Column flux gage

Figure 5-7. Schematic showing the locations of column heat flux gauges.

48 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

Table 5-1. Heat flux gauge location.

Designation Channel Type Location" Orientation

Hf 1 346 Total Near Floor: (500 cm, 6 cm, 13 cm) Upward

HF2 347 Total Near Floor: (550 cm, 3 cm, 13 cm) Upward

HCoHW 343 Total Column: (297 cm, 85cm, 327cm) Northward

HCoHF 342 Total Column: (304 cm, 77cm, 346 cm) Eastward

HCoLW 345 Total Column: (297 cm, 85cm, 102 cm) Northward

HCoLF 344 Total Column: (304 cm, 77cm, 92 cm) Eastward

HIFU 338 Total Station: (458 cm, 56 cm, 330 cm) Upward

HIRU 339 Radiometer Station: (458 cm, 44 cm, 330 cm) Upward

H2FD 340 Total Station: (458 cm, 23 cm, 315 cm) Downward

H2RD 341 Radiometer Station: (458 cm, 35 cm. 315 cm) Downward

HCel 60 Total Celling (349 cm, 28 cm , 382 cm) Downward

HCe2 61 Total Ceiling (399 cm, 28 cm , 382 cm) Downward

HCe3 62 Total Ceiling (474 cm, 28 cm , 382 cm) Downward

HCe4 63 Total Ceiling (650 cm, 28 cm , 382 cm) Downward

TSLl 299-302 Slug Upper Layer: (102 cm, 53 cm, 329 cm) Not directional TSL2 303-306 Slug Upper Layer: (212 cm, 53 cm, 291 cm) Not directional TSL3 307-310 Slug Upper Layer: (337 cm, 53 cm, 329 cm) Not directional TSL4 311-314 Slug Upper Layer: (397 cm, 53 cm, 329 cm) Not directional TSL5 315-318 Slug Upper Layer: (550 cm. 53 cm, 329 cm) Not directional

a. Also listed in Appendix C.

Four flux gauges ( 1 .6 cm diameter; total flux; rated at 150 kW/m") were mounted flush with the ceiling

surface near the compartment centerline. They were positioned at locations in the east half of the

compartment, beginning approximately above the upstream edge of the fire pan and continuing toward

the east wall (see HCel through HCe4 in Fig. 5^). These gauges were mounted within stainless steel block holders that also provided soot purge nozzles.

The positions of the slug calorimeters in the test compartment, designated as TSLl, TSL2, TSL3, TSL4,

and TSL5 are shown in Figs. 5-3 and 5^. They were arrayed in a line 53 cm to the north of the

compartment centerline but covering much of the east-west length of the compartment. Note that four of

the slugs were at the mid-height of the trusses and one, denoted TSL2, was at the height of the lowest

element of the trusses.

All of the flux measuring devices have diverse objects in their fields of view. The actual field of view

depends, of course, on the radiation path length in the fire atmosphere and that, in turn, depends on the

soot concentration, mainly in the hot upper layer (which varied with fuel type). The FDS model attempts

to take all of this "clutter" into account in characterizing the radiation field, though it cannot do so with

sufficient spatial resolution to precisely describe the shapes of some of the smaller objects.

NISTNCSTAR 1-5B, WTC Investigation 49
Chapter 5

5.2.2 Coolant System for Heat Flux Gauges

Figure 5-8 shows a schematic of the closed flow system used to circulate hot water through the gauges

and radiometers. Three of these systems were used to serve all of the heat flux gauges. The water was

heated inside a 1 14 L electric water heater tank. The tank contained two 4.5 kW heating components

controlled by thermostats to heat the water to an operating temperature between 70 °C and 80 °C. The

tank had a maximum operating pressure of 1 ,03 1 kPa, which was also the pressure limit of the gauges and

radiometers. Water was pumped from the tank using a self-priming, bronze gear pump coupled to a 373 W (I/2 hp) motor. The pump was rated for fluids at temperatures up to 99 °C, and had an adjustable

pressure relief valve set to approximately 687 kPa. The water was pumped through a 50 |im polyester

string wound filter to remove any contaminants that could clog the lines. The filter was contained in a

stainless steel housing with a temperature rating of 135 °C and a pressure rating of 1,718 kPa. An

aluminum manifold distributed the main flow into five separate flows, which were regulated using five

flow meters with a pressure rating of 1 ,7 1 8 kPa and a temperature rating of 12 1 °C. The flows were kept

above 0.9 L/min, the gauge manufacturer's recommended minimum flow. Copper lines were used to transport the heated water to individual flux gauges and radiometers and back to the tank where it was

reheated to the operating temperature. The copper lines were insulated using foam plastic and fiberglass

insulation to prevent heat loss from the lines to the surroundings.

return lines from radiometers

and flux gages

water lines to radiometers

and flux gages

water heater water pump filter

Figure 5-8. Cooling water system for radiometers and flux gauges.

50 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

5.3 DATA REDUCTION FOR THE SLUG CALORIMETERS

As discussed above, the output data from the slug calorimeters consisted of four individual measures of

the temperature within each slug, in all cases on the mass mean radius.^ For simplicity and tractability,

the data analysis discussed below assumes that the temperature within a slug is uniform, and thus, the

analysis uses the average of these measurements. As was noted above, there is necessarily a radial temperature gradient within a cylinder subjected to a constant (and spatially uniform) heat flux on its surface but that can be factored out, to a first approximation, by focusing on the temperature at the mass mean radius within the cylinder. In general, there can be asymmetrical gradients, as well, due to the non- uniform heat flux field impinging on the cylinder surface. Higher thermal conductivity minimizes all internal temperature differences; thus, nickel, which can withstand exposure to high temperatures, was

chosen as the slug material. Its thermal conductivity is about one fourth that of copper. Temperature

differences among the four readings did always exist since the flux on the slug surface was not spatially

uniform. In general, the differences increased with increases in the average flux to the calorimeter

surface, but they were always less than 5 °C. The impHcations of these differences are discussed below in the section on measurement uncertainties.

Given the assumption of a uniform temperature within a cylindrical slug subjected to a spatially uniform

heat flux on its cylindrical surface (adiabatic end surfaces), a heat balance leads to the following:

P^y,CAT)^ = q,,,A^.-A^s^.c7 (5-1)

at

where Pc is the density of the cylinder material, Vc is the cylinder volume, Cc(T) is the temperature

dependent specific heat of the cylinder material, T is the temperature, qtot is the total heat flux (radiation

plus convection) absorbed by unit area of the cylinder surface, Ac is the area of the cylinder (exclusive of

the ends), 8c is the emissivity of the cylinder surface and a is the Stefan-Boltzmann radiation constant.

The temperature dependent heat capacity of nickel 200 is describable by a polynomial, and all the other

parameters in the above equation are known. The equation can be solved for the item of interest, qtot:

q„„-(p^DJ4)CAT)^ +^,aT' (5-2)

Here (Dc/4), where Dc is the cylinder diameter, has been substituted for (Vc/Ac). This is the form used in the data reduction. Since the average temperature (as a function of time) is used, the qtot inferred is the

average heat flux over the cylinder surface.

' In a few cases, one of the thermocouples in a slug failed to give useable data due to poor contact in the ceramic connectors on

top of the compartment. The analyzed information was then based on the average of the remaining three thermocouples in that

slug.

NISTNCSTAR 7-56, WTC 51 Investigation
Chapter 5

The actual application of this equation to a data file was done with SigmaPlot software, using its transform capability. Thus, at each time step the four temperatures from within the slug were first averaged, then the average temperature history was fitted with a running cubic spline transform which

also calculated the time-derivative of this spline fit. The derivative had high frequency noise, which was

removed with a low pass filter transform with the bandpass set so that it just began to visibly affect the

starting transient at the beginning of a test.^ The heat capacity term was then calculated and all terms were combined in accord with the above equation to yield qtot at a given time.

Nickel is a magnetic material with a Curie point of 358 °C at which it becomes non-magnetic (Toulakian

and Buyco 1970). This is an order-disorder transition that causes an abrapt step downward in the heat

capacity. In principle, the data reduction process above can handle this transition. Its ability to do so was

tested successfully using temperature data generated from a numerical solution of Eq. 5-1; the data

reduction procedure recovered the imposed incident heat flux history right through the Curie point and

beyond to within ± 1 percent except within a few degrees of the Curie point. Localized noise in the

computed flux near the Curie point occurred upon processing the data, but this disappeared above the

Curie point.

5.4 UNCERTAINTIES IN THE HEAT FLUX MEASUREMENTS

For the total heat flux gauges, the output voltage was measured with the data acquisition system and

converted via a calibration to a heat flux reading. In the absence of soot accumulation effects (due to the use of a nitrogen purge to remove soot),^ the sources of uncertainty are in the voltage reading process and

in the calibration. The latter, in turn, has contributions from uncertainties in the calibration process itself and in the applicability of the calibration done at 75 °C to a gauge temperature that is somewhat different

from this value. The uncertainty due to the coolant water temperature variation (estimated as less than ±

0.5 kW/m", independent of incident heat flux, for the worst case) easily dominates over that due to voltage measurement uncertainty. However, the dominant uncertainty with this type of gauge is in the

calibration process itself. A recent round robin test (Pitts 2003) of one such gauge at five fire facihties

internationally (using a variety of calibration methods) indicated a standard deviation of about ± 3 percent, or ± 3 kW/m" at a flux of 100 kW/m".

This type of heat flux gauge is calibrated with a purely radiant source, and the round robin mentioned

above used a variety of such sources. In the measurements reported here, a mix of convective and

radiative heating is expected. At present, there is no widely used technique for calibrating this type of

heat flux gauge with a purely convective heat input so as to ascertain how differently it may respond to pure convection. The complexities attendant with convective heating imply a larger uncertainty. For all of the measurements, the dominant portion of the total flux to the gauges is radiation. However, in mixed

radiation plus convection situations, the expanded uncertainty in the total heat flux is expected to be

higher than for pure radiation; it is estimated to be at least ±10 percent (with an expansion factor of two).

Carslaw and Jaeger { 1 959) show that the time response for a nickel slug of this size to reach a steady state rate of increase to a constant surface flux is 20 s to 30 s. Thus, any variations appreciably faster than this are not of interest and are filtered out.

^ The nitrogen soot purge greatly altered the fiux gauge reading while it was on, but the gauge recovered from this in less than

1 s. The purge also briefly disrupted temperature readings in its neighborhood, but the total volume of purge gas emitted per purge cycle from all the purge tubes was less than 0.3 percent of the gas volume of the upper layer.

52 NISTNCSTAR 1-5B. WTC Investigation
Heat Flux Measurements

The soot purge nozzle on each of the total flux gauges in the upper layer blocks a portion of its field of

view by superimposing an end-on view of a piece of tubing. The fraction of the field of view that is blocked (altered) by this is quite small, however, estimated at 0.6 percent.

The radiometer measurements have two other sources of uncertainty: a less than hemispherical field of

view and a purge gas whose temperature could not be kept fully under control since the lead tubing was

heated by the fire impinging on the flux measurement station. Both of these were systematic, rather than

random, sources of error. As was noted previously, the radiometers had a 160 degree, not a 180 degree,

field of view, presumably due to limitations in machining the ellipsoidal cavity. This translates to about

97 percent of the available radiation (assuming a uniform radiation field). Countering this was a conx ective input to the sensor due to an increasingly hot nitrogen purge caused mainly by the heating of

the flux station shell containing the gauges. Ideally, the nitrogen purge would have stayed at the same

temperamre as the sensor itself (approximately 75 °C) so that it would have transferred no heat to it. However, in Test 5, where the flux station shell became very hot, the nitrogen purge reached about

250 °C to 300 °C by the time the gauge signal leads shorted (approximately 2,000 s into the test). The

heat transfer coefficient between the purge gas and the sensor within the elhpsoidal cavity was estimated

to be that for flow over a sphere of the same radius as the cavity. With this, the convective input to the

radiometer sensor was estimated as less than 4 kW/m" for Test 5. The two radiometers gave no useful

data in Test 6. This convective effect was minimal (< 0.2 kW/m') in the first four tests that were much

shorter in duration.

The radiometers could not be calibrated at NIST, so the manufacturer's calibration was used. In the

roundrobin study of a single total flux gauge mentioned above, the manufacturer's calibration fell well

within the range of variation of the other lab-to-lab variations. It is probable that the same calibration

expanded uncertainty (± 6 percent) applies here as well.

The total heat flux levels derived from the temperature measurements in the slug calorimeters used other

implicitly or explicitly measured quantities: slug diameter, slug density and slug heat capacity as a function of temperature. A formal analysis (Holman 1970) of the propagation of uncertainties from each

of these sources shows temperature uncertainty to contribute 90 percent of the total, while heat capacity

contributes most of the remainder. This is for a case where the standard deviation among the four

thermocouple measurements was ±4.5 °C, among the largest seen. In the first three tests, the temperature

variation was less than half of this value. The corresponding flux uncertainty for the 4.5 °C standard

deviation was ± 6 kW/m~ at a time when the calculated total flux was above 80 kW/m".

There is a systematic error in the slug calorimeter analysis due to the implicit assumption that the surface

temperature of the slug equals the average temperature of the slug. The calculated re-radiation term (last

term in Eq. 5-2) is thus less than in reality, since the surface must be hotter than the average bulk

temperature. If the average surface temperature was in reality 25 °C above the average slug temperature

(a value suggested by ANSYS software solutions for the transient heat conduction problem) when the

latter is at 350 °C , then the radiation term was underestimated by 1 .5 kW/m"^ , and the total flux to the

slug was also underestimated by this amount. This error was appreciably less in the first three tests in which the average slug temperamre by the end of the test was much less. Another systematic error is the

neglect of heat transfer through the ends of the cylinder (conduction through the insulation and through

the support bolts). These two leakage paths (inward) are estimated to be comparable in size and to result

in an overestimate by about 4 percent when the calculated average heat flux is 25 kW/m^. This source of

error helped to counter the error induced by underestimating the cylinder surface temperature. Overall,

NISTNCSTAR 1-5B, WTC Investigation 53
Chapter 5

the slug data have an expanded uncertainty for the 3 MW tests (Tests 4, 5, and 6) estimated to be

approximately 16 percent to 20 percent; the uncertainty for the 2 MW Tests (1 tlirough 3) is approximately half this value.

The uncertainty estimates for the three types of heat flux measuring devices are summarized in Table 5-2.

Table 5-2. Estimated certainties in heat flux measurements.

Measurement Device Estimate of Expanded Uncertainty

Total heat flux gauge ± 10%

Radiometer ±6%

Slug calorimeter ± 16% to 20%

5.5 DISCUSSION OF HEAT FLUX MEASUREMENTS

Initially, all wall and other solid object temperatures are cold (at room temperature) as is the gas in the room. The fire plume mixes in, fills and heats the space above the soffit level of the exhaust openings in a time on the order of the volume of this space divided by the plume volumetric flow (5 s to 10 s). As the

gas temperature there rises, the ceiling (and upper wall) surface temperatures in this volume follow fairly

closely due to the insulating character of the Marinite lining. The subsequent rise in ceiling and wall

temperatures occurs on a time scale on the order of 1 ,000 s, representative of the thermal relaxation time

of the 2.54 cm thick Marinite. Even though the ceiling temperatures in Test 5, for exainple, rose only

about 100 °C in the last 2,000 s of the test, this results in nearly 50 percent increase in radiative flux from

the Marinite ceiling. In this enclosure, all of the surfaces are coupled by these rising radiative fluxes

though the rates of rise will not be equal. An additional comphcation is the optical thickness of the

smoke layer. Different flux gauges see different surfaces and smoke volumes and have differing view

factors with respect to the fire plume. They also have differing contributions from convective heating.

With this temperature history in mind, consider first a typical set of measurements from one test. For this we chose Test 5, which was a 3 MW heptane fire with a 50 min duration.

The two floor gauges (total flux) in Fig. 5-9 have a relatively simple behavior to interpret since

convective effects should be minimal (though not zero) in the cool lower gas layer. Tiine zero was the

start of the fire; the rapid decay at about 3,050 s (here and in the other figures below) was due to the

abrupt termination of the fire. As soon as the fire was ignited, these two gauges, with their wide views of

the fire plume, jumped to substantial readings as a result of direct plume radiation. Subsequently, both

climbed monotonically and substantially, leveling only after about 2,500 s; the difference between the two

gauge readings, due principally to their differing distances from the fire plume, grew only slightly. Both

had appreciable fluctuation levels, reflecting the highly transient nature of the radiative processes from

the turbulent fire plume. Nearly the entire subsequent climb in flux above the initial level was attributable

to radiation, froin the room surfaces and from the soot in the hot upper gas layer. Some fraction was due,

however, to convection as the lower layer became warmer.

54 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

0 r 1 I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I I FTi"' 0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure 5-9. Total heat flux recorded by two flux gauges 15 cm above the floor,

just downstream of the fire pan in Test 5.

The convective component is difficult to estimate accurately for all of the total flux gauges. All of these

gauges respond to a combination of incident radiation and convection:

^,o, = ^raJ + ^ C^gas ' Tgaug^ ( 5-3 )

where qwiis the total heat flux to which the flux gauge responds, qrad is the radiative component of the flux

and the second term on the right in Eq. 5-3 is the convective flux seen by the gauge. Note that, as mentioned previously, it is dependent on the gauge surface temperature, Tgauge, which is typically lower

than that of the surface in which the gauge is embedded. Thus the convective component of the gauge

signal is somewhat exaggerated, but this effect is accounted for in the FDS model comparison (see

Chapter 10). The difficultly in estimating the convective contribution, both here and in the FDS model, is the proper choice of the value of the heat transfer coefficient, h, since the flows are not simple. It is necessary to select a situation that resembles that in which the gauge is placed and for which a heat

transfer correlation is available.

For the floor flux gauges above, one can estimate^ the convective flux component of the total flux to these

gauges using the local gas temperature and the assumption of buoyant convection locally as < 2 kW/m".

Here, the buoyant convection to the cooled gauge surface (Williams and Smith 1978) exceeds the

estimated forced convection for turbulent flow over a flat plate (Roshenow and Choi 1961). Since these

estimates are based on the gas temperatures achieved after an extended test period (approximately

300 °C). the estimates for the shorter experiments (Tests 1 through 4) will be less.

* Convective fluxes are estimated here and below from one or more of the heat transfer correlations that appear to most closely

fit the specific configuration under consideration. Any velocity estimates that go into these calculations are in the range

suggested by the FDS calculations. Gas temperatures come from the nearest aspirated thermocouple.

NISTNCSTAR 1-5B, WTC Investigation 55
Chapter 5

Figure 5-10 shows the total heat fluxes seen by the four gauges set flush into the ceiling in Test 5. The

strong spikes (in this case eveiy 120 s) were due to the soot purge flow blowing across the gauge faces. The fluctuations in the data were again due to the turbulence in the fire plume whose eddies had an

influence even at a substantial distance from the plume.*^ The fluctuations were greatest for heat flux

gauge HCe3, which was just downstream of the fire pan and most sensitive to plume position variations.

The ceiling heat fluxes started at moderately low levels (especially HCe4), but all immediately started climbing monotonically. Only by the end of the more than 3,000 s test did the flux seen by these gauges

appear to be leveling off fully. The overall upward trend was again due to the heating of the upper layer

and the compartment surfaces. Here, the convective heat transfer component to the "cold" flux gauges was secondary to radiation; it was estimated, on the assumption that the plume produces a stagnation

point flow (Kays 1966) at some distance from these gauges to have been in the 5 kW/m" to 10 kW/m"

range by the end of the test when the upper layer gases were at their hottest. The assumption of a turbulent flow over a flat plate (Roshenow and Choi 1961) gives comparable numbers. In the shorter tests (or early in the long tests), where the gas temperature was somewhat lower, Eq. 5-3 indicates that the

convective component would be less.

200

I

I I I I

I I I I I I

I I I I

I

I I I Ir

I

I

I

I 11 I I

It I

I

I jl

I

I I I I

I

I I I I

Time (s)

Figure 5-10. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 5. The periodic spikes are due to the soot purge operation.

^ The eddies cause not only convective heat flux variations but also radiative flux variations.

56 NISTNCSTAR 1-5B. WTC Investigation
Heat Flux Measurements

Figure 5-1 1 shows the total heat fluxes recorded by the four flux gauges in structures adjacent to the

vertical column in Test 5. The periodic spikes were due to the soot purge operation on the two higher

gauges. As with the preceding results, the trend was monotonically upward with time, approaching an asymptotic value by the end of the test. The underlying cause of this behavior was the same progressive

heating of the compartment surfaces. The two gauges facing the north compartment wall (HCoHW and

HCoLW) rather than the fire plume received the lowest fluxes, as one would expect.'^' The higher of

these two gauges had a good view of the wall (and, to some extent, the ceiling) in the upper layer region

where its surface would have radiated more strongly. In addition, this "cold" gauge received an estimated

10 kW/m' to 15 kW/m" as a convective input in the latter half of the test (assuming natural convection to a vertical plate as the dominant flow mode [Williams and Smith 1978]). A comparable convective input

applies to the other high gauge (HCoHF); this, plus its enhanced radiative input, explains the fact that it had the highest total flux readings of the column gauges. The convective inputs to the lower column flux

gauges are estimated, using the same assumptions, to have been in the 3 kW/m" to 5 kW/m' range. Again, these fluxes would have been lower earlier in this test or in the shorter tests since the gas

temperature was lower.

The total heat flux inferred from the average temperatures (vs. time) in the five slug calorimeters are shown in Figs. 5-12 through 5-16. Only Figs. 5-12 and 5-13 show the fiill test duration. The others

show just under 1 .000 s of inferred flux results; this is close enough to the long term flux levels that one

can discern the variations among the five slugs. This, in turn, indicates how the fluxes would vary along

the length of the horizontal steel trusses in the compartment. Slugs 1 and 2 yielded nearly identical flux

histories (slug 2 was 5 percent to 10 percent higher for the last 2,000 s of the test) and had the lowest flux

values of all the slugs, though for somewhat differing reasons. Slug 1 was the farthest from the fire plume and had the smallest radiative view factor with respect to it." Slug 2 was closer to the fire, thereby

receiving more plume radiation, but it was also lower, being at the height of the bottom element of the

trusses; it thus was in a lower temperature gas region and also would be expected to receive less soot

radiation on its lower surface.

The immediate readings on the order of 1 0 1< W/m" do not see the fire plume directly, so they are responding to fire plume

radiation reflected from the Marinite. " The slugs receive heat (radiation plus convection) only on their cylindrical surfaces, and the axis of each cylinder was parallel

to the trusses.

NISTNCSTAR 1-5B, WTC Investigation 57
Chapter 5

160

CME

X

LL

fo

CD X

o

I I I I II I I I

I

I I I I

I

I I I I I I I Q HCoHF

-0— HCoHW

HCoLF

--X--HCOLW

1500 2000 2500

Time (s)

4000

Figure 5-11. Total heat flux recorded by the four flux gauges in structures adjacent to

the vertical column in Test 5. The periodic spikes are due to the soot purge operation on

the two higher gauges.

3000 3500

Figure 5-12. Average total heat flux to the cylindrical surface of the

slug calorimeter 1 In Test 5.

58 NISTNCSTAR 1-5B. WTC Investigation
Heat Flux Measurements

-1 1 1 1—r-

"to Q.

CD

WTC Phase 1 Test #5

Average Total Heat Flux on Slug 2 Surface

Calculated from Average Temperature vs Time

-| 1 1 1

r 1

1 1

1 I

I I

1 I I 1 r- 500 1000 1500 2000 2500 —I<- 3000

Time (s)

Figure 5-13. Average total heat flux to the cylindrical surface of

slug calorimeter 2 in Test 5.

80

60

I

I

1 1

1

I

1 I

1 1

1

1 1 1 rSlug Calorimeter #3

Comparison of Data From Replicate Tests 5 and 6

1000

Time (s)

Figure 5-14. Comparison of the average total heat flux to the cylindrical

surface of slug calorimeter 3 in Tests 5 and 6.

NISTNCSTAR 1-5B, WTC Investigation 59
Chapter 5

80

r= 60

03

CD X

40

20

c

0) g

o

_c

-4—

»

O

ICDcn

TO

03><

* CDQ.

CO

Slug Calorimeter #4

Comparison of Data from Replicate Tests 5 and 6

A ^.

200 400 600 800 1000

Time (s)

Figure 5-15. Comparison of the average total heat flux to the

cylindrical surface of slug calorimeter 4 in Tests 5 and 6.

80

H= 60 -

CD

CD X

40 -

20

c

CD gq

_c

o

ICD

CD

L_

CD><

I

°

"cD Q.

C/D

-T 1 1 1 1I I r I 1 [ r [ I 1 I

1

[ I 1 1 I 1 I WTC Phase 1 Test #5

Average Total Heat Flux on Slug 5 Surface

Calculated from Average Temperature vs Time

1000

Time (s)

Figure 5-16. Comparison of the average total heat flux to the

cylindrical surface of slug calorimeter 5 in Tests 5 and 6.

60 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

The early transient in the slug fluxes was due to their time response and, to a lesser extent, the filtering of

the temperature gradient calculated during the data reduction. The residual noise in the data was partly a

real heat flux variation and partly a result of heat conduction processes in the slugs interacting with the

calculated average slug temperature. Figures 5-12 and 5-13 also show a weak spike at about 1,300 s that

is due to the abrupt change in the heat capacity of nickel at its Curie point. The spike is a result of the

data treatment and is not real.

Slugs 3 and 4 were both relatively close to the fire plume and, thus, received significantly (approximately

20 percent to 30 percent) higher heat fluxes than slugs 1 and 2. The fire plume tended to bend toward the

outlet side of the compartment, putting it somewhat closer to slug 4, which showed a slightly higher flux

(e.g., at 800 s. approximately 60 kW/m" vs. about 50 kW/m' for slug 3). The results from the slug

calorimeter measurements were consistent with the measurement of total heat flux in the upper layer. In

this sense, the slug data were useful in confirming the accuracy of the total heat flux gauge measurements.

Inspection of Fig. 5-16 shows that slug 5 began with a low flux reading (comparable to slug 2 but lower

than slug 3). then ended with a reading comparable to slug 4. This suggests the increasing importance

with time of radiation from the hot ceiling and walls, especially on the exit end of the compartment

towards which the bulk of the hot gas flow was moving. The maximum contribution of convection to these total heat fluxes is estimated to have been in the 4 kW/m" to 6 kW/m" range (Williams and Smith

1978; Roshenow and Choi 1961).

Finally, consider the results from the flux measurement station, shown in Fig. 5-17. There were two pairs

of flux gauges, each pair consisting of a radiometer and a total heat flux gauge, approximately side by

side, one pair looking directly up at the ceiling, and the other pair looking directly downward to one side

of the fire plume. Assuming that the components of each pair were exposed to the same radiative input,

the difference in their readings is attributable to convective heating of the total flux gauge. The total flux

looking downward was, as expected, the largest of the four readings, with an asymptotic value of about

90 kW/m'. The downward-looking radiometer failed at about 1,700 s. Prior to failure, the flux was

approximately 30 kW/m" below the total flux reading in the same direction. This implies a convective

component of 30 kW/m", which is difficult to rationalize. Using an available correlation for turbulent

convection to a cylinder (Eckert and Drake 1972), taking into account the difference in local heat transfer between the upstream and downstream convection coefficients, one can calculate an expected convective

flux of 8 kW/m' to 12 kW/m'. The real flow is, of course, more complex than is assumed by the available

correlation but it is difficult to see how the convection was evidently so enhanced. The problem is quite

comparable for the other pair of flux gauges (facing the ceiling). One can calculate an expected

convective contribution of 4 kW/m" to 6 kW/m", but the data indicate a convective contribution of about

20 kW/m". There are no checks on this issue elsewhere in the enclosure; it is possible that the convective

fluxes elsewhere are also greater than the calculated values reported above. On the other hand, it is also

possible that the radiometers do not read this complex radiation field accurately.

NISTNCSTAR 1-5B, WTC Investigation 61
Chapter 5

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure 5-17. Heat fluxes recorded by the radiometers and total flux gauges in the flux

measurement station during Test 5. The periodic spikes are due to the soot purge

operation.

5.5.1 Repeatability

Repeatability was examined in Tests 5 and 6, which both used 3 MW heptane fires within nominally

identical compartment set-ups. In general, the heat flux measurements tracked very close to each other

(differences approximately 5 percent or less) with the greatest departure being at slugs 3 and 4; see

Figs. 5-14 and 5-15. The source of these larger differences is not clear, but may have been related to average plume position.

It was intended that the average plume impingement position on the ceiling would be determined by an

array of thennocouples on the ceiling surface. These formed a cross, centered above the downstream

edge of the fire pan. The data proved to be insufficient to pin down the plume position in a definite manner. In Test 5, the mean impingement position was approximately 0.5 m downstream (toward flow

exit) of the downstream edge of the fire pan and more than 0.5 m south of the compartment centerline. In Test 6, the mean impingement was 0.5 m to 1.0 m downstream of the downstream edge of the fire pan

and approximately 0.5 m south of the compartment centerline. Thus, the plume did move appreciably in the replicate tests, possibly because of differing thermal warp patterns in the fire pan itself.

5.5.2 Effect of Fire Size

All of the above results were for a 3 MW heptane fire. In Test 1, the fire size for the same fuel was

2 MW. The effect of this 50 percent fire size reduction was considerable. At an equivalent time

(approximately 800 s, since Test 1 was much shorter than Test 5), upper gas layer and ceiling

62 NISTNCSTAR 1-5B. WTC Investigation
Heat Flux Measurements

temperatures were 200 °C to 300 °C less with the smaller fire (e.g., 500 °C vs. 700 °C).'- This, in turn, strongly decreased the total heat fluxes seen by the various measuring devices. For example. Fig. 5-18

shows the ceiling heat fluxes for gauge HCe2 (directly over the fire pan) for Tests 1 (2 MW) and

5 (3 MW). The 2 MW fire yielded a ceiling flux of less than half than that of the 3 MW fire. Similarly,

Fig. 5—19 shows a comparison of the results from slug 1 for the two fire sizes. Slug 1 was well away

from the fire in the upstream end of the compartment; as explained above, it responded mainly to the gas

and solid surface temperamres around it. All of the other flux measuring devices showed comparable

changes in their results (including those in the lower layer). The temperature difference easily explains

the greater than a factor of two heat flux differences in a situation like this where radiation dominated

the total heat fluxes. Even the relatively small difference in fire size between Tests 2 and 3 (2.4 MW

vs. 2 MW) led to substantial heat flux differences: e.g., 40 kW/m" vs. 25 kW/m" at 400 s for HCe2 and

20 kW/m" vs. 13 kW/m" at 380 s for slug 1. Thus, fire size has a potent influence on the heat flux

field impinging on structural components in the hot upper layer of a compartment and on surfaces

elsewhere in the compartment. This is consistent with previous results fi^om other laboratories

(Wakamatsu et al. 2003).

5.5.3 Effect of Soot

The effect of fuel soot yield was explored in switching from heptane to a mixture of heptane plus toluene.

The effect on measured heat fluxes can be seen by comparing results from Test 1 (heptane) and Test 3 (60 percent heptane/ 40 percent toluene), where the fire size was 2.0 MW for both (see Table 2-1). Soot

yield increased both the emissivity and the absorptivity in this layer. The former should increase heat

fluxes to the surfaces while the latter could trap heat in the gas and prevent ceiling and floor, for example,

from seeing each other radiatively. Inspection of the aspirated thermocouple records from the two tests showed minimal differences, which appear to be within the test repeatability. Comparison of the flux

readings from the various devices gives mixed results. For the slug calorimeters, only slug 4 differed

significantly with the heptane fire giving approximately 20 percent higher readings at this location close

to the fire plume. (This comparison is in the last half of the tests; the difference was greater than

30 percent at some earlier times.) The ceiling gauges all gave quite comparable results, in spite of some

being in the same area as slugs 3 and 4. The column gauges tended to read slightly higher (approximately

10 percent) for the heptane fire. The floor gauges read 10 percent to 1 5 percent higher for the heptane

fire. The largest difference was for the downward looking radiometer in the flux measurement station

above the fire: the flux from the heptane fire was 60 percent higher. Other station readings were higher

for the heptane fire, but appreciably less so. The overall trend for the flux gauges suggests that soot from

the toluene fire makes the fire plume opaque.

Equation 5-3 implies that the convective portion of the heat fluxes would have been reduced only about 30 percent, but the convective portion was, as discussed above, typically much less than radiation.

NISTNCSTAR 1-5B, WTC Investigation 63
Chapter 5

CME

100

80

—i—I—I—

I—I—i—I—I—I—r- Tests 1 and 5

1^

3 MW ' 1 «

I

'

I I, II I

I "

«

' Vi I

I

I "''V' w \

I "J

if 60

03

CD X

£ 40

CD g

'o

_c3 20 o

I0 100 200 300 400 500 600 700 800

Time (s)

Figure 5-18. Effect of fire size in Tests 1 and 5 (2 MW and 3 MW, respectively) on total

heat flux at ceiling gauge HCe2, directly above the fire pan.

60

X^ 50 H

CO0X

c

CD g

CDO)

CD

i_CD><

40 -

30 -

K 20-

10

Comparison of Total Heat Flux to Slug 1 for 2 MW and 3 MW Heptane Fires

200 400 600

Time (s)

800 1000

Figure 5-19. Effect of fire size in Tests 1 and 5 (2 MW and 3 MW, respectively) on the

recorded total heat flux at slug calorimeter 1, upstream of the fire plume.

64 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

5.6 REFERENCES

ASTM International. 2002. ASTM E 457-96 (2002), Standard Test Methodfor Measuring HeatTransfer Rate Using a Thermal Capacitance (Shig) Calorimeter. West Conshohocken, PA.

Carslaw. H.. and J. Jaeger. 1959. Conduction ofHeat in Solids. 2nd edition, p. 203. Oxford University

Press, London.

Eckert. E.. and R. Drake. 1972. Analysis ofHeat and Mass Transfer, p. 406, McGraw-Hill NY.

Holman. J. 1970. Experimental Methodsfor Engineers, 3rd edition, p. 44. McGraw-Hill, NY.

Kays. W. 1966. Convective Heat and Mass Transfer, pp. 209ff, McGraw-Hill. NY.

Pitts. W. 2003. NIST Building and Fire Research Laboratory, personal communication. April.

Roshenow. W.. and H. Choi. 1961. Heat, Mass and Momentum Transfer, p. 198, Prentice-Hall.

Engelwood Clifffs, NJ.

Toulakian. Y.. and E. Buyco. 1970. Specific Heat. Metallic Elements and Alloys. Volume 4 of

Thermophysical Properties of Matter. Plenum Press, NY.

Wakamatsu. T.. Y. Hasemi, K. Kagiya. and K. Kamikawa. 2003. Proceedings ofthe Seventh

International Symposium on Eire Safety Science, pp. 1 099- 1110.

Williams, G., and K. Smith. 1978. Marks Mechanical Engineer's Handbook, Eighth Edition, p. 4-68,

McGraw-Hill, NY.

NISTNCSTAR 1-5B. WTC Investigation 65
Chapter 5

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66 NISTNCSTAR 1-5B, WTC Investigation
Chapter 6

Smoke, CO2, CO, and O2 Concentrations

6.1 INSTRUMENTATION

Smoke/gas samples were extracted from the compartment at two locations for subsequent analysis of the

volume (or mass) fraction of the smoke aerosol and three gas species CO2, CO, and O2. The first location

labeled with the prefix "Sm" (see Fig. 6-1) was in the northwestern quadrant of the enclosure near the

ceiling at (x. 0.59 m, 3.24 m) with two closely spaced probes: probe 1 at x = 1 .63 m and probe 2 at X = 1.38 m. Each probe contained a filter. The flow system allowed bypass flow of air, flow through

probe 1 . or flow through probe 2. Downstream of the filter, the volume fractions of CO and CO2 were

monitored. This arrangement allowed the collection of two sequential filter samples. The second

location labeled "E" was at the east central location near the ceiling at (6.04 m, 1.56 m, 3.65 m). At this location the volume fractions of CO, CO2, and O2 were monitored, but there was no smoke aerosol

collection.

Figure 6-1 is a schematic diagram of the sampling system at the Sm location. The smoke was sampled at 3 L/min through 2 m of unheated section of 0.95 cm diameter stainless steel tubing and then a 60 cm long

section heated to 60 °C by a water jacket. This length of tubing allowed the incoming flow, which was as hot as 500 °C, to equilibrate to about 60 °C before being collected on a particulate filter located just

downstream of the water jacket. After the filter, the gas flowed thj"ough wet ice and dry ice traps to remove moisture, an absolute filter (99.9 percent efficient), a diaphragm pump, and another absolute filter to protect the mass flow controller. The flow was split after the flow controller with 1 L/min at a pressure

of about 100 kPa gauge to both CO and CO2 nondispersive infrared gas analyzers. The remaining flow was transported to the exhaust system. The total time from the start of sampling to a nearly constant output (within 10 percent of steady value) was about 45 s. This included about a 9 s response time of the

gas analyzers.

The sampling system at E was simplified relative to Sm because there was no smoke aerosol collection.

Thus there was no heated water jacket, heated filter holder, flow controller, or HEPA filter just before the

flow controller. Instead of a wet ice trap and dry ice trap in series, two dry ice traps were used in parallel to remove moisture from the 50 L/min sampling flow. The high flow was used to minimize the delay

time from the 8 m long copper sampling tube (0.95 cm OD tubing). The lower flow for the filter collection was required to allow thermal equilibration of the high temperature smoke with the water

jacket (60 °C) before reaching the filter. The total delay/response time at location E was about 15 s compared to 45 s at the Sm location. From the exhaust of the pump, a flow of 1 L/min each was directed

to the same type of CO and CO2 analyzer as above plus a paramagnetic type oxygen analyzer.

The effluent was continuously sampled at the E location throughout a test and the concentrations were

determined as volume fractions of CO2, CO, and O2 on a dry basis. Two hundred data points were

averaged over a 1 s period, and this average was recorded by the data acquisition system.

NISTNCSTAR 7-56, WTC Investigation 67
Chapter 6

Stainless Steel

heated filter holder

sampling sampling tube pre-weighed filter

l"'®* j

60 °C water jacket

^

t

TCs

to pump and

heated bath

TC

valves

-(8)—< bypass flow

HEPA filtei>L

to exhaust

flow

controller

CO2 analyzer

flow meters—

CO analyzer

Figure 6-1. Schematic of apparatus for measuring the smoke yield and smoke

concentration at location "Sm." The sampling tube extends from the inlet

through the water jacket to the filter holder.

For the Sm location, the smoke was sampled for an interval of 60 s or 120 s, assuring adequate collection

of particulate on the filter/tubing wall to make a quantitative smoke mass measurement. Two such

samples were collected during each test. As the gases flowed through the system, the CO2 and CO

analyzers increased to peak values about 45 s after the fiher collection was started. After the filter sample was collected, the valve position was changed so that clean air purged the remaining gases through the

analyzers. This way essentially all of the CO and CO2 passing through the filter also passed through the

gas analyzers. This was crucial for determining the species yield based on the carbon balance method as discussed below.

68 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

6.2 MEASUREMENT UNCERTAINTY FOR GAS ANALYZER RESULTS

All three gas analyzers have the same major sources of uncertainty: the uncertainty in the span gas, instnmient noise, and instrument drift. The values of these uncertainties are summarized in Table 6-1

The noise in the gas analyzer output, u,„ was computed as a standard deviation in the instrument response

to the calibration gas after correcting for long term drift. The value of ii„ was 1 x 1 0"'' for the oxygen

anah'zer and a factor of 2 greater for the CO and CO2 analyzers.

Gas

.Analyzer

Nominal

Volume

Fraction

Calibration

Gas Volume

Fraction

Noise,

Un

Calibration

Gas Uncert.,

Us

Total

Uncert.,

Ut

Drift/

30 min.

Expanded

Uncert.,

0: 0.15 0.2095 1.0 X

10"-' 5.0 X

10"^ 3.6 X10"^ 10.0 X 10"' 7.0 X 10^

CO: 0.04 0.1800 2.0 X

10"-' 1.8 X 10"^ 4.0 X

10"-' 4.0 X 10"' 8.0 X 10^

CO 0.0015 0.0700 2.0 X

10"-' 7.0 X 10"^ 2.5 X 10"' 5.0 X 10"' 11.0 X 10"'

Ambient air was used as the span gas for the oxygen analyzer. The volume fraction of oxygen in dried

ambient air was 0.20950. Based on scientific judgment, the estimated standard uncertainty for the span

gas, Us, was 5.0 x 10"^. For the CO2 and CO span gases, the estimated standard uncertainties were

1 percent of the volume fraction of the calibration gas.

The volume fraction of the gas, x, is proportional to the instrument voltage output.

(6-1)

where was the volume fi"action, and and V are the voltage outputs for the calibration gas and the

measured gas, respectively. The total uncertainty in x, u„ was computed using the law of propagation of

uncertainties. The result is conveniently expressed in terms of the relative uncertainty in the volume

fraction, u,(x)/x\

u, (x) + V

+

Vs

1/2

(6-2)

It is assumed that the voltage noise is the same for both the calibration and the measurements during the

test. For this case the ratios uy/Vand uvJVs are equal to u„lx and ujxs. The computed values for u, for typical concentrations are given in Table 6-1

There is one additional uncertainty component: the drift in the analyzer output. The drift in the output of

the gas analyzers was computed from repeat measurements over a 30 min period based on a hnear

regression of the data. The drift uncertainty, w^, was 10 x 10 "'' for the oxygen analyzer and about a factor of two smaller for the other gas analyzers (see Table 6-1).

NISTNCSTAR 1-5B, WTC Investigation 69
Chapter 6

A conservative estimate of the combined uncertainty, u,., was obtained by computing the root-sum-of

squares of z/, and

; =[w;+w;f' (6-3)

The expanded uncertainty, U, corresponding to the 95 percent confidence interval, is obtained as twice Uc. Table 6-1 shows the computed U for an oxygen volume fraction of 0.15, a CO2 volume fraction of 0.04,

and a CO volume fraction of 0.0015. The first two values are representative of the values observed for the smaller tests, and the value for CO was about the largest observed for the heptane-toluene bums. For

the relatively high O2 and CO2 concentrations, the dominant source of the uncertainty is the calibration

gas, while for the low CO concentration, the drift tenn is the dominant uncertainty. In Table 6-2, the

quantities AO(;^Q^ and AO^^ , which are the volume fractions of CO2 and CO of the gas minus the ambient

background volume fractions of these gases, are tabulated along with their uncertainties for the samples

collected at the Sm port for five tests. The reported values are the averages for the steady portion of the two

smoke sampling periods for each test. The expanded uncertainty for each quantity is also included.

Table 6-2. Mean value and expanded uncertainty for gas and smoke volume fractions

at the Sm sampling location.

1 estff- Probe# Fuel Period, s AO^o^i U AO^o ±U (IV1,± U), p/m^ 3)s± U

1-1 Heptane 60-180 0.039 1.1 X 10" 0.141 0.78 X 10"' ± 0.001 ± 1.1 X

10"" ±0.010 ±0.07 X10"'

1-2 Heptane 270-390 0.039 2.7 X 10"" 0.283 1.57 X 10

' ±0.001 ± 1.1 X

10"" ±0.018 ±0.13 X

10"'

2-1 60 % heptane 90-150 0.050 12.5 X 10"" 1.58 8.8 X

10"' 40 % toluene ±0.001 ± 1.3 X

10"" ±0.15 ± 0.9 X10"'

2-2 60 % heptane 300-360 0.051 14.5 X

10"" 1.85 10.3 X 10

' 40 % toluene ±0.001 ± 1.4 X

10"" ±0.17 1.1 X10"'

3-1 60 % heptane 90-150 0.038 8.1 X 10"" 1.01 5.6 X 10"' 40 % toluene ±0.001 ± 1.2 X 10

" ±0.08 ±0.5 X 10"'

3-2 60 % heptane 300-360 0.038 9.6 X 10"" 1.44 8.0 X 10"' 40 % toluene ±0.001 ± 1.2 X

10"" ± 0.12 ±0.8 X10"'

4-1 Heptane 90-210 0.069 18.0 X

10"" 0.176 0.98 X 10"' ±0.001 ± 1.5 X

10"" ±0.011 ±0.08 X10"'

4-2 Heptane 510-630 0.089 5.9 X 10"" 0.134 0.74 X 10"' ± 0.002 ± 1.1 X

10"" ±0.010 ±0.07 X10"'

5-1 Heptane 90-210 0.068 16.0 X

10"" 0.214 1.19 X10"' ±0.001 ± 1.4 X

10"" ±0.014 ±0.10 X 10

'

5-2 Heptane 510-630 0.077 6.8 X 10

" 0.213 1.18 X

10"' ± 0.002 ± 1.2 X

10'" ±0.016 ±0.11 X10"'

70 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

6.3 SMOKE AND GASEOUS YIELDS

The smoke yield was determined by the carbon balance method. This method required determination of

the ratio of the smoke mass in a given volume to the total mass of carbon in the form of gas or particulate

in the same volume. This was accomplished by dividing the smoke mass collected by the sum of the

smoke mass and the mass of carbon contained in the fornis of CO and CO2. Laboratory scale studies with

ethene indicate that for free burning conditions the contributions of unbumed vapor phase hydrocarbons

account for about 2 percent or less of the carbon (Leonard et al. 1994). The unbumed hydrocarbons were

neglected in this analysis. The equation for calculating the smoke yield (j'^) in terms of the CO2 and CO

concentrations is given by:

[ nis + 12uY^{A0coAti) + A0co{t,))At] ' (6^)

The quantity / is the carbon mass fraction of the fuel, is the mass of the smoke sample per second

collected on a filter and on the wall of the sampling probe, n is the number of moles of gas sampled, the constant 12 represents the molar mass of carbon in grams, and A 1 is the 1 second sampling interval. The

quantities ^^co. ) ^^co ) ^^e i'*' samples of the CO2 and CO gas volume fraction taken

during the test minus the ambient background volume fractions of these gases.

The yield for CO is given by a similar formula:

f /7 28^AOc^(r,)A/

yco=

' (6-5)

where 28 is the molecular weight of CO. The expression for v^^^ is identical to vco, except the 28 is replaced with 44 (molecular weight of CO2), and the CO volume fraction term in the numerator is replaced with the CO2 volume fraction. The results for the yields of smoke, CO, and CO2 are given in Table 6-4.

The measured mass concentrations of soot correspond to the values at the sample location at the gas

temperature measured by a thermocouple near the inlet. The value was computed using the following

formula:

m = (6-6) ^^'^smoke ) where V is the volume of smoke sampled at temperature Tsmoke- The volume is corrected for the moisture

removed by the ice bath by estimating the amount of water vapor at the inlet as a stoichiometric factor

times the volume fraction of CO2. The stoichiometric factor, SF, is the ratio of the moles of H2O

produced to the number of moles of CO2 produced based on complete combustion. This factor is 1.143

for heptane and 0.868 for the heptane-toluene mixture. The equation for the corrected volume is given

by:

=F/(l-5FxO^^J (6-7)

NISTNCSTAR 1-5B, WTC Investigation 71
Chapter 6

The volume fraction of soot, (l>s, was obtained by dividing the mass concentration by the density of soot, which was taken to be 1.8 g/cm\ This represents the volume of the condensed phase soot per volume of

gas at the sampling point. The results for the smoke concentration and volume fraction of smoke are given in Table 6-2.

6.4 MEASUREMENT UNCERTAINTY FOR SOOT, CO2, AND CO YIELDS

The law of propagation of uncertainty was used to compute the total uncertainty in the smoke yield

arising from the quantities appearing in Eq. 6-4. The contributions from smoke and CO in the denominator were neglected since they represent a small percent of the total carbon (less than 15 percent).

The resulting expression for the total standard relative uncertainty in the smoke yield is given by:

+ Y+ +

nl/2

'CO, AOCO, )

(6-8)

In this equation, the relative uncertainty in the sum of CO2 volume fractions was estimated as the

uncertainty in the average CO2 concentration, w^^^ , divided by the average value.

The uncertainty in the smoke mass, , includes contributions from the measurement of the mass of

smoke collected on the filter and on the walls of the tubing before the filter. The filter mass ranged from

about 2 mg to 10 mg, and the wall mass was in the range of 10 percent to 30 percent of the filter mass. For the filter, the estimated standard uncertainty for weighing was ± 0.007 mg and for particle penetration

through the filter was ± 0.02 /;?/. These contributions were combined in quadrature to obtain the

uncertainty associated the mass on the filter. The tubing walls are cleaned three times to remove the

deposited soot, m„ . The standard uncertainty in m„ was estimated as ± 0.10 m,! based on weighing

uncertainty of the cleaning cloth and on the residual smoke mass not removed from the tubing surface. The total uncertainty for the smoke mass, zy,„ , was computed by combining in quadrature the

uncertainties in W/ and 7/7,,

.

There was uncertainty in/for the heptane blend and the heptane/toluene mixture because the

compositions of the mixtures are not exactly known. In these cases, the uncertainty in/was taken to be

1 percent of the estimated value (/ = 0.84, heptane blend; / = 0.87, 40 percent toluene. 60 percent heptane

blend by volume).

The molar flow, h , was obtained from the measured pressure, temperature, and volumetric flow using the

ideal equation of state:

n = PV

'rt

(6-9)

The uncertainties in P, T, and V were ± 0.3 kPa, ± 1 °C, and 1.1 x lO'"* L/s, respectively. The volumetric

flow uncertainty includes an uncertainty in the flow calibration and an estimated leakage rate. Using the

law of propagation of uncertainty, the combined uncertainty for the molar flow, , was ± 4.9 x 10"^ moles/s for a nominal flow of 2.18 x 10""^ moles/s.

72 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

The last pieces of information for computing values of w„.( V J is ^^co. associated uncertainty,

which was given in Table 6-2. The values of the uncertainties listed above are given in Table 6-3 for one

specific test, and the resuhs for the smoke yields together with the expanded uncertainty (with a coverage

factor of two) are given in Table 6-4.

Table 6-3. Major contributors to the uncertainty in the smoke yield.

\'ariable Typical Value

Relative Standard

Uncertainty'

Mass on the filter 1 .594 mg 0.020

Mass on cloth 0.366 mg 0.10

Total mass of smoke 1.960 mg 0.025 Carbon fraction of fuel 0.84 0.01 Molar flow rate 2.19 X 10"- mole/s 0.022

Volume fraction of CO: - bgd. value 0.0393 0.010

Smoke yield 0.0135 0.036'

a. This value, the relative total standard uncertainty', was computed as the root-sum-squares of the uncertainties in the mass of smoke, carbon fraction, molar flov\', and CO2 volume fraction.

Table 6-4. Mean value and expar

(on a dry basis

ided uncertainty for the smoke and gas yields

at the Sm sampling location.

Test#-

Probe# Fuel

Sample

Period (s) v,± U yco ± U yco, ± ^

1-1 Heptane 60-180 0.014 ±0.001 0.006 ± 0.006 3.02 ±0.06

1-2 Heptane 270-390 0.028 ± 0.002 0.013 ±0.005 2.96 ±0.06

2-1 60 % heptane

40 % toluene

90-150 0.12 ±0.01 0.045 ± 0.005 2.70 ±0.07

2-2 60 % heptane

40 % toluene

300-360 0.14±0.01 0.048 ± 0.005 2.60 ±0.08

3-1 60 " 0 heptane

40 % toluene

90-150 0.089 ± 0.006 0.038 ±0.06 2.80 ±0.07

3-2 60 % heptane

40 % toluene

300-360 0.13 ± 0.01 0.042 ± 0.005 2.64 ±0.07

4-1 Heptane 90-210 0.012 ±0.001 0.049 ± 0.004 2.96 ±0.06

4-2 Heptane 510-630 0.0088 ± 0.0007 0.013 ±0.003 3.03 ± 0.06

5-1 Heptane 90-210 0.015 ±0.001 0.045 ± 0.004 2.96 ± 0.06

5-2 Heptane 510-630 0.015^0.001 0.018 ±0.003 3.00 ±0.06

The major terms contributing to the uncertainty in the CO yield are the uncertainty in the CO volume

fraction, which affects the numerator of Eq. 6-5, and the volume fraction of CO2, which was the

dominant term in the denominator. The relative total standard uncertainty in the CO yield, Utr(yco), was

equal to the root-sum-square of the relative standard uncertainties for CO and COt measurements.

^n(yco) = U{co) + uUm))" (6-10)

NISTNCSTAR 1-5B, WTC Investigation 73
Chapter 6

The relative uncertainty for the CO and CO2 measurements are obtained from Table 6-2. The expanded

uncertainty was computed based on Eq. 6-8, using the same general approach as above. The results are given in Table 6-4.

The uncertainty in the yield of CO2 was small, because the numerator and denominator of the yield

equation were dominated by the volume fraction of CO2. Making use of the dominance of the CO2 term

in the denominator, the following approximate expression for was obtained:

44 , 44 44

yco,=—f-^ys-—yco

Applying the law of the propagation of uncertainties and a coverage factor of two, the expanded

uncertainty was computed for each test based on the uncertainties in/ v,, and vco. The results are given

in Table 6^.

As seen from Eq. 6-6, the uncertainty in the mass concentration of smoke results from the uncertainty in the mass of smoke collected, in the uncertainty in the temperature, which was estimated as ± 1.5 percent

of the absolute temperature of the smoke, and the uncertainty in the corrected volume, which was

estimated to be ± 2.7 percent for the 60 s smoke sample and ± 1 .8 percent for the 120 s sample. The large

uncertainty in the smoke temperature results from the effect of radiant transport on the thermocouple.

The dominant factors affecting the volume of smoke sampled are the timing uncertainty (± 1 .4 s), the

flow calibration (± 1 percent), and uncertainty in the water vapor correction (± 1 percent). The expanded

uncertainty in the mass concentration of smoke was computed via the law of propagation of uncertainty

for Eq. 6-6. The results are given in Table 6-2.

The uncertainty in the soot volume fraction measurements are a quadrature sum of the uncertainty in the mass concentration and the uncertainty in the soot density, which was estimated to be ± 2.5 percent of the

nominal value. These results are also presented in Table 6-2.

6.5 EXPERIMENTAL OBSERVATIONS

Figure 6-2 shows the oxygen reading at the E location as a function of time for Test 1. The oxygen

volume fraction decreased rapidly from a value of about 0.21 to 0. 16 and then fluctuated about this value

for the remainder of the test. The standard deviation of the oxygen volume fraction computed over a small time interval (about 10 s) during the nominally steady period was about 0.001. This value was more than two orders of magnitude larger than the standard uncertainty associated with the noise of the

oxygen analyzer responding to ambient air. This was a result of the fluctuations in the oxygen

concentration in the enclosure. The actual fluctuations in the oxygen concentration were greater than

those recorded by the oxygen analyzer because the analyzer signal was smoothed by the 9 s response time

of the instruments.

Representative plots are presented, and key features of the tests are described, regarding the following

topics: repeatability, effect of smoke concentration, and the effect of heat release rate on the low and high

sooting fuel.

74 NISTNCSTAR 1-5B. WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

.s

0.15 O

0.1

I I I I I I I 'I I ' ' 'I I I I I I I I I I I I

I I I

0 100 200 300 400

Time (s)

500 600 700 800

Figure 6-2. Oxygen volume fraction (on a dry basis) sampled at location E

plotted as a function of time for Test 1

6.6 REPEATABILITY OF TESTS

Tests 4 and 5 were basically repeat tests with heptane as the fuel at a heat release rate near 3 MW. As

mentioned above, the gases were sampled continuously at the E location but only for 2 min intervals at the Sm location. The results for O2 and CO2 (Figs. 6-3 and 6-4) were typically within 10 percent for all the data. Over a period of 800 s, the CO: volume fraction increased by 10 percent to 20 percent, and the

O2 volume fraction decreased by a similar amount. The CO2 and temperature were slightly higher for Test 4 and the O2 slightly lower. These results are consistent with the slightly higher heat release rate measured in Test 4. A consistency test was made for the O2 and CO2 measurements by computing the ratio of the volume

fraction of CO2 to the decrease in the volume fraction of the O2. For the CO2 volume fraction, the

ambient CO2 volume fraction was subtracted. For complete combustion of heptane, this CO2/O2 ratio (on

a dry basis) was 0.64. The measured values ranged from about 0.67 to 0.69 (Fig. 6-5). The relatively

close agreement and the absence of large drifts provide validation of the gas analyzer output.

Some of the difference between experiment and theory could also be due to incomplete products of

combustion. The ratio for Test 5 was nearly constant at a value of about 0.67 over an 800 s period with

significantly less fluctuation than either the O2 or CO2 measurement. Increases in O2 concentration were

accompanied by decreases in the CO2 concentration. A clear example of this effect is seen in Figs. 6-3

and 6-4 at a time of about 360 s.

The CO results (Fig. 6-6) were similar for Tests 4 and 5, considering the low concentration and higher

measurement uncertainty. The CO concentration was a factor of 1.5 to 2.5 larger at the smoke sampling

port (Sm) compared to the ceiling port (E) to the east of the fire. This was compared to a factor of 1 . 1 to

1 .2 for the case of CO2. This same qualitative effect of a higher CO concentration at the Sm sampling

port was observed for all the other tests except for the 2 MW heptane fire without insulation on the steel. In this case, the CO was near the noise level of the instrument. The higher CO concentration at the Sm

port may result from a difference in the combustion environment on one side of the enclosure versus the

other. Another possibility is a performance issue with the CO analyzers. The measured CO volume

NISTNCSTAR 1-5B, WTC Investigation 75
Chapter 6

fraction varied from about 2 x 10"* to 2 x 10"^ compared to an estimated expanded uncertainty of

1.1 X lO'"*. The linearity of the CO analyzer output was verified at a concentration as low as 1.6 x 10'\

which was comparable to the observed concentration at 200 s. It is unlikely that this difference was a gas

analyzer issue, but such an effect cannot be ruled out without additional study.

.a

-t-«

O

0.25

0.2

0.15

0.1

0.05

1 1 1

1 1 1 1 1-—I

1 1 11 1 1

1 r WTC PHASE 1 TEST 4 &

E02-4

I I I J I I I I I I L

0 200 400 600 800

Time (s)

Figure 6-3. Comparison of the oxygen volume fraction (on a dry basis) for Tests 4 and 5.

0.12

0.1

0.08 h

'

I I —•—SMC02-4 ——EC02-4 —-O—- SMC02-5 —B— EC02-5

T

1 11 1 1

1

1 1 1

1 J— WTC PHASE 1 TEST 4 & 51

400 600

Time (s)

Figure 6-4. Comparison of CO2 volume fraction (on a dry basis) for Tests 4 and 5. Results are presented for the E sampling location and the Sm location, for which the gas

is sampled twice for 120 s each.

76 NIST NCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

O

ao

:§uO

I

>

0.76

0.74

0.72

0.64

T r-| n n—|—i—i—1—r-|—m—1 1 |—n—rn—

|—1—ri—r-|—rn—rn—

[—1—i—r-r

WTC Phase 1 Test 4 &5

Test 4

Tests

J i_l i_i i_i 1—1 i__i i_l i_i I I I i_i i_j l_i i_i I I I I I I I I I I I

1 00 200 300 400 500

Time (s)

600 700 800

Figure 6-5. Ratio of volume fraction of CO2 minus the background value to the decrease

in the O2 volume fraction for Tests 4 and 5 at sample location E (on a dry basis).

0.002

0.0015

0.001

8 0.0005

1 1

1

1 1 r

1 1 1

1 1 1 1 1

WTC PHASE 1 TEST 4 i& 5

-•—SMCO-4 —-G-— SMCO-5

H—ECO-4 —a— ECO-5

100 400

Time (s)

600 800

Figure 6-6. Comparison of CO volume fraction (on a dry basis) for Tests 4 and 5.

The soot yields for Test 4 were 0.013 and 0.0092 and for Test 5 were 0.015 and 0.015. These results are similar to the free bum results (refer to NIST NCSTAR 1-5C), although one of the Test 4 results was

unexpectedly low in value. The soot yield does not exhibit the large sensitivity to the sampling time that was observed by CO, so it appears that a constant soot yield equal to the free bum value was not a bad

approximation for this case.

NIST NCSTAR 1-5B. WTC Investigation 77
Chapter 6

6.7 EFFECT OF SMOKE CONCENTRATION

Tests 1 and 3, both with a heat release rate of about 2 MW, involved a relatively clean burning heptane

blend (Test 1) and the sootier heptane/toluene mixture (Test 3). The results for the concentrations of CO2

and O2 for the two tests typically agreed within 5 percent of their values. The CO2 concentration reached

a constant value shortly after ignition as illustrated in Fig. 6-7. The O2 concentration behaved in a similar

fashion. This steady behavior differed from the time dependent behavior for these species for the 3 MW

fires. The CO2 concentration was about 15 percent larger at the Sm port compared to the concentration at the E port. A similar difference was found in Tests 4 and 5. The temperature near the Sm port increased

from about 300 °C to 450 °C during the test as shown in Fig. 6-8.

0.06

0.05

•2 0,04

I

0.03 1

o

0.01

0

0 100 200 300 400 500 600 700 800

Tim e ( s

)

Figure 6-7. CO2 and O2 volume fraction (on a dry basis for low (Test 1) and high sooting

(Test 3).

For the Test 3 bum of the heptane/toluene mixture, the yields of both CO and soot were about five times

larger than for Test 1. The CO volume fraction at both sampling locations (Fig. 6-9) increased slightly

(10 percent to 20 percent) during Test 3, while the CO volume fraction increased by more than a factor of

two for heptane. The soot yield for the second sample collected for the heptane bum was also found to approximately double the first value, while for the heptane/toluene mixture, the second value increased by

about 40 percent.

78 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

I I I I II

I I I I I I I I I

I I I I II

I I I I

I

I I I I

I

I

1 WTC PHASE 1 TEST 1 & 3

200 - S

H .

100 ;

-

Q

I I I I I I 1 I I I I I I I I I I I I I I I I I I I I I 1 1 I I I I I I I r I I

0 100 200 300 400 500 600 700 800

Time (s)

Figure 6-8. Temperature as a function of time near the Sm probes

for low (Test 1) and high sooting (Test 3).

6.8 EFFECT OF HEAT RELEASE RATE FOR A LOW SOOTING FUEL

The effect of the heat release rate on species concentrations can be examined by comparing Test 1 and

Test 4. As seen in Figs. 6-9 and 6-10, the O2 and CO: volume fraction in Test 1 approached a constant

value, whereas in Test 4. the concentrations continued to change as discussed in Sec. 6.6. The larger fire resulted in nearly a 200 °C higher temperature. At 800 s, the temperature near the Sm port was 660 °C

for Test 4, as compared to 460 °C for Test 1.

The trends in the CO concentration were the most complex as illustrated in Fig. 6-11. First, consider the

CO data at location E. The larger fire (Test 4) initially had more than 10 times the CO concentration of

the smaller fire (Test 1 ); however, by 800 s, the situation reversed and the CO concentration was larger

for the smaller fire. For the smaller fire, the smoke yield also increased with time and approximately

doubled from 200 s to 400 s. Such an increase was not surprising, since the fire plume entrained an

increasing amount of combustion products as the smoke layer developed. The surprising result was the

very large CO volume fraction at the test start in the larger fire. One would expect the CO yield to be

similar to the free bum results, but the value was about 5 times greater for Test 4 as compared to the free bum (Table 6-5) (Hamins et al. 2003). This suggests a drastic difference in the fire conditions. The one

difference from the free bum fires was the inclusion of sprayed fire-resistive material (SFRM) around the

steel. It is possible that the SFRM may have been responsible for the increased amount of CO observed.

The high initial amount of CO was seen in both tests (see Fig. 6-6) involving the insulated steel, and it was observed at both sampling points.

NISTNCSTAR 1-5B, WTC Investigation 79
Chapter 6

0.0012

0.001

I

0.0008

T I jI I I r~] I I I I I

I I I 1 II I I I

I

I I I I I I n I I I—I—I—I—

I—I—I—I—

r

WTC PHASE 1 TESTl & 3 , —#—SMCO-1 —e—ECO-i -m - SMCO-3 —-Et— ECO-3 , WH

0.0006

? 0.0004 O

0.0002 MM

1 1I 1 1

1

' I 1

1 11

I I I I I I

0 100 200 300 400 500 600 700 800

TiiTie(s)

Figure 6-9. CO volume fraction (on a dry basis) as a function of time for low (Test 1)

and high sooting (Test 3).

o

0.25

0.15

0.05

I I I I [

I I I I II I I I

I

I I I I I I I I I

I

I I I I

I

I I I I

I

I I I WTC PHASE 1 TEST 1 & 4

O—- E02-4

_L 1 I I I 1 I [ I I I I I I ' I I ' I I ' I I I I I I I

0 100 200 300 400 500 600 700 800

Time (s)

Figure 6-10. The oxygen volume fraction (on a dry basis) for a lower (Test 1)

and a higher (Test 4) heat release rate.

80 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

Time (s)

Figure 6-1 1 . The CO2 volume fraction (on a dry basis) for a lower (Test 1 ) and

a higher (Test 4) heat release rate.

Table 6-5. Mean value and expanded uncertainty for the smoke and

gas yields (on a dry basis) for free burns.

Fuel yco ± l; yco, ± LI

Heptane 0.015 ^ 0.002 < 0.008 3.0±0.1

60 % heptane

40 % toluene

0.11 :i0.02 0.042 ±0.016 2.7 ±0.02

Source: Hamins et al. 2003.

6.9 EFFECT OF THE FIRE HEAT RELEASE RATE ON A HIGH SOOTING

FUEL

Tests 2 and 3 also provided information on the effect of the heat release rate, but in these cases the

heptane/toluene mixture produced a high smoke concentration (Fig. 6-12). The heat release rate was

varied by about 20 percent from 2.4 MW for Test 2 to 2.0 MW for Test 3. The insulated steel was not

present during the heptane/toluene fires. The trends in terms of the magnitude and time dependence for the CO:, O2, and temperature are qualitatively similar to that for Tests 1 and 4. The one major difference was the trend regarding the CO concentration. As seen in Fig. 6-13, the CO concentration increased

slightly with time for both tests with, possibly, a greater increase for the larger fire. The CO results after 400 s for the E port were not valid because the fuel flow was stopped at this time and the pump turned off. The time dependence of the CO was not similar to either Test 1 or Test 4. Perhaps the higher CO

production from this sooty flame made it less sensitive to small changes in the combustion conditions.

NISTNCSTAR 1-5B. WTC Investigation 81
Chapter 6

0.002

0.0015

0.001

8 0.0005

-]—I—I—1—I—I—I—.- WTC PHASE 1 TEST 1 & 4 -I

800

Figure 6-12. CO volume fraction (on a dry basis) for a lower (Test 1) and

a higher (Test 4) heat release rate.

0.002 I—'—I—I—I—

I—'—'—'—I—i—I—I—I—I—

I—I—I—I—I—

I—I—I—I—

T

Time (s)

Figure 6-13. CO volume fraction (on a dry basis) for high sooting tests with a low

(Test 3) and a high (Test 2) heat release rate.

82 NISTNCSTAR 1-5B, WTC Investigation
Smoke, CO2, CO, and O2 Concentrations

6.10 REFERENCES

Blevins, L. G., and W. M. Pitts. 1999. Fire Safen> Journal 33, 239-259, and National Institute of

Standards and Technology, Gaithersburg, MD, NISTIR 6310, April. Available at http://fire.mst.gov/bfrlpubs/fire99/PDF/f99022.pdf.

Hamins, A., A. Maranghides, and G. Mulholland. 2003. The Global Combustion Behavior of 1 MW to 3 MWHydrocarbon Spray Fires Burning in an Open Environment, NISTIR 7013, National Institute of Standards and Technology, Gaithersburg, MD, June.

He, Y. P., A. Fernando, and M. C. Luo. 1998. Fire Safety Journal, 31, 19-38.

McCaffrey, B.J., and G. Heskestad. 1976. Combust. Flame, 26, 125-127.

Pitts, W. M., E. Braun, R. D. Peacock, H. E. Mitler, E. L. Johnsson, P. A. Reneke, and L. G. Blevins.

1998. Temperature Uncertaintiesfor Bare-Bead and Aspirated Thermocouple Measurements in Fire

Environments and National Institute of Standards and Technology, Gaithersburg, MD, NISTIR 6242,

October. Available at http://fire.nist.gov/bfrlpubs/fire03/PDF/fD3052.pdf

NISTNCSTAR 1-5B, WTC Investigation 83
Chapter 6

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84 NISTNCSTAR 1-5B, WTC Investigation
Chapter 7

Velocity Measurements

7.1 INSTRUMENTATION

Bidirectional probes were used to measure gas velocities at the inlets and outlets of the compartment as a check on the computational model. These instruments were selected because they are durable and have

the necessary measurement characteristics. McCaffrey and Heskestad (1976) showed that bidirectional

probes are capable of measuring low velocities and are relatively insensitive to the direction of flow.

They found the velocity (mean speed) measurement to be accurate to within 10 percent, provided that the

velocity probe was aligned to within ± 50 degrees of the mean velocity direction.

The bidirectional probe measurement involves relating the differential pressure across the probe to the

local Reynolds number, Re, of the gas. The calibration for the probe is typically presented in terms of the

amplification factor K:

where, D, /u, Ap and are the probe diameter, gas viscosity, differential pressure across the probe, and

dynamic pressure of the gas, respectively, and is a function of Reynolds number (Re):

The velocity of the gas, V, and gas density, p, both appear in the amplification factor and the Reynolds

number. Thus, determination of the gas velocity from the measured data can be an iterative procedure.

For the measurements reported here, the probe dimensions were chosen to ensure that the Reynolds

numbers obtained would be in the asymptotic regime (Re > 600) to minimize the measurement

uncertainty and eliminate the need for an iterative procedure.

Calibrations relating the amplification factor, K, to the Reynolds number have been reported by several

researchers. The original calibration by McCaffrey and Heskestad (1976) found that K decreases from

approximately = 1 .5 aX Re ~ 0 to K ~ 1.1 atRe- 2000. An asymptotic value of AT = 1 .08 was reported

for large values of Re. In contrast, Kent and Schneider (1987) reported that K decreases with decreasing

Re for Re < 600. These researchers reported an asymptotic value of = 1 .07, which they claim is valid

for all values of Re > 600. The data of Liu et al. (1990), which correspond to bidirectional probes of four

different sizes, indicate that the amplification factor neither increases nor decreases for low values of Re. However, the data indicate that the uncertainty in the calibration increases significantly for small Re. They reported a mean value of 1 . 18 for A^ for Re > 600 and noted that this corresponds to a difference of

about 7 percent from the data of McCaffrey and Heskestad (1976). For these calculations, the asymptotic

value of K (equal to 1.08 as reported by McCaffrey and Heskestad [1976]) was used, and the uncertainty

in this value was taken as 10 percent.

(7-1)

Re^pDV I ^ (7-2)

NISTNCSTAR 1-5B. WTC Investigation 85
Chapter 7

The three investigations discussed above also included measurements of the effect of yaw angle, 9, on the

velocity measurement. McCaffrey and Heskestad (1976) reported that the mean velocity was within

±10 percent provided that the approaching flow direction was known to be within approximately

50 degrees of the axis of the probe. This conclusion was based upon data obtained at a Reynolds number

of 3400. Kent and Schneider (1987) explored the effect of yaw angle at Re = 900 and concluded that their

data were consistent with those of McCaffrey and Heskestad (1976). Liu et al. (1990) observed a

variation of approximately 15 percent in the value of the amplification factor over the range of yaw angles

0 < 6 < 60 degrees. Their data correspond to Reynolds numbers ranging from 300 to 550.

Two conclusions can be drawn from the published data. First, valid measurements of the mean gas speed

can be obtained provided the probe is aligned to within 50 degrees of the mean flow direction. The

bidirectional probe measurement corresponds to a total velocity and cannot be decomposed into

individual components. The probes were aligned perpendicular to the plane defined by the compartment

openings. The flow predicted by the simulations confirmed that this was within 50 degrees of the mean

flow direction. Second, the uncertainty is very large for low values of Re, and the measurement should be

limited to regions of the velocity field where Re > 600. For the probes used here, that corresponds to mean velocities of V^in ~ 0.5 m s"' for 7= 20 °C and Vmm ~ 3 m s"' for T= 500 °C. The probe dimensions were chosen to ensure the measurements would be in the high Reynolds number regime and thereby

provide reliable data.

Bare-bead thermocouples (Type-K) were used to measure the temperature of the gas. Data obtained from

bare-bead thermocouples requires a correction for radiative heating effects on the thermocouple.

Aspirated thennocouples (Type-K) were used in conjunction with the bare-bead thermocouples to measure the temperature of the gas. The aspirated themiocouple data provide a check for the radiation

coiTection and was useful in the estimate of measurement uncertainties.

7.1.1 Sampling Locations

The bidirectional probes and their corresponding thermocouples were placed in the inlet and outlet openings of the compartment. During Test 1 , there were six probes located in the compartment inlet and

eight probes at the outlet. For Tests 2 through 6, an additional two probes were added to the outlet so

there were six probes at inlet locations and ten at outlet locations. A larger number of measurement

locations were used at the outlet because the gradient of velocity was expected to be larger there. Table 7-1 lists the horizontal and vertical positions of the bidirectional probes.

86 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements

Table 7-1. Bidirectional probe measurement locations.

Inlet Measurements

Probe Y location z location (from floor)

Inlet 1 0.33 m 0.71 m

Inlet 2 0.33 m 0.46 m

Inlet 3 0.33 m 0.25 m

Inlet 4 1.27 m 0.71 m

Inlet 5 1.27 m 0.46 m

Inlet 6 1.27 m 0.25 m

Outlet Measurements

Probe y location z location (from floor)

Outlet 1 0.33 m 2.61 m

Outlet 2 0.33 m 2.36 m

Outlet 3 0.33 m 2.15 m

Outlet 4 0.60 m 2.61 m

Outlet 5 0.60 m 2.15 m

Outlet 6 1.26 m 2.61 m

Outlet 7 1.26 m 2.36 m

Outlet 8 1.26 m 2.15 m

Outlet 9 (excluding Test 1) -1.26 m 2.61 m

Outlet 10 (excluding Test 1

)

-1.26 m 2.15 m

7.1.2 Thermocouple Radiation Correction

Radiative exchange affects the bare-bead thermocouple signal because there is a difference between the

local gas temperature and the temperature of the thermocouple bead as discussed in Section 4.2. Thus, a correction was applied to the temperature measurements. It was necessary to consider the flux of energy

to and from the thermocouple due to radiation and convection. The thermocouple can exchange energy

with the surrounding gas by radiation and convection. An energy balance on the thermocouple bead sets the net flux on the bead equal to the radiative and convective exchange:

Q,A,=A,asT,UAi,h(T, -T^J (7-3)

NISTNCSTAR 1-5B, WTC Investigation 87
Chapter 7

Here, is the incident radiation flux on the themiocouple bead , estimated based on FDS model

predictions. At is the surface area of the thermocouple bead, Ti, is the temperature of the themiocouple

bead, and Tgas is the free stream temperature of the gas. The emissivity, e, is assumed to be equal to 1.0, and (7 is the Stefan-Boltzmann constant. The convective heat transfer coefficient, h, is obtained from a semi-empirical correlation for the Nusselt number, which relates the convective heat transport from the

thermocouple, assumed to be spherical, to the local Reynolds number. Re, and Prandtl number, Pr, of the

fluid:

Nu = ^= 2 + 0.6Re'"-Pr'" (7-4)

Here, Nu is the Nusselt number, r is the bead radius, and kg is the thermal conductivity of the gas. To

solve for the free stream gas temperature, Eq. 7-3 is rewritten as:

T =T +^ 9jl

h

(7-5)

From Eq. 7^, the convective heat transfer coefficient, h is a function of Re, Pr, and k„, which all depend on r„,„. An initial value for T„„, is assumed and an iterative process is used to solve for the correct gas

temperature. These equations were developed using the study by Blevins and Pitts (1999), which

contains additional information on radiation corrections for bare-bead thermocouples.

7.2 MEASUREMENT UNCERTAINTY

Estimation of the uncertainty in the velocity measurements involves estimation of the uncertainty of

the various components of the measurement and the effect that each component has on the overall

uncertainty. Computation of the overall uncertainty involved uncertainties for the pressure transducer

calibrations, temperature measurements and the gas density calculation. To account for the various

components that contribute to the overall uncertainty, the equations used to analyze the data are considered and a propagation-of-error analysis is performed using the following equation:

wJv) =

dv

dAP

V

u:{AP)+

dp

dv

dK

V

nl/2 Km (7-6)

Table 7-2 summarizes the estimates of the uncertainty of the various components and of the overall

uncertainty of the velocity.

88 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements

Table 7-2. Velocity measurement uncertainty.

Parameter

Inlet Outlet

Mean \'alue Uncertainty, Mean Value Uncertainty, Uc Molec. Wt'

0.029 kg/mol Negligible 0.029 kg/mol 0.002 kg/mol

^(Eq. 7-1)

1.08 0.108 1.08 0.108

T

21 °C 3 °C 400 °C 50 °C

AP

1.5 Pa 0.51 Pa 7 Pa 0.59 Pa

P

1.20 kg/m' 0.012 kgW 0.52 kgW 0.053 kg/m'

V

1.7 m s 0.29 mJs 5.6 m/s 0.57 m/s

a. Molec. Wt. is an estimate of the average molecular weight of the gases near the probe.

7.3 DISCUSSION AND RESULTS

Velocity measurements were taken at both the inlet and the outlet of the enclosure as identified in Table 7-1 . Plots of the velocities and temperatures measured at each of the inlet and outlet locations

are included in Appendix D of this report, while representative data are discussed in this section. The

velocities measured at the inlet increased in the seconds immediately following ignition and then

remained fairly steady. Figure 7-1 shows the velocity measurement of inlet probe 1 during Test 1.

Its behavior is typical of the results measured at all inlet locations for all six tests. All of the inlet velocities measured had mean steady values between 1.5 m/s and 2.1 m/s with a standard deviation of

approximately ± 0.3 m/s. To eliminate radiation effects on the measurements, the inlet gas temperature was assumed to remain at its initial ambient temperature throughout the test duration. This temperature

approximation was found to have a minimal effect on the velocity measurement, as the velocity calculated

using the assumed inlet temperature varied by less than 5.0 percent from the velocity calculated using the

bare-bead thermocouple temperature.

WTC Phasel Test 1 Velocity Inlet 1 i

:

I I I . . 1 t , I I , > , I > ^ > I > I 1 ' 0 200 400 600 800 1000 1200

Time (s)

Figure 7-1 . Inlet velocity measurements for Test 1 , inlet velocity probe 1

2

1.5

o^ 1 >

0.5

NISTNCSTAR 1-5B, WTC Investigation 89
Chapter 7

As expected, most of the outlet velocities rose from the beginning of the test to several minutes into the

test, and many continued to increase tliroughout the test. The velocities varied with both the heat release

rate ( Q) and the location. For measurements at the same location, the velocity was greater when the heat

release rate was higher. This is demonstrated in Fig. 7-2, in which the velocity at location 1 in Test 1 {Q = 2 MW) was significantly smaller than in Test 4 ( ^ = 3.2 MW). The mean velocity for Test 1 increased from approximately 6.0 m/s to approximately 6.5 m/s during the test. The increase in mean

velocity during Test 4 was much steeper, again starting near 6.0 m/s, but reaching nearly 10 m/s over the same time period as Test 1 . The increased velocity for higher heat release rate ( Q) tests was consistent

for all locations. Another example is shown in Fig. 7-3, where the velocity at location 10 for Test 3 iQ = 2 MW) was lower than for Test 5(^ = 3 MW).

10

_o>

WTC Phasel

^ Velocity Outlet 1 ..A^MkMlesiA

mm

\,MkiLl

Test 1

200 400 600 800

Time (s) 1000 1200

Figure 7-2. Outlet velocity at location 1 for Test 1 and Test 4.

10

n 1I—1—

r

WTC Phasel Test 5

I

' Outlet 1

I 400 800 1200 1600 2000 2400 2800 3200

Time (s) Figure 7-3. Outlet velocity at location 10 for Test 3 and Test 5.

90 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements

The measurements indicate that the gas velocity was greatest at the top of the vent and decreased with

height. Figure 7^ shows an example of the measured outlet velocities for probes at locations 1, 2, and 3 during Test 5. These three probes were aligned vertically. Location 1 was at the uppennost location,

61 cm above the vent opening, location 2 was in the middle, 35.5 cm above the vent opening, and

location 3 was the lowest location, 1 5 cm above the vent opening. The velocity at location 1 was greatest,

and decreased for lower locations in the vent. This trend for probes at 1,2, and 3 was consistent for all of

the tests. In general, the upper probes recorded higher velocities than the lower ones for the other sets of

probes aligned vertically. Some of the locations, however, saw more flucmations and there was some

amount of overlapping. Figure 7-5 illustrates another example of probes aligned vertically during Test 5, this time ^^'ith location 6 as the uppermost location, location 7 in the middle, and location 8 as the lowest.

The velocity at location 7 was consistently higher than the velocity at location 8. The velocity at 6

(shown with dotted line), however, overlapped the other velocities during the early part of the test, before

finally surpassing the other velocities about 25 min into the test. The velocity at location 6 deviated from

the mean much more than the other measurement locations, fluctuating by about ± 1 m/s or more

compared to about ± 0.5 m/s for the other locations. The velocity at location 6 behaved in this manner for

all of the tests.

A comparison of Figs. 7-4 and 7-5 shows that the outlet velocities were not symmetric about the

enclosure, which was consistent with the non-symmetric appearance of the fire. The probes at locations 6. 7 and 8 were positioned at the same heights as probes 1, 2 and 3, respectively. But, as seen

in Fig. 7-5 for Test 5, the velocities at location 7, approximately 7 m/s, were greater than the velocities at location 2, approximately 5.5 m/s, and the velocities measured at location 8, approximately 5 m/s, were

greater than those at location 3, less than 4 m/s, even though these probes were at the same heights in the

outlet. The \ elocity at location 6 was consistently lower than location 1 during the first half of the test, and the velocity at location 6 fluctuated much more than at location 1.

Figure 7-6 is an example of non-symmetry in measurements across the enclosure outlet. The figure

shows the velocities for Test 3 at the lowest location, 15 cm above the vent opening. Location 8 was

closest to the south wall, followed by location 5, location 3, and location 10, which was closest to the

north wall. As shown in Fig. 7-6, the velocities were higher at the outermost locations, 8 and 10, than

at inner locations 5 and 3. This trend of lower velocities at the inner locations for the lower row of probes was evident during all of the tests, although the magnitudes of the velocities varied.

The opposite was seen for the probes at the uppermost location in the outlet, 6 1 cm above the opening.

The probes near the center of the compartment, locations 4 and 1, measured higher velocities than the

probes at the outermost locations of the compartment, at locations 6 and 9. This is shown in Fig. 7-7,

which plots the mean velocities for probes at outlet locations 6, 4, 1, and 9 for Test 3. Mean velocities are plotted because the measurements were all within 2 m/s of each other and often overlapped. Similar

results were obtained in all six tests. With the exception of Test 2, the velocity at location 9 was lower

than the other velocities at this height in the outlet.

NISTNCSTAR 1-5B, WTC Investigation 91
Q I I i :

\ I I I I I I I I I I I ! I I I I I I 1 I I I I

I I I I 0 400 800 1200 1600 2000 2400 2800 3200

Time (s) Figure 7-4. Outlet velocities for Test 5 at probe locations 1, 2, and 3.

NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements

— 3

C/3 1

o

o

CD >

WTC Phasel Test 3

2 r1 r

n 1 r

Outlet 8

Outlet 10

Outlet 5

Outlet 3 J

200 400 600 800

Time (s) 1000 1200

Figure 7-6. Outlet velocities for probes located 15 cm above vent opening, Test 3. Outlets 8 and 10 are the probes toward the sides of the compartment,

while outlets 5 and 3 are closer to the center of the compartment.

7 r

1 1

' 1

1

i

1 1

1 i

1

1 1 1 1 1

1 r

P WTC Phasel Test 3

0 200 400 600 800 1000

Time (s)

Figure 7-7. Outlet velocities for probes located 61 cm above vent opening in Test 3. Outlets 6 and 9 are toward the side, while outlets 4 and 1 are closer to the compartment center.

NISTNCSTAR 1-5B, WTC Investigation 93
Chapter 7

7.4 RADIATION CORRECTION

The outlet velocities were computed using a gas temperature value that was corrected for radiation effects using values of Qi, taken from the FDS simulations. The corrected gas temperature T„„, was found by

reducing the measured bare-bead temperature Tf, by a correction AT according to Tt, - AT = Tg„,. The

tenn AT is from Eq. 7-3, which can be rewritten as:

h h

(7-7)

Equation 7-7 illustrates the separation of the temperature correction into a flux term (Qi/h) and a radiative

loss term (aeTi/h). The temperature correction is an approximation since it uses the predicted flux from

the model simulations. Figure 7-8 shows the difference between the measured thermocouple bead

temperature, Ti„ and the corrected gas temperature, at outlet location 9 for Test 6. Here, the

corrected gas temperature differs from the measured bead temperature by up to 140 °C. This particular

example represents one of the largest temperature corrections for all of the data. Most temperature

corrections were 50 °C or less.

900

800

700

; 600

CDQ.E

<D 1- 500

400

300

' '

I —

WTC Phasel Test 6

Outlet 9

~i 1 1 1 1 ]

i

1

i 1

1 r Measured

Corrected

_1 I I I I L_ 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure 7-8. Comparison of the measured temperature and corrected temperature

in Test 6 at outlet 9.

Figure 7-9 is a plot of the magnitudes of the two components of the temperature correction from Eq. 7-5.

The flux component is larger than the radiative component and increases significantly as the test progresses, causing a greater temperature correction for the latter part of the test. Aspirated thermocouples were used in addition to the many bare-bead thermocouples because aspirated thermocouples have lower

measurement uncertainty. The expanded uncertainty in the aspirated thermocouple measurements was

approximately 7 °C (Blevins and Pitts 1999). Aspirated thennocouple measurements were taken at outlet location 5. Additional aspirated thermocouple measurements were attempted at outlet location 4, but due

to equipment failure measurements were not successful. Figure 7-10 compares the temperatures

measured with the aspirated thermocouple to the corrected and uncorrected temperature measurements

94 NISTNCSTAR 1-5B. WTC Investigation
Velocity Measurements

using the bare-bead thermocouple at outlet location 5 during Test 5. As expected, the temperatures

measured by the aspirated thermocouple were lower than those measured directly by the bare-bead

thermocouples. As shown in Fig. 7-10, the aspirated thermocouple temperature measurements were

within 50 °C of the bare-bead temperatures corrected for radiation. Data for the other five tests included

in Appendix D show that the temperatures recorded by the aspirated thermocouples were within

approximately 50 °C of the corrected temperamres for all of the tests. The value of these differences was

taken as the uncertainty in the bare-bead temperature measurements and is consistent with uncertainties

estimated for the thermocouple trees reported in Chapter 4.

300

250

3 200

03

L—I

150

I-

(D 2. 100 o<

WTC Phase 1 Test 6

Outlet 9 Temperature Correction

1600 2000 2400

Time (s) 3200

Figure 7-9. Comparison of the component of temperature correction due to incoming

flux and the component of temperature correction due to radiative loss.

850 —'—I —'—'

'

'—'

'—

I—I—I—I—I—I—I—I—I—I—^—I—

I—I

'—1—

I—I

'—I

t WTC Phasel Test 5

800 7 Outlet 5 Temperatures Uncorrected bare-bead -

Time (s)

Figure 7-10. Comparison of bare-bead temperature measurements

with aspirated temperature measurements in Test 5 at outlet 5.

NISTNCSTAR 1-5B. WTC Investigation 95
Chapter 7

The difference between the aspirated thermocouple reading and the corrected temperature is due to using an estimated value for the radiant flux when computing the corrected temperatures. As described in Section 7.1 .2, model predictions of the radiant flux on the thennocouple bead were used in the corrections

because actual flux measurements at these locations were not available. The corrected bare-bead

temperature measurements were used to compute the velocity because aspirated thermocouple data was

available at only one location. The 50 °C uncertainty in the temperature had only a small effect on the

velocity computations, as shown in Fig. 7-11. Here, the mean velocity computed using the corrected

bare-bead thermocouple temperature is compared with the mean velocity calculated using the aspirated

temperatures. The resulting difference in the mean velocity was less than 6 percent. The temperature

uncertainty contributed to the uncertainty of the velocity measurement as described in Sec. 7.2 and listed

in Table 7-2.

5

4

>.

"oo

2

>

1

0

0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure 7-11. Comparison of the mean velocity computed using bare-bead

temperature measurements with the mean velocity computed using

aspirated temperature measurements in Test 5 at outlet 5.

7.5 REFERENCES

Blevins, L.G., and W.M. Pitts. 1999. Fire Safety Journal, 33, 239-259, 1999; or see NISTIR 6310, April.

Available at: http://fire.nist.gov/bfrlpubs/fire99/PDF/f99022.pdf.

Kent, L.A., and M.E. Schneider. 1987. ISA Tramactiom, 26, 25-32.

Lm, C.Y., Y.W. Wong, W.K. Chan, and T.C. Gan. 1990. Experiments in Fluids, 9, 354-356.

McCaffrey, B.J. , and G. Heskestad. 1976. Combust. Flame, 26, 125-127.

' ' ' I' ' '

I

' ' '

I WTC Phasel Test 5

- Velocity Outlet 5

Corrected Bare-bead TC

Aspirated TC

J I 1 I .

I I I I i I ' I I I 1 ) Ii I I 1 I I I : i :I L

96 NISTNCSTAR 1-5B, WTC Investigation
Chapter 8

Compartment Surface Temperature Measurements

8.1 INSTRUMENTATION

The compartment surfaces were composed of two 1.27 cm sheets of calcium silicate board (referred to as Marinite) that were tightly held together by screws. The temperature of the iimer compartment surface was measured at select locations on the ceiling and walls using thermocouples. The exact position of the

thermocouples is given in the list of instrumentation (see Appendix B). Approximately 25 gauges were

located on the surface of the ceiling, while 6 gauges were located between the two 1.27 cm Marinite

layers in the ceiling.

The thermocouples were type K 30 gauge (0.25 mm wire diameter). The thermocouple beads were

created by remo\'ing approximately 1 cm of insulation from the Chromel and Alumel lead wires and spot

welding them together. The mean bead diameter was approximately two times the wire diameter. Each

thermocouple bead was attached to the Marinite surface by spring-loading it (bending it near the end

so the bead pushed against the surface) and lightly peening the bead into the surface. The insulated

thermocouple leads were attached to the Marinite by a staple approximately 2 cm from the bead.

Figure 8-1 is a photograph of a thermocouple mounted on the Marinite surface.

Source: NIST. Figure 8-1. A type K thermocouple

attached to the Marinite surface.

8.2 MEASUREMENT UNCERTAINTY

An uncertainty analysis associated with the thermocouple measurement of the Marinite surface

temperature is presented. The analysis is similar to that presented in Chapter 9 of this report on the

uncertainty in the measurement of the steel surface temperature using thermocouples. Two contributors

to the uncertainty were considered. The first is the inherent expanded uncertainty associated with a type K thermocouple, which is approximately 5 °C (Omega 2000). The second contributor is related to the systematic difference in the measurement affected by the presence of the thermocouple itself Although it may be possible to develop an analytic expression for the systematic measurement error as a function of time, this would necessarily entail the solution of a non-linear, unsteady problem, which is not

straight forward, or require the inclusion of possibly overly restrictive simplifying assumptions. Although

NISTNCSTAR 1-5B, WTC Investigation 97
Chapter 8

an analytic steady-state solution may be simple, the transient analysis presented here is not so. In

addition, the transient problem is the most important as it leads to a much larger systematic measurement

error than the steady-state solution.

The error in the measurement arises from the fact that the actual measurement location is at the

thermocouple bead and not at the Marinite surface. Of course, the surface temperature of the Marinite

influences the thermocouple bead temperature, but there is a difference between the temperature of the

bead and the surface due to the various heat transfer processes that occur. Several mechanisms contribute

to the temperature measured at the thermocouple bead, including radiative and convective heat transfer from the hot gas layer to the bead and lead wires of the thermocouple, as well as conductive transfer from

the bead to the Marinite, from the lead wires to the bead, and along the lead wires.

The systematic measurement error was considered using a finite-element code (ANSYS 2003) for the

simplified geometry presented in Fig. 8-2. An idealized spherical thermocouple bead was peened to the

surface of a flat Marinite slab taken to be 2.5 cm thick. Thennocouple lead wires were attached to the

bead and extend away from the Marinite surface. The lead wire diameter was 0.25 mm (30 gauge), and

the bead diameter was taken to be twice the diameter of the wire leads, consistent with measurements of

the thermocouple used in the experiment. One-half of the bead was embedded in the Marinite. Although

composed of two different metals, the thermocouple was assumed to consist of only one material to simplify the computation through application of symmetry conditions. Muhiple calculations, using each

of the component materials, were performed to assess the vahdity of this assumption.

Representative thermophysical properties of the materials comprising a Type K thermocouple were based on literature values (Omega 2000). Representative thermophysical properties for the Marinite were taken

from a testing report (Taylor 2003). The properties of interest at ambient temperature are listed in Table 8-1. The temperature dependence of the Marinite properties was incorporated into the analysis, but was neglected for the thermocouple materials. Replicate calculations were performed using a different set of thermophysical properties for the Marinite as suggested by the manufacturer, with negligible effect on

the results. The emissivities of the Marinite and thermocouple surfaces were taken to equal one, assuming

that the surfaces were rapidly covered with soot during the experiment.

98 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

Table 8-1. Thermophysical properties of Marinite, Chromel, and Alumel.

D6nsit\' Hpat Canacitv Conductivitv

Material (kg/m') (J/kgK) (W/m K)

Marinite'' 669 778-1.223 0.111-0.198

Alumel^ 8,600 523 29.7 Chromel^ 8.730 448 19.2

a. Taylor 2003.

b. Omega 2000.

The temperature of the thermocouple bead was taken as an average across the entire bead. The lead wires were taken sufficiently long so that the end condition could be assumed to be at ambient conditions

outside of the test compartment. The first 1 cm of lead wire from the thermocouple bead was left uninsulated as in the experiment, while the remaining length was covered with a perfect insulator. In the

calculations, heat conduction through the leads was balanced by conduction into the Marinite, and the

leads were heated by radiation.

To test the efficacy of the model developed here for the systematic measurement error, the calculation

results were compared to experimental measurements. Tests 2 and 5 were selected as representative of

the fires observ ed during the experimental sequence. Test 2 was a 2.4 MW heptane/toluene fire, whereas

Test 5 was a 3.0 MW heptane fire. Aspirated thermocouple measurements (see Chapter 4) were used to estimate the incident heat flux to the ceiling by assuming that heat transfer was dominated by radiation

from a sooty upper gas layer. Convection was neglected because heat transfer occurs predominantly by

radiation for gas temperatures of this magnitude. Reradiation from the exposed Marinite surface and the

thermocouple was considered. To obtain appropriate boundary conditions for the unexposed back surface

of the Marinite, a Biot number {Bi) was computed:

Because of the low thermal conductivity of the Marinite, Bi » 1, implying that convection of heat away

from the unexposed back surface provided a thermal resistance small compared with the heat conduction

through the Marinite. It was reasonable to assume a constant temperature on this surface, equal to the

ambient temperature.

The results of the simulation for Test 2 are presented in Figs. 8-3 and 8-4. Figure 8-3 plots the measured

aspirated gas temperature at location (604 cm, 53 cm, 365 cm), the measured ceiling surface

thermocouple temperamre at (612 cm, 60 cm, 382 cm), and predicted thermocouple temperatures for two

values of the gas emittance (Siegel and Howell 2002). For an emittance of 0.6, the predicted

thermocouple temperature resembles that observed during the experiment at early times. As the test proceeded, the themiocouple temperature was underpredicted. For a gas emittance of 0.8, the predicted

thermocouple temperature was high for early times, but the data late in the test were more closely

predicted. These two values provide upper and lower bounds for the upper gas layer emittance during

Test 2.

NISTNCSTAR 1-5B, WTC Investigation 99
Chapter 8

600 7

Gas Temperature

• Measure thermocouple

Predicted thermocouple,

emittance = 0.6 —Predicted thermocouple,

emittance = 0.8

100 200

Time ( s )

300 400

Figure 8-3. Comparison of the measured and predicted ceiling surface thermocouple

temperature in Test 2. The measured gas temperature is also shown.

Oo

CD

Q.

E

CD

100

75

50

25

emittance - 0.6

emittance = 0.8

100 200 300

Time ( s )

400

Figure 8-4. The simulated thermocouple temperature minus the simulated Marinite

surface temperature for two values of the upper layer emittance in Test 2.

Figure 8-4 presents the uncertainty in the measurements by plotting the difference between the predicted

thermocouple temperature and the predicted Marinite surface temperature for bounding values of the

upper layer emittance. At early times, there is a significant difference between the two temperatures, on

the order of 75 °C to 85 °C. This difference decayed as the test proceeded until it obtained a value of

nearly 35 °C when the fuel flow was stopped. The reason for the temperature difference was due to the

heat that was rapidly transferred to the bead early in the test, which was unable to rapidly diffuse into the

100 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

Marinite because of its highly insulating properties. The relatively large thermal mass of the Marinite (as

compared to the thermocouple) exhibited a slower heating and, thus, there was a lag in the Marinite

surface temperamre as compared to the thermocouple.

Figures 8-5 and 8-6 present the results for simulation of Test 5. In Fig. 8-5, the measured gas

temperatures, the measured thermocouple temperatures, and the predicted thermocouple temperatures are

plotted. Predicted thermocouple temperatures for a gas layer emittance of 0.8 and 1 .0 provide upper and

lower bounds for the upper gas layer emittance. The 0.8 and 1.0 emittance values adequately match the measurements for early and later times in the experiment, respectively.

Predicted thermocouple,

emittance = 0.8 Predicted tliermocouple, emittance = 1 .0

600 1200 1800

Time ( s )

2400 3000

Figure 8-5. Comparison of the measured and predicted ceiling surface thermocouple

temperatures for Test 5. The measured gas temperature is also shown.

120

•emittance = 0.8

emittance = 1 .0

I

1200 1800 Time ( s )

2400 3000

Figure 8-6. The simulated thermocouple temperature minus the simulated Marinite

surface temperature for two values of the upper layer emittance in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 101
Chapter 8

Figure 8-6 plots the difference between the thermocouple and Marinite surface obtained from the

simulations. As in Test 2, significant differences are observed early in the experiment, with values on the

order of 100 °C. The differences are greater for Test 5, as the fire was more intense and the upper layer

temperatures were larger. Again, there is a gradual decay in the difference as time progresses, obtaining a value of 13 °C after long times. This difference is due to the competition between the various heat

transfer mechanisms that are present. Because the Marinite has one surface exposed to the ambient, there

will be some conductive heat transfer across its thickness. Conservation of energy dictates that the

radiative loss from the Marinite surface must be less than the net incident heat flux to account for this conduction, lowering the exposed surface temperature of the Marinite. A similar mechanism is present

for the thermocouple, but because it is in contact with the hot surface of the Marinite and because of its small mass, the effect will be significantly less. This disparity results in the steady state temperature

difference.

In summary, Marinite surface thermocouple measurements were found to possess a high degree of

systematic error during initial phases of the fire exposure due to the relatively large insulating properties

of the Marinite and the low thermal mass of the thermocouple. The error diminished as steady state conditions were approached. After long times, the error was as httle as 13 °C. If the systematic error is treated like a measurement uncertainty, then the total expanded uncertainty is bounded by -8 °C and

-18 °C. For times at which there are large temporal changes in the incident flux, it is difficult to quantify

the systematic error due to uncertainty in the boundary conditions. This is exemplified by the significant

differences between the model and measurements seen for early times in the experiment in Fig. 8-3.

Therefore, tests of model predictions for the Marinite are most meaningful at later times in the

experiments, when systematic measurement error was small. This will be considered in Chapter 11, when

measurement results are compared to simulations.

8.3 DISCUSSION OF RESULTS

Figure 8-7 shows the measured ceiling temperature as a function of time at 15 locations along the

centerline (x axis) of the compartment during Test 1 . The temperatures rapidly rose from ambient to greater than 100 °C within 15 s. The temperatures varied as a function of location on the ceiling. After

the rapid initial rise, the temperatures increased in a near-linear fashion until the fuel flow was closed, and

the measured temperatures dropped precipitously. Data were acquired for a significant period after the

fuel flow was stopped. This information may also be useful for validation of heat transfer sub-models

without the complications of the presence of a fire plume. The scatter in the data varied from 3 °C to 30 °C and was largest for positions above the fire and smallest for locations away from the fire. Regular

fluctuations were observed for a few locations that were affected by the Ni purging of nearby total heat

flux gauges (see Chapter 5). This was particularly noticeable after the fire was stopped.

Figure 8-8 shows the temperatures between the Marinite layers as a ftjnction of time during Test 1 . The

temperatures measured between the Marinite layers obtained a maximum about 160 s after the fuel flow was stopped (at 860 s), driven by conduction from the relatively high, but falling surface temperatures even after the fire was stopped. The temperatures varied as a function of location, but were much lower

than the surface temperatures.

102 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

Oo

a3

500 -

400 -

300 -

c 200 -

100 -

x=228 cm

x=279 cm

x~oU3 cm

x=36P cm, x=399 cm

x=414 cm

x=444 cm

x=466 cm

x=477 cm

x=492 cm

x=507 cm

x=552 cm

(TC95)

{TC93)

(TC92)

fTC91 '< (TC90)

(TC88)

(TC86)

(TC66)

(TC71

)

(TC72)

(TC73)

(TC76)

x=612 cm (TCSOj

500 1 000 1 500

Time (s)

2000 2500

Figure 8-7. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 1.

250

I—

I—I—I—I—I—

I—I—I—I—I—

I—r—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

0 r I I I I i I I I I I I I I I I

I I I I I I I I I I I I L 0 500 1000 1500 2000 2500

Time (s)

Figure 8-8. The measured temperatures between the two Marinite sheets in the

ceiling along the compartment centerline (x axis) as a function of time at several locations during Test 1.

NISTNCSTAR 1-5B, WTC Investigation 103
Chapter 8

Figure 8-9 uses a portion of the temperature data shown in Figs. 8-7 and 8-8. Figure 8-9 shows the peak

ceihng temperatures as a function of location on the centerline (x axis) 860 s after fire initiation, when the

temperatures were at their peak values and the fuel flow was stopped. The highest temperatures were

measured almost above the fuel pan, extending from x = 342 cm to 442 cm (see Fig. 2-5). Figure 8-9

also shows the peak temperature measured between the Marinite layers 1 ,020 s after fire initiation. The

value of the peak temperatures decreased with distance away from the fiiel pan, following the ceiling

surface temperatures.

ai

CTJ ^ 350

I- 300

CT5 250

300 400

X(cm)

700

Figure 8-9. The peak temperatures on the ceiling surface and sandwiched between

Marinite ceiling layers along the compartment centerline (x axis) during Test 1.

The ceiling surface temperature also varied along the y axis, in the north-south direction, due to the

leaning of the fire toward the south wall. The reasons for asymmetry of the fire may be attributed to asymmetry in the placement of compartment contents such as the column and the instrumentation

support, which were both located on the northern half of the compartment. Figure 8-10 shows the time

varying ceiling surface temperatures for various locations on the y axis at x = 456 cm during Test 5. Similar results were seen in the other tests. The peak values of the ceiling surface temperature at X = 456 cm (shown in Fig. 8-10 at 2,950 s) are re-plotted in Fig. 8-1 1 as a function of y location. The

results at X = 612 cm are also shown. Both sets of measurements are consistent with the observation that the flame was leaning toward the south wall. Similar results were obtained for the other tests.

Figures 8-12 through 8-21 show the peak surface and middle ceiling temperatures as a function of time

after fire initiation for Tests 2 through 6. The plots qualitatively resemble Figs. 8-7 and 8-8. The highest

peak temperatures were attained during Tests 5 and 6, when the fire duration was approximately 50 min,

the longest of all of the tests.

104 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

O

1000

800 -

600 -

^ 400 -'

200

2000 3000

Time (s)

5000

Figure 8-10. The measured ceiling surface temperatures as a function of time at seven

locations along the y axis at x = 456 cm during Test 5.

3 600

oSa.

m

03 Q_

1000

800

400

200

-|—I—I—I—I—

I—I—I—I—I—I—I—I—I—I—I—I—I—i—I—I—I—I—

I—1—1—I—I—I—

I—I—1—I—I—

Test 5

-•--X = 612 cm

-©—X = 456 cm

0 1—1—I I

I I 'I I 1—1 I I I I I I I I I I I I I I I I I I I 1 I—I 1 I -90 -60 -30 0 30 60 90

y (cm)

Figure 8-11. The measured peak ceiling surface temperatures as a function of location

along the y axis at two x locations (x = 456 cm and x = 612 cm) during Test 5.

NISTNCSTAR 1-5B, WTC Investigation 105
Chapter 8

Oo

x=228 cm (TC95)

x=279 cm (TC93)

x=309 cm (TC92)

x=^369cm (TC91)

x=399 cm (TC90)

x^414 cm (TC88'

x=444 cm (TC86;

x~'456

x=477 cm (TC71

)

x=492 cm (TC72)

x=507 cm (TC73;

=522 cm (TC75;

x=552 cm (TC76)

1000

Time (s

2000

Figure 8-12. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 2.

0)

200

150

Oo

0) S 100 - CO

C2.

50 -

x=399 cm (TC89)

x=412 cm (TC65)

x=522 cm (TC74)

x=612 cm (TC79)

500 1000

Time (s)

1500 2000

Figure 8-13. The measured temperatures between the two Marinite sheets in the

ceiNng along the compartment centerline (x axis) as a function of time at several

locations during Test 2.

106 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

Oop

500 -

400

300

200 -

100 -

x=228 cm (TC95)

- x=279 cm (TC93)

• x^309 cm (TC92)

x=363 cm (TC91

- x=399 cm (TC90)

x=414 cm {TC88)

x=444 cm (TC86)

x=477 cm (TC71

- x=492 cm (TC72)

- x=507 cm (TC73)

x=i>z2 cm \ \\.,>.'oi x=552 cm (TC76)

1 1 1 1 11

1 1 1 1 .

1 1 1 1 1 1

1 1 1 1

1

1 1 1 1

1

1 1

^7

^^^^^

:f

^^^^^^ x=61 2 cm (TC80)

r ~\

" i

_ f

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 500 1000 1500 2000

Time (s)

2500 3000 3500

Figure 8-14. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 3.

oo

05

It— Qia.£

03

250 T—I—I—I—

I—I—I—I—1—

I—I—I—I—I—I—

200 -

150 -

100 -

50 -

x=399 cm (TC89) M

x=412cm (TC65) M

x=522 cm (TC74) M

x=612 cm (TC79;

500 1000 1500 2000

Time (s)

2500 3000 3500

Figure 8-15. The measured temperatures between the two Marinite sheets in the

ceiling along the compartment centerline (x axis) as a function of time at several

locations during Test 3.

NISTNCSTAR 1-5B, WTC Investigation 107
Chapter 8

800

700

600

o

500

\u

-•— (C 400

CD

£3. E 300

(D 1- 200

100

0

-p-i—I—I—r-|—i—j—rn—

|—rn—i—i—|—i—i—i—r-|—i—i—i—

r

Test 4

x==228 cm (TC95)

- x==279 cm (TC93)

X-^369 cm {TC91

)

x==399 cm (TC90)

x=-414 cm (TC88)

x==444 cm (TC86)

X-"456 cm ( 1 C66)

x==477 cm (TC71)

x==492 cm (TC72)

X-=507 cm (TC73)

x==552 cm (TC76)

cm fi'G/V •

x==612 cm (TC80)

I I i__L J L_l L _l_l I I l_l I 1_L. 0 500 1000 1 500 2000

Time (s)

2500 3000 3500

Figure 8-16. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 4.

400

oo

03a.E

0)

1500 2000

Time (s)

3500

Figure 8-17. The measured temperatures between the two Marinite sheets in the

ceiling along the compartment centerline (x axis) as a function of time at several

locations during Test 4.

108 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

Oo

CO

5

1000

800 -

600

400

200

1000 2000 3000

Time (s) 4000 5000

Figure 8-18. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 5.

oo

0)

£2. £

0)

2000 3000

Time (s)

5000

Figure 8-19. The measured temperatures between the two Marinite sheets in the

ceiling along the compartment centerline (x axis) as a function of time at several

locations during Test 5.

NISTNCSTAR 1-5B, WTC Investigation 109
Chapter 8

O

53

<x<

1000

800 -

600

^ 400

200

"I—I—I—I—

I—I—I—I—I—I—I—I—I—

r

Teste , ,

1—I—I—

I—I—1—I—

r

x==228 cm (TC95)

x==279 cm (TC93)

x~-30y cm iTu92)

x==369 cm (TC91

)

x==399 cm (TC90)

x==414 cm (TC88)

x==444 cm (TC86)

x==477 cm (TC71

)

AYx=^607 cm (TC73}

=552 cm (TC76)

X:^^12 cm rJCSO)

0 t2 I I I I I 1 I I I I I 1 I I I I I I I I I I I I I L 0 1000 2000 3000 4000 5000

Time (s)

J I I 1

I I 6000

Figure 8-20. The measured ceiling temperatures as a function of time at 15 locations

along the compartment centerline (x axis) during Test 6.

1 I I I I r I II I I

o

x=399 cm (TC89)

x=412 cm (TC65)

x=522 cm (TC74i

x=612 cm (TC79)

2000 3000 4000

Time (s)

Figure 8-21. The measured temperatures between the two Marinite sheets in the

ceiling along the compartment centerline (x axis) as a function of time at several

locations during Test 6.

110 NISTNCSTAR 1-5B. WTC Investigation
Compartment Surface Temperature Measurements

8.4 REPEATABILITY

Figure 8-22 compares the temperatures measured along the compartment centerline (x axis) on the ceihng

surface and sandwiched between the two Marinite layers, 3,000 s and 3,080 s after fire initiation,

respectively, during Tests 5 and 6, which represent repeat experiments. These times were selected to compare the tests because the duration of Test 6 was 50 min (3,000 s), which was slightly shorter than

Test 5. The fuel used was heptane, and the heat release rate was 3 MW in both experiments. The

repeatability of the ceiling temperature measurements should be considered in terms of the repeatability

of the gas-phase temperamres in the upper layer of the compartment in Tests 5 and 6 (see Chapter 4),

which drive the values of the compartment ceiling temperature. The average value of the expanded

uncertainty limits in the gas-phase temperature was 21 °C and 23 °C from Table 4^ (Chapter 4) for Tests 5 and 6, respectively. The average of the absolute value of the difference between the measured

ceiling temperatures in Tests 5 and 6 was approximately 25 °C for the surface and 24 °C between the

Marinite layers. These differences were within the combined expanded gas-phase uncertainties. These

results in concert with the measurement uncertainty described in Sec. 8.1 above demonstrate that the

ceiling temperature measurements were repeatable within experimental uncertainty.

Oo

5

1000

800

600

400

200 -

-|—I—I—I—

I—I—I—I—

• Test 5 o Test 6

m Test 5 M

n Test 6 M

T—I—I—I—

I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—

I—I—I—

O

c

n

J I I I I 1_ 100 200 300 400

X(cm)

500 600 700

Figure 8-22. Comparison of the temperatures on the ceiling surface and sandwiched

between the two Marinite ceiling layers 3,000 s and 3,080 s after fire initiation,

respectively, along the compartment centerline (x axis) during Tests 5 and 6.

8.5 EFFECT OF HEAT RELEASE RATE

Figure 8-23 compares the temperatures measured along the compartment centerline (x axis) on the ceiling

surface and sandwiched between the Marinite layers, 830 s and 1,035 s after fire initiation, respectively,

during Tests 1 and 5. These times were selected to compare the tests because the duration of Test 1 was

only 14 min (840 s) as compared to 50 min for Test 5. Tests 1 and 5 were fires burning heptane with

Test 1 having a heat release rate of 2 MW as compared to 3 MW in Test 5. As expected, the temperatures were larger in the higher heat release rate fire.

NISTNCSTAR 1-5B, WTC Investigation 111
Chapter 8

o

CDa.E

800

700

600

500

400

300

200

100

T—i—I—I—

I—I—I—I—I

I

I—I—I—I—

I—I—I—I—I—

I—I—I I A, K' '

'—I

I

I—TT—

T

Surface

O CD Q3^-0 O

o o

• Test 1 o Test 5 m Test 1 M

a Test 5 fv1

H

n D

Sandwiched

II 1 1 I I III J I I I i_l I I I

I I I I I II i_ _1 I

I 1I I L. 100 200 300 400

X(cm)

500 600 700

Figure 8-23. Comparison of the temperatures on the ceiling surface and sandwiched

between the two Marinite ceiling layers 830 s and 1,035 s after fire initiation,

respectively, along the compartment centerline (x axis) during Tests 1 and 5.

Figure 8-24 compares the temperatures measured along the compartment centerline (x axis) on the ceiling

surface and sandwiched between the two Marinite layers, 350 s and 660 s after fire initiation,

respectively, during Tests 2 and 3. These times were selected to compare the tests because the duration of

Test 2 was 6 min (350 s), which was shorter than Test 3. Tests 2 and 3 were fires burning the

heptane/toluene fijel mixture with Test 2 having a heat release rate of 2.4 MW as compared to 2.0 MW in Test 3. As seen in Fig. 8-24, the ceiling temperatures were larger in the higher heat release rate fire. As

expected, the temperatures between the Marinite layers had smaller measured differences than the surface

temperatures.

700

o

0)

E2. E

600

500

400

300

200

100 -

1 1 1 1

I 1 1 1 1 1 • Test 2 o Testa

Test 2 M

n Test 3 M

Surface <^

O ^ o

Sandwiched Cl

f5 U

1

I I I I )I I 1 I [ 1 1 [ I I I I I I I I I I 1 I I I 1 I I I I 1 1 100 200 300 400

X (cm)

500 600 700

Figure 8-24. Comparison of the temperatures on the ceiling surface and sandwiched

between the two Marinite ceiling layers 350 s and 660 s after fire initiation,

respectively, along the compartment centerline (x axis) during Tests 2 and 3.

112 NISTNCSTAR 1-5B, WTC Investigation
Compartment Surface Temperature Measurements

8.6 EFFECT OF FUEL TYPE

Figure 8-25 compares the temperatures measured along the compartment centerline (x axis) on the ceiling

surface and sandwiched between the two Marinite layers, 830 s and 1,040 s, after fire initiation,

respectively, during Tests 1 and 3. The first time (830 s) was selected to compare the ceiling surface

temperature for the two tests because the duration of Test 1 was 840 s (14 min), which was shorter than

Test 3. The second time (1.040 s) was selected to compare the sandwiched temperatures. Both tests had

heat release rates of 2 MW. The fuel was heptane in Test 1, whereas it was a mixture of heptane and

toluene in Test 3. The difference in the measured ceiling temperatures between Tests 1 and 3 was about

7 °C, consistent with the small temperature differences between Tests 1 and 3 observed for the East tree gas-phase thermocouple measurements for these two tests shown in Fig. 4-1 1 . These results show that

fuel type had little influence on the ceiling temperatures.

600

500

400

300

—1—I—I—I—

I—I—I—I—r—

I—1—I—I—I—I—I—I—I—I—

I—I—I—I—I—

I—I—1—I—I—

I—I—I—I—I—]

o

oo

0}

0)

£2. 200 • Test 1 o Test 3

100 Test 1 M

n Test 3 M

± I I I

I t I I 100 200 300 400

X (cm)

500 600 700

Figure 8-25. Comparison of the temperatures on the ceiling surface and sandwiched

between the two Marinite ceiling layers 830 s and 1,040 s after fire initiation,

respectively, along the compartment centerline (x axis) during Tests 1 and 3.

8.7 REFERENCES

ANSYS, Inc. 2003. ANSYS Release 8.0 Documentation.

Omega Engineering, Inc. 2000. Omega Temperature Handbook, Technical Reference Section,

Section Z.

Siegel, R., and J.R. Howell. 2002. Thermal Radiation Heat Transfer, Taylor and Francis, NY.

Taylor, R.E., H. Groot, and J. Ferrier. 2003. Thermophysical Properties ofPVC, PE, and Marinite,

Report No. TPRL 2958, Thermophysical Properties Research Laboratory, Inc., April.

NISTNCSTAR 1-5B, WTC Investigation 113
Chapter 8

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114 NISTNCSTAR 1-5B, WTC Investigation
Chapter 9

Temperatures of Steel Components

9.1 OBJECTIVE

The time \ arying temperatures of the steel components were measured to determine the thermal response

of structural steel components to controlled room fire conditions, including flame immersion, and to establish a data set to validate predictions of the temperature rise of insulated and uninsulated structural

steel components. The components were tested in an uninsulated (unprotected bare) state in Tests 1, 2, and 3 and with an insulating coating of fireproofmg, or fibrous sprayed fire -resistive material (SFRM), in Tests 4, 5, and 6. The SFRM was applied in two nominal thicknesses: 17 mm or 34 mm. The steel components were not loaded or tested for strength; these tests were conducted primarily to measure the

heat transfer to the steel. Chapter 1 1 compares the finite-element analysis predictions of the steel temperatures with the measurements presented here.

This chapter contains a description of the experimental measurements, an analysis of uncertainty,

tabulation of key results, a statistical summary of the applied SFRM thickness on the steel components,

and representative plots of the experimental results, including thermocouple measurements of the surface

temperature of the bare and coated steel components during Tests 3 and 5. A complete record of the exact

locations of the thermocouples on the steel components and the SFRM coating is documented in Appendix F. The complete record of time dependent measurements for Tests 1, 2, 3, 5, and 6 is presented

in Appendix G. Test 4 was truncated after approximately 15 min due to a malfunction of the

thermocouples that occurred a few minutes after the start of the test. As no useful steel temperature data was obtained during Test 4, it is not included in this chapter or Appendix G.

9.2 EXPERIMENTAL DESCRIPTION

Steel components were selected to meet several criteria:

• Section thicknesses were sized so that 600 °C steel temperatures could be obtained with a fire heat release rate ( 0 of 2 MW to 3 MW in less than an hour,

• Components were sized to be similar to some of those in World Trade Center (WTC) 1 and

WTC 2,

• A simplebar was included as a baseline case for analysis purposes, and

Various types of structural components were selected to measure the heating effect of fires on

horizontal and vertical steel components.

Three types of steel components were selected, a bar, a hollow steel tubular column, and a bar joist truss, and are shown in Figs. 9-1 through 9-3. The bars were nominally 3 m long ( 10 ft) and 25 mm ( 1 in.) in diameter; the columns were 0.26 m (10 in.) by 0.36 m (14 in.) tubular steel sections with a nominal 6 mm

( % in.) wall thickness; and the trusses were 4.6 m (15 ft) long and 0.8 m (2.6 ft) deep with 64 mm to

NISTNCSTAR 1-5B. WTC Investigation 115
Chapter 9

76 mm (2.5 in. to 3 in.) double angles for the top and bottom chords and 25 mm (1 in.) bar for the web

members.

Figure 9-1. Bar dimensions.

3.35 m

6.31 mm

6.25 mm

T

0.36 m

0.26 m

1

Figure 9-2. Tubular column dimensions.

4.6 m

0.8 m

Web Bar

25.6

mm

Top Chord Angles

6.4 mm 76 mm

Bottom Chord Angles

6.4 mm—

<

6.5 mm I

76 mm 6.3 mm

I

64 mm

^—

64 mm

Figure 9-3. Truss dimensions.

116 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Bare-bead Type K 30-gauge thermocouples (TC) were selected to record the temperature time-histories at the steel and SFRM surfaces. Prior to application of SFRM to the steel components for Tests 4, 5, and 6, the thermocouple beads were spot welded to the bare steel surface of all components. The thennocouple

locations on the bar, column, and truss components are illustrated in Figs. 9-4a, 9-5a, 9-5b, 9-5c, 9-6a,

and 9-6b. The thermocouples that were to be positioned just under the SFRM surface were placed

against the steel surface (under the SFRM) and supported on a wire tie at the intended thermocouple

location. The wire tie minimized disruption to the SFRM after it cured and allowed sufficient

thermocouple wire length for inserting the thermocouple bead just under the SFRM surface by hand

while minimizing exposure of the thermocouple wire to the gas temperatures. Figure 9-6c shows a thermocouple wire supported on a wire tie, with the bead placed just under the SFRM surface. It is estimated that these thermocouple beads were within 1 mm to 5 mm of the surface. The wire tie was

not considered to be a significant source of heat transfer to the steel since the wire only contacted the

steel at occasional spots or comers where it was wrapped and twisted to provide support for the

thermocouple wire.

The SFRM product used in the WTC and these experiments was BLAZE-SHIELD DC/F, which was

manufactured and supplied by Isolatek International. BLAZE-SHIELD DC/F is a sprayed single-package

factory controlled premix, based on a mixture of mineral wool fibers and cement binders. The

manufacturer reported average density is 208 kg/m^ (13 Ib/ft^) minimum and its thermal conductivity is 0.042 W/m-K at 24 °C. These values were confirmed by a testing laboratory, which reported a density of

240 kg/m- and a thermal conductivity of 0.046 W/m-K (NIST NCSTAR 1-6A). Uncertainty in the

thermal properties impacts the prediction of the thermal-structural response of the insulated steel components. The SFRM was applied by an experienced applicator from Isolatek. A nominal thickness of

SFRM was specified for each steel component, either 17 mm or 34 mm. The applicator took considerable

care to apply an even coating of the specified thickness of SFRM to the steel components. Figures 9-6d

and 9-6e show the SFRM application process and steel components after the application. The 34 mm

SFRM was applied in two coats. The first coating adhered to the bare steel and the second coating was

applied the following day to obtain the required thickness. The SFRM was then allowed to dry in place

for four weeks, as required by the manufacturer, before proceeding with thickness measurements,

placement of thermocouple beads just under the SFRM surface, and verifying the thermocouple locations

and function prior to testing.

A pin thickness gauge, specifically designed for SFRM products, was used to measure the applied

coatings. This type of gauge can measure to the nearest 1 min; its measurement accuracy is 0.5 mm.

Figures 9^b, 9-5b, and 9-6f show the location of the thermocouples and SFRM thickness measurements

for steel components used in Tests 5 and 6. (The steel temperature data collected during Test 4 was

disregarded due to problems with the thermocouple signals during the test, so the thermocouple and

SFRM data are not presented.) The SFRM measurement statistics are summarized in Table 9-1. The

SFRM mean thickness, was greater than the specified nominal thickness by 20 percent to 35 percent

typically, but ranged in value from 5 percent to 42 percent greater. The standard deviation, a, assuming a normal distribution of thicknesses, ranged from 3 m to 8.2 m. The standard deviation was influenced by

the steel shape and the lumpy nature of the fibrous SFRM; a smooth surface cannot be obtained when the

product is sprayed (no tamping was allowed). The coefficient of variation (COV), where COV = a/fi, ranged from 0.17 to 0.27 for the bars, columns, and trusses, with the exception of the column in Test 5 which had a COV = 0.07. This lower COV occurred for the condition in which a thicker SFRM coat was

NIST NCSTAR 1-5B, WTC Investigation 117
Chapter 9

applied to a smooth steel surface. The variation in SFRM thickness for these steel components should be

considered as a lower bound of variance that would be expected for field applications.

30 cm

B1

122 cm

B2

183 cm

B3

274 cm

84

(South) C

TCs were placed facing downward toward the floor. Bar ends were used for handling and wrapped with Kaowool batting for testing.

(North)

Figure 9-4a. Thermocouple locations for bars.

la 1am 2a 2am 3a

i 1 i i i

30 cm 122 cm 183 cm

(8> (^—^ —

3am 4a

i i

274 cm

-0 ^ ^

"-"J V - — „— .V. . . ^ ............ .. : 1, . . ... , .^.^ ^ —^ ^—0— ^

t t tit t

lb Ibm 2b 2bm 3b 3bn 4b

SFRM

SFRM thickness was measured for two locations at each TC and at midpoints between the

TCs.

Figure 9-4b. SFRM measurement locations for bars.

118 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

369 cm

213 cm

77 cm

TC Labels

CU (North)

CU4-S

CU1-S

CM (West)

CU2-S

(East)

CL

CU3-S

(South)

TCs were placed on each face at the three elevations

shown. TC locations are unique for Test 5 as shown.

Figure 9-5a. Thermocouple locations for columns.

335.3 cm

320.0 cm

167.6 cm

30.5 cm

0 cm _

Measurement

Locations*^)

CU

C\Jr

TC

lb \ 1 la

J8l

SFRM

CM

CMn

CL

4a (g

4 (2

2b

i8r

3a

Measurements were taken for the TC locations for Test 5 as shown.

Figure 9-5b. SFRM measurement locations for column in Test 5 and 6.

NISTNCSTAR 1-5B, WTC Investigation 119
Chapter 9

Source: NIST. Figure 9-6b. Thermocouple bead spot welded onto the bottom chord of a

steel truss with the wire running close to the steel surface.

120 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Source: NIST. Figure 9-6c. Color photograph of a thermocouple wire supported on a wire tie, with the

thermocouple bead placed just under the SFRM surface.

Source: NIST. Figure 9-6d. Application of SFRM product to a truss.

NISTNCSTAR 1-5B, WTC Investigation 121
Source: NIST. Figure 9-6e. Steel components with SFRM product prior to testing.

122

Figure 9-6f. SFRM measurement locations for trusses.

NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Table 9-1. Statistical summary of SFRM thickness for steel components

in Tests 5 and 6. Applied Thickness

Specified Standard

Thickness IV^ n n n

iviean Deviation

Component (cm) (cm) (cm) COV

Test 5 Bar 1.91 2.30 0.55 0.24 Column 3.81 4.10 0.30 0.07

Truss A 1.91 2.69 0.73 0.27

Truss B 3.81 4.05 0.82 0.20 Test 6 Bar 1.91 2.53 0.46 0.18 Column 1.91 2.14 0.35 0.17

Truss A 1.91 2.60 0.69 0.26

Truss B 1.91 2.56 0.69 0.27

9.2.1 SFRM Thickness

The procedures to measure and evaluate SFRM thickness followed ASTM E 605-93, the Standard

Test Method for Thickness and Density of Sprayed Fire-Resistive Material (SFRM) Applied to Structural

Components, which is reproduced in italics below.

Selected areas to be measuredfor thickness shall be a predetermined, repetitive pattern to ensure

obtaining representative average thickness. Determine the thickness by inserting the penetrating pin of

the thickness gauge perpendicular to and through the SFRM. to the substrate. When the point ofthe pin

touches the substrate, move the sliding disk to the SFRMsurface with sufficientforce on the disk to register the average plane ofthe surface. Withdraw the gauge to read the thickness in approximately

1 mm (1/16 in.) increments as shown by the position ofthe sliding clip indicator.

For the purpose ofaveraging measurements, any measurement 6 mm (1/4 in.) or more, over the required

design thickness, shall be recorded as the design thickness plus 6 mm. No individual measured thickness

shall be more than 6 mm less, or more than 25 percent less, than the required design thickness.

Specific fire resistance rating criteria for beams, trusses, and columns may allow for a reduced thickness on flange tips. These thicknesses are to be averaged apartfrom other sections ofthe structural member.

An item (deck, column, beam, orjoist) is deficient if:

• An individual measured thickness is more than 6 mm (1/4 in.) less, or more than 25 percent

less, than the requiredfire resistance design thickness.

• The calculated average thickness ofthe SFRM is less than that required by the design:

NISTNCSTAR 1-5B, WTC Investigation 123
Chapter 9

The applicable building code governs. Consult the applicable building codefor exact requirements and

tolerances.

The International Building Code (IBC 2000) and the National Fire Protection Association in NFPA 5000

(NFPA 2003) both reference ASTM E 605 for determining adequate SFRM thickness. IBC 2000 further

states that the minimum allowable thickness for a design thickness greater than or equal to 2.5 cm (1 in.)

is the design thickness minus 0.6 cm (% in.). Similarly, for a design thickness less than 2.5 cm (1 in.), the

minimum allowable individual thickness is 75 percent of the design thickness.

Applying the IBC criteria to SFRM measurements that were less than the specified thickness would give

a minimum allowable thickness of 1.43 cm for the specified 1.91 cm coating and a minimum allowable

thickness of 3.21 cm for the specified 3.81 cm coating. Individual thickness measurements are shown in Tables 9-2 through 9-7. Values that were less than the specified nominal thickness, but greater than the

minimum allowable thickness, are highlighted in the tables. Bolded values were less than the minimum

allowable thickness.

Tab e9-2. SFRMt lickness for bars in Tests 5 and 6. Test 5 Bar

(cm) a b bn.

Test 6

Bar (cm) a b bm

Specified 1.91 cm Specified 1.91 cm

Bl 3.02 2.70 1.91 2.54 Bl 1.91 3.18 2.06 X B2 2.70 2.70 2.86 3.02 B2 2.86 2.70 2.54 X 83 1.91 2.06 1.91 1.75 B3 2.22 3.18 2.06 X B4 1.91 1.27 NA NA B4 2.22 2.86 NA NA

GOV = 0.24; |a = 2.30 cm; o = 0.55 cm GOV = 0. 18; |i = 2.53 cm; o = 0.46 cm

Key: "x" indicates that measurements were not taken; NA, not applicable. Note: Bold measurements were one standard deviation or more below specification; highlighted measurements were below specification.

124 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Table 9-3. SFRMthickness for columns in Tests 5 and 6. TestS

Column

(cm) la lb Im

Test 6

Column

(cm) la lb Im

Specified 3.81 cm Specified 1.91 cm

PI I 1 V_ (J 1 4.29 4.29 4.29 PI 1 1 1.91 1.59 2.54

4.44 4.13 4.06 PI 1 0 2.86 2.70 1.91 PI I ^ 4.13 4.13 3.81 PICU1 1D 2.54 1.91 1.91 PI I A 4.44 3.96 4.44 PICU1 /I4 2.54 1.91 1.91

PN/1 1 4.13 4.29 4.13 p\/1 1 ClVl 1 1.91 2.54 2.22 C\A 1 y^lvl L 3.81 4.44 3.81 CiVl / 2.54 1.91 2.06

4.44 4.44 4.13 CM i 1.43 1.75 1.91

4.44 4.06 4.13 CM 4 2.54 1.91 2 22

CL 1 4.06 3.33 X CL 1 1.75 1.91 X CL2 3.81 4.13 X CL2 2.38 2.06 X CL3 4.13 3.18 X CL3 2.38 2.54 X CL4 3.96 4.06 X CL4 1.91 2.38 X COV = 0.07; ^1 = 4.10 cm; a = 0.30 cm COV = 0.17; ^ = 2.14 cm; a = 0.35 cm

Key: "x" indicates that measurements were not taken. Note: Bold measurements were one standard deviation or more below specification; Highlighted measurements were

below specification.

Table 9-4. SFRM thickness for truss A In Test 5. Tests

Truss A

(cm) a b c d e f o

t> h

Specified 1.91 cm TU 1 3.02 2.54 3.49 3.18 3.02 2.54 3.02 3.18

2 2.54 2.38 2.06 3.81 2.54 2.54 2.54 3.18

3 1.27 2.54 3.33 3.33 2.70 2.54 1.91 2.70

4 2.70 2.86 1.75 2.54 1.75 3.18 2.38 TM 1 2.86 2 22 1.75 1.75 2.22 1.27 NA NA

2 2.86 2.86 2.54 1.91 2.54 2.70 NA NA

1.91 1.91 2.54 1.75 3.81 3.81 NA NA

4 2.08 3.81 3.81 2.70 2.72 5.08 NA NA

TL 1 2.86 2.70 3.49 4.13 2.22 2.06 3.18 1.91

2 3.81 3.81 1.27 1.75 2.54 3.18 3.02 3.18

3 3.02 2.86 1.91 3.18 3.18 1.91 3.33 4.13

4 3.18 1.59 1.59 1.91 2.54 2.54 3.02 3.18 COV = 0.27; n = 2.69 cm; o = 0.73 cm

Note: Bold measurements were one standard deviation or more below specification; highlighted measurements were below specification.

NISTNCSTAR 1-5B, WTC Investigation 125
Chapter 9

Table 9-5. SFRM thickness for truss B in Test 5. Tests

Truss B

(cm) a b c d e f a

to h

Specified 3.81 cm TU 1 2.54 2.54 2.38 4.06 3.49 3.18 3.81 5.08

2 4.76 3.18 3.49 3.81 3.18 3.81 3.81 3.81

3 3.81 5.72 1.91 3.81 3.96 2.54 3.81 5.72 4 5.72 4.06 3.96 4.44 3.96 3.18 4.44 4.92 TM 1 2.54 3.81 4.44 4.92 4.92 5.24 NA NA

2 3.81 3.81 3.81 4.44 5.72 4.44 NA NA

3 4.44 4.13 4.44 3.96 5.08 5.72 NA NA

4 3.81 4.44 4.06 4.44 5.08 5.24 NA NA

TL 1 4.44 4.44 4.29 3.81 3.96 2.70 5.72 5.40

2 3.81 3.49 3.18 4.44 4.13 3.81 4.44 4.13

3 4.92 3.81 3.18 3.81 3.96 3.18 3.81 3.81 4 3.18 3.81 3.18 4.13 4.13 3.18 4.44 4.44

COV = 0.20; IX = 4.05 cm; o = 0.82 cm

Key: NA, not applicable. Note: Bold measurements were one standard deviation or more below specification; highlighted measurements were

below specification.

Table 9-6. SFRM thickness for truss A in Test 6. Test 6

Truss A

(cm) a b c d e f g h

Orange Specified 1.91 cm TU 1 3.18 2.70 1.91 3.18 1.91 1.27 1.91 2.86

2 2.86 1.43 2.38 2.70 3.18 1.91 1.59 2.54

3 3.18 3.18 1.91 3.18 3.81 2.38 3.18 3.81 4 2.54 1.91 1.59 1.91 2.38 1.43 1.91 2.86 TM 1 3.81 4.76 3.02 3.18 3.18 3.18 NA NA

2 2.86 3.18 2.38 3.18 3.33 2.86 NA NA

3 3.18 2.86 1.91 2.38 2.86 3.18 NA NA

4 2.86 3.18 2.70 2.38 3.81 3.02 NA NA

TL 1 2.70 1.91 2.22 2.22 2.06 1.27 2.54 2.54

2 3.33 3.18 1.91 3.81 2.54 1.11 1.43 1.91

3 3.02 2.22 2.22 2.70 2.38 1.59 2.22 2.86 4 2.86 2.86 2.06 3.18 2.70 1.75 2.54 2.70 COV = 0.26; |.i = 2.60 cm; o = 0.69 cm

Key: NA, not applicable. Note: Bold measurements were one standard deviation or more below specification; highlighted measurements were

below specification.

126 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Table 9-7. SFRM thickness for truss B in Test 6. Test 6 Truss B

a b c d e f g h

\^ p 1 low Specified 1.91 cm

1.43 2.54 1.91 3.18 2.54 1.27 1.91 2.86

2 1.91 2.54 1.59 2.54 3.18 1.27 1.91 1.91

-IJ 3.18 2.54 1.91 2.54 2.86 1.27 2.54 2.70 A 2.38 1.91 1.59 2.70 1.91 1.27 1.91 2.38 TM 1 3.81 3.18 3.18 2.06 3.81 3.18 NA NA

3.18 3.81 3.18 2.86 3.96 3.81 NA NA

3.18 2.54 2.38 2.70 3.18 3.18 NA NA

A 3.18 3.18 2.38 2.70 3.18 2.38 NA NA

TL 1 2.54 1.91 1.59 2.70 3.02 1.91 2.86 3.18

2 2.54 1.43 1.43 2.86 3.81 1.91 2.54 2.54

3.18 2.38 1.27 2.70 2.86 4.06 2.54 2.38 4 3.18 2.22 1.91 2.22 2.70 2.22 3.18 3.18 COV = 0.27; |j = 2.56 cm: a = 0.69 cm

Key: N.A.. not applicable Note: Bold measurements were one standard deviation or more below specification; highlighted measurements were below specification.

The measurements which fall below the minimum allowable thickness are highlighted and are listed in Tables 9-8 and 9-9 below (refer to Figs. 9-4b, 9-5b, and 9-6f for measurement locations):

Table 9-8. SFRM thickness measurements below allowable in Test 5.

Component

Number below Allowable/

Total Number

Percent below

Allowable Location

Bar 1/14 7% B4

Column 1/32 3 % CL3

Truss A 3/88 3 % TU3 'a' TMl 'f TL2 'c' Truss B 1 7/96 17% TUl, TM1,TL4 'a' TUl, TU2 'b' TU1,TU3, TL2,TL3, TL4 'c' TU2 at 'e' TUl, TU3, TU4, TLl, TL3, TL4 T

NISTNCSTAR 1-5B, WTC Investigation 127
Chapter 9

Table 9-9. SFRM thickness measurements below allowable in Test 6.

Component

llUIIIUcI UCIUW /VIIU W ttUIC/ Total Numberrcicciii uciuw

Allowable Location

Bar None 0 B4

Column None 0

Truss A 3/88 3 % TUl, TLl, TL2, T

Truss B 5/88 6% TUl, TU2, TU3, TU4, T, TL3, 'c'

Most of the thickness measurements that were below the minimum allowable criteria occurred at the

flange tips (locations 'c' and T for the trusses), which is recognized as a difficult area to apply SFRM

coatings (see Note 1 of ASTM E 605, above). If the deficient flange tip measurements are removed from

consideration, there remains one deficient thickness location in truss A and six in truss B for Test 5 and none for Test 6. These steel components were carefully prepared in controlled conditions. They were

tested without any attempt to address the noted thickness deficiencies.

The insulated steel components were handled carefully when they were placed in the test chamber. The

trusses had short angle stubs attached to the center web member for lifting and positioning with a forklift. No thermocouples were attached to this web member. Once the trusses were placed within the chamber,

the handling stubs and tmss support ends were wrapped with Kaowool batting to prevent heat transfer

into the truss through the bare steel areas, as the Kaowool batting thickness and thermal resistance

properties were greater than that of the SFRM coating. The columns were lifted by grabbing removable

bars inserted through holes that were pre-drilled through the columns. The column was seated on an

elevated support designed to locate the column so that the top thermocouple was placed 15 cm below the

ceiling surface. The top and bottom of the columns were seated on Maranite supports to prevent heat

transfer through the interior of the column. After an insulated column was positioned in the test chamber,

the lifting bar was removed and the holes were filled with Kaowool batting. The bars were lifted by hand

into position on the bottom chord of the two trusses (see Fig. 9-7a, 9-7b, 9-7c, and 9-7d). The bare ends were wrapped with Kaowool.

128 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

North

West

187.5 cm 111.5 cm i_

-263.0 cm—

42.0 cm 47.5 cm

236.0 cm i i

35.5 cm

12.5 cm

51.5 cm'

.150.0. cm

Loc 1

.Loc 1 l(Truss A)

Loc 2 (Truss B)

Loc 2

South (Bar A) (Bar B) Bar height from ceiling: 86 cm

Column TC locations from ceiling: 13 cm, 169 cm, 305 cm

Truss top chord location from ceiling: 15 cm

East

Figure 9-7a. Steel component locations for Tests 1, 2, and 3.

West

t

187.5 cm 111.5cm

i_

North

236.0 cm

-263.0 cm -><

35.5 cm

123.5 cm

51.5 cm'

Loc 1

South

Bar height from ceiling: 86 cm

Column TC locations from ceiling: 13 cm, 169 cm, 305 cm

Truss top chord location from ceiling: 15 cm

1 Loc 1

Loc 2

East

Figure 9-7b. Steel component locations for Tests 5 and 6.

NISTNCSTAR 1-5B, WTC Investigation 129
Chapter 9

130 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Source: NIST. Figure 9-7d. Insulated steel components in test compartment for Tests 5 and 6.

9.3 DESCRIPTION OF THE TEMPERATURE MEASUREMENTS

The temperature measurements at the surface of the SFRM and the steel components are key for determining the amount of heat that is absorbed and conducted by these materials.

The controlling physical parameters of the solid and the density, p, the specific heat, c, and the thermal

conductivity, k, can be combined mto thermal diffusivity a(= k/pc) (Holman 1976). The larger the value

of a, the faster heat will diffuse through the material. A large thermal diffusivity could occur with a high

value of thermal conductivity (k) or a low value of thermal heat capacity (pc).

Steel and SFRM thermal properties vary with temperature, where c increases with temperature for both

materials but, as the temperature increases, k decreases for steel, and k increases for SFRM . Steel density decreases slightly as thermal expansion increases the volume with increasing temperature, but this effect is often negligible for computing changes in thermal diffiisivity.

While steel can be treated as a uniform solid, fibrous SFRM is a non-uniform composite material made up

of fibers, cement, and other admixtures. When the packaged dry mix is combined with water and sprayed

onto the steel surface, it dries and cures in place with air pockets distributed throughout the material.

Furthermore, the SFRM composite material releases moisture as it is heated. While heat is transferred

through the composite by a combination of conduction and radiation, reasonable results have been

obtained by modeling the heat flow through SFRM solely with conduction heat transfer. Modeling

analysis of heat transfer through the steel and SFRM is described in Chapter 1 1

NISTNCSTAR 1-5B, WTC Investigation 131
Chapter 9

9.4 MEASUREMENT UNCERTAINTY

An uncertainty analysis associated with the thennocouple measurement of the steel surface temperature is presented. Two contributions to the uncertainty are considered. The first contribution is the inherent

uncertainty associated with a type K thermocouple, taken as approximately 5 °C (Omega 2000). The

second contribution is the systematic and inherent difference between the measured temperature and the

actual surface temperature due to the experimental configuration and the presence of a thermocouple

attached to the surface of interest. This contribution may be further broken into two separate components, an error or offset in the measurement and an uncertainty of the measurement.

The error in the steel surface temperature measurement arises from the fact that the presence of the

thermocouple bead and lead wires inherently influence the veracity of the measurement. While the

thennocouple bead temperature is strongly influenced by the surface temperature of the steel, there is a difference due to the various heat transfer processes that occur. The condition of bare (uninsulated) steel leads to an upper bound in the temperature measurement as heat transfer to the bead and lead wires of the

thermocouple from the hot gas layer via convection and radiation, as well as conductive transfer from the

bead to the steel, from the lead wires to the bead, and along the lead wires all contribute to the actual

temperature of the thermocouple bead.

Uncertainty in the measurement is from the distribution of temperature across the thermocouple bead. In thermoelectric thermometry, the electromotive force (e.m.f.) is dependent upon the material and the flow

of heat from a region of one temperature to a region at a different temperature through the Seebeck effect (Bentley 1998). Joining two different materials whose Seebeck coefficient differs creates a thermocouple. The voltage measured across the open leads of the two materials is due to the temperature

at the junction of the materials. For a beaded thennocouple, therefore, the measured temperature

depends, in reality, on the geometry of the bead, the associated contact surfaces between the disparate

metals of the thermocouple, and the temperature distribution across these contact surfaces. As it is not

possible to ascertain the precise material distribution across the bead, it is possible that the measured

temperature may be from any location within the bead. Typically, these distributions are small and can be

neglected in comparison to other more dominant sources of uncertainty.

9.4.1 Finite-Element Analysis

Analysis of the systematic measurement enor was considered using a finite-element code (ANSYS 2003)

in a manner similar to that presented in Chapter 8 for thennocouple measurements of the temperature of

the Marinite ceiling. Although an analytic steady-state solution to this problem may be simple, an

analytic transient analysis is complex and is beyond the scope of this report. Furthermore, the transient

problem is relatively more important than the steady state problem because it has a larger systematic

measurement error than the steady-state problem.

132 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

The simplified geometry considered here is presented in Fig. 9-8. The analysis is analogous to that considered for a thermocouple attached to an insulating wall material (see Fig. 8-2). An idealized

hemispherical thermocouple bead is attached to the surface of a flat steel plate taken to be 2.5 cm thick. Thermocouple lead wires were attached to the bead and extend away from the steel surface. The lead

wire diameter was 0.25 mm (30 gauge), and the bead diameter was initially taken to be twice the diameter

of the wire leads, consistent with the thermocouples used in the experiment. Although composed of two

different metals, the thermocouple was assumed to consist of only one material so as to ease the

computational burden through the application of symmetry conditions. Multiple calculations, using each

of the component materials, were performed to assess the validity of this assumption.

Figure 9-8. Simplified configuration used for the numerical analysis.

Representative thermophysical properties of the materials present were based on literature values. The

temperature dependence of properties for Chromel and Alumel, the two materials comprising a Type K

thermocouple, was neglected in the analysis; however, the temperature dependence of steel was

considered (refer to NIST NCSTAR 1-3E). The properties of interest are listed in Table 9-10. For steel, they are presented at room temperature. The emissivities of the steel and thermocouple surfaces were

taken to equal one. assuming that the surfaces were rapidly covered with soot during the experiment.

Table 9-10. Thermophysical properties of steel, Chromel, and Alumel.

Densit> Heat Capacity Conductivity

Material (kg/m') (J/kg K) (W/m K)

Steel' 7,850 435 47.6

Alumel^ 8,600 523 29.7 Chromel' 8,730 448 19.2

a. Refer to NIST NCSTAR 1-3E, Physical Properties oj Striiciurul Steels.

b. Omega 2000.

The temperature of the thermocouple bead was taken as the average bead temperature. The distribution across the thermocouple bead and in the cross-section of the lead wires was neglected. This treatment presumes an idealized temperature measurement. The difference in temperature between the

thermocouple bead attached to the steel plate and the steel surface some distance away from the

NIST NCSTAR 1-5B, WTC Investigation 133
Chapter 9

thermocouple is denoted as the systematic error in the measurement. For these simulations, the lead wires were taken sufficiently long so that the end condition could be assumed to be at ambient conditions

outside of the test compartment. The first 1 cm of lead wire off the thermocouple bead was considered to be uninsulated, whereas the remaining length was shielded with a perfect insulator.

9.4.2 Simulation Results

At time zero in the calculation, the steel plate and the thennocouple were immersed in a bath of hot gas

simulating the upper layer of the compartment in the experiment. A uniform radiative flux similar in magnitude to that observed in the experiment (20 kW/m" to 80 kW/m") was applied to exposed surfaces,

including the steel and the thennocouple, and the transient temperature field was calculated. In the

experimental configuration, the thermocouple lead wires from the bead were held to the steel surface. The close proximity to the steel surface led to shadowing of portions of the lead wires from the full radiation field emitted by the hot upper gas layer. To approximate this, the radiative flux was only taken

to be incident on half of the surface of the wire leads. Re-radiation from the surfaces was considered, but

convective heat transfer was neglected. It is expected that similar results would be found if convection was considered in the analysis, although a detennination of the localized convective heat transfer

coefficients adds complexity to the model problem. Furthermore, at the temperatures reahzed in most fire situations, heat transfer from convection is typically much smaller than from radiation.

The simulation results showed that the thermocouple lead wires increased in temperature at a faster rate than did the surface of the steel. This can be attributed to the small thermal mass of the lead wire as compared to that of the steel. The rapid increase in temperature established a conductive flux along the

wire to the thermocouple bead, which in turn increased the temperature of the bead. The high thermal

conductivity of the steel helped dissipate the added heat flowing to the bead; nonetheless, a temperature

gradient was established between the surface temperature of the steel (away from the thennocouple) and

the thermocouple bead. The temperature difference arose from the conductive flux through the lead

wires. As time progressed and the temperature of the steel plate and thermocouple approached steady

state, the conductive flux diminished, and the temperature difference decreased to near zero at steady

state. A phase plot representing the systematic error associated with the measurement is presented in Fig. 9-9. The thennocouple temperature, plotted on the horizontal axis, is presented as a function of the

difference in temperature between the steel surface and the thermocouple temperature, plotted on the

vertical axis. For the various heat fluxes considered, as described above, there is an initial rise in the enor

observed early in the simulation. Maximum temperature differences of approximately 0.5 °C, 1 °C, 2 °C,

and 4 °C were found for fluxes of 10 kW/m", 20 kW/m% 40 kW/m" and 80 kW/m', respectively. The

differences remained nearly steady for some time and then rapidly decreased as steady state was reached.

134 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

4

900

Thermocouple Temperature ( °C )

Figure 9-9. Calculated measurement error for thermocouples on uninsulated steel.

As stated above, both materials of a Type K thermocouple were considered in independent analyses. The

resulting calculated temperature differences between the thermocouple bead and the steel surface were

found to be similar (within 10 percent) for both materials, justifying the earlier assumption that approximated the thermocouple using a single material. As expected, differences were slightly larger for simulations using Alumel as compared to Chromel due to Alumel's higher thermal conductivity. The

larger thermal conductivity increased the thermal flux from the wire to the bead, and resulted in a larger

temperature difference.

Variation of the bead diameter was also considered. For simulations utilizing a bead diameter equal to four times the wire diameter, an increase of approximately 1 0 percent in the systematic error was

calculated.

For very early times in the experiment, the thermocouple and steel surfaces may not be completely

covered with soot. The emissivities would then be closer to the true emissivities of the respective

materials, 0.8 for steel and 0.2 for Alumel (Smyth et al. 1985). The fluxes to the wire leads would be

diminished, reducing the rate of temperature rise, and decreasing the associated conductive flux between

the wire and the bead. The temperature differences between the thermocouple and the steel would be less than that described above as would the variation in temperature across the bead, reducing the uncertainty

in the measurements.

In summary, the systematic error in the bare steel measurements is minimized as the temperature reaches

steady state in the experiment and as the gradients between the thermocouple bead and the steel approach

zero. For elevated steel temperatures (T>300 °C), the analysis shows that the systematic error in the

temperature measurement at steady state due to the experimental configuration is less than 4 °C even for

high flux conditions (80 kW/m"). If the systematic error is treated as a measurement uncertainty, then the

total expanded measurement uncertainty is bounded by +1 °C and -9 °C. A correction to the temperature

NISTNCSTAR 1-5B, WTC Investigation 135
Chapter 9

measurements could be applied, provided that estimates of the local time varying heat flux are available.

The value of the correction, however, is relatively small for conditions of interest here.

9.5 MEASUREMENT UNCERTAINTY OF THE SFRM SURFACE

TEMPERATURE

The dominant contributor to the uncertainty in the measurement of the SFRM is related to the exact

location of the thermocouple bead. The coating is fibrous by nature, and the applied coating was uneven. Thennocouples were positioned onto the SFRM by pushing the lead wires and bead parallel to the steel surface and into the SFRM. A post-experiment assessment of the positions of the theraiocouple bead

relative to the surface of the coating showed that the thermocouple bead extended as far as 5 mm below

the surface of the SFRM. The uncertainty in thermocouple position can lead to a systematic error in the

measurement of the surface temperature. Calculations using the ANSYS finite-element software show

that the magnitude of the error varies as a function of time. At early times in the calculation, the

temperature gradient near the surface of the SFRM is very high, and the uncertainty in the position of the

bead can lead to temperature measurements that are 30 percent less than the actual surface temperature.

This error diminishes with time as the SFRM heats and the temperature gradient decreases. After long

times, for example, the calculations showed that the systematic measurement error decreased to approximately 70 °C for the insulated bar exposed to 1 ,000 °C in the upper layer. Values of 7 percent less than the actual surface temperature are taken as representative of the systematic expanded uncertainty.

The inherent thermocouple uncertainty of 5 °C provides an upper bound for the uncertainty, assuming

that the thermocouples were positioned below the surface of the SFRM. Post-experiment inspection of

representative thermocouples suggests that this was the case.

9.6 DESCRIPTION OF EXPERIMENTAL RESULTS

Tests 1, 2, and 3 subjected the same set of two bars, one column, and two trusses to direct exposure of

heat flux and temperatures from a fuel pan fire. In all three tests, the thermal response of these

components to flame proximity or immersion was similar. The steel began to heat immediately upon

initiation of the fire and to cool immediately after the fire was extinguished. The recorded response

showed no lag time between the steel and heated gas for transferring radiant heat from the higher to lower

temperature mediums.

Tests 5 and 6 exposed separate sets of one bar, one column, and two trusses that were insulated with

SFRM to direct exposure of heat flux and temperatures from a fuel pan fire. In both tests, the thermal

response of these components to flame proximity or immersion was similar. The SFRM surface began to heat immediately upon initiation of the fire and to cool immediately after the fire was extinguished. The

protected steel responded more slowly, taking approximately 50 min with 1.91 cm (0.75 in.) of SFRM to reach 600 °C temperatures as compared to the 7 min when the steel was bare and subjected to similar

3 MW fires. Steel temperatures continued to rise approximately 25 °C to 50 °C for a 5 min to 10 min

period after the fire was extinguished, before cooling began to take place by radiating heat through the

insulation back into the relatively cooler gases.

Tables 9-1 1 and 9-12 summarize the minimum and maximum peak temperatures recorded during each

test. Table 3-5 in Chapter 3 lists the mean heat release rate ( Q ) and duration for each test. The

maximum steel temperatures occurred near the fire pan in areas where there was flame impingement on

136 NISTNCSTAR1-5B,WTC Investigation
Temperatures of Steel Components

the bars (locations B3 and B4 in Fig. 9^a) and trusses (location 3 or 4 in Fig. 9-6a) and on column

surfaces facing the fire pan. The tests were stopped when one or more thermocouples on the steel surface

reached approximately 600 °C.

Table 9-1 1 lists temperatures recorded during Tests 1, 2, and 3 of bare steel components. Truss B at location TL3 recorded the highest temperature in each test (615 °C to 705 °C), as this location was

located directly over the fire pan and was often immersed in flames. The sequence of maximum

temperatures (from highest to lowest peak temperature) is the same for each test—truss B, truss A, bar A,

bar B, and the column—indicating that radiation from the fire is the primary heat transfer mechanism.

The trusses and bars were immersed in a hot gas layer near the ceiling of the test chamber (see Chapter 4),

and their peak minimum temperatures were within 60 °C of each other. Note that the minimum

temperatures were higher (330 °C to 390 °C) for Tests 1 to 3, which had a nominal ^ of 2 MW for approximately 15 min, than the minimum temperatures recorded for Test 2 (235 °C to 285 °C), which had

a higher nominal ^ of 3 MW for 6 min. The column components measured similar temperature ranges

(maximum-minimum) of 150 °C to 170 °C.

Table 9-1 1 . Summary of peak 1temperatures recorded during 1rests 1, 2, and 3.

Component

Time

(s)

Maximum

Temp

( ^)

TC

i^aoei

Minimum

Temp

( ^)

TC

i^aoei

Temp

Range

( ^)

Ttict 1 i ] A fnin\ 1 cM 1 ^ IH.J lilill^ 870

Bar A 525 B4 365 Bl 160

Bar B 440 B4 365 Bl 75 Coluinn 390 CL2 240 CLl 150

Truss A 560 TL3 360 TMl 200

Truss B 615 TL3 370 TMl 240

Test 2 (6.3 min) 380

Bar A 525 B3 235 Bl 190

BarB 370 B4 260 Bl 110 Column 320 CU3 150 CLl,

CL4

170

Truss A 600 TL3 270 TMl 230

Truss B 705 TL3 285 TMl 420

Testa (16.2 min) 970

Bar A 535 B4 330 Bl 205

BarB 435 B4 350 Bl 85

Column 380 CU3 225 CLl 155

Truss A 545 TU4 375 TMl 170

Truss B 635 TL3 390 TMl 245

NISTNCSTAR 1-5B, WTC Investigation 137
Chapter 9

Table 9-12. Summary of peak temperatures recorded during Tests 5 and 6.

Component

Time

(s)

Maximum

Temp

( C)

Temp

Range at Location

(°C)

TC

Label

Minimum

Temp

( C)

Temp

Range at Location

(°C)

TC

Label

Temp

Range

( C)

1 est 5 (j U min) Bar

1 .y ! cm brRM

845

380

B4 680

255

B2 165

Bar

bteel

465 B3 425 B2 40

Column

J.S 1 cm brKM

760

560

CU3 425

305

CLl 335

Column

bteel

200 CU2 120 CL4 80

Truss A

1.91 cm SrRM

835

285

TU4 625

225

TLl 210

Truss A

Steel

550 TM4

TI 14 400 TUl 150

Truss B

3.81 cm SrRM

870

355

TU4 660

485

TUl 210

Truss B

steel

415 TM4

TT 14 175 TL2 240

1 est 6 (50.2 mm) O A 1 A

Bar

1 .91 cm SFRM

865

355

B4 730

285

Bl 135

Bar

Steel

510 B4 445 Bl 65

Column

1 .91 cm SFRM

745

360

CM2 495

220

CL4 150

Column

Steel

375 CUl 275 CLl 100

Truss A

1.91 cm SrRM

950

285

TM4 670

300

TUl 280

Truss A

oieei

665 TU4 370 TLl 295

Truss B

1.91 cm SFRM

930

250

TM4 655

240

TMl 275

Truss B

Steel

680 TU4 415 TLl 265

Note: Steel temperatures for Tests 5 and 6 continued to rise approximately 25 °C to 50 °C after the ftiel flow was halted.

138 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Figures 9-10 through 9-14 show the temperature histories recorded during Test 3 for all the

thermocouples. The steel temperatures all rose rapidly for the first 8 min to 10 min at a nearly linear rate. As the steel reached 300 °C to 400 °C, the temperature rise continued at a slightly decreased rate. The

recorded thermocouple temperatures for bar B and truss A subject to flame immersion are 50 °C to 75 °C

higher than similar locations on the other bar and truss and tend to have a high frequency signal imposed

upon the average temperature. The columns were located next to the fire pan, but the airflow through the

test chamber resulted in the fire leaning toward the exhaust outlet, away from the columns, so that the

columns were heated primarily by radiation. The sides of the column facing the fire were heated more

quickly, with thermocouples at the ceiling and floor level facing the fire pan recording maximum

temperamres. There was a marked drop in steel temperatures when the fuel flow was shut off. As soon as the compartment gas temperatures fell, the steel temperature began to decrease as heat was transferred

through radiation and convection to the surroundings.

600

500 —TB1SA — TB2SA

0) fQ.

E

400

200

300

100

0

— TB1SA —TB2SA

TB3SA — TB4SA

TB3SA

TB4SA

Time (sec)

Figure 9-10. Temperature history for bar A in Test 3.

NISTNCSTAR 1-5B, WTC Investigation 139
Chapter 9

600

CNJ CO

CD LO

c\J ro

a> CO

CO 00 O)

•r- CM CO

t£) LD ^ o CM ro

lO tD

CO CSl <- LO ID

00 CT) O CM CO CMCNJCNJCvjCMCNCMCNCMrOrOcOrOCO

Time (sec)

Figure 9-11. Temperature history for column in Test 3.

600 -I

140

Figure 9-12. Temperature history for truss A in Test 3.

NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Figure 9-13. Temperature history for bar in Test 5.

900

800

700

600

i- 400

Time (sec)

Figure 9-14. Temperature history for column in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 141
Chapter 9

Table 9-12 lists temperatures of SFRM protected components recorded during Tests 5 and 6. Tests 5 and

6 are discussed separately to address the differences in the SFRM coatings. In Test 5, truss A (1.91 cm

SFRM) had peak maximum steel temperatures at locations TM4 and TU4, which were near, or immersed,

in flames. The peak maximum steel temperature on truss B (3.81 cm SFRM) was 135 °C less at the same

lateral (y axis) locations. The temperatures on the insulation surface rose sharply at the beginning of the

test and were close to their peak temperatures after 5 min to 10 min, whereas the steel surface

temperatures rose more gradually. The maximum SFRM surface temperature was 835 °C to 870 °C at TU4 for both trusses. The peak minimum SFRM surface temperatures were similar (625 °C to 660 °C),

but the steel minimum temperatures differed by 225 °C (400 °C vs. 175 °C). The difference in steel temperatures appears to be primarily due to the difference in SFRM thickness, as the SFRM temperatures were within 5 percent. The bar (1.91 cm SFRM) temperatures are similar to truss A, with the exception of

the peak maximum steel temperature (85 °C less). The column (3.81 cm SFRM) had markedly lower

SFRM and steel temperatures than the trusses and bar. The peak maximum steel temperatures recorded at column location CU3, in the hot gas layer near the ceiling and facing the fire, were 100 °C to 300 °C less than the other components.

In Test 6, truss A (1.91 cm SFRM) and tmss B (1.91 cm SFRM) recorded similar peak maximum and

minimum temperatures for the steel and SFRM at locations TM4 and TU4. The bar ( 1 .91 cm SFRM)

temperatures are similar to the bar in Test 5. The column ( 1 .91 cm SFRM) had similar SFRM

temperatures but higher steel temperatures than the column in Test 5. Figures 9-15 through 9-18

illustrate how the insulated steel components responded to the fires. The SFRM surface thermocouples

responded to the gas temperatures quickly, reaching their peak temperatures in 5 min to 10 min. The steel temperature rise was essentially uniform for all thermocouples until the steel temperatures reach 100 °C,

when the rate of temperature rise increases for all thermocouples, but at different rates. The SFRM

surface temperatures showed an immediate response to the drop in compartment gas temperatures.

The steel temperatures continued to rise approximately 25 °C to 50 °C after the fuel flow was shut. Figures 9-19 and 9-20 show the steel components before and during Test 6, with flame impingement,

through the exhaust opening.

500 ( —

600 1200 24U0 3IIUU 36(IU

Time (sec)

-TUISA

-TU2SA TU3SA

TU4SA -TMISA

-TM2SA

-TM4SA

-TLISA TL2SA

TL3SA

TL4SA

TUIIA

TU2IA

-TU3IA TU4IA TM21A

TM3IA TM4IA TLUA

-TL2IA

-TL3IA TL4IA Figure 9-15. Temperature history for truss A in Test 5.

142 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

0 600 1200 1800 2400

Time (sec)

3000 3600 4200

1 4800

TCU2S TCU3S TCU4S TCM1S TCM2S TCM3S TCM4S TCL1S TCL2S

TCL3S TCL4S TCU1I TCU2I TCU3I TCU4I TCM2I TCM3I TCM4I

TCL1I TCL2I TCL3I TCL4I

Figure 9-16. Temperature history for column in Test 5.

600 1200 1800 2400 3000

Time (sec)

3600 4200 4800

-TUISA

TISSA

-TI11A

-TU2SA TU3SA

TL4SA TUIIA

.TL21A TL3IA

-TU4SA

TIJ2IA

-TL4IA

-TM1SA

TUSIA

-TM2SA

TU4IA

-T1VI4SA -TM2IA-TL1SA

-TIVI3IA

TL2SA

j -TM4IA !

Figure 9-17. Temperature history for truss A in Test 5.

NISTNCSTAR 1-5B, W/TC Investigation 143
Chapter 9

0 600 1200 1 800 2400 3000 3600 4200 4800

Time (sec)

TU1SB TU2SB TU3SB TLMSB TM1SB TM2SB TM3SB TM4SB

TL1SB TL2SB TL3SB TL4SB TUIIB TU2IB TU3IB "TLWIB TMIIB —TM2IB TM3B TM4IB ^ . —TLaB .TL3B "n.4IB

Figure 9-18. Temperature history for truss B in Test 5.

Source: NIST. Figure 9-19. View of truss and bar components before Test 6.

144 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

Source: NIST.Figure 9-20. Flame impingement on trusses and bar during Test 6.

Table 9-13 lists the thermal lag in Tests 5 and 6 for the various insulated steel components. The time lag was defined as the difference between the time when the maximum temperature was reached in the

SFRM. which occurred when the fuel was shut and the fire stopped, and the time when the maximum

temperamre in the steel components was reached. The data shows that the SFRM thickness, SFRM

temperatures, and steel temperatures all affected the thermal wave or the rate at which heat traveled

through the SFRM. For components with a nominal 1 .91 cm SFRM thickness, the time lag ranged from

100 s to 800 s. For components with a nominal 3.81 cm SFRM thickness, the time lag ranged from

1,000 s to over 1 ,900 s. Further inspection of the table shows that larger temperature differences between

the steel and SFRM also had a shorter time lag. This reflects the fact that conduction is directly

dependent upon the temperature gradient between two points.

Table 9-13. Thermal lag in insulated steel components during Test s 5 and 6.

Test Component

SFRM Thickness

(cm)

Time Lag

(s) SFRM Temp(°C)

Steel Temp

(°C)

5

Bar 1.91 200-300 650-850 400-550

Column 3.81 1,000-1,900 650-900 200-400

Truss A 3.81 1,900+ 425-775 100-200

Truss B 1.91 400-600 650-850 400-500

6

Bar 1.91 300-800 650-925 375-675

Column 1.91 100-400 650-925 400-700

Truss A 1.91 600-1,000 500-750 275-375

Truss B 1.91 400-500 700-850 450-550

NISTNCSTAR 1-5B, WTC Investigation 145
Chapter 9

9.7 SUMMARY OF RESULTS

Insulated and uninsulated steel components were exposed to compartment fires with varying heat release

rates, fuel type, and duration. The surface temperature of bare steel trusses that were subject to impingement or proximity of flames from fires with a nominal 3 MW heat release rate reached 600 °C

within 15 min. In comparison, the insulated steel trusses with SFRM protection (1.91 cm nominal

thickness) reached 100 °C to 200 °C and the truss with 3.81 cm SFRM reached 50 °C to 100 °C after 15 min of exposure to a 3 MW heat release rate fire. After 50 min of exposure to a 3 MW nominal heat

release rate, the trusses with SFRM protection ( 1 .91 cm nominal thickness) reached 550 °C to 680 °C and

the truss with 3.81 cm SFRM protection reached 415 °C.

There is also a strong influence of SFRM thickness on steel temperatures. A significant variation in the

thickness of the SFRM coating was measured. The 1.91 cm specified thickness average was measured to be 2.2 cm to 2.7 cm between components with a COV ranging from 0. 1 7 to 0.27. The 3.81 cm specified

thickness average was measured to be 4.05 cm to 4.10 cm between components with a COV ranging from

0.07 to 0.20. The lower COV values were measured for the columns; the higher COV values occurred for

the bars and angles, which are more difficult shapes for unifonn application of the SFRM product. These

SFRM measurements may be considered as lower bound estimates of the variance that would be expected

of field applications.

The uninsulated steel temperatures reached their maximum values approximately when the fuel flow was

cut and the fire stopped. The same was true for the insulated steel components, except that the maximum

temperatures were reached after the fuel was stopped due to lag of the thermal wave through the SFRM.

The thennal lag varied between 100 s and 800 s for a nominal 1.91 cm SFRM thickness and 1,000 s to over 1 ,900 s for a nominal 3.81 cm SFRM thickness, depending upon the SFRM and steel temperatures

and gradients.

The experimental results showed:

• SFRM coatings significantly delay the rise of steel temperatures when exposed to hot gases

and flame impingement.

• The surface of the SFRM reached the gas temperatures within a few minutes after the start of

the test. The SFRM temperatures closely matched the gas temperature variation along the

chamber length.

• For uninsulated (bare) steel components, temperature variations existed along the length of

the component throughout the test.

• For insulated components, the highest temperatures reached at steel surfaces were

significantly lower than the temperature at the outside face of the insulation material, even at locations of flame impingement.

• Although steel has a high conductivity relative to the SFRM product, temperatures did not become uniform along the steel component lengths in either the bare or insulated state, indicating that the local gas temperature and the SFRM played a significant role in the

temperature rise.

146 NISTNCSTAR 1-5B, WTC Investigation
Temperatures of Steel Components

• The surface temperature of uninsulated steel trusses that were subject to impingement or proximity of flames from fires with a nominal 3 MW heat release rate reached 600 °C within

15 min. In comparison, the insulated steel trusses with SFRM protection (1.91 cm nominal

thickness) reached 100 °C to 200 °C and the truss with 3.81 cm SFRM reached 50 °C to 100 °C after 15 min of exposure to a 3 MW heat release rate fire. The steel surface of an

insulated steel truss with 1.91 cm (nominal) and 3.81 cm (nominal) SFRM obtained a temperature of 680 °C and 415 °C, respectively, after 50 min of exposure to a 3 MW fire.

• For the insulated steel components, the steel surface temperatures obtained a maximum

sometime after the fire was stopped due to lag of the thermal wave through the SFRM. A

thermal lag of between 100 s and 800 s for the nominal 1.91 cm SFRM thickness and 1,000 s to over 1,900 s for the nominal 3.81 cm SFRM thickness was observed, depending on the

values of the temperature and the temperature gradients within the SFRM and the steel.

Chapter 1 1 compares the experimental measurements reported here with the FDS/FSI/ANSYS modeling

results for the thermal behavior of the bare and insulated steel components, hi particular, the measured

steel and SFRM surface temperatures in Tests 1 and 5 are quantitatively compared with the modeling

results.

9.8 REFERENCES

ANSYS, Inc. 2003. ANSYS Release 8.0 Documentation.

American Society for Testing and Materials. 2000. ASTM E 605-93, Standard Test Methodsfor

Thickness and Density ofSprayed Fire-Resistive Material (SFRM) Applied to Structural Components,

West Conshohocken, PA.

Bentley, R. E. 1998. Handbook of Temperature Measurement, Volume 3: Theoiy and Practice of

Thermoelectric Thermometiy. Springer-Verlag Singapore Pte. Ltd.

Hamins, A., A. Maranghides, and G. Mulholland. 2003. The Global Combustion Behavior of 1 MWto 3 MWHydrocarbon Spray' Fires Burning in an Open Environment, NISTIR 7013, National Institute of Standards and Technology, Gaithersburg, MD, June.

Holman, J. P. 1976. Heat Transfer, Fourth Edition, McGraw-Hill Book Company.

IBC (International Building Code). 2000. International Code Council, Falls Church, VA.

Kay, T.R.. B.R. Kirby, and R.R. Preston. 1996. Fire Safety J. 26, 327-350.

NFPA (National Fire Protection Association). 2003. NFPA 5000: Building Construction and Safety

Code, 2003 edition, Quincy, MA.

Omega Engineering, Inc. 2000. Omega Temperature Handbook, Technical Reference Section,

Section Z.

NISTNCSTAR 1-5B, WTC Investigation 147
Chapter 9 ^

Smyth, K., J.H. Miller, R.C. Dorfman, W.G. Mallard, and R.J. Santoro. 1985. Combust. Flame 62,

157-181.

148 NISTNCSTAR 1-5B. WTC Investigation
Chapter 10

Simulation of Compartment Fire Phenomena

10.1 INTRODUCTION

In cooperation with the fire protection engineering community, a computational fire model, the Fire

Dynamics Simulator (FDS), has been developed at National Institute of Standards and Technology

(NIST) to study fire behavior and to evaluate the performance of fire protection systems in buildings. The

software was released into the public domain in 2000, and since then has been used for a wide variety of

analyses by fire protection engineers. Throughout its development, the model has been compared with a variety of fire experiments to ensure that the predictions are accurate. Most often, the experiments are not

necessarily designed solely for the purpose of model validation, and as a resuh there are often several

feamres in the tests that are difficult to model. It is then difficuh to assess the overall accuracy of the

model because the uncertainty in modeling the difficult features dominates the error. The large-scale

experiments described in this report were designed specifically to test the accuracy of FDS.

The fire, itself, was relatively easy to model, and the heat release rate was measured. The compartment was heavily instrumented so that all of the energy from the fire could be accounted for and reported in terms of conductive losses to walls, convective flux through openings, etc. With the large number of

measurements, it was possible to go beyond the traditional point by point comparison and discover why

the model either over-predicted or under-predicted a given measurement. It was possible to compare the

transport of energy, starting with the combustion of fiiel and ending with effluent exiting into a large

hood. Based on these integrated quantities, discrepancies in heat flux and gas concentration predictions

could be tied to errors in the overall energy budget, allowing assessment of the accuracy of various

components within the model.

Model calculations were performed before testing began to guide the design of the compartment and also

to provide a baseline set of "blind" predictions. A complete description of the model can be found in McGrattan (2004). Briefly, FDS is a computational fluid dynamics code that solves the Navier-Stokes

equations in low Mach number, or thermally-expandable, form. The transport algorithm is based on large

eddy simulation techniques; radiation is modeled using a gray gas approximation, and a finite volume

method is used to solve the radiation transport equation. Combustion is modeled using a mixture fraction

approach, in which a single transport equation is solved for a scalar variable representing the fraction of

gas originating in the fuel stream.

1 0.2 description of the calculations

The geometry of the compartment was relatively simple and is described in detail in Chapter 2. Figure 2-5 is a schematic drawing of the compartment contents. Figures 2-1 through 2^ are photographs of the compartment. The enclosure was rectangular as were the vents and most of the

obstructions. A uniform numerical grid, whose cells were 10 cm on a side, was chosen based on a grid

resolution study (NIST NCSTAR 1-5F). Figure 10-1 is a snapshot of a simulation showing the isometric

NISTNCSTAR 1-5B, WTC Investigation 149
Chapter 10

surface of stoichiometric mixture fraction, which is a representation of the flame surface. The figure also

shows the major geometric features of the compartment as seen from the south wall looking toward the

north. The inlets are evident on the left and the outlets on the right. It is useful to compare the snapshot

with a photograph such as Figs. 2-2 and 2-3. Figure 10-2 shows the temperature profile through the

centerline plane for one of the simulations as seen from the south wall looking toward the north.

Figure 10-1. Snapshot of numerical simulation of Test 1 showing the fire as seen from a cutaway view looking towards the north wall.

temp C

rz5

25.0

Figure 10-2. Snapshot of the numerical simulation of Test 1 showing the fire

and a contour plot of gas temperature along the centerline of the compartment

as seen from a cutaway view looking towards the north wall.

The compartment walls and ceiling were made of 2.54 cm (1 in.) thick Marinite 1, a product of

BNZ Materials, Inc. The manufacturer-provided thermal properties were given as a density of 737 kg/m'

a conductivity of 0.12 W/m/K. The specific heat range was given as 1,172 J/kg-K at 93 °C to

150 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Ptienomena

1,423 XTcg-K at 425 °C. These values differed somewhat from the values measured by Taylor et al. (2003). which are listed in Table 8-1. The calculation results were not sensitive to these differences as discussed below.

The steel used in the column and truss flanges for the experiments was 0.64 cm ( 1/4 in.) thick as described in Chapter 9. The density of the steel was assumed to be 7,860 kg/m\ and its specific heat

450 J/kg-K (NIST NCSTAR 1-3E). The steel was assumed in the FDS model to be thermally-thin; thus, no thermal conductivity was used. Note that FDS performs a simple one-dimensional calculation of the

steel temperature to be used as a boundary condition in the calculation. More detailed calculations of the

thermal behavior of the steel are discussed in Chapter 1 1 of this report.

Two fuels were used in the tests. The properties of the fuels were obtained from measurements made on a

series of unconfmed bums (Hamins 2003). The first fuel was a blend of heptane isomers. The soot yield was set at a constant value equal to 1.5 percent (Hamins et al. 2003). The present version of FDS does

not adjust the soot yield based on compartment ventilation or combustion efficiency. The second fuel was a mixture (40 percent/60 percent by mass) of toluene, CyH^, and heptane as discussed in Chapters 2 and 3. Because FDS only considers the burning of a single hydrocarbon fuel, the mixture was taken to be

C7H12 with a soot yield of 1 1 .2 percent. The radiative fraction for the heptane blend is 0.44; for the

heptane/toluene mixture it is 0.39. FDS does not use the measured radiative fraction of the fuel directly. Instead. FDS computes the radiative loss from the fire based on the temperature and gas composition.

This calculation is subject to considerable uncertainty because of the fourth power dependence on

temperature. As a consequence, FDS uses a user-prescribed lower bound for the radiative fraction, which

in this case was 0.35. This means that the simulated fire radiates at least 35 percent of its energy. For the

simulations reported here, the radiative fraction was computed to be about 0.40. A sensitivity analysis,

described in Sec. 10.4.6, shows that the calculation results were insensitive to the prescribed lower bound

of the radiative fraction. The FDS Technical Reference Guide (McGrattan 2004) provides a detailed

discussion of how the output parameters are interpreted in the simulations.

The heat release rate of the simulated burner was set to that which was measured in the experiments. No

attempt was made to model the spray burner. The heat release rate was assumed to be uniformly

distributed over a 1 m by 1 m area in the fuel pan, consistent with observations of the hydrocarbon spray

fire. With the exception of Tests 2 and 3, the target heat release rates for the experiments were achieved

by adjusting the fuel flow rate to account for combustion inefficiencies that were measured in a series of

unconfined bums. During the simulation, values of temperature, heat flux, and gas species concentrations were reported as 5 s averages, consistent with the time response of the various instmments (see

Table 2-3). Linear interpolation was used to approximate values between the 10 cm grid cells.

10.3 DISCUSSION OF MODEL SIMULATIONS AND MEASUREMENTS

In this section, the experimental measurements and the numerical simulations are presented. Comparison

of the simulation results with the measurements should be considered in terms of the measurement

uncertainty (which is summarized in Table 2-3) and a sensitivity analysis of the modehng results. In Sec. 10.4, a sensitivity analysis is presented, and comparisons are made with the measurements.

NIST NCSTAR 1-5B, WTC Investigation 151
Chapter 10

10.3.1 Gas Temperatures

Upper and lower layer gas temperatures were measured in the experiments using both aspirated and barebead thermocouples, as discussed in Chapter 4. For the purpose of model validation, the aspirated

thermocouple measurements were used because measurement uncertainty was relatively small (7 °C) as the influence of external thermal radiation on the measurement was greatly reduced. Two vertical arrays of thermocouples were simulated, one on the inlet (west) side, near the centerline, and one on the outlet

(east) side. Figures 10-3 through 10-12 display the time histories of both measured and predicted

compartment gas temperatures. Note that in some of the experiments, the pumps drawing air through the

aspiration tubes were not turned on until some time during the test. In these cases, the measured

temperatures were noticeably lower than the predicted values.

U

;-(

o

B

oH

600

500 -

400 -

300 -

200 -

100 -

WTC Phase 1, Test 1, West Aspirated TCs

34 cm (Exp)

365 cm (FDS)

215 cm (FDS)

34 cm (FDS)

200 400 600 800 1000 1200

Time (s)

Figure 10-3. Comparison of gas temperatures, Test 1, intake side of compartment.

The pumps connected to the aspirated thermocouples began to operate at 180 s.

152 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

U

600

500 -

400 -

« 300

<L> a, H 200

100 -

0 200

WTC Phase 1, Test 1, East Aspirated TCs

• ••* * *•* Jy^ ^

365 cm (Exp)

215 cm (Exp)

35 cm (Exp)

s / j ^ 1 / 365 cm(FDS)

/ / s / /

*> / / 215cm(FDS)

^ / /

% / /

/ /

\ / /

s If1 1

\ / /

35 cm(FDS)

1 1 1 1 1 1 400 600 800 1000 1200

Time (s)

Figure 10-4. Comparison of gas temperatures, Test 1, exhaust side of compartment.

The pumps connected to the aspirated thermocouples began to operate at 180 s.

o.

600

500 -

400 -

300

200 -

100 -

0

WTC Phase 1, Test 2, West Aspirated TCs

0

365 cm (Exp)

215 cm (Exp)

34 cm (Exp)

365 cm (EDS)

215 cm (EDS)

34 cm (EDS)

200 400 600 800 1000 1200

Time (s)

Figure 10-5. Comparison of gas temperatures, Test 2, intake side of compartment.

NISTNCSTAR 1-5B, WTC Investigation 153
Chapter 10

Time (s)

Figure 10-6. Comparison of gas temperatures, Test 2, exhaust side of compartment.

154 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

600

500 -

400 U

3

300

200 -

100 -

400 600 800 1000 1200

Time (s)

Figure 10-8. Comparison of gas temperatures, Test 3, exhaust side of compartment.

800

600

/—

\

f 400 -

53

200 -

0

WTC Phase 1, Test 4, West Aspirated TCs

0 200 400 600 800 1000 1200

Time (s)

Figure 10-9. Comparison of gas temperatures, Test 4, intake side of compartment.

NISTNCSTAR 1-5B, WTC Investigation 155
Chapter 10

U

P

3

u

r1000

800 -

600 -

400 -

200 -

WTC Phase 1, Test 4, East Aspirated TCs 365 cm (Exp)

215 cm (Exp)

35 cm (Exp)

365 cm (FDS)

215 cm (FDS)

35 cm (FDS)

400 600 800 1000 1200

Time (s)

Figure 10-10. Comparison of gas temperatures, Test 4, exhaust side of compartment.

U

(D I

(D H

1000

800 -

600 -

400

200 -

0

WTC Phase 1, Test 5, West Aspirated TCs

365 cm (FDS)

215 cm (FDS)

34 cm (FDS)

1000 2000 3000

Time (s)

Figure 1 0-1 1 . Comparison of gas temperatures, Test 5, intake side of compartment.

156 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

1000

800 -

WTC Phase 1, Test 5, East Aspirated TCs

365 cm (Exp)

215 cm (Exp)

35 cm (Exp)

365 cm (EDS)

215 cm (EDS)

35 cm (EDS)

2000 3000

Time (s)

Figure 10-12. Comparison of gas temperatures, Test 5, exhaust side of compartment.

10.3.2 Heat Flux to Solid Objects

Heat flux gauges, some measuring total heat flux, some radiative, were positioned at various locations in both the upper and lower layers as discussed in Chapter 5. Heat fluxes to the upper and lower sections of

the column are shown in Eigs. 10-13 through 10-22. The large spikes in these figures were due to the Nt

purge used to remove soot from the total heat flux gauges. Comparisons between measured and predicted

heat fluxes at the floor are shown in Eigs. 10-23 through 10-27.

It is worth noting that EDS predicted "cold wall" heat fluxes to the total flux gauges, thereby adding some

uncertainty in the comparison due to convective heat transfer. It did so by using estimated convective

heat transfer coefficients applied to each of the gauges. None of the flow situations in the compartment was a particularly clean example of a textbook case for which a heat transfer correlation is available. The

flow velocity and gas temperature data that go into the correlations came from the model itself, not from

experimental measurements. The uncertainties were difficult to quantify, but convection was typically a minor contributor to the total heat flux that the gauges recorded.

NISTNCSTAR 1-5B, WTC Investigation 157
Chapter 10

0 —=f 1 1 1 1 1 0 200 400 600 800 1000 1200

Time (s)

Figure 10-13. Comparison of heat fluxes to column, Test 1. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

0 200 400 600 800 1000 1200

Time (s)

Figure 10-14. Comparison of heat fluxes to column, Test 1. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

158 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

X

3

X

C 3lumn Heat Flux Gauges, East Facing

High East (Exp)

Low East (Exp)

High East (EDS)

Low East (EDS)

200 400 600

Time (s)

800 1000 1200

Figure 10-15. Comparison of heat fluxes to column, Test 2, The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

50

40

30 -

X

3

Z 20

X

10

WTC Phase 1, Ttst '.I, Column Heat Flux Gauges, North Facing

High North (Exp)

Low North (Exp)

High North (FDS)

Low North (FDS)

200 400 600

Time (s)

800 1000 1200

Figure 10-16. Comparison of heat fluxes to column, Test 2. The periodic spikes

In the measurements were due to the influence of a nitrogen purge.

NISTNCSTAR 1-5B, WTC Investigation 159
Chapter 10

0

1 1 1 1 1 r 0 200 400 600 800 1000 1200

Time (s)

Figure 10-17. Comparison of heat fluxes to column, Test 3. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

High North (Exp)

Low North (Exp)

0 200 400 600 800 1000 1200

Time (s)

Figure 10-18. Comparison of heat fluxes to column, Test 3. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

160 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

0 I

1 1 1 1 1 0 200 400 600 800 1000 1200

Time (s)

Figure 10-19. Comparison of heat fluxes to column, Test 4. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

1200

Time (s)

Figure 10-20. Comparison of heat fluxes to column, Test 4. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

NISTNCSTAR 1-5B. WTC Investigation 161
Chapter 10

'a

100

80 -

Time (s)

Figure 10-21. Comparison of heat fluxes to column, Test 5. The periodic spikes

in the measurements were due to the influence of a nitrogen purge.

100

3

1000 2000 3000

Time (s)

Figure 10-22. Comparison of heat fluxes to column, Test 5. The periodic spikes

In the measurements were due to the influence of a nitrogen purge.

162 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

35

30 - ^

25 -

I—

r15 H

10 -

5

-

Phase 1, Test 2, Floor Heat Flux Gauges

0

500 cm (Exp)

550 cm (Exp)

500 cm (FDS)

550 cm (FDS)

200 400 600 800 1000

Time (s)

Figure 10-24. Comparison of heat fluxes to floor targets, Test 2.

1200

NiSTNCSTAR 1-5B, WTC Investigation 163
Chapter 10

S

I

X

S 10

30

25 -

20 -

15 -

Phase 1, Test 3, Floor Heat Flux Gauges

0

(Exp)

550 cm (Exp)

500 cm (FDS)

550 cm (FDS)

200 400 600 800 1000 1200

Time (s)

Figure 10-25. Comparison of heat fluxes to floor targets, Test 3.

80

X

60 -

40

20

Phase 1 , Test 4, Floor Heat Flux Gauges

0 200

500 cm (Exp)

550 cm (Exp)

500 cm (FDS)

550 cm (FDS)

400 600 800 1000 1200

Time (s)

Figure 10-26. Comparison of heat fluxes to floor targets, Test 4.

164 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

120

Phase 1, Test 5, Floor Heat Flux Gauges

100 -

80 -

500 cm (Exp)

550 cm (Exp)

500 cm (FDS)

550 cm (FDS)

0

0 1000 2000 3000

Time (s)

Figure 10-27. Comparison of heat fluxes to floor targets, Test 5.

10.3.3 Compartment Surface Temperatures

Bare-bead thermocouples were attached to the ceiling, walls, and other objects within the compartment as described in Chapter 8. In addition, several thermocouples were sandwiched between the two sheets of

Marinite that made up the outer wall. These thermocouples were placed to test the model's prediction of

conductive losses through the walls, an important component of the accounting of the overall energy

budget. Ceiling surface temperatures and the corresponding inner temperatures are shown in Figs. 10-28

through 10-37. The three points chosen for comparison were along the centerline of the compartment:

directly over the fire (x = 399 cm), 0.5 m east of the fire (x = 456 cm), and beyond the fire plume towards

the outlet (x = 612 cm).

NISTNCSTAR1-5B.WTC Investigation 165
Chapter 10

600

WTC Phase 1 , Test 1 , Ceiling Surface Temperatures

U

500 -

400 -

2 300 sH 200 A

100 -

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

612 cm (FDS)

200 400 600 800 1000

Time (s)

Figure 10-28. Comparison of ceiling temperatures, Test 1.

1200

166 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

800

600 - U

2 400 H

a.S

H

200 -

WTC Phase 1, Test 2, Ceiling Surface Temperatures

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

(FDS)

200 400 600 800 1000

Time (s)

Figure 10-30. Comparison of ceiling temperatures, Test 2.

1200

220 -|

200 -

180 -

160 - G 140 -

120 -

100 - c

c

80 - H

60 -

40 -

20 -

0

-

WTC Phase 1, Test 2, Inner Wall Temperatures

200 400 800

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

612 cm (FDS)

600 1000

Time (s)

Figure 10-31. Comparison of inner celling temperatures, Test 2.

i200

NISTNCSTAR 1-5B, WTC Investigation 167
Chapter 10

168 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

U

p

3

o

a,o

1200

1000

800 -

600

400 -

200 -

WTC Phase 1 , Test 4, Ceiling Surface Temperatures

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

612 cm (FDS)

200 400 600 800 1000

Time (s)

Figure 10-34. Comparison of ceiling temperatures, Test 4.

1200

500

400 -

300

U

p

CO

P 200 -

100 -

WTC Phase 1 , Test 4, Inner Wall Temperatures

0 200 400 800

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

612 cm (FDS)

600 1000

Time (s)

Figure 10-35. Comparison of inner ceiling temperatures, Test 4.

1200

NISTNCSTAR 1-5B, WTC Investigation 169
Chapter 10

U

a

o

1200

1000

800 -

600 -

400

200 -

2000

Time (s)

Figure 10-36. Comparison of ceiling temperatures, Test 5.

700

600 -

500 -

400

u

3

S 300

(L) G. 00 -

100 -

0

-

WTC Phase 1, Test 5, Inner Wall Temperatures

1000

399 cm (Exp)

456 cm (Exp)

612 cm (Exp)

399 cm (FDS)

456 cm (FDS)

612 cm (FDS)

2000 3000

Time (s)

Figure 10-37. Comparison of inner ceiling temperatures, Test 5.

170 NISTNCSTAR 1-5B. WTC Investigation
Simulation of Compartment Fire Phenomena

10.3.4 Gas Species Concentrations

FDS uses a mixture fraction combustion model, meaning that all gas species within the compartment were assumed to be functions of a single scalar variable. FDS solves only one transport equation for this variable, and reports gas concentrations at any given point at any given time by extracting its value from a pre-computed table. For the major species, like carbon dioxide and oxygen, the calculation indicated how

well FDS predicts the bulk transport of combustion products throughout the space. For minor species,

like carbon monoxide. FDS at the present time, does not account for changes in combustion efficiency,

relying only on a fixed yield of CO from the combustion product. In reality, the generation rate of CO

changes depending on the ventilation conditions in the compartment. Figures 10-38 through 10-47

present comparisons of species concentration predictions with measurements. As part of the

measurement procedure, water vapor was removed from the gas samples (see Chapter 6). For comparison

with the measurements, a correction was made to the simulated gas concentrations to present them on a dry basis.

oo

0.25

0.20 -

F. 0.15 -

0.10 -

o

> 0.05

0.00 -

WTC Phase 1, Test 1, Ox}'gen and CO2 Concentrations

CO2 (Ex-p)

O2 (Exp)

CO2 (FDS)

O2 (FDS)

'

1 1 1 1 1 1 1 0 200 400 600 800 1000 1200

Time (s)

Figure 10-38. Comparison of gas concentrations on a dry basis, Test 1.

NISTNCSTAR 1-5B, WTC Investigation 111
Chapter 10

0.0010

0.0008 -

o

0.0006

C3

a^ 0.0004 o>

0.0002

0.0000

0 200

WTC Phase 1 , Test 1 , CO Concentration

CO (Exp)

CO(FDS)

400 600 800 1000

Time (s)

Figure 10-39. Comparison of gas concentrations on a dry basis, Test 1

1200

0.25

0.20 -

F. 0.15 - o

?3

0.10

0.05 -

0.00 -

>00 400 600

Time (s)

\ATC Phase 1, Test 2, Oxygen and CO2 Concentrations

C02(Exp)

02(Exp)

C02(FDS)

Oo (EDS)

1 1 1 1 1 1 800 1000 1200

Figure 10-40. Comparison of gas concentrations on a dry basis, Test 2.

172 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

0 200 400 600 800 1000 1200

Time (s)

Figure 10-41. Comparison of gas concentrations on a dry basis, Test 2.

T 1 1 1 1 1 1 0 200 400 600 800 1000 1200 Time (s)

Figure 10-42. Comparison of gas concentrations on a dry basis, Test 3.

NISTNCSTAR 1-5B, WTC Investigation 173
Chapter 10

0.0014 -

0.0012 -

I

0.0010

£ 0.0008 -

s^ 0.0006 - o>

0.0004 -

0.0002 -

0.0000

WTC Phase 1, Test 3, CO Concentration

0 100

CO (Exp)

CO (FDS)

400 600 800 1000

Time (s)

Figure 10-43. Comparison of gas concentrations on a dry basis, Test 3

1200

0.25

0.20

0.15 -

0.10 -

o> 0.05

0.00

WTC Phase 1 , Test 4, Oxygen and CO7 Concentrations

CO2 (Exp)

O2 (Exp)

CO2 (FDS)

O2 (FDS)

200 400 600

Time (s)

800 1000 1200

Figure 10-44. Comparison of gas concentrations on a dry basis, Test 4.

174 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

0.0010

0.0008

c

og 0.0006

\^^ 0.0004 o>

0.0002 -

0.0000 J

T

WTlI Phase 1 ^iTest 4, CO Concentration

0 200

CO (Exp)

CO (FDS)

400 600 800 1000

Time (s)

Figure 10-45. Comparison of gas concentrations on a dry basis, Test 4

1200

0.25

0.20 -

Z 0.15

0.00 -H

\ATC Phase 1 , Test 5, Oxx-gen and CO9 Concentrations

0

CO2 (Exp)

O2 (Exp)

CO2 (FDS)

O2 (FDS)

1000 2000 3000

Time (s)

Figure 10-46. Comparison of gas concentrations on a dry basis, Test 5.

NISTNCSTAR 1-5B, WTC Investigation 175
Chapter 10

0.0014 -

0.0012 -

g 0.0010 - o£ 0.0008 -

JB 0.0006 - o>

0.0004 -

0.0002 -

WTC Phase 1, Test 5, CO Concentration

0.0000 —

r

0

CO (Exp)

CO (FDS)

1000 3000

10.3.5

2000

Time (s)

Figure 10-47. Comparison of gas concentrations on a dry basis, Test 5.

Inlet and Outlet Velocities

Velocity probes were placed in both the inlet and outlet vents. Comparisons of experimental measurement

and model prediction for Test 1 are shown in Figs. 10-48 through 10-50. Station 1 was located 10 cm

below the top of the vent just south of the compartment centerline. Station 2 was shghtly above the

midpoint of the vent. Station 3 was 15 cm above the base of the vent.

t/5 I 4

o

CD

WTC Phase 1, Inlet and Outlet Velocit\^ at Station 1

Inlet (Exp)

Outlet (Exp)

Inlet (FDS)

Outlet (FDS)

400 600 800 1000 1200

Time (s)

Figure 1 0-48. Comparison of velocity near top of vents, inlet and outlet in Test 1

.

176 NISTNCSTAR 1-5B. WTC Investigation
Simulation of Compartment Fire Plienomena

0 200 400

Time (s)

Figure 1 0-49. Comparison of velocity near center of vents, inlet and outlet in Test 1

0 200 400 600 800 1000 1200

Time (s)

Figure 10-50. Comparison of velocity near bottom of vents, inlet and outlet in Test 1,

NISTNCSTAR 1-5B, WTC Investigation 177
Chapter 10

10.4 SENSITIVITY ANALYSIS AND COMPARISON TO MEASUREMENTS

Overall, the agreement between the numerical predictions and the measurements was within experimental

uncertainty. This was due to the fact that FDS predicted the upper layer temperature in most cases to within a few percent of the experiment. From the measurements of the gas temperature and the heat fluxes

in the upper layer, it appeared that the emissivity was nearly 1 as steady state burning occurred. For

example, this was seen to occur after approximately 1,500 s in Test 5 as seen in Fig. 8-5 (Sec. 8.2). At

earlier times, the emissivity was less than 1, but was monotonically increasing. The heat flux to the walls

and objects within the upper layer was dependent on the upper layer temperature according to q = 8oT'*,

where T is the temperature, a is the Stefan-Boltzmann constant and s is the emissivity of the gas. This was a useful result because it tied the upper layer temperature to the heat flux to structural components in the upper layer. Of course, all the measurement locations needed to be examined to deteraiine where it could be assumed that the emissivity was effectively unity. To better quantify the accuracy of the

predictions, the sensitivity of the model output to the heat release rate was considered. The next sections

consider the effect of the heat release rate on the FDS calculated gas temperature, heat flux to solid

objects, compartment surface temperature, gas species concentrations, and inlet and outlet velocities.

10.4.1 Sensitivity of Gas Temperatures to Heat Release Rate

Accurate prediction of the upper layer of the compartment temperature depends on an accurate

prescription of the heat release rate ( O ) of the fire. According to an empirical correlation by McCaffrey,

Quintiere, and Harkleroad (Walton and Thomas 2003), the rise in the upper layer gas temperature AT^ in a compartment is related to the overall Q by the relation:

AT. = 6.85

1/3 o-

^0 V^O^A^T-

(10-1)

where: = area of openings (m"^

)

//(, = height of openings (m)

AJ- = total area of compartment surfaces (m^

)

A: = thermal conductivity of walls (kW/m/K)

S = wall thickness (m)

Equation 10-1 shows that the temperature rise is proportional to Q to the 2/3 power. The reported

expanded uncertainty in the heat release rate measurement was 1 1 percent (see Chapter 3). From

Eq. 10-1, an 1 1 percent uncertainty in Q corresponds to a 7 percent uncertainty in the temperature rise (= 2/3 X 11 percent). For upper layer temperatures of approximately 600 °C, this translates to roughly

±40 °C. The difference in upper layer temperatures (Figs. 10-3 through 10-12) ranged from 5 °C to

178 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

20 °C. consistent with the sensitivity of the model to uncertainty in Q . Even though there were

uncertainties in the aspirated thermocouple measurement of the temperature of approximately 7 °C, the

discrepancy between measurement and prediction can be explained solely in terms of model uncertainty

due to uncertainty^- in the Q measurement.

The McCaffi-ey, Quintiere, and Harkleroad correlation is based on fire test data with temperatures not exceeding 600 °C. In some of the experiments conducted here, however, the upper layer temperature was

as high as 750 °C. For that reason, a simulation of Test 5 was re-run using a 5 percent increase in Q

The results showed a 4.5 percent increase in the upper layer temperature, which wass higher than

predicted by the correlation. Although larger than the hypothetical increase, this difference was not

surprising given the simplifications made to derive Eq. 10-1 and the fact that it was not substantiated for higher temperamres. The sensitivity calculation confirms the idea that the upper layer temperature is nearly proportional to Q , consistent w ith a widely accepted principle of fire protection engineering.

10.4.2 Sensitivity of Heat Flux to Solid Objects to Heat Release Rate

A comparison of simulations with measurements showed that the heat flux onto surfaces in the upper

layer was very nearly given by aT' where the value of a is 5.67 x 10"" kW/m'-K' and Tis the

temperature in degrees K. In other words, the emissivity of the upper layer gases was nearly 1, not

surprising given the high level of soot. Given a standard relative uncertainty in the upper layer

temperature rise of 4.5 percent (see previous section), the standard uncertainty in absolute temperature at 600 °C was 3.1 percent, leading to an estimate for the standard relative uncertainty in heat flux of

4 X 3.1 percent = 12 percent.

Figures 10-13 through 10-22 compare the measured and predicted heat flux to the upper and lower faces

of the vertical column. The relative standard (expansion factor of 1) uncertainty of the measurement was

approximately 5.4 percent (see Table 5-1). The combined relative model and measurement uncertainty

could then be estimated as 13 percent (~ V[5.4" + 12"]). In most instances, the difference between

measurement and prediction was within 10 percent, confirming that the model was within experimental

uncertainty'.

Figures 10-23 through 10-27 compare the heat flux measurements and predictions at targets on the floor

of the compartment. Because the targets were not within the upper layer, the simple a T estimate of the

heat flux no longer applies and the assessment of experimental uncertainty was more difficult because it could not be traced back to the uncertainty in the Q . Nevertheless, in most cases, the difference between

measurement and prediction was within 10 percent. The measurement uncertainty was approximately

10 percent as discussed in Chapter 5. In cases where the difference was larger, as in Tests 4 and 5, it was

possible that the model did not capture the near-field influence of the fire itself One of the floor targets was near the fire pan, and subtle differences between the real and simulated fire could have made a

substantial impact on the heat flux to a nearby target. Warping of the fire pan may also have played a

role.

10.4.3 Sensitivity of Compartment Surface Temperature to Heat Release Rate

The accuracy of the ceiling surface temperatures was a function of the distance away from the plume

impingement point. Predictions of surface temperatures at points outside of a circle of roughly 1 m radius

NISTNCSTAR 1-5B, WTC Investigation 179
Chapter 10

agreed with measurements to about 5 percent. Predictions within the plume impingement zone had worse

agreement with experiment and were accurate to about 20 percent. This is not surprising since the point

of plume impingement changed throughout the test. Although the experiments were designed to be

symmetric along the centerline of the compartment, the inclusion of the column and instrumentation

tunnels on one side of the compartment apparently caused the flow in the compartment to be asymmetric,

and the fire was observed to lean towards the trusses, which were positioned south of the centerline

(foreground of Fig. 10-1). The simulation captured the asymmetry as well, but it did not appear to be as pronounced as in the experiment. As a result, the ceiling surface temperature predictions above the fire were noticeably higher than the experiment. Comparison of the temperatures measured between the two

sheets of Marinite showed that the temperature-temperature history occurred at a similar rate, although

the shape of the curves was somewhat different. This was attributed to uncertainty in thermophysical

material properties and inaccuracy of the ceiling temperature simulations.

10.4.4 Sensitivity of Gas Species Concentration to Heat Release Rate

The comparison of oxygen and carbon dioxide prediction and measurement on a dry basis at one

sampling point (Figs. 10-38 through 10-47) indicated that FDS was handled the bulk transport of gases

well. In the first tliree tests, the CO2 prediction is indistinguishable from the experiment, and the oxygen

prediction differed from the measurements by about 5 percent. The expanded measurement uncertainty was approximately 2 percent for CO2 and 5 percent for O2 as discussed in Chapter 6. In the simulation,

both oxygen and CO2 were tied to the mixture fraction, a single scalar for which FDS solves a transport

equation. The prediction of CO is not expected to agree with the experiment since it is assumed in the

model that its yield is constant. In reality, the yield of CO will change, usually increasing as the room

fills with combustion products.

10.4.5 Sensitivity of Inlet and Outlet Velocities to Heat Release Rate

The comparison of inlet and outlet velocities revealed that the model predicted well the mass flow of air into the compartment, but it could not predict the details of the outlet velocity profile due to a lack of

resolution. The model used 10 cm grid cells, which meant that seven cells in the vertical direction

spanned the outlet vent. Whereas the calculated inlet velocity was fairly uniform over the height of the

vent, the outlet velocity varied by at least 3 m/s from top to bottom. Plus, the position of the probes

relative to the depth of the opening mattered since the probes were designed to measure the velocity

within a certain range of angular displacement.

10.4.6 Model Sensitivity to Parameters other than Heat Release Rate

In the discussion above, it was shown that the FDS predictions of upper layer temperatures and heat

fluxes agreed with the experiment if model sensitivity and measurement uncertainty were considered. It was assumed that the sensitivity of the simulation was based solely on the uncertainty of the heat release

rate measurement. Thus, it was shown how sensitive the upper layer temperature and heat flux

measurements were to the heat release rate. In the numerical simulations there are dozens of input

parameters prescribed by the model user. Often there is no way to assess the sensitivity of these

parameters except by numerical experiment; that is, running the model with small changes to the base

parameters to see what effect these have on the predictions.

180 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Ptienomena

There are two types of input parameters: numerical and physical. Physical parameters describe the wall

materials, fuel properties, reaction stoichiometry, etc. Numerical parameters describe how the calculation

is to be performed. The most important numerical parameters were the number of grid cells in each

coordinate direction. CFD models solve an approximate form of the conservation equations of mass, momenmm and energy. The error associated with the discretization of partial derivatives on a discrete

grid is a function of the size of the grid cells and the type of differencing used. FDS uses second-order

accurate approximations of both the temporal and spatial derivatives of the Navier-Stokes equations,

meaning that the discretization error is proportional to the square of the cell size. In other words reducing

the grid cell size by a factor of 2 reduces the discretization error by a factor of four. However, it also

increases the computing time by a factor of 1 6 (a factor of two for the temporal and each spatial

dimension). Clearly, there is a point of diminishing returns as one refines the numerical mesh.

Determining what size grid cell to use in any given calculation is known as a grid sensitivity study.

Sensitivity to Grid Size

The simulations of these experiments were performed on a grid whose cells were 10 cm cubes. However,

the simulations of the fires in the World Trade Center were performed with grid cells of dimension 0.5 m

by 0.5 m by 0.4 m high. The selection of the mesh size may impact the calculation resuUs for the configuration smdied in this report. An obvious problem is that many of the obstructions in the test compartment cannot be described well on a 0.5 m grid, especially the small vents on either side of the space. A remedy for the problem is to maintain in the coarse calculations as many geometrical features as

possible. For example, the exit vents in the coarse calculation have approximately the same total area and

soffit height as in the refined calculations. The trusses, bars and column were removed from the coarse

calculation, but the barrier in front of the fire remains. Figures 10-51 and 10-52 present a comparison of

the predicted temperatures from the coarse calculation of Test 5 with the experimental measurements.

The results are ver>' encouraging, although it should be emphasized that as a rule coarsening a numerical

grid to the extent done here will not yield such good results.' The reason for the good agreement is that in the experiment, a very stable, uniform layer of hot gases was established in the upper layer of the

compartment over the course of an hour. Given a fixed Q , and a comparable convective flux of energy

through the vents, energy conservation dictates that the compartment temperatures also ought to be

comparable between calculation and experiment. Even two-zone lumped parameter models that are often

used in fire protection engineering would yield a good prediction in this case due to the uniformity of the

upper and lower layers within the compartment. In a stably stratified compartment with fairly uniform

upper and lower layer, both a CFD model and zone model will make similar predictions of temperature

because both assume conserv ation of mass and energy. The behavior of the flow field as predicted by the

momentum conservation equation is not as important in this case, so long as the enthalpy flow through

the compartment vents is captured by both the fine and coarse meshes. However, the prediction of the temperature and heat flux to various objects in the compartment near the fire plume were subject to more

uncertainty than suggested by the results shown in Figs. 10-51 and 10-52 because the flow dynamics

have been simplified by the coarse mesh. This is less of a concern for the simulations done in conjunction

with the WTC investigation because the heat flux to the structural components is being computed using

the upper layer temperature field. Indeed, the FDS simulations of the WTC fires do not include the floor trusses because these objects cannot be resolved on the coarse grid.

' Simulation of the Jiour-long experiment on the coarse grid required 30 min on a Pentium 4, 240 MHz personal computer.

NISTNCSTAR 1-5B, WTC Investigation 181
Chapter 10

U

o;-! 3

o

SoH

1000

800 -

600 -

400

200 -

WTC Phase 1, Test 5, West Aspirated TCs (Coarse Grid)

365 cm (Exp)

215 cm (Exp)

34 cm (Exp)

365 cm (FDS)

215 cm (FDS)

34 cm (FDS)

1000 2000

Time (s)

3000

Figure 10-51. Coarse grid prediction of gas temperatures on the intake side of the

compartment for Test 5.

u

o

OH

1000

800 -

600 -

400 -

200 -

WTC Phase 1, Test 5, East Aspirated TCs (Coarse Grid)

365 cm (Exp)

215 cm (Exp)

35 cm (Exp)

365 cm (FDS)

215 cm (FDS)

34 cm (FDS)

1000 2000 3000

Time (s)

Figure 10-52. Coarse grid prediction of gas temperatures on the exhaust side of the

compartment for Test 5.

182 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Phenomena

Beyond the numerical grid, there are numerous physical parameters that the user must prescribe to simulate the experiments. First, there are the physical dimensions of the compartment plus the objects

within. Because one of the objectives of the experiments was model validation, the geometry was very

simple. With a 1 0 cm numerical grid, the compartment was modeled to an accuracy of ±5 cm. and there were no detectable differences in results based on small adjustments made to objects to fit the nearest grid

cell. With the 50 cm grid, however, there were significant differences in the results if proper attention was not paid to adjusting the vents to maintain the area and soffit height. A 20 cm shift in the height of

the top of the \'ent lead to a change in the hot upper layer depth of the same amount, resulting in changes

of upper layer temperature on the order of 100 °C. Based on these findings, the cell size for the WTC

simulations was selected to maintain the window area and height. The external columns of WTC 1 and

WTC 2 were spaced by the designers 1 .016 m (40 in.) apart. The horizontal dimensions of the grid cells used were 50 cm, yielding an alternating pattern of solid column, vent, column, vent, etc.

Sensitivity to Radiative Fraction

Along with the Q . another parameter of importance to fire models is the radiative loss from the fire. Room-scale fires t\'pically radiate about 1/3 of their energy, while the other 2/3 of the energy rises to form

the thermal plume. Often the radiative loss fi'action is a user input to fire models because a prediction of

it requires a fairly detailed model of the combustion processes. FDS uses a partial approach to the

problem. Based on the predicted fire temperature and chemical composition, it computes the source term

in the radiative transport equation, thus predicting the radiative loss from the fire. However, since the

temperature and chemical composition within the fire were subject to uncertainty due to the coarseness of

the numerical grid, the algorithm uses a user-prescribed radiative loss fraction instead of its own

prediction, if that prediction is lower than what the user thinks it ought to be. In short, the user has the

power to over-ride the FDS prediction if it is warranted. To ensure that the results of the simulations were

not sensitive to the user-prescribed radiative loss fraction, several simulations were performed with

prescribed loss fractions between 0.3 and 0.4. There was no discemable difference in the results. Again,

this is not always the case. Compartment geometry is important, particularly the ceiling height. Also, in these tests, the hot upper layer is as important a radiator as the fire. The prescribed loss fraction only

pertains to the fire.

Sensitivity to Thermal Properties of the Compartment Surface Material

Other physical parameters set by the user include the thermal properties of the wall material. In this case, the walls were lined with a 2.5 cm (1 in.) layer of calcium silicate board (Marinite). The manufacturer

provided the thermal properties of the material, including the thermal conductivity. To determine the

sensitivity of the model to this parameter, a simulation was performed with the thermal conductivity

doubled from its listed value of 0.12 W/m-K to 0.24 W/m-K. Measurements by Taylor et al. (2003) show

that this is an order of magnitude larger than the uncertainty in the thermal conductivity of the Marinite.

The predicted upper layer temperatures of Test 5, the hour-long 3 MW fire, decreased only 4 percent due

to the doubling of the conductivity. This suggests that these simulations were not especially sensitive to the thermal properties of the wall materials, although this is not always true. In this case, of the 3 MW of

energy generated in the compartment by the fire, more than half flowed out the vents, either as thermal

radiation or hot gases. The increase in conductivity means that more energy was being conducted through

the walls, but since the walls were conducting only a small amount of the total energy, the increased loss

NISTNCSTAR 1-5B, WTC Investigation 183
Chapter 10

did not affect the overall compartment temperature appreciably. If the compartment were of different

dimension, aspect ratio, or ventilation, the thermal conductivity might play more of a role.

Sensitivity to Smoke Yield

Parameters associated with the combustion process were also tested. The smoke yields of the heptane and

the heptane/toluene mixture were obtained during free bums of the two fuels. The smoke yield is the fraction of fuel mass that is converted into soot via the combustion process. In a ventilation-limited fire, the soot yield will usually increase due to the decrease in combustion efficiency. What effect can the

increased soot yield have on the temperature predictions? The soot yield for the heptane fires was

measured to be 0.014. Increasing this value to 0.10 produced a I percent increase in the upper layer

temperature prediction. The reason for this is that the upper layer is nearly optically thick, that is, the

emissivity of the smoke is nearly 1 . Adding more soot to the layer only blackens what is already

essentially black. The radiative heat transfer within the layer and to external targets is not affected.

10.5 COUPLING WITH THERMAL-STRUCTURAL MODELS

The computed values of the upper layer temperature, absorption coefficient, depth of the hot layer and the

lower layer temperature were computed at a representative slice at a particular height above the floor,

recorded at each (x, y) grid point, and saved in an ASCII file for utilization by the NIST Fire Structure

Interface (FSI) (Prasad and Baum 2005), which used this infonnation as input in the detennination of heat

flux boundary conditions for finite-element models that simulated the temperature distribution in the

structural steel components (see discussion in Chapter 11). The FSI calculated the transfer of radiant and

convective heat from FDS to a coupled, transient, three-dimensional finite-element model for the thermal

response of structural components. The structural components were either simple (e.g., bare steel) or complex (e.g., insulation-coated steel). A detailed description of the definitions of the absorption

coefficient and the depth of the hot layer are given in detail in the FDS user's guide (McGrattan 2004).

To provide an estimate of model sensitivity on the calculated temperature distribution, FDS calculation

results were also performed for varying heat release rate and grid resolution. The results of those

calculations are discussed in Chapter 1 1

.

10.6 SUMMARY

A series of geometrically and physically complex experiments were conducted to test the NIST FDS fire model. For the FSI methodology that was applied to detennine the temperature of structural steel that was in the fire compartment, the upper layer temperature calculated by FDS was a critical parameter. In

general, most predictions were within 10 percent, except for some near-field quantities (20 percent). In

particular,

• FDS predictions of the upper layer gas temperature were within 10 percent of the

measurements. This difference can be attributed to the relative expanded uncertainty in the

HRR measurement.

• FDS predictions were in agreement with measured gas velocities at the compartment inlets.

184 NISTNCSTAR 1-5B, WTC Investigation
Simulation of Compartment Fire Plienomena

• Because of limited spatial resolution at the outlet window, the steep gradient in velocity was

not captured by the model, but the integrated mass flux was.

• FDS predictions were in agreement with the measured gas-phase volume fraction of O2 and

CO2 in the upper layer.

• FDS predicted the leaning of the fire plume caused by asymmetric obstructions in the

compartment, but underestimated the extent of the leaning. This adversely impacted

FSI/ANSYS predictions of the thermal behavior of structural components at some locations near the fire.

• FDS predictions of heat flux in the upper layer were within 10 percent of the measurements.

This difference can be attributed to the relative standard uncertainty in the HRR

measurement.

• FDS predictions of the heat flux to the floor and the column were less accurate (within

20 percent) for surfaces facing the fire than surfaces facing away. This may be attributed to issues associated with prediction of the flame leaning and the radiative absorption of the

surrounding gases.

10.7 REFERENCES

Hamins, A., A. Maranghides, and G. Mulholland. 2003. The Global Combustion Behavior of 1 MWto 3

MIV Hydrocarbon Spray Fires Burning in an Open Environment, NISTIR 7013, National Institute of

Standards and Technology, Gaithersburg, MD, June.

McGrattan. K.B., ed. 2004. Fire Dynamics Simulator (Version 4), Technical Reference Guide. NIST

Special Publication 1018. National Institute of Standards and Technology, Gaithersburg, MD, July.

McGrattan, K.B., and Forney, G.P. 2004. FDS User's Guide. NIST Special Publication 1019. National

Institute of Standards and Technology, Gaithersburg, MD.

Prasad, K., and H.R. Baum. 2005. "Coupled Fire Dynamics and Thermal Response of Complex Building

Structures," Proc. Combustion Institute 30, 2255-2262.

Taylor, R.E., H. Groot, and J. Ferrier. 2003. Thermophysical Properties ofPVC, PE, and Marinite,

Report No. TPRL 2958, Thermophysical Properties Research Laboratory, Inc., April.

Walton, W.D., and P. H. Thomas. 2003. Estimating Temperatures in Compartment Fires. Fire

Protection Handbook, 3'^ Ed. National Fire Protection Association, Quincy, MA.

NISTNCSTAR 1-5B, WTC Investigation 185
Chapter 10

\

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186 NISTNCSTAR 1-5B, WTC Investigation
Chapter 11

Simulation of the Thermal Response of

Structural Components

11.1 INTRODUCTION

The large-scale experiments measured the thermal response of structural steel components within a steel frame compartment lined with calcium silicate boards. As described in Chapter 3, the fire was generated

using liquid hydrocarbon fiiels with nominal heat release rates of approximately 2 MW to 3 MW. The

layout of the steel components within the compartment and the location of the temperature measurements

on the structural components are described in detail in Chapter 9. Four steel components, consisting of

two trusses, one thin-walled tubular column, and a simple rod, were placed in the compartment. The steel components were uninsulated in the first three experiments and were coated with a fibrous sprayed fire- resistive material (SFRM) in the last three experiments. The temperatures of the surface of the steel components and the SFRM were measured and served as a data set to validate the prediction of the

temperature rise of the structural steel components. In this chapter, a comparison is made between the

simulated and measured temperatures on the steel and the SFRM surfaces.

11.2 methodology

Simulations of the effects of severe fires on the structural integrity of buildings requires a close coupling

between the gas-phase enthalpy release and transport phenomena and the stress analysis in the load

bearing materials. The connection between the two is established primarily through the interaction of the

heat transfer between the solid and gas phase with the conduction of heat through the structural components. This process is difficult in large, geometrically complex buildings due to the wide disparity

in length and time scales that must be accounted for in the simulations. Typical length scales range from

60 m, characteristic of the WTC buildings, to 1 cm or 2 cm, characteristic of structural components and

insulation. Typical Fire Dynamics Simulator (FDS) time steps are on the order of 10"^'

s, whereas total simulation times for World Trade Center (WTC) 1 are about 10'' s. These disparities create a challenge

both in terms of modeling and computational issues. The National Insfitute of Standards and Technology

(NIST) Fire Strucmre Interface (FSI) has been developed to overcome these difficulties in the analysis of

the collapse of the WTC buildings (Prasad and Baum 2005; NIST NCSTAR 1-5G).

NISTNCSTAR 1-5B, WTC Investigation 187
Chapter 1

1

The FSI uses information predicted by the NIST FDS, described in Chapter 10 of this report, as boundary

conditions for radiative and convective heat transfer to sub-grid scale structural components (Prasad and

Baum 2005). The FSI uses a simple radiative transport model that assumes that the compartment is locally divided into a hot, soot laden upper layer and a cool relatively clear lower layer. The properties of

the two layers are extracted from temporal averages of the results obtained from FDS. Explicit fomiulae

for the heat flux were obtained as a function of temperature, hot layer depth, soot concentration, and

orientation of each structural element. These formulae were used to generate realistic thermal boundary

conditions for the coupled transient three-dimensional finite-element code. The code was then used to generate solutions for the heating of the complex structural assemblies. The goal of this portion of the

study is to compare the numerically predicted temperature profiles within the structural steel components

and the SFRM with the experimental measurements.

The computed upper layer temperature, absorption coefficient, depth of the hot layer, and ambient

temperature, as computed by FDS, were recorded at 30 s intervals for grid points in the upper layer and

stored as an ASCII text file for computing the combined radiative and convective fluxes incident on the

structural components. The time constant for the temperature of the steel was on the order of minutes, so

the 30 s time interval was sufficiently resolved. The file was subsequently read into the ANSYS (2004)

finite-element program. Two major simplifications were introduced so that the radiative fluxes could be

applied to the thermal analysis of structures. First, the concept of a gray gas whose properties are independent of spectral wavelength was employed. The second simplification was that the enclosure

geometry induced a vertically stratified distribution of temperature and combustion products. As the fire developed in the enclosure, the hot layer had spatial variations in the above properties that varied much

more rapidly in the vertical than in the horizontal direction. The radiative transport equations can be

simplified, and the radiative flux to any surface can be computed. These radiative fluxes are a function of

the location and orientation of the structural element. The radiative flux also depends on the upper layer

temperature, the absorption coefficient, the hot layer depth, and the lower layer temperature that were

obtained from the fire dynamics simulation in the form of an input file and were stored as a function of

space and time. Figures 1 1-1 and 1 1-2 are representations of the FDS output values for these parameters

after 500 s in Test 1. The radiative flux to the horizontal surface was obtained in tenns of Exponential

Integrals En(z), whereas that to a vertical surface was obtained in tenns of the rth iterated integral of the

Modified Bessel function Ko(z) as described in detail in Prasad and Baum (2005). The numerical solution

and post-processing procedures are also described in Prasad and Baum (2005).

The convective fluxes incident on the structural components were also considered and were obtained from

the temperatures of the upper layer and the ambient temperature (Prasad and Baum 2005). Structural

components in the hot layer were subject to convective fluxes, with bulk temperature values equal to the

local instantaneous value of the temperature in the hot layer. Structural components not in the hot layer were subject to convective fluxes, with bulk temperature values equal to the local instantaneous value of

the ambient temperature. A convective heat transfer coefficient value of 25 W/m"-K was assumed for computing the convective fluxes to the structural components (Eurocode 1994).

188 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Slice u-tmp C

565

520

Figure 11-1. FDS simulation of the upper layer temperature on the plane 3.01 m above

the floor (top figure) and the layer interface height (bottom figure) after 500 s in Test 1.

NISTNCSTAR 1-5B, WTC Investigation 189
Chapter 1

1

190 NISTNCSTAR 1-5B, WTC Investigation
Simulation of tlie Thermal Response of Structural Components

In order to perform a thermal analysis, the thermophysical properties of the steel and SFRM were needed,

including the density, thermal conductivity, and heat capacity. These quantities are generally functions of

temperamre. The steel components were constructed of A572 steel, which has thermophysical properties

ver>' similar to A242 steel (NIST NCSTAR 1-3E). The thermophysical properties for A242 steel were

obtained from NIST NCSTAR 1-3E. Figure 1 1-3 shows the A242 steel thermophysical properties as a fimction of temperature. The specific heat increases with temperature, whereas the thermal conductivity

decreases as a function of temperature. Figure 1 1^ shows the thermophysical properties obtained by

Harmathy (1983) for the BLAZE-SHIELD D GT SFRM used in the experiments. The thermophysical

properties for Marinite, shown in Fig. 1 1-5, were obtained from Taylor et al. (2003).

mod

Figure 11-3. Temperature dependent thermophysical properties for A242 steel used in

the thermal analysis.

= xn.i

ucc- 4»>

coc

3*30

i: ii

lUC 4

*»x ^ «Sf) «W

Titi

Figure 11-4. Temperature dependent thermophysical properties of the SFRM used in the

experiments and thermal analysis.

:3.>

uoo

lose

loan «>

w»

eo«»

7y)

...»—

Tesperature (K) rature 'Ki Figure 11-5. Temperature dependent thermophysical properties for Marinite.

NIST NCSTAR 1-5B, WTC Investigation 191
Chapter 1

1

In order to perform the thermal analysis, initial and boundary conditions were applied on the structural components. The initial temperature of the steel and the SFRM was assumed to be 300 K. The radiative

and convective flux boundary conditions, as computed from the FSI plane layer analysis, were applied to the external faces of the structural components. The external faces were also covered with surface effect components to allow for re-radiation back to the ambient atmosphere. The ends of the bar were insulated

for experimental purposes, and therefore, a zero flux boundary condition was applied in the computations.

The initial time step was set to 1.0 s, while the minimum and maximum time steps were set to 0.1 s and

20.0 s, respectively.

The finite-element model ANSYS 8.0 was used to model the themial response of the structural components. Figure 1 1-6 is a finite-element representation of the 3.05 m (10.0 ft) long steel bar. The bar was meshed with "SOLID 70" brick components (an element with eight nodes each with a single degree

of freedom). The element was used primarily for a tliree-dimensional transient themial analysis. The bar was divided into 30 uniform divisions along its length. Figure 1 1-7 shows the components through a cross-section of the steel bar and SFRM. The bar diameter was 2.54 cm ( 1 .00 in.). The thickness of the

SFRM on the various components tested in the experiments was reported in Table 9-1. In Test 5, for example, the SFRM had a mean thickness of 2.30 cm ± 0.55 cm. In the calculations, the SFRM thickness was assumed to be uniform over the entire length of the component. A sensitivity analysis presented in Sec. 1 1.4.3 was performed to assess the validity of this assumption. The elements in Fig. 1 1-7 have been

color coded with the components in blue representing the steel and those in violet representing the fire protective coating. Simulations were performed with approximately 6 to 10 elements through the

thickness of the fire resistive coating. Surface elements were used to model the re-radiation back to the

ambient atmosphere as well as to apply convective flux boundary conditions to the structural components.

Figure 1 1-8 shows a typical calculated temperature distribution through a bar. The uncertainty in the

predicted themial response of the stmctural steel components is discussed in Sec. 1 1.4 in terms of model

sensitivity to uncertainties of various input parameters used in FDS and the FSI models.

Figure 11-6. Finite-element model for the 3.05 m (10.0 ft) long steel bar, showing the

elements distributed uniformly along the length of the bar.

192 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

NISTNCSTAR 1-5B, WTC Investigation 193
Chapter 1

1

1 1 .3 COMPARISON OF SIMULATIONS WITH MEASUREMENTS

The simulations were compared to two representative experiments, namely Tests 1 and 5. Test 1 was a

2 MW heptane fire with a 15 min duration in which the steel components were bare (uncoated). Test 5 was a 3 MW heptane fire with a 50 min duration in which the steel components were insulated with

various thicknesses of SFRM (see Chapter 9). The comparisons between simulations and measurements

focused on the peak temperatures attained by the steel rather than the amount of time it took to attain a

particular temperature. This approach was taken because the peak steel temperatures varied widely from

location to location and from test to test, making selection of a specific reference temperature impractical. In addition, the temperature rise in a given period of time was considered more important from the

perspective of structural stability than from a study of component response.

11.3.1 Testi

Figures 1 1-9 and 11-10 compare the numerical simulations (symbols) with the measured steel surface

temperatures (solid lines) at four locations on bar A and bar B. respectively, during Test 1. The location

of the origin was taken as the north end of the bar (at the top of the bar in the figures). A detailed

discussion of the steel temperature measurements, including the uncertainty, which was bounded from

-9 °C to +1 °C, is presented in Chapter 9 of this report. The shape of the simulated time-temperature

results was similar to the measurements. Both curves increased in a monotonic fashion from ambient

values at time equal to zero, to finite values, which obtained a maximum when the fuel was stopped, and

subsequently decreased, also in a monotonic manner. The general character of these results was

representative of all of the tests. For most locations, the absolute difference between the numerical

predictions and the experimental data was less than 20 °C at any time, whereas for other locations the

difference was as large as 100 °C. This was because the asymmetry of the fire plume was not precisely

predicted by FDS (see Fig. 8-11), so the difference between the temperature measurements and the model

simulations differed along the bar at some locations. For the WTC fires that were severe enough to cause

structural damage, the fire had a far wider spatial extent, and the calculation results will be less sensitive

to the exact location of the fire plume.

Here, and subsequently in this chapter, the relative difference between the measured and simulated

temperatures were compared by normalizing the temperature difference by the average of the measured

and predicted temperature change from ambient conditions. The relative difference between the peak

measured and simulated temperatures was approximately less than 20 percent for bar A and less than

10 percent for bar B. These differences are discussed in terms of model and measurement uncertainty in Sec. 1 1.4 below.

194 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

GOO

ExpGrimGntal Data:

3000

Figure 11-9. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations on bar A in Test 1.

600

O

^ 300

z>

COQ-200

a>

100.

j ExperimGnts

1 ft

3l Data; - 4 ft. - 6 ft. 9 ft.

NufTiGrical RGSuits

a

1 ft. 4 ft. 5 ft.

1 1 1 1 1 1 1 1 1 1 1 1

, , , , 1 1 1 1

9 ft.

, , , , , , , , 00 500 1000 1500

Time fs) 2000 2500 5000

Figure 11-10. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations on bar B in Test 1.

NISTNCSTAR 1-5B, WTC Investigation 195
Chapter 1

1

Figure 1 1-1 1 shows a perspective and a cross sectional view of the finite-element model that was used to model the thermal response of the column. The cross sectional view shows the steel column with a unifomi SFRM wrapped around it. A portion of the ceiling and floor were also included in the model in an effort to capture the radiative exchange between the column and these elements. Figure 11-12 shows

typical calculation results for the temperature (in K) plotted as isocontours on the surface of the SFRM

and the Marinite ceiling 2,000 s into the simulation of Test 5.

Figure 11-11. Perspective and top view of the finite-element model of

the steel column with SFRM.

Figure 11-12. Temperature contours plotted on the surface of the SFRM

and the ceiling after 2,000 s in Test 5.

196 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Figures 11-13 and 11-14 compare the numerical predictions with measurements of the steel surface

temperature for locations on the North, South, East, and West faces of the bare column during Test 1 at heights of 3.69 m and 2.13 m above the floor, respectively. The location of the column relative to the

compartment is shown in Figs. 1 1-1 and 1 1-2 and is schematically shown in Figs. 2-5 and 5-3 in Chapters 2 and 5, respectively. In Figs. 1 1-13 and 1 1-14, both the numerical predictions and the

experimental measurements have been color coded for the four faces of the column. (The west face of the column faced the compartment inlet.) The highest predicted and measured temperatures occurred on the

south face of the column. At one location on the south face (3.69 m above the floor), the largest

difference between the simulations and measurements was immediately after the start of the test, when the

temperatures were still relatively low. but the rate of temperature change was large. The largest

temperature difference generally occurred, however, just before the fuel flow was stopped. In Test 1 , this was 900 s after ignition. For most locations, the relative differences between the peak measured and

simulated temperamres was less than 5 percent, whereas for some locations the differences were as large as 10 percent. These results are discussed in terms of model and measurement uncertainty in Sec. 1 1.4 below.

Figure 11-13. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 3.69 m above the floor on the column in Test 1,

NISTNCSTAR 1-5B, WTC Investigation 197
Chapter 11

500

450

400

350

O 300

CDD 250

-+-» 2

g_200

E

«^ 150

100

50

0

"^^00 " 0

" ^ 500 iOOO ^lIOO 2000 2500 3000

Time (s) Figure 1 1-14. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 2.13 m above the floor on the column In Test 1.

Figures 1 1-15 and 1 1-16 show the finite-element representation of the truss used in the thermal analysis.

Figure 1 1-15 is a cross-sectional view of the truss, showing the top and bottom steel flanges (L-shaped

sections). Typical calculation results are shown in Fig. 1 1-17 plotted as temperature contours on the

surface of the SFRM for truss A 1,000 s into Test 5.

Figure 11-15. Finite-element model of the insulated steel truss, the

ceiling and the floor used in the thermal analysis of Test 5.

Experimental Data

North Face

East Face

South Face

:

; - West Face

if/f

Numerical Results

North Face

East Face

South Face

-1111 1 1 1 1 1 1

1 1

1

1 1 1 1 1 1 1 1 1

1 1

198 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Figure 11-16. Finite-element representation of the insulated steel truss (blue), the

SFRM (violet), and the ceiling (red) used in the thermal analysis of Test 5. H 300 S 400

-00

Figure 11-17. Temperature contours (in K) on the ceiling, floor, and the surface of the

SFRM on truss A after 1,000 s in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 199
Chapter 1

1

Figures 1 1-18 through 1 1-23 compare the numerical simulations and the temperature measurements for various locations on the surface of the steel trusses during Test 1. Figures 1 1-18 through 1 1-20 show

results at various locations for three heights (2.89 m, 3.29 m and 3.70 m) on truss A. Figures 1 1-21

through 1 1-23 show analogous results for truss B. The locations of the thermocouples are illustrated in Figs. 9-6a and 9-6b. For most locations, the maximum difference between the measurements and the

simulations was less than 10 percent, which typically occurred when the fuel flow was stopped. For a small number of locations, the difference was as large as 27 percent.

650

600

550

500

450

o 400

350

"c5 u.0 300

Q.

em 250

200

150

10050

00

Experimental Data

1000 1500 Time (s) 3000

Figure 11-18. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 2.89 m above the floor on truss A in Test 1.

1000 1500 Time(s)

3000

Figure 11-19. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 3.29 m above the floor on truss A in Test 1.

200 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Figure 11-20. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 3.70 m above the floor on truss A in Test 1.

700

Experimental Data

m 0 500 1000 1500 2000 2500 3000

Time (s)

Figure 1 1-21 . Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 2.89 m above the floor on truss B in Test 1.

NISTNCSTAR 1-5B, WTC Investigation 201
Chapter 1

1

Figure 11-22. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 3.29 m above the floor on truss B in Test 1.

650

1000 1500

Tjme(s)

3000

Figure 11-23. Comparison of numerical simulations with measurements for the steel

surface temperature at four locations 3.70 m above the floor on truss B in Test 1.

202 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

11.3.2 Tests

Figures 1 1-24 through 1 1-37 compare the numerical simulations and the temperature measurements

for various locations on the surface of the steel components and on the SFRM surface in Test 5. Figures 1 1-24, 1 1-26 through 1 1-28, and 1 1-32 through 1 1-34 compare the simulations with

measurements at locations on the surface of the SFRM for the bar, column and truss. Although these

results have less direct bearing on the stability of a steel component subjected to a fire, they are of interest because the SFRM surface temperature is related to the steel surface temperature. For most locations, the maximum difference between the measurements and the simulations of the SFRM surface temperature

typically occurred during the first 600 s after ignition, when the measurement uncertainty was largest (see

Chapter 8). During the steady burning period, 1,000 s after ignition, the temperature difference was

topically less than 10 percent, although it was as large as 20 percent to 30 percent in a small number of

cases.

The rate of temperature rise on the steel was significantly reduced when the SFRM was present.

Figures 1 1-25, 1 1-29 through 1 1-31, and 1 1-35 through 1 1-37 compare the simulations with

measurements at locations on the steel surface for the bar, column, and truss in Test 5. For the bar and the

column, the maximum difference between the measurements and the simulations was less than

15 percent, which generally occurred shortly after the fuel flow was stopped. For truss A, the maximum

difference between the measurements and the simulations was typically less than 20 percent. At one

location, however, the temperamre difference was as large as approximately 30 percent. The reason for the magnitude of this difference appears to be subtleties associated with predicting the exact location of

the flame, as mentioned previously. Other reasons for the differences between the numerical simulations

and the temperamre measurements are discussed in Sec. 1 1.4 in ternis of the uncertainty in the

measurements and a sensitivity' analysis of the model.

1000

800

— 600 O,

0)k_2 400

<DCL

r—C

200

-20 000

Expen mental Data

1000 2000 30OO Time (s) 4000 5CX)0 6CX)0

Figure 11-24. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations on the insulated bar in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 203
Chapter 1

1

Figure 11-25. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations on the insulated bar in Test 5.

800

600 -

o

400 - 2

0

Q.E

.CD 200 -

Experimental Data

North Face

East Face

South Face West Face

2000 3000 Tlme(s)

6000

Figure 11-26. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 0.77 m above the floor on the

insulated steel column in Test 5.

204 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

800

600

O

Q.

400

200

Experimental Data ——East Face

South Face

West Face

Numerical Results

North Face

East Face

South Face - ' Wfest Face

-1000 1000 2000 3000

Time(s)

4000 5000 6000

Figure 1 1-27. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 2.13 m above the floor on the

insulated steel column in Test 5.

800

600

o

o

• TOi_O

Q.

E

o

400 -

200

Experimental Data

North Face

East Face

South Face

West Face

Numerical Results

North Face

East Face

South Face

\Afest Face

JL -1000 1000 2000 3000

Time (s) 4000 5000 6000

Figure 11-28. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 3.69 m above the floor on the

insulated steel column in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 205
Chapter 1

1

300

250

200

O

o

Q.

E

o

I150

100

50

-51

Experimental Data

North Face —-—— East Face

South Face

West Face

000

NurrKjrica! Results

North Face

East Face

South Face

V\fest Face

_L J I I I I L 1000 2000 3000

Time (s) 4000 5000 6000

Figure 11-29. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 0.77 m above the floor on the

insulated column in Test 5.

300

250 -

200

O

^ 150

03i_

Q)|-100

50

Experimental Data

North Face

East Face

South Face

West Face

Numerical Results

North Face

East Face

South Face

= \Afest Face

_L _L TOOO 1000 2000 3000

Time (s) 4000 5000 6000

Figure 1 1-30. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 2.13 m above the floor on the

insulated column in Test 5.

206 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

300

^0 -

200 - o

(D

0)Q.

150 -

100 -

50 -

Experimental Data

North Face

East Face

South Face West Face

_L _L

North Face

East Face

South Face

\Afest Face

I L

000 0 1000 2000 3000 4000 5000 6000

Time (s)

Figure 1 1-31 . Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 3.69 m above the floor on the

insulated steel column in Test 5.

1200

Experimental Data

TU-1

. TU-2

TU-3

TU-4

Numerical Results

TU-1

TU - 2

TU-3

TU-4

2000 3000

Time (s) 6000

Figure 11-32. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 3.70 m above the floor on truss A in

Test 5.

NISTNCSTAR 7-56, WTC Investigation 207
Chapter 11

1200

1000

800

O

CD I

600

Q.

i5B 400

200

0 -

Experimental Data

4 TM -

1

TM-2

TM-3

TM-4

Numerical Results

TM-1

TM-2

TM-3

TM-4

_L 1000 2000 3000

Time (s) 4000 5000 6000

Figure 11-33. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 3.29 m above the floor on truss A in

Test 5.

1200

1000 -

800 - o

0) 5 600

£^ 400

200 -

0 -

2000 3000

Time (s) 6000

Figure 11-34. Comparison of numerical simulations with measurements for the

temperature of the SFRM surface at four locations 2.89 m above the floor on truss A in

Test 5.

208 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

600 -

9.400

o

Q.

i 200

Experimental Data

^TU-1

TU-2

TU-3

TU-4

Numerical Results

TU-1

TU-2

TU-3

TU-4

_L _L 1000 2000 3000

Time (s) 4000 5000 6000

Figure 11-35. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 3.70 m above the floor on truss A in

Test 5.

600 -

9-400

k—Q.o 200

Experimental Data TM-1

"nvi-2

^TM-4

Numerical Results

TM-1

TM-2

TlVl-3 nvi-4

1000 2000 3000

Tinne (s) 4000 5000 6000

Figure 11-36. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 3.29 m above the floor on truss A in

Test 5.

NISTNCSTAR 1-5B, WTC Investigation 209
Chapter 1

1

Figure 1 1-37. Comparison of numerical simulations with measurements for the

temperature of the steel surface at four locations 2,89 m above the floor on

truss A in Test 5.

The levels of agreement between the calculated and the measured values of the peak temperature for the

various steel components in Tests 1 and 5 are summarized in Tables 1 1-1 and 1 1-2, respectively. The

table lists the difference between the peak values of the measured and the stimulated steel surface

temperatures, nonnalized by the average of those temperatures on the bare or insulated steel surface. The

peak temperatures typically occurred 850 s after ignition in Test 1 (see, for example, Fig. 1 1-10) and

3,300 s after ignition in Test 5 (see, for example, Fig. 1 1-36). In terms of thermal load on a structural element, both the maximum temperature and the temperature gradient may be important. The data in the

tables represent the percentage agreement for the hottest and coolest temperature locations on each

element. A negative percentage in the tables indicates that the numerical simulation was less than the

measured value, and a positive percentage indicates that the simulation was greater than the measured

value. The average of the absolute values of the percentages is given at the bottom of each table. The

average of the percentages are shown in parenthesis. These values were negative for all four cases, indicating that the numerical predictions consistently underpredicted the measurements by a finite, but

small, percentage. For the uninsulated steel components in Test 1 , the average of the absolute value of

the differences was less than 8 percent, although some individual differences were as large as 12 percent

to 1 5 percent for the hottest temperature locations and as large as 1 6 percent to 20 percent for the coolest

temperature locations. For the steel components with SFRM in Test 5, the average of the absolute values

of the differences was less than 20 percent, although some individual differences were as large as 20 percent to 30 percent for the hottest temperature locations and as large as 30 percent to 33 percent for the coolest temperature locations.

210 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Table 11-1. Percentage difference between peak values of the measured and the

simulated steel surface temperatures at the hottest and coolest locations on various

uninsulated components in Test 1 Element (\'ertical Position) Hottest Location Coolest Location

Bar A -10 % 3 %

Bar B 8 % -2 %

Column (2.13 m abo\ e floor) 2 % -2 %

Column (3.69 m abo\ e floor) -3 % -6 %

Truss A (2.89 m above floor) -4 % 2 %

Truss A (3.29 m above floor) -6% -8 %

Truss A (3.70 m abox e floor) 1 % -10%

Truss B (2.89 m above floor) -12 % 10%

Truss B (3.29 m above floor) -15 % -20 %

Truss B (3.70 m above floor) -6% -16%

Absolute \'alue of Average (Average) 6.7 % (-4.5 %) 7.9% (-5.1 %)

Table 11-2. Percentage difference between peak values of the measured and the

simulated steel temperatures for the hottest and coolest locations on the steel surface of

various components with SFRM in 1rest 5. Element Hottest Location Coolest Location

Bar' 8% 3 %

Column (0.77 m above floor) -20 % 3 %

Column (2.13 m above floor) -11 % -13 %

Column (3.69 m above floor) -30 % 25 %

Truss A (2.89 m above floor) 15 % -15 %

Truss A (3.29 m above floor) -15 % -33 %

Truss A (3.70 m above floor) -10% -30 %

Absolute Value of A\erage (Average) 16% (-9%) 20% (-11 %)

a. Only one bar was tested in Test 5.

The temperatures of the steel surfaces in Tests 1 and 5 were controlled by the transient heat flux boundary

conditions, which were driven mainly by the upper layer gas temperature. The magnitude of the disagreement shown in Tables 1 1-l and 1 1-2 cannot be explained solely by uncertainty in the

measurement of the surface temperature of the uninsulated steel in Test 1, which was approximately

2 percent (bounded by +1 °C and -9 °C), nor by the uncertainty in the measurement of the surface

temperature of the steel with the SFRM in Test 5, which was approximately 1 percent (±5 °C) for the

highest temperature measured.

Table 1 1-2 shows the percentage difference between peak values of the measured and the simulated steel surface temperatures at the hottest and coolest locations on various uninsulated components in Test 5. The

percentage differences were much higher than analogous comparisons for Test 1 (shown in Table 1 1-1). Test 1 and Test 5 were somewhat similar. Although both tests used heptane as a fuel. Test 5 had a higher

heat release rate (3 MW rather than 2 MW) and was longer in duration (50 min rather than 14 min) than Test 1 . Other experimental parameters did not change from Test 1 to Test 5. The agreement between

NISTNCSTAR 1-5B, WTC Investigation 211
Chapter 1

1

FDS simulations of the upper layer gas temperature and the experimental measurements was very similar

for Test 1 and Test 5. A review of the data in Chapter 10 shows that the percentage difference in the peak

value of the upper layer gas-phase temperatures on the intake and exhaust sides of the compartment was

2 percent and 5 percent in Test 5 (see Figs. 10-1 1 and 10-12), and 5 percent and 3 percent in Test 1 (see

Figs. 10-3 and 10-^). The difference in the nature of the agreement between the FSI and the

measurements for Test 1 and Test 5 is attributed to the effects of the SFRM, which was present in Test 5, but not in Test 1 . Variation in the SFRM thickness was highlighted in Chapter 9 in which the coefficient

of variation was reported to be 24 percent, 7 percent, and 27 percent for the bars, column, and truss A in Test 5, respectively (see Tables 9-2, 9-3, and 9^a.). The uncertainty in SFRM thickness can have a

large impact on the predicted steel temperatures.

11.4 MODEL SENSITIVITY

Model sensitivity was considered in an effort to understand the effect of uncertainty of the various input

parameters used in the NIST FDS and FSI models on the calculated thermal response of the structural

steel components. This information is useful for understanding the differences between the simulated and

measured temperature behavior of the steel members in the fire tests. A key focus of the sensitivity study was the effect of the gas-phase heat release rate and the FDS grid resolution on the calculated thennal

response of the structural components. The FSI used the upper layer temperature, absorption coefficient,

layer height, and the lower layer temperature obtained from FDS simulations as model input. All of these

parameters are controlled by the fire heat release rate. Other effects, such as the impact of the SFRM

thickness on the calculated temperature, are also considered below.

To determine the effect of variation of the fire heat release rate on the calculated thermal response of the

structural steel components, the resuUs of the FDS calculations presented in Chapter 10 for the case of a 10 percent higher heat release rate and reduced grid resolution for Test Iwere considered. The results of

the FDS calculations for upper layer temperature, absorption coefficient, layer interface height, and the

lower layer temperature are shown in Figs. 1 1-38 through 1 1—41 . These simulation results should be

compared to the results for the baseline case shown in Figs. 1 1-1 and 1 1-2. For the increased heat

release rate, the FDS results showed that the upper layer temperature increased (see Sec. 10.4.1 for a complete discussion), whereas the absolution coefficient, the layer height, and the lower layer

temperatures exhibited only small changes. For the coarse grid FDS calculation, all of the parameters

exhibited significant changes. For example, the upper layer temperature decreased approximately 100 °C,

and the absorption coefficient decreased approximately 10 percent. The coarse grid calculations tended to create upper layer temperature and layer height distributions that were nearly uniform. These two sets of

FDS results were used as input in the NIST FSI to predict the thermal behavior of the steel components.

212 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Time: 500.0 Figure 1 1-38. FDS simulation of the upper layer temperature on a plane 3.01 m above the

floor after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure)

and a crude mesh with 50 cm resolution (bottom figure).

NISTNCSTAR 1-5B, WTC Investigation 213
Chapter 1

1

Slice layer m

2.50

2.25

2.00 ^'

1.75 f

Figure 11-39. FDS simulation of the layer interface height after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure) and a crude mesh with 50 cm

resolution (bottom figure).

214 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

Slice, u-kap

Time 500.0

Figure 11-40. FDS simulation of the absorption coefficient on a plane 3.01 m above the

floor after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure)

and a crude mesh with 50 cm resolution (bottom figure).

NISTNCSTAR 1-5B, WTC Investigation 215
Chapter 1

1

Slice l-tmp C

Figure 11-41. FDS simulation of the lower layer temperature after 500 s in Test 1 using a heat release rate increased by 10 percent (top figure) and a crude mesh with

50 cm resolution (bottom figure).

216 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

11.4.1 Sensitivity to Heat Release Rate

Figures 1 1-42 and 1 1-43 show the predicted steel temperatures for bars A and B in Test 1, respectively,

corresponding to the FDS simulation with a heat release rate 10 percent higher than the baseline value. The computed transient steel temperatures for the higher heat release rate fire were similar in shape to those computed using the baseline heat release rate (shown in Figs. 1 1-9 and 1 1-10), but the rate of the

temperature increase was always larger. This led to simulated temperatures that were 5 percent to 21 percent higher than values in the baseline case, depending on location, which either improved or worsened agreement with the measurements.

Figures 1 1^44 and 1 1-45 show the predicted steel surface temperature at two different heights on the

column in Test 1 for a 10 percent higher heat release rate than that used in the baseline simulations shown

in Figs. 1 1-13 and 1 1-14. In this case, the elevated steel temperatures were paradoxically lower for the

increased gas-phase heat release rate. This occurred because the flame tended to bend away from the

column for the simulated fire with the higher heat release rate.

600

o

2 300

I" 200

100

-100

0 0 • J3 _ tJ-D a oa3 DD

vP°

oD

o

OD0

Num(Bfical Results

c

: s

1ft.

c 4ft.

D 6 ft.

, .

-

i9 ft.

1

1 1 1

, , 1 500 500 1000 1500

Time (s) 2000 2500 3000

Figure 1 1-42. Numerical predictions of the steel surface temperature on bar A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline.

NIST NCSTAR 1-5B, WTC Investigation 217
Chapter 1

1

600

500

400

g

2^ 300

"to k_

CD|-200

100

-100,500

,-, - 0 Ij 0 0

S D

0 n n

0 ttiS

Odo

JL

Numerical Results

a 1ft.

o 4ft.

0 6ft. 9 ft.

500 1000 1500

Time (s) 2000 2500 3000

Figure 11-43. Numerical predictions of the steel surface temperature on bar B in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline.

500

O 300 b

i:

50

-300

r

~1 -t- ^^P$^

i. ^ --^

J^ -

i

- s - J^l

: ^

*

Numerical Results

fstafth Face

s East Face

•J South Face

u V\fes1 Face

500 1000 1500 Time (s) 2030 2500 3030

Figure 11-44. Numerical predictions of the steel surface temperature for the column in Test 1 at 3.69 m above the floor for an FDS calculation performed with a heat release rate 10 percent above baseline.

218 NISTNCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

500

400

O 300 3 250 1

E^ 150

100 b

50

0

00

#fp^

Jo n

Numerical Results

North Face

; East Face

< South Face

= Vvfest Face

- . , , , 1 . . : . 1 : . 1 .

. 1 1 1 500 1000 15C Time (s)

2)(X) 25{X) 3000

Figure 11-45. Numerical predictions of steel surface temperature for the column in

Test 1 at 2.13 m above the floor for an FDS calculation performed with a heat release rate 10 percent above baseline.

Figures 1 1^6 through 1 1^8 show the simulated steel surface temperature on truss A in Test 1 for a 10 percent higher heat release rate than that used in the baseline calculations (which are shown in Figs. 1 1-17 through 1 1-19, respectively). The simulations showed that the temperature of the steel on

truss A increased 4 percent to 29 percent for the higher heat release rate, depending on the exact location.

O

Q.

(U -

c

: o n

- U U

- g a

- c^n ? 0

" on

- ZD

a'- ^ a Numerical Results

71.-1 TL-2

= TL-3

r. ll-A

: 1 1 1 1 . I , : . , Time (s)

Figure 11-46. Numerical predictions of steel surface temperature at four locations

2.89 m above the floor on truss A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline.

NISTNCSTAR 1-5B. WTC Investigation 219
Chapter 1

1

550

500

450 F400 - O 350 B

D 300

R250

E^ 200

150

100

50

500

' n o a o- CI a

0 8

ft Numerical R^ults

o TM-1

0 oa ..TM , a t t! "nvi-3

ai I - s

500 1(X10 1500 Time (s) 2000 2500 3000

Figure 11-47. Numerical predictions of steel surface temperature at four locations

3.29 m above the floor on truss A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline.

550

500

450

400

§350

^ 300

ro^250

E^ 200

150

100

50

%oo

^o--*^_, ^ i

^ ,^ -P ; i ! ° n \

i

j ; o

o D

D- D -

I '" g cP^ j .

I 3 D D-D 1 D 3 1 |_ 0~D 1 : ° a° 1

i

i D ; : J uZ-^-

o

r-a

Numerical Results

n TU-2 1

id

1 nJ-4 1

: S

"

1 1 1 1 1 1 1 1 1 1 1 1 1 1 ) 1 i 1 1 1 I 1 I 1 ] 1 1 1 1 1 1 1 1 500 1000 1500

Time (s) 2000 2500 3000

Figure 11-48. Numerical simulations of the steel surface temperature at four locations

3.70 m above the floor on truss A in Test 1 for an FDS calculation performed with a heat release rate 10 percent above baseline.

220 NISTNCSTAR 1-5B, WTC Investigation
Simulation of tlie Thermal Response of Structural Components

Table 1 1-3 summarizes the FSI simulation results shown in Figs. 1 1-42 through 1 1-^8. The table lists the percentage difference in the peak value of the simulated temperature profiles between the baseline

simulations and those with a 10 percent higher heat release rate (and consequent 9 percent increase in the

upper layer temperature) at the hottest and coolest locations on the steel surface of various uninsulated

components in Test 1 . The increased heat release rate and upper layer temperature led to changes in the

simulated steel surface temperature of 21 percent on average at the hottest locations on each of the

components, with a range from -38 percent to +29 percent. Table 1 1-3 shows that the percentage change

at the coolest locations was smaller on average than at the hottest locations.

Table 11-3. Percentage difference in the peak value of temperatures between the

baseline simulations and those with a 10 percent higher heat release rate at the hottest

and coolest locations on the steel surface of various uninsulated components In Test 1. Element (X'ertical Position) Hottest Location Coolest Location

Bar A 5 % 5 %

BarB 21 % 14%

Column (3.69 m above floor) -33 % -22 %

Column (2.13 m above floor) -38 % -26 %

Truss A (2.89 m above floor) 20% 5 %

Truss A (3.29 m above floor) 29 % 8 %

Truss A (3.70 m above floor) 4% 4%

Absolute Value of .Average 21 % 12 %

The increase in the maximum calculated steel temperatures seen in Figs. 1 1-42 through 1 1-^5 for locations on the bars, column, and truss A should be compared with the change in the upper layer gas

temperatures calculated by FDS, which increased by 9 percent on average for the 10 percent increase in the heat release rate above baseline (see Sec. 1 0.4. 1 ). The change in the simulated steel surface

temperature did not change linearly with the upper layer gas temperature. The difference in the

temperature increases on the steel can be attributed to subtle changes in the character of the simulated fire, including its physical location, its size, and its extent, all of which impacted the calculated flux onto the

structural steel components and, subsequently, the calculated temperature distribution in the steel components.

Chapter 1 0 showed that the FDS simulations were in agreement with upper layer gas temperature

measurements. The absolute difference between the measured and (FDS) simulated upper layer gas

temperatures for Tests 1 and 5 was 4 percent on average (see Figs. 10-3 through 10-12) for locations near

the east and west walls of the compartment. FDS tended to underpredict the gas temperature on the

intake side of the compartment (west) and overpredict the gas temperature on the exhaust side of the

compartment. The differences could be accounted for by a combination of measurement uncertainty and

sensitivity of FDS to various physical and numerical parameters, most importantly the input heat release

rate, which had an expanded (2a) uncertainty of 1 1 percent.

The difference between the measured and predicted steel surface temperatures (see Table 1 1-1) cannot be

explained solely by measurement uncertainty, which was estimated as approximately 5 °C (or 1 percent to 2 percent depending on the absolute value of the steel surface temperature). The difference, however, can

be explained through consideration of FSI model sensitivity to uncertainty in physical input parameters,

NISTNCSTAR 1-5B, WTC Investigation 221
Chapter 1

1

the most important of which was the heat release rate. A sensitivity study showed that a 10 percent

increase in the heat release rate led to a 21 percent higher steel surface temperature on average at the

hottest locations, as compared to the baseline results for Test 1 . The FSI simulation exhibited a non-linear

sensitivity of the steel temperature to the fire heat release rate and associated upper layer temperatures.

The simulated thermal behavior of the steel was also sensitive to the proximity of the structural components to the fire. Chapter 10 describes how FDS did a better job predicting the heat flux to locations far from the fire, because the size, spatial extent, and precise location of the fire had less influence on the results.

11.4.2 Sensitivity to Grid Size

Figures 1 1-49 and 1 1-50 show the predicted steel temperature for bars A and B in Test 1 using the FDS

coarse grid simulations. These results are of interest because the WTC fire calculations (NIST

NCSTAR 1-5F) must be completed using a coarse grid due to computational costs. The computed steel temperatures using the coarse mesh were similar in shape to those computed using the finer mesh shown

in Figs. 1 1-9 and 1 1-10, but the maximum calculated temperatures were shifted by approximately

10 percent to either larger or smaller values. The coarse mesh calculations displayed uniform

temperatures along the length of the bar, which was in contrast to the fine mesh calculations, which led to

finite temperature differences (15 percent) along the length of the bar. The calculations show that FDS

grid resolution has a finite effect on the thermal-structural analysis, although it is not greater than the

effect of model sensitivity to various model input parameters.

500

400

2 300

03P-200

/ ,

. J ] §

i

aa s

a

I i!

^ 1 ft. o 4 ft. o 6 ft. o 9 ft.

1 [ 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

, , , , -500 0 500 1000 1500 2000 2500 3000

Time (s)

Figure 11-49. Predicted numerical results for steel temperature for bar A in Test 1. These

results correspond to an FDS calculation performed with a crude mesh

(50 cm resolution).

222 NIST NCSTAR 1-5B, WTC Investigation
Simulation of the Thermal Response of Structural Components

B %

! _ Numerical Results

c 1 ft.

: ; E 4ft.

! a 6ft.

I i 0 9ft.

J I I l_l I I I L-l I I I I I \ I I I I I l_ '"-'Boo 0 500 1000 1500 2000 2500 3000 Time (s)

Figure 11-50. Predicted numerical results for steel temperature on bar B in Test 1. These results correspond to an FDS calculation performed with a crude mesh

(50 cm resolution).

1 1.5 SENSITIVITY TO THE THICKNESS AND THERMO-PHYSICAL

PROPERTIES OF THE SPRAYED FIRE-RESISTIVE MATERIAL

There were additional factors that may have influenced the thermal response of the steel when the SFRM

was present, including variability in the thickness of the SFRM and the thermo-physical properties of the

steel and the SFRM. The optical properties of the steel and the SFRM, however, probably did not play an

important role, as soot rapidly coated the steel components upon ignition of the fire, as is typical of fires burning materials that produce copious amounts of soot.

The SFRM thickness was not uniform about a structural element, even a simple symmetric element such

as a bar. Table 9-2 shows that on the bars, for example, the SFRM thickness was not uniform about all sides of the bar—even at a single location along the axis of the bar. The nonunifonn thickness led to three-dimensional effects that impacted the thermal behavior of the SFRM, effectively increasing the

uncertainty in the simulated steel temperature underneath the SFRM. The importance of this effect can be

characterized by the coefficient of variation (COV) of the SFRM thickness (the ratio of the standard

deviation to the mean thickness; see Sec. 9.2) and is bounded by the maximum and minimum values of

the SFRM thickness. The coefficient of variation ranged from 0.17 to 0.27 for the bars, columns, and

trusses.

The SFRM thickness in the calculations presented here was based on this mean value, which was

assumed to be uniform over the entire length of the bar. In Test 5 for example, the mean thickness of the

SFRM on the bars was measured as 2.3 1 cm and the COV was 0.24 (see Table 9-2). Figure 1 1-5 1 shows

the numerical simulation of the steel surface temperature for bar A in Test 5, assuming a 25 percent

500

400

O

2 300

|-200

100

•in

NISTNCSTAR 1-5B, WTC Investigation 223
Chapter 1

1

decrease in the thickness of the SFRM. The resuhing temperatures increased by approximately 10

percent as compared to the baseline results shown in Fig. 1 1-25. Table 9-2 also shows that locally, the

SFRM thickness on some sections of the bars was smaller than the mean by as much as 45 percent. This

indicates that the variation in the calculated steel surface temperature was as much as 20 percent simply

from uncertainty in the SFRM thickness. NIST NCSTAR 1-5G and NIST NCSTAR 1-6A present a comprehensive sensitivity study of the role of the variation of the SFRM thickness on its effectiveness.

600

^ 400

Si

o3Q.o 200

0

0 1000 2000 3000 4000 5000

Time (s)

Figure 1 1-51 . Numerical predictions of steel surface temperature for bar A in

Test 5 assuming a SFRM thickness 25 percent below baseline.

The effect of uncertainty in the thennal conductivity of the SFRM will also play a role. Typical

uncertainty in the detemiination of themial conductivity using the ASTM standard method is not better than ±6 percent (ASTM 2000). For insulation materials like BLAZE-SHIELD DC/F SFRM, the

measurement uncertainty may be considerably larger, although it has not been quantified. According to the manufacturer, the thermal conductivity for BLAZE-SHIELD DC/F was 0.042 W/m-K at 24 °C, which was quite different from Hannathy's temperature-dependent data shown in Fig. 1 1-4 that was used in the

baseline calculations reported here. A third set of thermophysical data was obtained for the SFRM (NIST

NCSTAR 1-6A). In that data set, the themial conductivity was determined to be about 10 percent higher

than the value reported by the manufacturer. The sensitivity of the calculation results to the

thenuophysical properties was investigated by considering the idealized situation where a 2.5 cm bar with

2.5 cm of insulation was exposed to a constant source of heat flux. The results showed a 7 percent lower

steel temperature for the calculations that used the Harmathy data as compared to the calculations that used the data from NIST NCSTAR 1-6A. The sensitivity of FSI to the thermophysical properties of the

SFRM increased uncertainty in the simulation (NIST NCSTAR 1-5G and NIST NCSTAR 1-6A). The

results suggest that the sensitivity of FSI to the thermophysical properties was significant, but much less so than the uncertainty associated with the SFRM thickness.

11.6 SUMMARY

The behavior of the simulated and measured uninsulated steel temperatures was similar. Both increased

in a monotonic fashion from ambient values at time equal to zero, to finite values that obtained a maximum approximately when the fuel flow and the fire stopped. The same was true for the insulated

224 NIST NCSTAR 1-5B,WTC Investigation
Simulation of tiie Thermal Response of Structural Components

steel components, except that the temperatures obtained a maximum after the fiiel was stopped due to lag

of the thermal wave through the SFRM. In both cases, the temperatures decreased after the fire was

stopped, also in a monotonic manner.

On average, the numerical predictions of the steel surface temperature were within 7 percent of the

experimental measurements for the case of bare steel and within 17 percent for the steel components with

SFRM with the maximum difference reading 38 percent. For the bare steel, a series of calculations

showed that the differences could be attributed to model sensitivity to the heat release rate in the fire model. For the insulated steel, the differences could be attributed to model sensitivity to the SFRM

coating thickness.

11.7 REFERENCES

ASTM International. 2000. Standard Test Methodfor Thermal Conductivity^ ofRefractories by Hot Wire

(Platinum Resistance Theimometer Technique), vol. 15.01. ASTM C 1 1 13-99, Annual Book of

ASTM Standards. West Conshohocken, PA.

Eurocode 4. 1994. Design of Composite Steel and Concrete Structures, ENV 1993-1-2: General, EC4,

Rules - Structural Fire Design. European Committee for Standardization, Brussels, Belgium.

Harmathy, T.Z. 1983. "Properties of Building Materials at Elevated Temperatures," paper NRCC-20956

(DBR-P-1080), Division of Building Research, National Research Council, Canada, (1080), pp. 72,

March.

Prasad, K., and H.R. Baum. 2005. "Coupled Fire Dynamics and Thermal Response of Complex Building

Structures," Proc. Combustion Institute 30, 2255-2262.

Taylor, R.E., H. Groot, and J. Terrier. 2003. Thermophysical Properties ofPVC, PE, and Marinite,

Report No. TPRL 2958, Thermophysical Properties Research Laboratory, Inc., April.

NISTNCSTAR 1-5B, WTC Investigation 225
Chapter 1

1

s

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226 NISTNCSTAR 1-5B, WTC Investigation
Chapter 12

Summary of Results

A series of experiments was conducted to assess the models that were developed to determine the

temperature rise of structural components within a compartment fire. The well-controlled large-scale

experiments described in this report were designed specifically to validate aspects of the National Institute of Standards and Technology (NIST) Fire Dynamic Simulator (FDS), the NIST Fire Structure Interface

(FSI), and the commercially available finite-element program ANSYS, which were used to simulate the

fires and their impact on the thermal behavior of structural components as applied to the investigation of

the World Trade Center (WTC) fire and collapse. Measurements were conducted to assess the accuracy

of each of the model calculations. The fire and thermal models exercised here were used to generate

input for a set of structural analyses to determine the collapse sequence of the WTC towers. The

experiments reported here provide a basis for understanding the uncertainties associated with these

models, and thereby, constraints on the limits of accuracy of the structural calculations.

Each experiment was heavily instrumented so that many aspects of the fire model were tested. Many

assumptions went into the design of the experiments, including details associated with the test matrix, the

experimental configuration, and the test conditions. It was impossible to recreate every feature of the

actual fire conditions in the WTC buildings, and that was not the primary goal of the work presented here. In this work, a number of important aspects of the WTC fires were recreated to compare model

predictions with measurements. The measurements led to findings regarding the thennal behavior of the

compartment and its contents, including the steel components, as they were exposed to fires of various

size, fuel sootiness, and proximity to the fire plume. Noteworthy resuhs included the following:

• The temperature and flux measurements showed that, in general, the experiments were very

repeatable.

• Most measured quantities (temperatures, gas concentrations, soot, fluxes, velocities, etc.) showed rapid changes during the first 100 s in the experiments and then were fairly steady or slowly varied, as the temperature of the compartment walls slowly increased.

• The value of the heat flux about the compartment was dependent on location and test conditions. For example, at the floor in front of the fire pan, the heat flux varied from

20 kW/m- to 90 kW/m'. In the upper layer near the ceiling, the heat flux was measured to be

as high as 90 kW/m".

• Higher temperatures (200 °C to 300 °C higher) were measured in the upper layer of fires with

larger heat release rates (3 MW versus 2 MW). In addition, fire size had a large influence on

the heat flux impinging on structural components in both the hot upper layer and elsewhere in

the compartment.

• Ceiling surface temperatures obtained values as high as 700 °C to 800 °C and 800 °C to 900 °C after 15 min and 50 min, respectively, of exposure to a 3 MW heat release rate fire with peak temperatures observed to occur above the fuel pan.

NISTNCSTAR 1-5B, WTC Investigation 227
Chapter 12

• The measured yields of soot and CO2 in the hot upper layer of the compartment were very

similar to those measured in the downstream plume of spray fires burning in the open. The

results for CO were more complex, with the yield varying with location and changing with

time, depending on the experimental conditions.

• Gas velocity measurements at the vents showed that the velocity was largest near the top of

the vent and decreased with height.

• The experimental results showed that the SFRM significantly delayed the rise in the steel temperature. The highest temperatures obtained at the steel surface were significantly lower

than the temperature on the outside face of the insulation material.

• The surface temperature of uninsulated steel trusses that were subject to impingement or proximity of flames from fires with a nominal 3 MW heat release rate reached 600 °C within

15 min. In comparison, the insulated steel trusses with SFRM protection (1.91 cm nominal

thickness) reached 100 °C to 200 °C and the truss with 3.81 cm SFRM reached 50 °C to 100 °C after 1 5 min of exposure to a 3 MW heat release rate fire. The steel surface of an

insulated steel truss with 1.91 cm (nominal) and 3.81 cm (nominal) SFRM obtained a temperature of 680 °C and 415 °C, respectively, after 50 min of exposure to a 3 MW fire.

• For the insulated steel components, the steel surface temperatures obtained a maximum

sometime after the fire was stopped due to lag of the thermal wave through the SFRM. A

thennal lag of between 100 s and 800 s for the nominal 1.91 cm SFRM thickness and 1,000 s to over 1,900 s for the nominal 3.81 cm SFRM thickness was observed, depending on the

values of the temperature and the temperature gradients within the SFRM and the steel.

Before testing began, model calculations were performed to guide the design of the compartment and to provide a baseline set of "blind'" predictions. The accuracy of the models was assessed in tenns of

sensitivity to a number of parameters. The sensitivity analysis yielded estimates of uncertainty for model

output that facilitated quantitative evaluation of the agreement between models and measurements. Using

a well-known correlation between compartment temperature and the heat release rate of the fire, it was

possible to propagate the uncertainty in the measurement of the heat release rate to the measurement of

the upper layer temperature, and then to the measurement of the heat flux to the walls and the structural components. These uncertainties provided bounds for the numerical simulations. A comparison of the

predictions with the measurements showed that:

• The FDS predictions fell within the uncertainty bounds of the gas temperature and heat flux

measurements. FDS predictions of the upper layer gas temperature agreed with

measurements to within 1 0 percent, which was fairly accurate considering model sensitivity

to input parameters such as the heat release rate. Since the heat flux to the walls and objects

within the upper layer is highly dependent on the upper layer temperature, these predictions were also accurate. The model accuracy decreased near the fire. Greater uncertainty in nearfield predictions was due to a combination of limited spatial resolution of the numerical grid,

simplification of the combustion and radiation processes, and differences between model and

actual burner geometries.

228 NISTNCSTAR 1-5B, WTC Investigation
Summary of Results

• The sootier-burning fuel led to similar temperature rise in the ceiling and the steel above the

fire plume, while the measured and predicted fluxes to the lower layer was smaller.

• FDS predicted asymmetry of the fire plume caused by obstructions to uniform flow through

the compartment, but did not accurately predict the extent of the asymmetry, which adversely

impacted the FSI prediction of structural components very close to the fire. Away from the

fire, small differences in the exact flame position were not as important.

• The behaviors of the simulated and the measured uninsulated steel temperatures were similar. Both increased from ambient values to finite values in a monotonic fashion that obtained a maximum approximately when the fiiel flow was stopped. The same was true for the

insulated steel components, except that the temperatures obtained a maximum after the fuel was stopped due to lag of the thermal wave through the SFRM.

• On average, the numerical predictions of the steel surface temperature were within 8 percent

of the experimental measurements for the case of bare steel and within 20 percent for the

steel components with SFRM. For the bare steel, a series of calculations showed that these

differences could be attributed to model sensitivity to the fire heat release rate used in the fire model. For the insulated steel, the differences could be attributed mainly to model sensitivity

to the SFRM coating thickness.

In addition to checking the accuracy of the model, the experiments served as a means to check the

sensitivity of the model to changes in input parameters, both numerical and physical. One of the most

important parameters prescribed by the model user is the size of the numerical grid. Simulations of all the

tests were performed with grid cells 10 cm by 10 cm by 10 cm. Several tests were also simulated using a coarse grid 50 cm by 50 cm by 40 cm, the same as was used to perform the simulations of the WTC fires. The upper layer temperature predictions did not differ appreciably from those of the fine grid simulation,

justifying the use of the coarse grid in the WTC simulations. Changes in various other physical

parameters also did not significantly change the results. This is important because the initial and

boundary conditions in the WTC simulations are far less certain than those of the experiments presented

here.

In summary, measurements in a series of specially designed, large-scale fire experiments were compared

with predictions from the FDS/FSI/ANSYS modeling methodology to quantify uncertainties and the

parameters to which the models are most sensitive. The sensitivity analysis for FDS was extended to the

steel temperature calculations, providing the bounds for quantifying the accuracy of the FSI/ANSYS

model. Quantitative disagreement between the models and the experiments were attributed to model

sensitivities to input parameters that had experimental uncertainties. The study provides the level of

confidence necessary to apply the modeling methodology to the investigation of the WTC disaster.

NISTNCSTAR 1-5B, WTC Investigation 229
Chapter 12

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230 NISTNCSTAR 1-5B, WTC Investigation
Appendix A

Selection of Fuels

A.1 COMPOSITION OF OFFICE FURNISHINGS

It was assumed that there were two generic types of furniture present in the World Trade Center (WTC)

offices, namely wood based or modem. The first type was predominantly composed of cellulosic

materials, such as books, conglomerate bookshelves, wooden desks, and lots of paper. Some amount of

thermoplastics would also be present. The second type of furniture considered was composed of materials

that were less cellulosic and more thermoplastic in composition. Table A-1 shows the assumed

composition of the two office types. Nylon is essentially ignored because that carpeting would play only

a minor role in the total heat release of an office.

Table A-1 . Assumed composition of office furnishings.

Office Type Cellulose ABS PolystA rene Wood-based 2/3 1/6 1/6 Modem 1/3 1/3 1/3 Key: ABS. copolymer of acrylonitrile-butadiene-styrene

The smoke yield is defined as the number of grams of smoke emitted per gram of fuel combusted. The

smoke yield from fires burning cellulose (representative of woody materials), copolymer of acrylonitrilebutadiene-styrene (ABS). and polyst>Tene are 0.015, 0.105, and 0.164, respectively. The smoke yield

from fires burning heptane and toluene are 0.037 and 0.178, respectively (Tewarson 2002). Jet A (fuel

from the aircraft) is composed essentially of kerosene, which has a smoke yield of 0.215.

Assuming a linear relationship between composition and smoke yield, a wood-based office would be

expected to have a smoke yield of:

2/3 (0.015) + 1/6 (0.105) +1/6 (0.164) = 0.056;

whereas, a modem type office would be expected to have a smoke yield of:

1/3 (0.015) + 1/3 (0.105) +1/3 (0.164) = 0.095.

This can be compared to a mixture of 50 percent heptane and 50 percent toluene (by mass) (assuming a

linear relation); the yields would be expected to have a comparable smoke yield:

'/2 (0.037) + '/2 (0.178) = 0.11,

which is somewhat larger than the smoke yield of the modem office type. A mixture of 60 percent

heptane and 40 percent toluene (by mass), assuming a linear relation would be expected to have a smoke

yield of:

0.6 (0.037) + 0.4 (0.178) = 0.0934,

NISTNCSTAR 7-56, WTC Investigation 231
Appendix A

which is similar to the smoke yield of the modern office type. To cover the range of soot yields during

the validation experiments, the fuels used were heptane and a mixture of 60 percent heptane and

40 percent toluene (by mass). Pure toluene would yield a fire too heavily laden with soot as compared to materials that are typically present in office buildings such as the WTC towers. Although jet fuel may

have been present during the WTC fires, the majority of the mass combusted during the WTC fires was

due to office furnishings. Heptane and toluene are miscible with highly similar boiling points (Hamins

1985). Preferential burning is not significant as the fuels were injected in the form of a spray.

A.2 REFERENCES

Hamins, A. 1985. The Structure and Extinction of Diffusion Flames, Ph.D. Dissertation, University of

California at San Diego.

Tewarson, A. 2002. "Generation of Heat and Chemical Compounds in Fires," pp. 3-82 to 3-161 in the

SFPE Handbook ofFire Protection Engineering, 3rd ed. (P.J. DiNenno, ed.), Society of Fire

Protection Engineers, Bethesda, MD.

232 NISTNCSTAR 1-5B. WTC Investigation
Appendix B

Instrument Locations

Table B-1 is a detailed list of instrumentation location and channel assignment in the data acquisition

system. The location of the instrument is given when a measurement was made within the experimental

compartment. Many measurements were made outside of the compartment (e.g., calorimetry) or were

used to monitor instrument function and do not have coordinate locations listed in the table. The

reference location (0.0.0) was the floor at the center of the west wall, as described in Chapter 2. The

location of the thermocouples on the steel components changed from test to test and is described in detail

in Chapter 9.

Table B-1. Instrument locations and channe

Category Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

Calorimetry 0 02 Exh Duct

1 02 Press Exh Duct

2 C02 Exh Duct

i LU2 Range hxn Duct

4 CO Exh Duct

5 CO Range Exh Duct

6 Pressure Exh Duct

7 Pressure Exh Duct

8 Pressure Exh Duct

9 Pressure Exh Duct

10 Pressure Exh Duct

11 Pressure Exh Duct

15 Small Test Bay

16 Med Test Bay

17 Large Test Bay

18 Gas Pressure Gas Line

19 Gas Press Supply

Voltage

Gas Line

22 Ambient TC Rack

24

TC for Nat

Gas

Gas Line

25 TC for Press Exh Duct

assignments

NISTNCSTAR 1-5B. WTC Investigation 233
Appendix B

Category Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

26 TC for Press Exh Duct

27 TC for Press Exh Duel

28 TC for Press Exh Duct

29 TC for Press Exh Duct

30 TC for Press Exh Duct

(jas

Analyzers

02 33 02 604.0 28.0 365.0

C02 34 C02 raw 604.0 28.0 365.0

35 C02 range 604.0 28.0 365.0

CO 36 CO raw 604.0 28.0 365.0

37 CO range 604.0 28.0 365.0

C02 38 C02 raw 138.0 54.0 324.0

39 C02 range 138.0 54.0 324.0

CO 40 CO raw 153.0 54.0 324.0

41 CO range 153.0 54.0 324.0

Pressure Inlet 42 Press Inlet -17.8 -33.0 71.1

43 Press Inlet -17.8 -33.0 45.7

44 Press Inlet -17.8 -33.0 25.4

45 Press Inlet -17.8 -59.7 71.1 46 Press Inlet -17.8 -59.7 25.4

47 Press Inlet -17.8 -127.0 71.1

48 Press Inlet -17.8 -127.0 45.7

49 Press Inlet -17.8 -127.0 25.4

Outlet 50 Press Outlet 721.4 33.0 261.6

51 Press Outlet 721.4 33.0 236.2

52 Press Outlet 721.4 33.0 215.9

53 Press Outlet 721.4 59.7 261.6

54 Press Outlet 721.4 59.7 215.9

55 Press Outlet 721.4 125.7 261.6

56 Press Outlet 721.4 125.7 236.2

57 Press Outlet 721.4 125.7 215.9

Near Metal 58 Press Heat Flux 435.0 17.0 320.0

59 Press Heat Flux 481.0 17.0 320.0

234 NISTNCSTAR 1-5B, WTC Investigation
Instrument Locations

Categon' Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

Ceiling Heat Flux Gauges 60 Ceiling 349.0 28.0 382.0

61 Ceiling 399.0 28.0 382.0

t 62 Ceiling 474.0 28.0 382.0

63 Ceiling 650.0 28.0 382.0

ThermocouplesNear Metal TC (asp) for Velocity

64 Asp TC Heat Flux 161.0 53.0 322.0

Weill Xt r^oiMrtn Tt^c 65 Bare TC

Ceiling

between

layers

456.0 0.0 383

66 Bare TC Ceiling 456.0 0.0 382

67 Bare TC Ceiling 456.0 30.0 382

68 Bare TC Ceiling 456.0 45.0 382

69 Bare TC

Ceiling

between

layers

456.0 60.0 383

70 Bare TC Ceiling 456.0 60.0 382

71 Bare TC Ceiling 477.0 0.0 382

72 Bare TC Ceiling 492.0 0.0 382

73 Bare TC Ceiling 507.0 0.0 382

74 Bare TC

Ceiling

between

layers

522.0 0.0 383

75 Bare TC Ceiling 522.0 0.0 382

76 Bare TC Ceiling 552.0 0.0 382

77 Bare TC Ceiling 582.0 0.0 382

78 Bare TC Ceiling 612.0 60.0 382

79 Bare TC

Ceiling

between

layers

612.0 0.0 383

80 Bare TC Ceiling 612.0 0.0 382

81 Bare TC Ceiling 612.0 -60.0 382

82 Bare TC Ceiling 456.0 -15.0 382

83 Bare TC Ceiling 456.0 -30.0 382

84 Bare TC

Ceiling

between

layers

456.0 -60.0 383

85 Bare TC Ceiling 456.0 -60.0 382

86 Bare TC Ceiling 444.0 0.0 382

87 Bare TC Ceiling 429.0 0.0 382

NISTNCSTAR 1-5B. WTC Investigation 235
Appendix B

Category Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

88 Bare TC Ceiling 414.0 0.0 382

89 Bare TC

Ceiling

between

layers

399.0 0.0 383

90 Bare TC Ceiling 399.0 0.0 382

91 Bare TC Ceiling 369.0 0.0 382

92 Bare TC Ceiling 309.0 0.0 382

93 Bare TC Ceiling 279.0 0.0 382

94 Bare TC Ceiling 228.0 -60.0 382.0

95 Bare TC Ceiling 228.0 0.0 382.0

96 Bare TC Ceiling 228.0 60.0 382.0

North Wall TCs 97 Bare TC

Ceiling

between

layers

459.0 181.0 377

98 Bare TC 459.0 180.0 377

99 Bare TC 459.0 181.0 367.0

100 Bare TC 459.0 180.0 367.0

101 Bare TC 459.0 181.0 357.0

102 Bare TC 459.0 180.0 357.0

103 Bare TC 519.0 180.0 367.0

104 Bare TC 579.0 181.0 367.0

105 Bare TC 579.0 180.0 367.0

106 Bare TC 579.0 180.0 357.0

107 Bare TC 399.0 180.0 367.0

108 Bare TC 339.0 181.0 367.0

109 Bare TC 339.0 180.0 367.0

110 Bare TC 339.0 180.0 357.0

Truss 1 Uninsulated TCs 113 Bare TC Steel Surf * *

114 Bare TC Steel Surf * * *

115 Bare TC Steel Surf *

116 Bare TC Steel Surf * * *

117 Bare TC Steel Surf * * *

118 Bare TC Steel Surf * *

119 Bare TC Steel Surf * *

120 Bare TC Steel Surf * * *

121 Bare TC Steel Surf * * *

122 Bare TC Steel Surf * * *

236 NISTNCSTAR 1-5B, WTC Investigation
Instrument Locations

Category Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

123 Bare TC Steel Surf * *

124 Bare TC Steel Surf * *

Truss 1 Insulated TCs 125 Bare TC Insul Surf * * *

126 Bare TC Insul Surf * *

127 Bare TC Insul Surf * *

128 Bare TC Insul Surf * *

129 Bare TC Insul Surf * * *

130 Bare TC Insul Surf *

1 131 Bare TC Insul Surf * *

132 Bare TC Insul Surf * * *

133 Bare TC Insul Surf *

134 Bare TC Insul Surf * *

135 Bare TC Insul Surf * *

136 Bare TC Insul Surf * *

Column Uninsulated TCs 137 Bare TC Steel Surf * * *

138 Bare TC Steel Surf

139 Bare TC Steel Surf * *

140 Bare TC Steel Surf * * *

141 Bare TC Steel Surf * * *

142 Bare TC Steel Surf *

143 Bare TC Steel Surf *

144 Bare TC Steel Surf * *

145 Bare TC Steel Surf * * *

146 Bare TC Steel Surf * *

147 Bare TC Steel Surf * *

148 Bare TC Steel Surf * *

Column Insulated TCs 149 Bare TC Insul Surf * * *

150 Bare TC Insul Surf * * *

151 Bare TC Insul Surf * *

152 Bare TC Insul Surf * *

153 Bare TC Insul Surf * *

154 Bare TC Insul Surf *

155 Bare TC Insul Surf * * *

156 Bare TC Insul Surf * * *

157 Bare TC Insul Surf * *

NISTNCSTAR 1-5B. WTC Investigation 237
Appendix B

Category Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cin) Y(cm) Z (cm)

158 Bare TC Insul Surf * *

159 Bare TC Insul Surf * * *

160 Bare TC Insul Surf * * *

Rod Uninsulated TCs 161 Bare TC Steel Surf * *

162 Bare TC Steel Surf * *

163 Bare TC Steel Surf * * *

164 Bare TC Steel Surf * * *

Rod Insulated TCs 165 Bare TC Insul Surf * * *

166 Bare TC Insul Surf * * *

167 Bare TC Insul Surf * * *

168 Bare TC Insul Surf * * *

Truss 2 Uninsulated TCs 169 Bare TC Steel Surf * * *

170 Bare TC Steel Surf * * *

171 Bare TC Steel Surf * * *

172 Bare TC Steel Surf * * *

173 Bare TC Steel Surf * * *

174 Bare TC Steel Surf * * *

175 Bare TC Steel Surf * *

176 Bare TC Steel Surf * *

177 Bare TC Steel Surf * * *

178 Bare TC Steel Surf * * *

179 Bare TC Steel Surf * * *

180 Bare TC Steel Surf * * *

Truss 2 Insulated TCs 181 Bare TC Insul Surf * * *

182 Bare TC Insul Surf * * *

183 Bare TC Insul Surf * * *

184 Bare TC Insul Surf * * *

185 Bare TC Insul Surf * * *

186 Bare TC Insul Surf * * *

187 Bare TC Insul Surf * *

188 Bare TC Insul Surf * * *

189 Bare TC Insul Surf * *

190 Bare TC Insul Surf * * *

191 Bare TC Insul Surf * * *

192 Bare TC Insul Surf * * *

238 NISTNCSTAR 1-5B, WTC Investigation
Instrument Locations

CategorA Channel Groups

Chan.

No. Instr. Desc. Loc. Desc. X(cm) Y (cm) Z(cm)

Inlet TCs(12bare&2asp) 193 Bare TC Inlet -25.4 -31.8 73.7

194 Bare TC Inlet -25.4 -31.8 63.5

195 Bare TC Inlet -25.4 -31.8 49.5

196 Bare TC Inlet -25.4 -31.8 30.5

197 Bare TC Inlet -25.4 -57.2 73.7

198 Bare TC Inlet -25.4 -57.2 63.5 199 Bare TC Inlet -25.4 -57.2 49.5 200 Bare TC Inlet -25.4 -57.2 30.5 201 Bare TC Inlet -25.4 -125.7 73.7 202 Bare TC Inlet -25.4 -125.7 63.5 203 Bare TC Inlet -25.4 -125.7 49.5 204 Bare TC Inlet -25.4 -125.7 30.5 205 Asp TC Inlet -17.8 -50.8 45.7 206 Asp TC Inlet -17.8 -1 19.4 30.5

Outlet TCs (30 bare & 4

asp)

207 Bare TC Outlet 711.2 -127.0 265.4

208 Bare TC Outlet 711.2 -127.0 256.5 209 Bare TC Outlet 71 1.2 -127.0 242.6

210 Bare TC Outlet 711.2 -127.0 223.5

211 Bare TC Outlet 711.2 -127.0 210.8

212 Bare TC Outlet 711.2 -57.2 265.4

213 Bare TC Outlet 711.2 -57.2 256.5 214 Bare TC Outlet 711.2 -57.2 242.6

215 Bare TC Outlet 711.2 -57.2 223.5

216 Bare TC Outlet 7II.2 -57.2 210.8

217 Bare TC Outlet 711.2 -35.6 265.4

218 Bare TC Outlet 711.2 -35.6 256.5

219 Bare TC Outlet 711.2 -35.6 242.6

220 Bare TC Outlet 711.2 -35.6 223.5

221 Bare TC Outlet 711.2 -35.6 210.8

222 Bare TC Outlet 711.2 35.6 265.4

223 Bare TC Outlet 711.2 35.6 256.5 224 Bare TC Outlet 711.2 35.6 242.6

225 Bare TC Outlet 711.2 35.6 223.5 226 Bare TC Outlet 711.2 35.6 210.8

227 Bare TC Outlet 711.2 57.2 265.4

NISTNCSTAR 1-5B, WTC Investigation 239
Appendix B

Category Channel Croups

Chan.

No. Instr. Desc. Loc. Desc. X (cm) Y (cm) Z(cm)

228 Bare TC Outlet 711.2 57.2 256.5

229 Bare TC Outlet 711.2 57.2 242.6

230 Bare TC Outlet 711.2 57.2 223.5

231 Bare TC Outlet 711.2 57.2 210.8

232 Bare TC Outlet 711.2 127.0 265.4

233 Bare TC Outlet 711.2 127.0 256.5

234 Bare TC Outlet 711.2 127.0 242.6

235 Bare TC Outlet 711.2 127.0 223.5

236 Bare TC Outlet 711.2 127.0 210.8

237 Asp TC Outlet 721.4 50.8 254.0

238 Asp TC Outlet 721.4 50.8 210.8

West Tree TCs (15 bare & 2 asp)

239 Bare TC West Tree 95.0 51.0 377.0

240 Bare TC West Tree 95.0 51.0 365.0

241 Bare TC West Tree 95.0 51.0 350.0

242 Bare TC West Tree 95.0 51.0 335.0

243 Bare TC West Tree 95.0 51.0 320.0

244 Bare TC West Tree 95.0 51.0 305.0

245 Bare TC West Tree 95.0 51.0 289.0

246 Bare TC West Tree 95.0 51.0 274.0

247 Bare TC West Tree 95.0 51.0 244.0

248 Bare TC West Tree 95.0 51.0 216.0

249 Bare TC West Tree 95.0 51.0 183.0 250 Bare TC West Tree 95.0 51.0 153.0

251 Bare TC West Tree 95.0 51.0 124.0 252 Bare TC West Tree 95.0 51.0 91.0

253 Bare TC West Tree 95.0 51.0 33.0 254 Asp TC West Tree 100.0 53.0 364.5

255 Asp TC West Tree 100.0 53.0 214.5

256 Asp TC West Tree 100.0 53.0 33.5

East Tree TCs ( 1 5 bare & 2 asp)

257 Bare TC East Tree 604.0 52.0 379.0

258 Bare TC East Tree 604.0 52.0 364.0

259 Bare TC East Tree 604.0 52.0 349.0

260 Bare TC East Tree 604.0 52.0 334.0

261 Bare TC East Tree 604.0 52.0 319.0

262 Bare TC East Tree 604.0 52.0 304.0

240 NISTNCSTAR 1-5B, WTC Investigation
Instrument Locations

Categon Channel Groups

Chan.

No. Instr. Desc. X (cm) Y (cm) Z (cm)

263 Bare TC East Tree 604.0 52.0 289.0

264 Bare TC East Tree 604.0 52.0 274.0

265 Bare TC East Tree 604.0 52.0 244.0

266 Bare TC East Tree 604.0 52.0 214.0

267 Bare TC East Tree 604.0 52.0 183.0 268 Bare TC East Tree 604.0 52.0 152.0 269 Bare TC East Tree 604.0 52.0 122.0 270 Bare TC Fast Tree 604.0 52.0 92.0

271 Bare TC East Tree 604.0 52.0 62.0 272 Bare TC East Tree 604.0 52.0 34.0 273 Bare TC East Tree 604.0 52.0 3.0 274 Asp TC East Tree 604.0 53.0 365.0

275 Asp TC East Tree 604.0 53.0 215.0

276 Asp TC East Tree 604.0 53.0 35.0

r^pilincT Hpat FliiY Tr^<; 277 Bare TC Ceiling 349.0 28.0 382.0

278 Bare TC Ceiling 399.0 28.0 382.0

279 Bare TC Ceiling 474.0 28.0 382.0

280 Bare TC Ceiling 650.0 28.0 382.0

Flux Station Hpat Fliiv TCs

285 Bare TC Flux St 2 458.0 56.0 330.0

286 Bare TC Flux St 2 458.0 44.0 330.0

287 Bare TC Flux St 2 458.0 23.0 315.0

288 Bare TC Flux St 2 458.0 35.0 315.0

Column Heat Flux TCs 289 Bare TC Column 304.0 77.0 346.0

290 Bare TC Column 297.0 85.0 327.0

291 Bare TC Column 304.0 77.0 92.0

292 Bare TC Column 297.0 85.0 102.0

Floor Heat Flux TCs 293 Bare TC Floor 500.0 6.0 13.0 294 Bare TC Floor 550.0 3.0 13.0

Slug 1 TCs 299 TC 102.0 53.0 329.0

300 TC 102.0 53.0 329.0

301 TC 102.0 53.0 329.0

302 TC 102.0 53.0 329.0

Slug 2 TCs 303 TC 212.0 53.0 291.0

304 TC 212.0 53.0 291.0

305 TC 212.0 53.0 291.0

306 TC 212.0 53.0 291.0

NISTNCSTAR 1-5B, WTC Investigation 241
Appendix B

Category Channel Groups

Cnan.

No. Instr. Desc. Loc. Desc. X (cm) Y (cm) Z (cm)

Slug 3 TCs 307 TC 337.0 53.0 329.0

308 TC 337.0 53.0 329.0

309 TC 337.0 53.0 329.0

310 TC 337.0 53.0 329.0

Slug 4 TCs 311 TC 397.0 53.0 329.0

312 TC 397.0 53.0 329.0

313 TC 397.0 53.0 329.0

314 TC 397.0 53.0 329.0

Slug 5 TCs 315 TC 550.0 53.0 329.0

316 TC 550.0 53.0 329.0

5 1 / 1 L

C C A A

->!>U.U C5 A53.0 329.0

jto 1

L

C C A A 550.0 CIA 53.0 O TA A 329.0

Near Metal TC (asp) for Velocity

325 Asp TC Heat Flux 604.0 53.0 365.0

Soot TCs 326 Bare TC Filter 327 Bare TC Filter 328 Bare TC Gas

329 Bare TC Gas

330 Bare TC Probe

331 Bare TC Probe

332 Bare TC Bath Inlet

0 0 1 333 Bare IL Bath Outlet Heat Mux

rlux Station 2 Heat tlux

Gauges

338 Flux St 2 458.0 56.0 330.0

339 Flux St 2 458.0 44.0 330.0

340 Flux St 2 458.0 23.0 315.0

341 Flux St 2 458.0 35.0 315.0

Column Heat Flux Gauges 342 Column 304.0 77.0 346.0

343 Column 297.0 85.0 327.0

344 Column 304.0 77.0 92.0

345 Column 297.0 85.0 102.0

Floor Heat Flux Gauges 346 Floor 500.0 6.0 13.0

347 Floor 550.0 3.0 13.0 Key: location of thermocouples on structural components varied from test to test. See Chapter 9 and Appendix F for detailed information.

242 NISTNCSTAR 1-5B, WTC Investigation
Appendix C

Gas-Phase Temperature Measurements

This appendix contains data on the gas-phase temperatures measured using aspirated and bare bead

thermocouples (see Figs. C-1 through C-15). Further infonnation on these measurements can be found

in Chapter 4 of this report.

CJ

500

400

300

200

100

—I 1 1 1

1 1 1 1 r Test 2

- •/.- IB

1—I—I—I—I—

ii-t- A- -A -A

_i I I 1 I 100

— West High —•— East High

West Mid • East Mid

- - A - - West Low *- East Low

A

- A- -A- -A- -A - A- A- III -A- - A - A- -A - A

200 300 400

Time (s)

Figure C-1 . Aspirated thermocouple temperatures as a function of time for Test 2.

u

500

400

300

E 200

100

0

.A -A-

_] J I I I L 200

West High

West Mid

West Low — - Eas Eas

t High

i Mid

- -A - t Lxn-v -

• • -A .. -A

• A- -A --A

' A - "

- A- -A

1 1 -A- A1 I -Zs- -A |- A ' A- -A- -A 7 A1 1 1 A- -A

1 1 400 600 800 1000

Time (s) Figure C-2. Aspirated thermocouple temperatures as a function of time for Test 3.

NISTNCSTAR 1-5B. WTC Investigation 243
Appendix C

800

700

600

G^

500

i 400

(L) a.§ 300

H

200

100

0

T 1 r Test 4

3---

I'i

32

,,~0—G\ \

\

\

e West High hast High

West Mid « East Mid

- - A - - West Low

A- A- -A- 'A - A - A- A- -A -A - A- A- -Ar -A - A'

^

-A - A- 'ZV

I I X I I I I I I I I I 1 I \ L_ 0 200 400 600 800

Time (s)

Figure C-3. Aspirated thermocouple temperatures as a function of time for Test 4.

2

a.

(U H

800

700

600 Ir iM'^'^^^'

^

500 m

m

400 I

300

200

100

0

"1 I I t [~ Test 6

^ vv CM n is^J i ««•

- R - - - West Mid - - -

-

- - East Mid

- - A - - West Low A East Low

J I I 'AI I

\ I I I II I L_ .A-'A— A--J I I I u

0 500 1000 1500 2000 2500 3000

Time (s)

Figure C-4. Aspirated thermocouple temperatures as a function of time for Test 6.

244 NISTNCSTAR 1-5B, WTC Investigation
Gas Phase Temperature Measurements

500 Test 1

400

'

300 rr-^^' " 1%

S 200

100

.si '^V? 4 ?

'

0

fTTREl

'rTRE2

TTR.E4

H TTRE5

" TTRIi?

»~" TTRE8

TTRE9

-A---- TTREIO ^ - - TTREl I TTREl

— TTRE13

<1- TTREI4

> TTREl

O TTREl 6

TTREl

J I L_ J I L 0 400 800 1200 1600 2000 2400

Time (s)

Figure C-5. East thermocouple tree temperatures as a function of time in Test 1

500

400 Ta

300

£ 200 -

100

0

—e— TTRWl

[Tj TTRW^

I. I.KAV,:

- -X TTRW4

t TTRW5

- -•- - TTRW7

TTRW8

•— TTRW9

----A---- TTRWIO

TTRWl I TTRWl

—^— TTRWl 3

- TTRWi4

- -C> - - TTRWl

0 200 400 600 800 1000 1200

Time (s)

Figure C-6. West thermocouple tree temperatures as a function of time in Test 2

NISTNCSTAR 7-56, WTC Investigation 245
Appendix C

500

_ 400

u

a 300

s

D.£

H

200 j

100

-A-f

TTREl

ti:r£4

TTRE5

- - TTRE7

TTRE8

TTRE9

TTREIO

1 iKtJ i TTREl 2 TTREl 3 TIRE I

4

TTREl

5

TIRE I

6

TTREl

7

J I L_ 0 200 400 600 800 1000 1200

Time (s)

Figure C-7. East thermocouple tree temperatures as a function of time in Test 2

500

400

300

u

'B

CO

E 200

H

100 t

0

e TTRWl

O TTRW2

TTRWS

X T"T-RW4

H TTRW5

FTRW7

TTRW8 — — TTRW9

A - TTRWIO

~ - T - - TTRWl 1 TTRW12 — TTRWl 3

t> TTRWl 6

I I I 0 400 800 1200 1600 2000 2400

Time (s)

Figure C-8. West thermocouple tree temperatures as a function of time in Test 3

246 NISTNCSTAR 1-5B, WTC Investigation
Gas Phase Temperature Measurements

p

3

500

400

300

E 200

100

TTREl

T"T'R.E2

TTR.E4

H TTRE5

i TTREfi

- TTRE7

TTRE8

TTRJE9

A TTREIO

TTREl 1

2d TTREl

— TTREl 3 <1-- TTREl

->-- TTREl

C^~^ TTREl

-*- TTREl

_1 I L. _l 1 L_ _l I L. 0 400 800 1200 1600 2000 2400

Time (s)

Figure C-9. East thermocouple tree temperatures as a function of time in Test 3

e— TTRWl

TTRW2

i:iR\v?

-X " TTRW4

TTRW5

-«--TTR.W7

TTRW8 — TTRW9

A TTRWIO

- - TTRWl 1 --^r-- TTRWl

^— TTRWl 3

<! - - TTRWI4

1> - - TTRWl

1200 2000 2400

Time (s)

Figure C-10. West thermocouple tree temperatures as a function of time in Test 4.

NISTNCSTAR 1-5B, WTC Irjvestigation 247
Appendix C

800

700

600

u^

500

I 400

I 300

200

100

0

~i—I

h Test 4

1 1 r T 1 r

-ATTREl

"n'R.E2

TTRE;

T"1:R.E4 TTRE5

TTREi

TTRE7

TTRE8

TTRE9

TTRE 10

TTREI I TTRE I

2

TTRE 13

TTRE 15 TTRhlfi

TTRE 17

1 \ I I I I I ! I I L. 0 400 800 1200 1600 2000 2400

Time (s)

Figure C-1 1 . East thermocouple tree temperatures as a function of time in Test 4.

800

700

600

500

400

se 300

H

200

100

0

e TTRW1

-Q TTRW2

<^ TTRW3

•^^ TTRW4

~i TTRW5

• TTRW8

TTRW9

-A---- TTRWIO

- TTRW!

ZJ -- TTRW12 — TTRW13

-^^ TTRW14

i>--TTRW16

I I I I I I I I I I I I I I I I I 0 800 1600 2400 3200 4000 4800 5600 6400

Time (s)

Figure C-1 2. West thermocouple tree temperatures as a function of time in Test 5.

248 NISTNCSTAR 1-5B, WTC Investigation
Gas Phase Temperature Measurements

800

U 600

o

o

400 -

200 -

0

—3— TTREl

a TTRE2

TTRF3

TTRE4

\ TTRE5

TTRE6

-TTRE7

TTRE8 — TTRE9

—A—- TTREIO

- - - TTREl 1 TTRE12 —^—TTRE13

- <5- TTRE14

-->-- TTRE15

0"~ TTREl 6

-O- TTRE17

_L 0 800 1600 2400 3200 4000 4800 5600 6400

Time (s)

Figure C-13. East thermocouple tree temperatures as a function of time in Test 5.

u

a.

1000

800

600

r Test 6

E 400

200

0

"T—i—I—

—8— TTREl

---E-- TTRE2

---X--- TTRE4

^ TTRE5

-•--TTRE7

--»-- TTRE8 — —TTRE9

-A - TTREIO

- - TTREll

-^r-^ TTRE12 —TTRE13

TTRE14

-i>- -TTRE15

C— TTREl 6

TTREl

J I I I I L 0 1000 2000 3000 4000 5000 6000 7000

Time (s) Figure C-14, East thermocouple tree temperatures as a function of time in Test 6.

NISTNCSTAR 1-5B, WTC Investigation 249
Appendix C

3000 4000 5000 6000 7000

Time (s)

Figure C-15. West thermocouple tree temperatures as a function of time In Test 6.

250 NISTNCSTAR 1-5B, WTC Investigation
Appendix D

Velocity Measurements at the Compartment Inlet and

Outlet

This appendix presents data on the inlet and outlet velocities and the corresponding corrected temperature

measurements that was not presented in Chapter 7 of this report (see Figs. D-1 through D^2). Further

information on these measurements can be found in Chapter 7.

WTC Phasel Test 1 Inlet Velocity Measurements

inlet 4 -- Iniet 5

Inlet 6

200 400 600 800

Time (s)

1000 1200

Figure D-1. Inlet velocities for Test 1.

304

302

TO 300

298

296 -

— I

:

1

' ' ' 1 1- WTC Phasel Test 1 Inlet Temperatures

- Inlet 1 - Inlet 2 - Inlet 3

-- inlet 4 - Inlet 5 - Inlet 6

200 400 600 800

Tinne (s)

1000 1200

Figure D-2. Inlet temperatures for Test 1.

NISTNCSTAR 1-5B, WTC Investigation 251
Appendix D

! 1 !;

1

\ 1 1 1

i

1 1:

\ 1

]— WTC Phase 1 Test 2

inlet Velocity Measurements

Hi.!

l&tiMi miy 0*f%^^^^^^^f \^^^^^^^

Inlet 1 Inlet 2

Inlets

Inlet 4

Intet 5

inlet 6

J I \ I I ^ L_ _l i 1 I I \ I I ! I 1 L_ 100 200 300 400

Time (s)

Figure D-3. Inlet velocities for Test 2.

500 600

305

302.5

CD

i_2 300

CDQ.

E

CD

297.5

295

WTC Phasel Test 2

Inlet Temperatures

T 1i 1

1 1 1 r

- Inlet 1 Inlet 2

Inlet 3

inlet 4

Iniet 5

Inlet 6

^1 a-

_J I \ I I I I \ I I 1 l_ I I I I I I I 100 200 300 400

Time (s)

Figure D-4. Inlet temperatures for Test 2.

500 600

252 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

2.25

^> 1.5 o

_o>

0.75 t

200 400 600 800 1000 1200 1400

Time (s)

Figure D-5. Inlet velocities for Test 3.

300 ' '

!

'

^ '

I WTC Phasel Test 3 inlet 1

0 200 400 600 800 1000 1200 1400

Time (s)

Figure D-6. Inlet temperatures for Test 3.

NISTNCSTAR 1-5B, WTC Investigation 253
Appendix D

600 800

Time (s)

1200

Figure D-7. Inlet velocities for Test 4.

I I I 1 I I i

!

I [ 1 I I I

i

1 \ I I I I I I I 0 200 400 600 800 1000 1200

Time (s)

Figure D-8. Inlet temperatures for Test 4.

254 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

3 r WTC Phasel Test 5

Inlet Velocity Measurements

0.6 inlet 1 Inlet 2

Inlet 3

.J ] I I I I I I I I I I L

In let 4

inlet 5

Inlet 6

_J \ I 1 L 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-9. Inlet velocities for Test 5.

340 1 1

1 ; 1 1 1 \ r WTC Phasel Test 5

Inlet Temperatures

330 -

290

Inlet 1 Inlet 2

Inlet 3

Inlet 4

inlet 5

Inlet 6

400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-10. Inlet temperatures for Test 5.

NISTNCSTAR 1-5B, WTC Investigation 255
Appendix D

1—I—I

1 1—I—I—I—

r

"1

i

i r

2.4

I

1.8

o

o^ 1.2

WTC Phasel Test 6

inlet Velocity Measurements

I !

0.6 Inlet 1 Inlet 2

Inlet 3

inlet 4 -

J I I

[ I I L_ 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-1 1 . Inlet velocities for Test 6.

340

330

a 320

0)1 310

CD H

WTC Phasel Test 6

Inlet Temperatures

300

290

iL.l:

Inlet 1 Inlet 2

Inlet 3

Inlet 4

inlet 5

inlet 6

400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-1 2. Inlet temperatures for Test 6.

256 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

10

oo

0) 4

WTC Phasel Test 1 Outlet Velocity Measurements

o r 11-

1

1 1 1—

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outlets H

>

rii;

.5

' h '

1 lit

_l I L _1 I L_ ..L_.. J 0 200 400 600 800

Time (s)

1000 1200

Figure D-13. Outlet velocities for Test 1, outlet locations 1, 2, 3, 4, and 5.

10

C/)

_o

CD >

WTC Phasel Test 1 Outlet Velocity Measurements

Outlet 6

Outlet 7

Outlet 8

J I L. 200 400 600 800

Time (s)

1000 1200

Figure D-14. Outlet velocities for Test 1, outlet locations 6, 7, and 8.

NISTNCSTAR 1-5B, WTC Investigation 257
Appendix D

800

700

V 600

I—3-I—

»

(0

i_

(UE 500

0)

400

300

-T ] T "T 1 1 1

1 1 r WTC Phasel Test 1 Outlet Temperatures

I.

4'

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outlet 5

_I ] L_

I. I %Mi f

200 400 600 800

Time (s)

1000 1200

Figure D-15. Outlet temperatures for Test 1, outlet locations 1, 2, 3, 4, and 5.

800

700

^ 600

CO

\—CDE 500

400

300

—1—i—I—^—1

\—I—\—I—\—

WTC Phasel Test 1 Outlet Temperatures

ii si' ['(J

, Iff 1 pi

If iMj I ^

; fir

r

Outlet 6

Outlet 7

Outlet 8

_1 i t u 200 400 600 800

Time (s)

1000 1200

Figure D-16. Outlet temperatures for Test 1, outlet locations 6, 7, and 8.

258 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

8 1

' I' '

\

<

I I

'

I

I I

1 1 \1 1 1 1

1 1 1f 11 i i

1 r

E WTC Phasel Test 2 Outlet 1

-_ 7 r Outlet Velocity Measurements Outlet 2 z

- Outlet 3 :

Time (s)

Figure D-17. Outlet velocities for Test 2, outlet locations 1, 2, 3, 4, and 5.

3 p—I—1—I—I—

I—I

'•—I—I

\—I

'—I—I—j—I—I—I—I—

j—I—I—I—I—]—I—i—I—I

I WTC Phasel Test 2 Outlet 6 \

0 100 200 300 400 500 600

Time (s)

Figure D-18. Outlet velocities for Test 2, outlet locations 6, 7, 8, 9, and 10.

NISTNCSTAR 1-5B, WTC Investigation 259
Appendix D

800

700 -

V 600

103E 500

0

n ^ 1 !] r—i

1 1

j 1

' 1 1

\ r- WTC Phasel Test 2

Outlet Temperatures

-1 1

1 r

400

300

lJiM \ VI , ,,,,

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outlet 5

_l 1 I I I I I I L_ I ' I J \ I I I I I I I L 100 200 300 400

Time (s)

500 600

Figure D-19. Outlet temperatures for Test 2, outlet locations 1, 2, 3, 4, and 5.

800

700

^ 600

TO

E 500

CD

400

300

n 1 1 1] i

1 1 1

[ 1

i 1 1 1

1 r WTC Phasel Test 2

Outlet Temperatures

/i i / A .^Wa^^^^ "'I'm «ii i!\ I

1—^

'—'—

Outlet 6

Outlet 7

Outlet 8

Outlet 9

Outlet 10

ii !

J \ I I I i \ \ I

I I i I L. _J I I L 100 200 300 400

Time (s)

500 600

Figure D-20. Outlet temperatures for Test 2, outlet locations 6, 7, 8, 9, and 10.

260 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

o

o

a; >

WTC Phasel Test 3

Outlet Velocity Measurements

n \ 1

i

1 r 1 ' ^

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outiet 5

7

6

5

4

3

2

1

0 200 400 600 800 1000 1200 1400

Time (s)

Figure D-21. Outlet velocities for Test 3, outlet locations 1, 2, 3, 4, and 5.

o

(D >

WTC Phasel Test 3

Outlet Velocity Measurements

Outlet 6

Outlet 7

Outiet 8

Outlet 9

Outlet 10

400 600 800

Time (s)

1000 1200 1400

Figure D-22. Outlet velocities for Test 3, outlet locations 6, 7, 8, 9, and 10.

NISTNCSTAR 1-5B, WTC Investigation 261
Appendix D

800

700

^ 600

i

fo

CD£ 500

CD

400

300

n ; 11

1 11

j

i 1 1

WTC Phasel Test 3

Outlet Temperatures

n 1 i

\ 1 r

1

J -Mi !

" f f

,- v. L

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outlet 5

1 Mw

m

J 1 I I I I I I L 0 200 400 600 800 1000 1200 1400

Time (s)

Figure D-23. Outlet temperatures for Test 3, outlet locations 1, 2, 3, 4, and 5.

800

700

^ 600

-4—

E 500

0)

400

300

T 1 1

1

1 1 1 1

1 r 1 ' ' '

I

' ' '

i WTC Phasel Test 3

Outlet Temperatures

"T 1 1 1 1 1 r

Outlet 6

Outlet 7

, Ai^^'/vA'^^i Outlet 8

Outlet 9

, t' A Outlet 10

0 200 400 600 800 1000 1200 1400

Time (s)

Figure D-24. Outlet temperatures for Test 3, outlet locations 6, 7, 8, 9, and 10.

262 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

0 200 400 600 800 1000 1200

Time (s)

Figure D-25. Outlet velocities for Test 4, outlet locations 1, 2, 3, 4, and 5.

NISTNCSTAR 1-5B, WTC Investigation 263
Appendix D

1100

1000

900 ^ 800

2 700

0)Q.

i 600

500

400

300

n 1 r WTC Phasel Test 4

Outlet Temperatures

n i

< 1 r—

Outlet 1 Outlet 2

Outlet 3

- Ouliet 4

Outlet 5

1 I

0 200 400 600 800

Time (s)

1000 1200

n ! 1

;

1 1 r "T \ T"

Figure D-27. Outlet temperatures for Test 4, outlet locations 1, 2, 3, 4, and 5.

1100

1000

900

- I ^

^

< i I

^

i

I WTC Phasel Test 4

Outlet Temperatures

800 -

I 700 ^Md^

Outlet 6

Outlet 7

Outlet 8

Outlet 9

Outlet 10

CDQ.

i 600

500

400 1

300

I j '•

!

0 200 400 600 800 1000 1200

Time (s)

Figure D-28. Outlet temperatures for Test 4, outlet locations 6, 7, 8, 9, and 10.

264 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

14

12

10

WTC Phasel Test 5

Outlet Velocities

Outlet 1 Outlet 2

Outlet 3

Outlet 4

Outlet 5

1' ; !

I

400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-29. Outlet velocities for Test 5, outlet locations 1, 2, 3, 4, and 5,

14 i—I—'—'—

I—'—'—I—I—I—I—I—

T" WTC Phasel Test 5

12 \- Outlet Velocities

-i—I—

I—I—I—I—I—;—i—I—I—I—I—

Outlet 6 Outlet 9

Outlet 7 Outlet 1 0 -\ Outlet 8

J \ I L. 0 400 800 1200 1600 2000 2400 2800 3200

Time (s) Figure D-30. Outlet velocities for Test 5, outlet locations 6, 7, 8, 9, and 10.

NISTNCSTAR 1-5B, WTC Investigation 265
Appendix D

1200

1000

800

Q)

CD

Q.

I- 600

400

1—i—I—

r

WTC Phasel Test 5

Outlet Temperatures

Outlet 1 Outlet 2

Outlet 3

J I I 1 I I I I 1 I I I I I L_ _l 1 I i 1 L_ 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-31. Outlet temperatures for Test 5, outlet locations 1, 2, 3, 4, and 5.

1200

1000

CD B

800

"fo k_

(UQ.

E

CD H 600

400

"T 1 ]

] 1 T WTC Phasel Test 5

Outlet Temperatures

Outlet 6

Outlet 7

Outlet 8

Outlet 9

Outlet 10

J 1 1 I I

I 1 I L_ 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-32. Outlet temperatures for Test 5, outlet locations 6, 7, 8, 9, and 10.

266 NISTNCSTAR 1-5B. WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-33. Outlet velocities for Test 6, outlet locations 1, 2, 3, 4, and 5.

14

12

10

WTC Phasel Test 6

Outlet Velocities

1—j—i—I—I—

I—1—I—I—]—I—I—i—I—I—i—i—I

\—\—

Outlets Outlet 9

Outlet? Outlet 10

Outlet 8

0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-34. Outlet velocities for Test 6, outlet locations 6, 7, 8, 9, and 10.

NISTNCSTAR 1-5B, WTC Investigation 267
Appendix D

1200

1000

800 i_ ^—

03

CDCl

0) I- 600

400

n—I\—]

•—I—I—

I—1—I—I—

]—I

1—I—

p

WTC Phasel Test 6

Outlet Temperatures

-1—1

1 1 r

Outlet 1 Outlet 2

Outlet 3

I I I I I I _l I I

I 1 J I L 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-35. Outlet temperatures for Test 6, outlet locations 1, 2, 3, 4, and 5.

1200

1000

'

' '

:

' ^ '

I

' ' '

I

' WTC Phasel Test 6

Outlet Temperatures

n i

1 1 1 1 1 1 r

800

CD

i_ _2

TO(DQ.

0) 1- 600

400 ^

Outlet 6

Outlet 7

Outlet 8

- Outlet 9

Outiet 10

J \

I I L_ 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-36. Outlet temperatures for Test 6, outlet locations 6, 7, 8, 9, and 10.

268 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

700 ~i I

I r WTC Phasel Test 1 Outlet 5 Temperatures

1 1 \ 1

1

1 1 r

Aspirated

Corrected Bare-bead

200 400 600 800 1000 1200

Time (s)

Figure D-37. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 1.

F

700

""

r

1

^

1 1 : 1

^

1

i r- WTC Phasel Test 2

Outlet 5 Temperatures

1 1 1 [ 1 1 r

^ 600

CD

^ 500 h

E

0

400

300 J I I I I I 1 I I L

Aspirated

Corrected Bare-bead

I I

1 L 0 100 200 300 400 500 600

Time (s)

Figure D-38. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 2.

NISTNCSTAR 1-5B, WTC Investigation 269
Appendix D

700

650

600

^

550

2 500

CDQ.

i 450

400

350

300

WTC Phasel Test 3

Outlet 5 Temperatures

Aspirated

Corrected Bare-bead

mi,

J , I L_ 200 400 600 800

Time (s)

1000 1200

Figure D-39. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 3.

800

WTC Phasel Test 4

Outlet 5 Temperatures

300

Aspirated

Corrected Bare-bead

_J 1 L 200 400 600 800

Time (s)

1000 1200

Figure D-40. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 4.

270 NISTNCSTAR 1-5B, WTC Investigation
Velocity Measurements at the Compartment Inlet and Outlet

800

WTC Phasel Test 5

Outlet 5 Temperatures

Aspirated

Corrected Bare-bead

_1 1 I L. _l I I L_ _l I I I I I L_

400

300

0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-41. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 5.

800

700

T i

1

1

1

1 1 11 1 \ 11 r WTC Phasel Test 6

Outlet 5 Temperatures

400 r

300

Aspirated

Corrected Bare-bead V;

_i I I 1^ _L 1 I I I I I L^..J I 0 400 800 1200 1600 2000 2400 2800 3200

Time (s)

Figure D-42. Aspirated and bare-bead thermocouple temperatures at outlet 5 for Test 6.

NISTNCSTAR 1-5B, WTC Investigation 271
Appendix D

This page intentionally left blank.

272 NISTNCSTAR 1-5B, WTC Investigation
Appendix E

Heat Flux Measurements

This appendix presents heat flux data that was not presented in Chapter 5 of this report (see Figs. E-1

through E-24). Further information on these measurements can be found in Chapter 5.

(0 X

o

35

30

25

20

15

10

5

0

rn—I—I— —i—

J I I I I I I I I I I I I I I I L 200 400 600

Time (s) 800 1000 1200

Figure E-1. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 1.

ICQ

^ 80

CDO

60

S 40

20

Test 1 HCel

HCe2

-o - HCe3

-X- - HCe4

200 400 600

Time (s) 800 1000 1200

Figure E-2. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 1. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 273
Appendix E

X

03

CD X

100

80

60

40

20

0

"1 1 r ^ - H2FU

H2RU

~o- H2FD

-X- - H2RD

Test 1

J£ I i I I 1 \ L. I I I 200 400 600

Time (s)

800 1000 1200

Figure E-3. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 1.

X

13

03OX

o

60

50

40

30

20

10

0

T 1 1

1

1 1 r ~i 11 1 r

- HCoHF

HCoHW

^

HCoLF

-X- HCoLW

Test 1

y;[;|i;ji.iUv.i.i^.- ^

,„,,..„... ...-^y

1

J ] u J I L_ 200 400 600

Time (s)

800 1000 1200

Figure E-4. Heat flux recorded by the two radiometers and two flux gauges on the

column in Test 1.

274 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

X

3

03

CD X

O

30

25

20

15

10

5

0

-5

Test 2 ^3— HF1

-B— HF2

_J 1 L_ 0 200 400 600

Time (s)

800 1000

Figure E-5. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 2.

X

3

CD X

o

120

100

80

60

40

20

0

T 1 r T 1 r

mm

Test 2

I

1 I i_

9

- e HCel

HCe2

^ - HCe3

- -X- - HCe4

J I 'I li L_ fir

0 200 400 600

Time (s)

800 1000

Figure E-6. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 2. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 275
Appendix E

X

03

CD X

o

60

50 r

40

30

20

10

0

~T 1 r n \ r

if

Test 2 HCoHF - HCoHW

-o — HCoLF

- -X - - HCoLW

J I I L Li

0 200 400 600

Time (s)

800 1000

Figure E-7. Total heat flux recorded by the four flux gauges in structures adjacent to the

vertical column in Test 2. The periodic spikes are due to the soot purge operation on the

two higher gauges.

X

3

CO

CD X

70

60

50

40

30

20

10

0

"T [ r

Test 2 ^- H2FU

&- H2RU

-o- H2FD

X- H2RD

I III

•i mm

_ I'

' . '-Mi

0 200 400 600

Time (s)

800 1000

Figure E-8. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 2.

276 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

X

CDOX

O

30

25

20

15

10

5

0

T I I

I

I I r

HF1

HF2

Tests

I ,,1

111

'

_l I 1 L_ 0 200 400 600

Time (s)

800 1000 1200

Figure E-9. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 3.

100

E

X

Z5

03

CD X

o

80 -

60

40 -

20 -

0

1 I I

I r Tests

1 1 1 1 1 1 r

G HCe1

HCe2

-o — HCe3

- -X - - HCe4

I I I I I I I I I U L. 400 600

Time (s)

800 1000 1200

Figure E-10. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 3. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 277
Appendix E

X

030) I

"to o

60

50

40

30

20

10

0

-e— HCoHF

-a— HCoHW

-o- HCoLF

-X- HCoLW

Tests

m

l- f I

I L

m

1 1

J I I L_ -J L 0 200 400 600

Time (s)

800 1000 1200

Figure E-11. Total heat flux recorded by the four flux gauges in structures adjacent to

the vertical column in Test 3. The periodic spikes are due to the soot purge operation on

the two higher gauges.

X

CO0X

70

60

50

40

30

20

10

0

"T ; r

-X

H2FU

H2RU

H2FD

H2RD

Tests

fill ^1 ' Tl J^'^^.X^^^Ji^.^Aiihilf^^^K^

1 1 (v|(l'^'

_l ! L 0 200 400 600

Time (s)

800 1000 1200

Figure E-12. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 3.

278 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

80

X

_2

LL

"to CD X

O

70 h

60

50

40

30

20

10

0

"1 1 r

-^-HF1 ^

HF2

"1 ' ' ' r

1 1 1

i

1 r

i! I I \mm\i

Test 4

V'

-1 I L. _l I L_ 200 400 600

Time (s)

800 1000 1200

X

o

60

40

20

0

T 1 r

Figure E-13. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 4.

160

140

^ 120

I 100

X

il 80

—e

HCel — HCe2

-o - HCeS

- -X - - HCe4

0 200 400 600

Time (s)

800 1000 1200

Figure E-14. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 4. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 279
Appendix E

X

030X

o

100

80

60

40

20

—e—- HCoHF

HCoHW

o - HCoLF

- -X - - HCoLW

~i 1 r Test 4

-'/ /X"*"

i > i

: [

.i^'X'

'^-^t-,

J \ I I \ 0 200 400 600

Time (s)

800 1000 1200

Figure E-15. Total heat flux recorded by the four flux gauges in structures adjacent to

the vertical column in Test 4. The periodic spikes are due to the soot purge operation on

the two higher gauges.

150

100

X

13

^03 50

X

0

~i 1 r T 1 1

1

i r T ] 1

1

1

i r

e H2FU

-B— H2RU

H2FD

-X- H2RD

Test 4

m

Jiif

""^ m

mm

_l I L J ! L.

(p

<>

200 400 600

Time (s)

800 1000 1200

Figure E-16. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 4.

280 NISTNCSTAR 1-5B, WTC Investigation
Heat Flux Measurements

120

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-17. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 5.

200

^

150 -

X

CO X

o

100

50

500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-18. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 5. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 281
Appendix E

X

L!_ fo

CD X

O

160

140

120

100

80

60

40

20

0

1—I—I—I—

I—i—i—I—I

\—

r

— HCoHF e

-B

HCoHW

o HCoLF

- - X - - HCoLW

Tests

1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-19. Total heat flux recorded by the four flux gauges in structures adjacent to

the vertical column in Test 5. The periodic spikes are due to the soot purge operation on

the two higher gauges.

350

300

250

200

150

n ! I I r T I I

1 I 1 r

H2FU

-E

H2RU

-0 - H2FD

- -X - - H2RD

Tests

500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-20. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 5. Gauges H2RD and H2RU were damaged during the test.

282 NISTNCSTAR 1-5B. WTC Investigation
Heat Flux Measurements

120

X

13

X

o

~I

:—i—

500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-21. Total heat flux recorded by the two flux gauges 15 cm above the floor, just

downstream of the fire pan in Test 6.

200

150 -

n: 100 u

03

o 50

0

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-22. Total heat flux recorded by the four flux gauges in the compartment ceiling

in Test 6. The periodic spikes are due to the soot purge operation.

NISTNCSTAR 1-5B, WTC Investigation 283
Appendix E

X

3

3

o

200

150

d 100

CD(U X

50

T "1 I

I—I—I

'—I—I—i—i

I—I—I—I—i—I—

r

HCoHF

HCoHW

o " HCoLF

X - - HCoLW

Teste

_1 [ I I I ] I I i I \ I L 0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-23. Total heat flux recorded by the four flux gauges in structures adjacent to

the vertical column in Test 6. The periodic spikes are due to the soot purge operation on

the two higher gauges.

300

250

\ 200

X

CD

0) X

n \ I 1

i

\—I I I

I—i—I—I—i—i—I

I—I—II—I—1—1—i—I—1—i—I

I—I—I—I—!—I—1—I—I—I—

r

H2FU

-B—- H2RU

O H2FD

- -X - H2RD

150

100

50

0 J 1 I I \ : I : L

'III \[.\ M

Test 6

M

500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Figure E-24. Heat flux recorded by the two radiometers and two flux gauges in the

measurement station in Test 6. Gauges H2FD, H2RU, and H2RD were damaged during

the test.

284 NISTNCSTAR 1-5B, WTC Investigation
Appendix F

Thermocouple Locations on Steel Components

Chapter 9 presents representative measurements of the temperature on the surface of the steel components

and the sprayed fire-resistive material (SFRM) applied to those components. This appendix documents

the location of the thermocouples on the surface of the SFRM and the steel components in terms of their position within the test compartment (Tables F-1 through F-5).

Table F-1. Thermocouple locations on bars in Tests 1, 2, and 3. Simple Bar Test Chamber Coordinates (cm)

TC label X Y Z

Location 1

B-1 350 108 295

B-2 350 17 295

B-3 350 -43 295

B-4 350 -134 295

Location 2 B-1 500 102 295

B-2 500 12 295

B-3 500 -49 295

B-4 500 -139 295

Table F-2. Thermocouple locations on bars in Tests 5 and 6. Simple Bar Test Chamber Coordinates (cm)

TC Label X Y Z

Location 1

B-1 350 150 295

B-2 350 58 295

B-3 350 -3 295

B-4 350 -94 295

NISTNCSTAR 1-5B. WTC Investigation 285
Appendix F

Table F-3. Thermocouple locations on columns.

Column

Face Orientation

1 North 2 East

3 South 4 West Test Chamber Coordinates (cm)

TC Label X Y Z

CU-1 281 69 369

CU-2 299 55 369

CU-3 281 41 369

CU-4 263 55 369

CM-1 281 69 213

CM-2 299 55 213

CM-3 281 41 213

CM-4 263 55 213

CL-1 281 69 77

CL-2 299 55 77 CL-3 281 41 77 CL-4 263 55 77

Key: CU, column upper; CM, column middle; CL, column lower.

Table F-4. Thermocouple locations on truss A (location 1). Truss A Test Chamber Coordinates (cm)

TC Label X Y Z

TU-1 164 -8 370

TU-2 313 -8 370

TU-3 388 -8 370

TU-4 537 -8 370

TM-1 199 -8 329

TM-2 335 -8 329

TM-3 365 -8 329

TM-4 502 -8 329

TL-1 250 -8 289

TL-2 300 -8 289

TL-3 401 -8 289

TL-4 451 -8 289

Key: TU, truss upper chord; TM, truss middle (web); TL, truss lower chord.

286 NISTNCSTAR 1-5B, WTC Investigation
Thermocouple Locations on Steel Components

Table F-5. Thermocouple locations on truss B (location 2).

Truss B Test Chamber Coordinates (cm)

TC Label X Y Z

TU-1 164 -56 370

TU-2 313 -56 370

TU-3 388 -56 370

TU-4 537 -56 370

TM-1 199 -56 329

TM-2 335 -56 329

TM-3 365 -56 329

TM-4 502 -56 329

TL-1 250 -56 289

TL-2 300 -56 289

TL-3 401 -56 289

TL-4 451 -56 289

Key: TU. truss upper chord; TM. truss middle (web); TL, truss lower chord.

NISTNCSTAR 1-5B, WTC Investigation 287
Appendix F

This page intentionally left blank.

288 NISTNCSTAR 1-5B, WTC Investigation
Appendix G

Temperature Histories of Bare and Coated Steel

Components

Chapter 9 presents representative measurements of the temperature on the surface of the steel components

and the sprayed fire-resistive material (SFRM) applied to those components. This appendix presents the

measured temperature histories recorded for steel components and not shown in Chapter 9.

G.1 BARE steel COMPONENTS

Tests 1, 2. and 3 exposed the same set of two bars, one column, and two trusses to direct exposure of heat

flux and temperatures from a fuel pan fire. In all three tests, the themial response of these components to flame proximity or immersion was similar. The steel began to heat immediately upon initiation of the fire and to cool immediately after the fire was extinguished. The recorded response showed no lag time

between the steel and gas for transferring radiant heat from the higher to lower temperature mediums.

Test 1 steel surface temperatures are shown in Sec. G.1.1, Figs. G-1 through G-5; Test 2 temperatures are shown in Sec. G.1.2, Figs. G-6 through G-10; Test 3 temperatures are shown in Sec. G.1.3, Figs. G-1

through G-1 5.

Bar A and both truss A and truss B were located over the fire pan, but the fire tended to lean with the airflow through the chamber and primarily immerse bar A and truss B in flames. The recorded

thermocouple temperatures for these components (1) are 50 °C to 75 °C higher than those recorded for bar B (Figs. G-2, G-7, and G-1 2) and truss A (Figs. G^, G-9, and G-14) and (2) have a high frequency

component imposed upon the average temperature. These behaviors can be seen for thennocouples TB3

and TB4 on the bars and thermocouples TL3, TU3, TL4, and TU2 on the trusses. See Figs. G-1, G-6,

and G-1 1 for bar A and Figs. G-5, G-10, G-1 5 for truss B. Note that thermocouples TM2 and TM4 on

truss B showed some irregularities in the cool down phase in Test 1 , after the fire was extinguished,

which was not evident in Test 2 but was more pronounced in Test 3. The cause of this in'egularity is not known.

The columns were located next to the fire pan, but the airflow through the test chamber resulted in the fire leaning toward the exhaust outlet and away from the columns, so that they were heated primarily by

radiation from the fire and the hot gas layer near the ceiling. The sides of the column facing the fire were

heated more quickly, with TCU3 (at the ceiling level facing south), TCU2 (at the ceiling level facing

east), and TCL2 (at the floor level facing east) recording maximum temperatures during the tests (Figs. G-3, G-8, and G-1 3).

NISTNCSTAR 1-5B, WTC Investigation 289
Appendix G

The thermocouple naming convention used in the figures is as follows:

For bars (i.e., TB4SB or TB3S):

TB Bars

1,2,3,4 Location of thermocouple along length

S Steel surface

A, B Bar number (for bare steel tests)

For columns (i.e., TCU3S or TCU2S):

TC Columns

U, M, L Upper, middle, or lower location with respect to elevation

1, 2, 3, 4 Location of thermocouple along length

S Steel surface

For trusses (i.e., TL2SA or TM3SB):

T Truss

U, M, L Upper, middle, or lower location with respect to elevation

1,2,3,4 Location of thennocouple along length

S Steel surface

A, B Bar or truss number

G.1.1 Bare Steel Components Exposed to a 2 MW Fire for 15 Min in Test 1

700

600

500

o 400

£^ 300

200

100

0 600 1200

Time (sec)

1800 2400

.TB1SA TB2SA TB3SA TB4SA

Figure G-1. Bare steel bar A in Test 1.

290 NISTNCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

TCU4S

-TCL1S

TB1SB TB2SB TB3SB TB4SB

Figure G-2. Bare steel bar B in Test 1.

600 1200

Time (sec)

1800

-TCU1S

-TCL4S

TCU3S

-TCL2S

-TCU2S

TCL3S

-TCM1S -TCM4S

2400

-TCM2S TCM3S

Figure G-3. Bare steel column in Test 1.

NiSTNCSTAR 1-5B, WTC Investigation 291
Appendix G

1200

Time (sec)

-TU4SA TU3SA TU2SA TU1SA TM4SA

.TL1SA TL4SA - TL3SA TL2SA

T1VI3SA —TM2SA TM1SA

Figure G-4. Bare steel truss A in Test 1.

600 1200 1800

Time (sec)

2400

TU1SB —TU2SB TU3SB TU4SB —TM1SB TM2SB TM3SB —TM4SB

TL1SB TL2SB TL3SB -—TL4SB

Figure G-5. Bare steel truss B in Test 1.

292 NISTNCSTAR 1-5B, WTC Investigation
Time- Tempera ture Histories of Bare and Coated Steel Components

G.1 .2 Bare Steel Components Exposed to a 3 MW Fire for 7 Min in Test 2

700

600

500

o 400

Q.E

® 300

200

100

0 100 200 300 400 500

Time (sec)

600 700

-TBISA TB2SA TB3SA TB4SA

Figure G-6. Bare steel bar A in Test 2.

800

Q.E

300

200

100

0

100 200 300 400 500

Time (sec)

600

.TB1SB TB2SB TB3SB TB4SB

Figure G-7. Bare steel bar B in Test 2.

700 800

NISTNCSTAR 1-5B. WTC Investigation 293
Appendix G

700

100

0 100

-TCU4S

-TCL1S

200 300 400 500

Time (sec)

600

-TCU1S

TCL4S

TCU3S

-TCL2S

-TCU2S

TCL3S

-TCM1S TCM4S

700

TCM2S

800

TCM3S

Figure G-8. Bare steel column in Test 2.

1 100

1 "

200 300 400

Time (sec)

1 500

. ,.

,| „

600

1 7001 800

nJ4SA

TL1SA

TUSSA

TL4SA

"nJ2SA - TL3SA .

........ TU1SA

TL2SA

TM4SA ^ nvissA TTVI2SA TMISA

Figure G-9. Bare steel truss A in Test 2.

294 NIST NCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

D 100 200 300 400 500 600 700 800

Time (sec) —TUISB -—"m2SB TU3SB -—-nj4SB TM1SB ~—TIVI2SB TM3SB - —nvi4SB —Ill SB -— TL2SB TLSSB TL4SB

Figure G-10. Bare steel truss B in Test 2.

G.1.3 Bare Steel Components Exposed to a 2 MW Fire for 15 Min in Test 3

600 1200

Time (sec)

1800 2400

TB1SA TB2SA TB3SA TB4SA

Figure G-1 1 . Bare steel bar A in Test 3.

NISTNCSTAR 1-5B, WTC Investigation 295
Appendix G

700 T

600

500

Figure G-1 2. Bare steel bar B in Test 3.

600 -

500

o 400

0 600 1200 1800 2400

Time (sec)

TCU4S TCU1S TCU3S TCU2S TCM1S TCM4S —TCM2S

TCM3S —TCL1S TCL4S TCL2S TCL3S

Figure G-1 3. Bare steel column in Test 3.

296 NISTNCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

600 1200 1800 2400

Time (sec)

UMSA TUSSA TIJ2SA TU1SA TM4SA TM3SA -—nvi2SA —nviiSA TL1SA TL4SA TLSSA TL2SA

1

Figure G-14. Bare steel truss A in Test 3.

Time (sec)

TU1SB TU2SB TU3SB ~ - TU4SB TM1SB -—TM2SB TM3SB TM4SB

TL1SB TL2SB TL3SB TL4SB

Figure G-15. Bare steel truss B in Test 3.

NISTNCSTAR 1-5B, WTC Investigation 297
Appendix G

G.2 INSULATED STEEL COMPONENTS

Tests 5 and 6 exposed separate sets of one bar, one column, and two trusses, insulated with sprayed fire- resistive materials (SFRM), to direct exposure of heat flux and temperatures from a fuel pan fire. Test 4 was truncated after approximately 1 5 min, due to a malfunction of the thennocouples that started a few

minutes after the tests began. As no useful steel temperature data was obtained, Test 4 data are not included in this report. Test 5 SFRM and steel surface temperatures are shown in Sec. G.2.1, Figs. G-16

tlirough G-19; Test 6 temperatures are shown in Sec. G.2.2, Figs. G-20 through G-23.

In both tests, the thermal response of these components to flame proximity or immersion was similar. The

SFRM surface began to heat immediately upon initiation of the fire, reaching close to their peak

temperatures in 5 min to 10 min, and to cool immediately after the fire was extinguished. The protected

steel responded more slowly, taking approximately 50 min with 1.91 cm (0.75 in.) of SFRM to reach

600 °C temperatures as compared to 7 min when bare steel was subjected to a similar 3 MW fire. The

steel temperature rise was essentially uniform for all thennocouples until the steel temperatures reached

100 °C, when the rate of temperature rise increased for all thermocouples, but at different rates. Steel

temperatures continued to rise approximately 25 °C to 50 °C after the fire was extinguished over a 5 min

to 1 0 min period, before cooling began to take place by transmitting heat through the insulation back into

the relatively cooler gases.

The bar in each test was located over the fire and recorded similar temperatures for the SFRM and steel surface. Thermocouples TBIS and TB4S in Test 5 (Fig. G-16) were damaged during the SFRM

application process. Thermocouples TB4 in Test 6 (Fig. G-20) showed some signal noise in the cool

down phase in Test 1 after the fire was extinguished. The cause of this noise is not known.

The columns were located next to the fire pan, but the fire leaned toward the exhaust outlet and away

from the columns, so that they were heated primarily by radiation from the hot gas layer near the ceiling. In Test 5 (Fig. G-17), the steel temperature rise was uniform for all thermocouples until the steel temperatures reached 100 °C, when the rate of temperature rise increased for all the thermocouples but at different rates. Note that thermocouple TCUIS in Test 5 was damaged during the SFRM application

process. In Test 6 (Fig. G-21), however, the steel temperatures increased at different rates during the

entire test. Reasons for the difference in the steel surface response between these two column tests are not apparent. (A short in the thermocouple signal appeared around 1,500 s to 1,600 s, but the recorded

temperatures appear to be reasonable after this event and have been assumed to be correct records of the

actual temperatures.) Test 5 steel temperatures at 50 min were approximately half of the steel temperatures in Test 6, reflecting the increased nominal thickness of SFRM from 1.91 cm to 3.81 cm

(0.75 in. to 1.5 in.). The upper portion of the columns near the ceiling heated more quickly than the rest of the column, due to the continual presence of a hot gas layer near the ceiling.

As in the bare steel tests, both truss A and truss B were located over the fire pan, but the fire primarily

immersed truss B in flames. The trusses with 1.91 cm (0.75 in.) of SFRM reached 100 °C in approximately 10 min (Figs. G-18, G-22, and G-23), and truss B in Test 5 with 3.81 cm (1.5 in.) of

SFRM in Test 5 reached 100 °C steel temperatures in 25 min (Fig. G-19), reflecting the effects of

increased SFRM thickness. Thermocouples TM3IB and TL3IB became erratic approximately 35 min into Test 5 (these thermocouples are on the SFRM surface and subject to flame immersion during the test). Some of the other SFRM surface thermocouples also recorded abrupt jumps in their temperatures around

298 NISTNCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

this time, but the steel temperatures appear to be unaffected. Thennocouple TM3SA for truss A in Test 5 was damaged during the SFRM apphcation process.

The thermocouple naming convention used in the figures is as follows:

For bars (i.e., TB4SB or TB3I):

TB Bars

1, 2, 3, 4 Location of thermocouple along length

S Steel surface

I Insulation (SFRM) surface

A, B Bar number (for bare steel tests)

For columns (i.e., TCU3S or TCU2I):

TC Columns

U, M, L Upper, middle, or lower location with respect to elevation

1, 2, 3, 4 Location of thermocouple along length

S Steel surface

I Insulation (SFRM) surface

For trusses (i.e., TL2SA or TM3IB):

T Truss

U, M, L Upper, middle, or lower location with respect to elevation

1, 2, 3, 4 Location of thermocouple along length

S Steel surface

I Insulation (SFRM) surface

A, B Bar or truss number

G.2.1 Insulated Steel Components Exposed to a 3 MW Fire for 50 Min in Test 5

1000 r-^

900 i

0 il- — —— ______ 1

0 600 1200 1800 2400 3000 3600 4200 4800

Time (sec)

TB2S TB3S TBI I TB2I TB3I TB4I

Figure G-16. Insulated (1.91 cm SFRM) steel bar in Test 5.

NISTNCSTAR 1-5B, WTC Investigation 299
Appendix G

1000

900

800

700 ~ 600 o

^ 500

0) *- 400

300 f

200

100

0

. ^^=^<=si.>-^TM"—'

600 1200 1800 2400 3000 3600 4200

Time (sec)

-TCU2S TCU3S TCU4S TCM1S TCM2S TCM3S TCM4S TCL1S

TCL3S TCL4S TCU1I TCU2I TCU3I TCU4I TCIV12I TCM3I

.TCL1I TCL2I TCL31 TCL4I

Figure G-17. Insulated (3.81 cm SFRM) steel column in Test 5.

4800

TCL2S

TCM4I

U 1

0 600 1200 1800 2400

Time (sec]

3000 3600 4200 4800

TU1SA TU2SA TU3SA nj4SA TM1SA TM2SA TM4SA —TL1SA TL2SA

TL3SA TL4SA TU1IA TU2IA nj3IA "nj4IA —TM2IA .—TM3IA T1VI4IA

^...._.U1IA TL2IA TL3IA TL4IA

Figure G-18. Insulated (1.91 cm SFRM) steel truss A in Test 5.

300 NISTNCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

i 1

600 1200 1800 2400 3000 3600 4200 4800

Time sec)

TU1SB —- "0)256 TU3SB --"nj4SB -—TMISB —-TM2SB TIVI3SB —-"nVI4SB

-—TL1SB TL2SB TL3SB TL4SB - TIJ11B "nJ2IB "nJ31B nj4iB TM11B —-"nVI2IB TM3IB -~ TM4IB -—TL1IB ~-TL2IB TL31B ~-"n.4iB

Figure G-19. Insulated (3.81 cm SFRM) steel truss B in Test 5.

G.2.2 Insulated Steel Components Exposed to a 3 MW Fire for 50 Min in Test 6

0 600 1200 1800 2400 3000 3600 4200 4800 5400 6000

Time (sec) TB1S TB2S TB3S —- TB4S TB1I —TB2I —TB3I TB4I

Figure G-20. Insulated (1.91 cm SFRM) steel bar in Test 6.

NISTNCSTAR 1-5B, WTC Investigation 301
Appendix G

1000

900

800

700 — 600

E

500

400

300

200

100

0 600 1200 1800 2400 3000 3600

Time (sec)

4200 4800 5400 6000

TCU1S ~ TCU2S TCU3S —TCU4S TCM1S -—TCM2S TCM3S -—TCM4S TCL1S TCL2S

TCL3S TCL4S TCU1I TCU2I TCU3I TCU4I TCM1I TCM2I TCM3I - TCLII

•-- TCL2I - TCL3I - TCL4I

Figure G-21. Insulated (1.91 cm SFRM) steel column in Test 6.

0 600 1200 1800 2400 3000 3600 Time (sec)

4200 4800 5400 6000

TUISA TU2SA TU3SA - ~"nj4SA TM1SA TM2SA T1VI3SA TM4SA TL1SA "n.2SA TL3SA TL4SA TU1IA TU2IA TU3IA TU4IA T1VI1IA _,TM2IA - ...71VI3IA TM4IA U1IA .TL3IA —-TL4W

Figure G-22. Insulated (1.91 cm SFRM) steel truss A in Test 6.

302 NISTNCSTAR 1-5B, WTC Investigation
Time-Temperature Histories of Bare and Coated Steel Components

1000 -r

0 600 1200 1800 2400 3000 3600 4200 4800 5400 6000

TifTB (sec)

TU1SB 1U2SB TUaSB _TU4SB TMISB TM2SB TMSSB 1M4SB —TLISB

U2SB TLSSB TL4SB TU1IB TU2IB TU3IB TIMIB "IM1IB TIV12IB

TM3IB imB TLIIB TL2IB 1L3B

Figure G-23. Insulated (1.91 cm SFRM) steel truss B in Test 6.

NISTNCSTAR 1-5B, WTC Investigation 303
Appendix G

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304 NISTNCSTAR 1-5B, WTC Investigation